DYNAMICS AND CONTROL OF MULTIBODY TETHERED SYSTEMS …
Transcript of DYNAMICS AND CONTROL OF MULTIBODY TETHERED SYSTEMS …
D Y N A M I C S A N D C O N T R O L OF M U L T I B O D Y T E T H E R E D SYSTEMS USING A N O R D E R - N F O R M U L A T I O N
S P I R O S K A L A N T Z I S
B Eng (Honours) McGill University Montreal Canada 1994
A T H E S I S S U B M I T T E D I N P A R T I A L F U L F I L M E N T O F T H E R E Q U I R E M E N T S F O R T H E D E G R E E O F
M A S T E R O F A P P L I E D S C I E N C E
in
The Faculty of Graduate Studies Department of Mechanical Engineering
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ABSTRACT
The equations of motion for a multibody tethered satellite system in three dishy
mensional Keplerian orbit are derived The model considers a multi-satellite system
connected in series by flexible tethers Both tethers and subsatellites are free to unshy
dergo three dimensional attitude motion together with deployment and retrieval as
well as longitudinal and transverse vibration for the tether The elastic deformations
of the tethers are discretized using the assumed mode method The tether attachment
points to the subsatellites are kept arbitrary and time varying The model is also cashy
pable of simulating the response of the entire system spinning about an arbitrary
axis as in the case of OEDIPUS-AC which spins about the nominal tether length
or the proposed BICEPS mission where the system cartwheels about the orbit norshy
mal The governing equations of motion are derived using a non-recursive order(N)
Lagrangian procedure which significantly reduces the computational cost associated
with the inversion of the mass matrix an important consideration for multi-satellite
systems Also a symbolic integration and coding package is used to evaluate modal
integrals thus avoiding their costly on-line numerical evaluation
Next versatility of the formulation is illustrated through its application to
two different tethered satellite systems of contemporary interest Finally a thruster
and momentum-wheel based attitude controller is developed using the Feedback Linshy
earization Technique in conjunction with an offset (tether attachment point) control
strategy for the suppression of the tethers vibratory motion using the optimal Linshy
ear Quadratic Gaussian-Loop Transfer Recovery method Both the controllers are
successful in stabilizing the system over a range of mission profiles
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T A B L E OF C O N T E N T S
A B S T R A C T bull i i
T A B L E O F C O N T E N T S i i i
L I S T O F S Y M B O L S v i i
L I S T O F F I G U R E S x i
A C K N O W L E D G E M E N T xv
1 I N T R O D U C T I O N 1
11 Prel iminary Remarks 1
12 Brief Review of the Relevant Literature 7
121 Mul t ibody O(N) formulation 7
122 Issues of tether modelling 10
123 Att i tude and vibration control 11
13 Scope of the Investigation 12
2 F O R M U L A T I O N O F T H E P R O B L E M 14
21 Kinematics 14
211 Prel iminary definitions and the itfl position vector 14
212 Tether flexibility discretization 14
213 Rotat ion angles and transformations 17
214 Inertial velocity of the ith link 18
215 Cyl indr ica l orbital coordinates 19
216 Tether deployment and retrieval profile 21
i i i
22 Kinetics and System Energy 22
221 Kinet ic energy 22
222 Simplification for rigid links 23
223 Gravitat ional potential energy 25
224 Strain energy 26
225 Tether energy dissipation 27
23 O(N) Form of the Equations of Mot ion 28
231 Lagrange equations of motion 28
232 Generalized coordinates and position transformation 29
233 Velocity transformations 31
234 Cyl indrical coordinate modification 33
235 Mass matrix inversion 34
236 Specification of the offset position 35
24 Generalized Control Forces 36
241 Preliminary remarks 36
242 Generalized thruster forces 37
243 Generalized momentum gyro torques 39
3 C O M P U T E R I M P L E M E N T A T I O N 42
31 Preliminary Remarks 42
32 Numerical Implementation 43
321 Integration routine 43
322 Program structure 43
33 Verification of the Code 46
iv
331 Energy conservation 4 6
332 Comparison with available data 50
4 D Y N A M I C S I M U L A T I O N 53
41 Prel iminary Remarks 53
42 Parameter and Response Variable Definitions 53
43 Stationkeeping Profile 56
44 Tether Deployment 66
45 Tether Retrieval 71
46 Five-Body Tethered System 71
47 B I C E P S Configuration 80
48 O E D I P U S Spinning Configuration 88
5 A T T I T U D E A N D V I B R A T I O N C O N T R O L 93
51 Att i tude Control 93
511 Prel iminary remarks 93
512 Controller design using Feedback Linearization Technique 94
513 Simulation results 96
52 Control of Tethers Elastic Vibrations 110
521 Prel iminary remarks 110
522 System linearization and state-space realization 1 1 0
523 Linear Quadratic Gaussian control
with Loop Transfer Recovery 114
524 Simulation results 121
v
6 C O N C L U D I N G R E M A R K S 1 2 4
61 Summary of Results 124
62 Recommendations for Future Study 127
B I B L I O G R A P H Y 128
A P P E N D I C E S
I T E N S O R R E P R E S E N T A T I O N O F T H E E Q U A T I O N S O F M O T I O N 133
11 Prel iminary Remarks 133
12 Mathematical Background 133
13 Forcing Function 135
II R E D U C E D E Q U A T I O N S O F M O T I O N 137
II 1 Preliminary Remarks 137
II2 Derivation of the Lagrangian Equations of Mot ion 137
v i
LIST OF SYMBOLS
A j intermediate length over which the deploymentretrieval profile
for the itfl l ink is sinusoidal
A Lagrange multipliers
EUQU longitudinal state and measurement noise covariance matrices
respectively
matrix containing mode shape functions of the t h flexible link
pitch roll and yaw angles of the i 1 link
time-varying modal coordinate for the i 1 flexible link
link strain and stress respectively
damping factor of the t h attitude actuator
set of attitude angles (fjj = ci
structural damping coefficient for the t h l ink EjJEj
true anomaly
Earths gravitational constant GMejBe
density of the ith link
fundamental frequency of the link longitudinal tether vibrashy
tion
A tethers cross-sectional area
AfBfCjDf state-space representation of flexible subsystem
Dj inertial position vector of frame Fj
Dj magnitude of Dj mdash mdash
Df) transformation matrix relating Dj and Ds
D r i inplane radial distance of the first l ink
DSi Drv6i DZ1 T for i = 1 and DX Dy poundgt 2 T for 1
m
Pi
w 0 i
v i i
Dx- transformation matrix relating D j and Ds-
DXi Dyi Dz- Cartesian components of D
DZl out-of-plane position component of first link
E Youngs elastic modulus for the ith l ink
Ej^ contribution from structural damping to the augmented complex
modulus E
F systems conservative force vector
FQ inertial reference frame
Fi t h link body-fixed reference frame
n x n identity matrix
alternate form of inertia matrix for the i 1 l ink
K A Z $i(k + dXi+1)
Kei t h link kinetic energy
M(q t) systems coupled mass matrix
Mma Mma Mmy- control moments in the pitch roll and yaw directions respec-
tively for the t h link
Mr fr rigid mass matrix and force vector respectively
Mred fred- Qred mass matrix force vector and generalized coordinate vector for
the reduced model respectively
Mt block diagonal decoupled mass matrix
symmetric decoupled mass matrix
N total number of links
0N) order-N
Pi(9i) column matrix [T^giTpg^T^gi) mdash
Q nonconservative generalized force vector
Qu actuator coupling matrix or longitudinal L Q R state weighting
matrix
v i i i
Rdm- inertial position of the t h link mass element drrii
RP transformation matrix relating qt and tf
Ru longitudinal L Q R input weighting matrix
RvRn Rd transformation matrices relating qt and q
S(f Mdc) generalized acceleration vector of coupled system q for the full
nonlinear flexible system
th uiirv l u w u u u maniA
th
T j i l ink rotation matrix
TtaTto control thrust in the pitch and roll direction for the im l ink
respectively
rQi maximum deploymentretrieval velocity of the t h l ink
Vei l ink strain energy
Vgi link gravitational potential energy
Rayleigh dissipation function arising from structural damping
in the ith link
di position vector to the frame F from the frame F_i bull
dc desired offset acceleration vector (i 1 l ink offset position)
drrii infinitesimal mass element of the t h l ink
dx^dy^dZi Cartesian components of d along the local vertical local horishy
zontal and orbit normal directions respectively
F-Q mdash
gi r igid and flexible position vectors of drrii fj + ltEraquolt5
ijk unit vectors 1 0 0 r 010-^ and 00 l r respectively
li length of the t h l ink
raj mass of the t h link
nfj total number of flexible modes for the i 1 l ink nXi + nyi + nZi
nq total number of generalized coordinates per link nfj + 7
nqq systems total number of generalized coordinates Nnq
ix
number of flexible modes in the longitudinal inplane and out-
of-plane transverse directions respectively for the i ^ link
number of attitude control actuators
- $ T
set of generalized coordinates for the itfl l ink which accounts for
interactions with adjacent links
qv---QtNT
set of coordinates for the independent t h link (not connected to
adjacent links)
rigid position of dm in the frame Fj
position of centre of mass of the itfl l ink relative to the frame Fj
i + $DJi
desired settling time of the j 1 attitude actuator
actuator force vector for entire system
flexible deformation of the link along the Xj yi and Zj direcshy
tions respectively
control input for flexible subsystem
Cartesian components of fj
actual and estimated state of flexible subsystem respectively
output vector of flexible subsystem
LIST OF FIGURES
1-1 A schematic diagram of the space platform based N-body tethered
satellite system 2
1-2 Associated forces for the dumbbell satellite configuration 3
1- 3 Some applications of the tethered satellite systems (a) multiple
communication satellites at a fixed geostationary location
(b) retrieval maneuver (c) proposed B I C E P S configuration 6
2- 1 Vector components of the i 1 and (i- l ) 1 chain links 15
2-2 Vector components in cylindrical orbital coordinates 20
2-3 Inertial position of subsatellite thruster forces 38
2- 4 Coupled force representation of a momentum-wheel on a rigid body 40
3- 1 Flowchart showing the computer program structure 45
3-2 Kinet ic and potentialenergy transfer for the three-body platform based
tethered satellite system (a) variation of kinetic and potential energy
(b) percent change in total energy of system 48
3-3 Kinet ic and potential energy transfer for the five-body platform based
tethered satellite system (a) variation of kinetic and potential energy
(b) percent change in total energy of system 49
3-4 Simulation results for the platform based three-body tethered
system originally presented in Ref[43] 51
3- 5 Simulation results for the platform based three-body tethered
system obtained using the present computer program 52
4- 1 Schematic diagram showing the generalized coordinates used to
describe the system dynamics 55
4-2 Stationkeeping dynamics of the three-body STSS configuration
without offset (a) attitude response (b) vibration response 57
4-3 Stationkeeping dynamics of the three-body STSS configuration
xi
with offset along the local vertical
(a) attitude response (b) vibration response 60
4-4 Stationkeeping dynamics of the three-body STSS configuration
with offset along the local horizontal
(a) attitude response (b) vibration response 62
4-5 Stationkeeping dynamics of the three-body STSS configuration
with offset along the orbit normal
(a) attitude response (b) vibration response 64
4-6 Stationkeeping dynamics of the three-body STSS configuration
wi th offset along the local horizontal local vertical and orbit normal
(a) attitude response (b) vibration response 67
4-7 Deployment dynamics of the three-body STSS configuration
without offset
(a) attitude response (b) vibration response 69
4-8 Deployment dynamics of the three-body STSS configuration
with offset along the local vertical
(a) attitude response (b) vibration response 72
4-9 Retrieval dynamics of the three-body STSS configuration
with offset along the local vertical and orbit normal
(a) attitude response (b) vibration response 74
4-10 Schematic diagram of the five-body system used in the numerical example 77
4-11 Stationkeeping dynamics of the five-body STSS configuration
without offset
(a) attitude response (b) vibration response 78
4-12 Deployment dynamics of the five-body STSS configuration
wi th offset along the local vertical
(a) attitude response (b) vibration response 81
4-13 Stationkeeping dynamics of the three-body B I C E P S configuration
x i i
with offset along the local vertical
(a) attitude response (b) vibration response 84
4-14 Cartwheeling dynamics with deployment for the three-body B I C E P S
with offset along the local vertical
(a) attitude response (b) vibration response 86
4-15 Spin dynamics (7 = l deg s ) of the three-body O E D I P U S configuration
with offset along the local vertical
(a) attitude response (b) vibration response 89
4- 16 Spin dynamics (7 = 10deg s ) of the three-body O E D I P U S configuration
wi th offset along the local vertical
(a) attitude response (b) vibration response 91
5- 1 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear rigid F L T controller with offset along the
local horizontal local vertical and orbit normal
(a) attitude and vibration response (b) control actuator time histories 98
5-2 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear flexible F L T controller with offset along the
local horizontal local vertical and orbit normal
(a) attitude and vibration response (b) control actuator time histories 101
5-3 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear rigid F L T controller with offset along
the local vertical
(a) attitude and vibration response (b) control actuator time histories 103
5-4 Deployment dynamics of the three-body STSS using the nonlinear
rigid F L T controller with offset along the local vertical
(a) attitude and vibration response (b) control actuator time histories 105
5-5 Retrieval dynamics of the three-body STSS using the non-linear
rigid F L T controller with offset along the local vertical and orbit normal
(a) attitude and vibration response (b) control actuator time histories 108
x i i i
5-6 Block diagram for the L Q G L T R estimator based compensator 115
5-7 Singular values for the L Q G and L Q G L T R compensator compared to target
return ratio (a) longitudinal design (b) transverse design 118
5-8 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear rigid F L T attitude controller and L Q G L T R
offset vibration controller (a) attitude and libration controller response
(b) vibration and offset response 122
xiv
A C K N O W L E D G E M E N T
Firs t and foremost I would like to express the most genuine gratitude to
my supervisor Prof V i n o d J M o d i whose unending guidance and encouragement
proved most invaluable during the course of my studies
I am also deeply indebted to Dr Satyabrata Pradhan for his limitless patience
and technical advice from which this work would not have been possible I would also
like to extend a word of appreciation towards Prof A r u n K Mis ra ( M c G i l l University)
for ini t iat ing my interest in space dynamics and control
Also I would like to thank my collegues Dr Sandeep Munshi M r Mathieu
Caron and M r Yuan Chen as well as Dr Mae Seto M r Gary L i m and M r Mark
Chu for their words of advice and helpful suggestions that heavily contributed to all
aspects of my studies
Final ly I would like to thank all my family and friends here in Vancouver and
back home in Montreal whose unending support made the past two years the most
memorable of my life
The research project was supported by the Natural Sciences and Engineering
Research Council s P G S - A Scholarship held by the author as well as by N S E R C grant
A-2181 held by Prof V J M o d i
xv
1 INTRODUCTION
11 Prel iminary Remarks
W i t h the ever changing demands of the worlds population one often wonders
about the commitment to the space program On the other hand the importance
of worldwide communications global environmental monitoring as well as long-term
habitation in space have demanded more ambitious and versatile satellites systems
It is doubtful that the Russian scientist Tsiolkovsky[l] when he first proposed the
uti l izat ion of the Earths gravity-gradient environment in the last century would have
envisioned the current and proposed applications of tethered satellite systems
A tethered satellite system consists of two or more subsatellites connected
to each other by long thin cables or tethers which are in tension (Figure 1-1) The
subsatellite which can also include the shuttle and space station are in orbit together
The first proposed use of tethers in space was associated with the rescue of stranded
astronauts by throwing a buoy on a tether and reeling it to the rescue vehicle
Prel iminary studies of such systems led to the discovery of the inherent instability
during the tether retrieval[2] Nevertheless the bir th of tethered systems occured
during the Gemini X I and X I I missions[3] in 1966 when a short tether (10-30m)
was used to generate the first artificial gravity environment in space (000015g) by
cartwheeling (spinning) the system about its center of mass
The mission also demonstrated an important use of tethers for gravity gradient
stabilization The force analysis of a simplified model illustrates this point Consider
the dumbbell satellite configuration orbiting about the centre of Ear th as shown in
Figure 1-2 Satellite 1 and satellite 2 are connected to each other by a tether with
1
Space Station (Satellite 1 )
2
Satellite N
Figure 1-1 A schematic diagram of the space platform based N-body tethered satellite system
2
Figure 1-2 Associated forces for the dumbbell satellite configuration
3
the whole system free to librate about the systems centre of mass C M The two
major forces acting on each body are due to the centrifugal and gravitational effects
Since body 1 is further away from the earth its gravitational force Fg^y is smaller
than that of body 2 ie Fgi lt Fg^ However its centrifugal force FCl is greater
then the centrifugal force experienced by the second satellite (FCl gt FC2) Moreover
for body 1 Fg^ lt Fc^ in contrast to Fg2 gt FC2 for body 2 Thus resolving the force
components along the tether there is an evident resultant tension force Ff present
in the tether Similarly adding the normal components of the two forces vectorially
results in a force FR which restores the system to its stable equil ibrium along the
local vertical These tension and gravity-gradient restoring forces are the two most
important features of tethered systems
Several milestone missions have flown in the recent past The U S A J a p a n
project T P E (Tethered Payload Experiment) [4] was one of the first to be launched
using a sounding rocket to conduct environmental studies The results of the T P E
provided support for the N A S A I t a l y shuttle based tethered satellite system referred
to as the TSS-1 mission[5] This experiment aimed to study the electrodynamic effects
of a shuttle borne satellite system with a conductive tether connection where an
electric current was induced in the tether as it swept through the Earth s magnetic
field Unfortunetely the two attempts in August 1992 and February 1996 resulted
in only partial success The former suffered from a spool failure resulting in a tether
deployment of only 256m of a planned 20km During the latter attempt a break
in the tethers insulation resulted in arcing leading to tether rupture Nevertheless
the information gathered by these attempts st i l l provided the engineers and scientists
invaluable information on the dynamic behaviour of this complex system
Canada also paved the way with novel designs of tethered system aimed
primari ly at the study of the ionosphere and Northern Lights (Aurora Borealis)
4
The O E D I P U S A and C (Observation of Electrified Distributions in the Ionospheric
P lasma - a Unique Strategy) missions[6] launched from a sounding rocket in January
1989 and November 1995 respectively provided insight into the complex dynamical
behaviour of two comparable mass satellites connected by a 1 km long spinning
tether
Final ly two experiments were conducted by N A S A called S E D S I and II
(Small Expendable Deployable System) [7] which hold the current record of 20 km
long tether In each of these missions a small probe (26 kg) was deployed from the
second stage of a D E L T A II rocket thus succesfully demonstrating the feasibility of
long tether deployment Retrieval of the tether which is significantly more difficult
has not yet been achieved
Several proposed applications are also currently under study These include the
study of earths upper atmosphere using probes lowered from the shuttle in a low earth
orbit ( L E O ) to the deployment and retrieval of satellites for servicing Even more
promising application concerns the generation of power for the proposed International
Space Station using conductive tethers The Canadian Space Agency (CSA) has also
proposed a dual-mass satellite system B I C E P S (Bl-static Canadian Experiment on
Plasmas in Space) to make simultaneous measurements at different locations in the
environment for correlation studies[8] A unique feature of the B I C E P S mission is
the deployment of the tether aided by the angular momentum of the system which
is ini t ia l ly cartwheeling about its orbit normal (Figure 1-3)
In addition to the three-body configuration multi-tethered systems have been
proposed for monitoring Earths environment in the global project monitored by
N A S A entitled Mission to Planet Ear th ( M T P E ) The proposed mission aims to
study pollution control through the understanding of the dynamic interactions be-
5
Retrieval of Satellite for Servicing
lt mdash
lt Satellite
Figure 1-3 Some applications of the tethered satellite systems (a) multiple comshymunication satellites at a fixed geostationary location (b) retrieval maneuver (c) proposed BICEPS configuration
6
tween the Earths atmosphere biosphere hydrosphere and cryosphere This wi l l be
accomplished with multiple tethered instrumentation payloads simultaneously soundshy
ing at different altitudes for spatial correlations Such mult ibody systems are considshy
ered to be the next stage in the evolution of tethered satellites Micro-gravity payload
modules suspended from the proposed space station for long-term experiments as well
as communications satellites with increased line-of-sight capability represent just two
of the numerous possible applications under consideration
12 Brief Review of the Relevant Literature
The design of multi-payload systems wi l l require extensive dynamic analysis
and parametric study before the final mission configuration can be chosen This
wi l l include a review of the fundamental dynamics of the simple two-body tethered
systems and how its dynamics can be extended to those of multi-body tethered system
in a chain or more generally a tree topology
121 Multibody 0(N) formulation
Many studies of two-body systems have been reported One of the earliest
contribution is by Rupp[9] who was the first to study planar dynamics of the simshy
plified Shuttle based Tethered Satellite System (STSS) His pioneering work led to
the discovery of pitch oscillation growth during the retrieval phase Later Baker[10]
advanced the investigation to the third dimension in addition to adding atmospheric
effects to the system A more complete dynamical model was later developed and
analyzed by M o d i and Misra[ l l 12] which included deployment and tether flexibility
A comprehensive survey of important developments and milestones in the area have
been compiled and reviewed by Mis ra and Modi[13] The major conclusions based on
the literature may be summarized as follows
bull the stationkeeping phase is stable
7
bull deployment can be unstable if the rate exceeds a crit ical speed
bull retrieval of the tether is inherently unstable
bull transverse vibrations can grow due to the Coriolis force induced during deshy
ployment and retrieval
Misra Amier and Modi[14] were one of the first to extend these results to the
three-satellite double pendulum case This simple model which includes a variable
length tether was sufficient to uncover the multiple equilibrium configurations of
the system Later the work was extended to include out-of-plane motion and a
preliminary reel-rate control law to regulate the librational motion[15] Kesmir i and
Misra[16] developed a general formulation for N-body tethered systems based on the
Lagrangian principle where three-dimensional motion and flexibility are accounted for
However from their work it is clear that as the number of payload or bodies increases
the computational cost to solve the forward dynamics also increases dramatically
Tradit ional methods of inverting the mass matrix M in the equation
ML+F = Q
to solve for the acceleration vector proved to be computationally costly being of the
order for practical simulation algorithms O(N^) refers to a mult ibody formulashy
tion algorithm whose computational cost is proportional to the cube of the number
of bodies N used It is therefore clear that more efficient solution strategies have to
be considered
Many improved algorithms have been proposed to reduce the number of comshy
putational steps required for solving for the acceleration vector Rosenthal[17] proshy
posed a recursive formulation based on the triangularization of the equations of moshy
tion to obtain an 0(N2) formulation He also demonstrates the use of a symbolic
manipulation program to reduce the final number of computations A recursive L a -
8
grangian formulation proposed by Book[18] also points out the relative inefficiency of
conventional formulations and proceeds to derive an algorithm which is also 0(N2)
A recursive formulation is one where the equations of motion are calculated in order
from one end of the system to the other It usually involves one or more passes along
the links to calculate the acceleration of each link (forward pass) and the constraint
forces (backward or inverse pass) Non-recursive algorithms express the equations of
motion for each link independent of the other Although the coupling terms st i l l need
to be included the algorithm is in general more amenable to parallel computing
A st i l l more efficient dynamics formulation is an O(N) algorithm Here the
computational effort is directly proportional to the number of bodies or links used
Several types of such algorithms have been developed over the years The one based
on the Newton-Euler approach has recursive equations and is used extensively in the
field of multi-joint robotics[1920] It should be emphasized that efficient computation
of the forward and inverse dynamics is imperative if any on-line (real-time) control
of the joint is to be achieved Hollerbach[21] proposed a recursive O(N) formulation
based on Lagranges equations of motion However his derivation was primari ly foshy
cused on the inverse dynamics and did not improve the computational efficiency of
the forward dynamics (acceleration vector) Other recursive algorithms include methshy
ods based on the principle of vir tual work[22] and Kanes equations of motion[23]
Keat[24] proposed a method based on a velocity transformation that eliminated the
appearance of constraint forces and was recursive in nature O n the other hand
K u r d i l a Menon and Sunkel[25] proposed a non-recursive algorithm based on the
Range-Space method which employs element-by-element approach used in modern
finite-element procedures Authors also demonstrated the potential for parallel comshy
putation of their non-recursive formulation The introduction of a Spatial Operator
Factorization[262728] which utilizes an analogy between mult ibody robot dynamics
and linear filtering and smoothing theory to efficiently invert the systems mass ma-
9
t r ix is another approach to a recursive algorithm Their results were further extended
to include the case where joints follow user-specified profiles[29] More recently Prad-
han et al [30] proposed a non-recursive Lagrangian algorithm for factorizing the mass
matrix leading to an O(N) formulation of the forward dynamics of the system A n
efficient O(N) algorithm where the number of bodies N in the system varies on-line
has been developed by Banerjee[31]
122 Issues of tether modelling
The importance of including tether flexibility has been demonstrated by several
researchers in the past However there are several methods available for modelling
tether flexibility Each of these methods has relative advantages and limitations
wi th regards to the computational efficiency A continuum model where flexible
motion is discretized using the assumed mode method has been proposed[3233] and
succesfully implemented In the majority of cases even one flexible mode is sufficient
to capture the dynamical behaviour of the tether[34] K i m and Vadali[35] proposed a
so-called bead model or lumped-mass approach that discretizes the tether using point
masses along its length However in order to accurately portray the motion of the
tether a significantly high number of beads are needed thus increasing the number of
computation steps Consequently this has led to the development of a hybrid between
the last two approaches[16] ie interpolation between the beads using a continuum
approach thus requiring less number of beads Finally the tether flexibility can also
be modelled using a finite-element approach[36] which is more suitable for transient
response analysis In the present study the continuum approach is adopted due to
its simplicity and proven reliability in accurately conveying the vibratory response
In the past the modal integrals arising from the discretization process have
been evaluted numerically In general this leads to unnecessary effort by the comshy
puter Today these integrals can be evaluated analytically using a symbolic in-
10
tegration package eg Maple V Subsequently they can be coded in F O R T R A N
directly by the package which could also result in a significant reduction in debugging
time More importantly there is considerable improvement in the computational effishy
ciency [16] especially during deployment and retrieval where the modal integral must
be evaluated at each time-step
123 Attitude and vibration control
In view of the conclusions arrived at by some of the researchers mentioned
above it is clear that an appropriate control strategy is needed to regulate the dyshy
namics of the system ie attitude and tether vibration First the attitude motion of
the entire system must be controlled This of course would be essential in the case of
a satellite system intended for scientific experiments such as the micro-gravity facilishy
ties aboard the proposed Internation Space Station Vibra t ion of the flexible members
wi l l have to be checked if they affect the integrity of the on-board instrumentation
Thruster as well as momentum-wheel approach[34] have been considered to
regulate the rigid-body motion of the end-satellite as well as the swinging motion
of the tether The procedure is particularly attractive due to its effectiveness over a
wide range of tether lengths as well as ease of implementation Other methods inshy
clude tension and length rate control which regulate the tethers tension and nominal
unstretched length respectively[915] It is usually implemented at the deployment
spool of the tether More recently an offset strategy involving time dependent moshy
tion of the tether attachment point to the platform has been proposed[3738] It
overcomes the problem of plume impingement created by the thruster control and
the ineffectiveness of tension control at shorter lengths However the effectiveness of
such a controller can become limited with an exceedingly long tether due to a pracshy
tical l imi t on the permissible offset motion In response to these two control issues a
hybrid thrusteroffset scheme has been proposed to combine the best features of the
11
two methods[39]
In addition to attitude control these schemes can be used to attenuate the
flexible response of the tether Tension and length rate control[12] as well as thruster
based algorithms[3340] have been proposed to this end M o d i et al [41] have sucshy
cessfully demonstrated the effectiveness of an offset strategy to damp undesired v i shy
brations Passive energy dissipative devices eg viscous dampers are also another
viable solution to the problem
The development of various control laws to implement the above mentioned
strategies has also recieved much attention A n eigenstructure assignment in conducshy
tion wi th an offset controller for vibration attenuation and momemtum wheels for
platform libration control has been developed[42] in addition to the controller design
from a graph theoretic approach[41] Also non-linear feedback methods such as the
Feedback Linearization Technique ( F L T ) that are more suitable for controlling highly
nonlinear non-autonomous coupled systems have also been considered[43]
It is important to point out that several linear controllers including the classhy
sic state feedback Linear Quadratic Regulator ( L Q R ) have received considerable
attention[3944] Moreover robust methods such as the Linear Quadratic Guas-
s ian Loop Transfer Recovery ( L Q G L T R ) method have also been developed and imshy
plemented on tethered systems[43] A more complete review of these algorithms as
well as others applied to tethered system has been presented by Pradhan[43]
13 Scope of the Investigation
The objective of the thesis is to develop and implement a versatile as well as
computationally efficient formulation algorithm applicable to a large class of tethered
satellite systems The distinctive features of the model can be summarized as follows
12
bull TV-body 0(N) tethered satellite system in a chain-type topology
bull the system is free to negotiate a general Keplerian orbit and permitted to
undergo three dimensional inplane and out-of-plane l ibrtational motion
bull r igid bodies constituting the system are free to execute general rotational
motion
bull three dimensional flexibility present in the tether which is discretized using
the continuum assumed-mode method with an arbitrary number of flexible
modes
bull energy dissipation through structural damping is included
bull capability to model various mission configurations and maneuvers including
the tether spin about an arbitrary axis
bull user-defined time dependent deployment and retrieval profiles for the tether
as well as the tether attachment point (offset)
bull attitude control algorithm for the tether and rigid bodies using thrusters and
momentum-wheels based on the Feedback Linearization Technique
bull the algorithm for suppression of tether vibration using an active offset (tether
attachment point) control strategy based on the optimal linear control law
( L Q G L T - R )
To begin with in Chapter 2 kinematics and kinetics of the system are derived
using the O(N) formulation methodology developed by Pradhan et al [30] Chapshy
ter 3 discusses issues related to the development of the simulation program and its
validation This is followed by a detailed parametric study of several mission proshy
files simulated as particular cases of the versatile formulation in Chapter 4 Next
the attitude and vibration controllers are designed and their effectiveness assessed in
Chapter 5 The thesis ends with concluding remarks and suggestions for future work
13
2 F O R M U L A T I O N OF T H E P R O B L E M
21 Kinematics
211 Preliminary definitions and the i ^ position vector
The derivation of the equations of motion begins with the definition of the
inertial position of an arbitrary link of the mult ibody system Let the i ^ link of the
system be free to translate and rotate in 3-D space From Figure 2-1 the position
vector Rdm^ to the mass element drrn on the i ^ link wi th reference to the inertial
frame FQ can be written as
Here D represents the inertial position of the t h body fixed frame Fj relative to FQ
ri = [xiyiZi]T is the rigid position vector of drrii wi th respect to F J and f(fi) is
the flexible deformation at fj also with respect to the frame Fj Bo th these vectors
are relative to Fj Note in the present model tethers (i even) are considered flexible
and X corresponds to the nominal position along the unstretched tether while a and
ZJ are by definition equal to zero On the other hand the rigid bodies referred to as
satellites (i odd) including the platform are taken to be rigid ie f(fj) = 0 T j is
the rotation matrix used to express body-fixed vectors with reference to the inertial
frame
212 Tether flexibility discretization
The tether flexibility is discretized with an arbitrary but finite number of
Di + Ti(i + f(i)) (21)
14
15
modes in each direction using the assumed-mode method as
Ui nvi
wi I I i = 1
nzi bull
(22)
where nXi nyi and n 2 ^ are the number of modes used in the X m and Z directions
respectively For the ] t h longitudinal mode the admissible mode shape function is
taken as 2 j - l
Hi(xiii)=[j) (23)
where l is the i ^ tether length[32] In the case of inplane and out-of-plane transverse
deflections the admissible functions are
ampyi(xili) = ampZixuk) = v ^ s i n JKXj
k (24)
where y2 is added as a normalizing factor In this case both the longitudinal and
transverse mode shapes satisfy the geometric boundary conditions for a simple string
in axial and transverse vibration
Eq(22) is recast into a compact matrix form as
ff(i) = $i(xili)5i(t) (25)
where $i(x l) is the matrix of the tether mode shapes defined as
0 0
^i(xiJi) =
^ X bull bull bull ^XA
0
0 0
0
G 5R 3 x n f
(26)
and nfj = nx- + nyi + nZi Note $J(XJJ) is a function of the spatial coordinate x
and not of yi or ZJ However it is also a function of the tether length parameter li
16
which is time-varying during deployment and retrieval For this reason care must be
exercised when differentiating $J(XJJ) with respect to time such that
d d d k) = Jtregi(Xi li) = ^ 7 $ t ( ^ i k)Xi + gf$i(xigt li)k- (2-7)
Since x represents the nominal rate of change of the position of the elemental mass
drrii it is equal to j (deploymentretrieval rate) Let t ing
( 3 d
dx~ + dT)^Xili^ ( 2 8 )
one arrives at
$i(xili) = $Di(xili)ii (29)
The time varying generalized coordinate vector 5 in Eq(25) is composed of
the longitudinal and transverse components ie
ltM) G s f t n f i x l (210)
where 5X^ 5y- and Sz^ are nx- x 1 ny x 1 and nz- x 1 size vectors and correspond to
$ X and $ Z respectively
213 Rotation angles and transformations
The matrix T j in Eq(21) represents a rotation transformation from the body
fixed frame Fj to the inertial frame FQ ie FQ = T j F j It is defined by the 3-2-1
ordered sequence of elementary rotations[46]
1 P i tch F[ = Cf ( a j )F 0
2 R o l l F = C f (3j)F = C7f (3j)Cf ( a j )F 0
3 Yaw F = C7( 7 i )F = C7(7l)cf^)Cf(ai)FQ = Q F 0
17
where
c f (A) =
-Sai Cai
0 0
and
C(7i ) =
0 -s0i] = 0 1 0
0 cpi
i 0 0 = 0
0
(211)
(212)
(213)
Here Eq(211) corresponds to a rotation of ctj (pitch) about the inertial axis ZQ (orbit
normal) of frame FQ resulting in a new intermediate frame F It is then followed by
a roll rotation fa about the axis y of F- giving a second intermediate frame F-
Final ly a spin rotation 7J about the axis z of F- is given resulting in the body
fixed frame Fj for the i lt l link Since all the three rotation matrices are orthogonal
it follows that FQ = C~1Fi = Cj F and hence
C CPi Cai ~ CH Sai + SrYi SPi Sai 5 7 i Sampi + Cli SPi Cci CPi Sai C1i Cai + 5 7 i SPi ~ S1 Ca + C 7 bull Sp Sai
(214) SliCPi CliCPi
Here Cx and Sx are abbreviations for cos(x) and sin(x) respectively
214 Inertial velocity of the ith link
mdash Lett ing pj = fj - f $j(5j and differentiating Eq(21) with respect to time gives
m- D j + T j f j + T^Si + T j pound + T j $ j 5 j (215)
Each of the terms appearing in Eq(215) must be addressed individually Since
fj represents the rigid body motion within the link it only has meaning for the
deployable tether and since yi = Zj = 0 for al l tether links (i even) fj = Ifi where 1
is the unit vector 1 0 0T For the case of rigid links (i odd) fj = 0
18
Evaluating the derivative of each angle in the rotation matrix gives
T^i9i = Taiai9i + Tppigi + T^fagi (216)
where Tx = J j I V Collecting the scaler angular velocity terms and defining the set
Vi = deg^gt A) 7i^ Eq(216) can be rewritten as
T i f t = Pi9i)fji (217)
where Pi(gi) is a column matrix defined by
Pi(9i)Vi = [Taigi^p^iT^gilffi (218)
Inserting Eq(29) and Eq(217) into Eq(215) and defining Sj = 2 + leads to
215 Cylindrical orbital coordinates
The Di term in Eq(219) describes the orbital motion of the t h l ink and
is composed of the three Cartesian coordinates DXi A ^ DZi However over the
cycle of one orbit DXi and Dyi vary dramatically by around the order of Earths
radius This must be avoided since large variations in the coordinates can cause
severe truncation errors during their numerical integration For this reason it is
more convenient to express the Cartesian components in terms of more stationary
variables This is readily accomplished using cylindrical coordinates
However it is only necessary to transform the first l inks orbital components
ie Lgti since only D is a generalized coordinates From Figure 2-2 it is apparent
(219)
that
(220)
19
Figure 2-2 Vector components in cylindrical orbital coordinates
20
Eq(220) can be rewritten in matrix form as
cos(0i) 0 0 s i n ( 0 i ) 0 0
0 0 1
Dri
h (221)
= DTlDSv
where Dri is the inplane radial distance 9 the true anomaly and DZl the out-of-
plane distance normal to the orbital plane The total derivative of D wi th respect
to time gives in column matrix form
1 1 1 1 J (222)
= D D l D 3 v
where [cos(^i) -Dnsm(8i) 0 ]
(223)
For all remaining links ie i = 2 N
cos(0i) - A - j S i n ^ i ) 0 sin(^i) D r i c o s ( 0 i ) 0
0 0 1
DTl = D D i = I d 3 i = 2N (224)
where is the n x n identity matrix and
2N (225)
216 Tether deployment and retrieval profile
The deployment of the tether is a critical maneuver for this class of systems
Cr i t i ca l deployment rates would cause the system to librate out of control Normally
a smooth sinusoidal profile is chosen (S-curve) However one would also like a long
tether extending to ten twenty or even a hundred kilometers in a reasonable time
say 3-4 orbits This is usually difficult or impossible to accomplish solely wi th one
S-curve profile Often a composite S-curve-Steady-S-curve profile is adopted Thus
21
the deployment scheme for the t h tether can be summarized as
k = Tr l - cos (^r(ti - to-)) 1 ltUlt h 2 r V A V
k = Vov tiilttiltt2i
k = IT j 1 - C deg S lkitl ^ ^ lt i lt ^
(226)
where to- tt are the ini t ia l and final times of the deployment or retrieval maneuver
respectively A t j = t - tQ mdash tt mdash t2- and the steady deployment velocity VQ is
calculated based on the continuity of l and l at the specified intermediate times t
and to For the case of
and
Ai = i(ti) - ^(t 0 -) = - ^ ( t 2 i ) (227)
2Ali + li(tf) -h(tQ)
V = lUli l^lL (228)
1 tf-tQ v y
A t = (229)
22 Kinetics and System Energy
221 Kinetic energy
The kinetic energy of the ith link is given by
Kei = lJm Rdrnfi^Ami (230)
Setting Eq(219) in a matrix-vector form gives the relation
4 ^ ^ PM) T i $ 4 TiSi]i[t (231)
22
where
m
k J
x l (232)
and nq = -nfj + 7 is the number of generalized coordinates in each link Inserting
Eq(231) into Eq(230) and integrating the kinetic energy for the t h l ink can be
rewritten as
1 T
(233)
where Mti is the links symmetric mass matrix
Mt =
m i D D T D D DDTPifgidmi) DDTTi J ^dm D n T Sidm rn
sym fP(9i)Pi(9i)dmi J P[g^T^idm J P ( pound ) T ^ m sym sym J ^f^idm f ^f^dmi sym sym sym J sfsidmi
pound Sfttiqxnq
(234)
and mi is the mass of the Ith link The kinetic energy for the entire system ie N
bodies can now be stated as
1 bull T bull 1 bull T Ke = J2 MtA = o laquo Mtqt 2 ^ -i
i=l (235)
where qt = qj^^ bull bull -qfNT a n d Mt is a block diagonal decoupled mass matrix
of the system with Mf- on its diagonal
(236) Mt =
0 0
0 Mt2
0
0 0
MH
0 0 0
0 0 0 bull MtN
222 Simplification for rigid links
The mass matrix given by Eq(234) simplifies considerably for the case of rigid
23
links and is given as
Mti
m D D l
T D D l DD^Piiffidmi) 0
sym 0
0
fP(fi)Pi(i)dmi 0 0 0
0 Id nfj
0 m~
i = 1 3 5 N
(237)
Moreover when mult iplying the matrices in the (11) block term of Eq(237) one
gets
D D l
1 D D l = 1 0 0 0 D n
2 0 (238) 0 0 1
which is a diagonal matrix However more importantly i t is independent of 6 and
hence is always non-singular ensuring that is always invertible
In the (12) block term of M ^ it is apparent that f fidmi is simply the
definition of the centre of mass of the t h rigid link and hence given by
To drrii mdash miff (239)
where fcmi is the position vector of the centre of mass of link i For the case of the
(22) block term a little more attention is required Expanding the integrand using
Eq(218) where ltj = fj gives
fj Ta Tafi fj T Q Tpfi f- T Q T 7 i f j
sym sym
Now since each of the terms of Eq(240) is a scaler or a 1 x 1 matrix it is therefore
equal to its trace ie sum of its diagonal elements thus
tr(ffTai
TTaifi) tr(ffTai
TTpfi) tr(ff Ta^T^fi)
fjT^Tpn rfTp^fi (240)
P(rj)Pj(fi) = syrri trifjTpTTpfi) t r ^ T ^ T ^
sym sym tr(ff T 7 T r 7 f j) J
(241)
Using two properties of the trace function[47] namely
tr(ABC) = tr(BCA) = trCAB) (242)
24
and
j tr(A) =tr(^J A^j (243)
where A B and C are real matrices it becomes quite clear that the (22) block term
of Eq(237) simplifies to
P(fi)Pii)dmi =
tr(Tai
TTaiImi) tTTai
TTpImi) t r (T Q r T 7 7 m ) trTaTTpImi) tr(Tp
TT7Jmi) sym sym sym
(244)
where Im- is an alternate form of the inertia tensor of the t h link which when
expanded is
Imlaquo = I nfi dm bull i sym
sym sym
J x^dmi J Xydmi XZdm sym J yi2dmi f yiZidm
sym J z^dmi J
xVi -L VH sym ijxxi + lyy IZZJ)2
(245)
Note this is in terms of moments of inertia and products of inertia of the t f l rigid
link A similar simplification can be made for the flexible case where T is replaced
with Qi resulting in a new expression for Im
223 Gravitational potential energy
The gravitational potential energy of the ith link due to the central force law
can be written as
V9i f dm f
-A mdash = -A bullgttradelti RdmA
dm
inn Lgti + T j ^ | (246)
where z = 3986 x 105 km 3s 2 (Earths gravitational constant) Expanding binomially
and retaining up to third order terms gives
Vgi ~ ~ D T 2D [J gfgidrrii + 23f T j gidm mdashDT
D2 i l I I Df Tt I grffdmiTiDi
(2-lt 7)
25
where D = D is the magnitude of the inertial position vector to the frame F The
integrals in the above expression are functions of the flexible mode shapes and can be
evaluated through symbolic manipulation using an existing package such as Maple
V In turn Maple V can translate the evaluated integrals into F O R T R A N and store
the code in a file Furthermore for the case of rigid bodies it can be shown quite
readily that
j gjgidrrii = j rjfidrrn = [IXXi + Iyy + IZZi)2 (248)
Similarly the other integrals in Eq(234) can be integrated symbolically
224 Strain energy
When deriving the elastic strain energy of a simple string in tension the
assumptions of high tension and low amplitude of transverse vibrations are generally
valid Consequently the first order approximation of the strain-displacement relation
is generally acceptable However for orbiting tethers in a weak gravitational-gradient
field neglecting higher order terms can result in poorly modelled tether dynamics
Geometric nonlinearities are responsible for the foreshortening effects present in the
tether These cannot be neglected because they account for the heaving motion along
the longitudinal axis due to lateral vibrations The magnitude of the longitudinal
oscillations diminish as the tether becomes shorter[35]
W i t h this as background the tether strain which is derived from the theory
of elastic vibrations[3248] is given by
(249)
where Uj V and W are the flexible deformations along the Xj and Z direction
respectively and are obtained from Eq(25) The square bracketed term in Eq(249)
represents the geometric nonlinearity and is accounted for in the analysis
26
The total strain energy of a flexible link is given by
Ve = I f CiEidVl (250)
Substituting the stress-strain relationship 0 = E(poundi the strain energy is now given
by
Vei=l-EiAi J l l d x h (251)
where E^ is the tethers Youngs modulus and A is the cross-sectional area The
tether is assumed to be uniform thus EA which is the flexural stiffness of the
tether is assumed to be constant Eq(251) is evaluated symbolically for arbitrary
number of modes
225 Tether energy dissipation
The evaluation of the energy dissipation due to tether deformations remains
a problem not well understood even today Here this complex phenomenon is repshy
resented through a simplified structural damping model[32] In addition the system
exhibits hysteresis which also must be considered This is accomplished using an
augmented complex Youngs modulus
EX^Ei+jEi (252)
where Ej^ is the contribution from the structural damping and j = The
augmented stress-strain relation is now given as
a = Epoundi = Ei(l+jrldi)ei (253)
where 77^ = EjE^ ltC 1 is defined as the structural damping coefficient determined
experimentally [49]
27
If poundi is a harmonic function with frequency UJQ- then
jet = (254) wo-
Substituting into Eq(253) and rearranging the terms gives
e + I mdash I poundi = o-i + ad (255)
where ad and a are the stress with and without structural damping respectively
Now the Rayleigh dissipation function[50] for the ith tether can be expressed as
(ii) Jo
_ 1 EjAiVdi rU (2-56^ 2 w 0 Jo
The strain rate ii is the time derivative of Eq(249) The generalized external force
due to damping can now be written as Qd = Q^ bull bull bull QdNT where
dWd
^ = - i p = 2 4 ( 2 5 7 )
= 0 i = 13 N
23 0(N) Form of the Equations of Mot ion
231 Lagrange equations of motion
W i t h the kinetic energy expression defined and the potential energy of the
whole system given by Pe = ]Cn=l( 5i+ ej) the equations of motion can be obtained
quite readily using the Lagrangian principle
where q is the set of generalized coordinates to be defined in a later section
Substituting Eqs(235247251 and 257) into Eq(258) leads to the familiar
matr ix form of the non-linear non-autonomous coupled equations of motion for the
28
system
Mq t)t+ F(q q t) = Q(q q t) (259)
where M(q t) is the nonlinear symmetric mass matrix and F(q q t) is the forcing
function which can be written in matrix form as
Q(ltfgt Qi 0 is the vector of non-conservative generalized forces including control inshy
puts acting on the system A detailed expansion of Eq(260) in tensor notation is
developed in Appendix I
Numerical solution of Eq(259) in terms of q would require inversion of the
mass matrix However M is a full matrix of size nqq x nqq where nqq = Nnq is the
total number of generalized coordinates Therefore direct inversion of M would lead
to a large number of computation steps of the order nqq^ or higher The objective
here is to develop an order-N O(N) algorithm for obtaining M _ 1 that minimizes
the computational effort This is accomplished through the following transformations
ini t ia l ly developed for the stationkeeping case by Pradhan et al [30]
232 Generalized coordinates and position transformation
During the derivation of the energy expressions for the system the focus was
on the decoupled system ie each link was considered independent of the others
Thus each links energy expression is also uncoupled However the constraints forces
between two links must be incorporated in the final equations of motion
Let
e Unqxl (261) m
be the set of coupled coordinates of the t h l ink such that q = qT q2 bull bull bull 0T i s the
29
systems generalized coordinates Let qt- be the set of auxiliary decoupled coordinates
defined by Eq(232) such that qt = qj^ qj^ Qtj^1 bull The only difference between
qtj and q is the presence of D against d is the position of F wi th respect to
the inertial frame FQ whereas d is defined as the offset position of F relative to and
projected on the frame It can be expressed as d = dXidyidZi^ For the
special case of link 1 D = d
From Figure 2-1 can written as
Di = A - l + T i _ i J i _ i + di) + T V i S i - i f t - i + dXi)8i- (262)
Denoting KAj = $j(Zj + dx- ) Eq(262) can be rewritten in summation form as
i ampi = H ( T i - 1 ^ + T i - l K A i - l - l + T j - l ^ - l ) bull ( 2 - 6 3 )
3 = 1
Introducing the index substitution k mdash j mdash 1 into Eq(263)
i-1 D = ^2 (Tfc_ilt4 + TkKAk6k + Tkilk) + T^di (264)
k=l
since d = D TQ = 1^ and IQ DQ SQ are null vectors Recasting Eq(264) in
matr ix form leads to i-1
Qt Jfe=l
where
and
For the entire system
T fc-1 0 TkKAk
0 0 0 0 0 0 0 0 0 0 0 0 -
r T i - i 0 0 0 0 0 0 0 0 0 G
0 0 0 1
Qt = RpQ
G R n lt x l
(265)
(266)
G Unq x l (267)
(268)
30
where
RP
R 1 0 0 R R 2 0 R RP
2 R 3
0 0 0
R RP R PN _ 1 RN J
(269)
Here RP is the lower block triangular matrix which relates qt and q
233 Velocity transformations
The two sets of velocity coordinates qti and qi are related by the following
transformations
(i) First Transformation
Differentiating Eq(262) with respect to time gives the inertial velocity of the
t h l ink as
D = A - l + Ti_iK + + K A i _ i ^ _ i
+ T j _ i ^ + Ti-ii + T i - i K A i - i ^ + T j - i K A ^ i ^ i (270)
(271) (
Defining K A D = dddXi+1KAi K A L = d d K A and hi = di+1 +lfi +KA^i
results in K A ^ i = K A D j i o ^ + K A L i _ i i _ i
= KADi_iiTdi + K A L j _ i ^ _ i
Inserting Eq(271) into Eq(270) and rearranging gives
= A _ + T i ^ d g + K A D ^ x ^ i ^ ^ + Pi-^hi-iH-i
+ T i _ i K A i _ i 5 _ i + T i_ i2 + K A L j _ i 5 j _ i 4 _ i (272)
Using Eq(272) to define recursively - D j _ i and applying the index substitution
c = j mdash 1 as in the case of the position transformation it follows that
Di = ^ T ^ i K D ^ ^ + Pk(hk)ffk + T f c K A f c ^ + TfcKLfc4 Jfc=l (273)
+ T j _ i K D j _ i lt i ii
31
where K D j _ i = + K A D ^ i ^ - i r 7 and K L j = i + K A L j 5 j Finally by following a
procedure similar to that used for the position transformation it can be shown that
for the entire system
t = Rvil (2-74)
where Rv is given by the lower block diagonal matrix
Here
Rv
Rf 0 0 R R$ 0 R Rl RJ
0 0 0
T3V nN-l
R bullN
e Mnqqxl
Ri =
T i _ i K D i _ i TiKAi T j K L j 0 0 0 0 bullo 0 0 0 0 0 0 0
e K n lt x l
RS
T i-1 0 0 o-0 0 0 0 0 0
0 0 0 1
G Mnq x l
(275)
(276)
(277)
(ii) Second Transformation
The second transformation is simply Eq(272) set into matrix form This leads
to the expression for qt as
(278)
where Id Pihi) T j K A j T j K L j 0 0 0 0 0 0
0 0
0 0
0 0
G $lnq x l
Thus for the entire system
(279)
pound = Rnjt + Rdq
= Vdnqq - R nrlRdi (280)
32
where
and
- 0 0 0 RI 0 0
Rn = 0 0
0 0
Rf 0 0 0 rgtd K2 0
Rd = 0 0 Rl
0 0 0
RN-1 0
G 5R n w x l
0 0 0 e 5 R N ^ X L
(281)
(282)
234 Cylindrical coordinate modification
Because of the use of cylindrical coorfmates for the first body it is necessary
to slightly modify the above matrices to account for this coordinate system For the
first link d mdash D = D ^ D S L thus
Qti = 01 =gt DDTigi (283)
where
D D T i
Eq(269) now takes the form
DTL 0 0 0 d 3 0 0 0 ID
0 0 nfj 0
RP =
P f D T T i R2 0 P f D T T i Rp
2 R3
iJfDTTi i^
x l
0 0 0
R PN-I RN
(284)
(285)
In the velocity transformation a similar modification is necessary since D =
DD^DS^ Mak ing the substitution it can be shown that Eq(276) and Eq(279) now
33
takes the form for i = 1
T)V _ pn _ r t j mdash rt^ mdash 0 0 0 0
0 0 0 0 0 0 0 0
(286)
The rest of the matrices for the remaining links ie i = 2 N are unchanged
235 Mass matrix inversion
Returning to the expression for kinetic energy substitution of Eq(274) into
Eq(235) gives the first of two expressions for kinetic energy as
Ke = -q RvT MtRv (287)
Introduction of Eq(280) into Eq(235) results in the second expression for Ke as
1X 2q (hnqq-RnrlRd Mt (idnqq - R nrlRd (288)
Thus there are two factorizations for the systems coupled mass matrix given as
M = RV MtRv
M (hnqq - RnrlRd Mt ( i d n q q - R n r 1 R d
(289)
(290)
Inverting Eq(290)
r-1 _ ( rgtd~^ M 1 = (Ra) l d n q q - Rn)Mf1 (Rd)~l(Idnqq - Rn) (291)
Since both Rd and Mt are block diagonal matrices their inverse is simply the inverse
34
of each block on the diagonal ie
Mr1
h 0 0
0 0
0 0
0 0 0
0
0
0
tN
(292)
Each block has the dimension nq x nq and is independent of the number of bodies
in the system thus the inversion of M is only linearly dependent on N ie an O(N)
operation
236 Specification of the offset position
The derivation of the equations of motion using the O(N) formulation requires
that the offset coordinate d be treated as a generalized coordinate However this
offset may be constrained or controlled to follow a prescribed motion dictated either by
the designer or the controller This is achieved through the introduction of Lagrange
multipliers[51] To begin with the multipliers are assigned to all the d equations
Thus letting f = F - Q Eq(259) becomes
Mq + f = P C A (293)
where A G K 3 - 7 V x l is the vector of Lagrange multipliers and Pc G s R n 9 x 3 A r is the
permutation matrix assigning the appropriate Aj to its corresponding d equation
Inverting M and pre-multiplying both sides by PcT gives
pcTM-pc
= ( dc + PdegTM-^f) (294)
where dc is the desired offset acceleration vector Note both dc and PcTM lf are
known Thus the solution for A has the form
A pcTM-lpc - l dc + PC1 M (295)
35
Now substituting Eq(295) into Eq(293) and rearranging the terms gives
q = S(f Ml) = -M~lf + M~lPc
which is the new constrained vector equation of motion with the offset specified by
dc Note that pre-multiplying M~l by P c T provides the rows of M _ 1 corresponding
to the d equations Similarly post-multiplying by Pc leads the columns of the matrix
corresponding to the d equations
24 Generalized Control Forces
241 Preliminary remarks
The treatment of non-conservative external forces acting on the system is
considered next In reality the system is subjected to a wide variety of environmental
forces including atmospheric drag solar radiation pressure Earth s magnetic field as
well as dissipative devices to mention a few However in the present model only
the effect of active control thrusters and momemtum-wheels is considered These
generalized forces wi l l be used to control the attitude motion of the system
In the derivation of the equations of motion using the Lagrangian procedure
the contribution of each force must be properly distributed among all the generalized
coordinates This is acheived using the relation
nu ^5 Qk=E Pern ^ (2-97)
m=l deg Q k
where Qk and qk are the kth elements of Q and ltf respectively[51] Fem is the
mth external force acting on the system at the location Rm from the inertial frame
and nu is the total number of actuators Derivation of the generalized forces due to
thrusters and momentum wheels is given in the next two subsections
pcTM-lpc (dc + PcTM- Vj (2-96)
36
242 Generalized thruster forces
One set of thrusters is placed on each subsatellite ie al l the rigid bodies
except the first (platform) as shown in Figure 2-3 They are capable of firing in the
three orthogonal directions From the schematic diagram the inertial position of the
thruster located on body 3 is given as
Defining
R3 = di + Tid2 + T2(hi + KA2lt52) + T 3 f c m 3
= d + dfY + T 3 f c m 3
df = Tjdj+i + Tj+1lj+ii + KAj+i5j+i)
the thruster position for body 5 is given as
(298)
(299)
R5 = di+ dfx + d 3 + T5fcm
Thus in general the inertial position for the thruster on the t h body can be given
(2100)
as i - 2
k=l i1 cm (2101)
Let the column matrix
[Q i iT 9 R-K=rR dqj lA n i
WjRi A R i (2102)
where i = 3 5 N and j mdash 1 2 i The vector derivative of R wi th respect to
the scaler components qk is stored in the kth column of Thus for the case of
three bodies Eq(297) for thrusters becomes
1 Q
Q3
3 -1
T 3 T t 2 (2103)
37
Satellite 3
mdash gt -raquo cm
D1=d1
mdash gt
a mdash gt
o
Figure 2-3 Inertial position of subsatellite thruster forces
38
where ft = 0TtaTt^T is the thrust vector for the t h l ink (tether i-1) Here
Ttn and Tta are the thrust forces along the inplane m and out-of-plane Z directions i Pi
respectively For the case of 5-bodies
Q
~Qgt1
0
0
( T ^ 2 ) (2104)
The result for seven or more bodies follows the pattern established by Eq(2104)
243 Generalized momentum gyro torques
When deriving the generalized moments arising from the momemtum-wheels
on each rigid body (satellite) including the platform it is easier to view the torques as
coupled forces From Figure 2-4 it is apparent that for the ith l ink the generalized
forces constituting the couple are
(2105)
where
Fei = Faij acting at ei = eXi
Fe2 = -Fpk acting at e 2 = C a ^ l
F e 3 = F7ik acting at e 3 = eyj
Expanding Eq(2105) it becomes clear that
3
Qk = ^ F e i
i=l
d
dqk (2106)
which is independent of Ri- Thus the moments Mma = 2FaieXi Mm^ = 2FpeXi
and Mmi = 2FlieVi Transforming the coordinates to the body fixed frame using
the rotation matrix the generalized force due to the momentum-wheels on link i can
39
Figure 2-4 Coupled force representation of a momentum-wheel on a rigid body
40
be written as
d d d Qi = Td bull ^ T i t M m a - Tik bull mdashT^Mmp + T ^ bull T J j M ^
QiMmi
where Mm = Mma M m Mmi T
lt3 T J ^ T - T i f c bull ^ T i pound T ^ - ^ T z j
(2107)
(2108)
Note combining the thruster forces and momentum-wheel torques a compact
expression for the generalized force vector for 5 bodies can be written as
Q
Q T 3 0 Q T 5 0
0 0 0
0 lt33T3 Qyen Q5
3T5 0
0 0 0 Q4T5 0
0 0 0 Q5
5T5
M M I
Ti 2 m 3
Quu (2109)
The pattern of Eq(2109) is retained for the case of N links
41
3 C O M P U T E R I M P L E M E N T A T I O N
31 Prel iminary Remarks
The governing equations of motion for the TV-body tethered system were deshy
rived in the last chapter using an efficient O(N) factorization method These equashy
tions are highly nonlinear nonautonomous and coupled In order to predict and
appreciate the character of the system response under a wide variety of disturbances
it is necessary to solve this system of differential equations Although the closed-
form solutions to various simplified models can be obtained using well-known analytic
methods the complete nonlinear response of the coupled system can only be obtained
numerically
However the numerical solution of these complicated equations is not a straightshy
forward task To begin with the equations are stiff[52] be they have attitude and
vibration response frequencies separated by about an order of magnitude or more
which makes their numerical integration suceptible to accuracy errors i f a properly
designed integration routine is not used
Secondly the factorization algorithm that ensures efficient inversion of the
mass matrix as well as the time-varying offset and tether length significantly increase
the size of the resulting simulation code to well over 10000 lines This raises computer
memory issues which must be addressed in order to properly design a single program
that is capable of simulating a wide range of configurations and mission profiles It
is further desired that the program be modular in character to accomodate design
variations with ease and efficiency
This chapter begins with a discussion on issues of numerical and symbolic in-
42
tegration This is followed by an introduction to the program structure Final ly the
validity of the computer program is established using two methods namely the conshy
servation of total energy and comparison of the response results for simple particular
cases reported in the literature
32 Numerical Implementation
321 Integration routine
A computer program coded in F O R T R A N was written to integrate the equashy
tions of motion of the system The widely available routine used to integrate the difshy
ferential equations D G E A R [ 5 3 ] is based on Gears predictor-corrector method[54]
It is well suited to stiff differential equations as i t provides automatic step-size adjustshy
ment and error-checking capabilities at each iteration cycle These features provide
superior performance over traditional 4^ order Runge-Kut ta methods
Like most other routines the method requires that the differential equations be
transformed into a first order state space representation This is easily accomplished
by letting
thus
X - ^ ^ f ^ ) =ltbullbull) (3 -2 )
where S(fMdc) is defined by Eq(296) Eq(32) represents the new set of 2nqq
first order differential equations of motion
322 Program structure
A flow chart representing the computer programs structure is presented in
Figure 3-1 The M A I N program is responsible for calling the integration routine It
also coordinates the flow of information throughout the program The first subroutine
43
called is INIT which initializes the system variables such as the constant parameters
(mass density inertia etc) and the ini t ial conditions I C (attitude angles orbital
motion tether elastic deformation etc) from user-defined data files Furthermore all
the necessary parameters required by the integration subroutine D G E A R (step-size
tolerance) are provided by INIT The results of the simulation ie time history of the
generalized coordinates are then passed on to the O U T P U T subroutine which sorts
the information into different output files to be subsequently read in by the auxiliary
plott ing package
The majority of the simulation running time is spent in the F C N subroutine
which is called repeatedly by D G E A R F C N generates the fv(Xi) vector defined in
Eq(32) It is composed of several subroutines which are responsible for the comshy
putation of all the time-varying matrices vectors and scaler variables derived in
Chapter 2 that comprise the governing equations of motion for the entire system
These subroutines have been carefully constructed to ensure maximum efficiency in
computational performance as well as memory allocation
Two important aspects in the programming of F C N should be outlined at this
point As mentioned earlier some of the integrals defined in Chapter 2 are notably
more difficult to evaluate manually Traditionally they have been integrated numershy
ically on-line (during the simulation) In some cases particularly stationkeeping the
computational cost is quite reasonable since the modal integrals need only be evalushy
ated once However for the case of deployment or retrieval where the mode shape
functions vary they have to be integrated at each time-step This has considerable
repercussions on the total simulation time Thus in this program the integrals are
evaluated symbolically using M A P L E V[55] Once generated they are translated into
F O R T R A N source code and stored in I N C L U D E files to be read by the compiler
Al though these files can be lengthy (1000 lines or more) for a large number of flexible
44
INIT
DGEAR PARAM
IC amp SYS PARAM
MAIN
lt
DGEAR
lt
FCN
OUTPUT
STATE ACTUATOR FORCES
MOMENTS amp THRUST FORCES
VIBRATION
r OFFS DYNA
ET MICS
Figure 3-1 Flowchart showing the computer program structure
45
modes they st i l l require less time to compute compared to their on-line evaluation
For the case of deployment the performance improvement was found be as high as
10 times for certain cases
Secondly the O(N) algorithm presented in the last section involves matrices
that contained mostly zero terms (RP Rv and Mplusmn) The sparse structure of these mashy
trices can be used advantageously to improve the computational efficiency by avoiding
multiplications of zero elements The result is a quick evaluation of fv(Xt) Moreshy
over there is a considerable saving of computer memory as zero elements are no
longer stored
Final ly the C O N T R O L subroutine which is part of F C N is designed to calshy
culate the appropriate control forces torques and offset dynamics which are required
to stabilize the librational ( A T T I T U D E ) and vibrational ( V I B R A T I O N ) motion of
the system
33 Verification of the Code
The simulation program was checked for its validity using two methods The
first verified the conservation of total energy in the absence of dissipation The second
approach was a direct comparison of the planar response generated by the simulation
program with those available in the literature A s numerical results for the flexible
three-dimensional A^-body model are not available one is forced to be content with
the validation using a few simplified cases
331 Energy conservation
The configuration considered here is that of a 3-body tethered system with
the following parameters for the platform subsatellite and tether
46
h =
1091430 -8135 328108 -8135 8646050 27116 328108 27116 8286760
kg bull m 2 platform inertia
kg bull m 2 end-satellite inertia 200 0 0
0 400 0 0 0 400
m = 90000 kg mass of the space station platform
m2 = 500 kg mass of the end-satellite
EfAt = 61645 N tether elastic stiffness
pt mdash 49 kg km tether linear density
The tether length is taken to remain constant at 10 km (stationkeeping case)
A n in i t ia l disturbance in pitch and roll of 2deg and 1deg respectively is given to both
the rigid bodies and the flexible tether In addition the tether is ini t ia l ly deflected
by 45 m in the longitudinal direction at its end and 05 m in both the inplane and
out-of-plane directions in the first mode The tether attachment point offset at the
platform (satellite 1) end is 1 m in all the three directions ie d2 = 11 lTm The
attachment point at the subsatellite (satellite 2) end is 1 m in the x direction ie
fcm3 = 10 0T
The variation in the kinetic and potential energies is presented for the above-
mentioned case in Figure 3-2(a) As seen from the plot there is a continuous mutual
exchange between the kinetic and potential energies however the variation in the
total energy remains zero (Figure 3-2b) It should be noted that after 1 orbit the
variation of both kinetic and potential energy does not return to 0 This is due to
the attitude and elastic motion of the system which shifts the centre of mass of the
tethered system in a non-Keplerian orbit
Similar results are presented for the 5-body case ie a double pendulum (Figshy
ure 3-3) Here the system is composed of a platform and two subsatellites connected
47
STSS configuration 3-Body a 1(0)=a 2(0)=a t(0)=2deg P1(0)=p2(0)=p(0)=1deg 5x(0)=45m 8y(0)=52(0)=05m dx(0)=dy(0)=d2(0)=1m Stationkeeping lt=10km
Variation of Kinetic and Potential Energy
(a) Time (Orbits) Percent Variation in Total Energy
OOEOh
-50E-12
-10E-11 I i i
0 1 2 (b) Time (Orbits)
Figure 3-2 Kinetic and potential energy transfer for the three-body platform based tethered satellite system (a) variation of kinetic and potential energy (b) percent change in total energy of system
48
STSS configuration 5-Body a 1(0)=a 2(0)=a t 1(0)=2deg a 3(0)=a t 2(0)=25deg P1(0)=p2(0)=p3(0)=pt1(0)=pt2(0)=1lt
5x(0)=5m 8y(0)=5z(0)=05m dx(0)=dy(0)=dz(0)=0 Stationkeeping 1 =12=1 Okm
(a)
20E6
10E6r-
00E0
-10E6
-20E6
Variation of Kinetic and Potential Energy
-
1 I I I 1 I I
I I
AP e ~_
i i 1 i i i i
00E0
UJ -50E-12 U J
lt
Time (Orbits) Percent Variation in Total Energy
-10E-11h
(b) Time (Orbits)
Figure 3 -3 Kinet ic and potential energy transfer for the five-body platform based tethered satellite system (a) variation of kinetic and potential enshyergy (b) percent change in total energy of system
49
in sequence by two tethers The two subsatellites have the same mass and inertia
ie 7713 = 7 7 i 2 a n d I3 = I21 a n d are separated by identical tether segments each 10
k m in length The platform has the same properties as before however there is no
tether attachment point offset present in the system (d = fcmi = 0)
In al l the cases considered energy was found to be conserved However it
should be noted that for the cases where the tether structural damping deployment
retrieval attitude and vibration control are present energy is no longer conserved
since al l these features add or subtract energy from the system
332 Comparison with available data
The alternative approach used to validate the numerical program involves
comparison of system responses with those presented in the literature A sample case
considered here is the planar system whose parameters are the same as those given
in Section 331 A n ini t ia l pitch disturbance of 2deg is imparted to both the platform
and the tether together with an ini t ial tether deformation of 08 m and 001 m in
the longitudinal and transverse directions respectively in the same manner as before
The tether length is held constant at 5 km and it is assumed to be connected to the
centre of mass of the rigid platform and subsatellite The response time histories from
Ref[43] are compared with those obtained using the present program in Figures 3-4
and 3-5 Here ap and at represent the platform and tether pitch respectively and B i
C i are the tethers longitudinal and transverse generalized coordinates respectively
A comparison of Figures 3-4 and 3-5 clearly shows a remarkable correlation
between the two sets of results in both the amplitude and frequency The same
trend persisted even with several other comparisons Thus a considerable level of
confidence is provided in the simulation program as a tool to explore the dynamics
and control of flexible multibody tethered systems
50
a p(0) = 2deg 0(0) = 2C
6(0) = 08 m
0(0) = 001 m
Stationkeeping L = 5 km
D p y = 0 D p 2 = 0
a oo
c -ooo -0 01
Figure 3 - 4 Simulation results for the platform based three-body tethered system originally presented in Ref[43]
51
STSS configuration 3-Body a1(0)=a2(0)=at(0)=2deg p1(0)=(32(0)=(3t(0)=0 8x(0)=08m 5y(0)=001m 5Z(0)=0 dx(0)=dy(0)=dz(0)=0 Stationkeeping l=5km
Satellite 1 Pitch Angle Tether Pitch Angle
i -i i- i i i i i i d
1 2 0 1 2 Time (Orbits) Time (Orbits)
Longitudinal Vibration
^MAAAAAAAWVWWWWVVVW
V bull i bull i I i J 00 05 10 15 20
Time (Orbits) Transverse Vibration
Time (Orbits)
Figure 3-5 Simulation results for the platform based three-body tethered system obtained using the present computer program
52
4 D Y N A M I C SIMULATION
41 Prel iminary Remarks
Understanding the dynamics of a system is critical to its design and develshy
opment for engineering application Due to obvious limitations imposed by flight
tests and simulation of environmental effects in ground based facilities space based
systems are routinely designed through the use of numerical models This Chapter
studies the dynamical response of two different tethered systems during deployment
retrieval and stationkeeping phases In the first case the Space platform based Tethshy
ered Satellite System (STSS) which consists of a large mass (platform) connected to
a relatively smaller mass (subsatellite) with a long flexible tether is considered The
other system is the O E D I P U S B I C E P S configuration involving two comparable mass
satellites interconnected by a flexible tether In addition the mission requirement of
spin about an arbitrary axis the tether axis for O E D I P U S - A C and cartwheeling or
spin about the orbit normal for the proposed B I C E P S mission is accounted for
A s mentioned in Chapter 2 the tether flexibility is modelled using the asshy
sumed mode discretization method Although the simulation program can account
for an arbitrary number of vibrational modes in each direction only the first mode
is considered in this study as it accounts for most of the strain energy[34] and hence
dominates the vibratory motion The parametric study considers single as well as
double pendulum type systems with offset of the tether attachment points The
systems stability is also discussed
42 Parameter and Response Variable Definitions
The system parameters used in the simulation unless otherwise stated are
53
selected as follows for the STSS configuration
1091430 -8135 328108 -8135 8646050 27116 328108 27116 8286760
h =
h = 200 0 0
0 400 0 0 0 400
bull m = 90000 kg (mass of the space platform)
kg bull m 2 (platform inertia) 760 J
kg bull m 2 (subsatellite inertia)
bull 7772 = 500 kg (mass of the subsatellite)
bull EtAt = 61645 N (tether stiffness)
bull Pt = 49 k g k m (tether density)
bull Vd mdash 05 (tether structural damping coefficient)
bull fcm^ = l 0 0 r m (tether attachment point at the subsatellite)
The response variables are defined as follows
bull ai0 satellite 1 (platform) pitch and roll angles respectively
bull CK2 2 satellite 2 (subsatellite) pitch and roll angles respectively
bull at fit- tether pitch and roll angles respectively
bull If tether length
bull d = dx dy dzT- tether attachment position relative to satellite 1 (platform)
mdash
bull S = 8X 8y Sz tether flexible generalized coordinates in the longitudinal x
inplane transverse y and out-of-plane transverse z directions respectively
The attitude angles and are measured with respect to the Local Ve r t i c a l -
Loca l Horizontal ( L V L H ) frame A schematic diagram illustrating these variable
is presented in Figure 4-1 The system is taken to be in a nominal circular orbit at
an altitude of 289 km with an orbital period of 903 minutes
54
Satellite 1 (Platform) Y a w
Figure 4-1 Schematic diagram showing the generalized coordinates used to deshyscribe the system dynamics
55
43 Stationkeeping Profile
To facilitate comparison of the simulation results a reference case based on
the three-body STSS system with zero offset (d = 0 r c m 3 = 0) and a tether length
of It = 20 km is considered first (Figure 4-2) The system is subjected to an ini t ial
disturbance in pitch and roll of 2deg and 1deg respectively to all the three bodies
In addition an ini t ia l longitudinal deflection of 14 m from the tethers unstretched
position is given together with a transverse inplane and out-of-plane deflection of 1
m at the tethers mid-length From the attitude response given in Figure 4-2(a) it is
apparent that the three bodies oscillate about their respective equilibrium positions
Since coupling between the individual rigid-body dynamics is absent (zero offset) the
corresponding librational frequencies are unaffected Satellite 1 (Platform) displays
amplitude modulation arising from the products of inertia The result is a slight
increase of amplitude in the pitch direction and a significantly larger decrease in
amplitude in the roll angle
Figure 4-2(b) clearly shows coupling of the tethers attitude motion with its
flexibility dynamics The tether vibrates in the axial direction about its equilibrium
position (at approximately 138 m) with two vibration frequencies namely 012 Hz
and 31 x 1 0 - 4 Hz The former is the first longitudinal flexible mode which is dissishy
pated through the structural damping while the latter arises from the coupling with
the pitch motion of the tether Similarly in the transverse direction there are two
frequencies of oscillations arising from the flexible mode and its coupling wi th the
attitude motion As expected there is no apparent dissipation in the transverse flexshy
ible motion This is attributed to the weak coupling between the longitudinal and
transverse modes of vibration since the transverse strain in only a second order effect
relative to the longitudinal strain where the dissipation mechanism is in effect Thus
there is very litt le transfer of energy between the transverse and longitudinal modes
resulting in a very long decay time for the transverse vibrations
56
STSS configuration 3-Body a1(0)=a2(0)=a(0)=2deg (31(0)=p2(0)=p(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=dz(0)=0 Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
0 1 2 3 4 5 0 1 2 3 4 5 Time (Orbits) Time (Orbits)
Satellite 2 Pitch Angle Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-2 Stationkeeping dynamics of the three-body STSS configuration withshyout offset (a) attitude response
57
STSS configuration 3-Body a1(0)=cx2(0)=cx(0)=2deg P1(0)=p2(0)=p(0)=1deg 5x(0)=14m8y(0)=82(0)=1m dx(0)=dy(0)=dz(0)=0 Stationkeeping lt=20km
Tether Longitudinal Vibration 16| 1 1 1 1 1 r-
Time (Orbits) Tether In-Plane Transverse Vibration
imdash 1 mdash 1 mdash 1 mdash 1 mdashr
r I I 1 1 1 I 1 1 1 1 I 1 bull ^ bull bull I I J 0 1 2 3 4 5
Time (Orbits) Tether Out-of-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-2 Stationkeeping dynamics of the three-body STSS configuration withshyout offset (b) vibration response
58
Introduction of the attachment point offset significantly alters the response of
the attitude motion for the rigid bodies W i t h a i m offset along the local vertical
at the platform (dx = 1 m) and fcm^ = 100 m coupling between the tether
and end-bodies is established by providing a lever arm from which the tether is able
to exert a torque that affects the rotational motion of the rigid end-bodies (Figure
4-3) Bo th the platform and subsatellite now oscillate at the same frequency as the
tether whose motion remains unaltered In the case of the platform there is also an
amplitude modulation due to its non-zero products of inertia as explained before
However the elastic vibration response of the tether remains essentially unaffected
by the coupling
Providing an offset along the local horizontal direction (dy = 1 m) results in
a more dramatic effect on the pitch and roll response of the platform as shown in
Figure 4-4(a) Now the platform oscillates about its new equilibrium position of
-90deg The roll motion is also significantly disturbed resulting in an increase to over
10deg in amplitude On the other hand the rigid body dynamics of the tether and the
subsatellite as well as the tether flexible motion remain the same (Figure 4-4b)
Figure 4-5 presents the response when the offset of 1 m at the platform end is
along the z direction ie normal to the orbital plane The result is a larger amplitude
pitch oscillation about the reference equilibrium position as shown in Figure 4-5(a)
while the roll equilibrium is now shifted to approximately 90deg Note there is little
change in the attitude motion of the tether and end-satellites however there is a
noticeable change in the out-of-plane transverse vibration of the tether (Figure 4-5b)
which may be due to the large amplitude platform roll dynamics
Final ly by setting a i m offset in all the three direction simultaneously the
equil ibrium position of the platform in the pitch and roll angle is altered by approxishy
mately 30deg (Figure 4-6a) However the response of the system follows essentially the
59
STSS configuration 3-Body a1(0)=cx2(0)=cxt(0)=2deg p1(0)=(32(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=1m dy(0)=dz(0)=0 Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
Time (Orbits) Time (Orbits) Tether Pitch Angle Tether Roll Angle
L bull I I bull bull lt bull bull bull raquo bull bull bull bull -I h I I I I i H 0 1 2 3 4 5 0 1 2 3 4 5
Time (Orbits) Time (Orbits) Satellite 2 Pitch Angle Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-3 Stationkeeping dynamics of the three-body STSS configuration with offset along the local vertical (a) attitude response
60
STSS configuration 3-Body a 1(0)=a 2(0)=a t(0)=2deg (31(0)=p2(0)=pt(0)=1deg 5x(0)=14m5y(0)=5z(0)=1m dx(0)=1mdy(0)=dz(0)=0 Stationkeeping lt=20km
Tether Longitudinal Vibration 1 6 i mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash mdash i mdash i mdash mdash mdash mdash mdash i mdash mdash lt -
Time (Orbits)
Tether In-Plane Transverse Vibration
y bull i i i i i i i i i i i i i i i i i 0 1 2 3 4 5
Time (Orbits)
Tether Out-of-Plane Transverse Vibration r 1 1 1 1 1 1 1 1 1 1 1 1 1 1 q
h i I i i I i i i i I i i i i 3
0 1 2 3 4 5 (b) Time (Orbits)
Figure 4-3 Stationkeeping dynamics of the three-body STSS configuration with offset along the local vertical (b) vibration response
61
STSS configuration 3-Body a1(0)=cx2(0)=o(0)=2deg (31(0)=p2(0)=p(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dz(0)=0 dy(0)=1m Stationkeeping lt=20km
Satellite 1 Pitch Angle i 1 i
Satellite 1 Roll Angle
1 2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 4 Time (Orbits)
Tether Roll Angle
CD
edT
1 2 3 Time (Orbits)
Satellite 2 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 2 Roll Angle
CD
C O
(a)
Figure 4-
1 2 3 4 Time (Orbits)
1 2 3 4 Time (Orbits)
4 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal (a) attitude response
62
STSS configuration 3-Body a1(0)=cx2(0)=ot(0)=2deg p1(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dz(0)=0 dy(0)=1m Stationkeeping lt=20km
Tether Longitudinal Vibration 16| 1 gt 1 bull 1 1 1
0 1 2 3 4 5
Time (Orbits) Tether In-Plane Transverse Vibration
i mdash 1 mdash 1 mdash 1 mdash 1 mdash i mdash mdash lt mdash 1 mdash 1 mdash i mdash mdash 1 mdash 1 mdash
V I I bull bull bull I i i i I i i 1 1 3
0 1 2 3 4 5 Time (Orbits)
Tether Out-of-Plane Transverse Vibration r 1 1 1 1 1 1 1 1 1 1 1 gt bull | q
V i 1 1 1 1 1 mdash 1 mdash 3
0 1 2 3 4 5 (b) Time (Orbits)
Figure 4-4 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal (b) vibration response
63
STSS configuration 3-Body a1(0)=o2(0)=at(0)=2deg (31(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=0 dz(0)=1m Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 4 Time (Orbits)
Tether Roll Angle
1 2 3 4 Time (Orbits)
Satellite 2 Pitch Angle
2 3 4 Time (Orbits)
Satellite 2 Roll Angle
(a)
Figure 4-
1 2 3 4 Time (Orbits)
2 3 4 Time (Orbits)
5 Stationkeeping dynamics of the three-body STSS configuration with offset along the orbit normal (a) attitude response
64
STSS configuration 3-Body a1(0)=cx2(0)=ot(0)=2deg P1(0)=p2(0)=(3(0)=1deg 8x(0)=14m8y(0)=82(0)=1m dx(0)=dy(0)=0 d2(0)=1m Stationkeeping lt=20km
Tether Longitudinal Vibration 1 6 i mdash r mdash i mdash mdash i mdash i mdash mdash i mdash i mdash mdash i mdash mdash i mdash 1 mdash 1 mdash i mdash mdash lt mdash lt mdash lt mdash r
Time (Orbits) Tether In-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-5 Stationkeeping dynamics of the three-body STSS configuration with offset along the orbit normal (b) vibration response
65
same trend with only minor perturbations in the transverse elastic vibratory response
of the tether (Figure 4-6b)
44 Tether Deployment
Deployment of the tether from an ini t ial length of 200 m to 20 km is explored
next Here the deployment length profile is critical to ensure the systems stability
It is not desirable to deploy the tether too quickly since that can render the tether
slack Hence the deployment strategy detailed in Section 216 is adopted The tether
is deployed over 35 orbits with the sinusoidal acceleration and deceleration occurring
over a 4 km length ( A ^ 2 ) with the remaining deployment at a constant velocity of
V0 = 146 ms
Initially the tether is stretched only slightly S = 1304 x 1 0 _ 3 0 0 m
with al l the other states of the system ie pitch roll and transverse displacements
remaining zero The response for the zero offset case is presented in Figure 4-7 For
the platform there is no longer any coupling with the tether hence it oscillates about
the equilibrium orientation determined by its inertia matrix However there is st i l l
coupling between the subsatellite and the tether since rcm^ mdash 100 m resulting
in complete domination of the subsatellite dynamics by the tether Consequently
the pitch motion ini t ial ly grows by almost 50deg but eventually subsides as the tether
deployment rate decreases It is interesting to note that the out-of-plane motion is not
affected by the deployment This is because the Coriolis force which is responsible
for the tether pitch motion does not have a component in the z direction (as it does
in the y direction) since it is always perpendicular to the orbital rotation (Q) and
the deployment rate (It) ie Q x It
Similarly there is an increase in the amplitude of the transverse elastic oscilshy
lations of the tether again due to the Coriolis effect (Figure 4-7b) Moreover as the
66
STSS configuration 3-Body a1(0)=a2(0)=cxt(0)=2deg P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m5y(0)=52(0)=1m dx(0)=dy(0)=d2(0)=1m Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
i i i i i i i i i i i 0 1 2 3 4 5 0 1 2 3 4 5
Time (Orbits) Time (Orbits) Tether Pitch Angle Tether Roll Angle
l i i i i -I h i i i i I i i i i l- i i i i l i i i i I i i- i i H 0 1 2 3 4 5 0 1 2 3 4 5
Time (Orbits) Time (Orbits) Satellite 2 Pitch Angle Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-6 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal local vertical and orbit normal (a) attitude response
67
STSS configuration 3-Body a1(0)=cc2(0)=a(0)=2deg p1(0)=p2(0)=(3t(0)=1deg 5x(0)=14m6y(0)=52(0)=1m dx(0)=dy(0)=d2(0)=1m Stationkeeping lt=20km
Tether Longitudinal Vibration 1 6 1 mdash i mdash i mdash i mdash mdash i mdash 1 mdash 1 mdash mdash mdash i mdash mdash 1 mdash 1 mdash 1 mdash i mdash 1 mdash 1 mdash 1 mdash 1 mdash r
0 1 2 3 4 5
Time (Orbits) Tether In-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-6 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal local vertical and orbit normal (b) vibration response
68
STSS configuration 3-Body a(0)=a2(0)=ot(0)=0 P1(0)=p2(0)=pt(0)=0 8X(0)=1304 x 103m 8y(0)=8z(0)=0 dx(0)=dy(0)=dz(0)=0 Deployment
l=02km to 20km in 35 Orbits
Tether Length Profile
E
Satellite 1 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 1 Roll Angle mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash [ mdash i mdash i mdash i mdash i mdash | mdash
1 2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 4 Time (Orbits)
Tether Roll Angle
CD
oo ax
i imdashimdashimdashImdashimdashimdashimdashr-
_ 1 _ L
1 2 3 4 Time (Orbits)
Satellite 2 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 2 Roll Angle
-50
(a) 1 2 3 4
Time (Orbits) 1 2 3 4
Time (Orbits)
Figure 4-7 Deployment dynamics of the three-body STSS configuration without offset (a) attitude response
69
STSS configuration 3-Body a1(0)=a2(0)=at(0)=0 p1(0)=p2(0)=(3t(0)=0 8X(0)=1304x 103m 8y(0)=52(0)=0 dx(0)=dy(0)=d2(0)=0 Deployment
l=02km to 20km in 35 Orbits
Tether Longitudinal Vibration
to
-02
(b)
Figure 4-7
2 3 Time (Orbits)
Tether In-Plane Transverse Vibration
0 1 1 bull 1 1
1 2 3 4 Time (Orbits)
Tether Out-of-Plane Transverse Vibration
5
1 1 1 1 1 i i | i i | ^
_ i i i i L -I
1 2 3 4 5 Time (Orbits)
Deployment dynamics of the three-body STSS configuration without offset (b) vibration response
70
tether elongates its longitudinal static equilibrium position also changes due to an
increase in the gravity-gradient tether tension It may be pointed out that as in the
case of attitude motion there are no out-of-plane tether vibrations induced
When the same deployment conditions are applied to the case of 1 m offset in
the x direction (Figure 4-8) coupling effects are re-introduced The platform pitch
exceeds -100deg during the peak deployment acceleration However as the tether pitch
motion subsides the platform pitch response returns to a smaller amplitude oscillation
about its nominal equilibrium On the other hand the platform roll motion grows
to over 20deg in amplitude in the terminal stages of deployment Longitudinal and
transverse vibrations of the tether remain virtually unchanged from the zero offset
case except for minute out-of-plane perturbations due to the platform librations in
roll
45 Tether Retrieval
The system response for the case of the tether retrieval from 20 km to 200 m
with an offset of 1 m in the x and z directions at the platform end is presented in
Figure 4-9 Here the Coriolis force induced during retrieval renders the tether unstable
(Figure 4-9a) Consequently the pitch motion of the platform is also destabilized
through coupling The z offset provides an additional moment arm that shifts the
roll equilibrium to 35deg however this degree of freedom is not destabilized From
Figure 4-9(b) it is apparent that the transverse mode of the tether vibration is also
disturbed producing a high frequency response in the final stages of retrieval This
disturbance is responsible for the slight increase in the roll motion for both the tether
and subsatellite
46 Five-Body Tethered System
A five-body chain link system is considered next To facilitate comparison with
71
STSS configuration 3-Body a(0)=a2(0)=at(0)=0 P1(0)=P2(0)=P(0)=0 5X(0)=1304 x 103m 8y(0)=8z(0)=0 dx(0)=1mdy(0)=d2(0)=0 Deployment
lt=02km to 20km in 35 Orbits
Tether Length Profile
E
Satellite 1 Pitch Angle
1 2 3 Time (Orbits)
Satellite 1 Roll Angle
bull bull bull i bull bull bull bull i
1 2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 Time (Orbits)
Tether Roll Angle
i i i i J I I i _
1 2 3 4 Time (Orbits)
Satellite 2 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 2 Roll Angle 3E-3h
D )
T J ^ 0 E 0 eg
CQ
-3E-3
(a) 1 2 3 4
Time (Orbits) 1 2 3 4
Time (Orbits)
Figure 4-8 Deployment dynamics of the three-body STSS configuration with offset along the local vertical (a) attitude response
7 2
STSS configuration 3-Body a1(0)=a2(0)=at(0)=0 p1(0)=P2(0)=(3t(0)=0 Sx(0)=1304 x 103m 5y(0)=5z(0)=0 dx(0)=1mdy(0)=dz(0)=0 Deployment
l=02km to 20km in 35 Orbits
Tether Longitudinal Vibration
Time (Orbits) Tether In-Plane Transverse Vibration
n 1 1 1 f
I I I _j I I I I I 1 1 1 1 1 2 3 4 5
Time (Orbits) Tether Out-of-Plane Transverse Vibration
J 1 I I 1 I I I I I I I I L _ _ I I d 1 2 3 4 5
Time (Orbits)
Deployment dynamics of the three-body STSS configuration with offset along the local vertical (b) vibration response
(b)
Figure 4-8
73
STSS configuration 3-Body a1(0)=a2(0)=at(0)=0 p(0)=P2(0)=(3t(0)=0 Sx(0)=14m8y(0)=82(0)=0 dy(0)=0dx(0)=d2(0)=1m Retrieval
l=20km to 02km in 35 Orbits
Satellite 1 Pitch Angle
(a) Time (Orbits)
Figure 4-9 Retrieval dynamics of the along the local vertical ar
Tether Length Profile
i i i i i i i i i i I 0 1 2 3
Time (Orbits) Satellite 1 Roll Angle
Time (Orbits) Tether Roll Angle
b_ i i I i I i i L _ J
0 1 2 3 Time (Orbits)
Satellite 2 Roll Angle _ mdash | r i i | 1 i i imdashmdashj i 1 r mdash i mdash
04- -
0 1 2 3 Time (Orbits)
three-body STSS configuration with offset d orbit normal (a) attitude response
74
STSS configuration 3-Body a1(0)=a2(0)=ocl(0)=0 (31(0)=f32(0)=pt(0)=0 8x(0)=14m8y(0)=6z(0)=0 dy(0)=0dx(0)=dz(0)=1m Retrieval
l=20km to 02km in 35 Orbits
Tether Longitudinal Vibration
E 10 to
Time (Orbits) Tether In-Plane Transverse Vibration
1 2 Time (Orbits)
Tether Out-of-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-9 Retrieval dynamics of the three-body STSS configuration with offset along the local vertical and orbit normal (b) vibration response
75
the three-body system dynamics studied earlier the chain is extended through the
addition of two bodies a tether and a subsatellite with the same physical properties
as before (Section 42) Thus the five-body system consists of a platform (satellite 1)
tether 1 subsatellite 1 (satellite 2) tether 2 and subsatellite 2 (satellite 3) as shown
in Figure 4-10 The offset of tether 1 at the platform-end is denoted by d2 while that
of tether 2 to subsatellite 1 as 04 In this numerical example fcm^ = fcm^ = 0 ie
tethers 1 and 2 are attached to the centre of mass of subsatellites 1 and 2 respectively
The system response for the case where the tether attachment points coincide
wi th the centres of mass of the rigid bodies ie zero offset (d2 = d mdash 0) is presented
in Figure 4-11 Here ogtt and at2 represent the pitch motion of tethers 1 and 2
respectively whereas 03 is the pitch motions of satellite 3 A similar convention
is adopted for the rol l angle 0 A s before the zero offset eliminates the coupling
between the tethers and satellites such that they are now free to librate about their
static equil ibrium positions However their is s t i l l mutual coupling between the two
tethers This coupling is present regardless of the offset position since each tether is
capable of transferring a force to the other The coupling is clearly evident in the pitch
response of the two tethers (Figure 4 - l l a ) O n the other hand the roll motion appears
uncoupled as the in i t ia l conditions in rol l for the two tethers are identical Thus the
motion is in phase and there is no transfer of energy through coupling A s expected
during the elastic response the tethers vibrate about their static equil ibrium positions
and mutually interact (Figure 4 - l l b ) However due to the variation of tension along
the tether (x direction) they do not have the same longitudinal equilibrium Note
relatively large amplitude transverse vibrations are present particularly for tether 1
suggesting strong coupling effects with the longitudinal oscillations
76
Figure 4-10 Schematic diagram of the five-body system used in the numerical example
77
STSS configuration 5-Body a1(0)=(x2(0)=at1(0)=2deg a3(0)=a12(0)=25deg p1(0)=r32(0)=p3(0)=pt1(0)=pt2(0)=r 82(0)=84(0)=505)05m d2(0)=d4(0)=000 Stationkeeping lt1=lu=10km Pitch Roll
Satellite 1 Pitch Angle Satellite 1 Roll Angle
o 1 2 Time (Orbits)
Tether 1 Pitch and Roll Angle
0 1 2 Time (Orbits)
Tether 2 Pitch and Roll Angle
Time (Orbits) Satellite 2 Pitch and Roll Angle
Time (Orbits) Satellite 3 Pitch and Roll Angle
Time (Orbits) Time (Orbits)
Figure 4-11 Stationkeeping dynamics of the five-body STSS configuration withshyout offset (a) attitude response
78
STSS configuration 5-Body a1(0)=a2(0)=oct1(0)=2deg a3(0)=at2(0)=25o
p1(0)=p2(0)=(33(0)=pt1(0)=pt2(0)=10
62(0)=54(0)=50505m d2(0)=d4(0)=000 Stationkeeping lt1=1 =1 Okm
Tether 1 Longitudinal Vibration Tether 2 Longitudinal Vibration i i 1 i 1 r 1 1 1 1
(b) Time (Orbits) Time (Orbits)
Figure 4-11 Stationkeeping dynamics of the five-body STSS configuration withshyout offset (b) vibration response
79
The system dynamics during deployment of both the tethers in the double-
pendulum configuration is presented in Figure 4-12 Deployment of each tether takes
place from 200 m to 20 km in 35 orbits The pitch motion of the system is similar
to that observed during the three-body case The Coriolis effect causes the tethers
to pitch through a large angle ( laquo 50deg ) which in turn disturbs the platform and
subsatellite due to the presence of offset along the local vertical On the other hand
the roll response for both the tethers is damped to zero as the tether deploys in
confirmation with the principle of conservation of angular momentum whereas the
platform response in roll increases to around 20deg Subsatellite 2 remains virtually
unaffected by the other links since the tether is attached to its centre of mass thus
eliminating coupling Finally the flexibility response of the two tethers presented in
Figure 4-12(b) shows similarity with the three-body case
47 BICEPS Configuration
The mission profile of the Bl-static Canadian Experiment on Plasmas in Space
(BICEPS) is simulated next It is presently under consideration by the Canadian
Space Agency To be launched by the Black Brant 2000500 rocket it would inshy
volve interesting maneuvers of the payload before it acquires the final operational
configuration As shown in Figure 1-3 at launch the payload (two satellites) with
an undeployed tether is spinning about the orbit normal (phase 1) The internal
dampers next change the motion to a flat spin (phase 2) which in turn provides
momentum for deployment (phase 3) When fully deployed the one kilometre long
tether will connect two identical satellites carrying instrumentation video cameras
and transmitters as payload
The system parameters for the BICEPS configuration are summarized below [59 0 0 1
bull I = I2 = 0 366 0 kg bull m 2 (satellite inertias) [ 0 0 392_
bull mi = 7772 = 200 kg (mass of the satellites)
80
STSS configuration 5-Body a(0)=cx2(0)=al1(0)=2deg a3(0)=c (0)=25o
P1(0)=P2(0)=p3(0)=f3t1(0)=Pt2(0)=r 52(0)=84(0)=1304x 10300m d2(0)=d4(0)=1G0m Deployment lt1=le=02km to 20km in 35 Orbits
Pitch Roll Satellite 1 Pitch Angle
Time (Orbits) Satellite 2 Pitch and Roll Angle T 1mdash1 1 1 I I I I | I I I 1 I | I I lt~r
Tether Length Profile i 1 1 1 1 i 1 1 1 1 i 1 1 1 1 i
i i i i i 0 1 2 3 4 5
Time (Orbits) Satellite 1 Roll Angle
0 1 2 3 4 5 Time (Orbits)
Tether 1 amp 2 Roll Angle L i | I I 1 I 1 1 -I
-Ih I- bull I I i i i i I i i i i 1
0 1 2 3 4 5 Time (Orbits)
Satellite 3 Pitch and Roll Angle i- i 1 1 i 1 1 1 1 i 1 1 1 1 i 1 1 1 1 i
L i i i i i i 1 i i mdash i mdash i I i imdashimdashLJ
0 1 2 3 4 5 Time (Orbits)
Figure 4-12 Deployment dynamics of the five-body S T S S configuration with offshyset along the local vertical (a) attitude response
81
STSS configuration 5-Body a1(0)=a2(0)=oct1(0)=2deg a3(0)=at2(0)=25deg P1(0)=P2(0)=P3(0)=Pt1(0)=pt2(0)=1o
82(0)=54(0)=1304 x 10-300m d2(0)=d4(0)=100m Deployment lt1=l2=02km to 20km in 35 Orbits
Tether 1 Longitudinal Vibration Tether 2 Longitudinal Vibration
Time (Orbits) Time (Orbits) Tether 1 Z Tranverse Vibration Tether 2 Z Tranverse Vibration
Figure 4-12 Deployment dynamics of the five-body STSS configuration with offshyset along the local vertical (b) vibration response
82
bull EtAt = 61645 N (tether stiffness)
bull pt = 30 k g k m (tether density)
bull It = 1 km (tether length)
bull rid = 1-0 (tether structural damping coefficient)
The system is in a circular orbit at a 289 km altitude Offset of the tether
attachment point to the satellites at both ends is taken to be 078 m in the x
direction The response of the system in the stationkeeping phase for a prescribed set
of in i t ia l conditions is given by Figure 4-13 As in the case of the STSS configuration
there is strong coupling between the tether and the satellites rigid body dynamics
causing the latter to follow the attitude of the tether However because of the smaller
inertias of the payloads there are noticeable high frequency modulations arising from
the tether flexibility Response of the tether in the elastic degrees of freedom is
presented in Figure 4-13(b) Note the longitudinal vibrations decay quite rapidly
(lt 05 orbits) due to the structural damping however it has vir tual ly no effect on
the transverse oscillations
The unique mission requirement of B I C E P S is the proposed use of its ini t ia l
angular momentum in the cartwheeling mode to aid in the deployment of the tethered
system The maneuver is considered next The response of the system during this
maneuver wi th an ini t ia l cartwheeling rate of 5 deg s is illustrated in Figure 4-14(a) The
in i t ia l tether length is taken to be 10 m and is deployed to 1 km There is an in i t ia l
increase in the pitch motion of the tether however due to the conservation of angular
momentum the cartwheeling rate decreases proportionally to the square of the change
in tether deployment rate The result is a rapid drop in the cartwheeling rate unti l
the system stops rotating entirely and simply oscillates about its new equilibrium
point Consequently through coupling the end-bodies follow the same trend The
roll response also subsides once the cartwheeling motion ceases However
83
BICEPS configuration 3-Body a1(0)=o2(0)=at(0)=2deg P1(0)=P2(0)=pt(0)=1deg 8X(0)=001 m 6y(0)=52(0)=01 m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping l=1 km
Satellite 1 Pitch Angle Satellite 1 Roll Angle - i 1 1 1 1 i - i q r 1 1 -gt
I- I i t i i i i I H 0 1 2 0 1 2
Time (Orbits) Time (Orbits) Tether Pitch Angle Tether Roll Angle
F mdash 1 mdash 1 mdash mdash 1 mdash i mdash 1 mdash 1 mdash mdash 1 mdash = i F mdash 1 mdash mdash mdash 1 mdash i mdash 1 mdash 1 mdash 1 mdash 1 mdash q
(a) Time (Orbits) Time (Orbits)
Figure 4-13 Stationkeeping dynamics of the three-body BICEPS configuration with offset along the local vertical (a) attitude response
84
BICEPS configuration 3-Body a1(0)=o2(0)=at(0)=2deg Pl(0)=p2(0)=(3t(0)=1deg 8X(0)=001 m 8y(0)=82(0)=01 m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping lt=1km
Tether Longitudinal Vibration 1E-2[
laquo = 5 E - 3 |
000 002
Time (Orbits) Tether In-Plane Transverse Vibration
Time (Orbits) Tether Out-of-Plane Transverse Vibration
Time (Orbits)
Figure 4-13 Stationkeeping dynamics of the three-body BICEPS configuration with offset along the local vertical (b) vibration response
85
BICEPS configuration 3-Body a1(0)=o2(0)=at(0)=2deg a1(0)=a2(0)=at(0)=57s Pl(0)=P2(0)=(3t(0)=1deg 8X(0)=115 x 103m 8y(0)=82(0)=0 dx(0)=078m dy(0)=d2(0)=0 Cartwheeling Deployment
lt=10m to 1km in 1 Orbit
Tether Length Profile
E
2000
cu T3
2000
co CD
2000
C D CD
T J
Satellite 1 Pitch Angle
1 2 Time (Orbits)
Satellite 1 Roll Angle
cn CD
33 cdl
Time (Orbits) Tether Pitch Angle
Time (Orbits) Tether Roll Angle
Time (Orbits) Satellite 2 Pitch Angle
1 2 Time (Orbits)
Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-14 Cartwheeling dynamics with deployment for the three-body BICEPS with offset along the local vertical (a) attitude response
86
BICEPS configuration 3-Body a1(0)=a2(0)=at(0)=2deg a (0)=a2(0)=a(0)=57s P1(0)=P2(0)=f3(0)=1deg 8X(0)=115x103m 8y(0)=8z(0)=0 dx(0)=078m dy(0)=d2(0)=0 Cartwheeling Deployment
l=1 Om to 1 km in 1 Orbit
5E-3h
co
0E0
025
000 E to
-025
002
-002
(b)
Figure 4-14
Tether Longitudinal Vibration
Time (Orbits) Tether In-Plane Transverse Vibration
Time (Orbits) Tether Out-of-Plane Transverse Vibration
Time (Orbits)
Cartwheeling dynamics with deployment for the three-body BICEPS with offset along the local vertical (b) vibration response
87
as the deployment maneuver is completed the satellites continue to exhibit a small
amplitude low frequency roll response with small period perturbations induced by
the tether elastic oscillations superposed on it (Figure 4-14b)
48 OEDIPUS Spinning Configuration
The mission entitled Observation of Electrified Distributions of Ionospheric
Plasmas - A Unique Strategy ( O E D I P U S ) also consists of a 1 k m tethered system with
two similar end-masses It was first flown in a suborbital flight in 1989 ( O E D I P U S A )
followed by a recent study in November 1995 ( O E D I P U S - C ) in which the University
of Br i t i sh Columbia was one of the participants Dynamically O E D I P U S represents a
unique system never encountered before Launched by a Black Brant rocket developed
at Br is to l Aerospace L t d the system ie the end-bodies together wi th the tether
spins about the longitudinal axis of the tether to achieve stabilized alignment wi th
Earth s magnetic field
The response of the system undergoing a spin rate of j = l deg s is presented
in Figure 4-15 using the same parameters as those outlined for the B I C E P S conshy
figuration in Section 47 Again there is strong coupling between the three bodies
(two satellites and tether) due to the nonzero offset The spin motion introduces an
additional frequency component in the tethers elastic response which is transferred
to the l ibrational motion of the satellites A s the spin rate increases to 1 0 deg s (Figure
4-16) the amplitude and frequency of the perturbations also increase however the
general character of the response remains essentially the same
88
OEDIPUS configuration 3-Body a1(0)=a2(0)=at(0)=2deg i(0)^[2(0H(0)=Ys p1(0)=p2(0)=p(0)=r 8X(0)=12 x 103m 8y(0)=82(0)=01m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping l=1 km
Satellite 1 Pitch Angle Satellite 1 Roll Angle mdash i 1 1 1 a r 1 1 1 r
(a) Time (Orbits) Time (Orbits)
Figure 4-15 Spin dynamics (7 = ldegs) of the three-body OEDIPUS configuration with offset along the local vertical (a) attitude response
89
OEDIPUS configuration 3-Body a1(0)=a2(0)=cct(0)=2deg Y1(0)=Y2(0)=Yt(0)=l7s (31(0)=p2(0)=pt(0)=1deg 8X(0)=12 x 103m 8y(0)=5z(0)=01 m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping lt=1 km
1E-2
to
OEO
E 00
to
Tether Longitudinal Vibration
005
Time (Orbits) Tether In-Plane Transverse Vibration
Time (Orbits) Tether Out-of-Plane Transverse Vibration
Time (Orbits)
Figure 4-15 Spin dynamics (7 = l deg s ) of the three-body O E D I P U S configuration with offset along the local vertical (b) vibration response
90
OEDIPUS configuration 3-Body a1(0)=a2(0)=a(0)=2deg Yi(0)=Y2(0)=rt(0)=107s 81(0)=p2(0)=pt(0)=1deg 8X(0)=12 x 103m 8y(0)=82(0)=01m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping lt=1km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-16 Spin dynamics (7= 10degs) of the three-body OEDIPUS configura tion with offset along the local vertical (a) attitude response
91
OEDIPUS configuration 3-Body a1(0)=a2(0)=ot(0)=2deg Yi(0)H(0)4(0)laquo107s P1(0)=P2(0)=pt(0)=1deg 8X(0)=12 x 103m 8y(0)=82(0)=01m dx(0)=078m dy(0)=dz(0)=0 Stationkeeping lt=1 km
Tether Longitudinal Vibration
Figure 4-16 Spin dynamics (7 = 10degs) of the three-body OEDIPUS configurashytion with offset along the local vertical (b) vibration response
92
5 ATTITUDE AND VIBRATION CONTROL
51 Att itude Control
511 Preliminary remarks
The instability in the pitch and roll motions during the retrieval of the tether
the large amplitude librations during its deployment and the marginal stability during
the stationkeeping phase suggest that some form of active control is necessary to
satisfy the mission requirements This section focuses on the design of an attitude
controller with the objective to regulate the librational dynamics of the system
As discussed in Chapter 1 a number of methods are available to accomplish
this objective This includes a wide variety of linear and nonlinear control strategies
which can be applied in conjuction with tension thrusters offset momentum-wheels
etc and their hybrid combinations One may consider a finite number of Linear T ime
Invariant (LTI) controllers scheduled discretely over different system configurations
ie gain scheduling[43] This would be relevant during deployment and retrieval of
the tether where the configuration is changing with time A n alternative may be
to use a Linear T ime Varying ( L T V ) model in conjuction with an adaptive control
scheme where on-line parametric identification may be used to advantage[56] The
options are vir tually limitless
O f course the choice of control algorithms is governed by several important
factors effective for time-varying configurations computationally efficient for realshy
time implementation and simple in character Here the thrustermomentum-wheel
system in conjuction with the Feedback Linearization Technique ( F L T ) is chosen as
it accounts for the complete nonlinear dynamics of the system and promises to have
good overall performance over a wide range of tether lengths It is well suited for
93
highly time-varying systems whose dynamics can be modelled accurately as in the
present case
The proposed control method utilizes the thrusters located on each rigid satelshy
lite excluding the first one (platform) to regulate the pitch and rol l motion of the
tether to which it is attached ie the thrusters located on satellite 2 (link 3) regshy
ulates the attitude motion of tether 1 (link 2) O n the other hand the l ibrational
motion of the rigid bodies (platform and subsatellite) is controlled using a set of three
momentum wheels placed mutually perpendicular to each other
The F L T method is based on the transformation of the nonlinear time-varying
governing equations into a L T I system using a nonlinear time-varying feedback Deshy
tailed mathematical background to the method and associated design procedure are
discussed by several authors[435758] The resulting L T I system can be regulated
using any of the numerous linear control algorithms available in the literature In the
present study a simple P D controller is adopted arid is found to have good perforshy
mance
The choice of an F L T control scheme satisfies one of the criteria mentioned
earlier namely valid over a wide range of tether lengths The question of compushy
tational efficiency of the method wi l l have to be addressed In order to implement
this controller in real-time the computation of the systems inverse dynamics must
be executed quickly Hence a simpler model that performs well is desirable Here
the model based on the rigid system with nonlinear equations of motion is chosen
Of course its validity is assessed using the original nonlinear system that accounts
for the tether flexibility
512 Controller design using Feedback Linearization Technique
The control model used is based on the rigid system with the governing equa-
94
tions of motion given by
Mrqr + fr = Quu (51)
where the left hand side represents inertia and other forces while the right hand side
is the generalized external force due to thrusters and momentum-wheels and is given
by Eq(2109)
Lett ing
~fr = fr~ QuU (52)
and substituting in Eq(296)
ir = s(frMrdc) (53)
Expanding Eq(53) it can be shown that
qr = S(fr Mrdc) - S(QU Mr6)u (5-4)
mdash = Fr - Qru
where S(QU Mr0) is the column matrix given by
S(QuMr6) = [s(Qu(l)Mr0) 5 (Q u ( 2 ) M r | 0 ) S(QU nu) M r | 0 ) ] (55)
and Qu-i) is the i ^ column of Qu Extract ing only the controlled equations from
Eq(54) ie the attitude equations one has
Ire = Frc ~~ Qrcu
(56)
mdash vrci
where vrc is the new control input required to regulate the decoupled linear system
A t this point a simple P D controller can be applied ie
Vrc = qrcd +KvQrcd-Qrc) + Kp(Qrcd~ Qrc)- (57)
Eq(57) represents the secondary controller with Eq(54) as the primary controller
Kp and Kv are the proportional and derivative gain matrices with qrcdi Qrcd
an(^ qrcd
95
as the desired acceleration velocity and position vector for the attitude angles of each
controlled body respectively Solving for u from Eq(56) leads to
u Qrc Frc ~ vrcj bull (58)
5 13 Simulation results
The F L T controller is implemented on the three-body STSS tethered system
The choice of proportional and derivative matrix gains is based on a desired settling
time of ts mdash 05 orbit and a damping factor of = 07 for both the pitch and roll
actuators in each body Given O and ts it can be shown that
-In ) 0 5 V T ^ r (59)
and hence kn bull mdash10
V (510)
where kp and kVj are the diagonal elements of the gain matrices Kp and Kv in
Eq(57) respectively[59]
Note as each body-fixed frame is referred directly to the inertial frame FQ
the nominal pitch equilibrium angle is zero only when measured from the L V L H frame
(OJJ mdash 0) However it is 6 when referred to FQ Hence the desired position vector
qrC(l is set equal to 0(t) ie the time-varying orbital angle for the pitch motion
and zero for the roll and yaw rotations such that
qrcd 37Vxl (511)
96
The desired velocity and acceleration are then given as
e U 3 N x l (512)
E K 3 7 V x l (513)
When the system is in a circular orbit qrc^ = 0
Figure 5-1 presents the controlled response of the STSS system defined in
Chapter 4 in the stationkeeping phase with offset d2 = 111T m and If = 20
km A s mentioned earlier the F L T controller is based on the nonlinear rigid model
Note the pitch and roll angles of the rigid bodies as well as of the tether are now
attenuated in less than 1 orbit (Figure 5-la) Consequently the longitudinal vibration
response of the tether is free of coupling arising from the librational modes of the
tether This leaves only the vibration modes which are eventually damped through
structural damping The pitch and roll dynamics of the platform require relatively
higher control moments Ma and Mpi respectively (Figure 5-lb) since they have to
overcome the extra moments generated by the offset of the tether attachment point
Furthermore the platform demands an additional moment to maintain its orientation
at zero pitch and roll angle since it is not the nominal equilibrium position The nonshy
zero static equilibrium of the platform arises due to its products of inertia On the
other hand the tether pitch and roll dynamics require only a small ini t ia l control
force of about plusmn 1 N which eventually diminishes to zero once the system is
97
and
o 0
Qrcj mdash
o o
0 0
Oiit) 0 0
1
N
N
1deg I o
STSS configuration 3-Body a1(0)=a2(0)=at(0)=2lt
P1(0)=p2(0)=pt(0)=1 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=dz(0)=1m Controlled Stationkeeping lt=20km
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
05 10 15 Time (Orbits)
Sat 2 Pitch and Roll Angles
20
CO CD
T J
cdT
cT
- mdash mdash 1 1 1 1 1 I - 1
Pitch -
A Roll -
A -
1 bull bull
-
1 Time (Orbits)
Tether Y Tranverse Vibration
1 2 Time (Orbits)
Tether Z Tranverse Vibration
gt to N
to
Time (Orbits) Time (Orbits) Longitudinal Vibration
145h
(a) Time (Orbits)
Figure 5-1 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local horizontal local vertical and orbit normal (a) attitude and vibration response
98
1
STSS configuration 3-Body a1(0)=ct2(0)=cct(0)=20
P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=d2(0)=1m Controlled Stationkeeping lt=20km
Sat 1 Pitch Control Moment Sat 1 Roll Control Moment
Figure 5-1 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local horizontal local vertical and orbit normal (b) control actuator time histories
99
stabilized Similarly for the end-body a very small control torque is required to
attenuate the pitch and roll response
If the complete nonlinear flexible dynamics model is used in the feedback loop
the response performance was found to be virtually identical with minor differences
as shown in Figure 5-2 For example now the pitch and roll response of the platform
is exactly damped to zero with no steady-state error as against the minor almost
imperceptible deviation for the case of the controller basedon the rigid model (Figure
5-la) In addition the rigid controller introduces additional damping in the transverse
mode of vibration where none is present when the full flexible controller is used
This is due to a high frequency component in the at motion that slowly decays the
transverse motion through coupling
As expected the full nonlinear flexible controller which now accounts the
elastic degrees of freedom in the model introduces larger fluctuations in the control
requirement for each actuator except at the subsatellite which is not coupled to the
tether since f c m 3 = 0 The high frequency variations in the pitch and roll control
moments at the platform are due to longitudinal oscillations of the tether and the
associated changes in the tension Despite neglecting the flexible terms the overall
controlled performance of the system remains quite good Hence the feedback of the
flexible motion is not considered in subsequent analysis
When the tether offset at the platform is restricted to only 1 m along the x
direction a similar response is obtained for the system However in this case the
steady-state error in the platforms rigid body motion is much smaller (Figure 5-3a)
In addition from Figure 5-3(b) the platforms control requirement is significantly
reduced
Figure 5-4 presents the controlled response of the STSS deploying a tether
100
STSS configuration 3-Body a1(0)=a2(0)=oct(0)=2deg p1(0)=P2(0)=p(0)=1deg 8x(0)=14m5y(0)=82(0)=1m dx(0)=dy(0)=dz(0)=1m Controlled Stationkeeping l=20km
Tether Pitch and Roll Angles
c n
T J
cdT
Sat 1 Pitch and Roll Angles 1 r
00 05 10 15 Time (Orbits)
Sat 2 Pitch and Roll Angles
c n CD
cdT a
o T J
CQ
- - I 1 1 1 I 1 1 1 1
P i t c h
R o l l
bull
1 -
0 1 2 Time (Orbits)
Tether Y Tranverse Vibration
0 1 2 Time (Orbits)
Tether Z Tranverse Vibration
E gt
10 N
CO
Time (Orbits) Time (Orbits) Longitudinal Vibration
145h
(a) Time (Orbits)
Figure 5-2 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear flexible FLT controller with offset along the local horizontal local vertical and orbit normal (a) attitude and vibration response
101
STSS configuration 3-Body a1(0)=a2(0)=cxt(0)=2deg 31(0)=p2(0)=p(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=d2(0)=1m Controlled Stationkeeping lt=20km
r i i i t i mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash mdash J 0 1 2 0 1 2
Time (Orbits) Time (Orbits) Tether Pitch Control Thrust Tether Roll Control Thruster
(b) Time (Orbits) Time (Orbits)
Figure 5-2 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear flexible FLT controller with offset along the local horizontal local vertical and orbit normal (b) control actuator time histories
102
STSS configuration 3-Body a1(0)=o2(0)=a(0)=2deg P1(0)=p2(0)=(3t(0)=1deg 8x(0)=14m5y(0)=6z(0)=1m dx(0)=1mdy(0)=dz(0)=0 Controlled Stationkeeping lt=20km
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
05 10 15 Time (Orbits)
Sat 2 Pitch and Roll Angles
CD
eg CQ
0 1 2 Time (Orbits)
Tether Y Tranverse Vibration
1 2 Time (Orbits)
Tether Z Tranverse Vibration
gt CO
Time (Orbits) Time (Orbits) Longitudinal Vibration
145h
(a) Time (Orbits)
Figure 5-3 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local vertical (a) attitude and vibration response
103
STSS configuration 3-Body a1(0)=o2(0)=al(0)=2o
P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=1mdy(0)=dz(0)=0 Controlled Stationkeeping lt=20km
Sat 1 Pitch Control Moment 11 1 1 r
Sat 1 Roll Control Moment
0 1 2 Time (Orbits)
Tether Pitch Control Thrust
0 1 2 Time (Orbits)
Sat 2 Pitch Control Moment
0 1 2 Time (Orbits)
Tether Roll Control Thruster
-07 o 1
Time (Orbits Sat 2 Roll Control Moment
1E-5
0E0F
Time (Orbits)
Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local vertical (b) control actuator time histories
104
STSS configuration 3-Body a1(0)=a2(0)=a(0)=2deg Pl(0)=p2(0)=pt(0)=1deg 6X(0)=1304 x 103m 8y(0)=82(0)=0 dx(0)=1mdy(0)=d2(0)=0 Controlled Deployment
l=02km to 20km in 35 Orbits
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
o 1 2 Time (Orbits)
Tether Y Tranverse Vibration
0 1 2 Time (Orbits)
Sat 2 Pitch and Roll Angles
CD CD
CM CO
Time (Orbits) Tether Z Tranverse Vibration
to
150
100
1 2 3 0 1 2
Time (Orbits) Time (Orbits) Longitudinal Vibration
to
(a)
50
2 3 Time (Orbits)
Figure 5 - 4 Deployment dynamics of the three-body STSS using the nonlinear rigid FLT controller with offset along the local vertical (a) attitude and vibration response
105
STSS configuration 3-Body a1(0)=oc2(0)=al(0)=20
P1(0)=p2(0)=pt(0)=1deg 8X(0)=1304 x 103m 5y(0)=52(0)=0 dx(0)=1mdy(0)=dz(0)=0 Controlled Deployment
l=02km to 20km in 35 Orbits
Tether Length Profile i 1
2 0 h
E 2L
Sat 1 Pitch Control Moment t mdash r mdash i mdash i mdash | mdash i mdash i mdash i mdash i mdash | mdash r mdash i mdash i mdash i mdash | mdash r mdash
0 5
1 2 3
Time (Orbits) Sat 1 Roll Control Moment
i i t | t mdash t mdash i mdash [ i i mdash r -
1 2
Time (Orbits) Tether Pitch Control Thrust
1 2 3 Time (Orbits)
Tether Roll Control Thruster
2 E - 5
1 2 3
Time (Orbits) Sat 2 Pitch Control Moment
1 2
Time (Orbits) Sat 2 Roll Control Moment
1 E - 5
E
O E O
Time (Orbits)
Deployment dynamics of the three-body STSS using the nonlinear rigid FLT controller with offset along the local vertical (b) control actuator time histories
106
from 02 km to 20 km in 35 orbits with a i m offset along the tether length A s before
the systems attitude motion is well regulated by the controller However the control
cost increases significantly for the pitch motion of the platform and tether due to the
Coriolis force induced by the deployment maneuver (Figure 5-4b) Initially there is a
sinusoidal increase in the pitch control requirement for both the tether and platform
as the tether accelerates to its constant velocity VQ Then the control requirement
remains constant for the platform at about 60 N m as opposed to the tether where the
thrust demand increases linearly Finally when the tether decelerates the actuators
control input reduces sinusoidally back to zero The rest of the control inputs remain
essentially the same as those in the stationkeeping case In fact the tether pitch
control moment is significantly less since the tether is short during the ini t ia l stages
of control However the inplane thruster requirement Tat acts as a disturbance and
causes 6y to nearly double from the uncontrolled value (Figure 4-8a) O n the other
hand the tether has no out-of-plane deflection
Final ly the effectiveness of the F L T controller during the crit ical maneuver of
retrieval from 20 km to 02 k m in 35 orbits is assessed in Figure 5-5 A s in the earlier
cases the system response in pitch and roll is acceptable however the controller is
unable to suppress the high frequency elastic oscillations induced in the tether by the
retrieval O f course this is expected as there is no active control of the elastic degrees
of freedom However the control of the tether vibrations is discussed in detail in
Section 52 The pitch control requirement follows a similar trend in magnitude as
in the case of deployment with only minor differences due to the addition of offset in
the out-of-plane direction ( d 2 = 1 0 1 ^ m) This is also responsible for the higher
roll moment requirement for the platform control (Figure 5-5b)
107
STSS configuration 3-Body a1(0)=X2(0)=a(0)=2deg P1(0)=p2(0)=pt(0)=1deg 5x(0)=14m8y(0)=52(0)=0 dx(0)=1mdy(0)=dz(0)=0 Controlled Retrieval
l=20km to 02km in 35 Orbits
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
0 1 2 Time (Orbits)
Sat 2 Pitch and Roll Angles
CO CD
cdT
CO CD
T J CM
CO
1 2 Time (Orbits)
Tether Y Tranverse Vibration
Time (Orbits) Tether Z Tranverse Vibration
150
100
1 2 3 0 1 2 Time (Orbits) Time (Orbits)
Longitudinal Vibration
CO
(a)
x 50
2 3 Time (Orbits)
Figure 5 -5 Retrieval dynamics of the three-body STSS using the non-linear rigid FLT controller with offset along the local vertical and orbit normal (a) attitude and vibration response
108
STSS configuration 3-Body a1(0)=ct2(0)=at(0)=2deg p1(0)=p2(0)=(3(0)=1deg 5x(0)=14m5y(0)=82(0)=0 dx(0)=1mdy(0)=dz(0)=0 Controlled Retrieval
l=20km to 02km in 35 Orbits
Tether Length Profile
E
1 2 Time (Orbits)
Tether Pitch Control Thrust
2E-5
1 2 Time (Orbits)
Sat 2 Pitch Control Moment
OEOft
1 2 3 Time (Orbits)
Sat 1 Roll Control Moment
1 2 3 Time (Orbits)
Tether Roll Control Thruster
1 2 Time (Orbits)
Sat 2 Roll Control Moment
1E-5 E z
2
0E0
(b) Time (Orbits) Time (Orbits)
Figure 5-5 Retrieval dynamics of the three-body STSS using the non-linear rigid FLT controller with offset along the local vertical and orbit normal (b) control actuator time histories
109
52 Control of Tethers Elastic Vibrations
521 Prel iminary remarks
The requirement of a precisely controlled micro-gravity environment as well
as the accurate release of satellites into their final orbit demands additional control
of the tethers vibratory motion in addition to its attitude regulation To that end
an active vibration suppression strategy is designed and implemented in this section
The strategy adopted is based on offset control ie time dependent variation
of the tethers attachment point at the platform (satellite 1) A l l the three degrees
of freedom of the offset motion are used to control both the longitudinal as well as
the inplane and out-of-plane transverse modes of vibration In practice the offset
controller can be implemented through the motion of a dedicated manipulator or a
robotic arm supporting the tether which in turn is supported by the platform The
focus here is on the control of elastic deformations during stationkeeping (ie fully
deployed tether fixed length) as it represents the phase when the mission objectives
are carried out
This section begins with the linearization of the systems equations of motion
for the reduced three body stationkeeping case This is followed by the design of
the optimal control algorithm Linear Quadratic Gaussian-Loop Transfer Recovery
( L Q G L T R ) based on the reduced model and its implementation on the full nonlinear
model Final ly the systems response in the presence of offset control is presented
which tends to substantiate its effectiveness
522 System linearization and state-space realization
The design of the offset controller begins with the linearization of the equations
of motion about the systems equilibrium position Linearization of extremely lengthy
(even in matrix form) highly nonlinear nonautonomous and coupled equations of
110
motion presents a challenging problem This is further complicated by the fact the
pitch angle ot is not referred to the L V L H frame Hence the equilibrium pitch angle
is not a constant but is equal to the instantaneous orbital angle 9(t) Two methods
are available to resolve this problem
One may use the non-stationary equations of motion in their present form and
derive a controller based on the Linear T ime Varying ( L T V ) system Alternatively
one can design a Linear T ime Invariant (LTI) controller based on a new set of reduced
governing equations representing the motion of the three body tethered system A l shy
though several studies pertaining to L T V systems have been reported[5660] the latter
approach is chosen because of its simplicity Furthermore the L T I controller design
is carried out completely off-line and thus the procedure is computationally more
efficient
The reduced model is derived using the Lagrangian approach with the genershy
alized coordinates given by
where lty and are the pitch and roll angles of link i relative to the L V L H frame
respectively 5X 8y and 5Z are the generalized coordinates associated with the flexible
tether deformations in the longitudinal inplane and out-of-plane transverse direcshy
tions respectively Only the first mode of vibration is considered in the analysis
The nonlinear nonautonomous and coupled equations of motion for the tethshy
ered system can now be written as
Qred = [ltXlPlltX2P2fixSy5za302gt] (514)
M redQred + fr mdash 0gt (515)
where Mred and frec[ are the reduced mass matrix and forcing term of the system
111
respectively They are functions of qred and qred in addition to the time varying offset
position d2 and its time derivatives d2 and d2 A detailed derivation of Eq(515)
is given in Appendix II A n additional consequence of referring the pitch motion to
a local frame is that the new reduced equations are now independent of the orbital
angle 9 and under the further assumption of the system negotiating a circular orbit
91 remains constant
The nonlinear reduced equations can now be linearized about their static equishy
l ibr ium position For all the generalized coordinates the equilibrium position is zero
wi th the exception of 5X which has a non-zero equilibrium 5XQ The offset position
d2) is linearized about its ini t ial position d2^ Linearizing Eq(515) and recasting
into matr ix form gives
Msqred + CsQred + KsQred + Md^2 + Cdd2 + Kdd2 + fs = 0 (516)
where Ms Cs Ks Md Cd Kd and s are constant matrices Mul t ip ly ing throughout
by Ms
1 gives
qr = -Ms 1Csqr - Ms
lKsqr
- Ms~lCdd22 - Ms-Kdd2 - Ms~Mdd2 - Mg~1 fs r-1 (517)
Defining ud = d2 and v = qTd^T Eq(517) can be rewritten as
Pu t t ing
v =
+
-M~lCs -M~Cd -1
0
-Mg~lMd
Id
0 v + -M~lKs -M~lKd
0 0
-M-lfs
0
MCv + MKv + Mlud +Fs
- 4 ) - ( gt the L T I equations of motion can be recast into the state-space form as
(518)
(519)
x =Ax + Bud + Fd (520)
112
where
A = MC MK Id12 o
24x24
and
Let
B = ^ U 5 R 2 4 X 3 -
Fd = I FJ ) e f t 2 4 1
mdash _ i _ -5
(521)
(522)
(523)
(524)
where x is the perturbation vector from the constant equilibrium state vector xeq
Substituting Eq(524) into Eq(520) gives the linear perturbation state equation
modelling the reduced tethered system as
bullJ x mdash x mdash Ax + Bud + (Fd + Axeq) (525)
It can be shown that for ideal control[60] Fd + Axeq = 0 thus giving the familiar
state-space equation
S=Ax + Bud (526)
The selection of the output vector completes the state space realization of
the system The output vector consists of the longitudinal deformation from the
equil ibrium position SXQ of the tether at xt = h the slope of the tether due to the
transverse deformation at xt = 0 and the offset position d2 from its in i t ia l position
d2Q Thus the output vector y is given by
K(o)sx
V
dx ~ dXQ dy ~ dyQ
d z - d z0 J
Jy
dx da XQ vo
d z - dZQ ) dy dyQ
(527)
113
where lt ( 0 ) = ddxi[$vxi)]Xi=o and lt ( 0 ) = ddxi[4gtw(xi)]Xi=0
523 Linear Quadratic Gaussian control with Loop Transfer Recovery
W i t h the linear state space model defined (Eqs526527) the design of the
controller can commence The algorithm chosen is the Linear Quadratic Gaussian
( L Q G ) estimator based optimal controller[6061] The L Q G is a widely used optimal
controller with its theoretical background well developped by many authors over the
last 25 years It involves of the design of the Kalman-Bucy Fi l ter ( K B F ) which
provides an estimate of the states x and a Linear Quadratic Regulator ( L Q R ) which
is separately designed assuming all the states x are known Both the L Q R and K B F
designs independently have good robustness properties ie retain good performance
when disturbances due to model uncertainty are included However the combined
L Q R and K B F designs ie the L Q G design may have poor stability margins in the
presence of model uncertainties This l imitat ion has led to the development of an L Q G
design procedure that improves the performance of the compensator by recovering
the full state feedback robustness properties at the plant input or output (Figure 5-6)
This procedure is known as the Linear Quadratic Guassian-Loop Transfer Recovery
( L Q G L T R ) control algorithm A detailed development of its theory is given in
Ref[61] and hence is not repeated here for conciseness
The design of the L Q G L T R controller involves the repeated solution of comshy
plicated matrix equations too tedious to be executed by hand Fortunenately the enshy
tire algorithm is available in the Robust Control Toolbox[62] of the popular software
package M A T L A B The input matrices required for the function are the following
(i) state space matrices ABC and D (D = 0 for this system)
(ii) state and measurement noise covariance matrices E and 0 respectively
(iii) state and input weighting matrix Q and R respectively
114
u
MODEL UNCERTAINTY
SYSTEM PLANT
y
u LQR CONTROLLER
X KBF ESTIMATOR
u
y
LQG COMPENSATOR
PLANT
X = f v ( X t )
COMPENSATOR
x = A k x - B k y
ud
= C k f
y
Figure 5-6 Block diagram for the LQGLTR estimator based compensator
115
A s mentionned earlier the main objective of this offset controller is to regu-mdash
late the tether vibration described by the 5 equations However from the response
of the uncontrolled system presented in Chapter 4 it is clear that there is a large
difference between the magnitude of the librational and vibrational frequencies This
separation of frequencies allow for the separate design of the vibration and attitude mdash mdash
controllers Thus only the flexible subsystem composed of the S and d equations is
required in the offset controller design Similarly there is also a wide separation of
frequencies between the longitudinal and transverse modes of vibrations permitt ing mdash
the decoupling of the flexible subsystem into a set of longitudinal (Sx and dx) and
transverse (5y Sz dy and dz) subsystems The appended offset system d must also
be included since it acts as the control actuator However it is important to note
that the inplane and out-of-plane transverse modes can not be decoupled because
their oscillation frequencies are of the same order The two flexible subsystems are
summarized below
(i) Longitudinal Vibra t ion Subsystem
This is defined by u mdash Auxu + Buudu
(528)
where xu = SxdX25xdX2T u d u = dX2 and from Eq(527)
0 0 $u(lt) 0 0 0 0 1
0 0 1 0 0 0 0 1
(529)
Au and Bu are the rows and columns of A and B respectively and correspond to the
components of xu
The L Q R state weighting matrix Qu is taken as
Qu =
bull1 0 0 o -0 1 0 0 0 0 1 0
0 0 0 10
(530)
116
and the input weighting matrix Ru = 1^ The state noise covariance matr ix is
selected as 1 0 0 0 0 1 0 0 0 0 4 0
LO 0 0 1
while the measurement noise covariance matrix is taken as
(531)
i o 0 15
(532)
Given the above mentionned matrices the design of the L Q G L T R compenshy
sator is computed using M A T L A B with the following 2 commands
(i) kfu = l q r c ( ^ ( i C bdquo d i a g m x ( E u e u ) )
(ii) [AkugtBkugtCkugtDku =^ryAuBuCuDukfuQuRur)
where r is a scaler
The first command is used to compute the Ka lman filter gain matrix kfu Once
kfu is known the second command is invoked returning the state space representation
of the L Q G L T R dynamic compensator to
(533) xu mdash Akuxu Bkuyu
where xu is the state estimate vector oixu The function l t ry performs loop transfer
recovery at the system output ie the return ratio at the output approaches that
of the K B F loop given by Cu(sld - Au)Kfu[61 This is achieved by choosing a
sufficently large scaler value r in l t ry such that the singular values of the return
ratio approach those of the target design For the longitudinal controller design
r mdash 5 x 10 5 Figure 5-7(a) compares the singular values of the recovered compensator
design and the non-recovered L Q G compensator design with respect to the target
Unfortunetaly perfect recovery is not possible especially at higher frequencies
117
100
co
gt CO
-50
-100 10
Longitudinal Compensator Singular Values i I I I N I I mdash i mdash i i i i n i i mdash i mdash i i i i inimdash imdashi i i i i i i i mdash i mdash i i M i I I
Target r = 5x10 5
r = 0 (LQG)
i i i i I I i ii i i i 1 1 1 1 i i i i i 1 1 1 1 ii i I I i i i i 11111 -3 10 -2
(a)
101 10deg Frequency (rads)
101 10
Transverse Compensator Singular Values
100
50
X3
gt 0
CO
-50
bull100 10
-imdashi i 111II i 111 ni 1mdashi I I I I I I I 1mdashi I I I N I I 1mdashi i M 111
Target r = 50 r = 0 (LQG)
m i i i i i i n i i i i I I I I I I I 1mdashi I I I I I I I 1mdashi i 111 n -3 10 -2
(b)
101 10deg Frequency (rads)
101 10
Figure 5-7 Singular values for the LQG and LQGLTR compensator compared to target return ratio (a) longitudinal design (b) transverse design
118
because the system is non-minimal ie it has transmission zeros with positive real
parts[63]
(ii) Transverse Vibra t ion Subsystem
Here Xy mdash AyXy + ByU(lv
Dv = Cvxv
with xv = 6y 8Z dV2 dZ2 Sy 6Z dV2 dZ2T] udv = dy2dZ2T and
(534)
Cy
0 0 0 0 (0) 0 0 0 0 0 0 0 0 0 0 0 0
0 0 0 0
0 0 0 0 0 0 0 0 0 0 0 0
0 0 0
h 0 0 0
0 0 0 lt ( 0 ) o 0
0 1 0 0 0 1
0 0 0
2 r
h 0 0
0 0 1 0 0 1
(535)
Av and Bv are the rows and columns of the A and B corresponding to the components
O f Xy
The state L Q R weighting matrix Qv is given as
Qv
10 0 0 0 0 0 0 0 0 10 0 0 0 0 0 0 0 0 10 4 0 0 0 0 0 0 0 0 10 4 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1
(536)
mdash
Ry =w The state noise covariance matrix is taken
1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 9 x 10 4 0
0 0 0 0 0 0 0 9 x 10 4
(537)
119
while the measurement noise covariance matrix is
1 0 0 0 0 1 0 0 0 0 5 x 10 4 0
LO 0 0 5 x 10 4
(538)
As in the development of the longitudinal controller the transverse offset conshy
troller can be designed using the same two commands (lqrc and l try) wi th the input
matrices corresponding to the transverse subsystem with r = 50 The result is the
transverse compensator system given by
(539)
where xv is the state estimates of xv The singular values of the target transfer
function is compared with those achieved by the L Q G and L Q G L T R in Figure 5-
7(b) It is apparent that the recovery is not as good as that for the longitudinal case
again due to the non-minimal system Moreover the low value of r indicates that
only a lit t le recovery in the transverse subsystem is possible This suggests that the
robustness of the L Q G design is almost the maximum that can be achieved under
these conditions
Denoting f y = f j f j r xf = x^x^T and ud = udu ^ T the longishy
tudinal compensator can be combined to give
xf =
$d =
Aku
0
0 Akv
Cku o
Cu 0
Bku
0 B Xf
Xf = Ckxf
0
kv
Hf = Akxf - Bkyf
(540)
Vu Vv 0 Cv
Xf = CfXf
Defining a permutation matrix Pf such that Xj = PfX where X is the state vector
of the original nonlinear system given in Eq(32) the compensator and full nonlinear
120
system equations can be combined as
X
(541)
Ckxf
A block diagram representation of Eq(541) is shown in Figure 5-6
524 Simulation results
The dynamical response for the stationkeeping STSS with a tether length of 20
km and ini t ia l offset position of 1 m in each direction is presented in Figure 5-8 when
both the F L T attitude controller and the L Q G L T R offset controller are activated As
expected the offset controller is successful in quickly damping the elastic vibrations
in the longitudinal and transverse direction (Figure 5-8b) However from Figure 5-
8(a) it is clear that the presence of offset control requires a larger control moment
to regulate the attitude of the platform This is due to the additional torque created
by the larger moment arm around 2 m and 125 m in the inplane and out-of-plane
directions respectively introduced by the offset control In addition the control
moment for the platform is modulated by the tethers transverse vibration through
offset coupling However this does not significantly affect the l ibrational motion of
the tether whose thruster force remains relatively unchanged from the uncontrolled
vibration case
Final ly it can be concluded that the tether elastic vibration suppression
though offset control in conjunction with thruster and momentum-wheel attitude
control presents a viable strategy for regulating the dynamics tethered satellite sysshy
tems
121
STSS configuration 3-Body a1(0)=a2(0)=at(0)=2deg P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m6y(0)=52(0)=1m dx(0)=dy(0)=dz(0)=1m Stationkeeping I = 20km LQGLTR Vibration Control
Sat 1 Pitch and Roll Angles Tether Pitch and Roll Angles mdash 1 1 1 1 1 1 I 1 1 lt 1 ~ ^ 1 ^ l _
Time (Orbits) Time (Orbits) Sat 1 Roll Control Moment Tether Roll Control Thrust
(a) Time (Orbits) Time (Orbits)
Figure 5-8 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT attitude controller and LQGLTR offset vibration controller (a) attitude and libration controller reshysponse
122
STSS configuration 3-Body a1(0)=a2(0)=a(0)=2deg p1(0)=(32(0)=Q(0)=1deg ox(0)=14mSy(0)=5z(0)=1m dx(0)=dy(0)=dz(0)=1m Stationkeeping lt = 20km LQGLTR Vibration Control
Tether X Offset Position
Time (Orbits) Tether Y Tranverse Vibration 20r
Time (Orbits) Tether Y Offset Position
t o
1 2 Time (Orbits)
Tether Z Tranverse Vibration Time (Orbits)
Tether Z Offset Position
(b)
Figure 5-8
Time (Orbits) Time (Orbits)
Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT attitude controller and LQGLTR offset vibration controller (b) vibration and offset response
123
6 C O N C L U D I N G R E M A R K S
61 Summary of Results
The thesis has developed a rather general dynamics formulation for a multi-
body tethered system undergoing three-dimensional motion The system is composed
of multiple rigid bodies connected in a chain configuration by long flexible tethers
The tethers which are free to undergo libration as well as elastic vibrations in three
dimensions are also capable of deployment retrieval and constant length stationkeepshy
ing modes of operation Two types of actuators are located on the rigid satellites
thrusters and momentum-wheels
The governing equations of motion are developed using a new Order(N) apshy
proach that factorizes the mass matrix of the system such that it can be inverted
efficiently The derivation of the differential equations is generalized to account for
an arbitrary number of rigid bodies The equations were then coded in FORTRAN
for their numerical integration with the aid of a symbolic manipulation package that
algebraically evaluated the integrals involving the modes shapes functions used to dis-
cretize the flexible tether motion The simulation program was then used to assess the
uncontrolled dynamical behaviour of the system under the influence of several system
parameters including offset at the tether attachment point stationkeeping deployshy
ment and retrieval of the tether The study covered the three-body and five-body
geometries recently flown OEDIPUS system and the proposed BICEPS configurashy
tion It represents innovation at every phase a general three dimensional formulation
for multibody tethered systems an order-N algorithm for efficient computation and
application to systems of contemporary interest
Two types of controllers one for the attitude motion and the other for the
124
flexible vibratory motion are developed using the thrusters and momentum-wheels for
the former and the variable offset position for the latter These controllers are used to
regulate the motion of the system under various disturbances The attitude controller
is developed using the nonlinear Feedback Linearization Technique ( F L T ) and is based
on the nonlinear but rigid model of the system O n the other hand the separation
of the longitudinal and transverse frequencies from those of the attitude response
allows for the development of a linear optimal offset controller using the robust Linear
Quadratic Gaussian-Loop Transfer Recovery ( L Q G L T R ) method The effectiveness
of the two controllers is assessed through their subsequent application to the original
nonlinear flexible model
More important original contributions of the thesis which have not been reshy
ported in the literature include the following
(i) the model accounts for the motion of a multibody chain-type system undershy
going librational and in the case of tethers elastic vibrational motion in all
the three directions
(ii) the dynamics formulation is based on a recursive O(N) Lagrangian algorithm
that efficiently computes the systems generalized acceleration vector
(iii) the development of an attitude controller based on the Feedback Linearizashy
tion Technique for a multibody system using thrusters and momentum-wheels
located on each rigid body
(iv) the design of a three degree of freedom offset controller to regulate elastic v i shy
brations of the tether using the Linear Quadratic Gaussian and Loop Transfer
Recovery method ( L Q G L T R )
(v) substantiation of the formulation and control strategies through the applicashy
tion to a wide variety of systems thus demonstrating its versatility
125
The emphasis throughout has been on the development of a methodology to
study a large class of tethered systems efficiently It was not intended to compile
an extensive amount of data concerning the dynamical behaviour through a planned
variation of system parameters O f course a designer can easily employ the user-
friendly program to acquire such information Rather the objective was to establish
trends based on the parameters which are likely to have more significant effect on
the system dynamics both uncontrolled as well as controlled Based on the results
obtained the following general remarks can be made
(i) The presence of the platform products of inertia modulates the attitude reshy
sponse of the platform and gives rise to non-zero equilibrium pitch and roll
angles
(ii) Offset of the tether attachment point significantly affects the equil ibrium orishy
entation of the system as well as its dynamics through coupling W i t h a relshy
atively small subsatellite the tether dynamics dominate the system response
with high frequency elastic vibrations modulating the librational motion On
the other hand the effect of the platform dynamics on the tether response is
negligible As can be expected the platform dynamics is significantly affected
by the offset along the local horizontal and the orbit normal
(iii) For a three-body system deployment of the tether can destabilize the platform
in pitch as in the case of nonzero offset However the roll motion remains
undisturbed by the Coriolis force Moreover deployment can also render the
tether slack if it proceeds beyond a critical speed
(iv) Uncontrolled retrieval of the tether is always unstable in pitch It also leads
to high frequency elastic vibrations in the tether
(v) The five-body tethered system exhibits dynamical characteristics similar to
those observed for the three-body case As can be expected now there are
additional coupling effects due to two extra bodies a rigid subsatellite and a
126
flexible tether
(vi) Cartwheeling motion of the B I C E P S configuration can also be used to adshy
vantage in the deployment of the tether However exceeding a crit ical ini t ia l
cartwheeling rate can result in a large tension in the tether which eventually
causes the satellite to bounce back rendering the tether slack
(vii) Spinning the tether about its nominal length as in the case of O E D I P U S
introduces high frequency transverse vibrations in the tether which in turn
affect the dynamics of the satellites
(viii) The F L T based controller is quite successful in regulating the attitude moshy
tion of both the tether and rigid satellites in a short period of time during
stationkeeping deployment as well as retrieval phases
(ix) The offset controller is successful in suppressing the elastic vibrations of the
tether wi thin the allowable l imits of the offset mechanism ( plusmn 2 0 m)
62 Recommendations for Future Study
A s can be expected any scientific inquiry is merely a prelude to further efforts
and insight Several possible avenues exist for further investigation in this field A
few related to the study presented here are indicated below
(i) inclusion of environmental effects particularly the free molecular reaction
forces as well as interaction with Earths magnetic field and solar radiation
(ii) addition of flexible booms attached to the rigid satellites providing a mechashy
nism for energy dissipation
(iii) validation of the new three dimensional offset control strategies using ground
based experiments
(iv) animation of the simulation results to allow visual interpretation of the system
dynamics
127
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[34] Pradhan S M o d i V J and Misra A K On the Inverse Control of the Tethered Satellite Systems The Journal of the Astronautical Sciences A A S Publications Office San Diego California U S A V o l 43 No 2 Apr i l - June 1995 pp 179-193
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[37] Pradhan S M o d i V J and Misra A K Simultaneous Control of Platform Att i tude and Tether Vibra t ion using Offset Strategy AIAAA AS Astrodynamics Specialist Conference San Diego C A U S A July 1996 Paper No 96-3570 pp 1-11
[38] Ruying F and Bainum P M Dynamics and Control of a Space Platform with a Tethered Subsatellite Journal of Guidance Control and Dynamics V o l 11 No 4 July-August 1988 pp 377-381
[39] M o d i V J Lakshmanan P K and Misra A K Offset Control of Tethered Satellite Systems Analysis and Experimental Verification Acta Astronautica V o l 21 No 5 1990 pp 283-294
[40] X u D M Misra A K and M o d i V J Thruster Augmented Act ive Cin t ro l of a Tethered Satellite System During its Retrieval Journal of Guidance Control and Dynamics V o l 9 No 6 November-December 1986 pp 663-672
[41] M o d i V J Pradhan S and Misra A K Off-Set Control of the Tethered Systems Using a Graph Theoretic Approach Acta Astronautica V o l 35 No
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6 1995 pp 373-384
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[43] Pradhan S Planar Dynamics and Control of Tethered Satellite Systems P h D Thesis Department of Mechanical Engineering The University of Br i t i sh Columbia December 1994
[44] Bainum P M and Kumar V K Optimal-Control of the Shuttle-Tethered System Acta Astronautica V o l 7 No 12 1980 pp 1333-1348
[46] Hughes P C Spacecraft Att i tude Dynamics John Wi ley amp Sons New York U S A 1992 Chapter 2 pp 7-38
[47] Stengel R F Optimal Control and Estimation Dover Publications Inc New York U S A 1994 p 26
[48] Nayfeh A H and Mook D T Nonlinear Oscillation John Wi ley and Sons New York U S A 1979 pp 485-500
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[50] Meirovitch L Elements of Vibration Analysis M c G r a w - H i l l Inc New York U S A 1986 pp 255-256
[51] Greenwood D T Principles of Dynamics Prentice Ha l l Inc Englewood Cliffs New Jersey U S A 1965 pp 252-275
[52] Boyce W E and D i P r i m a R C Elemental Differential Equations and Boundshy
ary Value Problems John Wi ley amp Sons Inc New York U S A 1986 pp 424-431
[53] IMSL Library Reference Manual Vol 1 I M S L Inc Houston Texas U S A 1980 pp D G E A R 1 - D G E A R 9
[54] Gear C W Numerical Initial Value Problems in Ordinary Differential Equashy
tions Prentice-Hall Englewood Cliffs New Jersey U S A 1971 pp 158-166
[55] Char B W Geddes K O Gonnet G H Leong B L Monagan M B and Wat t S M First Leaves A Tutorial Introduction to Maple V Springer-Verlag New York U S A 1992
[56] Tsakalis K S and Ioannou P A Linear Time-Varying Systems Control and
Adaptation Prentice-Hall Inc Englewood Cliffs New Jersey U S A 1993 pp 148-180
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131
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[60] Athans M The Role and Use of the Stochastic Linear-Quadrat ic-Guassian Problem in Control System Design IEEE Transactions on Automatic Control V o l A C - 1 6 No 6 December 1971 pp 529-552
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132
A P P E N D I X I TENSOR R E P R E S E N T A T I O N OF T H E
EQUATIONS OF M O T I O N
11 Prel iminary Remarks
The derivation of the equations of motion for the multibody tethered system
involves the product of several matrices as well as the derivative of matrices and
vectors with respect to other vectors In order to concisely code these relationships
in FORTRAN the matrix equations of motion are expressed in tensor notation
Tensor mathematics in general is an extremely powerful tool which can be
used to solve many complex problems in physics and engineering[6465] However
only a few preliminary results of Cartesian tensor analysis are required here They
are summarized below
12 Mathematical Background
The representation of matrices and vectors in tensor notation simply involves
the use of indices referring to the specific elements of the matrix entity For example
v = vk (k = lN) (Il)
A = Aij (i = lNj = lM)
where vk is the kth element of vector v and Aj is the element on the t h row and
3th column of matrix A It is clear from Eq(Il) that exactly one index is required to
completely define an entire vector whereas two independent indices are required for
matrices For this reason vectors and matrices are known as first-order and second-
order tensors respectively Scaler variables are known as zeroth-order tensors since
in this case an index is not required
133
Matrix operations are expressed in tensor notation similar to the way they are
programmed using a computer language such as FORTRAN or C They are expressed
in a summation notation known as Einstein notation For example the product of a
matrix with a vector is given as
w = Av (12)
or in tensor notation W i = ^2^2Aij(vkSjk)
4-1 ltL3gt j
where 8jk is the Kronecker delta Dropping the summation symbol the matrix-vector
product can be expressed compactly as
Wi = AijVj = Aimvm (14)
Here j is a dummy index since it appears twice in a single term and hence can
be replaced by any other index Note that since the resulting product is a vector
only one index is required namely i that appears exactly once on both sides of the
equation Similarly
E = aAB + bCD (15)
or Ej mdash aAimBmj + bCimDmj
~ aBmjAim I bDmjCim (16)
where A B C D and E are second-order tensors (matrices) and a and b are zeroth-
order tensors (scalers) In addition once an expression is in tensor form it can be
treated similar to a scaler and the terms can be rearranged
The transpose of a matrix is also easily described using tensors One simply
switches the position of the indices For example let w = ATv then in tensor-
notation
Wi = AjiVj (17)
134
The real power of tensor notation is in its unambiguous expression of terms
containing partial derivatives of scalers vectors or matrices with respect to other
vectors For example the Jacobian matrix of y = f(x) can be easily expressed in
tensor notation as
di-dx-j-1 ( L 8 )
mdash which is a second-order tensor as expected If f(x) = Ax where A is constant then
f - a )
according to the current definition If C = A(x)B(x) then the time derivative of C
is given by
C = Ax)Bx) + A(x)Bx) (110)
or in tensor notation
Cij = AimBmj + AimBmj (111)
= XhiAimfcBmj + AimBmj^k)
which is more clear than its equivalent vector form Note that A^^ and Bmjk are
third-order tensors that can be readily handled in F O R T R A N
13 Forcing Function
The forcing function of the equations of motion as given by Eq(260) is
- u 1-SdM- dqtdPe 9 f i m F(q qt) = Mq - -q mdashq + mdashmdash - Qd (260)
This expression can be converted into tensor form to facilitate its implementation
into F O R T R A N source code
Consider the first term Mq From Eq(289) the mass matrix M is given by
T M mdash Rv MtRv which in tensor form is
Mtj = RZiM^R^j (112)
135
Taking the time derivative of M gives
Mij = qsRv
ni^MtnrnRv
mj + Rv
niMtnTn^Rv
mj + Rv
niMtnmRv
mjs) (113)
The second term on the right hand side in Eq(260) is given by
-JPdMu 1 bdquo T x
2Q ~cWq = 2Qs s r k Q r ^ ^
Expanding Msrk leads to
M src = RnskMtnmRmr + RnsMtnmkRmr + RnsMtnmRmrk- (L15)
Finally the potential energy term in Eq(260) is also expanded into tensor
form to give dPe = dqt dPe
dq dq dqt = dq^QP^
and since ^ = Rp(q)q then
Now inserting Eqs(I13-I17) into Eq(260) and rearranging the tensor form of the
forcing term can how be stated as
Fk(q^q^)=(RP
sk + R P n J q n ) ^ - - Q d k
+ QsQr | (^Rlks ~ 2 M t n m R m r
+ (^RnkMtnms ~ 2~RnsMtnmgtk^ R m r
(^RnkRmrs ~ ~2RnsRmrk^j
where Fk(q q t) represents the kth component of F
(118)
136
A P P E N D I X II R E D U C E D EQUATIONS OF M O T I O N
II 1 Prel iminary Remarks
The reduced model used for the design of the vibration controller is represented
by two rigid end-bodies capable of three-dimensional attitude motion interconnected
with a fixed length tether The flexible tether is discretized using the assumed-
mode method with only the fundamental mode of vibration considered to represent
longitudinal as well as in-plane and out-of-plane transverse deformations In this
model tether structural damping is neglected Finally the system is restricted to a
nominal circular orbit
II 2 Derivation of the Lagrangian Equations of Mot ion
Derivation of the equations of motion for the reduced system follows a similar
procedure to that for the full case presented in Chapter 2 However in this model the
straightforward derivation of the energy expressions for the entire system is undershy
taken ignoring the Order(N) approach discussed previously Furthermore in order
to make the final governing equations stationary the attitude motion is now referred
to the local LVLH frame This is accomplished by simply defining the following new
attitude angles
ci = on + 6
Pi = Pi (HI)
7t = 0
where and are the pitch and roll angles respectively Note that spin is not
considered in the reduced model hence 7J = 0 Substituting Eq(IIl) into Eq(214)
137
gives the new expression for the rotation matrix as
T- CpiSa^ Cai+61 S 0 S a + e i
-S 0 c Pi
(II2)
W i t h the new rotation matrix defined the inertial velocity of the elemental
mass drrii o n the ^ link can now be expressed using Eq(215) as
(215)
Since the tether length is assumed to be fixed T j f j and Tj$jlt^ of Eq(215) are
identically zero Also defining
(II3) Vi = I Pi i
Rdm- f deg r the reduced system is given by
where
and
Rdm =Di + Pi(gi)m + T^iSi
Pi(9i)m= [Tai9i T0gi T^] fj
(II4)
(II5)
(II6)
D1 = D D l D S v
52 = 31 + P(d2)m +T[d2
D3 = D2 + l2 + dX2Pii)m + T2i6x
Note that $ j for i = 2 is defined in Section 212 whereas for i = 1 and 3 it is
the null matrix Since only one mode is used for each flexible degree of freedom
5i = [6Xi6yi6Zi]T for i = 2
The kinetic energy of the system can now be written as
i=l Z i = l Jmi RdmRdmdmi (UJ)
138
Expanding Kampi using Eq(II4)
KH = miDD + 2DiP
i j gidmi)fji + 2D T ltMtrade^
+tf J P^(9i)Pi9i)d^l + 2mT j PFmTi^dm^ (II8)
where the integrals are evaluated using the procedure similar to that described in
Chapter 2 The gravitational and strain energy expressions are given by Eq(247)
and Eq(251) respectively using the newly defined rotation matrix T in place of
TV
Substituting the kinetic and potential energy expressions in
d ( 8Ke d(Ke - Pe) i bull i Q ^ 0 (II9)
d t dqred d(ired
with
Qred = [ algt l gt a 2 2 A 2 A 2 lt ^ 2 Q 3 3 ] (5-14)
and d2 regarded as a time-varying parameter the nonlinear non-autonomous equashy
tions of motion for the reduced system can finally be expressed as
MredQred + fred = reg- (5-15)
139
In presenting this thesis in partial fulfilment of the requirements for an advanced
degree at the University of British Columbia I agree that the Library shall make it
freely available for reference and study I further agree that permission for extensive
copying of this thesis for scholarly purposes may be granted by the head of my
department or by his or her representatives It is understood that copying or
publication of this thesis for financial gain shall not be allowed without my written
permission
Department of
The University of British Columbia Vancouver Canada
DE-6 (288)
ABSTRACT
The equations of motion for a multibody tethered satellite system in three dishy
mensional Keplerian orbit are derived The model considers a multi-satellite system
connected in series by flexible tethers Both tethers and subsatellites are free to unshy
dergo three dimensional attitude motion together with deployment and retrieval as
well as longitudinal and transverse vibration for the tether The elastic deformations
of the tethers are discretized using the assumed mode method The tether attachment
points to the subsatellites are kept arbitrary and time varying The model is also cashy
pable of simulating the response of the entire system spinning about an arbitrary
axis as in the case of OEDIPUS-AC which spins about the nominal tether length
or the proposed BICEPS mission where the system cartwheels about the orbit norshy
mal The governing equations of motion are derived using a non-recursive order(N)
Lagrangian procedure which significantly reduces the computational cost associated
with the inversion of the mass matrix an important consideration for multi-satellite
systems Also a symbolic integration and coding package is used to evaluate modal
integrals thus avoiding their costly on-line numerical evaluation
Next versatility of the formulation is illustrated through its application to
two different tethered satellite systems of contemporary interest Finally a thruster
and momentum-wheel based attitude controller is developed using the Feedback Linshy
earization Technique in conjunction with an offset (tether attachment point) control
strategy for the suppression of the tethers vibratory motion using the optimal Linshy
ear Quadratic Gaussian-Loop Transfer Recovery method Both the controllers are
successful in stabilizing the system over a range of mission profiles
ii
T A B L E OF C O N T E N T S
A B S T R A C T bull i i
T A B L E O F C O N T E N T S i i i
L I S T O F S Y M B O L S v i i
L I S T O F F I G U R E S x i
A C K N O W L E D G E M E N T xv
1 I N T R O D U C T I O N 1
11 Prel iminary Remarks 1
12 Brief Review of the Relevant Literature 7
121 Mul t ibody O(N) formulation 7
122 Issues of tether modelling 10
123 Att i tude and vibration control 11
13 Scope of the Investigation 12
2 F O R M U L A T I O N O F T H E P R O B L E M 14
21 Kinematics 14
211 Prel iminary definitions and the itfl position vector 14
212 Tether flexibility discretization 14
213 Rotat ion angles and transformations 17
214 Inertial velocity of the ith link 18
215 Cyl indr ica l orbital coordinates 19
216 Tether deployment and retrieval profile 21
i i i
22 Kinetics and System Energy 22
221 Kinet ic energy 22
222 Simplification for rigid links 23
223 Gravitat ional potential energy 25
224 Strain energy 26
225 Tether energy dissipation 27
23 O(N) Form of the Equations of Mot ion 28
231 Lagrange equations of motion 28
232 Generalized coordinates and position transformation 29
233 Velocity transformations 31
234 Cyl indrical coordinate modification 33
235 Mass matrix inversion 34
236 Specification of the offset position 35
24 Generalized Control Forces 36
241 Preliminary remarks 36
242 Generalized thruster forces 37
243 Generalized momentum gyro torques 39
3 C O M P U T E R I M P L E M E N T A T I O N 42
31 Preliminary Remarks 42
32 Numerical Implementation 43
321 Integration routine 43
322 Program structure 43
33 Verification of the Code 46
iv
331 Energy conservation 4 6
332 Comparison with available data 50
4 D Y N A M I C S I M U L A T I O N 53
41 Prel iminary Remarks 53
42 Parameter and Response Variable Definitions 53
43 Stationkeeping Profile 56
44 Tether Deployment 66
45 Tether Retrieval 71
46 Five-Body Tethered System 71
47 B I C E P S Configuration 80
48 O E D I P U S Spinning Configuration 88
5 A T T I T U D E A N D V I B R A T I O N C O N T R O L 93
51 Att i tude Control 93
511 Prel iminary remarks 93
512 Controller design using Feedback Linearization Technique 94
513 Simulation results 96
52 Control of Tethers Elastic Vibrations 110
521 Prel iminary remarks 110
522 System linearization and state-space realization 1 1 0
523 Linear Quadratic Gaussian control
with Loop Transfer Recovery 114
524 Simulation results 121
v
6 C O N C L U D I N G R E M A R K S 1 2 4
61 Summary of Results 124
62 Recommendations for Future Study 127
B I B L I O G R A P H Y 128
A P P E N D I C E S
I T E N S O R R E P R E S E N T A T I O N O F T H E E Q U A T I O N S O F M O T I O N 133
11 Prel iminary Remarks 133
12 Mathematical Background 133
13 Forcing Function 135
II R E D U C E D E Q U A T I O N S O F M O T I O N 137
II 1 Preliminary Remarks 137
II2 Derivation of the Lagrangian Equations of Mot ion 137
v i
LIST OF SYMBOLS
A j intermediate length over which the deploymentretrieval profile
for the itfl l ink is sinusoidal
A Lagrange multipliers
EUQU longitudinal state and measurement noise covariance matrices
respectively
matrix containing mode shape functions of the t h flexible link
pitch roll and yaw angles of the i 1 link
time-varying modal coordinate for the i 1 flexible link
link strain and stress respectively
damping factor of the t h attitude actuator
set of attitude angles (fjj = ci
structural damping coefficient for the t h l ink EjJEj
true anomaly
Earths gravitational constant GMejBe
density of the ith link
fundamental frequency of the link longitudinal tether vibrashy
tion
A tethers cross-sectional area
AfBfCjDf state-space representation of flexible subsystem
Dj inertial position vector of frame Fj
Dj magnitude of Dj mdash mdash
Df) transformation matrix relating Dj and Ds
D r i inplane radial distance of the first l ink
DSi Drv6i DZ1 T for i = 1 and DX Dy poundgt 2 T for 1
m
Pi
w 0 i
v i i
Dx- transformation matrix relating D j and Ds-
DXi Dyi Dz- Cartesian components of D
DZl out-of-plane position component of first link
E Youngs elastic modulus for the ith l ink
Ej^ contribution from structural damping to the augmented complex
modulus E
F systems conservative force vector
FQ inertial reference frame
Fi t h link body-fixed reference frame
n x n identity matrix
alternate form of inertia matrix for the i 1 l ink
K A Z $i(k + dXi+1)
Kei t h link kinetic energy
M(q t) systems coupled mass matrix
Mma Mma Mmy- control moments in the pitch roll and yaw directions respec-
tively for the t h link
Mr fr rigid mass matrix and force vector respectively
Mred fred- Qred mass matrix force vector and generalized coordinate vector for
the reduced model respectively
Mt block diagonal decoupled mass matrix
symmetric decoupled mass matrix
N total number of links
0N) order-N
Pi(9i) column matrix [T^giTpg^T^gi) mdash
Q nonconservative generalized force vector
Qu actuator coupling matrix or longitudinal L Q R state weighting
matrix
v i i i
Rdm- inertial position of the t h link mass element drrii
RP transformation matrix relating qt and tf
Ru longitudinal L Q R input weighting matrix
RvRn Rd transformation matrices relating qt and q
S(f Mdc) generalized acceleration vector of coupled system q for the full
nonlinear flexible system
th uiirv l u w u u u maniA
th
T j i l ink rotation matrix
TtaTto control thrust in the pitch and roll direction for the im l ink
respectively
rQi maximum deploymentretrieval velocity of the t h l ink
Vei l ink strain energy
Vgi link gravitational potential energy
Rayleigh dissipation function arising from structural damping
in the ith link
di position vector to the frame F from the frame F_i bull
dc desired offset acceleration vector (i 1 l ink offset position)
drrii infinitesimal mass element of the t h l ink
dx^dy^dZi Cartesian components of d along the local vertical local horishy
zontal and orbit normal directions respectively
F-Q mdash
gi r igid and flexible position vectors of drrii fj + ltEraquolt5
ijk unit vectors 1 0 0 r 010-^ and 00 l r respectively
li length of the t h l ink
raj mass of the t h link
nfj total number of flexible modes for the i 1 l ink nXi + nyi + nZi
nq total number of generalized coordinates per link nfj + 7
nqq systems total number of generalized coordinates Nnq
ix
number of flexible modes in the longitudinal inplane and out-
of-plane transverse directions respectively for the i ^ link
number of attitude control actuators
- $ T
set of generalized coordinates for the itfl l ink which accounts for
interactions with adjacent links
qv---QtNT
set of coordinates for the independent t h link (not connected to
adjacent links)
rigid position of dm in the frame Fj
position of centre of mass of the itfl l ink relative to the frame Fj
i + $DJi
desired settling time of the j 1 attitude actuator
actuator force vector for entire system
flexible deformation of the link along the Xj yi and Zj direcshy
tions respectively
control input for flexible subsystem
Cartesian components of fj
actual and estimated state of flexible subsystem respectively
output vector of flexible subsystem
LIST OF FIGURES
1-1 A schematic diagram of the space platform based N-body tethered
satellite system 2
1-2 Associated forces for the dumbbell satellite configuration 3
1- 3 Some applications of the tethered satellite systems (a) multiple
communication satellites at a fixed geostationary location
(b) retrieval maneuver (c) proposed B I C E P S configuration 6
2- 1 Vector components of the i 1 and (i- l ) 1 chain links 15
2-2 Vector components in cylindrical orbital coordinates 20
2-3 Inertial position of subsatellite thruster forces 38
2- 4 Coupled force representation of a momentum-wheel on a rigid body 40
3- 1 Flowchart showing the computer program structure 45
3-2 Kinet ic and potentialenergy transfer for the three-body platform based
tethered satellite system (a) variation of kinetic and potential energy
(b) percent change in total energy of system 48
3-3 Kinet ic and potential energy transfer for the five-body platform based
tethered satellite system (a) variation of kinetic and potential energy
(b) percent change in total energy of system 49
3-4 Simulation results for the platform based three-body tethered
system originally presented in Ref[43] 51
3- 5 Simulation results for the platform based three-body tethered
system obtained using the present computer program 52
4- 1 Schematic diagram showing the generalized coordinates used to
describe the system dynamics 55
4-2 Stationkeeping dynamics of the three-body STSS configuration
without offset (a) attitude response (b) vibration response 57
4-3 Stationkeeping dynamics of the three-body STSS configuration
xi
with offset along the local vertical
(a) attitude response (b) vibration response 60
4-4 Stationkeeping dynamics of the three-body STSS configuration
with offset along the local horizontal
(a) attitude response (b) vibration response 62
4-5 Stationkeeping dynamics of the three-body STSS configuration
with offset along the orbit normal
(a) attitude response (b) vibration response 64
4-6 Stationkeeping dynamics of the three-body STSS configuration
wi th offset along the local horizontal local vertical and orbit normal
(a) attitude response (b) vibration response 67
4-7 Deployment dynamics of the three-body STSS configuration
without offset
(a) attitude response (b) vibration response 69
4-8 Deployment dynamics of the three-body STSS configuration
with offset along the local vertical
(a) attitude response (b) vibration response 72
4-9 Retrieval dynamics of the three-body STSS configuration
with offset along the local vertical and orbit normal
(a) attitude response (b) vibration response 74
4-10 Schematic diagram of the five-body system used in the numerical example 77
4-11 Stationkeeping dynamics of the five-body STSS configuration
without offset
(a) attitude response (b) vibration response 78
4-12 Deployment dynamics of the five-body STSS configuration
wi th offset along the local vertical
(a) attitude response (b) vibration response 81
4-13 Stationkeeping dynamics of the three-body B I C E P S configuration
x i i
with offset along the local vertical
(a) attitude response (b) vibration response 84
4-14 Cartwheeling dynamics with deployment for the three-body B I C E P S
with offset along the local vertical
(a) attitude response (b) vibration response 86
4-15 Spin dynamics (7 = l deg s ) of the three-body O E D I P U S configuration
with offset along the local vertical
(a) attitude response (b) vibration response 89
4- 16 Spin dynamics (7 = 10deg s ) of the three-body O E D I P U S configuration
wi th offset along the local vertical
(a) attitude response (b) vibration response 91
5- 1 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear rigid F L T controller with offset along the
local horizontal local vertical and orbit normal
(a) attitude and vibration response (b) control actuator time histories 98
5-2 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear flexible F L T controller with offset along the
local horizontal local vertical and orbit normal
(a) attitude and vibration response (b) control actuator time histories 101
5-3 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear rigid F L T controller with offset along
the local vertical
(a) attitude and vibration response (b) control actuator time histories 103
5-4 Deployment dynamics of the three-body STSS using the nonlinear
rigid F L T controller with offset along the local vertical
(a) attitude and vibration response (b) control actuator time histories 105
5-5 Retrieval dynamics of the three-body STSS using the non-linear
rigid F L T controller with offset along the local vertical and orbit normal
(a) attitude and vibration response (b) control actuator time histories 108
x i i i
5-6 Block diagram for the L Q G L T R estimator based compensator 115
5-7 Singular values for the L Q G and L Q G L T R compensator compared to target
return ratio (a) longitudinal design (b) transverse design 118
5-8 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear rigid F L T attitude controller and L Q G L T R
offset vibration controller (a) attitude and libration controller response
(b) vibration and offset response 122
xiv
A C K N O W L E D G E M E N T
Firs t and foremost I would like to express the most genuine gratitude to
my supervisor Prof V i n o d J M o d i whose unending guidance and encouragement
proved most invaluable during the course of my studies
I am also deeply indebted to Dr Satyabrata Pradhan for his limitless patience
and technical advice from which this work would not have been possible I would also
like to extend a word of appreciation towards Prof A r u n K Mis ra ( M c G i l l University)
for ini t iat ing my interest in space dynamics and control
Also I would like to thank my collegues Dr Sandeep Munshi M r Mathieu
Caron and M r Yuan Chen as well as Dr Mae Seto M r Gary L i m and M r Mark
Chu for their words of advice and helpful suggestions that heavily contributed to all
aspects of my studies
Final ly I would like to thank all my family and friends here in Vancouver and
back home in Montreal whose unending support made the past two years the most
memorable of my life
The research project was supported by the Natural Sciences and Engineering
Research Council s P G S - A Scholarship held by the author as well as by N S E R C grant
A-2181 held by Prof V J M o d i
xv
1 INTRODUCTION
11 Prel iminary Remarks
W i t h the ever changing demands of the worlds population one often wonders
about the commitment to the space program On the other hand the importance
of worldwide communications global environmental monitoring as well as long-term
habitation in space have demanded more ambitious and versatile satellites systems
It is doubtful that the Russian scientist Tsiolkovsky[l] when he first proposed the
uti l izat ion of the Earths gravity-gradient environment in the last century would have
envisioned the current and proposed applications of tethered satellite systems
A tethered satellite system consists of two or more subsatellites connected
to each other by long thin cables or tethers which are in tension (Figure 1-1) The
subsatellite which can also include the shuttle and space station are in orbit together
The first proposed use of tethers in space was associated with the rescue of stranded
astronauts by throwing a buoy on a tether and reeling it to the rescue vehicle
Prel iminary studies of such systems led to the discovery of the inherent instability
during the tether retrieval[2] Nevertheless the bir th of tethered systems occured
during the Gemini X I and X I I missions[3] in 1966 when a short tether (10-30m)
was used to generate the first artificial gravity environment in space (000015g) by
cartwheeling (spinning) the system about its center of mass
The mission also demonstrated an important use of tethers for gravity gradient
stabilization The force analysis of a simplified model illustrates this point Consider
the dumbbell satellite configuration orbiting about the centre of Ear th as shown in
Figure 1-2 Satellite 1 and satellite 2 are connected to each other by a tether with
1
Space Station (Satellite 1 )
2
Satellite N
Figure 1-1 A schematic diagram of the space platform based N-body tethered satellite system
2
Figure 1-2 Associated forces for the dumbbell satellite configuration
3
the whole system free to librate about the systems centre of mass C M The two
major forces acting on each body are due to the centrifugal and gravitational effects
Since body 1 is further away from the earth its gravitational force Fg^y is smaller
than that of body 2 ie Fgi lt Fg^ However its centrifugal force FCl is greater
then the centrifugal force experienced by the second satellite (FCl gt FC2) Moreover
for body 1 Fg^ lt Fc^ in contrast to Fg2 gt FC2 for body 2 Thus resolving the force
components along the tether there is an evident resultant tension force Ff present
in the tether Similarly adding the normal components of the two forces vectorially
results in a force FR which restores the system to its stable equil ibrium along the
local vertical These tension and gravity-gradient restoring forces are the two most
important features of tethered systems
Several milestone missions have flown in the recent past The U S A J a p a n
project T P E (Tethered Payload Experiment) [4] was one of the first to be launched
using a sounding rocket to conduct environmental studies The results of the T P E
provided support for the N A S A I t a l y shuttle based tethered satellite system referred
to as the TSS-1 mission[5] This experiment aimed to study the electrodynamic effects
of a shuttle borne satellite system with a conductive tether connection where an
electric current was induced in the tether as it swept through the Earth s magnetic
field Unfortunetely the two attempts in August 1992 and February 1996 resulted
in only partial success The former suffered from a spool failure resulting in a tether
deployment of only 256m of a planned 20km During the latter attempt a break
in the tethers insulation resulted in arcing leading to tether rupture Nevertheless
the information gathered by these attempts st i l l provided the engineers and scientists
invaluable information on the dynamic behaviour of this complex system
Canada also paved the way with novel designs of tethered system aimed
primari ly at the study of the ionosphere and Northern Lights (Aurora Borealis)
4
The O E D I P U S A and C (Observation of Electrified Distributions in the Ionospheric
P lasma - a Unique Strategy) missions[6] launched from a sounding rocket in January
1989 and November 1995 respectively provided insight into the complex dynamical
behaviour of two comparable mass satellites connected by a 1 km long spinning
tether
Final ly two experiments were conducted by N A S A called S E D S I and II
(Small Expendable Deployable System) [7] which hold the current record of 20 km
long tether In each of these missions a small probe (26 kg) was deployed from the
second stage of a D E L T A II rocket thus succesfully demonstrating the feasibility of
long tether deployment Retrieval of the tether which is significantly more difficult
has not yet been achieved
Several proposed applications are also currently under study These include the
study of earths upper atmosphere using probes lowered from the shuttle in a low earth
orbit ( L E O ) to the deployment and retrieval of satellites for servicing Even more
promising application concerns the generation of power for the proposed International
Space Station using conductive tethers The Canadian Space Agency (CSA) has also
proposed a dual-mass satellite system B I C E P S (Bl-static Canadian Experiment on
Plasmas in Space) to make simultaneous measurements at different locations in the
environment for correlation studies[8] A unique feature of the B I C E P S mission is
the deployment of the tether aided by the angular momentum of the system which
is ini t ia l ly cartwheeling about its orbit normal (Figure 1-3)
In addition to the three-body configuration multi-tethered systems have been
proposed for monitoring Earths environment in the global project monitored by
N A S A entitled Mission to Planet Ear th ( M T P E ) The proposed mission aims to
study pollution control through the understanding of the dynamic interactions be-
5
Retrieval of Satellite for Servicing
lt mdash
lt Satellite
Figure 1-3 Some applications of the tethered satellite systems (a) multiple comshymunication satellites at a fixed geostationary location (b) retrieval maneuver (c) proposed BICEPS configuration
6
tween the Earths atmosphere biosphere hydrosphere and cryosphere This wi l l be
accomplished with multiple tethered instrumentation payloads simultaneously soundshy
ing at different altitudes for spatial correlations Such mult ibody systems are considshy
ered to be the next stage in the evolution of tethered satellites Micro-gravity payload
modules suspended from the proposed space station for long-term experiments as well
as communications satellites with increased line-of-sight capability represent just two
of the numerous possible applications under consideration
12 Brief Review of the Relevant Literature
The design of multi-payload systems wi l l require extensive dynamic analysis
and parametric study before the final mission configuration can be chosen This
wi l l include a review of the fundamental dynamics of the simple two-body tethered
systems and how its dynamics can be extended to those of multi-body tethered system
in a chain or more generally a tree topology
121 Multibody 0(N) formulation
Many studies of two-body systems have been reported One of the earliest
contribution is by Rupp[9] who was the first to study planar dynamics of the simshy
plified Shuttle based Tethered Satellite System (STSS) His pioneering work led to
the discovery of pitch oscillation growth during the retrieval phase Later Baker[10]
advanced the investigation to the third dimension in addition to adding atmospheric
effects to the system A more complete dynamical model was later developed and
analyzed by M o d i and Misra[ l l 12] which included deployment and tether flexibility
A comprehensive survey of important developments and milestones in the area have
been compiled and reviewed by Mis ra and Modi[13] The major conclusions based on
the literature may be summarized as follows
bull the stationkeeping phase is stable
7
bull deployment can be unstable if the rate exceeds a crit ical speed
bull retrieval of the tether is inherently unstable
bull transverse vibrations can grow due to the Coriolis force induced during deshy
ployment and retrieval
Misra Amier and Modi[14] were one of the first to extend these results to the
three-satellite double pendulum case This simple model which includes a variable
length tether was sufficient to uncover the multiple equilibrium configurations of
the system Later the work was extended to include out-of-plane motion and a
preliminary reel-rate control law to regulate the librational motion[15] Kesmir i and
Misra[16] developed a general formulation for N-body tethered systems based on the
Lagrangian principle where three-dimensional motion and flexibility are accounted for
However from their work it is clear that as the number of payload or bodies increases
the computational cost to solve the forward dynamics also increases dramatically
Tradit ional methods of inverting the mass matrix M in the equation
ML+F = Q
to solve for the acceleration vector proved to be computationally costly being of the
order for practical simulation algorithms O(N^) refers to a mult ibody formulashy
tion algorithm whose computational cost is proportional to the cube of the number
of bodies N used It is therefore clear that more efficient solution strategies have to
be considered
Many improved algorithms have been proposed to reduce the number of comshy
putational steps required for solving for the acceleration vector Rosenthal[17] proshy
posed a recursive formulation based on the triangularization of the equations of moshy
tion to obtain an 0(N2) formulation He also demonstrates the use of a symbolic
manipulation program to reduce the final number of computations A recursive L a -
8
grangian formulation proposed by Book[18] also points out the relative inefficiency of
conventional formulations and proceeds to derive an algorithm which is also 0(N2)
A recursive formulation is one where the equations of motion are calculated in order
from one end of the system to the other It usually involves one or more passes along
the links to calculate the acceleration of each link (forward pass) and the constraint
forces (backward or inverse pass) Non-recursive algorithms express the equations of
motion for each link independent of the other Although the coupling terms st i l l need
to be included the algorithm is in general more amenable to parallel computing
A st i l l more efficient dynamics formulation is an O(N) algorithm Here the
computational effort is directly proportional to the number of bodies or links used
Several types of such algorithms have been developed over the years The one based
on the Newton-Euler approach has recursive equations and is used extensively in the
field of multi-joint robotics[1920] It should be emphasized that efficient computation
of the forward and inverse dynamics is imperative if any on-line (real-time) control
of the joint is to be achieved Hollerbach[21] proposed a recursive O(N) formulation
based on Lagranges equations of motion However his derivation was primari ly foshy
cused on the inverse dynamics and did not improve the computational efficiency of
the forward dynamics (acceleration vector) Other recursive algorithms include methshy
ods based on the principle of vir tual work[22] and Kanes equations of motion[23]
Keat[24] proposed a method based on a velocity transformation that eliminated the
appearance of constraint forces and was recursive in nature O n the other hand
K u r d i l a Menon and Sunkel[25] proposed a non-recursive algorithm based on the
Range-Space method which employs element-by-element approach used in modern
finite-element procedures Authors also demonstrated the potential for parallel comshy
putation of their non-recursive formulation The introduction of a Spatial Operator
Factorization[262728] which utilizes an analogy between mult ibody robot dynamics
and linear filtering and smoothing theory to efficiently invert the systems mass ma-
9
t r ix is another approach to a recursive algorithm Their results were further extended
to include the case where joints follow user-specified profiles[29] More recently Prad-
han et al [30] proposed a non-recursive Lagrangian algorithm for factorizing the mass
matrix leading to an O(N) formulation of the forward dynamics of the system A n
efficient O(N) algorithm where the number of bodies N in the system varies on-line
has been developed by Banerjee[31]
122 Issues of tether modelling
The importance of including tether flexibility has been demonstrated by several
researchers in the past However there are several methods available for modelling
tether flexibility Each of these methods has relative advantages and limitations
wi th regards to the computational efficiency A continuum model where flexible
motion is discretized using the assumed mode method has been proposed[3233] and
succesfully implemented In the majority of cases even one flexible mode is sufficient
to capture the dynamical behaviour of the tether[34] K i m and Vadali[35] proposed a
so-called bead model or lumped-mass approach that discretizes the tether using point
masses along its length However in order to accurately portray the motion of the
tether a significantly high number of beads are needed thus increasing the number of
computation steps Consequently this has led to the development of a hybrid between
the last two approaches[16] ie interpolation between the beads using a continuum
approach thus requiring less number of beads Finally the tether flexibility can also
be modelled using a finite-element approach[36] which is more suitable for transient
response analysis In the present study the continuum approach is adopted due to
its simplicity and proven reliability in accurately conveying the vibratory response
In the past the modal integrals arising from the discretization process have
been evaluted numerically In general this leads to unnecessary effort by the comshy
puter Today these integrals can be evaluated analytically using a symbolic in-
10
tegration package eg Maple V Subsequently they can be coded in F O R T R A N
directly by the package which could also result in a significant reduction in debugging
time More importantly there is considerable improvement in the computational effishy
ciency [16] especially during deployment and retrieval where the modal integral must
be evaluated at each time-step
123 Attitude and vibration control
In view of the conclusions arrived at by some of the researchers mentioned
above it is clear that an appropriate control strategy is needed to regulate the dyshy
namics of the system ie attitude and tether vibration First the attitude motion of
the entire system must be controlled This of course would be essential in the case of
a satellite system intended for scientific experiments such as the micro-gravity facilishy
ties aboard the proposed Internation Space Station Vibra t ion of the flexible members
wi l l have to be checked if they affect the integrity of the on-board instrumentation
Thruster as well as momentum-wheel approach[34] have been considered to
regulate the rigid-body motion of the end-satellite as well as the swinging motion
of the tether The procedure is particularly attractive due to its effectiveness over a
wide range of tether lengths as well as ease of implementation Other methods inshy
clude tension and length rate control which regulate the tethers tension and nominal
unstretched length respectively[915] It is usually implemented at the deployment
spool of the tether More recently an offset strategy involving time dependent moshy
tion of the tether attachment point to the platform has been proposed[3738] It
overcomes the problem of plume impingement created by the thruster control and
the ineffectiveness of tension control at shorter lengths However the effectiveness of
such a controller can become limited with an exceedingly long tether due to a pracshy
tical l imi t on the permissible offset motion In response to these two control issues a
hybrid thrusteroffset scheme has been proposed to combine the best features of the
11
two methods[39]
In addition to attitude control these schemes can be used to attenuate the
flexible response of the tether Tension and length rate control[12] as well as thruster
based algorithms[3340] have been proposed to this end M o d i et al [41] have sucshy
cessfully demonstrated the effectiveness of an offset strategy to damp undesired v i shy
brations Passive energy dissipative devices eg viscous dampers are also another
viable solution to the problem
The development of various control laws to implement the above mentioned
strategies has also recieved much attention A n eigenstructure assignment in conducshy
tion wi th an offset controller for vibration attenuation and momemtum wheels for
platform libration control has been developed[42] in addition to the controller design
from a graph theoretic approach[41] Also non-linear feedback methods such as the
Feedback Linearization Technique ( F L T ) that are more suitable for controlling highly
nonlinear non-autonomous coupled systems have also been considered[43]
It is important to point out that several linear controllers including the classhy
sic state feedback Linear Quadratic Regulator ( L Q R ) have received considerable
attention[3944] Moreover robust methods such as the Linear Quadratic Guas-
s ian Loop Transfer Recovery ( L Q G L T R ) method have also been developed and imshy
plemented on tethered systems[43] A more complete review of these algorithms as
well as others applied to tethered system has been presented by Pradhan[43]
13 Scope of the Investigation
The objective of the thesis is to develop and implement a versatile as well as
computationally efficient formulation algorithm applicable to a large class of tethered
satellite systems The distinctive features of the model can be summarized as follows
12
bull TV-body 0(N) tethered satellite system in a chain-type topology
bull the system is free to negotiate a general Keplerian orbit and permitted to
undergo three dimensional inplane and out-of-plane l ibrtational motion
bull r igid bodies constituting the system are free to execute general rotational
motion
bull three dimensional flexibility present in the tether which is discretized using
the continuum assumed-mode method with an arbitrary number of flexible
modes
bull energy dissipation through structural damping is included
bull capability to model various mission configurations and maneuvers including
the tether spin about an arbitrary axis
bull user-defined time dependent deployment and retrieval profiles for the tether
as well as the tether attachment point (offset)
bull attitude control algorithm for the tether and rigid bodies using thrusters and
momentum-wheels based on the Feedback Linearization Technique
bull the algorithm for suppression of tether vibration using an active offset (tether
attachment point) control strategy based on the optimal linear control law
( L Q G L T - R )
To begin with in Chapter 2 kinematics and kinetics of the system are derived
using the O(N) formulation methodology developed by Pradhan et al [30] Chapshy
ter 3 discusses issues related to the development of the simulation program and its
validation This is followed by a detailed parametric study of several mission proshy
files simulated as particular cases of the versatile formulation in Chapter 4 Next
the attitude and vibration controllers are designed and their effectiveness assessed in
Chapter 5 The thesis ends with concluding remarks and suggestions for future work
13
2 F O R M U L A T I O N OF T H E P R O B L E M
21 Kinematics
211 Preliminary definitions and the i ^ position vector
The derivation of the equations of motion begins with the definition of the
inertial position of an arbitrary link of the mult ibody system Let the i ^ link of the
system be free to translate and rotate in 3-D space From Figure 2-1 the position
vector Rdm^ to the mass element drrn on the i ^ link wi th reference to the inertial
frame FQ can be written as
Here D represents the inertial position of the t h body fixed frame Fj relative to FQ
ri = [xiyiZi]T is the rigid position vector of drrii wi th respect to F J and f(fi) is
the flexible deformation at fj also with respect to the frame Fj Bo th these vectors
are relative to Fj Note in the present model tethers (i even) are considered flexible
and X corresponds to the nominal position along the unstretched tether while a and
ZJ are by definition equal to zero On the other hand the rigid bodies referred to as
satellites (i odd) including the platform are taken to be rigid ie f(fj) = 0 T j is
the rotation matrix used to express body-fixed vectors with reference to the inertial
frame
212 Tether flexibility discretization
The tether flexibility is discretized with an arbitrary but finite number of
Di + Ti(i + f(i)) (21)
14
15
modes in each direction using the assumed-mode method as
Ui nvi
wi I I i = 1
nzi bull
(22)
where nXi nyi and n 2 ^ are the number of modes used in the X m and Z directions
respectively For the ] t h longitudinal mode the admissible mode shape function is
taken as 2 j - l
Hi(xiii)=[j) (23)
where l is the i ^ tether length[32] In the case of inplane and out-of-plane transverse
deflections the admissible functions are
ampyi(xili) = ampZixuk) = v ^ s i n JKXj
k (24)
where y2 is added as a normalizing factor In this case both the longitudinal and
transverse mode shapes satisfy the geometric boundary conditions for a simple string
in axial and transverse vibration
Eq(22) is recast into a compact matrix form as
ff(i) = $i(xili)5i(t) (25)
where $i(x l) is the matrix of the tether mode shapes defined as
0 0
^i(xiJi) =
^ X bull bull bull ^XA
0
0 0
0
G 5R 3 x n f
(26)
and nfj = nx- + nyi + nZi Note $J(XJJ) is a function of the spatial coordinate x
and not of yi or ZJ However it is also a function of the tether length parameter li
16
which is time-varying during deployment and retrieval For this reason care must be
exercised when differentiating $J(XJJ) with respect to time such that
d d d k) = Jtregi(Xi li) = ^ 7 $ t ( ^ i k)Xi + gf$i(xigt li)k- (2-7)
Since x represents the nominal rate of change of the position of the elemental mass
drrii it is equal to j (deploymentretrieval rate) Let t ing
( 3 d
dx~ + dT)^Xili^ ( 2 8 )
one arrives at
$i(xili) = $Di(xili)ii (29)
The time varying generalized coordinate vector 5 in Eq(25) is composed of
the longitudinal and transverse components ie
ltM) G s f t n f i x l (210)
where 5X^ 5y- and Sz^ are nx- x 1 ny x 1 and nz- x 1 size vectors and correspond to
$ X and $ Z respectively
213 Rotation angles and transformations
The matrix T j in Eq(21) represents a rotation transformation from the body
fixed frame Fj to the inertial frame FQ ie FQ = T j F j It is defined by the 3-2-1
ordered sequence of elementary rotations[46]
1 P i tch F[ = Cf ( a j )F 0
2 R o l l F = C f (3j)F = C7f (3j)Cf ( a j )F 0
3 Yaw F = C7( 7 i )F = C7(7l)cf^)Cf(ai)FQ = Q F 0
17
where
c f (A) =
-Sai Cai
0 0
and
C(7i ) =
0 -s0i] = 0 1 0
0 cpi
i 0 0 = 0
0
(211)
(212)
(213)
Here Eq(211) corresponds to a rotation of ctj (pitch) about the inertial axis ZQ (orbit
normal) of frame FQ resulting in a new intermediate frame F It is then followed by
a roll rotation fa about the axis y of F- giving a second intermediate frame F-
Final ly a spin rotation 7J about the axis z of F- is given resulting in the body
fixed frame Fj for the i lt l link Since all the three rotation matrices are orthogonal
it follows that FQ = C~1Fi = Cj F and hence
C CPi Cai ~ CH Sai + SrYi SPi Sai 5 7 i Sampi + Cli SPi Cci CPi Sai C1i Cai + 5 7 i SPi ~ S1 Ca + C 7 bull Sp Sai
(214) SliCPi CliCPi
Here Cx and Sx are abbreviations for cos(x) and sin(x) respectively
214 Inertial velocity of the ith link
mdash Lett ing pj = fj - f $j(5j and differentiating Eq(21) with respect to time gives
m- D j + T j f j + T^Si + T j pound + T j $ j 5 j (215)
Each of the terms appearing in Eq(215) must be addressed individually Since
fj represents the rigid body motion within the link it only has meaning for the
deployable tether and since yi = Zj = 0 for al l tether links (i even) fj = Ifi where 1
is the unit vector 1 0 0T For the case of rigid links (i odd) fj = 0
18
Evaluating the derivative of each angle in the rotation matrix gives
T^i9i = Taiai9i + Tppigi + T^fagi (216)
where Tx = J j I V Collecting the scaler angular velocity terms and defining the set
Vi = deg^gt A) 7i^ Eq(216) can be rewritten as
T i f t = Pi9i)fji (217)
where Pi(gi) is a column matrix defined by
Pi(9i)Vi = [Taigi^p^iT^gilffi (218)
Inserting Eq(29) and Eq(217) into Eq(215) and defining Sj = 2 + leads to
215 Cylindrical orbital coordinates
The Di term in Eq(219) describes the orbital motion of the t h l ink and
is composed of the three Cartesian coordinates DXi A ^ DZi However over the
cycle of one orbit DXi and Dyi vary dramatically by around the order of Earths
radius This must be avoided since large variations in the coordinates can cause
severe truncation errors during their numerical integration For this reason it is
more convenient to express the Cartesian components in terms of more stationary
variables This is readily accomplished using cylindrical coordinates
However it is only necessary to transform the first l inks orbital components
ie Lgti since only D is a generalized coordinates From Figure 2-2 it is apparent
(219)
that
(220)
19
Figure 2-2 Vector components in cylindrical orbital coordinates
20
Eq(220) can be rewritten in matrix form as
cos(0i) 0 0 s i n ( 0 i ) 0 0
0 0 1
Dri
h (221)
= DTlDSv
where Dri is the inplane radial distance 9 the true anomaly and DZl the out-of-
plane distance normal to the orbital plane The total derivative of D wi th respect
to time gives in column matrix form
1 1 1 1 J (222)
= D D l D 3 v
where [cos(^i) -Dnsm(8i) 0 ]
(223)
For all remaining links ie i = 2 N
cos(0i) - A - j S i n ^ i ) 0 sin(^i) D r i c o s ( 0 i ) 0
0 0 1
DTl = D D i = I d 3 i = 2N (224)
where is the n x n identity matrix and
2N (225)
216 Tether deployment and retrieval profile
The deployment of the tether is a critical maneuver for this class of systems
Cr i t i ca l deployment rates would cause the system to librate out of control Normally
a smooth sinusoidal profile is chosen (S-curve) However one would also like a long
tether extending to ten twenty or even a hundred kilometers in a reasonable time
say 3-4 orbits This is usually difficult or impossible to accomplish solely wi th one
S-curve profile Often a composite S-curve-Steady-S-curve profile is adopted Thus
21
the deployment scheme for the t h tether can be summarized as
k = Tr l - cos (^r(ti - to-)) 1 ltUlt h 2 r V A V
k = Vov tiilttiltt2i
k = IT j 1 - C deg S lkitl ^ ^ lt i lt ^
(226)
where to- tt are the ini t ia l and final times of the deployment or retrieval maneuver
respectively A t j = t - tQ mdash tt mdash t2- and the steady deployment velocity VQ is
calculated based on the continuity of l and l at the specified intermediate times t
and to For the case of
and
Ai = i(ti) - ^(t 0 -) = - ^ ( t 2 i ) (227)
2Ali + li(tf) -h(tQ)
V = lUli l^lL (228)
1 tf-tQ v y
A t = (229)
22 Kinetics and System Energy
221 Kinetic energy
The kinetic energy of the ith link is given by
Kei = lJm Rdrnfi^Ami (230)
Setting Eq(219) in a matrix-vector form gives the relation
4 ^ ^ PM) T i $ 4 TiSi]i[t (231)
22
where
m
k J
x l (232)
and nq = -nfj + 7 is the number of generalized coordinates in each link Inserting
Eq(231) into Eq(230) and integrating the kinetic energy for the t h l ink can be
rewritten as
1 T
(233)
where Mti is the links symmetric mass matrix
Mt =
m i D D T D D DDTPifgidmi) DDTTi J ^dm D n T Sidm rn
sym fP(9i)Pi(9i)dmi J P[g^T^idm J P ( pound ) T ^ m sym sym J ^f^idm f ^f^dmi sym sym sym J sfsidmi
pound Sfttiqxnq
(234)
and mi is the mass of the Ith link The kinetic energy for the entire system ie N
bodies can now be stated as
1 bull T bull 1 bull T Ke = J2 MtA = o laquo Mtqt 2 ^ -i
i=l (235)
where qt = qj^^ bull bull -qfNT a n d Mt is a block diagonal decoupled mass matrix
of the system with Mf- on its diagonal
(236) Mt =
0 0
0 Mt2
0
0 0
MH
0 0 0
0 0 0 bull MtN
222 Simplification for rigid links
The mass matrix given by Eq(234) simplifies considerably for the case of rigid
23
links and is given as
Mti
m D D l
T D D l DD^Piiffidmi) 0
sym 0
0
fP(fi)Pi(i)dmi 0 0 0
0 Id nfj
0 m~
i = 1 3 5 N
(237)
Moreover when mult iplying the matrices in the (11) block term of Eq(237) one
gets
D D l
1 D D l = 1 0 0 0 D n
2 0 (238) 0 0 1
which is a diagonal matrix However more importantly i t is independent of 6 and
hence is always non-singular ensuring that is always invertible
In the (12) block term of M ^ it is apparent that f fidmi is simply the
definition of the centre of mass of the t h rigid link and hence given by
To drrii mdash miff (239)
where fcmi is the position vector of the centre of mass of link i For the case of the
(22) block term a little more attention is required Expanding the integrand using
Eq(218) where ltj = fj gives
fj Ta Tafi fj T Q Tpfi f- T Q T 7 i f j
sym sym
Now since each of the terms of Eq(240) is a scaler or a 1 x 1 matrix it is therefore
equal to its trace ie sum of its diagonal elements thus
tr(ffTai
TTaifi) tr(ffTai
TTpfi) tr(ff Ta^T^fi)
fjT^Tpn rfTp^fi (240)
P(rj)Pj(fi) = syrri trifjTpTTpfi) t r ^ T ^ T ^
sym sym tr(ff T 7 T r 7 f j) J
(241)
Using two properties of the trace function[47] namely
tr(ABC) = tr(BCA) = trCAB) (242)
24
and
j tr(A) =tr(^J A^j (243)
where A B and C are real matrices it becomes quite clear that the (22) block term
of Eq(237) simplifies to
P(fi)Pii)dmi =
tr(Tai
TTaiImi) tTTai
TTpImi) t r (T Q r T 7 7 m ) trTaTTpImi) tr(Tp
TT7Jmi) sym sym sym
(244)
where Im- is an alternate form of the inertia tensor of the t h link which when
expanded is
Imlaquo = I nfi dm bull i sym
sym sym
J x^dmi J Xydmi XZdm sym J yi2dmi f yiZidm
sym J z^dmi J
xVi -L VH sym ijxxi + lyy IZZJ)2
(245)
Note this is in terms of moments of inertia and products of inertia of the t f l rigid
link A similar simplification can be made for the flexible case where T is replaced
with Qi resulting in a new expression for Im
223 Gravitational potential energy
The gravitational potential energy of the ith link due to the central force law
can be written as
V9i f dm f
-A mdash = -A bullgttradelti RdmA
dm
inn Lgti + T j ^ | (246)
where z = 3986 x 105 km 3s 2 (Earths gravitational constant) Expanding binomially
and retaining up to third order terms gives
Vgi ~ ~ D T 2D [J gfgidrrii + 23f T j gidm mdashDT
D2 i l I I Df Tt I grffdmiTiDi
(2-lt 7)
25
where D = D is the magnitude of the inertial position vector to the frame F The
integrals in the above expression are functions of the flexible mode shapes and can be
evaluated through symbolic manipulation using an existing package such as Maple
V In turn Maple V can translate the evaluated integrals into F O R T R A N and store
the code in a file Furthermore for the case of rigid bodies it can be shown quite
readily that
j gjgidrrii = j rjfidrrn = [IXXi + Iyy + IZZi)2 (248)
Similarly the other integrals in Eq(234) can be integrated symbolically
224 Strain energy
When deriving the elastic strain energy of a simple string in tension the
assumptions of high tension and low amplitude of transverse vibrations are generally
valid Consequently the first order approximation of the strain-displacement relation
is generally acceptable However for orbiting tethers in a weak gravitational-gradient
field neglecting higher order terms can result in poorly modelled tether dynamics
Geometric nonlinearities are responsible for the foreshortening effects present in the
tether These cannot be neglected because they account for the heaving motion along
the longitudinal axis due to lateral vibrations The magnitude of the longitudinal
oscillations diminish as the tether becomes shorter[35]
W i t h this as background the tether strain which is derived from the theory
of elastic vibrations[3248] is given by
(249)
where Uj V and W are the flexible deformations along the Xj and Z direction
respectively and are obtained from Eq(25) The square bracketed term in Eq(249)
represents the geometric nonlinearity and is accounted for in the analysis
26
The total strain energy of a flexible link is given by
Ve = I f CiEidVl (250)
Substituting the stress-strain relationship 0 = E(poundi the strain energy is now given
by
Vei=l-EiAi J l l d x h (251)
where E^ is the tethers Youngs modulus and A is the cross-sectional area The
tether is assumed to be uniform thus EA which is the flexural stiffness of the
tether is assumed to be constant Eq(251) is evaluated symbolically for arbitrary
number of modes
225 Tether energy dissipation
The evaluation of the energy dissipation due to tether deformations remains
a problem not well understood even today Here this complex phenomenon is repshy
resented through a simplified structural damping model[32] In addition the system
exhibits hysteresis which also must be considered This is accomplished using an
augmented complex Youngs modulus
EX^Ei+jEi (252)
where Ej^ is the contribution from the structural damping and j = The
augmented stress-strain relation is now given as
a = Epoundi = Ei(l+jrldi)ei (253)
where 77^ = EjE^ ltC 1 is defined as the structural damping coefficient determined
experimentally [49]
27
If poundi is a harmonic function with frequency UJQ- then
jet = (254) wo-
Substituting into Eq(253) and rearranging the terms gives
e + I mdash I poundi = o-i + ad (255)
where ad and a are the stress with and without structural damping respectively
Now the Rayleigh dissipation function[50] for the ith tether can be expressed as
(ii) Jo
_ 1 EjAiVdi rU (2-56^ 2 w 0 Jo
The strain rate ii is the time derivative of Eq(249) The generalized external force
due to damping can now be written as Qd = Q^ bull bull bull QdNT where
dWd
^ = - i p = 2 4 ( 2 5 7 )
= 0 i = 13 N
23 0(N) Form of the Equations of Mot ion
231 Lagrange equations of motion
W i t h the kinetic energy expression defined and the potential energy of the
whole system given by Pe = ]Cn=l( 5i+ ej) the equations of motion can be obtained
quite readily using the Lagrangian principle
where q is the set of generalized coordinates to be defined in a later section
Substituting Eqs(235247251 and 257) into Eq(258) leads to the familiar
matr ix form of the non-linear non-autonomous coupled equations of motion for the
28
system
Mq t)t+ F(q q t) = Q(q q t) (259)
where M(q t) is the nonlinear symmetric mass matrix and F(q q t) is the forcing
function which can be written in matrix form as
Q(ltfgt Qi 0 is the vector of non-conservative generalized forces including control inshy
puts acting on the system A detailed expansion of Eq(260) in tensor notation is
developed in Appendix I
Numerical solution of Eq(259) in terms of q would require inversion of the
mass matrix However M is a full matrix of size nqq x nqq where nqq = Nnq is the
total number of generalized coordinates Therefore direct inversion of M would lead
to a large number of computation steps of the order nqq^ or higher The objective
here is to develop an order-N O(N) algorithm for obtaining M _ 1 that minimizes
the computational effort This is accomplished through the following transformations
ini t ia l ly developed for the stationkeeping case by Pradhan et al [30]
232 Generalized coordinates and position transformation
During the derivation of the energy expressions for the system the focus was
on the decoupled system ie each link was considered independent of the others
Thus each links energy expression is also uncoupled However the constraints forces
between two links must be incorporated in the final equations of motion
Let
e Unqxl (261) m
be the set of coupled coordinates of the t h l ink such that q = qT q2 bull bull bull 0T i s the
29
systems generalized coordinates Let qt- be the set of auxiliary decoupled coordinates
defined by Eq(232) such that qt = qj^ qj^ Qtj^1 bull The only difference between
qtj and q is the presence of D against d is the position of F wi th respect to
the inertial frame FQ whereas d is defined as the offset position of F relative to and
projected on the frame It can be expressed as d = dXidyidZi^ For the
special case of link 1 D = d
From Figure 2-1 can written as
Di = A - l + T i _ i J i _ i + di) + T V i S i - i f t - i + dXi)8i- (262)
Denoting KAj = $j(Zj + dx- ) Eq(262) can be rewritten in summation form as
i ampi = H ( T i - 1 ^ + T i - l K A i - l - l + T j - l ^ - l ) bull ( 2 - 6 3 )
3 = 1
Introducing the index substitution k mdash j mdash 1 into Eq(263)
i-1 D = ^2 (Tfc_ilt4 + TkKAk6k + Tkilk) + T^di (264)
k=l
since d = D TQ = 1^ and IQ DQ SQ are null vectors Recasting Eq(264) in
matr ix form leads to i-1
Qt Jfe=l
where
and
For the entire system
T fc-1 0 TkKAk
0 0 0 0 0 0 0 0 0 0 0 0 -
r T i - i 0 0 0 0 0 0 0 0 0 G
0 0 0 1
Qt = RpQ
G R n lt x l
(265)
(266)
G Unq x l (267)
(268)
30
where
RP
R 1 0 0 R R 2 0 R RP
2 R 3
0 0 0
R RP R PN _ 1 RN J
(269)
Here RP is the lower block triangular matrix which relates qt and q
233 Velocity transformations
The two sets of velocity coordinates qti and qi are related by the following
transformations
(i) First Transformation
Differentiating Eq(262) with respect to time gives the inertial velocity of the
t h l ink as
D = A - l + Ti_iK + + K A i _ i ^ _ i
+ T j _ i ^ + Ti-ii + T i - i K A i - i ^ + T j - i K A ^ i ^ i (270)
(271) (
Defining K A D = dddXi+1KAi K A L = d d K A and hi = di+1 +lfi +KA^i
results in K A ^ i = K A D j i o ^ + K A L i _ i i _ i
= KADi_iiTdi + K A L j _ i ^ _ i
Inserting Eq(271) into Eq(270) and rearranging gives
= A _ + T i ^ d g + K A D ^ x ^ i ^ ^ + Pi-^hi-iH-i
+ T i _ i K A i _ i 5 _ i + T i_ i2 + K A L j _ i 5 j _ i 4 _ i (272)
Using Eq(272) to define recursively - D j _ i and applying the index substitution
c = j mdash 1 as in the case of the position transformation it follows that
Di = ^ T ^ i K D ^ ^ + Pk(hk)ffk + T f c K A f c ^ + TfcKLfc4 Jfc=l (273)
+ T j _ i K D j _ i lt i ii
31
where K D j _ i = + K A D ^ i ^ - i r 7 and K L j = i + K A L j 5 j Finally by following a
procedure similar to that used for the position transformation it can be shown that
for the entire system
t = Rvil (2-74)
where Rv is given by the lower block diagonal matrix
Here
Rv
Rf 0 0 R R$ 0 R Rl RJ
0 0 0
T3V nN-l
R bullN
e Mnqqxl
Ri =
T i _ i K D i _ i TiKAi T j K L j 0 0 0 0 bullo 0 0 0 0 0 0 0
e K n lt x l
RS
T i-1 0 0 o-0 0 0 0 0 0
0 0 0 1
G Mnq x l
(275)
(276)
(277)
(ii) Second Transformation
The second transformation is simply Eq(272) set into matrix form This leads
to the expression for qt as
(278)
where Id Pihi) T j K A j T j K L j 0 0 0 0 0 0
0 0
0 0
0 0
G $lnq x l
Thus for the entire system
(279)
pound = Rnjt + Rdq
= Vdnqq - R nrlRdi (280)
32
where
and
- 0 0 0 RI 0 0
Rn = 0 0
0 0
Rf 0 0 0 rgtd K2 0
Rd = 0 0 Rl
0 0 0
RN-1 0
G 5R n w x l
0 0 0 e 5 R N ^ X L
(281)
(282)
234 Cylindrical coordinate modification
Because of the use of cylindrical coorfmates for the first body it is necessary
to slightly modify the above matrices to account for this coordinate system For the
first link d mdash D = D ^ D S L thus
Qti = 01 =gt DDTigi (283)
where
D D T i
Eq(269) now takes the form
DTL 0 0 0 d 3 0 0 0 ID
0 0 nfj 0
RP =
P f D T T i R2 0 P f D T T i Rp
2 R3
iJfDTTi i^
x l
0 0 0
R PN-I RN
(284)
(285)
In the velocity transformation a similar modification is necessary since D =
DD^DS^ Mak ing the substitution it can be shown that Eq(276) and Eq(279) now
33
takes the form for i = 1
T)V _ pn _ r t j mdash rt^ mdash 0 0 0 0
0 0 0 0 0 0 0 0
(286)
The rest of the matrices for the remaining links ie i = 2 N are unchanged
235 Mass matrix inversion
Returning to the expression for kinetic energy substitution of Eq(274) into
Eq(235) gives the first of two expressions for kinetic energy as
Ke = -q RvT MtRv (287)
Introduction of Eq(280) into Eq(235) results in the second expression for Ke as
1X 2q (hnqq-RnrlRd Mt (idnqq - R nrlRd (288)
Thus there are two factorizations for the systems coupled mass matrix given as
M = RV MtRv
M (hnqq - RnrlRd Mt ( i d n q q - R n r 1 R d
(289)
(290)
Inverting Eq(290)
r-1 _ ( rgtd~^ M 1 = (Ra) l d n q q - Rn)Mf1 (Rd)~l(Idnqq - Rn) (291)
Since both Rd and Mt are block diagonal matrices their inverse is simply the inverse
34
of each block on the diagonal ie
Mr1
h 0 0
0 0
0 0
0 0 0
0
0
0
tN
(292)
Each block has the dimension nq x nq and is independent of the number of bodies
in the system thus the inversion of M is only linearly dependent on N ie an O(N)
operation
236 Specification of the offset position
The derivation of the equations of motion using the O(N) formulation requires
that the offset coordinate d be treated as a generalized coordinate However this
offset may be constrained or controlled to follow a prescribed motion dictated either by
the designer or the controller This is achieved through the introduction of Lagrange
multipliers[51] To begin with the multipliers are assigned to all the d equations
Thus letting f = F - Q Eq(259) becomes
Mq + f = P C A (293)
where A G K 3 - 7 V x l is the vector of Lagrange multipliers and Pc G s R n 9 x 3 A r is the
permutation matrix assigning the appropriate Aj to its corresponding d equation
Inverting M and pre-multiplying both sides by PcT gives
pcTM-pc
= ( dc + PdegTM-^f) (294)
where dc is the desired offset acceleration vector Note both dc and PcTM lf are
known Thus the solution for A has the form
A pcTM-lpc - l dc + PC1 M (295)
35
Now substituting Eq(295) into Eq(293) and rearranging the terms gives
q = S(f Ml) = -M~lf + M~lPc
which is the new constrained vector equation of motion with the offset specified by
dc Note that pre-multiplying M~l by P c T provides the rows of M _ 1 corresponding
to the d equations Similarly post-multiplying by Pc leads the columns of the matrix
corresponding to the d equations
24 Generalized Control Forces
241 Preliminary remarks
The treatment of non-conservative external forces acting on the system is
considered next In reality the system is subjected to a wide variety of environmental
forces including atmospheric drag solar radiation pressure Earth s magnetic field as
well as dissipative devices to mention a few However in the present model only
the effect of active control thrusters and momemtum-wheels is considered These
generalized forces wi l l be used to control the attitude motion of the system
In the derivation of the equations of motion using the Lagrangian procedure
the contribution of each force must be properly distributed among all the generalized
coordinates This is acheived using the relation
nu ^5 Qk=E Pern ^ (2-97)
m=l deg Q k
where Qk and qk are the kth elements of Q and ltf respectively[51] Fem is the
mth external force acting on the system at the location Rm from the inertial frame
and nu is the total number of actuators Derivation of the generalized forces due to
thrusters and momentum wheels is given in the next two subsections
pcTM-lpc (dc + PcTM- Vj (2-96)
36
242 Generalized thruster forces
One set of thrusters is placed on each subsatellite ie al l the rigid bodies
except the first (platform) as shown in Figure 2-3 They are capable of firing in the
three orthogonal directions From the schematic diagram the inertial position of the
thruster located on body 3 is given as
Defining
R3 = di + Tid2 + T2(hi + KA2lt52) + T 3 f c m 3
= d + dfY + T 3 f c m 3
df = Tjdj+i + Tj+1lj+ii + KAj+i5j+i)
the thruster position for body 5 is given as
(298)
(299)
R5 = di+ dfx + d 3 + T5fcm
Thus in general the inertial position for the thruster on the t h body can be given
(2100)
as i - 2
k=l i1 cm (2101)
Let the column matrix
[Q i iT 9 R-K=rR dqj lA n i
WjRi A R i (2102)
where i = 3 5 N and j mdash 1 2 i The vector derivative of R wi th respect to
the scaler components qk is stored in the kth column of Thus for the case of
three bodies Eq(297) for thrusters becomes
1 Q
Q3
3 -1
T 3 T t 2 (2103)
37
Satellite 3
mdash gt -raquo cm
D1=d1
mdash gt
a mdash gt
o
Figure 2-3 Inertial position of subsatellite thruster forces
38
where ft = 0TtaTt^T is the thrust vector for the t h l ink (tether i-1) Here
Ttn and Tta are the thrust forces along the inplane m and out-of-plane Z directions i Pi
respectively For the case of 5-bodies
Q
~Qgt1
0
0
( T ^ 2 ) (2104)
The result for seven or more bodies follows the pattern established by Eq(2104)
243 Generalized momentum gyro torques
When deriving the generalized moments arising from the momemtum-wheels
on each rigid body (satellite) including the platform it is easier to view the torques as
coupled forces From Figure 2-4 it is apparent that for the ith l ink the generalized
forces constituting the couple are
(2105)
where
Fei = Faij acting at ei = eXi
Fe2 = -Fpk acting at e 2 = C a ^ l
F e 3 = F7ik acting at e 3 = eyj
Expanding Eq(2105) it becomes clear that
3
Qk = ^ F e i
i=l
d
dqk (2106)
which is independent of Ri- Thus the moments Mma = 2FaieXi Mm^ = 2FpeXi
and Mmi = 2FlieVi Transforming the coordinates to the body fixed frame using
the rotation matrix the generalized force due to the momentum-wheels on link i can
39
Figure 2-4 Coupled force representation of a momentum-wheel on a rigid body
40
be written as
d d d Qi = Td bull ^ T i t M m a - Tik bull mdashT^Mmp + T ^ bull T J j M ^
QiMmi
where Mm = Mma M m Mmi T
lt3 T J ^ T - T i f c bull ^ T i pound T ^ - ^ T z j
(2107)
(2108)
Note combining the thruster forces and momentum-wheel torques a compact
expression for the generalized force vector for 5 bodies can be written as
Q
Q T 3 0 Q T 5 0
0 0 0
0 lt33T3 Qyen Q5
3T5 0
0 0 0 Q4T5 0
0 0 0 Q5
5T5
M M I
Ti 2 m 3
Quu (2109)
The pattern of Eq(2109) is retained for the case of N links
41
3 C O M P U T E R I M P L E M E N T A T I O N
31 Prel iminary Remarks
The governing equations of motion for the TV-body tethered system were deshy
rived in the last chapter using an efficient O(N) factorization method These equashy
tions are highly nonlinear nonautonomous and coupled In order to predict and
appreciate the character of the system response under a wide variety of disturbances
it is necessary to solve this system of differential equations Although the closed-
form solutions to various simplified models can be obtained using well-known analytic
methods the complete nonlinear response of the coupled system can only be obtained
numerically
However the numerical solution of these complicated equations is not a straightshy
forward task To begin with the equations are stiff[52] be they have attitude and
vibration response frequencies separated by about an order of magnitude or more
which makes their numerical integration suceptible to accuracy errors i f a properly
designed integration routine is not used
Secondly the factorization algorithm that ensures efficient inversion of the
mass matrix as well as the time-varying offset and tether length significantly increase
the size of the resulting simulation code to well over 10000 lines This raises computer
memory issues which must be addressed in order to properly design a single program
that is capable of simulating a wide range of configurations and mission profiles It
is further desired that the program be modular in character to accomodate design
variations with ease and efficiency
This chapter begins with a discussion on issues of numerical and symbolic in-
42
tegration This is followed by an introduction to the program structure Final ly the
validity of the computer program is established using two methods namely the conshy
servation of total energy and comparison of the response results for simple particular
cases reported in the literature
32 Numerical Implementation
321 Integration routine
A computer program coded in F O R T R A N was written to integrate the equashy
tions of motion of the system The widely available routine used to integrate the difshy
ferential equations D G E A R [ 5 3 ] is based on Gears predictor-corrector method[54]
It is well suited to stiff differential equations as i t provides automatic step-size adjustshy
ment and error-checking capabilities at each iteration cycle These features provide
superior performance over traditional 4^ order Runge-Kut ta methods
Like most other routines the method requires that the differential equations be
transformed into a first order state space representation This is easily accomplished
by letting
thus
X - ^ ^ f ^ ) =ltbullbull) (3 -2 )
where S(fMdc) is defined by Eq(296) Eq(32) represents the new set of 2nqq
first order differential equations of motion
322 Program structure
A flow chart representing the computer programs structure is presented in
Figure 3-1 The M A I N program is responsible for calling the integration routine It
also coordinates the flow of information throughout the program The first subroutine
43
called is INIT which initializes the system variables such as the constant parameters
(mass density inertia etc) and the ini t ial conditions I C (attitude angles orbital
motion tether elastic deformation etc) from user-defined data files Furthermore all
the necessary parameters required by the integration subroutine D G E A R (step-size
tolerance) are provided by INIT The results of the simulation ie time history of the
generalized coordinates are then passed on to the O U T P U T subroutine which sorts
the information into different output files to be subsequently read in by the auxiliary
plott ing package
The majority of the simulation running time is spent in the F C N subroutine
which is called repeatedly by D G E A R F C N generates the fv(Xi) vector defined in
Eq(32) It is composed of several subroutines which are responsible for the comshy
putation of all the time-varying matrices vectors and scaler variables derived in
Chapter 2 that comprise the governing equations of motion for the entire system
These subroutines have been carefully constructed to ensure maximum efficiency in
computational performance as well as memory allocation
Two important aspects in the programming of F C N should be outlined at this
point As mentioned earlier some of the integrals defined in Chapter 2 are notably
more difficult to evaluate manually Traditionally they have been integrated numershy
ically on-line (during the simulation) In some cases particularly stationkeeping the
computational cost is quite reasonable since the modal integrals need only be evalushy
ated once However for the case of deployment or retrieval where the mode shape
functions vary they have to be integrated at each time-step This has considerable
repercussions on the total simulation time Thus in this program the integrals are
evaluated symbolically using M A P L E V[55] Once generated they are translated into
F O R T R A N source code and stored in I N C L U D E files to be read by the compiler
Al though these files can be lengthy (1000 lines or more) for a large number of flexible
44
INIT
DGEAR PARAM
IC amp SYS PARAM
MAIN
lt
DGEAR
lt
FCN
OUTPUT
STATE ACTUATOR FORCES
MOMENTS amp THRUST FORCES
VIBRATION
r OFFS DYNA
ET MICS
Figure 3-1 Flowchart showing the computer program structure
45
modes they st i l l require less time to compute compared to their on-line evaluation
For the case of deployment the performance improvement was found be as high as
10 times for certain cases
Secondly the O(N) algorithm presented in the last section involves matrices
that contained mostly zero terms (RP Rv and Mplusmn) The sparse structure of these mashy
trices can be used advantageously to improve the computational efficiency by avoiding
multiplications of zero elements The result is a quick evaluation of fv(Xt) Moreshy
over there is a considerable saving of computer memory as zero elements are no
longer stored
Final ly the C O N T R O L subroutine which is part of F C N is designed to calshy
culate the appropriate control forces torques and offset dynamics which are required
to stabilize the librational ( A T T I T U D E ) and vibrational ( V I B R A T I O N ) motion of
the system
33 Verification of the Code
The simulation program was checked for its validity using two methods The
first verified the conservation of total energy in the absence of dissipation The second
approach was a direct comparison of the planar response generated by the simulation
program with those available in the literature A s numerical results for the flexible
three-dimensional A^-body model are not available one is forced to be content with
the validation using a few simplified cases
331 Energy conservation
The configuration considered here is that of a 3-body tethered system with
the following parameters for the platform subsatellite and tether
46
h =
1091430 -8135 328108 -8135 8646050 27116 328108 27116 8286760
kg bull m 2 platform inertia
kg bull m 2 end-satellite inertia 200 0 0
0 400 0 0 0 400
m = 90000 kg mass of the space station platform
m2 = 500 kg mass of the end-satellite
EfAt = 61645 N tether elastic stiffness
pt mdash 49 kg km tether linear density
The tether length is taken to remain constant at 10 km (stationkeeping case)
A n in i t ia l disturbance in pitch and roll of 2deg and 1deg respectively is given to both
the rigid bodies and the flexible tether In addition the tether is ini t ia l ly deflected
by 45 m in the longitudinal direction at its end and 05 m in both the inplane and
out-of-plane directions in the first mode The tether attachment point offset at the
platform (satellite 1) end is 1 m in all the three directions ie d2 = 11 lTm The
attachment point at the subsatellite (satellite 2) end is 1 m in the x direction ie
fcm3 = 10 0T
The variation in the kinetic and potential energies is presented for the above-
mentioned case in Figure 3-2(a) As seen from the plot there is a continuous mutual
exchange between the kinetic and potential energies however the variation in the
total energy remains zero (Figure 3-2b) It should be noted that after 1 orbit the
variation of both kinetic and potential energy does not return to 0 This is due to
the attitude and elastic motion of the system which shifts the centre of mass of the
tethered system in a non-Keplerian orbit
Similar results are presented for the 5-body case ie a double pendulum (Figshy
ure 3-3) Here the system is composed of a platform and two subsatellites connected
47
STSS configuration 3-Body a 1(0)=a 2(0)=a t(0)=2deg P1(0)=p2(0)=p(0)=1deg 5x(0)=45m 8y(0)=52(0)=05m dx(0)=dy(0)=d2(0)=1m Stationkeeping lt=10km
Variation of Kinetic and Potential Energy
(a) Time (Orbits) Percent Variation in Total Energy
OOEOh
-50E-12
-10E-11 I i i
0 1 2 (b) Time (Orbits)
Figure 3-2 Kinetic and potential energy transfer for the three-body platform based tethered satellite system (a) variation of kinetic and potential energy (b) percent change in total energy of system
48
STSS configuration 5-Body a 1(0)=a 2(0)=a t 1(0)=2deg a 3(0)=a t 2(0)=25deg P1(0)=p2(0)=p3(0)=pt1(0)=pt2(0)=1lt
5x(0)=5m 8y(0)=5z(0)=05m dx(0)=dy(0)=dz(0)=0 Stationkeeping 1 =12=1 Okm
(a)
20E6
10E6r-
00E0
-10E6
-20E6
Variation of Kinetic and Potential Energy
-
1 I I I 1 I I
I I
AP e ~_
i i 1 i i i i
00E0
UJ -50E-12 U J
lt
Time (Orbits) Percent Variation in Total Energy
-10E-11h
(b) Time (Orbits)
Figure 3 -3 Kinet ic and potential energy transfer for the five-body platform based tethered satellite system (a) variation of kinetic and potential enshyergy (b) percent change in total energy of system
49
in sequence by two tethers The two subsatellites have the same mass and inertia
ie 7713 = 7 7 i 2 a n d I3 = I21 a n d are separated by identical tether segments each 10
k m in length The platform has the same properties as before however there is no
tether attachment point offset present in the system (d = fcmi = 0)
In al l the cases considered energy was found to be conserved However it
should be noted that for the cases where the tether structural damping deployment
retrieval attitude and vibration control are present energy is no longer conserved
since al l these features add or subtract energy from the system
332 Comparison with available data
The alternative approach used to validate the numerical program involves
comparison of system responses with those presented in the literature A sample case
considered here is the planar system whose parameters are the same as those given
in Section 331 A n ini t ia l pitch disturbance of 2deg is imparted to both the platform
and the tether together with an ini t ial tether deformation of 08 m and 001 m in
the longitudinal and transverse directions respectively in the same manner as before
The tether length is held constant at 5 km and it is assumed to be connected to the
centre of mass of the rigid platform and subsatellite The response time histories from
Ref[43] are compared with those obtained using the present program in Figures 3-4
and 3-5 Here ap and at represent the platform and tether pitch respectively and B i
C i are the tethers longitudinal and transverse generalized coordinates respectively
A comparison of Figures 3-4 and 3-5 clearly shows a remarkable correlation
between the two sets of results in both the amplitude and frequency The same
trend persisted even with several other comparisons Thus a considerable level of
confidence is provided in the simulation program as a tool to explore the dynamics
and control of flexible multibody tethered systems
50
a p(0) = 2deg 0(0) = 2C
6(0) = 08 m
0(0) = 001 m
Stationkeeping L = 5 km
D p y = 0 D p 2 = 0
a oo
c -ooo -0 01
Figure 3 - 4 Simulation results for the platform based three-body tethered system originally presented in Ref[43]
51
STSS configuration 3-Body a1(0)=a2(0)=at(0)=2deg p1(0)=(32(0)=(3t(0)=0 8x(0)=08m 5y(0)=001m 5Z(0)=0 dx(0)=dy(0)=dz(0)=0 Stationkeeping l=5km
Satellite 1 Pitch Angle Tether Pitch Angle
i -i i- i i i i i i d
1 2 0 1 2 Time (Orbits) Time (Orbits)
Longitudinal Vibration
^MAAAAAAAWVWWWWVVVW
V bull i bull i I i J 00 05 10 15 20
Time (Orbits) Transverse Vibration
Time (Orbits)
Figure 3-5 Simulation results for the platform based three-body tethered system obtained using the present computer program
52
4 D Y N A M I C SIMULATION
41 Prel iminary Remarks
Understanding the dynamics of a system is critical to its design and develshy
opment for engineering application Due to obvious limitations imposed by flight
tests and simulation of environmental effects in ground based facilities space based
systems are routinely designed through the use of numerical models This Chapter
studies the dynamical response of two different tethered systems during deployment
retrieval and stationkeeping phases In the first case the Space platform based Tethshy
ered Satellite System (STSS) which consists of a large mass (platform) connected to
a relatively smaller mass (subsatellite) with a long flexible tether is considered The
other system is the O E D I P U S B I C E P S configuration involving two comparable mass
satellites interconnected by a flexible tether In addition the mission requirement of
spin about an arbitrary axis the tether axis for O E D I P U S - A C and cartwheeling or
spin about the orbit normal for the proposed B I C E P S mission is accounted for
A s mentioned in Chapter 2 the tether flexibility is modelled using the asshy
sumed mode discretization method Although the simulation program can account
for an arbitrary number of vibrational modes in each direction only the first mode
is considered in this study as it accounts for most of the strain energy[34] and hence
dominates the vibratory motion The parametric study considers single as well as
double pendulum type systems with offset of the tether attachment points The
systems stability is also discussed
42 Parameter and Response Variable Definitions
The system parameters used in the simulation unless otherwise stated are
53
selected as follows for the STSS configuration
1091430 -8135 328108 -8135 8646050 27116 328108 27116 8286760
h =
h = 200 0 0
0 400 0 0 0 400
bull m = 90000 kg (mass of the space platform)
kg bull m 2 (platform inertia) 760 J
kg bull m 2 (subsatellite inertia)
bull 7772 = 500 kg (mass of the subsatellite)
bull EtAt = 61645 N (tether stiffness)
bull Pt = 49 k g k m (tether density)
bull Vd mdash 05 (tether structural damping coefficient)
bull fcm^ = l 0 0 r m (tether attachment point at the subsatellite)
The response variables are defined as follows
bull ai0 satellite 1 (platform) pitch and roll angles respectively
bull CK2 2 satellite 2 (subsatellite) pitch and roll angles respectively
bull at fit- tether pitch and roll angles respectively
bull If tether length
bull d = dx dy dzT- tether attachment position relative to satellite 1 (platform)
mdash
bull S = 8X 8y Sz tether flexible generalized coordinates in the longitudinal x
inplane transverse y and out-of-plane transverse z directions respectively
The attitude angles and are measured with respect to the Local Ve r t i c a l -
Loca l Horizontal ( L V L H ) frame A schematic diagram illustrating these variable
is presented in Figure 4-1 The system is taken to be in a nominal circular orbit at
an altitude of 289 km with an orbital period of 903 minutes
54
Satellite 1 (Platform) Y a w
Figure 4-1 Schematic diagram showing the generalized coordinates used to deshyscribe the system dynamics
55
43 Stationkeeping Profile
To facilitate comparison of the simulation results a reference case based on
the three-body STSS system with zero offset (d = 0 r c m 3 = 0) and a tether length
of It = 20 km is considered first (Figure 4-2) The system is subjected to an ini t ial
disturbance in pitch and roll of 2deg and 1deg respectively to all the three bodies
In addition an ini t ia l longitudinal deflection of 14 m from the tethers unstretched
position is given together with a transverse inplane and out-of-plane deflection of 1
m at the tethers mid-length From the attitude response given in Figure 4-2(a) it is
apparent that the three bodies oscillate about their respective equilibrium positions
Since coupling between the individual rigid-body dynamics is absent (zero offset) the
corresponding librational frequencies are unaffected Satellite 1 (Platform) displays
amplitude modulation arising from the products of inertia The result is a slight
increase of amplitude in the pitch direction and a significantly larger decrease in
amplitude in the roll angle
Figure 4-2(b) clearly shows coupling of the tethers attitude motion with its
flexibility dynamics The tether vibrates in the axial direction about its equilibrium
position (at approximately 138 m) with two vibration frequencies namely 012 Hz
and 31 x 1 0 - 4 Hz The former is the first longitudinal flexible mode which is dissishy
pated through the structural damping while the latter arises from the coupling with
the pitch motion of the tether Similarly in the transverse direction there are two
frequencies of oscillations arising from the flexible mode and its coupling wi th the
attitude motion As expected there is no apparent dissipation in the transverse flexshy
ible motion This is attributed to the weak coupling between the longitudinal and
transverse modes of vibration since the transverse strain in only a second order effect
relative to the longitudinal strain where the dissipation mechanism is in effect Thus
there is very litt le transfer of energy between the transverse and longitudinal modes
resulting in a very long decay time for the transverse vibrations
56
STSS configuration 3-Body a1(0)=a2(0)=a(0)=2deg (31(0)=p2(0)=p(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=dz(0)=0 Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
0 1 2 3 4 5 0 1 2 3 4 5 Time (Orbits) Time (Orbits)
Satellite 2 Pitch Angle Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-2 Stationkeeping dynamics of the three-body STSS configuration withshyout offset (a) attitude response
57
STSS configuration 3-Body a1(0)=cx2(0)=cx(0)=2deg P1(0)=p2(0)=p(0)=1deg 5x(0)=14m8y(0)=82(0)=1m dx(0)=dy(0)=dz(0)=0 Stationkeeping lt=20km
Tether Longitudinal Vibration 16| 1 1 1 1 1 r-
Time (Orbits) Tether In-Plane Transverse Vibration
imdash 1 mdash 1 mdash 1 mdash 1 mdashr
r I I 1 1 1 I 1 1 1 1 I 1 bull ^ bull bull I I J 0 1 2 3 4 5
Time (Orbits) Tether Out-of-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-2 Stationkeeping dynamics of the three-body STSS configuration withshyout offset (b) vibration response
58
Introduction of the attachment point offset significantly alters the response of
the attitude motion for the rigid bodies W i t h a i m offset along the local vertical
at the platform (dx = 1 m) and fcm^ = 100 m coupling between the tether
and end-bodies is established by providing a lever arm from which the tether is able
to exert a torque that affects the rotational motion of the rigid end-bodies (Figure
4-3) Bo th the platform and subsatellite now oscillate at the same frequency as the
tether whose motion remains unaltered In the case of the platform there is also an
amplitude modulation due to its non-zero products of inertia as explained before
However the elastic vibration response of the tether remains essentially unaffected
by the coupling
Providing an offset along the local horizontal direction (dy = 1 m) results in
a more dramatic effect on the pitch and roll response of the platform as shown in
Figure 4-4(a) Now the platform oscillates about its new equilibrium position of
-90deg The roll motion is also significantly disturbed resulting in an increase to over
10deg in amplitude On the other hand the rigid body dynamics of the tether and the
subsatellite as well as the tether flexible motion remain the same (Figure 4-4b)
Figure 4-5 presents the response when the offset of 1 m at the platform end is
along the z direction ie normal to the orbital plane The result is a larger amplitude
pitch oscillation about the reference equilibrium position as shown in Figure 4-5(a)
while the roll equilibrium is now shifted to approximately 90deg Note there is little
change in the attitude motion of the tether and end-satellites however there is a
noticeable change in the out-of-plane transverse vibration of the tether (Figure 4-5b)
which may be due to the large amplitude platform roll dynamics
Final ly by setting a i m offset in all the three direction simultaneously the
equil ibrium position of the platform in the pitch and roll angle is altered by approxishy
mately 30deg (Figure 4-6a) However the response of the system follows essentially the
59
STSS configuration 3-Body a1(0)=cx2(0)=cxt(0)=2deg p1(0)=(32(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=1m dy(0)=dz(0)=0 Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
Time (Orbits) Time (Orbits) Tether Pitch Angle Tether Roll Angle
L bull I I bull bull lt bull bull bull raquo bull bull bull bull -I h I I I I i H 0 1 2 3 4 5 0 1 2 3 4 5
Time (Orbits) Time (Orbits) Satellite 2 Pitch Angle Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-3 Stationkeeping dynamics of the three-body STSS configuration with offset along the local vertical (a) attitude response
60
STSS configuration 3-Body a 1(0)=a 2(0)=a t(0)=2deg (31(0)=p2(0)=pt(0)=1deg 5x(0)=14m5y(0)=5z(0)=1m dx(0)=1mdy(0)=dz(0)=0 Stationkeeping lt=20km
Tether Longitudinal Vibration 1 6 i mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash mdash i mdash i mdash mdash mdash mdash mdash i mdash mdash lt -
Time (Orbits)
Tether In-Plane Transverse Vibration
y bull i i i i i i i i i i i i i i i i i 0 1 2 3 4 5
Time (Orbits)
Tether Out-of-Plane Transverse Vibration r 1 1 1 1 1 1 1 1 1 1 1 1 1 1 q
h i I i i I i i i i I i i i i 3
0 1 2 3 4 5 (b) Time (Orbits)
Figure 4-3 Stationkeeping dynamics of the three-body STSS configuration with offset along the local vertical (b) vibration response
61
STSS configuration 3-Body a1(0)=cx2(0)=o(0)=2deg (31(0)=p2(0)=p(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dz(0)=0 dy(0)=1m Stationkeeping lt=20km
Satellite 1 Pitch Angle i 1 i
Satellite 1 Roll Angle
1 2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 4 Time (Orbits)
Tether Roll Angle
CD
edT
1 2 3 Time (Orbits)
Satellite 2 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 2 Roll Angle
CD
C O
(a)
Figure 4-
1 2 3 4 Time (Orbits)
1 2 3 4 Time (Orbits)
4 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal (a) attitude response
62
STSS configuration 3-Body a1(0)=cx2(0)=ot(0)=2deg p1(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dz(0)=0 dy(0)=1m Stationkeeping lt=20km
Tether Longitudinal Vibration 16| 1 gt 1 bull 1 1 1
0 1 2 3 4 5
Time (Orbits) Tether In-Plane Transverse Vibration
i mdash 1 mdash 1 mdash 1 mdash 1 mdash i mdash mdash lt mdash 1 mdash 1 mdash i mdash mdash 1 mdash 1 mdash
V I I bull bull bull I i i i I i i 1 1 3
0 1 2 3 4 5 Time (Orbits)
Tether Out-of-Plane Transverse Vibration r 1 1 1 1 1 1 1 1 1 1 1 gt bull | q
V i 1 1 1 1 1 mdash 1 mdash 3
0 1 2 3 4 5 (b) Time (Orbits)
Figure 4-4 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal (b) vibration response
63
STSS configuration 3-Body a1(0)=o2(0)=at(0)=2deg (31(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=0 dz(0)=1m Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 4 Time (Orbits)
Tether Roll Angle
1 2 3 4 Time (Orbits)
Satellite 2 Pitch Angle
2 3 4 Time (Orbits)
Satellite 2 Roll Angle
(a)
Figure 4-
1 2 3 4 Time (Orbits)
2 3 4 Time (Orbits)
5 Stationkeeping dynamics of the three-body STSS configuration with offset along the orbit normal (a) attitude response
64
STSS configuration 3-Body a1(0)=cx2(0)=ot(0)=2deg P1(0)=p2(0)=(3(0)=1deg 8x(0)=14m8y(0)=82(0)=1m dx(0)=dy(0)=0 d2(0)=1m Stationkeeping lt=20km
Tether Longitudinal Vibration 1 6 i mdash r mdash i mdash mdash i mdash i mdash mdash i mdash i mdash mdash i mdash mdash i mdash 1 mdash 1 mdash i mdash mdash lt mdash lt mdash lt mdash r
Time (Orbits) Tether In-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-5 Stationkeeping dynamics of the three-body STSS configuration with offset along the orbit normal (b) vibration response
65
same trend with only minor perturbations in the transverse elastic vibratory response
of the tether (Figure 4-6b)
44 Tether Deployment
Deployment of the tether from an ini t ial length of 200 m to 20 km is explored
next Here the deployment length profile is critical to ensure the systems stability
It is not desirable to deploy the tether too quickly since that can render the tether
slack Hence the deployment strategy detailed in Section 216 is adopted The tether
is deployed over 35 orbits with the sinusoidal acceleration and deceleration occurring
over a 4 km length ( A ^ 2 ) with the remaining deployment at a constant velocity of
V0 = 146 ms
Initially the tether is stretched only slightly S = 1304 x 1 0 _ 3 0 0 m
with al l the other states of the system ie pitch roll and transverse displacements
remaining zero The response for the zero offset case is presented in Figure 4-7 For
the platform there is no longer any coupling with the tether hence it oscillates about
the equilibrium orientation determined by its inertia matrix However there is st i l l
coupling between the subsatellite and the tether since rcm^ mdash 100 m resulting
in complete domination of the subsatellite dynamics by the tether Consequently
the pitch motion ini t ial ly grows by almost 50deg but eventually subsides as the tether
deployment rate decreases It is interesting to note that the out-of-plane motion is not
affected by the deployment This is because the Coriolis force which is responsible
for the tether pitch motion does not have a component in the z direction (as it does
in the y direction) since it is always perpendicular to the orbital rotation (Q) and
the deployment rate (It) ie Q x It
Similarly there is an increase in the amplitude of the transverse elastic oscilshy
lations of the tether again due to the Coriolis effect (Figure 4-7b) Moreover as the
66
STSS configuration 3-Body a1(0)=a2(0)=cxt(0)=2deg P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m5y(0)=52(0)=1m dx(0)=dy(0)=d2(0)=1m Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
i i i i i i i i i i i 0 1 2 3 4 5 0 1 2 3 4 5
Time (Orbits) Time (Orbits) Tether Pitch Angle Tether Roll Angle
l i i i i -I h i i i i I i i i i l- i i i i l i i i i I i i- i i H 0 1 2 3 4 5 0 1 2 3 4 5
Time (Orbits) Time (Orbits) Satellite 2 Pitch Angle Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-6 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal local vertical and orbit normal (a) attitude response
67
STSS configuration 3-Body a1(0)=cc2(0)=a(0)=2deg p1(0)=p2(0)=(3t(0)=1deg 5x(0)=14m6y(0)=52(0)=1m dx(0)=dy(0)=d2(0)=1m Stationkeeping lt=20km
Tether Longitudinal Vibration 1 6 1 mdash i mdash i mdash i mdash mdash i mdash 1 mdash 1 mdash mdash mdash i mdash mdash 1 mdash 1 mdash 1 mdash i mdash 1 mdash 1 mdash 1 mdash 1 mdash r
0 1 2 3 4 5
Time (Orbits) Tether In-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-6 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal local vertical and orbit normal (b) vibration response
68
STSS configuration 3-Body a(0)=a2(0)=ot(0)=0 P1(0)=p2(0)=pt(0)=0 8X(0)=1304 x 103m 8y(0)=8z(0)=0 dx(0)=dy(0)=dz(0)=0 Deployment
l=02km to 20km in 35 Orbits
Tether Length Profile
E
Satellite 1 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 1 Roll Angle mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash [ mdash i mdash i mdash i mdash i mdash | mdash
1 2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 4 Time (Orbits)
Tether Roll Angle
CD
oo ax
i imdashimdashimdashImdashimdashimdashimdashr-
_ 1 _ L
1 2 3 4 Time (Orbits)
Satellite 2 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 2 Roll Angle
-50
(a) 1 2 3 4
Time (Orbits) 1 2 3 4
Time (Orbits)
Figure 4-7 Deployment dynamics of the three-body STSS configuration without offset (a) attitude response
69
STSS configuration 3-Body a1(0)=a2(0)=at(0)=0 p1(0)=p2(0)=(3t(0)=0 8X(0)=1304x 103m 8y(0)=52(0)=0 dx(0)=dy(0)=d2(0)=0 Deployment
l=02km to 20km in 35 Orbits
Tether Longitudinal Vibration
to
-02
(b)
Figure 4-7
2 3 Time (Orbits)
Tether In-Plane Transverse Vibration
0 1 1 bull 1 1
1 2 3 4 Time (Orbits)
Tether Out-of-Plane Transverse Vibration
5
1 1 1 1 1 i i | i i | ^
_ i i i i L -I
1 2 3 4 5 Time (Orbits)
Deployment dynamics of the three-body STSS configuration without offset (b) vibration response
70
tether elongates its longitudinal static equilibrium position also changes due to an
increase in the gravity-gradient tether tension It may be pointed out that as in the
case of attitude motion there are no out-of-plane tether vibrations induced
When the same deployment conditions are applied to the case of 1 m offset in
the x direction (Figure 4-8) coupling effects are re-introduced The platform pitch
exceeds -100deg during the peak deployment acceleration However as the tether pitch
motion subsides the platform pitch response returns to a smaller amplitude oscillation
about its nominal equilibrium On the other hand the platform roll motion grows
to over 20deg in amplitude in the terminal stages of deployment Longitudinal and
transverse vibrations of the tether remain virtually unchanged from the zero offset
case except for minute out-of-plane perturbations due to the platform librations in
roll
45 Tether Retrieval
The system response for the case of the tether retrieval from 20 km to 200 m
with an offset of 1 m in the x and z directions at the platform end is presented in
Figure 4-9 Here the Coriolis force induced during retrieval renders the tether unstable
(Figure 4-9a) Consequently the pitch motion of the platform is also destabilized
through coupling The z offset provides an additional moment arm that shifts the
roll equilibrium to 35deg however this degree of freedom is not destabilized From
Figure 4-9(b) it is apparent that the transverse mode of the tether vibration is also
disturbed producing a high frequency response in the final stages of retrieval This
disturbance is responsible for the slight increase in the roll motion for both the tether
and subsatellite
46 Five-Body Tethered System
A five-body chain link system is considered next To facilitate comparison with
71
STSS configuration 3-Body a(0)=a2(0)=at(0)=0 P1(0)=P2(0)=P(0)=0 5X(0)=1304 x 103m 8y(0)=8z(0)=0 dx(0)=1mdy(0)=d2(0)=0 Deployment
lt=02km to 20km in 35 Orbits
Tether Length Profile
E
Satellite 1 Pitch Angle
1 2 3 Time (Orbits)
Satellite 1 Roll Angle
bull bull bull i bull bull bull bull i
1 2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 Time (Orbits)
Tether Roll Angle
i i i i J I I i _
1 2 3 4 Time (Orbits)
Satellite 2 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 2 Roll Angle 3E-3h
D )
T J ^ 0 E 0 eg
CQ
-3E-3
(a) 1 2 3 4
Time (Orbits) 1 2 3 4
Time (Orbits)
Figure 4-8 Deployment dynamics of the three-body STSS configuration with offset along the local vertical (a) attitude response
7 2
STSS configuration 3-Body a1(0)=a2(0)=at(0)=0 p1(0)=P2(0)=(3t(0)=0 Sx(0)=1304 x 103m 5y(0)=5z(0)=0 dx(0)=1mdy(0)=dz(0)=0 Deployment
l=02km to 20km in 35 Orbits
Tether Longitudinal Vibration
Time (Orbits) Tether In-Plane Transverse Vibration
n 1 1 1 f
I I I _j I I I I I 1 1 1 1 1 2 3 4 5
Time (Orbits) Tether Out-of-Plane Transverse Vibration
J 1 I I 1 I I I I I I I I L _ _ I I d 1 2 3 4 5
Time (Orbits)
Deployment dynamics of the three-body STSS configuration with offset along the local vertical (b) vibration response
(b)
Figure 4-8
73
STSS configuration 3-Body a1(0)=a2(0)=at(0)=0 p(0)=P2(0)=(3t(0)=0 Sx(0)=14m8y(0)=82(0)=0 dy(0)=0dx(0)=d2(0)=1m Retrieval
l=20km to 02km in 35 Orbits
Satellite 1 Pitch Angle
(a) Time (Orbits)
Figure 4-9 Retrieval dynamics of the along the local vertical ar
Tether Length Profile
i i i i i i i i i i I 0 1 2 3
Time (Orbits) Satellite 1 Roll Angle
Time (Orbits) Tether Roll Angle
b_ i i I i I i i L _ J
0 1 2 3 Time (Orbits)
Satellite 2 Roll Angle _ mdash | r i i | 1 i i imdashmdashj i 1 r mdash i mdash
04- -
0 1 2 3 Time (Orbits)
three-body STSS configuration with offset d orbit normal (a) attitude response
74
STSS configuration 3-Body a1(0)=a2(0)=ocl(0)=0 (31(0)=f32(0)=pt(0)=0 8x(0)=14m8y(0)=6z(0)=0 dy(0)=0dx(0)=dz(0)=1m Retrieval
l=20km to 02km in 35 Orbits
Tether Longitudinal Vibration
E 10 to
Time (Orbits) Tether In-Plane Transverse Vibration
1 2 Time (Orbits)
Tether Out-of-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-9 Retrieval dynamics of the three-body STSS configuration with offset along the local vertical and orbit normal (b) vibration response
75
the three-body system dynamics studied earlier the chain is extended through the
addition of two bodies a tether and a subsatellite with the same physical properties
as before (Section 42) Thus the five-body system consists of a platform (satellite 1)
tether 1 subsatellite 1 (satellite 2) tether 2 and subsatellite 2 (satellite 3) as shown
in Figure 4-10 The offset of tether 1 at the platform-end is denoted by d2 while that
of tether 2 to subsatellite 1 as 04 In this numerical example fcm^ = fcm^ = 0 ie
tethers 1 and 2 are attached to the centre of mass of subsatellites 1 and 2 respectively
The system response for the case where the tether attachment points coincide
wi th the centres of mass of the rigid bodies ie zero offset (d2 = d mdash 0) is presented
in Figure 4-11 Here ogtt and at2 represent the pitch motion of tethers 1 and 2
respectively whereas 03 is the pitch motions of satellite 3 A similar convention
is adopted for the rol l angle 0 A s before the zero offset eliminates the coupling
between the tethers and satellites such that they are now free to librate about their
static equil ibrium positions However their is s t i l l mutual coupling between the two
tethers This coupling is present regardless of the offset position since each tether is
capable of transferring a force to the other The coupling is clearly evident in the pitch
response of the two tethers (Figure 4 - l l a ) O n the other hand the roll motion appears
uncoupled as the in i t ia l conditions in rol l for the two tethers are identical Thus the
motion is in phase and there is no transfer of energy through coupling A s expected
during the elastic response the tethers vibrate about their static equil ibrium positions
and mutually interact (Figure 4 - l l b ) However due to the variation of tension along
the tether (x direction) they do not have the same longitudinal equilibrium Note
relatively large amplitude transverse vibrations are present particularly for tether 1
suggesting strong coupling effects with the longitudinal oscillations
76
Figure 4-10 Schematic diagram of the five-body system used in the numerical example
77
STSS configuration 5-Body a1(0)=(x2(0)=at1(0)=2deg a3(0)=a12(0)=25deg p1(0)=r32(0)=p3(0)=pt1(0)=pt2(0)=r 82(0)=84(0)=505)05m d2(0)=d4(0)=000 Stationkeeping lt1=lu=10km Pitch Roll
Satellite 1 Pitch Angle Satellite 1 Roll Angle
o 1 2 Time (Orbits)
Tether 1 Pitch and Roll Angle
0 1 2 Time (Orbits)
Tether 2 Pitch and Roll Angle
Time (Orbits) Satellite 2 Pitch and Roll Angle
Time (Orbits) Satellite 3 Pitch and Roll Angle
Time (Orbits) Time (Orbits)
Figure 4-11 Stationkeeping dynamics of the five-body STSS configuration withshyout offset (a) attitude response
78
STSS configuration 5-Body a1(0)=a2(0)=oct1(0)=2deg a3(0)=at2(0)=25o
p1(0)=p2(0)=(33(0)=pt1(0)=pt2(0)=10
62(0)=54(0)=50505m d2(0)=d4(0)=000 Stationkeeping lt1=1 =1 Okm
Tether 1 Longitudinal Vibration Tether 2 Longitudinal Vibration i i 1 i 1 r 1 1 1 1
(b) Time (Orbits) Time (Orbits)
Figure 4-11 Stationkeeping dynamics of the five-body STSS configuration withshyout offset (b) vibration response
79
The system dynamics during deployment of both the tethers in the double-
pendulum configuration is presented in Figure 4-12 Deployment of each tether takes
place from 200 m to 20 km in 35 orbits The pitch motion of the system is similar
to that observed during the three-body case The Coriolis effect causes the tethers
to pitch through a large angle ( laquo 50deg ) which in turn disturbs the platform and
subsatellite due to the presence of offset along the local vertical On the other hand
the roll response for both the tethers is damped to zero as the tether deploys in
confirmation with the principle of conservation of angular momentum whereas the
platform response in roll increases to around 20deg Subsatellite 2 remains virtually
unaffected by the other links since the tether is attached to its centre of mass thus
eliminating coupling Finally the flexibility response of the two tethers presented in
Figure 4-12(b) shows similarity with the three-body case
47 BICEPS Configuration
The mission profile of the Bl-static Canadian Experiment on Plasmas in Space
(BICEPS) is simulated next It is presently under consideration by the Canadian
Space Agency To be launched by the Black Brant 2000500 rocket it would inshy
volve interesting maneuvers of the payload before it acquires the final operational
configuration As shown in Figure 1-3 at launch the payload (two satellites) with
an undeployed tether is spinning about the orbit normal (phase 1) The internal
dampers next change the motion to a flat spin (phase 2) which in turn provides
momentum for deployment (phase 3) When fully deployed the one kilometre long
tether will connect two identical satellites carrying instrumentation video cameras
and transmitters as payload
The system parameters for the BICEPS configuration are summarized below [59 0 0 1
bull I = I2 = 0 366 0 kg bull m 2 (satellite inertias) [ 0 0 392_
bull mi = 7772 = 200 kg (mass of the satellites)
80
STSS configuration 5-Body a(0)=cx2(0)=al1(0)=2deg a3(0)=c (0)=25o
P1(0)=P2(0)=p3(0)=f3t1(0)=Pt2(0)=r 52(0)=84(0)=1304x 10300m d2(0)=d4(0)=1G0m Deployment lt1=le=02km to 20km in 35 Orbits
Pitch Roll Satellite 1 Pitch Angle
Time (Orbits) Satellite 2 Pitch and Roll Angle T 1mdash1 1 1 I I I I | I I I 1 I | I I lt~r
Tether Length Profile i 1 1 1 1 i 1 1 1 1 i 1 1 1 1 i
i i i i i 0 1 2 3 4 5
Time (Orbits) Satellite 1 Roll Angle
0 1 2 3 4 5 Time (Orbits)
Tether 1 amp 2 Roll Angle L i | I I 1 I 1 1 -I
-Ih I- bull I I i i i i I i i i i 1
0 1 2 3 4 5 Time (Orbits)
Satellite 3 Pitch and Roll Angle i- i 1 1 i 1 1 1 1 i 1 1 1 1 i 1 1 1 1 i
L i i i i i i 1 i i mdash i mdash i I i imdashimdashLJ
0 1 2 3 4 5 Time (Orbits)
Figure 4-12 Deployment dynamics of the five-body S T S S configuration with offshyset along the local vertical (a) attitude response
81
STSS configuration 5-Body a1(0)=a2(0)=oct1(0)=2deg a3(0)=at2(0)=25deg P1(0)=P2(0)=P3(0)=Pt1(0)=pt2(0)=1o
82(0)=54(0)=1304 x 10-300m d2(0)=d4(0)=100m Deployment lt1=l2=02km to 20km in 35 Orbits
Tether 1 Longitudinal Vibration Tether 2 Longitudinal Vibration
Time (Orbits) Time (Orbits) Tether 1 Z Tranverse Vibration Tether 2 Z Tranverse Vibration
Figure 4-12 Deployment dynamics of the five-body STSS configuration with offshyset along the local vertical (b) vibration response
82
bull EtAt = 61645 N (tether stiffness)
bull pt = 30 k g k m (tether density)
bull It = 1 km (tether length)
bull rid = 1-0 (tether structural damping coefficient)
The system is in a circular orbit at a 289 km altitude Offset of the tether
attachment point to the satellites at both ends is taken to be 078 m in the x
direction The response of the system in the stationkeeping phase for a prescribed set
of in i t ia l conditions is given by Figure 4-13 As in the case of the STSS configuration
there is strong coupling between the tether and the satellites rigid body dynamics
causing the latter to follow the attitude of the tether However because of the smaller
inertias of the payloads there are noticeable high frequency modulations arising from
the tether flexibility Response of the tether in the elastic degrees of freedom is
presented in Figure 4-13(b) Note the longitudinal vibrations decay quite rapidly
(lt 05 orbits) due to the structural damping however it has vir tual ly no effect on
the transverse oscillations
The unique mission requirement of B I C E P S is the proposed use of its ini t ia l
angular momentum in the cartwheeling mode to aid in the deployment of the tethered
system The maneuver is considered next The response of the system during this
maneuver wi th an ini t ia l cartwheeling rate of 5 deg s is illustrated in Figure 4-14(a) The
in i t ia l tether length is taken to be 10 m and is deployed to 1 km There is an in i t ia l
increase in the pitch motion of the tether however due to the conservation of angular
momentum the cartwheeling rate decreases proportionally to the square of the change
in tether deployment rate The result is a rapid drop in the cartwheeling rate unti l
the system stops rotating entirely and simply oscillates about its new equilibrium
point Consequently through coupling the end-bodies follow the same trend The
roll response also subsides once the cartwheeling motion ceases However
83
BICEPS configuration 3-Body a1(0)=o2(0)=at(0)=2deg P1(0)=P2(0)=pt(0)=1deg 8X(0)=001 m 6y(0)=52(0)=01 m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping l=1 km
Satellite 1 Pitch Angle Satellite 1 Roll Angle - i 1 1 1 1 i - i q r 1 1 -gt
I- I i t i i i i I H 0 1 2 0 1 2
Time (Orbits) Time (Orbits) Tether Pitch Angle Tether Roll Angle
F mdash 1 mdash 1 mdash mdash 1 mdash i mdash 1 mdash 1 mdash mdash 1 mdash = i F mdash 1 mdash mdash mdash 1 mdash i mdash 1 mdash 1 mdash 1 mdash 1 mdash q
(a) Time (Orbits) Time (Orbits)
Figure 4-13 Stationkeeping dynamics of the three-body BICEPS configuration with offset along the local vertical (a) attitude response
84
BICEPS configuration 3-Body a1(0)=o2(0)=at(0)=2deg Pl(0)=p2(0)=(3t(0)=1deg 8X(0)=001 m 8y(0)=82(0)=01 m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping lt=1km
Tether Longitudinal Vibration 1E-2[
laquo = 5 E - 3 |
000 002
Time (Orbits) Tether In-Plane Transverse Vibration
Time (Orbits) Tether Out-of-Plane Transverse Vibration
Time (Orbits)
Figure 4-13 Stationkeeping dynamics of the three-body BICEPS configuration with offset along the local vertical (b) vibration response
85
BICEPS configuration 3-Body a1(0)=o2(0)=at(0)=2deg a1(0)=a2(0)=at(0)=57s Pl(0)=P2(0)=(3t(0)=1deg 8X(0)=115 x 103m 8y(0)=82(0)=0 dx(0)=078m dy(0)=d2(0)=0 Cartwheeling Deployment
lt=10m to 1km in 1 Orbit
Tether Length Profile
E
2000
cu T3
2000
co CD
2000
C D CD
T J
Satellite 1 Pitch Angle
1 2 Time (Orbits)
Satellite 1 Roll Angle
cn CD
33 cdl
Time (Orbits) Tether Pitch Angle
Time (Orbits) Tether Roll Angle
Time (Orbits) Satellite 2 Pitch Angle
1 2 Time (Orbits)
Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-14 Cartwheeling dynamics with deployment for the three-body BICEPS with offset along the local vertical (a) attitude response
86
BICEPS configuration 3-Body a1(0)=a2(0)=at(0)=2deg a (0)=a2(0)=a(0)=57s P1(0)=P2(0)=f3(0)=1deg 8X(0)=115x103m 8y(0)=8z(0)=0 dx(0)=078m dy(0)=d2(0)=0 Cartwheeling Deployment
l=1 Om to 1 km in 1 Orbit
5E-3h
co
0E0
025
000 E to
-025
002
-002
(b)
Figure 4-14
Tether Longitudinal Vibration
Time (Orbits) Tether In-Plane Transverse Vibration
Time (Orbits) Tether Out-of-Plane Transverse Vibration
Time (Orbits)
Cartwheeling dynamics with deployment for the three-body BICEPS with offset along the local vertical (b) vibration response
87
as the deployment maneuver is completed the satellites continue to exhibit a small
amplitude low frequency roll response with small period perturbations induced by
the tether elastic oscillations superposed on it (Figure 4-14b)
48 OEDIPUS Spinning Configuration
The mission entitled Observation of Electrified Distributions of Ionospheric
Plasmas - A Unique Strategy ( O E D I P U S ) also consists of a 1 k m tethered system with
two similar end-masses It was first flown in a suborbital flight in 1989 ( O E D I P U S A )
followed by a recent study in November 1995 ( O E D I P U S - C ) in which the University
of Br i t i sh Columbia was one of the participants Dynamically O E D I P U S represents a
unique system never encountered before Launched by a Black Brant rocket developed
at Br is to l Aerospace L t d the system ie the end-bodies together wi th the tether
spins about the longitudinal axis of the tether to achieve stabilized alignment wi th
Earth s magnetic field
The response of the system undergoing a spin rate of j = l deg s is presented
in Figure 4-15 using the same parameters as those outlined for the B I C E P S conshy
figuration in Section 47 Again there is strong coupling between the three bodies
(two satellites and tether) due to the nonzero offset The spin motion introduces an
additional frequency component in the tethers elastic response which is transferred
to the l ibrational motion of the satellites A s the spin rate increases to 1 0 deg s (Figure
4-16) the amplitude and frequency of the perturbations also increase however the
general character of the response remains essentially the same
88
OEDIPUS configuration 3-Body a1(0)=a2(0)=at(0)=2deg i(0)^[2(0H(0)=Ys p1(0)=p2(0)=p(0)=r 8X(0)=12 x 103m 8y(0)=82(0)=01m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping l=1 km
Satellite 1 Pitch Angle Satellite 1 Roll Angle mdash i 1 1 1 a r 1 1 1 r
(a) Time (Orbits) Time (Orbits)
Figure 4-15 Spin dynamics (7 = ldegs) of the three-body OEDIPUS configuration with offset along the local vertical (a) attitude response
89
OEDIPUS configuration 3-Body a1(0)=a2(0)=cct(0)=2deg Y1(0)=Y2(0)=Yt(0)=l7s (31(0)=p2(0)=pt(0)=1deg 8X(0)=12 x 103m 8y(0)=5z(0)=01 m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping lt=1 km
1E-2
to
OEO
E 00
to
Tether Longitudinal Vibration
005
Time (Orbits) Tether In-Plane Transverse Vibration
Time (Orbits) Tether Out-of-Plane Transverse Vibration
Time (Orbits)
Figure 4-15 Spin dynamics (7 = l deg s ) of the three-body O E D I P U S configuration with offset along the local vertical (b) vibration response
90
OEDIPUS configuration 3-Body a1(0)=a2(0)=a(0)=2deg Yi(0)=Y2(0)=rt(0)=107s 81(0)=p2(0)=pt(0)=1deg 8X(0)=12 x 103m 8y(0)=82(0)=01m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping lt=1km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-16 Spin dynamics (7= 10degs) of the three-body OEDIPUS configura tion with offset along the local vertical (a) attitude response
91
OEDIPUS configuration 3-Body a1(0)=a2(0)=ot(0)=2deg Yi(0)H(0)4(0)laquo107s P1(0)=P2(0)=pt(0)=1deg 8X(0)=12 x 103m 8y(0)=82(0)=01m dx(0)=078m dy(0)=dz(0)=0 Stationkeeping lt=1 km
Tether Longitudinal Vibration
Figure 4-16 Spin dynamics (7 = 10degs) of the three-body OEDIPUS configurashytion with offset along the local vertical (b) vibration response
92
5 ATTITUDE AND VIBRATION CONTROL
51 Att itude Control
511 Preliminary remarks
The instability in the pitch and roll motions during the retrieval of the tether
the large amplitude librations during its deployment and the marginal stability during
the stationkeeping phase suggest that some form of active control is necessary to
satisfy the mission requirements This section focuses on the design of an attitude
controller with the objective to regulate the librational dynamics of the system
As discussed in Chapter 1 a number of methods are available to accomplish
this objective This includes a wide variety of linear and nonlinear control strategies
which can be applied in conjuction with tension thrusters offset momentum-wheels
etc and their hybrid combinations One may consider a finite number of Linear T ime
Invariant (LTI) controllers scheduled discretely over different system configurations
ie gain scheduling[43] This would be relevant during deployment and retrieval of
the tether where the configuration is changing with time A n alternative may be
to use a Linear T ime Varying ( L T V ) model in conjuction with an adaptive control
scheme where on-line parametric identification may be used to advantage[56] The
options are vir tually limitless
O f course the choice of control algorithms is governed by several important
factors effective for time-varying configurations computationally efficient for realshy
time implementation and simple in character Here the thrustermomentum-wheel
system in conjuction with the Feedback Linearization Technique ( F L T ) is chosen as
it accounts for the complete nonlinear dynamics of the system and promises to have
good overall performance over a wide range of tether lengths It is well suited for
93
highly time-varying systems whose dynamics can be modelled accurately as in the
present case
The proposed control method utilizes the thrusters located on each rigid satelshy
lite excluding the first one (platform) to regulate the pitch and rol l motion of the
tether to which it is attached ie the thrusters located on satellite 2 (link 3) regshy
ulates the attitude motion of tether 1 (link 2) O n the other hand the l ibrational
motion of the rigid bodies (platform and subsatellite) is controlled using a set of three
momentum wheels placed mutually perpendicular to each other
The F L T method is based on the transformation of the nonlinear time-varying
governing equations into a L T I system using a nonlinear time-varying feedback Deshy
tailed mathematical background to the method and associated design procedure are
discussed by several authors[435758] The resulting L T I system can be regulated
using any of the numerous linear control algorithms available in the literature In the
present study a simple P D controller is adopted arid is found to have good perforshy
mance
The choice of an F L T control scheme satisfies one of the criteria mentioned
earlier namely valid over a wide range of tether lengths The question of compushy
tational efficiency of the method wi l l have to be addressed In order to implement
this controller in real-time the computation of the systems inverse dynamics must
be executed quickly Hence a simpler model that performs well is desirable Here
the model based on the rigid system with nonlinear equations of motion is chosen
Of course its validity is assessed using the original nonlinear system that accounts
for the tether flexibility
512 Controller design using Feedback Linearization Technique
The control model used is based on the rigid system with the governing equa-
94
tions of motion given by
Mrqr + fr = Quu (51)
where the left hand side represents inertia and other forces while the right hand side
is the generalized external force due to thrusters and momentum-wheels and is given
by Eq(2109)
Lett ing
~fr = fr~ QuU (52)
and substituting in Eq(296)
ir = s(frMrdc) (53)
Expanding Eq(53) it can be shown that
qr = S(fr Mrdc) - S(QU Mr6)u (5-4)
mdash = Fr - Qru
where S(QU Mr0) is the column matrix given by
S(QuMr6) = [s(Qu(l)Mr0) 5 (Q u ( 2 ) M r | 0 ) S(QU nu) M r | 0 ) ] (55)
and Qu-i) is the i ^ column of Qu Extract ing only the controlled equations from
Eq(54) ie the attitude equations one has
Ire = Frc ~~ Qrcu
(56)
mdash vrci
where vrc is the new control input required to regulate the decoupled linear system
A t this point a simple P D controller can be applied ie
Vrc = qrcd +KvQrcd-Qrc) + Kp(Qrcd~ Qrc)- (57)
Eq(57) represents the secondary controller with Eq(54) as the primary controller
Kp and Kv are the proportional and derivative gain matrices with qrcdi Qrcd
an(^ qrcd
95
as the desired acceleration velocity and position vector for the attitude angles of each
controlled body respectively Solving for u from Eq(56) leads to
u Qrc Frc ~ vrcj bull (58)
5 13 Simulation results
The F L T controller is implemented on the three-body STSS tethered system
The choice of proportional and derivative matrix gains is based on a desired settling
time of ts mdash 05 orbit and a damping factor of = 07 for both the pitch and roll
actuators in each body Given O and ts it can be shown that
-In ) 0 5 V T ^ r (59)
and hence kn bull mdash10
V (510)
where kp and kVj are the diagonal elements of the gain matrices Kp and Kv in
Eq(57) respectively[59]
Note as each body-fixed frame is referred directly to the inertial frame FQ
the nominal pitch equilibrium angle is zero only when measured from the L V L H frame
(OJJ mdash 0) However it is 6 when referred to FQ Hence the desired position vector
qrC(l is set equal to 0(t) ie the time-varying orbital angle for the pitch motion
and zero for the roll and yaw rotations such that
qrcd 37Vxl (511)
96
The desired velocity and acceleration are then given as
e U 3 N x l (512)
E K 3 7 V x l (513)
When the system is in a circular orbit qrc^ = 0
Figure 5-1 presents the controlled response of the STSS system defined in
Chapter 4 in the stationkeeping phase with offset d2 = 111T m and If = 20
km A s mentioned earlier the F L T controller is based on the nonlinear rigid model
Note the pitch and roll angles of the rigid bodies as well as of the tether are now
attenuated in less than 1 orbit (Figure 5-la) Consequently the longitudinal vibration
response of the tether is free of coupling arising from the librational modes of the
tether This leaves only the vibration modes which are eventually damped through
structural damping The pitch and roll dynamics of the platform require relatively
higher control moments Ma and Mpi respectively (Figure 5-lb) since they have to
overcome the extra moments generated by the offset of the tether attachment point
Furthermore the platform demands an additional moment to maintain its orientation
at zero pitch and roll angle since it is not the nominal equilibrium position The nonshy
zero static equilibrium of the platform arises due to its products of inertia On the
other hand the tether pitch and roll dynamics require only a small ini t ia l control
force of about plusmn 1 N which eventually diminishes to zero once the system is
97
and
o 0
Qrcj mdash
o o
0 0
Oiit) 0 0
1
N
N
1deg I o
STSS configuration 3-Body a1(0)=a2(0)=at(0)=2lt
P1(0)=p2(0)=pt(0)=1 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=dz(0)=1m Controlled Stationkeeping lt=20km
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
05 10 15 Time (Orbits)
Sat 2 Pitch and Roll Angles
20
CO CD
T J
cdT
cT
- mdash mdash 1 1 1 1 1 I - 1
Pitch -
A Roll -
A -
1 bull bull
-
1 Time (Orbits)
Tether Y Tranverse Vibration
1 2 Time (Orbits)
Tether Z Tranverse Vibration
gt to N
to
Time (Orbits) Time (Orbits) Longitudinal Vibration
145h
(a) Time (Orbits)
Figure 5-1 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local horizontal local vertical and orbit normal (a) attitude and vibration response
98
1
STSS configuration 3-Body a1(0)=ct2(0)=cct(0)=20
P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=d2(0)=1m Controlled Stationkeeping lt=20km
Sat 1 Pitch Control Moment Sat 1 Roll Control Moment
Figure 5-1 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local horizontal local vertical and orbit normal (b) control actuator time histories
99
stabilized Similarly for the end-body a very small control torque is required to
attenuate the pitch and roll response
If the complete nonlinear flexible dynamics model is used in the feedback loop
the response performance was found to be virtually identical with minor differences
as shown in Figure 5-2 For example now the pitch and roll response of the platform
is exactly damped to zero with no steady-state error as against the minor almost
imperceptible deviation for the case of the controller basedon the rigid model (Figure
5-la) In addition the rigid controller introduces additional damping in the transverse
mode of vibration where none is present when the full flexible controller is used
This is due to a high frequency component in the at motion that slowly decays the
transverse motion through coupling
As expected the full nonlinear flexible controller which now accounts the
elastic degrees of freedom in the model introduces larger fluctuations in the control
requirement for each actuator except at the subsatellite which is not coupled to the
tether since f c m 3 = 0 The high frequency variations in the pitch and roll control
moments at the platform are due to longitudinal oscillations of the tether and the
associated changes in the tension Despite neglecting the flexible terms the overall
controlled performance of the system remains quite good Hence the feedback of the
flexible motion is not considered in subsequent analysis
When the tether offset at the platform is restricted to only 1 m along the x
direction a similar response is obtained for the system However in this case the
steady-state error in the platforms rigid body motion is much smaller (Figure 5-3a)
In addition from Figure 5-3(b) the platforms control requirement is significantly
reduced
Figure 5-4 presents the controlled response of the STSS deploying a tether
100
STSS configuration 3-Body a1(0)=a2(0)=oct(0)=2deg p1(0)=P2(0)=p(0)=1deg 8x(0)=14m5y(0)=82(0)=1m dx(0)=dy(0)=dz(0)=1m Controlled Stationkeeping l=20km
Tether Pitch and Roll Angles
c n
T J
cdT
Sat 1 Pitch and Roll Angles 1 r
00 05 10 15 Time (Orbits)
Sat 2 Pitch and Roll Angles
c n CD
cdT a
o T J
CQ
- - I 1 1 1 I 1 1 1 1
P i t c h
R o l l
bull
1 -
0 1 2 Time (Orbits)
Tether Y Tranverse Vibration
0 1 2 Time (Orbits)
Tether Z Tranverse Vibration
E gt
10 N
CO
Time (Orbits) Time (Orbits) Longitudinal Vibration
145h
(a) Time (Orbits)
Figure 5-2 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear flexible FLT controller with offset along the local horizontal local vertical and orbit normal (a) attitude and vibration response
101
STSS configuration 3-Body a1(0)=a2(0)=cxt(0)=2deg 31(0)=p2(0)=p(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=d2(0)=1m Controlled Stationkeeping lt=20km
r i i i t i mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash mdash J 0 1 2 0 1 2
Time (Orbits) Time (Orbits) Tether Pitch Control Thrust Tether Roll Control Thruster
(b) Time (Orbits) Time (Orbits)
Figure 5-2 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear flexible FLT controller with offset along the local horizontal local vertical and orbit normal (b) control actuator time histories
102
STSS configuration 3-Body a1(0)=o2(0)=a(0)=2deg P1(0)=p2(0)=(3t(0)=1deg 8x(0)=14m5y(0)=6z(0)=1m dx(0)=1mdy(0)=dz(0)=0 Controlled Stationkeeping lt=20km
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
05 10 15 Time (Orbits)
Sat 2 Pitch and Roll Angles
CD
eg CQ
0 1 2 Time (Orbits)
Tether Y Tranverse Vibration
1 2 Time (Orbits)
Tether Z Tranverse Vibration
gt CO
Time (Orbits) Time (Orbits) Longitudinal Vibration
145h
(a) Time (Orbits)
Figure 5-3 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local vertical (a) attitude and vibration response
103
STSS configuration 3-Body a1(0)=o2(0)=al(0)=2o
P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=1mdy(0)=dz(0)=0 Controlled Stationkeeping lt=20km
Sat 1 Pitch Control Moment 11 1 1 r
Sat 1 Roll Control Moment
0 1 2 Time (Orbits)
Tether Pitch Control Thrust
0 1 2 Time (Orbits)
Sat 2 Pitch Control Moment
0 1 2 Time (Orbits)
Tether Roll Control Thruster
-07 o 1
Time (Orbits Sat 2 Roll Control Moment
1E-5
0E0F
Time (Orbits)
Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local vertical (b) control actuator time histories
104
STSS configuration 3-Body a1(0)=a2(0)=a(0)=2deg Pl(0)=p2(0)=pt(0)=1deg 6X(0)=1304 x 103m 8y(0)=82(0)=0 dx(0)=1mdy(0)=d2(0)=0 Controlled Deployment
l=02km to 20km in 35 Orbits
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
o 1 2 Time (Orbits)
Tether Y Tranverse Vibration
0 1 2 Time (Orbits)
Sat 2 Pitch and Roll Angles
CD CD
CM CO
Time (Orbits) Tether Z Tranverse Vibration
to
150
100
1 2 3 0 1 2
Time (Orbits) Time (Orbits) Longitudinal Vibration
to
(a)
50
2 3 Time (Orbits)
Figure 5 - 4 Deployment dynamics of the three-body STSS using the nonlinear rigid FLT controller with offset along the local vertical (a) attitude and vibration response
105
STSS configuration 3-Body a1(0)=oc2(0)=al(0)=20
P1(0)=p2(0)=pt(0)=1deg 8X(0)=1304 x 103m 5y(0)=52(0)=0 dx(0)=1mdy(0)=dz(0)=0 Controlled Deployment
l=02km to 20km in 35 Orbits
Tether Length Profile i 1
2 0 h
E 2L
Sat 1 Pitch Control Moment t mdash r mdash i mdash i mdash | mdash i mdash i mdash i mdash i mdash | mdash r mdash i mdash i mdash i mdash | mdash r mdash
0 5
1 2 3
Time (Orbits) Sat 1 Roll Control Moment
i i t | t mdash t mdash i mdash [ i i mdash r -
1 2
Time (Orbits) Tether Pitch Control Thrust
1 2 3 Time (Orbits)
Tether Roll Control Thruster
2 E - 5
1 2 3
Time (Orbits) Sat 2 Pitch Control Moment
1 2
Time (Orbits) Sat 2 Roll Control Moment
1 E - 5
E
O E O
Time (Orbits)
Deployment dynamics of the three-body STSS using the nonlinear rigid FLT controller with offset along the local vertical (b) control actuator time histories
106
from 02 km to 20 km in 35 orbits with a i m offset along the tether length A s before
the systems attitude motion is well regulated by the controller However the control
cost increases significantly for the pitch motion of the platform and tether due to the
Coriolis force induced by the deployment maneuver (Figure 5-4b) Initially there is a
sinusoidal increase in the pitch control requirement for both the tether and platform
as the tether accelerates to its constant velocity VQ Then the control requirement
remains constant for the platform at about 60 N m as opposed to the tether where the
thrust demand increases linearly Finally when the tether decelerates the actuators
control input reduces sinusoidally back to zero The rest of the control inputs remain
essentially the same as those in the stationkeeping case In fact the tether pitch
control moment is significantly less since the tether is short during the ini t ia l stages
of control However the inplane thruster requirement Tat acts as a disturbance and
causes 6y to nearly double from the uncontrolled value (Figure 4-8a) O n the other
hand the tether has no out-of-plane deflection
Final ly the effectiveness of the F L T controller during the crit ical maneuver of
retrieval from 20 km to 02 k m in 35 orbits is assessed in Figure 5-5 A s in the earlier
cases the system response in pitch and roll is acceptable however the controller is
unable to suppress the high frequency elastic oscillations induced in the tether by the
retrieval O f course this is expected as there is no active control of the elastic degrees
of freedom However the control of the tether vibrations is discussed in detail in
Section 52 The pitch control requirement follows a similar trend in magnitude as
in the case of deployment with only minor differences due to the addition of offset in
the out-of-plane direction ( d 2 = 1 0 1 ^ m) This is also responsible for the higher
roll moment requirement for the platform control (Figure 5-5b)
107
STSS configuration 3-Body a1(0)=X2(0)=a(0)=2deg P1(0)=p2(0)=pt(0)=1deg 5x(0)=14m8y(0)=52(0)=0 dx(0)=1mdy(0)=dz(0)=0 Controlled Retrieval
l=20km to 02km in 35 Orbits
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
0 1 2 Time (Orbits)
Sat 2 Pitch and Roll Angles
CO CD
cdT
CO CD
T J CM
CO
1 2 Time (Orbits)
Tether Y Tranverse Vibration
Time (Orbits) Tether Z Tranverse Vibration
150
100
1 2 3 0 1 2 Time (Orbits) Time (Orbits)
Longitudinal Vibration
CO
(a)
x 50
2 3 Time (Orbits)
Figure 5 -5 Retrieval dynamics of the three-body STSS using the non-linear rigid FLT controller with offset along the local vertical and orbit normal (a) attitude and vibration response
108
STSS configuration 3-Body a1(0)=ct2(0)=at(0)=2deg p1(0)=p2(0)=(3(0)=1deg 5x(0)=14m5y(0)=82(0)=0 dx(0)=1mdy(0)=dz(0)=0 Controlled Retrieval
l=20km to 02km in 35 Orbits
Tether Length Profile
E
1 2 Time (Orbits)
Tether Pitch Control Thrust
2E-5
1 2 Time (Orbits)
Sat 2 Pitch Control Moment
OEOft
1 2 3 Time (Orbits)
Sat 1 Roll Control Moment
1 2 3 Time (Orbits)
Tether Roll Control Thruster
1 2 Time (Orbits)
Sat 2 Roll Control Moment
1E-5 E z
2
0E0
(b) Time (Orbits) Time (Orbits)
Figure 5-5 Retrieval dynamics of the three-body STSS using the non-linear rigid FLT controller with offset along the local vertical and orbit normal (b) control actuator time histories
109
52 Control of Tethers Elastic Vibrations
521 Prel iminary remarks
The requirement of a precisely controlled micro-gravity environment as well
as the accurate release of satellites into their final orbit demands additional control
of the tethers vibratory motion in addition to its attitude regulation To that end
an active vibration suppression strategy is designed and implemented in this section
The strategy adopted is based on offset control ie time dependent variation
of the tethers attachment point at the platform (satellite 1) A l l the three degrees
of freedom of the offset motion are used to control both the longitudinal as well as
the inplane and out-of-plane transverse modes of vibration In practice the offset
controller can be implemented through the motion of a dedicated manipulator or a
robotic arm supporting the tether which in turn is supported by the platform The
focus here is on the control of elastic deformations during stationkeeping (ie fully
deployed tether fixed length) as it represents the phase when the mission objectives
are carried out
This section begins with the linearization of the systems equations of motion
for the reduced three body stationkeeping case This is followed by the design of
the optimal control algorithm Linear Quadratic Gaussian-Loop Transfer Recovery
( L Q G L T R ) based on the reduced model and its implementation on the full nonlinear
model Final ly the systems response in the presence of offset control is presented
which tends to substantiate its effectiveness
522 System linearization and state-space realization
The design of the offset controller begins with the linearization of the equations
of motion about the systems equilibrium position Linearization of extremely lengthy
(even in matrix form) highly nonlinear nonautonomous and coupled equations of
110
motion presents a challenging problem This is further complicated by the fact the
pitch angle ot is not referred to the L V L H frame Hence the equilibrium pitch angle
is not a constant but is equal to the instantaneous orbital angle 9(t) Two methods
are available to resolve this problem
One may use the non-stationary equations of motion in their present form and
derive a controller based on the Linear T ime Varying ( L T V ) system Alternatively
one can design a Linear T ime Invariant (LTI) controller based on a new set of reduced
governing equations representing the motion of the three body tethered system A l shy
though several studies pertaining to L T V systems have been reported[5660] the latter
approach is chosen because of its simplicity Furthermore the L T I controller design
is carried out completely off-line and thus the procedure is computationally more
efficient
The reduced model is derived using the Lagrangian approach with the genershy
alized coordinates given by
where lty and are the pitch and roll angles of link i relative to the L V L H frame
respectively 5X 8y and 5Z are the generalized coordinates associated with the flexible
tether deformations in the longitudinal inplane and out-of-plane transverse direcshy
tions respectively Only the first mode of vibration is considered in the analysis
The nonlinear nonautonomous and coupled equations of motion for the tethshy
ered system can now be written as
Qred = [ltXlPlltX2P2fixSy5za302gt] (514)
M redQred + fr mdash 0gt (515)
where Mred and frec[ are the reduced mass matrix and forcing term of the system
111
respectively They are functions of qred and qred in addition to the time varying offset
position d2 and its time derivatives d2 and d2 A detailed derivation of Eq(515)
is given in Appendix II A n additional consequence of referring the pitch motion to
a local frame is that the new reduced equations are now independent of the orbital
angle 9 and under the further assumption of the system negotiating a circular orbit
91 remains constant
The nonlinear reduced equations can now be linearized about their static equishy
l ibr ium position For all the generalized coordinates the equilibrium position is zero
wi th the exception of 5X which has a non-zero equilibrium 5XQ The offset position
d2) is linearized about its ini t ial position d2^ Linearizing Eq(515) and recasting
into matr ix form gives
Msqred + CsQred + KsQred + Md^2 + Cdd2 + Kdd2 + fs = 0 (516)
where Ms Cs Ks Md Cd Kd and s are constant matrices Mul t ip ly ing throughout
by Ms
1 gives
qr = -Ms 1Csqr - Ms
lKsqr
- Ms~lCdd22 - Ms-Kdd2 - Ms~Mdd2 - Mg~1 fs r-1 (517)
Defining ud = d2 and v = qTd^T Eq(517) can be rewritten as
Pu t t ing
v =
+
-M~lCs -M~Cd -1
0
-Mg~lMd
Id
0 v + -M~lKs -M~lKd
0 0
-M-lfs
0
MCv + MKv + Mlud +Fs
- 4 ) - ( gt the L T I equations of motion can be recast into the state-space form as
(518)
(519)
x =Ax + Bud + Fd (520)
112
where
A = MC MK Id12 o
24x24
and
Let
B = ^ U 5 R 2 4 X 3 -
Fd = I FJ ) e f t 2 4 1
mdash _ i _ -5
(521)
(522)
(523)
(524)
where x is the perturbation vector from the constant equilibrium state vector xeq
Substituting Eq(524) into Eq(520) gives the linear perturbation state equation
modelling the reduced tethered system as
bullJ x mdash x mdash Ax + Bud + (Fd + Axeq) (525)
It can be shown that for ideal control[60] Fd + Axeq = 0 thus giving the familiar
state-space equation
S=Ax + Bud (526)
The selection of the output vector completes the state space realization of
the system The output vector consists of the longitudinal deformation from the
equil ibrium position SXQ of the tether at xt = h the slope of the tether due to the
transverse deformation at xt = 0 and the offset position d2 from its in i t ia l position
d2Q Thus the output vector y is given by
K(o)sx
V
dx ~ dXQ dy ~ dyQ
d z - d z0 J
Jy
dx da XQ vo
d z - dZQ ) dy dyQ
(527)
113
where lt ( 0 ) = ddxi[$vxi)]Xi=o and lt ( 0 ) = ddxi[4gtw(xi)]Xi=0
523 Linear Quadratic Gaussian control with Loop Transfer Recovery
W i t h the linear state space model defined (Eqs526527) the design of the
controller can commence The algorithm chosen is the Linear Quadratic Gaussian
( L Q G ) estimator based optimal controller[6061] The L Q G is a widely used optimal
controller with its theoretical background well developped by many authors over the
last 25 years It involves of the design of the Kalman-Bucy Fi l ter ( K B F ) which
provides an estimate of the states x and a Linear Quadratic Regulator ( L Q R ) which
is separately designed assuming all the states x are known Both the L Q R and K B F
designs independently have good robustness properties ie retain good performance
when disturbances due to model uncertainty are included However the combined
L Q R and K B F designs ie the L Q G design may have poor stability margins in the
presence of model uncertainties This l imitat ion has led to the development of an L Q G
design procedure that improves the performance of the compensator by recovering
the full state feedback robustness properties at the plant input or output (Figure 5-6)
This procedure is known as the Linear Quadratic Guassian-Loop Transfer Recovery
( L Q G L T R ) control algorithm A detailed development of its theory is given in
Ref[61] and hence is not repeated here for conciseness
The design of the L Q G L T R controller involves the repeated solution of comshy
plicated matrix equations too tedious to be executed by hand Fortunenately the enshy
tire algorithm is available in the Robust Control Toolbox[62] of the popular software
package M A T L A B The input matrices required for the function are the following
(i) state space matrices ABC and D (D = 0 for this system)
(ii) state and measurement noise covariance matrices E and 0 respectively
(iii) state and input weighting matrix Q and R respectively
114
u
MODEL UNCERTAINTY
SYSTEM PLANT
y
u LQR CONTROLLER
X KBF ESTIMATOR
u
y
LQG COMPENSATOR
PLANT
X = f v ( X t )
COMPENSATOR
x = A k x - B k y
ud
= C k f
y
Figure 5-6 Block diagram for the LQGLTR estimator based compensator
115
A s mentionned earlier the main objective of this offset controller is to regu-mdash
late the tether vibration described by the 5 equations However from the response
of the uncontrolled system presented in Chapter 4 it is clear that there is a large
difference between the magnitude of the librational and vibrational frequencies This
separation of frequencies allow for the separate design of the vibration and attitude mdash mdash
controllers Thus only the flexible subsystem composed of the S and d equations is
required in the offset controller design Similarly there is also a wide separation of
frequencies between the longitudinal and transverse modes of vibrations permitt ing mdash
the decoupling of the flexible subsystem into a set of longitudinal (Sx and dx) and
transverse (5y Sz dy and dz) subsystems The appended offset system d must also
be included since it acts as the control actuator However it is important to note
that the inplane and out-of-plane transverse modes can not be decoupled because
their oscillation frequencies are of the same order The two flexible subsystems are
summarized below
(i) Longitudinal Vibra t ion Subsystem
This is defined by u mdash Auxu + Buudu
(528)
where xu = SxdX25xdX2T u d u = dX2 and from Eq(527)
0 0 $u(lt) 0 0 0 0 1
0 0 1 0 0 0 0 1
(529)
Au and Bu are the rows and columns of A and B respectively and correspond to the
components of xu
The L Q R state weighting matrix Qu is taken as
Qu =
bull1 0 0 o -0 1 0 0 0 0 1 0
0 0 0 10
(530)
116
and the input weighting matrix Ru = 1^ The state noise covariance matr ix is
selected as 1 0 0 0 0 1 0 0 0 0 4 0
LO 0 0 1
while the measurement noise covariance matrix is taken as
(531)
i o 0 15
(532)
Given the above mentionned matrices the design of the L Q G L T R compenshy
sator is computed using M A T L A B with the following 2 commands
(i) kfu = l q r c ( ^ ( i C bdquo d i a g m x ( E u e u ) )
(ii) [AkugtBkugtCkugtDku =^ryAuBuCuDukfuQuRur)
where r is a scaler
The first command is used to compute the Ka lman filter gain matrix kfu Once
kfu is known the second command is invoked returning the state space representation
of the L Q G L T R dynamic compensator to
(533) xu mdash Akuxu Bkuyu
where xu is the state estimate vector oixu The function l t ry performs loop transfer
recovery at the system output ie the return ratio at the output approaches that
of the K B F loop given by Cu(sld - Au)Kfu[61 This is achieved by choosing a
sufficently large scaler value r in l t ry such that the singular values of the return
ratio approach those of the target design For the longitudinal controller design
r mdash 5 x 10 5 Figure 5-7(a) compares the singular values of the recovered compensator
design and the non-recovered L Q G compensator design with respect to the target
Unfortunetaly perfect recovery is not possible especially at higher frequencies
117
100
co
gt CO
-50
-100 10
Longitudinal Compensator Singular Values i I I I N I I mdash i mdash i i i i n i i mdash i mdash i i i i inimdash imdashi i i i i i i i mdash i mdash i i M i I I
Target r = 5x10 5
r = 0 (LQG)
i i i i I I i ii i i i 1 1 1 1 i i i i i 1 1 1 1 ii i I I i i i i 11111 -3 10 -2
(a)
101 10deg Frequency (rads)
101 10
Transverse Compensator Singular Values
100
50
X3
gt 0
CO
-50
bull100 10
-imdashi i 111II i 111 ni 1mdashi I I I I I I I 1mdashi I I I N I I 1mdashi i M 111
Target r = 50 r = 0 (LQG)
m i i i i i i n i i i i I I I I I I I 1mdashi I I I I I I I 1mdashi i 111 n -3 10 -2
(b)
101 10deg Frequency (rads)
101 10
Figure 5-7 Singular values for the LQG and LQGLTR compensator compared to target return ratio (a) longitudinal design (b) transverse design
118
because the system is non-minimal ie it has transmission zeros with positive real
parts[63]
(ii) Transverse Vibra t ion Subsystem
Here Xy mdash AyXy + ByU(lv
Dv = Cvxv
with xv = 6y 8Z dV2 dZ2 Sy 6Z dV2 dZ2T] udv = dy2dZ2T and
(534)
Cy
0 0 0 0 (0) 0 0 0 0 0 0 0 0 0 0 0 0
0 0 0 0
0 0 0 0 0 0 0 0 0 0 0 0
0 0 0
h 0 0 0
0 0 0 lt ( 0 ) o 0
0 1 0 0 0 1
0 0 0
2 r
h 0 0
0 0 1 0 0 1
(535)
Av and Bv are the rows and columns of the A and B corresponding to the components
O f Xy
The state L Q R weighting matrix Qv is given as
Qv
10 0 0 0 0 0 0 0 0 10 0 0 0 0 0 0 0 0 10 4 0 0 0 0 0 0 0 0 10 4 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1
(536)
mdash
Ry =w The state noise covariance matrix is taken
1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 9 x 10 4 0
0 0 0 0 0 0 0 9 x 10 4
(537)
119
while the measurement noise covariance matrix is
1 0 0 0 0 1 0 0 0 0 5 x 10 4 0
LO 0 0 5 x 10 4
(538)
As in the development of the longitudinal controller the transverse offset conshy
troller can be designed using the same two commands (lqrc and l try) wi th the input
matrices corresponding to the transverse subsystem with r = 50 The result is the
transverse compensator system given by
(539)
where xv is the state estimates of xv The singular values of the target transfer
function is compared with those achieved by the L Q G and L Q G L T R in Figure 5-
7(b) It is apparent that the recovery is not as good as that for the longitudinal case
again due to the non-minimal system Moreover the low value of r indicates that
only a lit t le recovery in the transverse subsystem is possible This suggests that the
robustness of the L Q G design is almost the maximum that can be achieved under
these conditions
Denoting f y = f j f j r xf = x^x^T and ud = udu ^ T the longishy
tudinal compensator can be combined to give
xf =
$d =
Aku
0
0 Akv
Cku o
Cu 0
Bku
0 B Xf
Xf = Ckxf
0
kv
Hf = Akxf - Bkyf
(540)
Vu Vv 0 Cv
Xf = CfXf
Defining a permutation matrix Pf such that Xj = PfX where X is the state vector
of the original nonlinear system given in Eq(32) the compensator and full nonlinear
120
system equations can be combined as
X
(541)
Ckxf
A block diagram representation of Eq(541) is shown in Figure 5-6
524 Simulation results
The dynamical response for the stationkeeping STSS with a tether length of 20
km and ini t ia l offset position of 1 m in each direction is presented in Figure 5-8 when
both the F L T attitude controller and the L Q G L T R offset controller are activated As
expected the offset controller is successful in quickly damping the elastic vibrations
in the longitudinal and transverse direction (Figure 5-8b) However from Figure 5-
8(a) it is clear that the presence of offset control requires a larger control moment
to regulate the attitude of the platform This is due to the additional torque created
by the larger moment arm around 2 m and 125 m in the inplane and out-of-plane
directions respectively introduced by the offset control In addition the control
moment for the platform is modulated by the tethers transverse vibration through
offset coupling However this does not significantly affect the l ibrational motion of
the tether whose thruster force remains relatively unchanged from the uncontrolled
vibration case
Final ly it can be concluded that the tether elastic vibration suppression
though offset control in conjunction with thruster and momentum-wheel attitude
control presents a viable strategy for regulating the dynamics tethered satellite sysshy
tems
121
STSS configuration 3-Body a1(0)=a2(0)=at(0)=2deg P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m6y(0)=52(0)=1m dx(0)=dy(0)=dz(0)=1m Stationkeeping I = 20km LQGLTR Vibration Control
Sat 1 Pitch and Roll Angles Tether Pitch and Roll Angles mdash 1 1 1 1 1 1 I 1 1 lt 1 ~ ^ 1 ^ l _
Time (Orbits) Time (Orbits) Sat 1 Roll Control Moment Tether Roll Control Thrust
(a) Time (Orbits) Time (Orbits)
Figure 5-8 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT attitude controller and LQGLTR offset vibration controller (a) attitude and libration controller reshysponse
122
STSS configuration 3-Body a1(0)=a2(0)=a(0)=2deg p1(0)=(32(0)=Q(0)=1deg ox(0)=14mSy(0)=5z(0)=1m dx(0)=dy(0)=dz(0)=1m Stationkeeping lt = 20km LQGLTR Vibration Control
Tether X Offset Position
Time (Orbits) Tether Y Tranverse Vibration 20r
Time (Orbits) Tether Y Offset Position
t o
1 2 Time (Orbits)
Tether Z Tranverse Vibration Time (Orbits)
Tether Z Offset Position
(b)
Figure 5-8
Time (Orbits) Time (Orbits)
Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT attitude controller and LQGLTR offset vibration controller (b) vibration and offset response
123
6 C O N C L U D I N G R E M A R K S
61 Summary of Results
The thesis has developed a rather general dynamics formulation for a multi-
body tethered system undergoing three-dimensional motion The system is composed
of multiple rigid bodies connected in a chain configuration by long flexible tethers
The tethers which are free to undergo libration as well as elastic vibrations in three
dimensions are also capable of deployment retrieval and constant length stationkeepshy
ing modes of operation Two types of actuators are located on the rigid satellites
thrusters and momentum-wheels
The governing equations of motion are developed using a new Order(N) apshy
proach that factorizes the mass matrix of the system such that it can be inverted
efficiently The derivation of the differential equations is generalized to account for
an arbitrary number of rigid bodies The equations were then coded in FORTRAN
for their numerical integration with the aid of a symbolic manipulation package that
algebraically evaluated the integrals involving the modes shapes functions used to dis-
cretize the flexible tether motion The simulation program was then used to assess the
uncontrolled dynamical behaviour of the system under the influence of several system
parameters including offset at the tether attachment point stationkeeping deployshy
ment and retrieval of the tether The study covered the three-body and five-body
geometries recently flown OEDIPUS system and the proposed BICEPS configurashy
tion It represents innovation at every phase a general three dimensional formulation
for multibody tethered systems an order-N algorithm for efficient computation and
application to systems of contemporary interest
Two types of controllers one for the attitude motion and the other for the
124
flexible vibratory motion are developed using the thrusters and momentum-wheels for
the former and the variable offset position for the latter These controllers are used to
regulate the motion of the system under various disturbances The attitude controller
is developed using the nonlinear Feedback Linearization Technique ( F L T ) and is based
on the nonlinear but rigid model of the system O n the other hand the separation
of the longitudinal and transverse frequencies from those of the attitude response
allows for the development of a linear optimal offset controller using the robust Linear
Quadratic Gaussian-Loop Transfer Recovery ( L Q G L T R ) method The effectiveness
of the two controllers is assessed through their subsequent application to the original
nonlinear flexible model
More important original contributions of the thesis which have not been reshy
ported in the literature include the following
(i) the model accounts for the motion of a multibody chain-type system undershy
going librational and in the case of tethers elastic vibrational motion in all
the three directions
(ii) the dynamics formulation is based on a recursive O(N) Lagrangian algorithm
that efficiently computes the systems generalized acceleration vector
(iii) the development of an attitude controller based on the Feedback Linearizashy
tion Technique for a multibody system using thrusters and momentum-wheels
located on each rigid body
(iv) the design of a three degree of freedom offset controller to regulate elastic v i shy
brations of the tether using the Linear Quadratic Gaussian and Loop Transfer
Recovery method ( L Q G L T R )
(v) substantiation of the formulation and control strategies through the applicashy
tion to a wide variety of systems thus demonstrating its versatility
125
The emphasis throughout has been on the development of a methodology to
study a large class of tethered systems efficiently It was not intended to compile
an extensive amount of data concerning the dynamical behaviour through a planned
variation of system parameters O f course a designer can easily employ the user-
friendly program to acquire such information Rather the objective was to establish
trends based on the parameters which are likely to have more significant effect on
the system dynamics both uncontrolled as well as controlled Based on the results
obtained the following general remarks can be made
(i) The presence of the platform products of inertia modulates the attitude reshy
sponse of the platform and gives rise to non-zero equilibrium pitch and roll
angles
(ii) Offset of the tether attachment point significantly affects the equil ibrium orishy
entation of the system as well as its dynamics through coupling W i t h a relshy
atively small subsatellite the tether dynamics dominate the system response
with high frequency elastic vibrations modulating the librational motion On
the other hand the effect of the platform dynamics on the tether response is
negligible As can be expected the platform dynamics is significantly affected
by the offset along the local horizontal and the orbit normal
(iii) For a three-body system deployment of the tether can destabilize the platform
in pitch as in the case of nonzero offset However the roll motion remains
undisturbed by the Coriolis force Moreover deployment can also render the
tether slack if it proceeds beyond a critical speed
(iv) Uncontrolled retrieval of the tether is always unstable in pitch It also leads
to high frequency elastic vibrations in the tether
(v) The five-body tethered system exhibits dynamical characteristics similar to
those observed for the three-body case As can be expected now there are
additional coupling effects due to two extra bodies a rigid subsatellite and a
126
flexible tether
(vi) Cartwheeling motion of the B I C E P S configuration can also be used to adshy
vantage in the deployment of the tether However exceeding a crit ical ini t ia l
cartwheeling rate can result in a large tension in the tether which eventually
causes the satellite to bounce back rendering the tether slack
(vii) Spinning the tether about its nominal length as in the case of O E D I P U S
introduces high frequency transverse vibrations in the tether which in turn
affect the dynamics of the satellites
(viii) The F L T based controller is quite successful in regulating the attitude moshy
tion of both the tether and rigid satellites in a short period of time during
stationkeeping deployment as well as retrieval phases
(ix) The offset controller is successful in suppressing the elastic vibrations of the
tether wi thin the allowable l imits of the offset mechanism ( plusmn 2 0 m)
62 Recommendations for Future Study
A s can be expected any scientific inquiry is merely a prelude to further efforts
and insight Several possible avenues exist for further investigation in this field A
few related to the study presented here are indicated below
(i) inclusion of environmental effects particularly the free molecular reaction
forces as well as interaction with Earths magnetic field and solar radiation
(ii) addition of flexible booms attached to the rigid satellites providing a mechashy
nism for energy dissipation
(iii) validation of the new three dimensional offset control strategies using ground
based experiments
(iv) animation of the simulation results to allow visual interpretation of the system
dynamics
127
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[38] Ruying F and Bainum P M Dynamics and Control of a Space Platform with a Tethered Subsatellite Journal of Guidance Control and Dynamics V o l 11 No 4 July-August 1988 pp 377-381
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132
A P P E N D I X I TENSOR R E P R E S E N T A T I O N OF T H E
EQUATIONS OF M O T I O N
11 Prel iminary Remarks
The derivation of the equations of motion for the multibody tethered system
involves the product of several matrices as well as the derivative of matrices and
vectors with respect to other vectors In order to concisely code these relationships
in FORTRAN the matrix equations of motion are expressed in tensor notation
Tensor mathematics in general is an extremely powerful tool which can be
used to solve many complex problems in physics and engineering[6465] However
only a few preliminary results of Cartesian tensor analysis are required here They
are summarized below
12 Mathematical Background
The representation of matrices and vectors in tensor notation simply involves
the use of indices referring to the specific elements of the matrix entity For example
v = vk (k = lN) (Il)
A = Aij (i = lNj = lM)
where vk is the kth element of vector v and Aj is the element on the t h row and
3th column of matrix A It is clear from Eq(Il) that exactly one index is required to
completely define an entire vector whereas two independent indices are required for
matrices For this reason vectors and matrices are known as first-order and second-
order tensors respectively Scaler variables are known as zeroth-order tensors since
in this case an index is not required
133
Matrix operations are expressed in tensor notation similar to the way they are
programmed using a computer language such as FORTRAN or C They are expressed
in a summation notation known as Einstein notation For example the product of a
matrix with a vector is given as
w = Av (12)
or in tensor notation W i = ^2^2Aij(vkSjk)
4-1 ltL3gt j
where 8jk is the Kronecker delta Dropping the summation symbol the matrix-vector
product can be expressed compactly as
Wi = AijVj = Aimvm (14)
Here j is a dummy index since it appears twice in a single term and hence can
be replaced by any other index Note that since the resulting product is a vector
only one index is required namely i that appears exactly once on both sides of the
equation Similarly
E = aAB + bCD (15)
or Ej mdash aAimBmj + bCimDmj
~ aBmjAim I bDmjCim (16)
where A B C D and E are second-order tensors (matrices) and a and b are zeroth-
order tensors (scalers) In addition once an expression is in tensor form it can be
treated similar to a scaler and the terms can be rearranged
The transpose of a matrix is also easily described using tensors One simply
switches the position of the indices For example let w = ATv then in tensor-
notation
Wi = AjiVj (17)
134
The real power of tensor notation is in its unambiguous expression of terms
containing partial derivatives of scalers vectors or matrices with respect to other
vectors For example the Jacobian matrix of y = f(x) can be easily expressed in
tensor notation as
di-dx-j-1 ( L 8 )
mdash which is a second-order tensor as expected If f(x) = Ax where A is constant then
f - a )
according to the current definition If C = A(x)B(x) then the time derivative of C
is given by
C = Ax)Bx) + A(x)Bx) (110)
or in tensor notation
Cij = AimBmj + AimBmj (111)
= XhiAimfcBmj + AimBmj^k)
which is more clear than its equivalent vector form Note that A^^ and Bmjk are
third-order tensors that can be readily handled in F O R T R A N
13 Forcing Function
The forcing function of the equations of motion as given by Eq(260) is
- u 1-SdM- dqtdPe 9 f i m F(q qt) = Mq - -q mdashq + mdashmdash - Qd (260)
This expression can be converted into tensor form to facilitate its implementation
into F O R T R A N source code
Consider the first term Mq From Eq(289) the mass matrix M is given by
T M mdash Rv MtRv which in tensor form is
Mtj = RZiM^R^j (112)
135
Taking the time derivative of M gives
Mij = qsRv
ni^MtnrnRv
mj + Rv
niMtnTn^Rv
mj + Rv
niMtnmRv
mjs) (113)
The second term on the right hand side in Eq(260) is given by
-JPdMu 1 bdquo T x
2Q ~cWq = 2Qs s r k Q r ^ ^
Expanding Msrk leads to
M src = RnskMtnmRmr + RnsMtnmkRmr + RnsMtnmRmrk- (L15)
Finally the potential energy term in Eq(260) is also expanded into tensor
form to give dPe = dqt dPe
dq dq dqt = dq^QP^
and since ^ = Rp(q)q then
Now inserting Eqs(I13-I17) into Eq(260) and rearranging the tensor form of the
forcing term can how be stated as
Fk(q^q^)=(RP
sk + R P n J q n ) ^ - - Q d k
+ QsQr | (^Rlks ~ 2 M t n m R m r
+ (^RnkMtnms ~ 2~RnsMtnmgtk^ R m r
(^RnkRmrs ~ ~2RnsRmrk^j
where Fk(q q t) represents the kth component of F
(118)
136
A P P E N D I X II R E D U C E D EQUATIONS OF M O T I O N
II 1 Prel iminary Remarks
The reduced model used for the design of the vibration controller is represented
by two rigid end-bodies capable of three-dimensional attitude motion interconnected
with a fixed length tether The flexible tether is discretized using the assumed-
mode method with only the fundamental mode of vibration considered to represent
longitudinal as well as in-plane and out-of-plane transverse deformations In this
model tether structural damping is neglected Finally the system is restricted to a
nominal circular orbit
II 2 Derivation of the Lagrangian Equations of Mot ion
Derivation of the equations of motion for the reduced system follows a similar
procedure to that for the full case presented in Chapter 2 However in this model the
straightforward derivation of the energy expressions for the entire system is undershy
taken ignoring the Order(N) approach discussed previously Furthermore in order
to make the final governing equations stationary the attitude motion is now referred
to the local LVLH frame This is accomplished by simply defining the following new
attitude angles
ci = on + 6
Pi = Pi (HI)
7t = 0
where and are the pitch and roll angles respectively Note that spin is not
considered in the reduced model hence 7J = 0 Substituting Eq(IIl) into Eq(214)
137
gives the new expression for the rotation matrix as
T- CpiSa^ Cai+61 S 0 S a + e i
-S 0 c Pi
(II2)
W i t h the new rotation matrix defined the inertial velocity of the elemental
mass drrii o n the ^ link can now be expressed using Eq(215) as
(215)
Since the tether length is assumed to be fixed T j f j and Tj$jlt^ of Eq(215) are
identically zero Also defining
(II3) Vi = I Pi i
Rdm- f deg r the reduced system is given by
where
and
Rdm =Di + Pi(gi)m + T^iSi
Pi(9i)m= [Tai9i T0gi T^] fj
(II4)
(II5)
(II6)
D1 = D D l D S v
52 = 31 + P(d2)m +T[d2
D3 = D2 + l2 + dX2Pii)m + T2i6x
Note that $ j for i = 2 is defined in Section 212 whereas for i = 1 and 3 it is
the null matrix Since only one mode is used for each flexible degree of freedom
5i = [6Xi6yi6Zi]T for i = 2
The kinetic energy of the system can now be written as
i=l Z i = l Jmi RdmRdmdmi (UJ)
138
Expanding Kampi using Eq(II4)
KH = miDD + 2DiP
i j gidmi)fji + 2D T ltMtrade^
+tf J P^(9i)Pi9i)d^l + 2mT j PFmTi^dm^ (II8)
where the integrals are evaluated using the procedure similar to that described in
Chapter 2 The gravitational and strain energy expressions are given by Eq(247)
and Eq(251) respectively using the newly defined rotation matrix T in place of
TV
Substituting the kinetic and potential energy expressions in
d ( 8Ke d(Ke - Pe) i bull i Q ^ 0 (II9)
d t dqred d(ired
with
Qred = [ algt l gt a 2 2 A 2 A 2 lt ^ 2 Q 3 3 ] (5-14)
and d2 regarded as a time-varying parameter the nonlinear non-autonomous equashy
tions of motion for the reduced system can finally be expressed as
MredQred + fred = reg- (5-15)
139
ABSTRACT
The equations of motion for a multibody tethered satellite system in three dishy
mensional Keplerian orbit are derived The model considers a multi-satellite system
connected in series by flexible tethers Both tethers and subsatellites are free to unshy
dergo three dimensional attitude motion together with deployment and retrieval as
well as longitudinal and transverse vibration for the tether The elastic deformations
of the tethers are discretized using the assumed mode method The tether attachment
points to the subsatellites are kept arbitrary and time varying The model is also cashy
pable of simulating the response of the entire system spinning about an arbitrary
axis as in the case of OEDIPUS-AC which spins about the nominal tether length
or the proposed BICEPS mission where the system cartwheels about the orbit norshy
mal The governing equations of motion are derived using a non-recursive order(N)
Lagrangian procedure which significantly reduces the computational cost associated
with the inversion of the mass matrix an important consideration for multi-satellite
systems Also a symbolic integration and coding package is used to evaluate modal
integrals thus avoiding their costly on-line numerical evaluation
Next versatility of the formulation is illustrated through its application to
two different tethered satellite systems of contemporary interest Finally a thruster
and momentum-wheel based attitude controller is developed using the Feedback Linshy
earization Technique in conjunction with an offset (tether attachment point) control
strategy for the suppression of the tethers vibratory motion using the optimal Linshy
ear Quadratic Gaussian-Loop Transfer Recovery method Both the controllers are
successful in stabilizing the system over a range of mission profiles
ii
T A B L E OF C O N T E N T S
A B S T R A C T bull i i
T A B L E O F C O N T E N T S i i i
L I S T O F S Y M B O L S v i i
L I S T O F F I G U R E S x i
A C K N O W L E D G E M E N T xv
1 I N T R O D U C T I O N 1
11 Prel iminary Remarks 1
12 Brief Review of the Relevant Literature 7
121 Mul t ibody O(N) formulation 7
122 Issues of tether modelling 10
123 Att i tude and vibration control 11
13 Scope of the Investigation 12
2 F O R M U L A T I O N O F T H E P R O B L E M 14
21 Kinematics 14
211 Prel iminary definitions and the itfl position vector 14
212 Tether flexibility discretization 14
213 Rotat ion angles and transformations 17
214 Inertial velocity of the ith link 18
215 Cyl indr ica l orbital coordinates 19
216 Tether deployment and retrieval profile 21
i i i
22 Kinetics and System Energy 22
221 Kinet ic energy 22
222 Simplification for rigid links 23
223 Gravitat ional potential energy 25
224 Strain energy 26
225 Tether energy dissipation 27
23 O(N) Form of the Equations of Mot ion 28
231 Lagrange equations of motion 28
232 Generalized coordinates and position transformation 29
233 Velocity transformations 31
234 Cyl indrical coordinate modification 33
235 Mass matrix inversion 34
236 Specification of the offset position 35
24 Generalized Control Forces 36
241 Preliminary remarks 36
242 Generalized thruster forces 37
243 Generalized momentum gyro torques 39
3 C O M P U T E R I M P L E M E N T A T I O N 42
31 Preliminary Remarks 42
32 Numerical Implementation 43
321 Integration routine 43
322 Program structure 43
33 Verification of the Code 46
iv
331 Energy conservation 4 6
332 Comparison with available data 50
4 D Y N A M I C S I M U L A T I O N 53
41 Prel iminary Remarks 53
42 Parameter and Response Variable Definitions 53
43 Stationkeeping Profile 56
44 Tether Deployment 66
45 Tether Retrieval 71
46 Five-Body Tethered System 71
47 B I C E P S Configuration 80
48 O E D I P U S Spinning Configuration 88
5 A T T I T U D E A N D V I B R A T I O N C O N T R O L 93
51 Att i tude Control 93
511 Prel iminary remarks 93
512 Controller design using Feedback Linearization Technique 94
513 Simulation results 96
52 Control of Tethers Elastic Vibrations 110
521 Prel iminary remarks 110
522 System linearization and state-space realization 1 1 0
523 Linear Quadratic Gaussian control
with Loop Transfer Recovery 114
524 Simulation results 121
v
6 C O N C L U D I N G R E M A R K S 1 2 4
61 Summary of Results 124
62 Recommendations for Future Study 127
B I B L I O G R A P H Y 128
A P P E N D I C E S
I T E N S O R R E P R E S E N T A T I O N O F T H E E Q U A T I O N S O F M O T I O N 133
11 Prel iminary Remarks 133
12 Mathematical Background 133
13 Forcing Function 135
II R E D U C E D E Q U A T I O N S O F M O T I O N 137
II 1 Preliminary Remarks 137
II2 Derivation of the Lagrangian Equations of Mot ion 137
v i
LIST OF SYMBOLS
A j intermediate length over which the deploymentretrieval profile
for the itfl l ink is sinusoidal
A Lagrange multipliers
EUQU longitudinal state and measurement noise covariance matrices
respectively
matrix containing mode shape functions of the t h flexible link
pitch roll and yaw angles of the i 1 link
time-varying modal coordinate for the i 1 flexible link
link strain and stress respectively
damping factor of the t h attitude actuator
set of attitude angles (fjj = ci
structural damping coefficient for the t h l ink EjJEj
true anomaly
Earths gravitational constant GMejBe
density of the ith link
fundamental frequency of the link longitudinal tether vibrashy
tion
A tethers cross-sectional area
AfBfCjDf state-space representation of flexible subsystem
Dj inertial position vector of frame Fj
Dj magnitude of Dj mdash mdash
Df) transformation matrix relating Dj and Ds
D r i inplane radial distance of the first l ink
DSi Drv6i DZ1 T for i = 1 and DX Dy poundgt 2 T for 1
m
Pi
w 0 i
v i i
Dx- transformation matrix relating D j and Ds-
DXi Dyi Dz- Cartesian components of D
DZl out-of-plane position component of first link
E Youngs elastic modulus for the ith l ink
Ej^ contribution from structural damping to the augmented complex
modulus E
F systems conservative force vector
FQ inertial reference frame
Fi t h link body-fixed reference frame
n x n identity matrix
alternate form of inertia matrix for the i 1 l ink
K A Z $i(k + dXi+1)
Kei t h link kinetic energy
M(q t) systems coupled mass matrix
Mma Mma Mmy- control moments in the pitch roll and yaw directions respec-
tively for the t h link
Mr fr rigid mass matrix and force vector respectively
Mred fred- Qred mass matrix force vector and generalized coordinate vector for
the reduced model respectively
Mt block diagonal decoupled mass matrix
symmetric decoupled mass matrix
N total number of links
0N) order-N
Pi(9i) column matrix [T^giTpg^T^gi) mdash
Q nonconservative generalized force vector
Qu actuator coupling matrix or longitudinal L Q R state weighting
matrix
v i i i
Rdm- inertial position of the t h link mass element drrii
RP transformation matrix relating qt and tf
Ru longitudinal L Q R input weighting matrix
RvRn Rd transformation matrices relating qt and q
S(f Mdc) generalized acceleration vector of coupled system q for the full
nonlinear flexible system
th uiirv l u w u u u maniA
th
T j i l ink rotation matrix
TtaTto control thrust in the pitch and roll direction for the im l ink
respectively
rQi maximum deploymentretrieval velocity of the t h l ink
Vei l ink strain energy
Vgi link gravitational potential energy
Rayleigh dissipation function arising from structural damping
in the ith link
di position vector to the frame F from the frame F_i bull
dc desired offset acceleration vector (i 1 l ink offset position)
drrii infinitesimal mass element of the t h l ink
dx^dy^dZi Cartesian components of d along the local vertical local horishy
zontal and orbit normal directions respectively
F-Q mdash
gi r igid and flexible position vectors of drrii fj + ltEraquolt5
ijk unit vectors 1 0 0 r 010-^ and 00 l r respectively
li length of the t h l ink
raj mass of the t h link
nfj total number of flexible modes for the i 1 l ink nXi + nyi + nZi
nq total number of generalized coordinates per link nfj + 7
nqq systems total number of generalized coordinates Nnq
ix
number of flexible modes in the longitudinal inplane and out-
of-plane transverse directions respectively for the i ^ link
number of attitude control actuators
- $ T
set of generalized coordinates for the itfl l ink which accounts for
interactions with adjacent links
qv---QtNT
set of coordinates for the independent t h link (not connected to
adjacent links)
rigid position of dm in the frame Fj
position of centre of mass of the itfl l ink relative to the frame Fj
i + $DJi
desired settling time of the j 1 attitude actuator
actuator force vector for entire system
flexible deformation of the link along the Xj yi and Zj direcshy
tions respectively
control input for flexible subsystem
Cartesian components of fj
actual and estimated state of flexible subsystem respectively
output vector of flexible subsystem
LIST OF FIGURES
1-1 A schematic diagram of the space platform based N-body tethered
satellite system 2
1-2 Associated forces for the dumbbell satellite configuration 3
1- 3 Some applications of the tethered satellite systems (a) multiple
communication satellites at a fixed geostationary location
(b) retrieval maneuver (c) proposed B I C E P S configuration 6
2- 1 Vector components of the i 1 and (i- l ) 1 chain links 15
2-2 Vector components in cylindrical orbital coordinates 20
2-3 Inertial position of subsatellite thruster forces 38
2- 4 Coupled force representation of a momentum-wheel on a rigid body 40
3- 1 Flowchart showing the computer program structure 45
3-2 Kinet ic and potentialenergy transfer for the three-body platform based
tethered satellite system (a) variation of kinetic and potential energy
(b) percent change in total energy of system 48
3-3 Kinet ic and potential energy transfer for the five-body platform based
tethered satellite system (a) variation of kinetic and potential energy
(b) percent change in total energy of system 49
3-4 Simulation results for the platform based three-body tethered
system originally presented in Ref[43] 51
3- 5 Simulation results for the platform based three-body tethered
system obtained using the present computer program 52
4- 1 Schematic diagram showing the generalized coordinates used to
describe the system dynamics 55
4-2 Stationkeeping dynamics of the three-body STSS configuration
without offset (a) attitude response (b) vibration response 57
4-3 Stationkeeping dynamics of the three-body STSS configuration
xi
with offset along the local vertical
(a) attitude response (b) vibration response 60
4-4 Stationkeeping dynamics of the three-body STSS configuration
with offset along the local horizontal
(a) attitude response (b) vibration response 62
4-5 Stationkeeping dynamics of the three-body STSS configuration
with offset along the orbit normal
(a) attitude response (b) vibration response 64
4-6 Stationkeeping dynamics of the three-body STSS configuration
wi th offset along the local horizontal local vertical and orbit normal
(a) attitude response (b) vibration response 67
4-7 Deployment dynamics of the three-body STSS configuration
without offset
(a) attitude response (b) vibration response 69
4-8 Deployment dynamics of the three-body STSS configuration
with offset along the local vertical
(a) attitude response (b) vibration response 72
4-9 Retrieval dynamics of the three-body STSS configuration
with offset along the local vertical and orbit normal
(a) attitude response (b) vibration response 74
4-10 Schematic diagram of the five-body system used in the numerical example 77
4-11 Stationkeeping dynamics of the five-body STSS configuration
without offset
(a) attitude response (b) vibration response 78
4-12 Deployment dynamics of the five-body STSS configuration
wi th offset along the local vertical
(a) attitude response (b) vibration response 81
4-13 Stationkeeping dynamics of the three-body B I C E P S configuration
x i i
with offset along the local vertical
(a) attitude response (b) vibration response 84
4-14 Cartwheeling dynamics with deployment for the three-body B I C E P S
with offset along the local vertical
(a) attitude response (b) vibration response 86
4-15 Spin dynamics (7 = l deg s ) of the three-body O E D I P U S configuration
with offset along the local vertical
(a) attitude response (b) vibration response 89
4- 16 Spin dynamics (7 = 10deg s ) of the three-body O E D I P U S configuration
wi th offset along the local vertical
(a) attitude response (b) vibration response 91
5- 1 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear rigid F L T controller with offset along the
local horizontal local vertical and orbit normal
(a) attitude and vibration response (b) control actuator time histories 98
5-2 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear flexible F L T controller with offset along the
local horizontal local vertical and orbit normal
(a) attitude and vibration response (b) control actuator time histories 101
5-3 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear rigid F L T controller with offset along
the local vertical
(a) attitude and vibration response (b) control actuator time histories 103
5-4 Deployment dynamics of the three-body STSS using the nonlinear
rigid F L T controller with offset along the local vertical
(a) attitude and vibration response (b) control actuator time histories 105
5-5 Retrieval dynamics of the three-body STSS using the non-linear
rigid F L T controller with offset along the local vertical and orbit normal
(a) attitude and vibration response (b) control actuator time histories 108
x i i i
5-6 Block diagram for the L Q G L T R estimator based compensator 115
5-7 Singular values for the L Q G and L Q G L T R compensator compared to target
return ratio (a) longitudinal design (b) transverse design 118
5-8 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear rigid F L T attitude controller and L Q G L T R
offset vibration controller (a) attitude and libration controller response
(b) vibration and offset response 122
xiv
A C K N O W L E D G E M E N T
Firs t and foremost I would like to express the most genuine gratitude to
my supervisor Prof V i n o d J M o d i whose unending guidance and encouragement
proved most invaluable during the course of my studies
I am also deeply indebted to Dr Satyabrata Pradhan for his limitless patience
and technical advice from which this work would not have been possible I would also
like to extend a word of appreciation towards Prof A r u n K Mis ra ( M c G i l l University)
for ini t iat ing my interest in space dynamics and control
Also I would like to thank my collegues Dr Sandeep Munshi M r Mathieu
Caron and M r Yuan Chen as well as Dr Mae Seto M r Gary L i m and M r Mark
Chu for their words of advice and helpful suggestions that heavily contributed to all
aspects of my studies
Final ly I would like to thank all my family and friends here in Vancouver and
back home in Montreal whose unending support made the past two years the most
memorable of my life
The research project was supported by the Natural Sciences and Engineering
Research Council s P G S - A Scholarship held by the author as well as by N S E R C grant
A-2181 held by Prof V J M o d i
xv
1 INTRODUCTION
11 Prel iminary Remarks
W i t h the ever changing demands of the worlds population one often wonders
about the commitment to the space program On the other hand the importance
of worldwide communications global environmental monitoring as well as long-term
habitation in space have demanded more ambitious and versatile satellites systems
It is doubtful that the Russian scientist Tsiolkovsky[l] when he first proposed the
uti l izat ion of the Earths gravity-gradient environment in the last century would have
envisioned the current and proposed applications of tethered satellite systems
A tethered satellite system consists of two or more subsatellites connected
to each other by long thin cables or tethers which are in tension (Figure 1-1) The
subsatellite which can also include the shuttle and space station are in orbit together
The first proposed use of tethers in space was associated with the rescue of stranded
astronauts by throwing a buoy on a tether and reeling it to the rescue vehicle
Prel iminary studies of such systems led to the discovery of the inherent instability
during the tether retrieval[2] Nevertheless the bir th of tethered systems occured
during the Gemini X I and X I I missions[3] in 1966 when a short tether (10-30m)
was used to generate the first artificial gravity environment in space (000015g) by
cartwheeling (spinning) the system about its center of mass
The mission also demonstrated an important use of tethers for gravity gradient
stabilization The force analysis of a simplified model illustrates this point Consider
the dumbbell satellite configuration orbiting about the centre of Ear th as shown in
Figure 1-2 Satellite 1 and satellite 2 are connected to each other by a tether with
1
Space Station (Satellite 1 )
2
Satellite N
Figure 1-1 A schematic diagram of the space platform based N-body tethered satellite system
2
Figure 1-2 Associated forces for the dumbbell satellite configuration
3
the whole system free to librate about the systems centre of mass C M The two
major forces acting on each body are due to the centrifugal and gravitational effects
Since body 1 is further away from the earth its gravitational force Fg^y is smaller
than that of body 2 ie Fgi lt Fg^ However its centrifugal force FCl is greater
then the centrifugal force experienced by the second satellite (FCl gt FC2) Moreover
for body 1 Fg^ lt Fc^ in contrast to Fg2 gt FC2 for body 2 Thus resolving the force
components along the tether there is an evident resultant tension force Ff present
in the tether Similarly adding the normal components of the two forces vectorially
results in a force FR which restores the system to its stable equil ibrium along the
local vertical These tension and gravity-gradient restoring forces are the two most
important features of tethered systems
Several milestone missions have flown in the recent past The U S A J a p a n
project T P E (Tethered Payload Experiment) [4] was one of the first to be launched
using a sounding rocket to conduct environmental studies The results of the T P E
provided support for the N A S A I t a l y shuttle based tethered satellite system referred
to as the TSS-1 mission[5] This experiment aimed to study the electrodynamic effects
of a shuttle borne satellite system with a conductive tether connection where an
electric current was induced in the tether as it swept through the Earth s magnetic
field Unfortunetely the two attempts in August 1992 and February 1996 resulted
in only partial success The former suffered from a spool failure resulting in a tether
deployment of only 256m of a planned 20km During the latter attempt a break
in the tethers insulation resulted in arcing leading to tether rupture Nevertheless
the information gathered by these attempts st i l l provided the engineers and scientists
invaluable information on the dynamic behaviour of this complex system
Canada also paved the way with novel designs of tethered system aimed
primari ly at the study of the ionosphere and Northern Lights (Aurora Borealis)
4
The O E D I P U S A and C (Observation of Electrified Distributions in the Ionospheric
P lasma - a Unique Strategy) missions[6] launched from a sounding rocket in January
1989 and November 1995 respectively provided insight into the complex dynamical
behaviour of two comparable mass satellites connected by a 1 km long spinning
tether
Final ly two experiments were conducted by N A S A called S E D S I and II
(Small Expendable Deployable System) [7] which hold the current record of 20 km
long tether In each of these missions a small probe (26 kg) was deployed from the
second stage of a D E L T A II rocket thus succesfully demonstrating the feasibility of
long tether deployment Retrieval of the tether which is significantly more difficult
has not yet been achieved
Several proposed applications are also currently under study These include the
study of earths upper atmosphere using probes lowered from the shuttle in a low earth
orbit ( L E O ) to the deployment and retrieval of satellites for servicing Even more
promising application concerns the generation of power for the proposed International
Space Station using conductive tethers The Canadian Space Agency (CSA) has also
proposed a dual-mass satellite system B I C E P S (Bl-static Canadian Experiment on
Plasmas in Space) to make simultaneous measurements at different locations in the
environment for correlation studies[8] A unique feature of the B I C E P S mission is
the deployment of the tether aided by the angular momentum of the system which
is ini t ia l ly cartwheeling about its orbit normal (Figure 1-3)
In addition to the three-body configuration multi-tethered systems have been
proposed for monitoring Earths environment in the global project monitored by
N A S A entitled Mission to Planet Ear th ( M T P E ) The proposed mission aims to
study pollution control through the understanding of the dynamic interactions be-
5
Retrieval of Satellite for Servicing
lt mdash
lt Satellite
Figure 1-3 Some applications of the tethered satellite systems (a) multiple comshymunication satellites at a fixed geostationary location (b) retrieval maneuver (c) proposed BICEPS configuration
6
tween the Earths atmosphere biosphere hydrosphere and cryosphere This wi l l be
accomplished with multiple tethered instrumentation payloads simultaneously soundshy
ing at different altitudes for spatial correlations Such mult ibody systems are considshy
ered to be the next stage in the evolution of tethered satellites Micro-gravity payload
modules suspended from the proposed space station for long-term experiments as well
as communications satellites with increased line-of-sight capability represent just two
of the numerous possible applications under consideration
12 Brief Review of the Relevant Literature
The design of multi-payload systems wi l l require extensive dynamic analysis
and parametric study before the final mission configuration can be chosen This
wi l l include a review of the fundamental dynamics of the simple two-body tethered
systems and how its dynamics can be extended to those of multi-body tethered system
in a chain or more generally a tree topology
121 Multibody 0(N) formulation
Many studies of two-body systems have been reported One of the earliest
contribution is by Rupp[9] who was the first to study planar dynamics of the simshy
plified Shuttle based Tethered Satellite System (STSS) His pioneering work led to
the discovery of pitch oscillation growth during the retrieval phase Later Baker[10]
advanced the investigation to the third dimension in addition to adding atmospheric
effects to the system A more complete dynamical model was later developed and
analyzed by M o d i and Misra[ l l 12] which included deployment and tether flexibility
A comprehensive survey of important developments and milestones in the area have
been compiled and reviewed by Mis ra and Modi[13] The major conclusions based on
the literature may be summarized as follows
bull the stationkeeping phase is stable
7
bull deployment can be unstable if the rate exceeds a crit ical speed
bull retrieval of the tether is inherently unstable
bull transverse vibrations can grow due to the Coriolis force induced during deshy
ployment and retrieval
Misra Amier and Modi[14] were one of the first to extend these results to the
three-satellite double pendulum case This simple model which includes a variable
length tether was sufficient to uncover the multiple equilibrium configurations of
the system Later the work was extended to include out-of-plane motion and a
preliminary reel-rate control law to regulate the librational motion[15] Kesmir i and
Misra[16] developed a general formulation for N-body tethered systems based on the
Lagrangian principle where three-dimensional motion and flexibility are accounted for
However from their work it is clear that as the number of payload or bodies increases
the computational cost to solve the forward dynamics also increases dramatically
Tradit ional methods of inverting the mass matrix M in the equation
ML+F = Q
to solve for the acceleration vector proved to be computationally costly being of the
order for practical simulation algorithms O(N^) refers to a mult ibody formulashy
tion algorithm whose computational cost is proportional to the cube of the number
of bodies N used It is therefore clear that more efficient solution strategies have to
be considered
Many improved algorithms have been proposed to reduce the number of comshy
putational steps required for solving for the acceleration vector Rosenthal[17] proshy
posed a recursive formulation based on the triangularization of the equations of moshy
tion to obtain an 0(N2) formulation He also demonstrates the use of a symbolic
manipulation program to reduce the final number of computations A recursive L a -
8
grangian formulation proposed by Book[18] also points out the relative inefficiency of
conventional formulations and proceeds to derive an algorithm which is also 0(N2)
A recursive formulation is one where the equations of motion are calculated in order
from one end of the system to the other It usually involves one or more passes along
the links to calculate the acceleration of each link (forward pass) and the constraint
forces (backward or inverse pass) Non-recursive algorithms express the equations of
motion for each link independent of the other Although the coupling terms st i l l need
to be included the algorithm is in general more amenable to parallel computing
A st i l l more efficient dynamics formulation is an O(N) algorithm Here the
computational effort is directly proportional to the number of bodies or links used
Several types of such algorithms have been developed over the years The one based
on the Newton-Euler approach has recursive equations and is used extensively in the
field of multi-joint robotics[1920] It should be emphasized that efficient computation
of the forward and inverse dynamics is imperative if any on-line (real-time) control
of the joint is to be achieved Hollerbach[21] proposed a recursive O(N) formulation
based on Lagranges equations of motion However his derivation was primari ly foshy
cused on the inverse dynamics and did not improve the computational efficiency of
the forward dynamics (acceleration vector) Other recursive algorithms include methshy
ods based on the principle of vir tual work[22] and Kanes equations of motion[23]
Keat[24] proposed a method based on a velocity transformation that eliminated the
appearance of constraint forces and was recursive in nature O n the other hand
K u r d i l a Menon and Sunkel[25] proposed a non-recursive algorithm based on the
Range-Space method which employs element-by-element approach used in modern
finite-element procedures Authors also demonstrated the potential for parallel comshy
putation of their non-recursive formulation The introduction of a Spatial Operator
Factorization[262728] which utilizes an analogy between mult ibody robot dynamics
and linear filtering and smoothing theory to efficiently invert the systems mass ma-
9
t r ix is another approach to a recursive algorithm Their results were further extended
to include the case where joints follow user-specified profiles[29] More recently Prad-
han et al [30] proposed a non-recursive Lagrangian algorithm for factorizing the mass
matrix leading to an O(N) formulation of the forward dynamics of the system A n
efficient O(N) algorithm where the number of bodies N in the system varies on-line
has been developed by Banerjee[31]
122 Issues of tether modelling
The importance of including tether flexibility has been demonstrated by several
researchers in the past However there are several methods available for modelling
tether flexibility Each of these methods has relative advantages and limitations
wi th regards to the computational efficiency A continuum model where flexible
motion is discretized using the assumed mode method has been proposed[3233] and
succesfully implemented In the majority of cases even one flexible mode is sufficient
to capture the dynamical behaviour of the tether[34] K i m and Vadali[35] proposed a
so-called bead model or lumped-mass approach that discretizes the tether using point
masses along its length However in order to accurately portray the motion of the
tether a significantly high number of beads are needed thus increasing the number of
computation steps Consequently this has led to the development of a hybrid between
the last two approaches[16] ie interpolation between the beads using a continuum
approach thus requiring less number of beads Finally the tether flexibility can also
be modelled using a finite-element approach[36] which is more suitable for transient
response analysis In the present study the continuum approach is adopted due to
its simplicity and proven reliability in accurately conveying the vibratory response
In the past the modal integrals arising from the discretization process have
been evaluted numerically In general this leads to unnecessary effort by the comshy
puter Today these integrals can be evaluated analytically using a symbolic in-
10
tegration package eg Maple V Subsequently they can be coded in F O R T R A N
directly by the package which could also result in a significant reduction in debugging
time More importantly there is considerable improvement in the computational effishy
ciency [16] especially during deployment and retrieval where the modal integral must
be evaluated at each time-step
123 Attitude and vibration control
In view of the conclusions arrived at by some of the researchers mentioned
above it is clear that an appropriate control strategy is needed to regulate the dyshy
namics of the system ie attitude and tether vibration First the attitude motion of
the entire system must be controlled This of course would be essential in the case of
a satellite system intended for scientific experiments such as the micro-gravity facilishy
ties aboard the proposed Internation Space Station Vibra t ion of the flexible members
wi l l have to be checked if they affect the integrity of the on-board instrumentation
Thruster as well as momentum-wheel approach[34] have been considered to
regulate the rigid-body motion of the end-satellite as well as the swinging motion
of the tether The procedure is particularly attractive due to its effectiveness over a
wide range of tether lengths as well as ease of implementation Other methods inshy
clude tension and length rate control which regulate the tethers tension and nominal
unstretched length respectively[915] It is usually implemented at the deployment
spool of the tether More recently an offset strategy involving time dependent moshy
tion of the tether attachment point to the platform has been proposed[3738] It
overcomes the problem of plume impingement created by the thruster control and
the ineffectiveness of tension control at shorter lengths However the effectiveness of
such a controller can become limited with an exceedingly long tether due to a pracshy
tical l imi t on the permissible offset motion In response to these two control issues a
hybrid thrusteroffset scheme has been proposed to combine the best features of the
11
two methods[39]
In addition to attitude control these schemes can be used to attenuate the
flexible response of the tether Tension and length rate control[12] as well as thruster
based algorithms[3340] have been proposed to this end M o d i et al [41] have sucshy
cessfully demonstrated the effectiveness of an offset strategy to damp undesired v i shy
brations Passive energy dissipative devices eg viscous dampers are also another
viable solution to the problem
The development of various control laws to implement the above mentioned
strategies has also recieved much attention A n eigenstructure assignment in conducshy
tion wi th an offset controller for vibration attenuation and momemtum wheels for
platform libration control has been developed[42] in addition to the controller design
from a graph theoretic approach[41] Also non-linear feedback methods such as the
Feedback Linearization Technique ( F L T ) that are more suitable for controlling highly
nonlinear non-autonomous coupled systems have also been considered[43]
It is important to point out that several linear controllers including the classhy
sic state feedback Linear Quadratic Regulator ( L Q R ) have received considerable
attention[3944] Moreover robust methods such as the Linear Quadratic Guas-
s ian Loop Transfer Recovery ( L Q G L T R ) method have also been developed and imshy
plemented on tethered systems[43] A more complete review of these algorithms as
well as others applied to tethered system has been presented by Pradhan[43]
13 Scope of the Investigation
The objective of the thesis is to develop and implement a versatile as well as
computationally efficient formulation algorithm applicable to a large class of tethered
satellite systems The distinctive features of the model can be summarized as follows
12
bull TV-body 0(N) tethered satellite system in a chain-type topology
bull the system is free to negotiate a general Keplerian orbit and permitted to
undergo three dimensional inplane and out-of-plane l ibrtational motion
bull r igid bodies constituting the system are free to execute general rotational
motion
bull three dimensional flexibility present in the tether which is discretized using
the continuum assumed-mode method with an arbitrary number of flexible
modes
bull energy dissipation through structural damping is included
bull capability to model various mission configurations and maneuvers including
the tether spin about an arbitrary axis
bull user-defined time dependent deployment and retrieval profiles for the tether
as well as the tether attachment point (offset)
bull attitude control algorithm for the tether and rigid bodies using thrusters and
momentum-wheels based on the Feedback Linearization Technique
bull the algorithm for suppression of tether vibration using an active offset (tether
attachment point) control strategy based on the optimal linear control law
( L Q G L T - R )
To begin with in Chapter 2 kinematics and kinetics of the system are derived
using the O(N) formulation methodology developed by Pradhan et al [30] Chapshy
ter 3 discusses issues related to the development of the simulation program and its
validation This is followed by a detailed parametric study of several mission proshy
files simulated as particular cases of the versatile formulation in Chapter 4 Next
the attitude and vibration controllers are designed and their effectiveness assessed in
Chapter 5 The thesis ends with concluding remarks and suggestions for future work
13
2 F O R M U L A T I O N OF T H E P R O B L E M
21 Kinematics
211 Preliminary definitions and the i ^ position vector
The derivation of the equations of motion begins with the definition of the
inertial position of an arbitrary link of the mult ibody system Let the i ^ link of the
system be free to translate and rotate in 3-D space From Figure 2-1 the position
vector Rdm^ to the mass element drrn on the i ^ link wi th reference to the inertial
frame FQ can be written as
Here D represents the inertial position of the t h body fixed frame Fj relative to FQ
ri = [xiyiZi]T is the rigid position vector of drrii wi th respect to F J and f(fi) is
the flexible deformation at fj also with respect to the frame Fj Bo th these vectors
are relative to Fj Note in the present model tethers (i even) are considered flexible
and X corresponds to the nominal position along the unstretched tether while a and
ZJ are by definition equal to zero On the other hand the rigid bodies referred to as
satellites (i odd) including the platform are taken to be rigid ie f(fj) = 0 T j is
the rotation matrix used to express body-fixed vectors with reference to the inertial
frame
212 Tether flexibility discretization
The tether flexibility is discretized with an arbitrary but finite number of
Di + Ti(i + f(i)) (21)
14
15
modes in each direction using the assumed-mode method as
Ui nvi
wi I I i = 1
nzi bull
(22)
where nXi nyi and n 2 ^ are the number of modes used in the X m and Z directions
respectively For the ] t h longitudinal mode the admissible mode shape function is
taken as 2 j - l
Hi(xiii)=[j) (23)
where l is the i ^ tether length[32] In the case of inplane and out-of-plane transverse
deflections the admissible functions are
ampyi(xili) = ampZixuk) = v ^ s i n JKXj
k (24)
where y2 is added as a normalizing factor In this case both the longitudinal and
transverse mode shapes satisfy the geometric boundary conditions for a simple string
in axial and transverse vibration
Eq(22) is recast into a compact matrix form as
ff(i) = $i(xili)5i(t) (25)
where $i(x l) is the matrix of the tether mode shapes defined as
0 0
^i(xiJi) =
^ X bull bull bull ^XA
0
0 0
0
G 5R 3 x n f
(26)
and nfj = nx- + nyi + nZi Note $J(XJJ) is a function of the spatial coordinate x
and not of yi or ZJ However it is also a function of the tether length parameter li
16
which is time-varying during deployment and retrieval For this reason care must be
exercised when differentiating $J(XJJ) with respect to time such that
d d d k) = Jtregi(Xi li) = ^ 7 $ t ( ^ i k)Xi + gf$i(xigt li)k- (2-7)
Since x represents the nominal rate of change of the position of the elemental mass
drrii it is equal to j (deploymentretrieval rate) Let t ing
( 3 d
dx~ + dT)^Xili^ ( 2 8 )
one arrives at
$i(xili) = $Di(xili)ii (29)
The time varying generalized coordinate vector 5 in Eq(25) is composed of
the longitudinal and transverse components ie
ltM) G s f t n f i x l (210)
where 5X^ 5y- and Sz^ are nx- x 1 ny x 1 and nz- x 1 size vectors and correspond to
$ X and $ Z respectively
213 Rotation angles and transformations
The matrix T j in Eq(21) represents a rotation transformation from the body
fixed frame Fj to the inertial frame FQ ie FQ = T j F j It is defined by the 3-2-1
ordered sequence of elementary rotations[46]
1 P i tch F[ = Cf ( a j )F 0
2 R o l l F = C f (3j)F = C7f (3j)Cf ( a j )F 0
3 Yaw F = C7( 7 i )F = C7(7l)cf^)Cf(ai)FQ = Q F 0
17
where
c f (A) =
-Sai Cai
0 0
and
C(7i ) =
0 -s0i] = 0 1 0
0 cpi
i 0 0 = 0
0
(211)
(212)
(213)
Here Eq(211) corresponds to a rotation of ctj (pitch) about the inertial axis ZQ (orbit
normal) of frame FQ resulting in a new intermediate frame F It is then followed by
a roll rotation fa about the axis y of F- giving a second intermediate frame F-
Final ly a spin rotation 7J about the axis z of F- is given resulting in the body
fixed frame Fj for the i lt l link Since all the three rotation matrices are orthogonal
it follows that FQ = C~1Fi = Cj F and hence
C CPi Cai ~ CH Sai + SrYi SPi Sai 5 7 i Sampi + Cli SPi Cci CPi Sai C1i Cai + 5 7 i SPi ~ S1 Ca + C 7 bull Sp Sai
(214) SliCPi CliCPi
Here Cx and Sx are abbreviations for cos(x) and sin(x) respectively
214 Inertial velocity of the ith link
mdash Lett ing pj = fj - f $j(5j and differentiating Eq(21) with respect to time gives
m- D j + T j f j + T^Si + T j pound + T j $ j 5 j (215)
Each of the terms appearing in Eq(215) must be addressed individually Since
fj represents the rigid body motion within the link it only has meaning for the
deployable tether and since yi = Zj = 0 for al l tether links (i even) fj = Ifi where 1
is the unit vector 1 0 0T For the case of rigid links (i odd) fj = 0
18
Evaluating the derivative of each angle in the rotation matrix gives
T^i9i = Taiai9i + Tppigi + T^fagi (216)
where Tx = J j I V Collecting the scaler angular velocity terms and defining the set
Vi = deg^gt A) 7i^ Eq(216) can be rewritten as
T i f t = Pi9i)fji (217)
where Pi(gi) is a column matrix defined by
Pi(9i)Vi = [Taigi^p^iT^gilffi (218)
Inserting Eq(29) and Eq(217) into Eq(215) and defining Sj = 2 + leads to
215 Cylindrical orbital coordinates
The Di term in Eq(219) describes the orbital motion of the t h l ink and
is composed of the three Cartesian coordinates DXi A ^ DZi However over the
cycle of one orbit DXi and Dyi vary dramatically by around the order of Earths
radius This must be avoided since large variations in the coordinates can cause
severe truncation errors during their numerical integration For this reason it is
more convenient to express the Cartesian components in terms of more stationary
variables This is readily accomplished using cylindrical coordinates
However it is only necessary to transform the first l inks orbital components
ie Lgti since only D is a generalized coordinates From Figure 2-2 it is apparent
(219)
that
(220)
19
Figure 2-2 Vector components in cylindrical orbital coordinates
20
Eq(220) can be rewritten in matrix form as
cos(0i) 0 0 s i n ( 0 i ) 0 0
0 0 1
Dri
h (221)
= DTlDSv
where Dri is the inplane radial distance 9 the true anomaly and DZl the out-of-
plane distance normal to the orbital plane The total derivative of D wi th respect
to time gives in column matrix form
1 1 1 1 J (222)
= D D l D 3 v
where [cos(^i) -Dnsm(8i) 0 ]
(223)
For all remaining links ie i = 2 N
cos(0i) - A - j S i n ^ i ) 0 sin(^i) D r i c o s ( 0 i ) 0
0 0 1
DTl = D D i = I d 3 i = 2N (224)
where is the n x n identity matrix and
2N (225)
216 Tether deployment and retrieval profile
The deployment of the tether is a critical maneuver for this class of systems
Cr i t i ca l deployment rates would cause the system to librate out of control Normally
a smooth sinusoidal profile is chosen (S-curve) However one would also like a long
tether extending to ten twenty or even a hundred kilometers in a reasonable time
say 3-4 orbits This is usually difficult or impossible to accomplish solely wi th one
S-curve profile Often a composite S-curve-Steady-S-curve profile is adopted Thus
21
the deployment scheme for the t h tether can be summarized as
k = Tr l - cos (^r(ti - to-)) 1 ltUlt h 2 r V A V
k = Vov tiilttiltt2i
k = IT j 1 - C deg S lkitl ^ ^ lt i lt ^
(226)
where to- tt are the ini t ia l and final times of the deployment or retrieval maneuver
respectively A t j = t - tQ mdash tt mdash t2- and the steady deployment velocity VQ is
calculated based on the continuity of l and l at the specified intermediate times t
and to For the case of
and
Ai = i(ti) - ^(t 0 -) = - ^ ( t 2 i ) (227)
2Ali + li(tf) -h(tQ)
V = lUli l^lL (228)
1 tf-tQ v y
A t = (229)
22 Kinetics and System Energy
221 Kinetic energy
The kinetic energy of the ith link is given by
Kei = lJm Rdrnfi^Ami (230)
Setting Eq(219) in a matrix-vector form gives the relation
4 ^ ^ PM) T i $ 4 TiSi]i[t (231)
22
where
m
k J
x l (232)
and nq = -nfj + 7 is the number of generalized coordinates in each link Inserting
Eq(231) into Eq(230) and integrating the kinetic energy for the t h l ink can be
rewritten as
1 T
(233)
where Mti is the links symmetric mass matrix
Mt =
m i D D T D D DDTPifgidmi) DDTTi J ^dm D n T Sidm rn
sym fP(9i)Pi(9i)dmi J P[g^T^idm J P ( pound ) T ^ m sym sym J ^f^idm f ^f^dmi sym sym sym J sfsidmi
pound Sfttiqxnq
(234)
and mi is the mass of the Ith link The kinetic energy for the entire system ie N
bodies can now be stated as
1 bull T bull 1 bull T Ke = J2 MtA = o laquo Mtqt 2 ^ -i
i=l (235)
where qt = qj^^ bull bull -qfNT a n d Mt is a block diagonal decoupled mass matrix
of the system with Mf- on its diagonal
(236) Mt =
0 0
0 Mt2
0
0 0
MH
0 0 0
0 0 0 bull MtN
222 Simplification for rigid links
The mass matrix given by Eq(234) simplifies considerably for the case of rigid
23
links and is given as
Mti
m D D l
T D D l DD^Piiffidmi) 0
sym 0
0
fP(fi)Pi(i)dmi 0 0 0
0 Id nfj
0 m~
i = 1 3 5 N
(237)
Moreover when mult iplying the matrices in the (11) block term of Eq(237) one
gets
D D l
1 D D l = 1 0 0 0 D n
2 0 (238) 0 0 1
which is a diagonal matrix However more importantly i t is independent of 6 and
hence is always non-singular ensuring that is always invertible
In the (12) block term of M ^ it is apparent that f fidmi is simply the
definition of the centre of mass of the t h rigid link and hence given by
To drrii mdash miff (239)
where fcmi is the position vector of the centre of mass of link i For the case of the
(22) block term a little more attention is required Expanding the integrand using
Eq(218) where ltj = fj gives
fj Ta Tafi fj T Q Tpfi f- T Q T 7 i f j
sym sym
Now since each of the terms of Eq(240) is a scaler or a 1 x 1 matrix it is therefore
equal to its trace ie sum of its diagonal elements thus
tr(ffTai
TTaifi) tr(ffTai
TTpfi) tr(ff Ta^T^fi)
fjT^Tpn rfTp^fi (240)
P(rj)Pj(fi) = syrri trifjTpTTpfi) t r ^ T ^ T ^
sym sym tr(ff T 7 T r 7 f j) J
(241)
Using two properties of the trace function[47] namely
tr(ABC) = tr(BCA) = trCAB) (242)
24
and
j tr(A) =tr(^J A^j (243)
where A B and C are real matrices it becomes quite clear that the (22) block term
of Eq(237) simplifies to
P(fi)Pii)dmi =
tr(Tai
TTaiImi) tTTai
TTpImi) t r (T Q r T 7 7 m ) trTaTTpImi) tr(Tp
TT7Jmi) sym sym sym
(244)
where Im- is an alternate form of the inertia tensor of the t h link which when
expanded is
Imlaquo = I nfi dm bull i sym
sym sym
J x^dmi J Xydmi XZdm sym J yi2dmi f yiZidm
sym J z^dmi J
xVi -L VH sym ijxxi + lyy IZZJ)2
(245)
Note this is in terms of moments of inertia and products of inertia of the t f l rigid
link A similar simplification can be made for the flexible case where T is replaced
with Qi resulting in a new expression for Im
223 Gravitational potential energy
The gravitational potential energy of the ith link due to the central force law
can be written as
V9i f dm f
-A mdash = -A bullgttradelti RdmA
dm
inn Lgti + T j ^ | (246)
where z = 3986 x 105 km 3s 2 (Earths gravitational constant) Expanding binomially
and retaining up to third order terms gives
Vgi ~ ~ D T 2D [J gfgidrrii + 23f T j gidm mdashDT
D2 i l I I Df Tt I grffdmiTiDi
(2-lt 7)
25
where D = D is the magnitude of the inertial position vector to the frame F The
integrals in the above expression are functions of the flexible mode shapes and can be
evaluated through symbolic manipulation using an existing package such as Maple
V In turn Maple V can translate the evaluated integrals into F O R T R A N and store
the code in a file Furthermore for the case of rigid bodies it can be shown quite
readily that
j gjgidrrii = j rjfidrrn = [IXXi + Iyy + IZZi)2 (248)
Similarly the other integrals in Eq(234) can be integrated symbolically
224 Strain energy
When deriving the elastic strain energy of a simple string in tension the
assumptions of high tension and low amplitude of transverse vibrations are generally
valid Consequently the first order approximation of the strain-displacement relation
is generally acceptable However for orbiting tethers in a weak gravitational-gradient
field neglecting higher order terms can result in poorly modelled tether dynamics
Geometric nonlinearities are responsible for the foreshortening effects present in the
tether These cannot be neglected because they account for the heaving motion along
the longitudinal axis due to lateral vibrations The magnitude of the longitudinal
oscillations diminish as the tether becomes shorter[35]
W i t h this as background the tether strain which is derived from the theory
of elastic vibrations[3248] is given by
(249)
where Uj V and W are the flexible deformations along the Xj and Z direction
respectively and are obtained from Eq(25) The square bracketed term in Eq(249)
represents the geometric nonlinearity and is accounted for in the analysis
26
The total strain energy of a flexible link is given by
Ve = I f CiEidVl (250)
Substituting the stress-strain relationship 0 = E(poundi the strain energy is now given
by
Vei=l-EiAi J l l d x h (251)
where E^ is the tethers Youngs modulus and A is the cross-sectional area The
tether is assumed to be uniform thus EA which is the flexural stiffness of the
tether is assumed to be constant Eq(251) is evaluated symbolically for arbitrary
number of modes
225 Tether energy dissipation
The evaluation of the energy dissipation due to tether deformations remains
a problem not well understood even today Here this complex phenomenon is repshy
resented through a simplified structural damping model[32] In addition the system
exhibits hysteresis which also must be considered This is accomplished using an
augmented complex Youngs modulus
EX^Ei+jEi (252)
where Ej^ is the contribution from the structural damping and j = The
augmented stress-strain relation is now given as
a = Epoundi = Ei(l+jrldi)ei (253)
where 77^ = EjE^ ltC 1 is defined as the structural damping coefficient determined
experimentally [49]
27
If poundi is a harmonic function with frequency UJQ- then
jet = (254) wo-
Substituting into Eq(253) and rearranging the terms gives
e + I mdash I poundi = o-i + ad (255)
where ad and a are the stress with and without structural damping respectively
Now the Rayleigh dissipation function[50] for the ith tether can be expressed as
(ii) Jo
_ 1 EjAiVdi rU (2-56^ 2 w 0 Jo
The strain rate ii is the time derivative of Eq(249) The generalized external force
due to damping can now be written as Qd = Q^ bull bull bull QdNT where
dWd
^ = - i p = 2 4 ( 2 5 7 )
= 0 i = 13 N
23 0(N) Form of the Equations of Mot ion
231 Lagrange equations of motion
W i t h the kinetic energy expression defined and the potential energy of the
whole system given by Pe = ]Cn=l( 5i+ ej) the equations of motion can be obtained
quite readily using the Lagrangian principle
where q is the set of generalized coordinates to be defined in a later section
Substituting Eqs(235247251 and 257) into Eq(258) leads to the familiar
matr ix form of the non-linear non-autonomous coupled equations of motion for the
28
system
Mq t)t+ F(q q t) = Q(q q t) (259)
where M(q t) is the nonlinear symmetric mass matrix and F(q q t) is the forcing
function which can be written in matrix form as
Q(ltfgt Qi 0 is the vector of non-conservative generalized forces including control inshy
puts acting on the system A detailed expansion of Eq(260) in tensor notation is
developed in Appendix I
Numerical solution of Eq(259) in terms of q would require inversion of the
mass matrix However M is a full matrix of size nqq x nqq where nqq = Nnq is the
total number of generalized coordinates Therefore direct inversion of M would lead
to a large number of computation steps of the order nqq^ or higher The objective
here is to develop an order-N O(N) algorithm for obtaining M _ 1 that minimizes
the computational effort This is accomplished through the following transformations
ini t ia l ly developed for the stationkeeping case by Pradhan et al [30]
232 Generalized coordinates and position transformation
During the derivation of the energy expressions for the system the focus was
on the decoupled system ie each link was considered independent of the others
Thus each links energy expression is also uncoupled However the constraints forces
between two links must be incorporated in the final equations of motion
Let
e Unqxl (261) m
be the set of coupled coordinates of the t h l ink such that q = qT q2 bull bull bull 0T i s the
29
systems generalized coordinates Let qt- be the set of auxiliary decoupled coordinates
defined by Eq(232) such that qt = qj^ qj^ Qtj^1 bull The only difference between
qtj and q is the presence of D against d is the position of F wi th respect to
the inertial frame FQ whereas d is defined as the offset position of F relative to and
projected on the frame It can be expressed as d = dXidyidZi^ For the
special case of link 1 D = d
From Figure 2-1 can written as
Di = A - l + T i _ i J i _ i + di) + T V i S i - i f t - i + dXi)8i- (262)
Denoting KAj = $j(Zj + dx- ) Eq(262) can be rewritten in summation form as
i ampi = H ( T i - 1 ^ + T i - l K A i - l - l + T j - l ^ - l ) bull ( 2 - 6 3 )
3 = 1
Introducing the index substitution k mdash j mdash 1 into Eq(263)
i-1 D = ^2 (Tfc_ilt4 + TkKAk6k + Tkilk) + T^di (264)
k=l
since d = D TQ = 1^ and IQ DQ SQ are null vectors Recasting Eq(264) in
matr ix form leads to i-1
Qt Jfe=l
where
and
For the entire system
T fc-1 0 TkKAk
0 0 0 0 0 0 0 0 0 0 0 0 -
r T i - i 0 0 0 0 0 0 0 0 0 G
0 0 0 1
Qt = RpQ
G R n lt x l
(265)
(266)
G Unq x l (267)
(268)
30
where
RP
R 1 0 0 R R 2 0 R RP
2 R 3
0 0 0
R RP R PN _ 1 RN J
(269)
Here RP is the lower block triangular matrix which relates qt and q
233 Velocity transformations
The two sets of velocity coordinates qti and qi are related by the following
transformations
(i) First Transformation
Differentiating Eq(262) with respect to time gives the inertial velocity of the
t h l ink as
D = A - l + Ti_iK + + K A i _ i ^ _ i
+ T j _ i ^ + Ti-ii + T i - i K A i - i ^ + T j - i K A ^ i ^ i (270)
(271) (
Defining K A D = dddXi+1KAi K A L = d d K A and hi = di+1 +lfi +KA^i
results in K A ^ i = K A D j i o ^ + K A L i _ i i _ i
= KADi_iiTdi + K A L j _ i ^ _ i
Inserting Eq(271) into Eq(270) and rearranging gives
= A _ + T i ^ d g + K A D ^ x ^ i ^ ^ + Pi-^hi-iH-i
+ T i _ i K A i _ i 5 _ i + T i_ i2 + K A L j _ i 5 j _ i 4 _ i (272)
Using Eq(272) to define recursively - D j _ i and applying the index substitution
c = j mdash 1 as in the case of the position transformation it follows that
Di = ^ T ^ i K D ^ ^ + Pk(hk)ffk + T f c K A f c ^ + TfcKLfc4 Jfc=l (273)
+ T j _ i K D j _ i lt i ii
31
where K D j _ i = + K A D ^ i ^ - i r 7 and K L j = i + K A L j 5 j Finally by following a
procedure similar to that used for the position transformation it can be shown that
for the entire system
t = Rvil (2-74)
where Rv is given by the lower block diagonal matrix
Here
Rv
Rf 0 0 R R$ 0 R Rl RJ
0 0 0
T3V nN-l
R bullN
e Mnqqxl
Ri =
T i _ i K D i _ i TiKAi T j K L j 0 0 0 0 bullo 0 0 0 0 0 0 0
e K n lt x l
RS
T i-1 0 0 o-0 0 0 0 0 0
0 0 0 1
G Mnq x l
(275)
(276)
(277)
(ii) Second Transformation
The second transformation is simply Eq(272) set into matrix form This leads
to the expression for qt as
(278)
where Id Pihi) T j K A j T j K L j 0 0 0 0 0 0
0 0
0 0
0 0
G $lnq x l
Thus for the entire system
(279)
pound = Rnjt + Rdq
= Vdnqq - R nrlRdi (280)
32
where
and
- 0 0 0 RI 0 0
Rn = 0 0
0 0
Rf 0 0 0 rgtd K2 0
Rd = 0 0 Rl
0 0 0
RN-1 0
G 5R n w x l
0 0 0 e 5 R N ^ X L
(281)
(282)
234 Cylindrical coordinate modification
Because of the use of cylindrical coorfmates for the first body it is necessary
to slightly modify the above matrices to account for this coordinate system For the
first link d mdash D = D ^ D S L thus
Qti = 01 =gt DDTigi (283)
where
D D T i
Eq(269) now takes the form
DTL 0 0 0 d 3 0 0 0 ID
0 0 nfj 0
RP =
P f D T T i R2 0 P f D T T i Rp
2 R3
iJfDTTi i^
x l
0 0 0
R PN-I RN
(284)
(285)
In the velocity transformation a similar modification is necessary since D =
DD^DS^ Mak ing the substitution it can be shown that Eq(276) and Eq(279) now
33
takes the form for i = 1
T)V _ pn _ r t j mdash rt^ mdash 0 0 0 0
0 0 0 0 0 0 0 0
(286)
The rest of the matrices for the remaining links ie i = 2 N are unchanged
235 Mass matrix inversion
Returning to the expression for kinetic energy substitution of Eq(274) into
Eq(235) gives the first of two expressions for kinetic energy as
Ke = -q RvT MtRv (287)
Introduction of Eq(280) into Eq(235) results in the second expression for Ke as
1X 2q (hnqq-RnrlRd Mt (idnqq - R nrlRd (288)
Thus there are two factorizations for the systems coupled mass matrix given as
M = RV MtRv
M (hnqq - RnrlRd Mt ( i d n q q - R n r 1 R d
(289)
(290)
Inverting Eq(290)
r-1 _ ( rgtd~^ M 1 = (Ra) l d n q q - Rn)Mf1 (Rd)~l(Idnqq - Rn) (291)
Since both Rd and Mt are block diagonal matrices their inverse is simply the inverse
34
of each block on the diagonal ie
Mr1
h 0 0
0 0
0 0
0 0 0
0
0
0
tN
(292)
Each block has the dimension nq x nq and is independent of the number of bodies
in the system thus the inversion of M is only linearly dependent on N ie an O(N)
operation
236 Specification of the offset position
The derivation of the equations of motion using the O(N) formulation requires
that the offset coordinate d be treated as a generalized coordinate However this
offset may be constrained or controlled to follow a prescribed motion dictated either by
the designer or the controller This is achieved through the introduction of Lagrange
multipliers[51] To begin with the multipliers are assigned to all the d equations
Thus letting f = F - Q Eq(259) becomes
Mq + f = P C A (293)
where A G K 3 - 7 V x l is the vector of Lagrange multipliers and Pc G s R n 9 x 3 A r is the
permutation matrix assigning the appropriate Aj to its corresponding d equation
Inverting M and pre-multiplying both sides by PcT gives
pcTM-pc
= ( dc + PdegTM-^f) (294)
where dc is the desired offset acceleration vector Note both dc and PcTM lf are
known Thus the solution for A has the form
A pcTM-lpc - l dc + PC1 M (295)
35
Now substituting Eq(295) into Eq(293) and rearranging the terms gives
q = S(f Ml) = -M~lf + M~lPc
which is the new constrained vector equation of motion with the offset specified by
dc Note that pre-multiplying M~l by P c T provides the rows of M _ 1 corresponding
to the d equations Similarly post-multiplying by Pc leads the columns of the matrix
corresponding to the d equations
24 Generalized Control Forces
241 Preliminary remarks
The treatment of non-conservative external forces acting on the system is
considered next In reality the system is subjected to a wide variety of environmental
forces including atmospheric drag solar radiation pressure Earth s magnetic field as
well as dissipative devices to mention a few However in the present model only
the effect of active control thrusters and momemtum-wheels is considered These
generalized forces wi l l be used to control the attitude motion of the system
In the derivation of the equations of motion using the Lagrangian procedure
the contribution of each force must be properly distributed among all the generalized
coordinates This is acheived using the relation
nu ^5 Qk=E Pern ^ (2-97)
m=l deg Q k
where Qk and qk are the kth elements of Q and ltf respectively[51] Fem is the
mth external force acting on the system at the location Rm from the inertial frame
and nu is the total number of actuators Derivation of the generalized forces due to
thrusters and momentum wheels is given in the next two subsections
pcTM-lpc (dc + PcTM- Vj (2-96)
36
242 Generalized thruster forces
One set of thrusters is placed on each subsatellite ie al l the rigid bodies
except the first (platform) as shown in Figure 2-3 They are capable of firing in the
three orthogonal directions From the schematic diagram the inertial position of the
thruster located on body 3 is given as
Defining
R3 = di + Tid2 + T2(hi + KA2lt52) + T 3 f c m 3
= d + dfY + T 3 f c m 3
df = Tjdj+i + Tj+1lj+ii + KAj+i5j+i)
the thruster position for body 5 is given as
(298)
(299)
R5 = di+ dfx + d 3 + T5fcm
Thus in general the inertial position for the thruster on the t h body can be given
(2100)
as i - 2
k=l i1 cm (2101)
Let the column matrix
[Q i iT 9 R-K=rR dqj lA n i
WjRi A R i (2102)
where i = 3 5 N and j mdash 1 2 i The vector derivative of R wi th respect to
the scaler components qk is stored in the kth column of Thus for the case of
three bodies Eq(297) for thrusters becomes
1 Q
Q3
3 -1
T 3 T t 2 (2103)
37
Satellite 3
mdash gt -raquo cm
D1=d1
mdash gt
a mdash gt
o
Figure 2-3 Inertial position of subsatellite thruster forces
38
where ft = 0TtaTt^T is the thrust vector for the t h l ink (tether i-1) Here
Ttn and Tta are the thrust forces along the inplane m and out-of-plane Z directions i Pi
respectively For the case of 5-bodies
Q
~Qgt1
0
0
( T ^ 2 ) (2104)
The result for seven or more bodies follows the pattern established by Eq(2104)
243 Generalized momentum gyro torques
When deriving the generalized moments arising from the momemtum-wheels
on each rigid body (satellite) including the platform it is easier to view the torques as
coupled forces From Figure 2-4 it is apparent that for the ith l ink the generalized
forces constituting the couple are
(2105)
where
Fei = Faij acting at ei = eXi
Fe2 = -Fpk acting at e 2 = C a ^ l
F e 3 = F7ik acting at e 3 = eyj
Expanding Eq(2105) it becomes clear that
3
Qk = ^ F e i
i=l
d
dqk (2106)
which is independent of Ri- Thus the moments Mma = 2FaieXi Mm^ = 2FpeXi
and Mmi = 2FlieVi Transforming the coordinates to the body fixed frame using
the rotation matrix the generalized force due to the momentum-wheels on link i can
39
Figure 2-4 Coupled force representation of a momentum-wheel on a rigid body
40
be written as
d d d Qi = Td bull ^ T i t M m a - Tik bull mdashT^Mmp + T ^ bull T J j M ^
QiMmi
where Mm = Mma M m Mmi T
lt3 T J ^ T - T i f c bull ^ T i pound T ^ - ^ T z j
(2107)
(2108)
Note combining the thruster forces and momentum-wheel torques a compact
expression for the generalized force vector for 5 bodies can be written as
Q
Q T 3 0 Q T 5 0
0 0 0
0 lt33T3 Qyen Q5
3T5 0
0 0 0 Q4T5 0
0 0 0 Q5
5T5
M M I
Ti 2 m 3
Quu (2109)
The pattern of Eq(2109) is retained for the case of N links
41
3 C O M P U T E R I M P L E M E N T A T I O N
31 Prel iminary Remarks
The governing equations of motion for the TV-body tethered system were deshy
rived in the last chapter using an efficient O(N) factorization method These equashy
tions are highly nonlinear nonautonomous and coupled In order to predict and
appreciate the character of the system response under a wide variety of disturbances
it is necessary to solve this system of differential equations Although the closed-
form solutions to various simplified models can be obtained using well-known analytic
methods the complete nonlinear response of the coupled system can only be obtained
numerically
However the numerical solution of these complicated equations is not a straightshy
forward task To begin with the equations are stiff[52] be they have attitude and
vibration response frequencies separated by about an order of magnitude or more
which makes their numerical integration suceptible to accuracy errors i f a properly
designed integration routine is not used
Secondly the factorization algorithm that ensures efficient inversion of the
mass matrix as well as the time-varying offset and tether length significantly increase
the size of the resulting simulation code to well over 10000 lines This raises computer
memory issues which must be addressed in order to properly design a single program
that is capable of simulating a wide range of configurations and mission profiles It
is further desired that the program be modular in character to accomodate design
variations with ease and efficiency
This chapter begins with a discussion on issues of numerical and symbolic in-
42
tegration This is followed by an introduction to the program structure Final ly the
validity of the computer program is established using two methods namely the conshy
servation of total energy and comparison of the response results for simple particular
cases reported in the literature
32 Numerical Implementation
321 Integration routine
A computer program coded in F O R T R A N was written to integrate the equashy
tions of motion of the system The widely available routine used to integrate the difshy
ferential equations D G E A R [ 5 3 ] is based on Gears predictor-corrector method[54]
It is well suited to stiff differential equations as i t provides automatic step-size adjustshy
ment and error-checking capabilities at each iteration cycle These features provide
superior performance over traditional 4^ order Runge-Kut ta methods
Like most other routines the method requires that the differential equations be
transformed into a first order state space representation This is easily accomplished
by letting
thus
X - ^ ^ f ^ ) =ltbullbull) (3 -2 )
where S(fMdc) is defined by Eq(296) Eq(32) represents the new set of 2nqq
first order differential equations of motion
322 Program structure
A flow chart representing the computer programs structure is presented in
Figure 3-1 The M A I N program is responsible for calling the integration routine It
also coordinates the flow of information throughout the program The first subroutine
43
called is INIT which initializes the system variables such as the constant parameters
(mass density inertia etc) and the ini t ial conditions I C (attitude angles orbital
motion tether elastic deformation etc) from user-defined data files Furthermore all
the necessary parameters required by the integration subroutine D G E A R (step-size
tolerance) are provided by INIT The results of the simulation ie time history of the
generalized coordinates are then passed on to the O U T P U T subroutine which sorts
the information into different output files to be subsequently read in by the auxiliary
plott ing package
The majority of the simulation running time is spent in the F C N subroutine
which is called repeatedly by D G E A R F C N generates the fv(Xi) vector defined in
Eq(32) It is composed of several subroutines which are responsible for the comshy
putation of all the time-varying matrices vectors and scaler variables derived in
Chapter 2 that comprise the governing equations of motion for the entire system
These subroutines have been carefully constructed to ensure maximum efficiency in
computational performance as well as memory allocation
Two important aspects in the programming of F C N should be outlined at this
point As mentioned earlier some of the integrals defined in Chapter 2 are notably
more difficult to evaluate manually Traditionally they have been integrated numershy
ically on-line (during the simulation) In some cases particularly stationkeeping the
computational cost is quite reasonable since the modal integrals need only be evalushy
ated once However for the case of deployment or retrieval where the mode shape
functions vary they have to be integrated at each time-step This has considerable
repercussions on the total simulation time Thus in this program the integrals are
evaluated symbolically using M A P L E V[55] Once generated they are translated into
F O R T R A N source code and stored in I N C L U D E files to be read by the compiler
Al though these files can be lengthy (1000 lines or more) for a large number of flexible
44
INIT
DGEAR PARAM
IC amp SYS PARAM
MAIN
lt
DGEAR
lt
FCN
OUTPUT
STATE ACTUATOR FORCES
MOMENTS amp THRUST FORCES
VIBRATION
r OFFS DYNA
ET MICS
Figure 3-1 Flowchart showing the computer program structure
45
modes they st i l l require less time to compute compared to their on-line evaluation
For the case of deployment the performance improvement was found be as high as
10 times for certain cases
Secondly the O(N) algorithm presented in the last section involves matrices
that contained mostly zero terms (RP Rv and Mplusmn) The sparse structure of these mashy
trices can be used advantageously to improve the computational efficiency by avoiding
multiplications of zero elements The result is a quick evaluation of fv(Xt) Moreshy
over there is a considerable saving of computer memory as zero elements are no
longer stored
Final ly the C O N T R O L subroutine which is part of F C N is designed to calshy
culate the appropriate control forces torques and offset dynamics which are required
to stabilize the librational ( A T T I T U D E ) and vibrational ( V I B R A T I O N ) motion of
the system
33 Verification of the Code
The simulation program was checked for its validity using two methods The
first verified the conservation of total energy in the absence of dissipation The second
approach was a direct comparison of the planar response generated by the simulation
program with those available in the literature A s numerical results for the flexible
three-dimensional A^-body model are not available one is forced to be content with
the validation using a few simplified cases
331 Energy conservation
The configuration considered here is that of a 3-body tethered system with
the following parameters for the platform subsatellite and tether
46
h =
1091430 -8135 328108 -8135 8646050 27116 328108 27116 8286760
kg bull m 2 platform inertia
kg bull m 2 end-satellite inertia 200 0 0
0 400 0 0 0 400
m = 90000 kg mass of the space station platform
m2 = 500 kg mass of the end-satellite
EfAt = 61645 N tether elastic stiffness
pt mdash 49 kg km tether linear density
The tether length is taken to remain constant at 10 km (stationkeeping case)
A n in i t ia l disturbance in pitch and roll of 2deg and 1deg respectively is given to both
the rigid bodies and the flexible tether In addition the tether is ini t ia l ly deflected
by 45 m in the longitudinal direction at its end and 05 m in both the inplane and
out-of-plane directions in the first mode The tether attachment point offset at the
platform (satellite 1) end is 1 m in all the three directions ie d2 = 11 lTm The
attachment point at the subsatellite (satellite 2) end is 1 m in the x direction ie
fcm3 = 10 0T
The variation in the kinetic and potential energies is presented for the above-
mentioned case in Figure 3-2(a) As seen from the plot there is a continuous mutual
exchange between the kinetic and potential energies however the variation in the
total energy remains zero (Figure 3-2b) It should be noted that after 1 orbit the
variation of both kinetic and potential energy does not return to 0 This is due to
the attitude and elastic motion of the system which shifts the centre of mass of the
tethered system in a non-Keplerian orbit
Similar results are presented for the 5-body case ie a double pendulum (Figshy
ure 3-3) Here the system is composed of a platform and two subsatellites connected
47
STSS configuration 3-Body a 1(0)=a 2(0)=a t(0)=2deg P1(0)=p2(0)=p(0)=1deg 5x(0)=45m 8y(0)=52(0)=05m dx(0)=dy(0)=d2(0)=1m Stationkeeping lt=10km
Variation of Kinetic and Potential Energy
(a) Time (Orbits) Percent Variation in Total Energy
OOEOh
-50E-12
-10E-11 I i i
0 1 2 (b) Time (Orbits)
Figure 3-2 Kinetic and potential energy transfer for the three-body platform based tethered satellite system (a) variation of kinetic and potential energy (b) percent change in total energy of system
48
STSS configuration 5-Body a 1(0)=a 2(0)=a t 1(0)=2deg a 3(0)=a t 2(0)=25deg P1(0)=p2(0)=p3(0)=pt1(0)=pt2(0)=1lt
5x(0)=5m 8y(0)=5z(0)=05m dx(0)=dy(0)=dz(0)=0 Stationkeeping 1 =12=1 Okm
(a)
20E6
10E6r-
00E0
-10E6
-20E6
Variation of Kinetic and Potential Energy
-
1 I I I 1 I I
I I
AP e ~_
i i 1 i i i i
00E0
UJ -50E-12 U J
lt
Time (Orbits) Percent Variation in Total Energy
-10E-11h
(b) Time (Orbits)
Figure 3 -3 Kinet ic and potential energy transfer for the five-body platform based tethered satellite system (a) variation of kinetic and potential enshyergy (b) percent change in total energy of system
49
in sequence by two tethers The two subsatellites have the same mass and inertia
ie 7713 = 7 7 i 2 a n d I3 = I21 a n d are separated by identical tether segments each 10
k m in length The platform has the same properties as before however there is no
tether attachment point offset present in the system (d = fcmi = 0)
In al l the cases considered energy was found to be conserved However it
should be noted that for the cases where the tether structural damping deployment
retrieval attitude and vibration control are present energy is no longer conserved
since al l these features add or subtract energy from the system
332 Comparison with available data
The alternative approach used to validate the numerical program involves
comparison of system responses with those presented in the literature A sample case
considered here is the planar system whose parameters are the same as those given
in Section 331 A n ini t ia l pitch disturbance of 2deg is imparted to both the platform
and the tether together with an ini t ial tether deformation of 08 m and 001 m in
the longitudinal and transverse directions respectively in the same manner as before
The tether length is held constant at 5 km and it is assumed to be connected to the
centre of mass of the rigid platform and subsatellite The response time histories from
Ref[43] are compared with those obtained using the present program in Figures 3-4
and 3-5 Here ap and at represent the platform and tether pitch respectively and B i
C i are the tethers longitudinal and transverse generalized coordinates respectively
A comparison of Figures 3-4 and 3-5 clearly shows a remarkable correlation
between the two sets of results in both the amplitude and frequency The same
trend persisted even with several other comparisons Thus a considerable level of
confidence is provided in the simulation program as a tool to explore the dynamics
and control of flexible multibody tethered systems
50
a p(0) = 2deg 0(0) = 2C
6(0) = 08 m
0(0) = 001 m
Stationkeeping L = 5 km
D p y = 0 D p 2 = 0
a oo
c -ooo -0 01
Figure 3 - 4 Simulation results for the platform based three-body tethered system originally presented in Ref[43]
51
STSS configuration 3-Body a1(0)=a2(0)=at(0)=2deg p1(0)=(32(0)=(3t(0)=0 8x(0)=08m 5y(0)=001m 5Z(0)=0 dx(0)=dy(0)=dz(0)=0 Stationkeeping l=5km
Satellite 1 Pitch Angle Tether Pitch Angle
i -i i- i i i i i i d
1 2 0 1 2 Time (Orbits) Time (Orbits)
Longitudinal Vibration
^MAAAAAAAWVWWWWVVVW
V bull i bull i I i J 00 05 10 15 20
Time (Orbits) Transverse Vibration
Time (Orbits)
Figure 3-5 Simulation results for the platform based three-body tethered system obtained using the present computer program
52
4 D Y N A M I C SIMULATION
41 Prel iminary Remarks
Understanding the dynamics of a system is critical to its design and develshy
opment for engineering application Due to obvious limitations imposed by flight
tests and simulation of environmental effects in ground based facilities space based
systems are routinely designed through the use of numerical models This Chapter
studies the dynamical response of two different tethered systems during deployment
retrieval and stationkeeping phases In the first case the Space platform based Tethshy
ered Satellite System (STSS) which consists of a large mass (platform) connected to
a relatively smaller mass (subsatellite) with a long flexible tether is considered The
other system is the O E D I P U S B I C E P S configuration involving two comparable mass
satellites interconnected by a flexible tether In addition the mission requirement of
spin about an arbitrary axis the tether axis for O E D I P U S - A C and cartwheeling or
spin about the orbit normal for the proposed B I C E P S mission is accounted for
A s mentioned in Chapter 2 the tether flexibility is modelled using the asshy
sumed mode discretization method Although the simulation program can account
for an arbitrary number of vibrational modes in each direction only the first mode
is considered in this study as it accounts for most of the strain energy[34] and hence
dominates the vibratory motion The parametric study considers single as well as
double pendulum type systems with offset of the tether attachment points The
systems stability is also discussed
42 Parameter and Response Variable Definitions
The system parameters used in the simulation unless otherwise stated are
53
selected as follows for the STSS configuration
1091430 -8135 328108 -8135 8646050 27116 328108 27116 8286760
h =
h = 200 0 0
0 400 0 0 0 400
bull m = 90000 kg (mass of the space platform)
kg bull m 2 (platform inertia) 760 J
kg bull m 2 (subsatellite inertia)
bull 7772 = 500 kg (mass of the subsatellite)
bull EtAt = 61645 N (tether stiffness)
bull Pt = 49 k g k m (tether density)
bull Vd mdash 05 (tether structural damping coefficient)
bull fcm^ = l 0 0 r m (tether attachment point at the subsatellite)
The response variables are defined as follows
bull ai0 satellite 1 (platform) pitch and roll angles respectively
bull CK2 2 satellite 2 (subsatellite) pitch and roll angles respectively
bull at fit- tether pitch and roll angles respectively
bull If tether length
bull d = dx dy dzT- tether attachment position relative to satellite 1 (platform)
mdash
bull S = 8X 8y Sz tether flexible generalized coordinates in the longitudinal x
inplane transverse y and out-of-plane transverse z directions respectively
The attitude angles and are measured with respect to the Local Ve r t i c a l -
Loca l Horizontal ( L V L H ) frame A schematic diagram illustrating these variable
is presented in Figure 4-1 The system is taken to be in a nominal circular orbit at
an altitude of 289 km with an orbital period of 903 minutes
54
Satellite 1 (Platform) Y a w
Figure 4-1 Schematic diagram showing the generalized coordinates used to deshyscribe the system dynamics
55
43 Stationkeeping Profile
To facilitate comparison of the simulation results a reference case based on
the three-body STSS system with zero offset (d = 0 r c m 3 = 0) and a tether length
of It = 20 km is considered first (Figure 4-2) The system is subjected to an ini t ial
disturbance in pitch and roll of 2deg and 1deg respectively to all the three bodies
In addition an ini t ia l longitudinal deflection of 14 m from the tethers unstretched
position is given together with a transverse inplane and out-of-plane deflection of 1
m at the tethers mid-length From the attitude response given in Figure 4-2(a) it is
apparent that the three bodies oscillate about their respective equilibrium positions
Since coupling between the individual rigid-body dynamics is absent (zero offset) the
corresponding librational frequencies are unaffected Satellite 1 (Platform) displays
amplitude modulation arising from the products of inertia The result is a slight
increase of amplitude in the pitch direction and a significantly larger decrease in
amplitude in the roll angle
Figure 4-2(b) clearly shows coupling of the tethers attitude motion with its
flexibility dynamics The tether vibrates in the axial direction about its equilibrium
position (at approximately 138 m) with two vibration frequencies namely 012 Hz
and 31 x 1 0 - 4 Hz The former is the first longitudinal flexible mode which is dissishy
pated through the structural damping while the latter arises from the coupling with
the pitch motion of the tether Similarly in the transverse direction there are two
frequencies of oscillations arising from the flexible mode and its coupling wi th the
attitude motion As expected there is no apparent dissipation in the transverse flexshy
ible motion This is attributed to the weak coupling between the longitudinal and
transverse modes of vibration since the transverse strain in only a second order effect
relative to the longitudinal strain where the dissipation mechanism is in effect Thus
there is very litt le transfer of energy between the transverse and longitudinal modes
resulting in a very long decay time for the transverse vibrations
56
STSS configuration 3-Body a1(0)=a2(0)=a(0)=2deg (31(0)=p2(0)=p(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=dz(0)=0 Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
0 1 2 3 4 5 0 1 2 3 4 5 Time (Orbits) Time (Orbits)
Satellite 2 Pitch Angle Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-2 Stationkeeping dynamics of the three-body STSS configuration withshyout offset (a) attitude response
57
STSS configuration 3-Body a1(0)=cx2(0)=cx(0)=2deg P1(0)=p2(0)=p(0)=1deg 5x(0)=14m8y(0)=82(0)=1m dx(0)=dy(0)=dz(0)=0 Stationkeeping lt=20km
Tether Longitudinal Vibration 16| 1 1 1 1 1 r-
Time (Orbits) Tether In-Plane Transverse Vibration
imdash 1 mdash 1 mdash 1 mdash 1 mdashr
r I I 1 1 1 I 1 1 1 1 I 1 bull ^ bull bull I I J 0 1 2 3 4 5
Time (Orbits) Tether Out-of-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-2 Stationkeeping dynamics of the three-body STSS configuration withshyout offset (b) vibration response
58
Introduction of the attachment point offset significantly alters the response of
the attitude motion for the rigid bodies W i t h a i m offset along the local vertical
at the platform (dx = 1 m) and fcm^ = 100 m coupling between the tether
and end-bodies is established by providing a lever arm from which the tether is able
to exert a torque that affects the rotational motion of the rigid end-bodies (Figure
4-3) Bo th the platform and subsatellite now oscillate at the same frequency as the
tether whose motion remains unaltered In the case of the platform there is also an
amplitude modulation due to its non-zero products of inertia as explained before
However the elastic vibration response of the tether remains essentially unaffected
by the coupling
Providing an offset along the local horizontal direction (dy = 1 m) results in
a more dramatic effect on the pitch and roll response of the platform as shown in
Figure 4-4(a) Now the platform oscillates about its new equilibrium position of
-90deg The roll motion is also significantly disturbed resulting in an increase to over
10deg in amplitude On the other hand the rigid body dynamics of the tether and the
subsatellite as well as the tether flexible motion remain the same (Figure 4-4b)
Figure 4-5 presents the response when the offset of 1 m at the platform end is
along the z direction ie normal to the orbital plane The result is a larger amplitude
pitch oscillation about the reference equilibrium position as shown in Figure 4-5(a)
while the roll equilibrium is now shifted to approximately 90deg Note there is little
change in the attitude motion of the tether and end-satellites however there is a
noticeable change in the out-of-plane transverse vibration of the tether (Figure 4-5b)
which may be due to the large amplitude platform roll dynamics
Final ly by setting a i m offset in all the three direction simultaneously the
equil ibrium position of the platform in the pitch and roll angle is altered by approxishy
mately 30deg (Figure 4-6a) However the response of the system follows essentially the
59
STSS configuration 3-Body a1(0)=cx2(0)=cxt(0)=2deg p1(0)=(32(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=1m dy(0)=dz(0)=0 Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
Time (Orbits) Time (Orbits) Tether Pitch Angle Tether Roll Angle
L bull I I bull bull lt bull bull bull raquo bull bull bull bull -I h I I I I i H 0 1 2 3 4 5 0 1 2 3 4 5
Time (Orbits) Time (Orbits) Satellite 2 Pitch Angle Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-3 Stationkeeping dynamics of the three-body STSS configuration with offset along the local vertical (a) attitude response
60
STSS configuration 3-Body a 1(0)=a 2(0)=a t(0)=2deg (31(0)=p2(0)=pt(0)=1deg 5x(0)=14m5y(0)=5z(0)=1m dx(0)=1mdy(0)=dz(0)=0 Stationkeeping lt=20km
Tether Longitudinal Vibration 1 6 i mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash mdash i mdash i mdash mdash mdash mdash mdash i mdash mdash lt -
Time (Orbits)
Tether In-Plane Transverse Vibration
y bull i i i i i i i i i i i i i i i i i 0 1 2 3 4 5
Time (Orbits)
Tether Out-of-Plane Transverse Vibration r 1 1 1 1 1 1 1 1 1 1 1 1 1 1 q
h i I i i I i i i i I i i i i 3
0 1 2 3 4 5 (b) Time (Orbits)
Figure 4-3 Stationkeeping dynamics of the three-body STSS configuration with offset along the local vertical (b) vibration response
61
STSS configuration 3-Body a1(0)=cx2(0)=o(0)=2deg (31(0)=p2(0)=p(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dz(0)=0 dy(0)=1m Stationkeeping lt=20km
Satellite 1 Pitch Angle i 1 i
Satellite 1 Roll Angle
1 2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 4 Time (Orbits)
Tether Roll Angle
CD
edT
1 2 3 Time (Orbits)
Satellite 2 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 2 Roll Angle
CD
C O
(a)
Figure 4-
1 2 3 4 Time (Orbits)
1 2 3 4 Time (Orbits)
4 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal (a) attitude response
62
STSS configuration 3-Body a1(0)=cx2(0)=ot(0)=2deg p1(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dz(0)=0 dy(0)=1m Stationkeeping lt=20km
Tether Longitudinal Vibration 16| 1 gt 1 bull 1 1 1
0 1 2 3 4 5
Time (Orbits) Tether In-Plane Transverse Vibration
i mdash 1 mdash 1 mdash 1 mdash 1 mdash i mdash mdash lt mdash 1 mdash 1 mdash i mdash mdash 1 mdash 1 mdash
V I I bull bull bull I i i i I i i 1 1 3
0 1 2 3 4 5 Time (Orbits)
Tether Out-of-Plane Transverse Vibration r 1 1 1 1 1 1 1 1 1 1 1 gt bull | q
V i 1 1 1 1 1 mdash 1 mdash 3
0 1 2 3 4 5 (b) Time (Orbits)
Figure 4-4 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal (b) vibration response
63
STSS configuration 3-Body a1(0)=o2(0)=at(0)=2deg (31(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=0 dz(0)=1m Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 4 Time (Orbits)
Tether Roll Angle
1 2 3 4 Time (Orbits)
Satellite 2 Pitch Angle
2 3 4 Time (Orbits)
Satellite 2 Roll Angle
(a)
Figure 4-
1 2 3 4 Time (Orbits)
2 3 4 Time (Orbits)
5 Stationkeeping dynamics of the three-body STSS configuration with offset along the orbit normal (a) attitude response
64
STSS configuration 3-Body a1(0)=cx2(0)=ot(0)=2deg P1(0)=p2(0)=(3(0)=1deg 8x(0)=14m8y(0)=82(0)=1m dx(0)=dy(0)=0 d2(0)=1m Stationkeeping lt=20km
Tether Longitudinal Vibration 1 6 i mdash r mdash i mdash mdash i mdash i mdash mdash i mdash i mdash mdash i mdash mdash i mdash 1 mdash 1 mdash i mdash mdash lt mdash lt mdash lt mdash r
Time (Orbits) Tether In-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-5 Stationkeeping dynamics of the three-body STSS configuration with offset along the orbit normal (b) vibration response
65
same trend with only minor perturbations in the transverse elastic vibratory response
of the tether (Figure 4-6b)
44 Tether Deployment
Deployment of the tether from an ini t ial length of 200 m to 20 km is explored
next Here the deployment length profile is critical to ensure the systems stability
It is not desirable to deploy the tether too quickly since that can render the tether
slack Hence the deployment strategy detailed in Section 216 is adopted The tether
is deployed over 35 orbits with the sinusoidal acceleration and deceleration occurring
over a 4 km length ( A ^ 2 ) with the remaining deployment at a constant velocity of
V0 = 146 ms
Initially the tether is stretched only slightly S = 1304 x 1 0 _ 3 0 0 m
with al l the other states of the system ie pitch roll and transverse displacements
remaining zero The response for the zero offset case is presented in Figure 4-7 For
the platform there is no longer any coupling with the tether hence it oscillates about
the equilibrium orientation determined by its inertia matrix However there is st i l l
coupling between the subsatellite and the tether since rcm^ mdash 100 m resulting
in complete domination of the subsatellite dynamics by the tether Consequently
the pitch motion ini t ial ly grows by almost 50deg but eventually subsides as the tether
deployment rate decreases It is interesting to note that the out-of-plane motion is not
affected by the deployment This is because the Coriolis force which is responsible
for the tether pitch motion does not have a component in the z direction (as it does
in the y direction) since it is always perpendicular to the orbital rotation (Q) and
the deployment rate (It) ie Q x It
Similarly there is an increase in the amplitude of the transverse elastic oscilshy
lations of the tether again due to the Coriolis effect (Figure 4-7b) Moreover as the
66
STSS configuration 3-Body a1(0)=a2(0)=cxt(0)=2deg P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m5y(0)=52(0)=1m dx(0)=dy(0)=d2(0)=1m Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
i i i i i i i i i i i 0 1 2 3 4 5 0 1 2 3 4 5
Time (Orbits) Time (Orbits) Tether Pitch Angle Tether Roll Angle
l i i i i -I h i i i i I i i i i l- i i i i l i i i i I i i- i i H 0 1 2 3 4 5 0 1 2 3 4 5
Time (Orbits) Time (Orbits) Satellite 2 Pitch Angle Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-6 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal local vertical and orbit normal (a) attitude response
67
STSS configuration 3-Body a1(0)=cc2(0)=a(0)=2deg p1(0)=p2(0)=(3t(0)=1deg 5x(0)=14m6y(0)=52(0)=1m dx(0)=dy(0)=d2(0)=1m Stationkeeping lt=20km
Tether Longitudinal Vibration 1 6 1 mdash i mdash i mdash i mdash mdash i mdash 1 mdash 1 mdash mdash mdash i mdash mdash 1 mdash 1 mdash 1 mdash i mdash 1 mdash 1 mdash 1 mdash 1 mdash r
0 1 2 3 4 5
Time (Orbits) Tether In-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-6 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal local vertical and orbit normal (b) vibration response
68
STSS configuration 3-Body a(0)=a2(0)=ot(0)=0 P1(0)=p2(0)=pt(0)=0 8X(0)=1304 x 103m 8y(0)=8z(0)=0 dx(0)=dy(0)=dz(0)=0 Deployment
l=02km to 20km in 35 Orbits
Tether Length Profile
E
Satellite 1 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 1 Roll Angle mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash [ mdash i mdash i mdash i mdash i mdash | mdash
1 2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 4 Time (Orbits)
Tether Roll Angle
CD
oo ax
i imdashimdashimdashImdashimdashimdashimdashr-
_ 1 _ L
1 2 3 4 Time (Orbits)
Satellite 2 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 2 Roll Angle
-50
(a) 1 2 3 4
Time (Orbits) 1 2 3 4
Time (Orbits)
Figure 4-7 Deployment dynamics of the three-body STSS configuration without offset (a) attitude response
69
STSS configuration 3-Body a1(0)=a2(0)=at(0)=0 p1(0)=p2(0)=(3t(0)=0 8X(0)=1304x 103m 8y(0)=52(0)=0 dx(0)=dy(0)=d2(0)=0 Deployment
l=02km to 20km in 35 Orbits
Tether Longitudinal Vibration
to
-02
(b)
Figure 4-7
2 3 Time (Orbits)
Tether In-Plane Transverse Vibration
0 1 1 bull 1 1
1 2 3 4 Time (Orbits)
Tether Out-of-Plane Transverse Vibration
5
1 1 1 1 1 i i | i i | ^
_ i i i i L -I
1 2 3 4 5 Time (Orbits)
Deployment dynamics of the three-body STSS configuration without offset (b) vibration response
70
tether elongates its longitudinal static equilibrium position also changes due to an
increase in the gravity-gradient tether tension It may be pointed out that as in the
case of attitude motion there are no out-of-plane tether vibrations induced
When the same deployment conditions are applied to the case of 1 m offset in
the x direction (Figure 4-8) coupling effects are re-introduced The platform pitch
exceeds -100deg during the peak deployment acceleration However as the tether pitch
motion subsides the platform pitch response returns to a smaller amplitude oscillation
about its nominal equilibrium On the other hand the platform roll motion grows
to over 20deg in amplitude in the terminal stages of deployment Longitudinal and
transverse vibrations of the tether remain virtually unchanged from the zero offset
case except for minute out-of-plane perturbations due to the platform librations in
roll
45 Tether Retrieval
The system response for the case of the tether retrieval from 20 km to 200 m
with an offset of 1 m in the x and z directions at the platform end is presented in
Figure 4-9 Here the Coriolis force induced during retrieval renders the tether unstable
(Figure 4-9a) Consequently the pitch motion of the platform is also destabilized
through coupling The z offset provides an additional moment arm that shifts the
roll equilibrium to 35deg however this degree of freedom is not destabilized From
Figure 4-9(b) it is apparent that the transverse mode of the tether vibration is also
disturbed producing a high frequency response in the final stages of retrieval This
disturbance is responsible for the slight increase in the roll motion for both the tether
and subsatellite
46 Five-Body Tethered System
A five-body chain link system is considered next To facilitate comparison with
71
STSS configuration 3-Body a(0)=a2(0)=at(0)=0 P1(0)=P2(0)=P(0)=0 5X(0)=1304 x 103m 8y(0)=8z(0)=0 dx(0)=1mdy(0)=d2(0)=0 Deployment
lt=02km to 20km in 35 Orbits
Tether Length Profile
E
Satellite 1 Pitch Angle
1 2 3 Time (Orbits)
Satellite 1 Roll Angle
bull bull bull i bull bull bull bull i
1 2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 Time (Orbits)
Tether Roll Angle
i i i i J I I i _
1 2 3 4 Time (Orbits)
Satellite 2 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 2 Roll Angle 3E-3h
D )
T J ^ 0 E 0 eg
CQ
-3E-3
(a) 1 2 3 4
Time (Orbits) 1 2 3 4
Time (Orbits)
Figure 4-8 Deployment dynamics of the three-body STSS configuration with offset along the local vertical (a) attitude response
7 2
STSS configuration 3-Body a1(0)=a2(0)=at(0)=0 p1(0)=P2(0)=(3t(0)=0 Sx(0)=1304 x 103m 5y(0)=5z(0)=0 dx(0)=1mdy(0)=dz(0)=0 Deployment
l=02km to 20km in 35 Orbits
Tether Longitudinal Vibration
Time (Orbits) Tether In-Plane Transverse Vibration
n 1 1 1 f
I I I _j I I I I I 1 1 1 1 1 2 3 4 5
Time (Orbits) Tether Out-of-Plane Transverse Vibration
J 1 I I 1 I I I I I I I I L _ _ I I d 1 2 3 4 5
Time (Orbits)
Deployment dynamics of the three-body STSS configuration with offset along the local vertical (b) vibration response
(b)
Figure 4-8
73
STSS configuration 3-Body a1(0)=a2(0)=at(0)=0 p(0)=P2(0)=(3t(0)=0 Sx(0)=14m8y(0)=82(0)=0 dy(0)=0dx(0)=d2(0)=1m Retrieval
l=20km to 02km in 35 Orbits
Satellite 1 Pitch Angle
(a) Time (Orbits)
Figure 4-9 Retrieval dynamics of the along the local vertical ar
Tether Length Profile
i i i i i i i i i i I 0 1 2 3
Time (Orbits) Satellite 1 Roll Angle
Time (Orbits) Tether Roll Angle
b_ i i I i I i i L _ J
0 1 2 3 Time (Orbits)
Satellite 2 Roll Angle _ mdash | r i i | 1 i i imdashmdashj i 1 r mdash i mdash
04- -
0 1 2 3 Time (Orbits)
three-body STSS configuration with offset d orbit normal (a) attitude response
74
STSS configuration 3-Body a1(0)=a2(0)=ocl(0)=0 (31(0)=f32(0)=pt(0)=0 8x(0)=14m8y(0)=6z(0)=0 dy(0)=0dx(0)=dz(0)=1m Retrieval
l=20km to 02km in 35 Orbits
Tether Longitudinal Vibration
E 10 to
Time (Orbits) Tether In-Plane Transverse Vibration
1 2 Time (Orbits)
Tether Out-of-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-9 Retrieval dynamics of the three-body STSS configuration with offset along the local vertical and orbit normal (b) vibration response
75
the three-body system dynamics studied earlier the chain is extended through the
addition of two bodies a tether and a subsatellite with the same physical properties
as before (Section 42) Thus the five-body system consists of a platform (satellite 1)
tether 1 subsatellite 1 (satellite 2) tether 2 and subsatellite 2 (satellite 3) as shown
in Figure 4-10 The offset of tether 1 at the platform-end is denoted by d2 while that
of tether 2 to subsatellite 1 as 04 In this numerical example fcm^ = fcm^ = 0 ie
tethers 1 and 2 are attached to the centre of mass of subsatellites 1 and 2 respectively
The system response for the case where the tether attachment points coincide
wi th the centres of mass of the rigid bodies ie zero offset (d2 = d mdash 0) is presented
in Figure 4-11 Here ogtt and at2 represent the pitch motion of tethers 1 and 2
respectively whereas 03 is the pitch motions of satellite 3 A similar convention
is adopted for the rol l angle 0 A s before the zero offset eliminates the coupling
between the tethers and satellites such that they are now free to librate about their
static equil ibrium positions However their is s t i l l mutual coupling between the two
tethers This coupling is present regardless of the offset position since each tether is
capable of transferring a force to the other The coupling is clearly evident in the pitch
response of the two tethers (Figure 4 - l l a ) O n the other hand the roll motion appears
uncoupled as the in i t ia l conditions in rol l for the two tethers are identical Thus the
motion is in phase and there is no transfer of energy through coupling A s expected
during the elastic response the tethers vibrate about their static equil ibrium positions
and mutually interact (Figure 4 - l l b ) However due to the variation of tension along
the tether (x direction) they do not have the same longitudinal equilibrium Note
relatively large amplitude transverse vibrations are present particularly for tether 1
suggesting strong coupling effects with the longitudinal oscillations
76
Figure 4-10 Schematic diagram of the five-body system used in the numerical example
77
STSS configuration 5-Body a1(0)=(x2(0)=at1(0)=2deg a3(0)=a12(0)=25deg p1(0)=r32(0)=p3(0)=pt1(0)=pt2(0)=r 82(0)=84(0)=505)05m d2(0)=d4(0)=000 Stationkeeping lt1=lu=10km Pitch Roll
Satellite 1 Pitch Angle Satellite 1 Roll Angle
o 1 2 Time (Orbits)
Tether 1 Pitch and Roll Angle
0 1 2 Time (Orbits)
Tether 2 Pitch and Roll Angle
Time (Orbits) Satellite 2 Pitch and Roll Angle
Time (Orbits) Satellite 3 Pitch and Roll Angle
Time (Orbits) Time (Orbits)
Figure 4-11 Stationkeeping dynamics of the five-body STSS configuration withshyout offset (a) attitude response
78
STSS configuration 5-Body a1(0)=a2(0)=oct1(0)=2deg a3(0)=at2(0)=25o
p1(0)=p2(0)=(33(0)=pt1(0)=pt2(0)=10
62(0)=54(0)=50505m d2(0)=d4(0)=000 Stationkeeping lt1=1 =1 Okm
Tether 1 Longitudinal Vibration Tether 2 Longitudinal Vibration i i 1 i 1 r 1 1 1 1
(b) Time (Orbits) Time (Orbits)
Figure 4-11 Stationkeeping dynamics of the five-body STSS configuration withshyout offset (b) vibration response
79
The system dynamics during deployment of both the tethers in the double-
pendulum configuration is presented in Figure 4-12 Deployment of each tether takes
place from 200 m to 20 km in 35 orbits The pitch motion of the system is similar
to that observed during the three-body case The Coriolis effect causes the tethers
to pitch through a large angle ( laquo 50deg ) which in turn disturbs the platform and
subsatellite due to the presence of offset along the local vertical On the other hand
the roll response for both the tethers is damped to zero as the tether deploys in
confirmation with the principle of conservation of angular momentum whereas the
platform response in roll increases to around 20deg Subsatellite 2 remains virtually
unaffected by the other links since the tether is attached to its centre of mass thus
eliminating coupling Finally the flexibility response of the two tethers presented in
Figure 4-12(b) shows similarity with the three-body case
47 BICEPS Configuration
The mission profile of the Bl-static Canadian Experiment on Plasmas in Space
(BICEPS) is simulated next It is presently under consideration by the Canadian
Space Agency To be launched by the Black Brant 2000500 rocket it would inshy
volve interesting maneuvers of the payload before it acquires the final operational
configuration As shown in Figure 1-3 at launch the payload (two satellites) with
an undeployed tether is spinning about the orbit normal (phase 1) The internal
dampers next change the motion to a flat spin (phase 2) which in turn provides
momentum for deployment (phase 3) When fully deployed the one kilometre long
tether will connect two identical satellites carrying instrumentation video cameras
and transmitters as payload
The system parameters for the BICEPS configuration are summarized below [59 0 0 1
bull I = I2 = 0 366 0 kg bull m 2 (satellite inertias) [ 0 0 392_
bull mi = 7772 = 200 kg (mass of the satellites)
80
STSS configuration 5-Body a(0)=cx2(0)=al1(0)=2deg a3(0)=c (0)=25o
P1(0)=P2(0)=p3(0)=f3t1(0)=Pt2(0)=r 52(0)=84(0)=1304x 10300m d2(0)=d4(0)=1G0m Deployment lt1=le=02km to 20km in 35 Orbits
Pitch Roll Satellite 1 Pitch Angle
Time (Orbits) Satellite 2 Pitch and Roll Angle T 1mdash1 1 1 I I I I | I I I 1 I | I I lt~r
Tether Length Profile i 1 1 1 1 i 1 1 1 1 i 1 1 1 1 i
i i i i i 0 1 2 3 4 5
Time (Orbits) Satellite 1 Roll Angle
0 1 2 3 4 5 Time (Orbits)
Tether 1 amp 2 Roll Angle L i | I I 1 I 1 1 -I
-Ih I- bull I I i i i i I i i i i 1
0 1 2 3 4 5 Time (Orbits)
Satellite 3 Pitch and Roll Angle i- i 1 1 i 1 1 1 1 i 1 1 1 1 i 1 1 1 1 i
L i i i i i i 1 i i mdash i mdash i I i imdashimdashLJ
0 1 2 3 4 5 Time (Orbits)
Figure 4-12 Deployment dynamics of the five-body S T S S configuration with offshyset along the local vertical (a) attitude response
81
STSS configuration 5-Body a1(0)=a2(0)=oct1(0)=2deg a3(0)=at2(0)=25deg P1(0)=P2(0)=P3(0)=Pt1(0)=pt2(0)=1o
82(0)=54(0)=1304 x 10-300m d2(0)=d4(0)=100m Deployment lt1=l2=02km to 20km in 35 Orbits
Tether 1 Longitudinal Vibration Tether 2 Longitudinal Vibration
Time (Orbits) Time (Orbits) Tether 1 Z Tranverse Vibration Tether 2 Z Tranverse Vibration
Figure 4-12 Deployment dynamics of the five-body STSS configuration with offshyset along the local vertical (b) vibration response
82
bull EtAt = 61645 N (tether stiffness)
bull pt = 30 k g k m (tether density)
bull It = 1 km (tether length)
bull rid = 1-0 (tether structural damping coefficient)
The system is in a circular orbit at a 289 km altitude Offset of the tether
attachment point to the satellites at both ends is taken to be 078 m in the x
direction The response of the system in the stationkeeping phase for a prescribed set
of in i t ia l conditions is given by Figure 4-13 As in the case of the STSS configuration
there is strong coupling between the tether and the satellites rigid body dynamics
causing the latter to follow the attitude of the tether However because of the smaller
inertias of the payloads there are noticeable high frequency modulations arising from
the tether flexibility Response of the tether in the elastic degrees of freedom is
presented in Figure 4-13(b) Note the longitudinal vibrations decay quite rapidly
(lt 05 orbits) due to the structural damping however it has vir tual ly no effect on
the transverse oscillations
The unique mission requirement of B I C E P S is the proposed use of its ini t ia l
angular momentum in the cartwheeling mode to aid in the deployment of the tethered
system The maneuver is considered next The response of the system during this
maneuver wi th an ini t ia l cartwheeling rate of 5 deg s is illustrated in Figure 4-14(a) The
in i t ia l tether length is taken to be 10 m and is deployed to 1 km There is an in i t ia l
increase in the pitch motion of the tether however due to the conservation of angular
momentum the cartwheeling rate decreases proportionally to the square of the change
in tether deployment rate The result is a rapid drop in the cartwheeling rate unti l
the system stops rotating entirely and simply oscillates about its new equilibrium
point Consequently through coupling the end-bodies follow the same trend The
roll response also subsides once the cartwheeling motion ceases However
83
BICEPS configuration 3-Body a1(0)=o2(0)=at(0)=2deg P1(0)=P2(0)=pt(0)=1deg 8X(0)=001 m 6y(0)=52(0)=01 m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping l=1 km
Satellite 1 Pitch Angle Satellite 1 Roll Angle - i 1 1 1 1 i - i q r 1 1 -gt
I- I i t i i i i I H 0 1 2 0 1 2
Time (Orbits) Time (Orbits) Tether Pitch Angle Tether Roll Angle
F mdash 1 mdash 1 mdash mdash 1 mdash i mdash 1 mdash 1 mdash mdash 1 mdash = i F mdash 1 mdash mdash mdash 1 mdash i mdash 1 mdash 1 mdash 1 mdash 1 mdash q
(a) Time (Orbits) Time (Orbits)
Figure 4-13 Stationkeeping dynamics of the three-body BICEPS configuration with offset along the local vertical (a) attitude response
84
BICEPS configuration 3-Body a1(0)=o2(0)=at(0)=2deg Pl(0)=p2(0)=(3t(0)=1deg 8X(0)=001 m 8y(0)=82(0)=01 m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping lt=1km
Tether Longitudinal Vibration 1E-2[
laquo = 5 E - 3 |
000 002
Time (Orbits) Tether In-Plane Transverse Vibration
Time (Orbits) Tether Out-of-Plane Transverse Vibration
Time (Orbits)
Figure 4-13 Stationkeeping dynamics of the three-body BICEPS configuration with offset along the local vertical (b) vibration response
85
BICEPS configuration 3-Body a1(0)=o2(0)=at(0)=2deg a1(0)=a2(0)=at(0)=57s Pl(0)=P2(0)=(3t(0)=1deg 8X(0)=115 x 103m 8y(0)=82(0)=0 dx(0)=078m dy(0)=d2(0)=0 Cartwheeling Deployment
lt=10m to 1km in 1 Orbit
Tether Length Profile
E
2000
cu T3
2000
co CD
2000
C D CD
T J
Satellite 1 Pitch Angle
1 2 Time (Orbits)
Satellite 1 Roll Angle
cn CD
33 cdl
Time (Orbits) Tether Pitch Angle
Time (Orbits) Tether Roll Angle
Time (Orbits) Satellite 2 Pitch Angle
1 2 Time (Orbits)
Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-14 Cartwheeling dynamics with deployment for the three-body BICEPS with offset along the local vertical (a) attitude response
86
BICEPS configuration 3-Body a1(0)=a2(0)=at(0)=2deg a (0)=a2(0)=a(0)=57s P1(0)=P2(0)=f3(0)=1deg 8X(0)=115x103m 8y(0)=8z(0)=0 dx(0)=078m dy(0)=d2(0)=0 Cartwheeling Deployment
l=1 Om to 1 km in 1 Orbit
5E-3h
co
0E0
025
000 E to
-025
002
-002
(b)
Figure 4-14
Tether Longitudinal Vibration
Time (Orbits) Tether In-Plane Transverse Vibration
Time (Orbits) Tether Out-of-Plane Transverse Vibration
Time (Orbits)
Cartwheeling dynamics with deployment for the three-body BICEPS with offset along the local vertical (b) vibration response
87
as the deployment maneuver is completed the satellites continue to exhibit a small
amplitude low frequency roll response with small period perturbations induced by
the tether elastic oscillations superposed on it (Figure 4-14b)
48 OEDIPUS Spinning Configuration
The mission entitled Observation of Electrified Distributions of Ionospheric
Plasmas - A Unique Strategy ( O E D I P U S ) also consists of a 1 k m tethered system with
two similar end-masses It was first flown in a suborbital flight in 1989 ( O E D I P U S A )
followed by a recent study in November 1995 ( O E D I P U S - C ) in which the University
of Br i t i sh Columbia was one of the participants Dynamically O E D I P U S represents a
unique system never encountered before Launched by a Black Brant rocket developed
at Br is to l Aerospace L t d the system ie the end-bodies together wi th the tether
spins about the longitudinal axis of the tether to achieve stabilized alignment wi th
Earth s magnetic field
The response of the system undergoing a spin rate of j = l deg s is presented
in Figure 4-15 using the same parameters as those outlined for the B I C E P S conshy
figuration in Section 47 Again there is strong coupling between the three bodies
(two satellites and tether) due to the nonzero offset The spin motion introduces an
additional frequency component in the tethers elastic response which is transferred
to the l ibrational motion of the satellites A s the spin rate increases to 1 0 deg s (Figure
4-16) the amplitude and frequency of the perturbations also increase however the
general character of the response remains essentially the same
88
OEDIPUS configuration 3-Body a1(0)=a2(0)=at(0)=2deg i(0)^[2(0H(0)=Ys p1(0)=p2(0)=p(0)=r 8X(0)=12 x 103m 8y(0)=82(0)=01m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping l=1 km
Satellite 1 Pitch Angle Satellite 1 Roll Angle mdash i 1 1 1 a r 1 1 1 r
(a) Time (Orbits) Time (Orbits)
Figure 4-15 Spin dynamics (7 = ldegs) of the three-body OEDIPUS configuration with offset along the local vertical (a) attitude response
89
OEDIPUS configuration 3-Body a1(0)=a2(0)=cct(0)=2deg Y1(0)=Y2(0)=Yt(0)=l7s (31(0)=p2(0)=pt(0)=1deg 8X(0)=12 x 103m 8y(0)=5z(0)=01 m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping lt=1 km
1E-2
to
OEO
E 00
to
Tether Longitudinal Vibration
005
Time (Orbits) Tether In-Plane Transverse Vibration
Time (Orbits) Tether Out-of-Plane Transverse Vibration
Time (Orbits)
Figure 4-15 Spin dynamics (7 = l deg s ) of the three-body O E D I P U S configuration with offset along the local vertical (b) vibration response
90
OEDIPUS configuration 3-Body a1(0)=a2(0)=a(0)=2deg Yi(0)=Y2(0)=rt(0)=107s 81(0)=p2(0)=pt(0)=1deg 8X(0)=12 x 103m 8y(0)=82(0)=01m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping lt=1km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-16 Spin dynamics (7= 10degs) of the three-body OEDIPUS configura tion with offset along the local vertical (a) attitude response
91
OEDIPUS configuration 3-Body a1(0)=a2(0)=ot(0)=2deg Yi(0)H(0)4(0)laquo107s P1(0)=P2(0)=pt(0)=1deg 8X(0)=12 x 103m 8y(0)=82(0)=01m dx(0)=078m dy(0)=dz(0)=0 Stationkeeping lt=1 km
Tether Longitudinal Vibration
Figure 4-16 Spin dynamics (7 = 10degs) of the three-body OEDIPUS configurashytion with offset along the local vertical (b) vibration response
92
5 ATTITUDE AND VIBRATION CONTROL
51 Att itude Control
511 Preliminary remarks
The instability in the pitch and roll motions during the retrieval of the tether
the large amplitude librations during its deployment and the marginal stability during
the stationkeeping phase suggest that some form of active control is necessary to
satisfy the mission requirements This section focuses on the design of an attitude
controller with the objective to regulate the librational dynamics of the system
As discussed in Chapter 1 a number of methods are available to accomplish
this objective This includes a wide variety of linear and nonlinear control strategies
which can be applied in conjuction with tension thrusters offset momentum-wheels
etc and their hybrid combinations One may consider a finite number of Linear T ime
Invariant (LTI) controllers scheduled discretely over different system configurations
ie gain scheduling[43] This would be relevant during deployment and retrieval of
the tether where the configuration is changing with time A n alternative may be
to use a Linear T ime Varying ( L T V ) model in conjuction with an adaptive control
scheme where on-line parametric identification may be used to advantage[56] The
options are vir tually limitless
O f course the choice of control algorithms is governed by several important
factors effective for time-varying configurations computationally efficient for realshy
time implementation and simple in character Here the thrustermomentum-wheel
system in conjuction with the Feedback Linearization Technique ( F L T ) is chosen as
it accounts for the complete nonlinear dynamics of the system and promises to have
good overall performance over a wide range of tether lengths It is well suited for
93
highly time-varying systems whose dynamics can be modelled accurately as in the
present case
The proposed control method utilizes the thrusters located on each rigid satelshy
lite excluding the first one (platform) to regulate the pitch and rol l motion of the
tether to which it is attached ie the thrusters located on satellite 2 (link 3) regshy
ulates the attitude motion of tether 1 (link 2) O n the other hand the l ibrational
motion of the rigid bodies (platform and subsatellite) is controlled using a set of three
momentum wheels placed mutually perpendicular to each other
The F L T method is based on the transformation of the nonlinear time-varying
governing equations into a L T I system using a nonlinear time-varying feedback Deshy
tailed mathematical background to the method and associated design procedure are
discussed by several authors[435758] The resulting L T I system can be regulated
using any of the numerous linear control algorithms available in the literature In the
present study a simple P D controller is adopted arid is found to have good perforshy
mance
The choice of an F L T control scheme satisfies one of the criteria mentioned
earlier namely valid over a wide range of tether lengths The question of compushy
tational efficiency of the method wi l l have to be addressed In order to implement
this controller in real-time the computation of the systems inverse dynamics must
be executed quickly Hence a simpler model that performs well is desirable Here
the model based on the rigid system with nonlinear equations of motion is chosen
Of course its validity is assessed using the original nonlinear system that accounts
for the tether flexibility
512 Controller design using Feedback Linearization Technique
The control model used is based on the rigid system with the governing equa-
94
tions of motion given by
Mrqr + fr = Quu (51)
where the left hand side represents inertia and other forces while the right hand side
is the generalized external force due to thrusters and momentum-wheels and is given
by Eq(2109)
Lett ing
~fr = fr~ QuU (52)
and substituting in Eq(296)
ir = s(frMrdc) (53)
Expanding Eq(53) it can be shown that
qr = S(fr Mrdc) - S(QU Mr6)u (5-4)
mdash = Fr - Qru
where S(QU Mr0) is the column matrix given by
S(QuMr6) = [s(Qu(l)Mr0) 5 (Q u ( 2 ) M r | 0 ) S(QU nu) M r | 0 ) ] (55)
and Qu-i) is the i ^ column of Qu Extract ing only the controlled equations from
Eq(54) ie the attitude equations one has
Ire = Frc ~~ Qrcu
(56)
mdash vrci
where vrc is the new control input required to regulate the decoupled linear system
A t this point a simple P D controller can be applied ie
Vrc = qrcd +KvQrcd-Qrc) + Kp(Qrcd~ Qrc)- (57)
Eq(57) represents the secondary controller with Eq(54) as the primary controller
Kp and Kv are the proportional and derivative gain matrices with qrcdi Qrcd
an(^ qrcd
95
as the desired acceleration velocity and position vector for the attitude angles of each
controlled body respectively Solving for u from Eq(56) leads to
u Qrc Frc ~ vrcj bull (58)
5 13 Simulation results
The F L T controller is implemented on the three-body STSS tethered system
The choice of proportional and derivative matrix gains is based on a desired settling
time of ts mdash 05 orbit and a damping factor of = 07 for both the pitch and roll
actuators in each body Given O and ts it can be shown that
-In ) 0 5 V T ^ r (59)
and hence kn bull mdash10
V (510)
where kp and kVj are the diagonal elements of the gain matrices Kp and Kv in
Eq(57) respectively[59]
Note as each body-fixed frame is referred directly to the inertial frame FQ
the nominal pitch equilibrium angle is zero only when measured from the L V L H frame
(OJJ mdash 0) However it is 6 when referred to FQ Hence the desired position vector
qrC(l is set equal to 0(t) ie the time-varying orbital angle for the pitch motion
and zero for the roll and yaw rotations such that
qrcd 37Vxl (511)
96
The desired velocity and acceleration are then given as
e U 3 N x l (512)
E K 3 7 V x l (513)
When the system is in a circular orbit qrc^ = 0
Figure 5-1 presents the controlled response of the STSS system defined in
Chapter 4 in the stationkeeping phase with offset d2 = 111T m and If = 20
km A s mentioned earlier the F L T controller is based on the nonlinear rigid model
Note the pitch and roll angles of the rigid bodies as well as of the tether are now
attenuated in less than 1 orbit (Figure 5-la) Consequently the longitudinal vibration
response of the tether is free of coupling arising from the librational modes of the
tether This leaves only the vibration modes which are eventually damped through
structural damping The pitch and roll dynamics of the platform require relatively
higher control moments Ma and Mpi respectively (Figure 5-lb) since they have to
overcome the extra moments generated by the offset of the tether attachment point
Furthermore the platform demands an additional moment to maintain its orientation
at zero pitch and roll angle since it is not the nominal equilibrium position The nonshy
zero static equilibrium of the platform arises due to its products of inertia On the
other hand the tether pitch and roll dynamics require only a small ini t ia l control
force of about plusmn 1 N which eventually diminishes to zero once the system is
97
and
o 0
Qrcj mdash
o o
0 0
Oiit) 0 0
1
N
N
1deg I o
STSS configuration 3-Body a1(0)=a2(0)=at(0)=2lt
P1(0)=p2(0)=pt(0)=1 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=dz(0)=1m Controlled Stationkeeping lt=20km
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
05 10 15 Time (Orbits)
Sat 2 Pitch and Roll Angles
20
CO CD
T J
cdT
cT
- mdash mdash 1 1 1 1 1 I - 1
Pitch -
A Roll -
A -
1 bull bull
-
1 Time (Orbits)
Tether Y Tranverse Vibration
1 2 Time (Orbits)
Tether Z Tranverse Vibration
gt to N
to
Time (Orbits) Time (Orbits) Longitudinal Vibration
145h
(a) Time (Orbits)
Figure 5-1 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local horizontal local vertical and orbit normal (a) attitude and vibration response
98
1
STSS configuration 3-Body a1(0)=ct2(0)=cct(0)=20
P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=d2(0)=1m Controlled Stationkeeping lt=20km
Sat 1 Pitch Control Moment Sat 1 Roll Control Moment
Figure 5-1 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local horizontal local vertical and orbit normal (b) control actuator time histories
99
stabilized Similarly for the end-body a very small control torque is required to
attenuate the pitch and roll response
If the complete nonlinear flexible dynamics model is used in the feedback loop
the response performance was found to be virtually identical with minor differences
as shown in Figure 5-2 For example now the pitch and roll response of the platform
is exactly damped to zero with no steady-state error as against the minor almost
imperceptible deviation for the case of the controller basedon the rigid model (Figure
5-la) In addition the rigid controller introduces additional damping in the transverse
mode of vibration where none is present when the full flexible controller is used
This is due to a high frequency component in the at motion that slowly decays the
transverse motion through coupling
As expected the full nonlinear flexible controller which now accounts the
elastic degrees of freedom in the model introduces larger fluctuations in the control
requirement for each actuator except at the subsatellite which is not coupled to the
tether since f c m 3 = 0 The high frequency variations in the pitch and roll control
moments at the platform are due to longitudinal oscillations of the tether and the
associated changes in the tension Despite neglecting the flexible terms the overall
controlled performance of the system remains quite good Hence the feedback of the
flexible motion is not considered in subsequent analysis
When the tether offset at the platform is restricted to only 1 m along the x
direction a similar response is obtained for the system However in this case the
steady-state error in the platforms rigid body motion is much smaller (Figure 5-3a)
In addition from Figure 5-3(b) the platforms control requirement is significantly
reduced
Figure 5-4 presents the controlled response of the STSS deploying a tether
100
STSS configuration 3-Body a1(0)=a2(0)=oct(0)=2deg p1(0)=P2(0)=p(0)=1deg 8x(0)=14m5y(0)=82(0)=1m dx(0)=dy(0)=dz(0)=1m Controlled Stationkeeping l=20km
Tether Pitch and Roll Angles
c n
T J
cdT
Sat 1 Pitch and Roll Angles 1 r
00 05 10 15 Time (Orbits)
Sat 2 Pitch and Roll Angles
c n CD
cdT a
o T J
CQ
- - I 1 1 1 I 1 1 1 1
P i t c h
R o l l
bull
1 -
0 1 2 Time (Orbits)
Tether Y Tranverse Vibration
0 1 2 Time (Orbits)
Tether Z Tranverse Vibration
E gt
10 N
CO
Time (Orbits) Time (Orbits) Longitudinal Vibration
145h
(a) Time (Orbits)
Figure 5-2 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear flexible FLT controller with offset along the local horizontal local vertical and orbit normal (a) attitude and vibration response
101
STSS configuration 3-Body a1(0)=a2(0)=cxt(0)=2deg 31(0)=p2(0)=p(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=d2(0)=1m Controlled Stationkeeping lt=20km
r i i i t i mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash mdash J 0 1 2 0 1 2
Time (Orbits) Time (Orbits) Tether Pitch Control Thrust Tether Roll Control Thruster
(b) Time (Orbits) Time (Orbits)
Figure 5-2 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear flexible FLT controller with offset along the local horizontal local vertical and orbit normal (b) control actuator time histories
102
STSS configuration 3-Body a1(0)=o2(0)=a(0)=2deg P1(0)=p2(0)=(3t(0)=1deg 8x(0)=14m5y(0)=6z(0)=1m dx(0)=1mdy(0)=dz(0)=0 Controlled Stationkeeping lt=20km
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
05 10 15 Time (Orbits)
Sat 2 Pitch and Roll Angles
CD
eg CQ
0 1 2 Time (Orbits)
Tether Y Tranverse Vibration
1 2 Time (Orbits)
Tether Z Tranverse Vibration
gt CO
Time (Orbits) Time (Orbits) Longitudinal Vibration
145h
(a) Time (Orbits)
Figure 5-3 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local vertical (a) attitude and vibration response
103
STSS configuration 3-Body a1(0)=o2(0)=al(0)=2o
P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=1mdy(0)=dz(0)=0 Controlled Stationkeeping lt=20km
Sat 1 Pitch Control Moment 11 1 1 r
Sat 1 Roll Control Moment
0 1 2 Time (Orbits)
Tether Pitch Control Thrust
0 1 2 Time (Orbits)
Sat 2 Pitch Control Moment
0 1 2 Time (Orbits)
Tether Roll Control Thruster
-07 o 1
Time (Orbits Sat 2 Roll Control Moment
1E-5
0E0F
Time (Orbits)
Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local vertical (b) control actuator time histories
104
STSS configuration 3-Body a1(0)=a2(0)=a(0)=2deg Pl(0)=p2(0)=pt(0)=1deg 6X(0)=1304 x 103m 8y(0)=82(0)=0 dx(0)=1mdy(0)=d2(0)=0 Controlled Deployment
l=02km to 20km in 35 Orbits
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
o 1 2 Time (Orbits)
Tether Y Tranverse Vibration
0 1 2 Time (Orbits)
Sat 2 Pitch and Roll Angles
CD CD
CM CO
Time (Orbits) Tether Z Tranverse Vibration
to
150
100
1 2 3 0 1 2
Time (Orbits) Time (Orbits) Longitudinal Vibration
to
(a)
50
2 3 Time (Orbits)
Figure 5 - 4 Deployment dynamics of the three-body STSS using the nonlinear rigid FLT controller with offset along the local vertical (a) attitude and vibration response
105
STSS configuration 3-Body a1(0)=oc2(0)=al(0)=20
P1(0)=p2(0)=pt(0)=1deg 8X(0)=1304 x 103m 5y(0)=52(0)=0 dx(0)=1mdy(0)=dz(0)=0 Controlled Deployment
l=02km to 20km in 35 Orbits
Tether Length Profile i 1
2 0 h
E 2L
Sat 1 Pitch Control Moment t mdash r mdash i mdash i mdash | mdash i mdash i mdash i mdash i mdash | mdash r mdash i mdash i mdash i mdash | mdash r mdash
0 5
1 2 3
Time (Orbits) Sat 1 Roll Control Moment
i i t | t mdash t mdash i mdash [ i i mdash r -
1 2
Time (Orbits) Tether Pitch Control Thrust
1 2 3 Time (Orbits)
Tether Roll Control Thruster
2 E - 5
1 2 3
Time (Orbits) Sat 2 Pitch Control Moment
1 2
Time (Orbits) Sat 2 Roll Control Moment
1 E - 5
E
O E O
Time (Orbits)
Deployment dynamics of the three-body STSS using the nonlinear rigid FLT controller with offset along the local vertical (b) control actuator time histories
106
from 02 km to 20 km in 35 orbits with a i m offset along the tether length A s before
the systems attitude motion is well regulated by the controller However the control
cost increases significantly for the pitch motion of the platform and tether due to the
Coriolis force induced by the deployment maneuver (Figure 5-4b) Initially there is a
sinusoidal increase in the pitch control requirement for both the tether and platform
as the tether accelerates to its constant velocity VQ Then the control requirement
remains constant for the platform at about 60 N m as opposed to the tether where the
thrust demand increases linearly Finally when the tether decelerates the actuators
control input reduces sinusoidally back to zero The rest of the control inputs remain
essentially the same as those in the stationkeeping case In fact the tether pitch
control moment is significantly less since the tether is short during the ini t ia l stages
of control However the inplane thruster requirement Tat acts as a disturbance and
causes 6y to nearly double from the uncontrolled value (Figure 4-8a) O n the other
hand the tether has no out-of-plane deflection
Final ly the effectiveness of the F L T controller during the crit ical maneuver of
retrieval from 20 km to 02 k m in 35 orbits is assessed in Figure 5-5 A s in the earlier
cases the system response in pitch and roll is acceptable however the controller is
unable to suppress the high frequency elastic oscillations induced in the tether by the
retrieval O f course this is expected as there is no active control of the elastic degrees
of freedom However the control of the tether vibrations is discussed in detail in
Section 52 The pitch control requirement follows a similar trend in magnitude as
in the case of deployment with only minor differences due to the addition of offset in
the out-of-plane direction ( d 2 = 1 0 1 ^ m) This is also responsible for the higher
roll moment requirement for the platform control (Figure 5-5b)
107
STSS configuration 3-Body a1(0)=X2(0)=a(0)=2deg P1(0)=p2(0)=pt(0)=1deg 5x(0)=14m8y(0)=52(0)=0 dx(0)=1mdy(0)=dz(0)=0 Controlled Retrieval
l=20km to 02km in 35 Orbits
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
0 1 2 Time (Orbits)
Sat 2 Pitch and Roll Angles
CO CD
cdT
CO CD
T J CM
CO
1 2 Time (Orbits)
Tether Y Tranverse Vibration
Time (Orbits) Tether Z Tranverse Vibration
150
100
1 2 3 0 1 2 Time (Orbits) Time (Orbits)
Longitudinal Vibration
CO
(a)
x 50
2 3 Time (Orbits)
Figure 5 -5 Retrieval dynamics of the three-body STSS using the non-linear rigid FLT controller with offset along the local vertical and orbit normal (a) attitude and vibration response
108
STSS configuration 3-Body a1(0)=ct2(0)=at(0)=2deg p1(0)=p2(0)=(3(0)=1deg 5x(0)=14m5y(0)=82(0)=0 dx(0)=1mdy(0)=dz(0)=0 Controlled Retrieval
l=20km to 02km in 35 Orbits
Tether Length Profile
E
1 2 Time (Orbits)
Tether Pitch Control Thrust
2E-5
1 2 Time (Orbits)
Sat 2 Pitch Control Moment
OEOft
1 2 3 Time (Orbits)
Sat 1 Roll Control Moment
1 2 3 Time (Orbits)
Tether Roll Control Thruster
1 2 Time (Orbits)
Sat 2 Roll Control Moment
1E-5 E z
2
0E0
(b) Time (Orbits) Time (Orbits)
Figure 5-5 Retrieval dynamics of the three-body STSS using the non-linear rigid FLT controller with offset along the local vertical and orbit normal (b) control actuator time histories
109
52 Control of Tethers Elastic Vibrations
521 Prel iminary remarks
The requirement of a precisely controlled micro-gravity environment as well
as the accurate release of satellites into their final orbit demands additional control
of the tethers vibratory motion in addition to its attitude regulation To that end
an active vibration suppression strategy is designed and implemented in this section
The strategy adopted is based on offset control ie time dependent variation
of the tethers attachment point at the platform (satellite 1) A l l the three degrees
of freedom of the offset motion are used to control both the longitudinal as well as
the inplane and out-of-plane transverse modes of vibration In practice the offset
controller can be implemented through the motion of a dedicated manipulator or a
robotic arm supporting the tether which in turn is supported by the platform The
focus here is on the control of elastic deformations during stationkeeping (ie fully
deployed tether fixed length) as it represents the phase when the mission objectives
are carried out
This section begins with the linearization of the systems equations of motion
for the reduced three body stationkeeping case This is followed by the design of
the optimal control algorithm Linear Quadratic Gaussian-Loop Transfer Recovery
( L Q G L T R ) based on the reduced model and its implementation on the full nonlinear
model Final ly the systems response in the presence of offset control is presented
which tends to substantiate its effectiveness
522 System linearization and state-space realization
The design of the offset controller begins with the linearization of the equations
of motion about the systems equilibrium position Linearization of extremely lengthy
(even in matrix form) highly nonlinear nonautonomous and coupled equations of
110
motion presents a challenging problem This is further complicated by the fact the
pitch angle ot is not referred to the L V L H frame Hence the equilibrium pitch angle
is not a constant but is equal to the instantaneous orbital angle 9(t) Two methods
are available to resolve this problem
One may use the non-stationary equations of motion in their present form and
derive a controller based on the Linear T ime Varying ( L T V ) system Alternatively
one can design a Linear T ime Invariant (LTI) controller based on a new set of reduced
governing equations representing the motion of the three body tethered system A l shy
though several studies pertaining to L T V systems have been reported[5660] the latter
approach is chosen because of its simplicity Furthermore the L T I controller design
is carried out completely off-line and thus the procedure is computationally more
efficient
The reduced model is derived using the Lagrangian approach with the genershy
alized coordinates given by
where lty and are the pitch and roll angles of link i relative to the L V L H frame
respectively 5X 8y and 5Z are the generalized coordinates associated with the flexible
tether deformations in the longitudinal inplane and out-of-plane transverse direcshy
tions respectively Only the first mode of vibration is considered in the analysis
The nonlinear nonautonomous and coupled equations of motion for the tethshy
ered system can now be written as
Qred = [ltXlPlltX2P2fixSy5za302gt] (514)
M redQred + fr mdash 0gt (515)
where Mred and frec[ are the reduced mass matrix and forcing term of the system
111
respectively They are functions of qred and qred in addition to the time varying offset
position d2 and its time derivatives d2 and d2 A detailed derivation of Eq(515)
is given in Appendix II A n additional consequence of referring the pitch motion to
a local frame is that the new reduced equations are now independent of the orbital
angle 9 and under the further assumption of the system negotiating a circular orbit
91 remains constant
The nonlinear reduced equations can now be linearized about their static equishy
l ibr ium position For all the generalized coordinates the equilibrium position is zero
wi th the exception of 5X which has a non-zero equilibrium 5XQ The offset position
d2) is linearized about its ini t ial position d2^ Linearizing Eq(515) and recasting
into matr ix form gives
Msqred + CsQred + KsQred + Md^2 + Cdd2 + Kdd2 + fs = 0 (516)
where Ms Cs Ks Md Cd Kd and s are constant matrices Mul t ip ly ing throughout
by Ms
1 gives
qr = -Ms 1Csqr - Ms
lKsqr
- Ms~lCdd22 - Ms-Kdd2 - Ms~Mdd2 - Mg~1 fs r-1 (517)
Defining ud = d2 and v = qTd^T Eq(517) can be rewritten as
Pu t t ing
v =
+
-M~lCs -M~Cd -1
0
-Mg~lMd
Id
0 v + -M~lKs -M~lKd
0 0
-M-lfs
0
MCv + MKv + Mlud +Fs
- 4 ) - ( gt the L T I equations of motion can be recast into the state-space form as
(518)
(519)
x =Ax + Bud + Fd (520)
112
where
A = MC MK Id12 o
24x24
and
Let
B = ^ U 5 R 2 4 X 3 -
Fd = I FJ ) e f t 2 4 1
mdash _ i _ -5
(521)
(522)
(523)
(524)
where x is the perturbation vector from the constant equilibrium state vector xeq
Substituting Eq(524) into Eq(520) gives the linear perturbation state equation
modelling the reduced tethered system as
bullJ x mdash x mdash Ax + Bud + (Fd + Axeq) (525)
It can be shown that for ideal control[60] Fd + Axeq = 0 thus giving the familiar
state-space equation
S=Ax + Bud (526)
The selection of the output vector completes the state space realization of
the system The output vector consists of the longitudinal deformation from the
equil ibrium position SXQ of the tether at xt = h the slope of the tether due to the
transverse deformation at xt = 0 and the offset position d2 from its in i t ia l position
d2Q Thus the output vector y is given by
K(o)sx
V
dx ~ dXQ dy ~ dyQ
d z - d z0 J
Jy
dx da XQ vo
d z - dZQ ) dy dyQ
(527)
113
where lt ( 0 ) = ddxi[$vxi)]Xi=o and lt ( 0 ) = ddxi[4gtw(xi)]Xi=0
523 Linear Quadratic Gaussian control with Loop Transfer Recovery
W i t h the linear state space model defined (Eqs526527) the design of the
controller can commence The algorithm chosen is the Linear Quadratic Gaussian
( L Q G ) estimator based optimal controller[6061] The L Q G is a widely used optimal
controller with its theoretical background well developped by many authors over the
last 25 years It involves of the design of the Kalman-Bucy Fi l ter ( K B F ) which
provides an estimate of the states x and a Linear Quadratic Regulator ( L Q R ) which
is separately designed assuming all the states x are known Both the L Q R and K B F
designs independently have good robustness properties ie retain good performance
when disturbances due to model uncertainty are included However the combined
L Q R and K B F designs ie the L Q G design may have poor stability margins in the
presence of model uncertainties This l imitat ion has led to the development of an L Q G
design procedure that improves the performance of the compensator by recovering
the full state feedback robustness properties at the plant input or output (Figure 5-6)
This procedure is known as the Linear Quadratic Guassian-Loop Transfer Recovery
( L Q G L T R ) control algorithm A detailed development of its theory is given in
Ref[61] and hence is not repeated here for conciseness
The design of the L Q G L T R controller involves the repeated solution of comshy
plicated matrix equations too tedious to be executed by hand Fortunenately the enshy
tire algorithm is available in the Robust Control Toolbox[62] of the popular software
package M A T L A B The input matrices required for the function are the following
(i) state space matrices ABC and D (D = 0 for this system)
(ii) state and measurement noise covariance matrices E and 0 respectively
(iii) state and input weighting matrix Q and R respectively
114
u
MODEL UNCERTAINTY
SYSTEM PLANT
y
u LQR CONTROLLER
X KBF ESTIMATOR
u
y
LQG COMPENSATOR
PLANT
X = f v ( X t )
COMPENSATOR
x = A k x - B k y
ud
= C k f
y
Figure 5-6 Block diagram for the LQGLTR estimator based compensator
115
A s mentionned earlier the main objective of this offset controller is to regu-mdash
late the tether vibration described by the 5 equations However from the response
of the uncontrolled system presented in Chapter 4 it is clear that there is a large
difference between the magnitude of the librational and vibrational frequencies This
separation of frequencies allow for the separate design of the vibration and attitude mdash mdash
controllers Thus only the flexible subsystem composed of the S and d equations is
required in the offset controller design Similarly there is also a wide separation of
frequencies between the longitudinal and transverse modes of vibrations permitt ing mdash
the decoupling of the flexible subsystem into a set of longitudinal (Sx and dx) and
transverse (5y Sz dy and dz) subsystems The appended offset system d must also
be included since it acts as the control actuator However it is important to note
that the inplane and out-of-plane transverse modes can not be decoupled because
their oscillation frequencies are of the same order The two flexible subsystems are
summarized below
(i) Longitudinal Vibra t ion Subsystem
This is defined by u mdash Auxu + Buudu
(528)
where xu = SxdX25xdX2T u d u = dX2 and from Eq(527)
0 0 $u(lt) 0 0 0 0 1
0 0 1 0 0 0 0 1
(529)
Au and Bu are the rows and columns of A and B respectively and correspond to the
components of xu
The L Q R state weighting matrix Qu is taken as
Qu =
bull1 0 0 o -0 1 0 0 0 0 1 0
0 0 0 10
(530)
116
and the input weighting matrix Ru = 1^ The state noise covariance matr ix is
selected as 1 0 0 0 0 1 0 0 0 0 4 0
LO 0 0 1
while the measurement noise covariance matrix is taken as
(531)
i o 0 15
(532)
Given the above mentionned matrices the design of the L Q G L T R compenshy
sator is computed using M A T L A B with the following 2 commands
(i) kfu = l q r c ( ^ ( i C bdquo d i a g m x ( E u e u ) )
(ii) [AkugtBkugtCkugtDku =^ryAuBuCuDukfuQuRur)
where r is a scaler
The first command is used to compute the Ka lman filter gain matrix kfu Once
kfu is known the second command is invoked returning the state space representation
of the L Q G L T R dynamic compensator to
(533) xu mdash Akuxu Bkuyu
where xu is the state estimate vector oixu The function l t ry performs loop transfer
recovery at the system output ie the return ratio at the output approaches that
of the K B F loop given by Cu(sld - Au)Kfu[61 This is achieved by choosing a
sufficently large scaler value r in l t ry such that the singular values of the return
ratio approach those of the target design For the longitudinal controller design
r mdash 5 x 10 5 Figure 5-7(a) compares the singular values of the recovered compensator
design and the non-recovered L Q G compensator design with respect to the target
Unfortunetaly perfect recovery is not possible especially at higher frequencies
117
100
co
gt CO
-50
-100 10
Longitudinal Compensator Singular Values i I I I N I I mdash i mdash i i i i n i i mdash i mdash i i i i inimdash imdashi i i i i i i i mdash i mdash i i M i I I
Target r = 5x10 5
r = 0 (LQG)
i i i i I I i ii i i i 1 1 1 1 i i i i i 1 1 1 1 ii i I I i i i i 11111 -3 10 -2
(a)
101 10deg Frequency (rads)
101 10
Transverse Compensator Singular Values
100
50
X3
gt 0
CO
-50
bull100 10
-imdashi i 111II i 111 ni 1mdashi I I I I I I I 1mdashi I I I N I I 1mdashi i M 111
Target r = 50 r = 0 (LQG)
m i i i i i i n i i i i I I I I I I I 1mdashi I I I I I I I 1mdashi i 111 n -3 10 -2
(b)
101 10deg Frequency (rads)
101 10
Figure 5-7 Singular values for the LQG and LQGLTR compensator compared to target return ratio (a) longitudinal design (b) transverse design
118
because the system is non-minimal ie it has transmission zeros with positive real
parts[63]
(ii) Transverse Vibra t ion Subsystem
Here Xy mdash AyXy + ByU(lv
Dv = Cvxv
with xv = 6y 8Z dV2 dZ2 Sy 6Z dV2 dZ2T] udv = dy2dZ2T and
(534)
Cy
0 0 0 0 (0) 0 0 0 0 0 0 0 0 0 0 0 0
0 0 0 0
0 0 0 0 0 0 0 0 0 0 0 0
0 0 0
h 0 0 0
0 0 0 lt ( 0 ) o 0
0 1 0 0 0 1
0 0 0
2 r
h 0 0
0 0 1 0 0 1
(535)
Av and Bv are the rows and columns of the A and B corresponding to the components
O f Xy
The state L Q R weighting matrix Qv is given as
Qv
10 0 0 0 0 0 0 0 0 10 0 0 0 0 0 0 0 0 10 4 0 0 0 0 0 0 0 0 10 4 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1
(536)
mdash
Ry =w The state noise covariance matrix is taken
1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 9 x 10 4 0
0 0 0 0 0 0 0 9 x 10 4
(537)
119
while the measurement noise covariance matrix is
1 0 0 0 0 1 0 0 0 0 5 x 10 4 0
LO 0 0 5 x 10 4
(538)
As in the development of the longitudinal controller the transverse offset conshy
troller can be designed using the same two commands (lqrc and l try) wi th the input
matrices corresponding to the transverse subsystem with r = 50 The result is the
transverse compensator system given by
(539)
where xv is the state estimates of xv The singular values of the target transfer
function is compared with those achieved by the L Q G and L Q G L T R in Figure 5-
7(b) It is apparent that the recovery is not as good as that for the longitudinal case
again due to the non-minimal system Moreover the low value of r indicates that
only a lit t le recovery in the transverse subsystem is possible This suggests that the
robustness of the L Q G design is almost the maximum that can be achieved under
these conditions
Denoting f y = f j f j r xf = x^x^T and ud = udu ^ T the longishy
tudinal compensator can be combined to give
xf =
$d =
Aku
0
0 Akv
Cku o
Cu 0
Bku
0 B Xf
Xf = Ckxf
0
kv
Hf = Akxf - Bkyf
(540)
Vu Vv 0 Cv
Xf = CfXf
Defining a permutation matrix Pf such that Xj = PfX where X is the state vector
of the original nonlinear system given in Eq(32) the compensator and full nonlinear
120
system equations can be combined as
X
(541)
Ckxf
A block diagram representation of Eq(541) is shown in Figure 5-6
524 Simulation results
The dynamical response for the stationkeeping STSS with a tether length of 20
km and ini t ia l offset position of 1 m in each direction is presented in Figure 5-8 when
both the F L T attitude controller and the L Q G L T R offset controller are activated As
expected the offset controller is successful in quickly damping the elastic vibrations
in the longitudinal and transverse direction (Figure 5-8b) However from Figure 5-
8(a) it is clear that the presence of offset control requires a larger control moment
to regulate the attitude of the platform This is due to the additional torque created
by the larger moment arm around 2 m and 125 m in the inplane and out-of-plane
directions respectively introduced by the offset control In addition the control
moment for the platform is modulated by the tethers transverse vibration through
offset coupling However this does not significantly affect the l ibrational motion of
the tether whose thruster force remains relatively unchanged from the uncontrolled
vibration case
Final ly it can be concluded that the tether elastic vibration suppression
though offset control in conjunction with thruster and momentum-wheel attitude
control presents a viable strategy for regulating the dynamics tethered satellite sysshy
tems
121
STSS configuration 3-Body a1(0)=a2(0)=at(0)=2deg P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m6y(0)=52(0)=1m dx(0)=dy(0)=dz(0)=1m Stationkeeping I = 20km LQGLTR Vibration Control
Sat 1 Pitch and Roll Angles Tether Pitch and Roll Angles mdash 1 1 1 1 1 1 I 1 1 lt 1 ~ ^ 1 ^ l _
Time (Orbits) Time (Orbits) Sat 1 Roll Control Moment Tether Roll Control Thrust
(a) Time (Orbits) Time (Orbits)
Figure 5-8 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT attitude controller and LQGLTR offset vibration controller (a) attitude and libration controller reshysponse
122
STSS configuration 3-Body a1(0)=a2(0)=a(0)=2deg p1(0)=(32(0)=Q(0)=1deg ox(0)=14mSy(0)=5z(0)=1m dx(0)=dy(0)=dz(0)=1m Stationkeeping lt = 20km LQGLTR Vibration Control
Tether X Offset Position
Time (Orbits) Tether Y Tranverse Vibration 20r
Time (Orbits) Tether Y Offset Position
t o
1 2 Time (Orbits)
Tether Z Tranverse Vibration Time (Orbits)
Tether Z Offset Position
(b)
Figure 5-8
Time (Orbits) Time (Orbits)
Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT attitude controller and LQGLTR offset vibration controller (b) vibration and offset response
123
6 C O N C L U D I N G R E M A R K S
61 Summary of Results
The thesis has developed a rather general dynamics formulation for a multi-
body tethered system undergoing three-dimensional motion The system is composed
of multiple rigid bodies connected in a chain configuration by long flexible tethers
The tethers which are free to undergo libration as well as elastic vibrations in three
dimensions are also capable of deployment retrieval and constant length stationkeepshy
ing modes of operation Two types of actuators are located on the rigid satellites
thrusters and momentum-wheels
The governing equations of motion are developed using a new Order(N) apshy
proach that factorizes the mass matrix of the system such that it can be inverted
efficiently The derivation of the differential equations is generalized to account for
an arbitrary number of rigid bodies The equations were then coded in FORTRAN
for their numerical integration with the aid of a symbolic manipulation package that
algebraically evaluated the integrals involving the modes shapes functions used to dis-
cretize the flexible tether motion The simulation program was then used to assess the
uncontrolled dynamical behaviour of the system under the influence of several system
parameters including offset at the tether attachment point stationkeeping deployshy
ment and retrieval of the tether The study covered the three-body and five-body
geometries recently flown OEDIPUS system and the proposed BICEPS configurashy
tion It represents innovation at every phase a general three dimensional formulation
for multibody tethered systems an order-N algorithm for efficient computation and
application to systems of contemporary interest
Two types of controllers one for the attitude motion and the other for the
124
flexible vibratory motion are developed using the thrusters and momentum-wheels for
the former and the variable offset position for the latter These controllers are used to
regulate the motion of the system under various disturbances The attitude controller
is developed using the nonlinear Feedback Linearization Technique ( F L T ) and is based
on the nonlinear but rigid model of the system O n the other hand the separation
of the longitudinal and transverse frequencies from those of the attitude response
allows for the development of a linear optimal offset controller using the robust Linear
Quadratic Gaussian-Loop Transfer Recovery ( L Q G L T R ) method The effectiveness
of the two controllers is assessed through their subsequent application to the original
nonlinear flexible model
More important original contributions of the thesis which have not been reshy
ported in the literature include the following
(i) the model accounts for the motion of a multibody chain-type system undershy
going librational and in the case of tethers elastic vibrational motion in all
the three directions
(ii) the dynamics formulation is based on a recursive O(N) Lagrangian algorithm
that efficiently computes the systems generalized acceleration vector
(iii) the development of an attitude controller based on the Feedback Linearizashy
tion Technique for a multibody system using thrusters and momentum-wheels
located on each rigid body
(iv) the design of a three degree of freedom offset controller to regulate elastic v i shy
brations of the tether using the Linear Quadratic Gaussian and Loop Transfer
Recovery method ( L Q G L T R )
(v) substantiation of the formulation and control strategies through the applicashy
tion to a wide variety of systems thus demonstrating its versatility
125
The emphasis throughout has been on the development of a methodology to
study a large class of tethered systems efficiently It was not intended to compile
an extensive amount of data concerning the dynamical behaviour through a planned
variation of system parameters O f course a designer can easily employ the user-
friendly program to acquire such information Rather the objective was to establish
trends based on the parameters which are likely to have more significant effect on
the system dynamics both uncontrolled as well as controlled Based on the results
obtained the following general remarks can be made
(i) The presence of the platform products of inertia modulates the attitude reshy
sponse of the platform and gives rise to non-zero equilibrium pitch and roll
angles
(ii) Offset of the tether attachment point significantly affects the equil ibrium orishy
entation of the system as well as its dynamics through coupling W i t h a relshy
atively small subsatellite the tether dynamics dominate the system response
with high frequency elastic vibrations modulating the librational motion On
the other hand the effect of the platform dynamics on the tether response is
negligible As can be expected the platform dynamics is significantly affected
by the offset along the local horizontal and the orbit normal
(iii) For a three-body system deployment of the tether can destabilize the platform
in pitch as in the case of nonzero offset However the roll motion remains
undisturbed by the Coriolis force Moreover deployment can also render the
tether slack if it proceeds beyond a critical speed
(iv) Uncontrolled retrieval of the tether is always unstable in pitch It also leads
to high frequency elastic vibrations in the tether
(v) The five-body tethered system exhibits dynamical characteristics similar to
those observed for the three-body case As can be expected now there are
additional coupling effects due to two extra bodies a rigid subsatellite and a
126
flexible tether
(vi) Cartwheeling motion of the B I C E P S configuration can also be used to adshy
vantage in the deployment of the tether However exceeding a crit ical ini t ia l
cartwheeling rate can result in a large tension in the tether which eventually
causes the satellite to bounce back rendering the tether slack
(vii) Spinning the tether about its nominal length as in the case of O E D I P U S
introduces high frequency transverse vibrations in the tether which in turn
affect the dynamics of the satellites
(viii) The F L T based controller is quite successful in regulating the attitude moshy
tion of both the tether and rigid satellites in a short period of time during
stationkeeping deployment as well as retrieval phases
(ix) The offset controller is successful in suppressing the elastic vibrations of the
tether wi thin the allowable l imits of the offset mechanism ( plusmn 2 0 m)
62 Recommendations for Future Study
A s can be expected any scientific inquiry is merely a prelude to further efforts
and insight Several possible avenues exist for further investigation in this field A
few related to the study presented here are indicated below
(i) inclusion of environmental effects particularly the free molecular reaction
forces as well as interaction with Earths magnetic field and solar radiation
(ii) addition of flexible booms attached to the rigid satellites providing a mechashy
nism for energy dissipation
(iii) validation of the new three dimensional offset control strategies using ground
based experiments
(iv) animation of the simulation results to allow visual interpretation of the system
dynamics
127
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132
A P P E N D I X I TENSOR R E P R E S E N T A T I O N OF T H E
EQUATIONS OF M O T I O N
11 Prel iminary Remarks
The derivation of the equations of motion for the multibody tethered system
involves the product of several matrices as well as the derivative of matrices and
vectors with respect to other vectors In order to concisely code these relationships
in FORTRAN the matrix equations of motion are expressed in tensor notation
Tensor mathematics in general is an extremely powerful tool which can be
used to solve many complex problems in physics and engineering[6465] However
only a few preliminary results of Cartesian tensor analysis are required here They
are summarized below
12 Mathematical Background
The representation of matrices and vectors in tensor notation simply involves
the use of indices referring to the specific elements of the matrix entity For example
v = vk (k = lN) (Il)
A = Aij (i = lNj = lM)
where vk is the kth element of vector v and Aj is the element on the t h row and
3th column of matrix A It is clear from Eq(Il) that exactly one index is required to
completely define an entire vector whereas two independent indices are required for
matrices For this reason vectors and matrices are known as first-order and second-
order tensors respectively Scaler variables are known as zeroth-order tensors since
in this case an index is not required
133
Matrix operations are expressed in tensor notation similar to the way they are
programmed using a computer language such as FORTRAN or C They are expressed
in a summation notation known as Einstein notation For example the product of a
matrix with a vector is given as
w = Av (12)
or in tensor notation W i = ^2^2Aij(vkSjk)
4-1 ltL3gt j
where 8jk is the Kronecker delta Dropping the summation symbol the matrix-vector
product can be expressed compactly as
Wi = AijVj = Aimvm (14)
Here j is a dummy index since it appears twice in a single term and hence can
be replaced by any other index Note that since the resulting product is a vector
only one index is required namely i that appears exactly once on both sides of the
equation Similarly
E = aAB + bCD (15)
or Ej mdash aAimBmj + bCimDmj
~ aBmjAim I bDmjCim (16)
where A B C D and E are second-order tensors (matrices) and a and b are zeroth-
order tensors (scalers) In addition once an expression is in tensor form it can be
treated similar to a scaler and the terms can be rearranged
The transpose of a matrix is also easily described using tensors One simply
switches the position of the indices For example let w = ATv then in tensor-
notation
Wi = AjiVj (17)
134
The real power of tensor notation is in its unambiguous expression of terms
containing partial derivatives of scalers vectors or matrices with respect to other
vectors For example the Jacobian matrix of y = f(x) can be easily expressed in
tensor notation as
di-dx-j-1 ( L 8 )
mdash which is a second-order tensor as expected If f(x) = Ax where A is constant then
f - a )
according to the current definition If C = A(x)B(x) then the time derivative of C
is given by
C = Ax)Bx) + A(x)Bx) (110)
or in tensor notation
Cij = AimBmj + AimBmj (111)
= XhiAimfcBmj + AimBmj^k)
which is more clear than its equivalent vector form Note that A^^ and Bmjk are
third-order tensors that can be readily handled in F O R T R A N
13 Forcing Function
The forcing function of the equations of motion as given by Eq(260) is
- u 1-SdM- dqtdPe 9 f i m F(q qt) = Mq - -q mdashq + mdashmdash - Qd (260)
This expression can be converted into tensor form to facilitate its implementation
into F O R T R A N source code
Consider the first term Mq From Eq(289) the mass matrix M is given by
T M mdash Rv MtRv which in tensor form is
Mtj = RZiM^R^j (112)
135
Taking the time derivative of M gives
Mij = qsRv
ni^MtnrnRv
mj + Rv
niMtnTn^Rv
mj + Rv
niMtnmRv
mjs) (113)
The second term on the right hand side in Eq(260) is given by
-JPdMu 1 bdquo T x
2Q ~cWq = 2Qs s r k Q r ^ ^
Expanding Msrk leads to
M src = RnskMtnmRmr + RnsMtnmkRmr + RnsMtnmRmrk- (L15)
Finally the potential energy term in Eq(260) is also expanded into tensor
form to give dPe = dqt dPe
dq dq dqt = dq^QP^
and since ^ = Rp(q)q then
Now inserting Eqs(I13-I17) into Eq(260) and rearranging the tensor form of the
forcing term can how be stated as
Fk(q^q^)=(RP
sk + R P n J q n ) ^ - - Q d k
+ QsQr | (^Rlks ~ 2 M t n m R m r
+ (^RnkMtnms ~ 2~RnsMtnmgtk^ R m r
(^RnkRmrs ~ ~2RnsRmrk^j
where Fk(q q t) represents the kth component of F
(118)
136
A P P E N D I X II R E D U C E D EQUATIONS OF M O T I O N
II 1 Prel iminary Remarks
The reduced model used for the design of the vibration controller is represented
by two rigid end-bodies capable of three-dimensional attitude motion interconnected
with a fixed length tether The flexible tether is discretized using the assumed-
mode method with only the fundamental mode of vibration considered to represent
longitudinal as well as in-plane and out-of-plane transverse deformations In this
model tether structural damping is neglected Finally the system is restricted to a
nominal circular orbit
II 2 Derivation of the Lagrangian Equations of Mot ion
Derivation of the equations of motion for the reduced system follows a similar
procedure to that for the full case presented in Chapter 2 However in this model the
straightforward derivation of the energy expressions for the entire system is undershy
taken ignoring the Order(N) approach discussed previously Furthermore in order
to make the final governing equations stationary the attitude motion is now referred
to the local LVLH frame This is accomplished by simply defining the following new
attitude angles
ci = on + 6
Pi = Pi (HI)
7t = 0
where and are the pitch and roll angles respectively Note that spin is not
considered in the reduced model hence 7J = 0 Substituting Eq(IIl) into Eq(214)
137
gives the new expression for the rotation matrix as
T- CpiSa^ Cai+61 S 0 S a + e i
-S 0 c Pi
(II2)
W i t h the new rotation matrix defined the inertial velocity of the elemental
mass drrii o n the ^ link can now be expressed using Eq(215) as
(215)
Since the tether length is assumed to be fixed T j f j and Tj$jlt^ of Eq(215) are
identically zero Also defining
(II3) Vi = I Pi i
Rdm- f deg r the reduced system is given by
where
and
Rdm =Di + Pi(gi)m + T^iSi
Pi(9i)m= [Tai9i T0gi T^] fj
(II4)
(II5)
(II6)
D1 = D D l D S v
52 = 31 + P(d2)m +T[d2
D3 = D2 + l2 + dX2Pii)m + T2i6x
Note that $ j for i = 2 is defined in Section 212 whereas for i = 1 and 3 it is
the null matrix Since only one mode is used for each flexible degree of freedom
5i = [6Xi6yi6Zi]T for i = 2
The kinetic energy of the system can now be written as
i=l Z i = l Jmi RdmRdmdmi (UJ)
138
Expanding Kampi using Eq(II4)
KH = miDD + 2DiP
i j gidmi)fji + 2D T ltMtrade^
+tf J P^(9i)Pi9i)d^l + 2mT j PFmTi^dm^ (II8)
where the integrals are evaluated using the procedure similar to that described in
Chapter 2 The gravitational and strain energy expressions are given by Eq(247)
and Eq(251) respectively using the newly defined rotation matrix T in place of
TV
Substituting the kinetic and potential energy expressions in
d ( 8Ke d(Ke - Pe) i bull i Q ^ 0 (II9)
d t dqred d(ired
with
Qred = [ algt l gt a 2 2 A 2 A 2 lt ^ 2 Q 3 3 ] (5-14)
and d2 regarded as a time-varying parameter the nonlinear non-autonomous equashy
tions of motion for the reduced system can finally be expressed as
MredQred + fred = reg- (5-15)
139
T A B L E OF C O N T E N T S
A B S T R A C T bull i i
T A B L E O F C O N T E N T S i i i
L I S T O F S Y M B O L S v i i
L I S T O F F I G U R E S x i
A C K N O W L E D G E M E N T xv
1 I N T R O D U C T I O N 1
11 Prel iminary Remarks 1
12 Brief Review of the Relevant Literature 7
121 Mul t ibody O(N) formulation 7
122 Issues of tether modelling 10
123 Att i tude and vibration control 11
13 Scope of the Investigation 12
2 F O R M U L A T I O N O F T H E P R O B L E M 14
21 Kinematics 14
211 Prel iminary definitions and the itfl position vector 14
212 Tether flexibility discretization 14
213 Rotat ion angles and transformations 17
214 Inertial velocity of the ith link 18
215 Cyl indr ica l orbital coordinates 19
216 Tether deployment and retrieval profile 21
i i i
22 Kinetics and System Energy 22
221 Kinet ic energy 22
222 Simplification for rigid links 23
223 Gravitat ional potential energy 25
224 Strain energy 26
225 Tether energy dissipation 27
23 O(N) Form of the Equations of Mot ion 28
231 Lagrange equations of motion 28
232 Generalized coordinates and position transformation 29
233 Velocity transformations 31
234 Cyl indrical coordinate modification 33
235 Mass matrix inversion 34
236 Specification of the offset position 35
24 Generalized Control Forces 36
241 Preliminary remarks 36
242 Generalized thruster forces 37
243 Generalized momentum gyro torques 39
3 C O M P U T E R I M P L E M E N T A T I O N 42
31 Preliminary Remarks 42
32 Numerical Implementation 43
321 Integration routine 43
322 Program structure 43
33 Verification of the Code 46
iv
331 Energy conservation 4 6
332 Comparison with available data 50
4 D Y N A M I C S I M U L A T I O N 53
41 Prel iminary Remarks 53
42 Parameter and Response Variable Definitions 53
43 Stationkeeping Profile 56
44 Tether Deployment 66
45 Tether Retrieval 71
46 Five-Body Tethered System 71
47 B I C E P S Configuration 80
48 O E D I P U S Spinning Configuration 88
5 A T T I T U D E A N D V I B R A T I O N C O N T R O L 93
51 Att i tude Control 93
511 Prel iminary remarks 93
512 Controller design using Feedback Linearization Technique 94
513 Simulation results 96
52 Control of Tethers Elastic Vibrations 110
521 Prel iminary remarks 110
522 System linearization and state-space realization 1 1 0
523 Linear Quadratic Gaussian control
with Loop Transfer Recovery 114
524 Simulation results 121
v
6 C O N C L U D I N G R E M A R K S 1 2 4
61 Summary of Results 124
62 Recommendations for Future Study 127
B I B L I O G R A P H Y 128
A P P E N D I C E S
I T E N S O R R E P R E S E N T A T I O N O F T H E E Q U A T I O N S O F M O T I O N 133
11 Prel iminary Remarks 133
12 Mathematical Background 133
13 Forcing Function 135
II R E D U C E D E Q U A T I O N S O F M O T I O N 137
II 1 Preliminary Remarks 137
II2 Derivation of the Lagrangian Equations of Mot ion 137
v i
LIST OF SYMBOLS
A j intermediate length over which the deploymentretrieval profile
for the itfl l ink is sinusoidal
A Lagrange multipliers
EUQU longitudinal state and measurement noise covariance matrices
respectively
matrix containing mode shape functions of the t h flexible link
pitch roll and yaw angles of the i 1 link
time-varying modal coordinate for the i 1 flexible link
link strain and stress respectively
damping factor of the t h attitude actuator
set of attitude angles (fjj = ci
structural damping coefficient for the t h l ink EjJEj
true anomaly
Earths gravitational constant GMejBe
density of the ith link
fundamental frequency of the link longitudinal tether vibrashy
tion
A tethers cross-sectional area
AfBfCjDf state-space representation of flexible subsystem
Dj inertial position vector of frame Fj
Dj magnitude of Dj mdash mdash
Df) transformation matrix relating Dj and Ds
D r i inplane radial distance of the first l ink
DSi Drv6i DZ1 T for i = 1 and DX Dy poundgt 2 T for 1
m
Pi
w 0 i
v i i
Dx- transformation matrix relating D j and Ds-
DXi Dyi Dz- Cartesian components of D
DZl out-of-plane position component of first link
E Youngs elastic modulus for the ith l ink
Ej^ contribution from structural damping to the augmented complex
modulus E
F systems conservative force vector
FQ inertial reference frame
Fi t h link body-fixed reference frame
n x n identity matrix
alternate form of inertia matrix for the i 1 l ink
K A Z $i(k + dXi+1)
Kei t h link kinetic energy
M(q t) systems coupled mass matrix
Mma Mma Mmy- control moments in the pitch roll and yaw directions respec-
tively for the t h link
Mr fr rigid mass matrix and force vector respectively
Mred fred- Qred mass matrix force vector and generalized coordinate vector for
the reduced model respectively
Mt block diagonal decoupled mass matrix
symmetric decoupled mass matrix
N total number of links
0N) order-N
Pi(9i) column matrix [T^giTpg^T^gi) mdash
Q nonconservative generalized force vector
Qu actuator coupling matrix or longitudinal L Q R state weighting
matrix
v i i i
Rdm- inertial position of the t h link mass element drrii
RP transformation matrix relating qt and tf
Ru longitudinal L Q R input weighting matrix
RvRn Rd transformation matrices relating qt and q
S(f Mdc) generalized acceleration vector of coupled system q for the full
nonlinear flexible system
th uiirv l u w u u u maniA
th
T j i l ink rotation matrix
TtaTto control thrust in the pitch and roll direction for the im l ink
respectively
rQi maximum deploymentretrieval velocity of the t h l ink
Vei l ink strain energy
Vgi link gravitational potential energy
Rayleigh dissipation function arising from structural damping
in the ith link
di position vector to the frame F from the frame F_i bull
dc desired offset acceleration vector (i 1 l ink offset position)
drrii infinitesimal mass element of the t h l ink
dx^dy^dZi Cartesian components of d along the local vertical local horishy
zontal and orbit normal directions respectively
F-Q mdash
gi r igid and flexible position vectors of drrii fj + ltEraquolt5
ijk unit vectors 1 0 0 r 010-^ and 00 l r respectively
li length of the t h l ink
raj mass of the t h link
nfj total number of flexible modes for the i 1 l ink nXi + nyi + nZi
nq total number of generalized coordinates per link nfj + 7
nqq systems total number of generalized coordinates Nnq
ix
number of flexible modes in the longitudinal inplane and out-
of-plane transverse directions respectively for the i ^ link
number of attitude control actuators
- $ T
set of generalized coordinates for the itfl l ink which accounts for
interactions with adjacent links
qv---QtNT
set of coordinates for the independent t h link (not connected to
adjacent links)
rigid position of dm in the frame Fj
position of centre of mass of the itfl l ink relative to the frame Fj
i + $DJi
desired settling time of the j 1 attitude actuator
actuator force vector for entire system
flexible deformation of the link along the Xj yi and Zj direcshy
tions respectively
control input for flexible subsystem
Cartesian components of fj
actual and estimated state of flexible subsystem respectively
output vector of flexible subsystem
LIST OF FIGURES
1-1 A schematic diagram of the space platform based N-body tethered
satellite system 2
1-2 Associated forces for the dumbbell satellite configuration 3
1- 3 Some applications of the tethered satellite systems (a) multiple
communication satellites at a fixed geostationary location
(b) retrieval maneuver (c) proposed B I C E P S configuration 6
2- 1 Vector components of the i 1 and (i- l ) 1 chain links 15
2-2 Vector components in cylindrical orbital coordinates 20
2-3 Inertial position of subsatellite thruster forces 38
2- 4 Coupled force representation of a momentum-wheel on a rigid body 40
3- 1 Flowchart showing the computer program structure 45
3-2 Kinet ic and potentialenergy transfer for the three-body platform based
tethered satellite system (a) variation of kinetic and potential energy
(b) percent change in total energy of system 48
3-3 Kinet ic and potential energy transfer for the five-body platform based
tethered satellite system (a) variation of kinetic and potential energy
(b) percent change in total energy of system 49
3-4 Simulation results for the platform based three-body tethered
system originally presented in Ref[43] 51
3- 5 Simulation results for the platform based three-body tethered
system obtained using the present computer program 52
4- 1 Schematic diagram showing the generalized coordinates used to
describe the system dynamics 55
4-2 Stationkeeping dynamics of the three-body STSS configuration
without offset (a) attitude response (b) vibration response 57
4-3 Stationkeeping dynamics of the three-body STSS configuration
xi
with offset along the local vertical
(a) attitude response (b) vibration response 60
4-4 Stationkeeping dynamics of the three-body STSS configuration
with offset along the local horizontal
(a) attitude response (b) vibration response 62
4-5 Stationkeeping dynamics of the three-body STSS configuration
with offset along the orbit normal
(a) attitude response (b) vibration response 64
4-6 Stationkeeping dynamics of the three-body STSS configuration
wi th offset along the local horizontal local vertical and orbit normal
(a) attitude response (b) vibration response 67
4-7 Deployment dynamics of the three-body STSS configuration
without offset
(a) attitude response (b) vibration response 69
4-8 Deployment dynamics of the three-body STSS configuration
with offset along the local vertical
(a) attitude response (b) vibration response 72
4-9 Retrieval dynamics of the three-body STSS configuration
with offset along the local vertical and orbit normal
(a) attitude response (b) vibration response 74
4-10 Schematic diagram of the five-body system used in the numerical example 77
4-11 Stationkeeping dynamics of the five-body STSS configuration
without offset
(a) attitude response (b) vibration response 78
4-12 Deployment dynamics of the five-body STSS configuration
wi th offset along the local vertical
(a) attitude response (b) vibration response 81
4-13 Stationkeeping dynamics of the three-body B I C E P S configuration
x i i
with offset along the local vertical
(a) attitude response (b) vibration response 84
4-14 Cartwheeling dynamics with deployment for the three-body B I C E P S
with offset along the local vertical
(a) attitude response (b) vibration response 86
4-15 Spin dynamics (7 = l deg s ) of the three-body O E D I P U S configuration
with offset along the local vertical
(a) attitude response (b) vibration response 89
4- 16 Spin dynamics (7 = 10deg s ) of the three-body O E D I P U S configuration
wi th offset along the local vertical
(a) attitude response (b) vibration response 91
5- 1 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear rigid F L T controller with offset along the
local horizontal local vertical and orbit normal
(a) attitude and vibration response (b) control actuator time histories 98
5-2 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear flexible F L T controller with offset along the
local horizontal local vertical and orbit normal
(a) attitude and vibration response (b) control actuator time histories 101
5-3 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear rigid F L T controller with offset along
the local vertical
(a) attitude and vibration response (b) control actuator time histories 103
5-4 Deployment dynamics of the three-body STSS using the nonlinear
rigid F L T controller with offset along the local vertical
(a) attitude and vibration response (b) control actuator time histories 105
5-5 Retrieval dynamics of the three-body STSS using the non-linear
rigid F L T controller with offset along the local vertical and orbit normal
(a) attitude and vibration response (b) control actuator time histories 108
x i i i
5-6 Block diagram for the L Q G L T R estimator based compensator 115
5-7 Singular values for the L Q G and L Q G L T R compensator compared to target
return ratio (a) longitudinal design (b) transverse design 118
5-8 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear rigid F L T attitude controller and L Q G L T R
offset vibration controller (a) attitude and libration controller response
(b) vibration and offset response 122
xiv
A C K N O W L E D G E M E N T
Firs t and foremost I would like to express the most genuine gratitude to
my supervisor Prof V i n o d J M o d i whose unending guidance and encouragement
proved most invaluable during the course of my studies
I am also deeply indebted to Dr Satyabrata Pradhan for his limitless patience
and technical advice from which this work would not have been possible I would also
like to extend a word of appreciation towards Prof A r u n K Mis ra ( M c G i l l University)
for ini t iat ing my interest in space dynamics and control
Also I would like to thank my collegues Dr Sandeep Munshi M r Mathieu
Caron and M r Yuan Chen as well as Dr Mae Seto M r Gary L i m and M r Mark
Chu for their words of advice and helpful suggestions that heavily contributed to all
aspects of my studies
Final ly I would like to thank all my family and friends here in Vancouver and
back home in Montreal whose unending support made the past two years the most
memorable of my life
The research project was supported by the Natural Sciences and Engineering
Research Council s P G S - A Scholarship held by the author as well as by N S E R C grant
A-2181 held by Prof V J M o d i
xv
1 INTRODUCTION
11 Prel iminary Remarks
W i t h the ever changing demands of the worlds population one often wonders
about the commitment to the space program On the other hand the importance
of worldwide communications global environmental monitoring as well as long-term
habitation in space have demanded more ambitious and versatile satellites systems
It is doubtful that the Russian scientist Tsiolkovsky[l] when he first proposed the
uti l izat ion of the Earths gravity-gradient environment in the last century would have
envisioned the current and proposed applications of tethered satellite systems
A tethered satellite system consists of two or more subsatellites connected
to each other by long thin cables or tethers which are in tension (Figure 1-1) The
subsatellite which can also include the shuttle and space station are in orbit together
The first proposed use of tethers in space was associated with the rescue of stranded
astronauts by throwing a buoy on a tether and reeling it to the rescue vehicle
Prel iminary studies of such systems led to the discovery of the inherent instability
during the tether retrieval[2] Nevertheless the bir th of tethered systems occured
during the Gemini X I and X I I missions[3] in 1966 when a short tether (10-30m)
was used to generate the first artificial gravity environment in space (000015g) by
cartwheeling (spinning) the system about its center of mass
The mission also demonstrated an important use of tethers for gravity gradient
stabilization The force analysis of a simplified model illustrates this point Consider
the dumbbell satellite configuration orbiting about the centre of Ear th as shown in
Figure 1-2 Satellite 1 and satellite 2 are connected to each other by a tether with
1
Space Station (Satellite 1 )
2
Satellite N
Figure 1-1 A schematic diagram of the space platform based N-body tethered satellite system
2
Figure 1-2 Associated forces for the dumbbell satellite configuration
3
the whole system free to librate about the systems centre of mass C M The two
major forces acting on each body are due to the centrifugal and gravitational effects
Since body 1 is further away from the earth its gravitational force Fg^y is smaller
than that of body 2 ie Fgi lt Fg^ However its centrifugal force FCl is greater
then the centrifugal force experienced by the second satellite (FCl gt FC2) Moreover
for body 1 Fg^ lt Fc^ in contrast to Fg2 gt FC2 for body 2 Thus resolving the force
components along the tether there is an evident resultant tension force Ff present
in the tether Similarly adding the normal components of the two forces vectorially
results in a force FR which restores the system to its stable equil ibrium along the
local vertical These tension and gravity-gradient restoring forces are the two most
important features of tethered systems
Several milestone missions have flown in the recent past The U S A J a p a n
project T P E (Tethered Payload Experiment) [4] was one of the first to be launched
using a sounding rocket to conduct environmental studies The results of the T P E
provided support for the N A S A I t a l y shuttle based tethered satellite system referred
to as the TSS-1 mission[5] This experiment aimed to study the electrodynamic effects
of a shuttle borne satellite system with a conductive tether connection where an
electric current was induced in the tether as it swept through the Earth s magnetic
field Unfortunetely the two attempts in August 1992 and February 1996 resulted
in only partial success The former suffered from a spool failure resulting in a tether
deployment of only 256m of a planned 20km During the latter attempt a break
in the tethers insulation resulted in arcing leading to tether rupture Nevertheless
the information gathered by these attempts st i l l provided the engineers and scientists
invaluable information on the dynamic behaviour of this complex system
Canada also paved the way with novel designs of tethered system aimed
primari ly at the study of the ionosphere and Northern Lights (Aurora Borealis)
4
The O E D I P U S A and C (Observation of Electrified Distributions in the Ionospheric
P lasma - a Unique Strategy) missions[6] launched from a sounding rocket in January
1989 and November 1995 respectively provided insight into the complex dynamical
behaviour of two comparable mass satellites connected by a 1 km long spinning
tether
Final ly two experiments were conducted by N A S A called S E D S I and II
(Small Expendable Deployable System) [7] which hold the current record of 20 km
long tether In each of these missions a small probe (26 kg) was deployed from the
second stage of a D E L T A II rocket thus succesfully demonstrating the feasibility of
long tether deployment Retrieval of the tether which is significantly more difficult
has not yet been achieved
Several proposed applications are also currently under study These include the
study of earths upper atmosphere using probes lowered from the shuttle in a low earth
orbit ( L E O ) to the deployment and retrieval of satellites for servicing Even more
promising application concerns the generation of power for the proposed International
Space Station using conductive tethers The Canadian Space Agency (CSA) has also
proposed a dual-mass satellite system B I C E P S (Bl-static Canadian Experiment on
Plasmas in Space) to make simultaneous measurements at different locations in the
environment for correlation studies[8] A unique feature of the B I C E P S mission is
the deployment of the tether aided by the angular momentum of the system which
is ini t ia l ly cartwheeling about its orbit normal (Figure 1-3)
In addition to the three-body configuration multi-tethered systems have been
proposed for monitoring Earths environment in the global project monitored by
N A S A entitled Mission to Planet Ear th ( M T P E ) The proposed mission aims to
study pollution control through the understanding of the dynamic interactions be-
5
Retrieval of Satellite for Servicing
lt mdash
lt Satellite
Figure 1-3 Some applications of the tethered satellite systems (a) multiple comshymunication satellites at a fixed geostationary location (b) retrieval maneuver (c) proposed BICEPS configuration
6
tween the Earths atmosphere biosphere hydrosphere and cryosphere This wi l l be
accomplished with multiple tethered instrumentation payloads simultaneously soundshy
ing at different altitudes for spatial correlations Such mult ibody systems are considshy
ered to be the next stage in the evolution of tethered satellites Micro-gravity payload
modules suspended from the proposed space station for long-term experiments as well
as communications satellites with increased line-of-sight capability represent just two
of the numerous possible applications under consideration
12 Brief Review of the Relevant Literature
The design of multi-payload systems wi l l require extensive dynamic analysis
and parametric study before the final mission configuration can be chosen This
wi l l include a review of the fundamental dynamics of the simple two-body tethered
systems and how its dynamics can be extended to those of multi-body tethered system
in a chain or more generally a tree topology
121 Multibody 0(N) formulation
Many studies of two-body systems have been reported One of the earliest
contribution is by Rupp[9] who was the first to study planar dynamics of the simshy
plified Shuttle based Tethered Satellite System (STSS) His pioneering work led to
the discovery of pitch oscillation growth during the retrieval phase Later Baker[10]
advanced the investigation to the third dimension in addition to adding atmospheric
effects to the system A more complete dynamical model was later developed and
analyzed by M o d i and Misra[ l l 12] which included deployment and tether flexibility
A comprehensive survey of important developments and milestones in the area have
been compiled and reviewed by Mis ra and Modi[13] The major conclusions based on
the literature may be summarized as follows
bull the stationkeeping phase is stable
7
bull deployment can be unstable if the rate exceeds a crit ical speed
bull retrieval of the tether is inherently unstable
bull transverse vibrations can grow due to the Coriolis force induced during deshy
ployment and retrieval
Misra Amier and Modi[14] were one of the first to extend these results to the
three-satellite double pendulum case This simple model which includes a variable
length tether was sufficient to uncover the multiple equilibrium configurations of
the system Later the work was extended to include out-of-plane motion and a
preliminary reel-rate control law to regulate the librational motion[15] Kesmir i and
Misra[16] developed a general formulation for N-body tethered systems based on the
Lagrangian principle where three-dimensional motion and flexibility are accounted for
However from their work it is clear that as the number of payload or bodies increases
the computational cost to solve the forward dynamics also increases dramatically
Tradit ional methods of inverting the mass matrix M in the equation
ML+F = Q
to solve for the acceleration vector proved to be computationally costly being of the
order for practical simulation algorithms O(N^) refers to a mult ibody formulashy
tion algorithm whose computational cost is proportional to the cube of the number
of bodies N used It is therefore clear that more efficient solution strategies have to
be considered
Many improved algorithms have been proposed to reduce the number of comshy
putational steps required for solving for the acceleration vector Rosenthal[17] proshy
posed a recursive formulation based on the triangularization of the equations of moshy
tion to obtain an 0(N2) formulation He also demonstrates the use of a symbolic
manipulation program to reduce the final number of computations A recursive L a -
8
grangian formulation proposed by Book[18] also points out the relative inefficiency of
conventional formulations and proceeds to derive an algorithm which is also 0(N2)
A recursive formulation is one where the equations of motion are calculated in order
from one end of the system to the other It usually involves one or more passes along
the links to calculate the acceleration of each link (forward pass) and the constraint
forces (backward or inverse pass) Non-recursive algorithms express the equations of
motion for each link independent of the other Although the coupling terms st i l l need
to be included the algorithm is in general more amenable to parallel computing
A st i l l more efficient dynamics formulation is an O(N) algorithm Here the
computational effort is directly proportional to the number of bodies or links used
Several types of such algorithms have been developed over the years The one based
on the Newton-Euler approach has recursive equations and is used extensively in the
field of multi-joint robotics[1920] It should be emphasized that efficient computation
of the forward and inverse dynamics is imperative if any on-line (real-time) control
of the joint is to be achieved Hollerbach[21] proposed a recursive O(N) formulation
based on Lagranges equations of motion However his derivation was primari ly foshy
cused on the inverse dynamics and did not improve the computational efficiency of
the forward dynamics (acceleration vector) Other recursive algorithms include methshy
ods based on the principle of vir tual work[22] and Kanes equations of motion[23]
Keat[24] proposed a method based on a velocity transformation that eliminated the
appearance of constraint forces and was recursive in nature O n the other hand
K u r d i l a Menon and Sunkel[25] proposed a non-recursive algorithm based on the
Range-Space method which employs element-by-element approach used in modern
finite-element procedures Authors also demonstrated the potential for parallel comshy
putation of their non-recursive formulation The introduction of a Spatial Operator
Factorization[262728] which utilizes an analogy between mult ibody robot dynamics
and linear filtering and smoothing theory to efficiently invert the systems mass ma-
9
t r ix is another approach to a recursive algorithm Their results were further extended
to include the case where joints follow user-specified profiles[29] More recently Prad-
han et al [30] proposed a non-recursive Lagrangian algorithm for factorizing the mass
matrix leading to an O(N) formulation of the forward dynamics of the system A n
efficient O(N) algorithm where the number of bodies N in the system varies on-line
has been developed by Banerjee[31]
122 Issues of tether modelling
The importance of including tether flexibility has been demonstrated by several
researchers in the past However there are several methods available for modelling
tether flexibility Each of these methods has relative advantages and limitations
wi th regards to the computational efficiency A continuum model where flexible
motion is discretized using the assumed mode method has been proposed[3233] and
succesfully implemented In the majority of cases even one flexible mode is sufficient
to capture the dynamical behaviour of the tether[34] K i m and Vadali[35] proposed a
so-called bead model or lumped-mass approach that discretizes the tether using point
masses along its length However in order to accurately portray the motion of the
tether a significantly high number of beads are needed thus increasing the number of
computation steps Consequently this has led to the development of a hybrid between
the last two approaches[16] ie interpolation between the beads using a continuum
approach thus requiring less number of beads Finally the tether flexibility can also
be modelled using a finite-element approach[36] which is more suitable for transient
response analysis In the present study the continuum approach is adopted due to
its simplicity and proven reliability in accurately conveying the vibratory response
In the past the modal integrals arising from the discretization process have
been evaluted numerically In general this leads to unnecessary effort by the comshy
puter Today these integrals can be evaluated analytically using a symbolic in-
10
tegration package eg Maple V Subsequently they can be coded in F O R T R A N
directly by the package which could also result in a significant reduction in debugging
time More importantly there is considerable improvement in the computational effishy
ciency [16] especially during deployment and retrieval where the modal integral must
be evaluated at each time-step
123 Attitude and vibration control
In view of the conclusions arrived at by some of the researchers mentioned
above it is clear that an appropriate control strategy is needed to regulate the dyshy
namics of the system ie attitude and tether vibration First the attitude motion of
the entire system must be controlled This of course would be essential in the case of
a satellite system intended for scientific experiments such as the micro-gravity facilishy
ties aboard the proposed Internation Space Station Vibra t ion of the flexible members
wi l l have to be checked if they affect the integrity of the on-board instrumentation
Thruster as well as momentum-wheel approach[34] have been considered to
regulate the rigid-body motion of the end-satellite as well as the swinging motion
of the tether The procedure is particularly attractive due to its effectiveness over a
wide range of tether lengths as well as ease of implementation Other methods inshy
clude tension and length rate control which regulate the tethers tension and nominal
unstretched length respectively[915] It is usually implemented at the deployment
spool of the tether More recently an offset strategy involving time dependent moshy
tion of the tether attachment point to the platform has been proposed[3738] It
overcomes the problem of plume impingement created by the thruster control and
the ineffectiveness of tension control at shorter lengths However the effectiveness of
such a controller can become limited with an exceedingly long tether due to a pracshy
tical l imi t on the permissible offset motion In response to these two control issues a
hybrid thrusteroffset scheme has been proposed to combine the best features of the
11
two methods[39]
In addition to attitude control these schemes can be used to attenuate the
flexible response of the tether Tension and length rate control[12] as well as thruster
based algorithms[3340] have been proposed to this end M o d i et al [41] have sucshy
cessfully demonstrated the effectiveness of an offset strategy to damp undesired v i shy
brations Passive energy dissipative devices eg viscous dampers are also another
viable solution to the problem
The development of various control laws to implement the above mentioned
strategies has also recieved much attention A n eigenstructure assignment in conducshy
tion wi th an offset controller for vibration attenuation and momemtum wheels for
platform libration control has been developed[42] in addition to the controller design
from a graph theoretic approach[41] Also non-linear feedback methods such as the
Feedback Linearization Technique ( F L T ) that are more suitable for controlling highly
nonlinear non-autonomous coupled systems have also been considered[43]
It is important to point out that several linear controllers including the classhy
sic state feedback Linear Quadratic Regulator ( L Q R ) have received considerable
attention[3944] Moreover robust methods such as the Linear Quadratic Guas-
s ian Loop Transfer Recovery ( L Q G L T R ) method have also been developed and imshy
plemented on tethered systems[43] A more complete review of these algorithms as
well as others applied to tethered system has been presented by Pradhan[43]
13 Scope of the Investigation
The objective of the thesis is to develop and implement a versatile as well as
computationally efficient formulation algorithm applicable to a large class of tethered
satellite systems The distinctive features of the model can be summarized as follows
12
bull TV-body 0(N) tethered satellite system in a chain-type topology
bull the system is free to negotiate a general Keplerian orbit and permitted to
undergo three dimensional inplane and out-of-plane l ibrtational motion
bull r igid bodies constituting the system are free to execute general rotational
motion
bull three dimensional flexibility present in the tether which is discretized using
the continuum assumed-mode method with an arbitrary number of flexible
modes
bull energy dissipation through structural damping is included
bull capability to model various mission configurations and maneuvers including
the tether spin about an arbitrary axis
bull user-defined time dependent deployment and retrieval profiles for the tether
as well as the tether attachment point (offset)
bull attitude control algorithm for the tether and rigid bodies using thrusters and
momentum-wheels based on the Feedback Linearization Technique
bull the algorithm for suppression of tether vibration using an active offset (tether
attachment point) control strategy based on the optimal linear control law
( L Q G L T - R )
To begin with in Chapter 2 kinematics and kinetics of the system are derived
using the O(N) formulation methodology developed by Pradhan et al [30] Chapshy
ter 3 discusses issues related to the development of the simulation program and its
validation This is followed by a detailed parametric study of several mission proshy
files simulated as particular cases of the versatile formulation in Chapter 4 Next
the attitude and vibration controllers are designed and their effectiveness assessed in
Chapter 5 The thesis ends with concluding remarks and suggestions for future work
13
2 F O R M U L A T I O N OF T H E P R O B L E M
21 Kinematics
211 Preliminary definitions and the i ^ position vector
The derivation of the equations of motion begins with the definition of the
inertial position of an arbitrary link of the mult ibody system Let the i ^ link of the
system be free to translate and rotate in 3-D space From Figure 2-1 the position
vector Rdm^ to the mass element drrn on the i ^ link wi th reference to the inertial
frame FQ can be written as
Here D represents the inertial position of the t h body fixed frame Fj relative to FQ
ri = [xiyiZi]T is the rigid position vector of drrii wi th respect to F J and f(fi) is
the flexible deformation at fj also with respect to the frame Fj Bo th these vectors
are relative to Fj Note in the present model tethers (i even) are considered flexible
and X corresponds to the nominal position along the unstretched tether while a and
ZJ are by definition equal to zero On the other hand the rigid bodies referred to as
satellites (i odd) including the platform are taken to be rigid ie f(fj) = 0 T j is
the rotation matrix used to express body-fixed vectors with reference to the inertial
frame
212 Tether flexibility discretization
The tether flexibility is discretized with an arbitrary but finite number of
Di + Ti(i + f(i)) (21)
14
15
modes in each direction using the assumed-mode method as
Ui nvi
wi I I i = 1
nzi bull
(22)
where nXi nyi and n 2 ^ are the number of modes used in the X m and Z directions
respectively For the ] t h longitudinal mode the admissible mode shape function is
taken as 2 j - l
Hi(xiii)=[j) (23)
where l is the i ^ tether length[32] In the case of inplane and out-of-plane transverse
deflections the admissible functions are
ampyi(xili) = ampZixuk) = v ^ s i n JKXj
k (24)
where y2 is added as a normalizing factor In this case both the longitudinal and
transverse mode shapes satisfy the geometric boundary conditions for a simple string
in axial and transverse vibration
Eq(22) is recast into a compact matrix form as
ff(i) = $i(xili)5i(t) (25)
where $i(x l) is the matrix of the tether mode shapes defined as
0 0
^i(xiJi) =
^ X bull bull bull ^XA
0
0 0
0
G 5R 3 x n f
(26)
and nfj = nx- + nyi + nZi Note $J(XJJ) is a function of the spatial coordinate x
and not of yi or ZJ However it is also a function of the tether length parameter li
16
which is time-varying during deployment and retrieval For this reason care must be
exercised when differentiating $J(XJJ) with respect to time such that
d d d k) = Jtregi(Xi li) = ^ 7 $ t ( ^ i k)Xi + gf$i(xigt li)k- (2-7)
Since x represents the nominal rate of change of the position of the elemental mass
drrii it is equal to j (deploymentretrieval rate) Let t ing
( 3 d
dx~ + dT)^Xili^ ( 2 8 )
one arrives at
$i(xili) = $Di(xili)ii (29)
The time varying generalized coordinate vector 5 in Eq(25) is composed of
the longitudinal and transverse components ie
ltM) G s f t n f i x l (210)
where 5X^ 5y- and Sz^ are nx- x 1 ny x 1 and nz- x 1 size vectors and correspond to
$ X and $ Z respectively
213 Rotation angles and transformations
The matrix T j in Eq(21) represents a rotation transformation from the body
fixed frame Fj to the inertial frame FQ ie FQ = T j F j It is defined by the 3-2-1
ordered sequence of elementary rotations[46]
1 P i tch F[ = Cf ( a j )F 0
2 R o l l F = C f (3j)F = C7f (3j)Cf ( a j )F 0
3 Yaw F = C7( 7 i )F = C7(7l)cf^)Cf(ai)FQ = Q F 0
17
where
c f (A) =
-Sai Cai
0 0
and
C(7i ) =
0 -s0i] = 0 1 0
0 cpi
i 0 0 = 0
0
(211)
(212)
(213)
Here Eq(211) corresponds to a rotation of ctj (pitch) about the inertial axis ZQ (orbit
normal) of frame FQ resulting in a new intermediate frame F It is then followed by
a roll rotation fa about the axis y of F- giving a second intermediate frame F-
Final ly a spin rotation 7J about the axis z of F- is given resulting in the body
fixed frame Fj for the i lt l link Since all the three rotation matrices are orthogonal
it follows that FQ = C~1Fi = Cj F and hence
C CPi Cai ~ CH Sai + SrYi SPi Sai 5 7 i Sampi + Cli SPi Cci CPi Sai C1i Cai + 5 7 i SPi ~ S1 Ca + C 7 bull Sp Sai
(214) SliCPi CliCPi
Here Cx and Sx are abbreviations for cos(x) and sin(x) respectively
214 Inertial velocity of the ith link
mdash Lett ing pj = fj - f $j(5j and differentiating Eq(21) with respect to time gives
m- D j + T j f j + T^Si + T j pound + T j $ j 5 j (215)
Each of the terms appearing in Eq(215) must be addressed individually Since
fj represents the rigid body motion within the link it only has meaning for the
deployable tether and since yi = Zj = 0 for al l tether links (i even) fj = Ifi where 1
is the unit vector 1 0 0T For the case of rigid links (i odd) fj = 0
18
Evaluating the derivative of each angle in the rotation matrix gives
T^i9i = Taiai9i + Tppigi + T^fagi (216)
where Tx = J j I V Collecting the scaler angular velocity terms and defining the set
Vi = deg^gt A) 7i^ Eq(216) can be rewritten as
T i f t = Pi9i)fji (217)
where Pi(gi) is a column matrix defined by
Pi(9i)Vi = [Taigi^p^iT^gilffi (218)
Inserting Eq(29) and Eq(217) into Eq(215) and defining Sj = 2 + leads to
215 Cylindrical orbital coordinates
The Di term in Eq(219) describes the orbital motion of the t h l ink and
is composed of the three Cartesian coordinates DXi A ^ DZi However over the
cycle of one orbit DXi and Dyi vary dramatically by around the order of Earths
radius This must be avoided since large variations in the coordinates can cause
severe truncation errors during their numerical integration For this reason it is
more convenient to express the Cartesian components in terms of more stationary
variables This is readily accomplished using cylindrical coordinates
However it is only necessary to transform the first l inks orbital components
ie Lgti since only D is a generalized coordinates From Figure 2-2 it is apparent
(219)
that
(220)
19
Figure 2-2 Vector components in cylindrical orbital coordinates
20
Eq(220) can be rewritten in matrix form as
cos(0i) 0 0 s i n ( 0 i ) 0 0
0 0 1
Dri
h (221)
= DTlDSv
where Dri is the inplane radial distance 9 the true anomaly and DZl the out-of-
plane distance normal to the orbital plane The total derivative of D wi th respect
to time gives in column matrix form
1 1 1 1 J (222)
= D D l D 3 v
where [cos(^i) -Dnsm(8i) 0 ]
(223)
For all remaining links ie i = 2 N
cos(0i) - A - j S i n ^ i ) 0 sin(^i) D r i c o s ( 0 i ) 0
0 0 1
DTl = D D i = I d 3 i = 2N (224)
where is the n x n identity matrix and
2N (225)
216 Tether deployment and retrieval profile
The deployment of the tether is a critical maneuver for this class of systems
Cr i t i ca l deployment rates would cause the system to librate out of control Normally
a smooth sinusoidal profile is chosen (S-curve) However one would also like a long
tether extending to ten twenty or even a hundred kilometers in a reasonable time
say 3-4 orbits This is usually difficult or impossible to accomplish solely wi th one
S-curve profile Often a composite S-curve-Steady-S-curve profile is adopted Thus
21
the deployment scheme for the t h tether can be summarized as
k = Tr l - cos (^r(ti - to-)) 1 ltUlt h 2 r V A V
k = Vov tiilttiltt2i
k = IT j 1 - C deg S lkitl ^ ^ lt i lt ^
(226)
where to- tt are the ini t ia l and final times of the deployment or retrieval maneuver
respectively A t j = t - tQ mdash tt mdash t2- and the steady deployment velocity VQ is
calculated based on the continuity of l and l at the specified intermediate times t
and to For the case of
and
Ai = i(ti) - ^(t 0 -) = - ^ ( t 2 i ) (227)
2Ali + li(tf) -h(tQ)
V = lUli l^lL (228)
1 tf-tQ v y
A t = (229)
22 Kinetics and System Energy
221 Kinetic energy
The kinetic energy of the ith link is given by
Kei = lJm Rdrnfi^Ami (230)
Setting Eq(219) in a matrix-vector form gives the relation
4 ^ ^ PM) T i $ 4 TiSi]i[t (231)
22
where
m
k J
x l (232)
and nq = -nfj + 7 is the number of generalized coordinates in each link Inserting
Eq(231) into Eq(230) and integrating the kinetic energy for the t h l ink can be
rewritten as
1 T
(233)
where Mti is the links symmetric mass matrix
Mt =
m i D D T D D DDTPifgidmi) DDTTi J ^dm D n T Sidm rn
sym fP(9i)Pi(9i)dmi J P[g^T^idm J P ( pound ) T ^ m sym sym J ^f^idm f ^f^dmi sym sym sym J sfsidmi
pound Sfttiqxnq
(234)
and mi is the mass of the Ith link The kinetic energy for the entire system ie N
bodies can now be stated as
1 bull T bull 1 bull T Ke = J2 MtA = o laquo Mtqt 2 ^ -i
i=l (235)
where qt = qj^^ bull bull -qfNT a n d Mt is a block diagonal decoupled mass matrix
of the system with Mf- on its diagonal
(236) Mt =
0 0
0 Mt2
0
0 0
MH
0 0 0
0 0 0 bull MtN
222 Simplification for rigid links
The mass matrix given by Eq(234) simplifies considerably for the case of rigid
23
links and is given as
Mti
m D D l
T D D l DD^Piiffidmi) 0
sym 0
0
fP(fi)Pi(i)dmi 0 0 0
0 Id nfj
0 m~
i = 1 3 5 N
(237)
Moreover when mult iplying the matrices in the (11) block term of Eq(237) one
gets
D D l
1 D D l = 1 0 0 0 D n
2 0 (238) 0 0 1
which is a diagonal matrix However more importantly i t is independent of 6 and
hence is always non-singular ensuring that is always invertible
In the (12) block term of M ^ it is apparent that f fidmi is simply the
definition of the centre of mass of the t h rigid link and hence given by
To drrii mdash miff (239)
where fcmi is the position vector of the centre of mass of link i For the case of the
(22) block term a little more attention is required Expanding the integrand using
Eq(218) where ltj = fj gives
fj Ta Tafi fj T Q Tpfi f- T Q T 7 i f j
sym sym
Now since each of the terms of Eq(240) is a scaler or a 1 x 1 matrix it is therefore
equal to its trace ie sum of its diagonal elements thus
tr(ffTai
TTaifi) tr(ffTai
TTpfi) tr(ff Ta^T^fi)
fjT^Tpn rfTp^fi (240)
P(rj)Pj(fi) = syrri trifjTpTTpfi) t r ^ T ^ T ^
sym sym tr(ff T 7 T r 7 f j) J
(241)
Using two properties of the trace function[47] namely
tr(ABC) = tr(BCA) = trCAB) (242)
24
and
j tr(A) =tr(^J A^j (243)
where A B and C are real matrices it becomes quite clear that the (22) block term
of Eq(237) simplifies to
P(fi)Pii)dmi =
tr(Tai
TTaiImi) tTTai
TTpImi) t r (T Q r T 7 7 m ) trTaTTpImi) tr(Tp
TT7Jmi) sym sym sym
(244)
where Im- is an alternate form of the inertia tensor of the t h link which when
expanded is
Imlaquo = I nfi dm bull i sym
sym sym
J x^dmi J Xydmi XZdm sym J yi2dmi f yiZidm
sym J z^dmi J
xVi -L VH sym ijxxi + lyy IZZJ)2
(245)
Note this is in terms of moments of inertia and products of inertia of the t f l rigid
link A similar simplification can be made for the flexible case where T is replaced
with Qi resulting in a new expression for Im
223 Gravitational potential energy
The gravitational potential energy of the ith link due to the central force law
can be written as
V9i f dm f
-A mdash = -A bullgttradelti RdmA
dm
inn Lgti + T j ^ | (246)
where z = 3986 x 105 km 3s 2 (Earths gravitational constant) Expanding binomially
and retaining up to third order terms gives
Vgi ~ ~ D T 2D [J gfgidrrii + 23f T j gidm mdashDT
D2 i l I I Df Tt I grffdmiTiDi
(2-lt 7)
25
where D = D is the magnitude of the inertial position vector to the frame F The
integrals in the above expression are functions of the flexible mode shapes and can be
evaluated through symbolic manipulation using an existing package such as Maple
V In turn Maple V can translate the evaluated integrals into F O R T R A N and store
the code in a file Furthermore for the case of rigid bodies it can be shown quite
readily that
j gjgidrrii = j rjfidrrn = [IXXi + Iyy + IZZi)2 (248)
Similarly the other integrals in Eq(234) can be integrated symbolically
224 Strain energy
When deriving the elastic strain energy of a simple string in tension the
assumptions of high tension and low amplitude of transverse vibrations are generally
valid Consequently the first order approximation of the strain-displacement relation
is generally acceptable However for orbiting tethers in a weak gravitational-gradient
field neglecting higher order terms can result in poorly modelled tether dynamics
Geometric nonlinearities are responsible for the foreshortening effects present in the
tether These cannot be neglected because they account for the heaving motion along
the longitudinal axis due to lateral vibrations The magnitude of the longitudinal
oscillations diminish as the tether becomes shorter[35]
W i t h this as background the tether strain which is derived from the theory
of elastic vibrations[3248] is given by
(249)
where Uj V and W are the flexible deformations along the Xj and Z direction
respectively and are obtained from Eq(25) The square bracketed term in Eq(249)
represents the geometric nonlinearity and is accounted for in the analysis
26
The total strain energy of a flexible link is given by
Ve = I f CiEidVl (250)
Substituting the stress-strain relationship 0 = E(poundi the strain energy is now given
by
Vei=l-EiAi J l l d x h (251)
where E^ is the tethers Youngs modulus and A is the cross-sectional area The
tether is assumed to be uniform thus EA which is the flexural stiffness of the
tether is assumed to be constant Eq(251) is evaluated symbolically for arbitrary
number of modes
225 Tether energy dissipation
The evaluation of the energy dissipation due to tether deformations remains
a problem not well understood even today Here this complex phenomenon is repshy
resented through a simplified structural damping model[32] In addition the system
exhibits hysteresis which also must be considered This is accomplished using an
augmented complex Youngs modulus
EX^Ei+jEi (252)
where Ej^ is the contribution from the structural damping and j = The
augmented stress-strain relation is now given as
a = Epoundi = Ei(l+jrldi)ei (253)
where 77^ = EjE^ ltC 1 is defined as the structural damping coefficient determined
experimentally [49]
27
If poundi is a harmonic function with frequency UJQ- then
jet = (254) wo-
Substituting into Eq(253) and rearranging the terms gives
e + I mdash I poundi = o-i + ad (255)
where ad and a are the stress with and without structural damping respectively
Now the Rayleigh dissipation function[50] for the ith tether can be expressed as
(ii) Jo
_ 1 EjAiVdi rU (2-56^ 2 w 0 Jo
The strain rate ii is the time derivative of Eq(249) The generalized external force
due to damping can now be written as Qd = Q^ bull bull bull QdNT where
dWd
^ = - i p = 2 4 ( 2 5 7 )
= 0 i = 13 N
23 0(N) Form of the Equations of Mot ion
231 Lagrange equations of motion
W i t h the kinetic energy expression defined and the potential energy of the
whole system given by Pe = ]Cn=l( 5i+ ej) the equations of motion can be obtained
quite readily using the Lagrangian principle
where q is the set of generalized coordinates to be defined in a later section
Substituting Eqs(235247251 and 257) into Eq(258) leads to the familiar
matr ix form of the non-linear non-autonomous coupled equations of motion for the
28
system
Mq t)t+ F(q q t) = Q(q q t) (259)
where M(q t) is the nonlinear symmetric mass matrix and F(q q t) is the forcing
function which can be written in matrix form as
Q(ltfgt Qi 0 is the vector of non-conservative generalized forces including control inshy
puts acting on the system A detailed expansion of Eq(260) in tensor notation is
developed in Appendix I
Numerical solution of Eq(259) in terms of q would require inversion of the
mass matrix However M is a full matrix of size nqq x nqq where nqq = Nnq is the
total number of generalized coordinates Therefore direct inversion of M would lead
to a large number of computation steps of the order nqq^ or higher The objective
here is to develop an order-N O(N) algorithm for obtaining M _ 1 that minimizes
the computational effort This is accomplished through the following transformations
ini t ia l ly developed for the stationkeeping case by Pradhan et al [30]
232 Generalized coordinates and position transformation
During the derivation of the energy expressions for the system the focus was
on the decoupled system ie each link was considered independent of the others
Thus each links energy expression is also uncoupled However the constraints forces
between two links must be incorporated in the final equations of motion
Let
e Unqxl (261) m
be the set of coupled coordinates of the t h l ink such that q = qT q2 bull bull bull 0T i s the
29
systems generalized coordinates Let qt- be the set of auxiliary decoupled coordinates
defined by Eq(232) such that qt = qj^ qj^ Qtj^1 bull The only difference between
qtj and q is the presence of D against d is the position of F wi th respect to
the inertial frame FQ whereas d is defined as the offset position of F relative to and
projected on the frame It can be expressed as d = dXidyidZi^ For the
special case of link 1 D = d
From Figure 2-1 can written as
Di = A - l + T i _ i J i _ i + di) + T V i S i - i f t - i + dXi)8i- (262)
Denoting KAj = $j(Zj + dx- ) Eq(262) can be rewritten in summation form as
i ampi = H ( T i - 1 ^ + T i - l K A i - l - l + T j - l ^ - l ) bull ( 2 - 6 3 )
3 = 1
Introducing the index substitution k mdash j mdash 1 into Eq(263)
i-1 D = ^2 (Tfc_ilt4 + TkKAk6k + Tkilk) + T^di (264)
k=l
since d = D TQ = 1^ and IQ DQ SQ are null vectors Recasting Eq(264) in
matr ix form leads to i-1
Qt Jfe=l
where
and
For the entire system
T fc-1 0 TkKAk
0 0 0 0 0 0 0 0 0 0 0 0 -
r T i - i 0 0 0 0 0 0 0 0 0 G
0 0 0 1
Qt = RpQ
G R n lt x l
(265)
(266)
G Unq x l (267)
(268)
30
where
RP
R 1 0 0 R R 2 0 R RP
2 R 3
0 0 0
R RP R PN _ 1 RN J
(269)
Here RP is the lower block triangular matrix which relates qt and q
233 Velocity transformations
The two sets of velocity coordinates qti and qi are related by the following
transformations
(i) First Transformation
Differentiating Eq(262) with respect to time gives the inertial velocity of the
t h l ink as
D = A - l + Ti_iK + + K A i _ i ^ _ i
+ T j _ i ^ + Ti-ii + T i - i K A i - i ^ + T j - i K A ^ i ^ i (270)
(271) (
Defining K A D = dddXi+1KAi K A L = d d K A and hi = di+1 +lfi +KA^i
results in K A ^ i = K A D j i o ^ + K A L i _ i i _ i
= KADi_iiTdi + K A L j _ i ^ _ i
Inserting Eq(271) into Eq(270) and rearranging gives
= A _ + T i ^ d g + K A D ^ x ^ i ^ ^ + Pi-^hi-iH-i
+ T i _ i K A i _ i 5 _ i + T i_ i2 + K A L j _ i 5 j _ i 4 _ i (272)
Using Eq(272) to define recursively - D j _ i and applying the index substitution
c = j mdash 1 as in the case of the position transformation it follows that
Di = ^ T ^ i K D ^ ^ + Pk(hk)ffk + T f c K A f c ^ + TfcKLfc4 Jfc=l (273)
+ T j _ i K D j _ i lt i ii
31
where K D j _ i = + K A D ^ i ^ - i r 7 and K L j = i + K A L j 5 j Finally by following a
procedure similar to that used for the position transformation it can be shown that
for the entire system
t = Rvil (2-74)
where Rv is given by the lower block diagonal matrix
Here
Rv
Rf 0 0 R R$ 0 R Rl RJ
0 0 0
T3V nN-l
R bullN
e Mnqqxl
Ri =
T i _ i K D i _ i TiKAi T j K L j 0 0 0 0 bullo 0 0 0 0 0 0 0
e K n lt x l
RS
T i-1 0 0 o-0 0 0 0 0 0
0 0 0 1
G Mnq x l
(275)
(276)
(277)
(ii) Second Transformation
The second transformation is simply Eq(272) set into matrix form This leads
to the expression for qt as
(278)
where Id Pihi) T j K A j T j K L j 0 0 0 0 0 0
0 0
0 0
0 0
G $lnq x l
Thus for the entire system
(279)
pound = Rnjt + Rdq
= Vdnqq - R nrlRdi (280)
32
where
and
- 0 0 0 RI 0 0
Rn = 0 0
0 0
Rf 0 0 0 rgtd K2 0
Rd = 0 0 Rl
0 0 0
RN-1 0
G 5R n w x l
0 0 0 e 5 R N ^ X L
(281)
(282)
234 Cylindrical coordinate modification
Because of the use of cylindrical coorfmates for the first body it is necessary
to slightly modify the above matrices to account for this coordinate system For the
first link d mdash D = D ^ D S L thus
Qti = 01 =gt DDTigi (283)
where
D D T i
Eq(269) now takes the form
DTL 0 0 0 d 3 0 0 0 ID
0 0 nfj 0
RP =
P f D T T i R2 0 P f D T T i Rp
2 R3
iJfDTTi i^
x l
0 0 0
R PN-I RN
(284)
(285)
In the velocity transformation a similar modification is necessary since D =
DD^DS^ Mak ing the substitution it can be shown that Eq(276) and Eq(279) now
33
takes the form for i = 1
T)V _ pn _ r t j mdash rt^ mdash 0 0 0 0
0 0 0 0 0 0 0 0
(286)
The rest of the matrices for the remaining links ie i = 2 N are unchanged
235 Mass matrix inversion
Returning to the expression for kinetic energy substitution of Eq(274) into
Eq(235) gives the first of two expressions for kinetic energy as
Ke = -q RvT MtRv (287)
Introduction of Eq(280) into Eq(235) results in the second expression for Ke as
1X 2q (hnqq-RnrlRd Mt (idnqq - R nrlRd (288)
Thus there are two factorizations for the systems coupled mass matrix given as
M = RV MtRv
M (hnqq - RnrlRd Mt ( i d n q q - R n r 1 R d
(289)
(290)
Inverting Eq(290)
r-1 _ ( rgtd~^ M 1 = (Ra) l d n q q - Rn)Mf1 (Rd)~l(Idnqq - Rn) (291)
Since both Rd and Mt are block diagonal matrices their inverse is simply the inverse
34
of each block on the diagonal ie
Mr1
h 0 0
0 0
0 0
0 0 0
0
0
0
tN
(292)
Each block has the dimension nq x nq and is independent of the number of bodies
in the system thus the inversion of M is only linearly dependent on N ie an O(N)
operation
236 Specification of the offset position
The derivation of the equations of motion using the O(N) formulation requires
that the offset coordinate d be treated as a generalized coordinate However this
offset may be constrained or controlled to follow a prescribed motion dictated either by
the designer or the controller This is achieved through the introduction of Lagrange
multipliers[51] To begin with the multipliers are assigned to all the d equations
Thus letting f = F - Q Eq(259) becomes
Mq + f = P C A (293)
where A G K 3 - 7 V x l is the vector of Lagrange multipliers and Pc G s R n 9 x 3 A r is the
permutation matrix assigning the appropriate Aj to its corresponding d equation
Inverting M and pre-multiplying both sides by PcT gives
pcTM-pc
= ( dc + PdegTM-^f) (294)
where dc is the desired offset acceleration vector Note both dc and PcTM lf are
known Thus the solution for A has the form
A pcTM-lpc - l dc + PC1 M (295)
35
Now substituting Eq(295) into Eq(293) and rearranging the terms gives
q = S(f Ml) = -M~lf + M~lPc
which is the new constrained vector equation of motion with the offset specified by
dc Note that pre-multiplying M~l by P c T provides the rows of M _ 1 corresponding
to the d equations Similarly post-multiplying by Pc leads the columns of the matrix
corresponding to the d equations
24 Generalized Control Forces
241 Preliminary remarks
The treatment of non-conservative external forces acting on the system is
considered next In reality the system is subjected to a wide variety of environmental
forces including atmospheric drag solar radiation pressure Earth s magnetic field as
well as dissipative devices to mention a few However in the present model only
the effect of active control thrusters and momemtum-wheels is considered These
generalized forces wi l l be used to control the attitude motion of the system
In the derivation of the equations of motion using the Lagrangian procedure
the contribution of each force must be properly distributed among all the generalized
coordinates This is acheived using the relation
nu ^5 Qk=E Pern ^ (2-97)
m=l deg Q k
where Qk and qk are the kth elements of Q and ltf respectively[51] Fem is the
mth external force acting on the system at the location Rm from the inertial frame
and nu is the total number of actuators Derivation of the generalized forces due to
thrusters and momentum wheels is given in the next two subsections
pcTM-lpc (dc + PcTM- Vj (2-96)
36
242 Generalized thruster forces
One set of thrusters is placed on each subsatellite ie al l the rigid bodies
except the first (platform) as shown in Figure 2-3 They are capable of firing in the
three orthogonal directions From the schematic diagram the inertial position of the
thruster located on body 3 is given as
Defining
R3 = di + Tid2 + T2(hi + KA2lt52) + T 3 f c m 3
= d + dfY + T 3 f c m 3
df = Tjdj+i + Tj+1lj+ii + KAj+i5j+i)
the thruster position for body 5 is given as
(298)
(299)
R5 = di+ dfx + d 3 + T5fcm
Thus in general the inertial position for the thruster on the t h body can be given
(2100)
as i - 2
k=l i1 cm (2101)
Let the column matrix
[Q i iT 9 R-K=rR dqj lA n i
WjRi A R i (2102)
where i = 3 5 N and j mdash 1 2 i The vector derivative of R wi th respect to
the scaler components qk is stored in the kth column of Thus for the case of
three bodies Eq(297) for thrusters becomes
1 Q
Q3
3 -1
T 3 T t 2 (2103)
37
Satellite 3
mdash gt -raquo cm
D1=d1
mdash gt
a mdash gt
o
Figure 2-3 Inertial position of subsatellite thruster forces
38
where ft = 0TtaTt^T is the thrust vector for the t h l ink (tether i-1) Here
Ttn and Tta are the thrust forces along the inplane m and out-of-plane Z directions i Pi
respectively For the case of 5-bodies
Q
~Qgt1
0
0
( T ^ 2 ) (2104)
The result for seven or more bodies follows the pattern established by Eq(2104)
243 Generalized momentum gyro torques
When deriving the generalized moments arising from the momemtum-wheels
on each rigid body (satellite) including the platform it is easier to view the torques as
coupled forces From Figure 2-4 it is apparent that for the ith l ink the generalized
forces constituting the couple are
(2105)
where
Fei = Faij acting at ei = eXi
Fe2 = -Fpk acting at e 2 = C a ^ l
F e 3 = F7ik acting at e 3 = eyj
Expanding Eq(2105) it becomes clear that
3
Qk = ^ F e i
i=l
d
dqk (2106)
which is independent of Ri- Thus the moments Mma = 2FaieXi Mm^ = 2FpeXi
and Mmi = 2FlieVi Transforming the coordinates to the body fixed frame using
the rotation matrix the generalized force due to the momentum-wheels on link i can
39
Figure 2-4 Coupled force representation of a momentum-wheel on a rigid body
40
be written as
d d d Qi = Td bull ^ T i t M m a - Tik bull mdashT^Mmp + T ^ bull T J j M ^
QiMmi
where Mm = Mma M m Mmi T
lt3 T J ^ T - T i f c bull ^ T i pound T ^ - ^ T z j
(2107)
(2108)
Note combining the thruster forces and momentum-wheel torques a compact
expression for the generalized force vector for 5 bodies can be written as
Q
Q T 3 0 Q T 5 0
0 0 0
0 lt33T3 Qyen Q5
3T5 0
0 0 0 Q4T5 0
0 0 0 Q5
5T5
M M I
Ti 2 m 3
Quu (2109)
The pattern of Eq(2109) is retained for the case of N links
41
3 C O M P U T E R I M P L E M E N T A T I O N
31 Prel iminary Remarks
The governing equations of motion for the TV-body tethered system were deshy
rived in the last chapter using an efficient O(N) factorization method These equashy
tions are highly nonlinear nonautonomous and coupled In order to predict and
appreciate the character of the system response under a wide variety of disturbances
it is necessary to solve this system of differential equations Although the closed-
form solutions to various simplified models can be obtained using well-known analytic
methods the complete nonlinear response of the coupled system can only be obtained
numerically
However the numerical solution of these complicated equations is not a straightshy
forward task To begin with the equations are stiff[52] be they have attitude and
vibration response frequencies separated by about an order of magnitude or more
which makes their numerical integration suceptible to accuracy errors i f a properly
designed integration routine is not used
Secondly the factorization algorithm that ensures efficient inversion of the
mass matrix as well as the time-varying offset and tether length significantly increase
the size of the resulting simulation code to well over 10000 lines This raises computer
memory issues which must be addressed in order to properly design a single program
that is capable of simulating a wide range of configurations and mission profiles It
is further desired that the program be modular in character to accomodate design
variations with ease and efficiency
This chapter begins with a discussion on issues of numerical and symbolic in-
42
tegration This is followed by an introduction to the program structure Final ly the
validity of the computer program is established using two methods namely the conshy
servation of total energy and comparison of the response results for simple particular
cases reported in the literature
32 Numerical Implementation
321 Integration routine
A computer program coded in F O R T R A N was written to integrate the equashy
tions of motion of the system The widely available routine used to integrate the difshy
ferential equations D G E A R [ 5 3 ] is based on Gears predictor-corrector method[54]
It is well suited to stiff differential equations as i t provides automatic step-size adjustshy
ment and error-checking capabilities at each iteration cycle These features provide
superior performance over traditional 4^ order Runge-Kut ta methods
Like most other routines the method requires that the differential equations be
transformed into a first order state space representation This is easily accomplished
by letting
thus
X - ^ ^ f ^ ) =ltbullbull) (3 -2 )
where S(fMdc) is defined by Eq(296) Eq(32) represents the new set of 2nqq
first order differential equations of motion
322 Program structure
A flow chart representing the computer programs structure is presented in
Figure 3-1 The M A I N program is responsible for calling the integration routine It
also coordinates the flow of information throughout the program The first subroutine
43
called is INIT which initializes the system variables such as the constant parameters
(mass density inertia etc) and the ini t ial conditions I C (attitude angles orbital
motion tether elastic deformation etc) from user-defined data files Furthermore all
the necessary parameters required by the integration subroutine D G E A R (step-size
tolerance) are provided by INIT The results of the simulation ie time history of the
generalized coordinates are then passed on to the O U T P U T subroutine which sorts
the information into different output files to be subsequently read in by the auxiliary
plott ing package
The majority of the simulation running time is spent in the F C N subroutine
which is called repeatedly by D G E A R F C N generates the fv(Xi) vector defined in
Eq(32) It is composed of several subroutines which are responsible for the comshy
putation of all the time-varying matrices vectors and scaler variables derived in
Chapter 2 that comprise the governing equations of motion for the entire system
These subroutines have been carefully constructed to ensure maximum efficiency in
computational performance as well as memory allocation
Two important aspects in the programming of F C N should be outlined at this
point As mentioned earlier some of the integrals defined in Chapter 2 are notably
more difficult to evaluate manually Traditionally they have been integrated numershy
ically on-line (during the simulation) In some cases particularly stationkeeping the
computational cost is quite reasonable since the modal integrals need only be evalushy
ated once However for the case of deployment or retrieval where the mode shape
functions vary they have to be integrated at each time-step This has considerable
repercussions on the total simulation time Thus in this program the integrals are
evaluated symbolically using M A P L E V[55] Once generated they are translated into
F O R T R A N source code and stored in I N C L U D E files to be read by the compiler
Al though these files can be lengthy (1000 lines or more) for a large number of flexible
44
INIT
DGEAR PARAM
IC amp SYS PARAM
MAIN
lt
DGEAR
lt
FCN
OUTPUT
STATE ACTUATOR FORCES
MOMENTS amp THRUST FORCES
VIBRATION
r OFFS DYNA
ET MICS
Figure 3-1 Flowchart showing the computer program structure
45
modes they st i l l require less time to compute compared to their on-line evaluation
For the case of deployment the performance improvement was found be as high as
10 times for certain cases
Secondly the O(N) algorithm presented in the last section involves matrices
that contained mostly zero terms (RP Rv and Mplusmn) The sparse structure of these mashy
trices can be used advantageously to improve the computational efficiency by avoiding
multiplications of zero elements The result is a quick evaluation of fv(Xt) Moreshy
over there is a considerable saving of computer memory as zero elements are no
longer stored
Final ly the C O N T R O L subroutine which is part of F C N is designed to calshy
culate the appropriate control forces torques and offset dynamics which are required
to stabilize the librational ( A T T I T U D E ) and vibrational ( V I B R A T I O N ) motion of
the system
33 Verification of the Code
The simulation program was checked for its validity using two methods The
first verified the conservation of total energy in the absence of dissipation The second
approach was a direct comparison of the planar response generated by the simulation
program with those available in the literature A s numerical results for the flexible
three-dimensional A^-body model are not available one is forced to be content with
the validation using a few simplified cases
331 Energy conservation
The configuration considered here is that of a 3-body tethered system with
the following parameters for the platform subsatellite and tether
46
h =
1091430 -8135 328108 -8135 8646050 27116 328108 27116 8286760
kg bull m 2 platform inertia
kg bull m 2 end-satellite inertia 200 0 0
0 400 0 0 0 400
m = 90000 kg mass of the space station platform
m2 = 500 kg mass of the end-satellite
EfAt = 61645 N tether elastic stiffness
pt mdash 49 kg km tether linear density
The tether length is taken to remain constant at 10 km (stationkeeping case)
A n in i t ia l disturbance in pitch and roll of 2deg and 1deg respectively is given to both
the rigid bodies and the flexible tether In addition the tether is ini t ia l ly deflected
by 45 m in the longitudinal direction at its end and 05 m in both the inplane and
out-of-plane directions in the first mode The tether attachment point offset at the
platform (satellite 1) end is 1 m in all the three directions ie d2 = 11 lTm The
attachment point at the subsatellite (satellite 2) end is 1 m in the x direction ie
fcm3 = 10 0T
The variation in the kinetic and potential energies is presented for the above-
mentioned case in Figure 3-2(a) As seen from the plot there is a continuous mutual
exchange between the kinetic and potential energies however the variation in the
total energy remains zero (Figure 3-2b) It should be noted that after 1 orbit the
variation of both kinetic and potential energy does not return to 0 This is due to
the attitude and elastic motion of the system which shifts the centre of mass of the
tethered system in a non-Keplerian orbit
Similar results are presented for the 5-body case ie a double pendulum (Figshy
ure 3-3) Here the system is composed of a platform and two subsatellites connected
47
STSS configuration 3-Body a 1(0)=a 2(0)=a t(0)=2deg P1(0)=p2(0)=p(0)=1deg 5x(0)=45m 8y(0)=52(0)=05m dx(0)=dy(0)=d2(0)=1m Stationkeeping lt=10km
Variation of Kinetic and Potential Energy
(a) Time (Orbits) Percent Variation in Total Energy
OOEOh
-50E-12
-10E-11 I i i
0 1 2 (b) Time (Orbits)
Figure 3-2 Kinetic and potential energy transfer for the three-body platform based tethered satellite system (a) variation of kinetic and potential energy (b) percent change in total energy of system
48
STSS configuration 5-Body a 1(0)=a 2(0)=a t 1(0)=2deg a 3(0)=a t 2(0)=25deg P1(0)=p2(0)=p3(0)=pt1(0)=pt2(0)=1lt
5x(0)=5m 8y(0)=5z(0)=05m dx(0)=dy(0)=dz(0)=0 Stationkeeping 1 =12=1 Okm
(a)
20E6
10E6r-
00E0
-10E6
-20E6
Variation of Kinetic and Potential Energy
-
1 I I I 1 I I
I I
AP e ~_
i i 1 i i i i
00E0
UJ -50E-12 U J
lt
Time (Orbits) Percent Variation in Total Energy
-10E-11h
(b) Time (Orbits)
Figure 3 -3 Kinet ic and potential energy transfer for the five-body platform based tethered satellite system (a) variation of kinetic and potential enshyergy (b) percent change in total energy of system
49
in sequence by two tethers The two subsatellites have the same mass and inertia
ie 7713 = 7 7 i 2 a n d I3 = I21 a n d are separated by identical tether segments each 10
k m in length The platform has the same properties as before however there is no
tether attachment point offset present in the system (d = fcmi = 0)
In al l the cases considered energy was found to be conserved However it
should be noted that for the cases where the tether structural damping deployment
retrieval attitude and vibration control are present energy is no longer conserved
since al l these features add or subtract energy from the system
332 Comparison with available data
The alternative approach used to validate the numerical program involves
comparison of system responses with those presented in the literature A sample case
considered here is the planar system whose parameters are the same as those given
in Section 331 A n ini t ia l pitch disturbance of 2deg is imparted to both the platform
and the tether together with an ini t ial tether deformation of 08 m and 001 m in
the longitudinal and transverse directions respectively in the same manner as before
The tether length is held constant at 5 km and it is assumed to be connected to the
centre of mass of the rigid platform and subsatellite The response time histories from
Ref[43] are compared with those obtained using the present program in Figures 3-4
and 3-5 Here ap and at represent the platform and tether pitch respectively and B i
C i are the tethers longitudinal and transverse generalized coordinates respectively
A comparison of Figures 3-4 and 3-5 clearly shows a remarkable correlation
between the two sets of results in both the amplitude and frequency The same
trend persisted even with several other comparisons Thus a considerable level of
confidence is provided in the simulation program as a tool to explore the dynamics
and control of flexible multibody tethered systems
50
a p(0) = 2deg 0(0) = 2C
6(0) = 08 m
0(0) = 001 m
Stationkeeping L = 5 km
D p y = 0 D p 2 = 0
a oo
c -ooo -0 01
Figure 3 - 4 Simulation results for the platform based three-body tethered system originally presented in Ref[43]
51
STSS configuration 3-Body a1(0)=a2(0)=at(0)=2deg p1(0)=(32(0)=(3t(0)=0 8x(0)=08m 5y(0)=001m 5Z(0)=0 dx(0)=dy(0)=dz(0)=0 Stationkeeping l=5km
Satellite 1 Pitch Angle Tether Pitch Angle
i -i i- i i i i i i d
1 2 0 1 2 Time (Orbits) Time (Orbits)
Longitudinal Vibration
^MAAAAAAAWVWWWWVVVW
V bull i bull i I i J 00 05 10 15 20
Time (Orbits) Transverse Vibration
Time (Orbits)
Figure 3-5 Simulation results for the platform based three-body tethered system obtained using the present computer program
52
4 D Y N A M I C SIMULATION
41 Prel iminary Remarks
Understanding the dynamics of a system is critical to its design and develshy
opment for engineering application Due to obvious limitations imposed by flight
tests and simulation of environmental effects in ground based facilities space based
systems are routinely designed through the use of numerical models This Chapter
studies the dynamical response of two different tethered systems during deployment
retrieval and stationkeeping phases In the first case the Space platform based Tethshy
ered Satellite System (STSS) which consists of a large mass (platform) connected to
a relatively smaller mass (subsatellite) with a long flexible tether is considered The
other system is the O E D I P U S B I C E P S configuration involving two comparable mass
satellites interconnected by a flexible tether In addition the mission requirement of
spin about an arbitrary axis the tether axis for O E D I P U S - A C and cartwheeling or
spin about the orbit normal for the proposed B I C E P S mission is accounted for
A s mentioned in Chapter 2 the tether flexibility is modelled using the asshy
sumed mode discretization method Although the simulation program can account
for an arbitrary number of vibrational modes in each direction only the first mode
is considered in this study as it accounts for most of the strain energy[34] and hence
dominates the vibratory motion The parametric study considers single as well as
double pendulum type systems with offset of the tether attachment points The
systems stability is also discussed
42 Parameter and Response Variable Definitions
The system parameters used in the simulation unless otherwise stated are
53
selected as follows for the STSS configuration
1091430 -8135 328108 -8135 8646050 27116 328108 27116 8286760
h =
h = 200 0 0
0 400 0 0 0 400
bull m = 90000 kg (mass of the space platform)
kg bull m 2 (platform inertia) 760 J
kg bull m 2 (subsatellite inertia)
bull 7772 = 500 kg (mass of the subsatellite)
bull EtAt = 61645 N (tether stiffness)
bull Pt = 49 k g k m (tether density)
bull Vd mdash 05 (tether structural damping coefficient)
bull fcm^ = l 0 0 r m (tether attachment point at the subsatellite)
The response variables are defined as follows
bull ai0 satellite 1 (platform) pitch and roll angles respectively
bull CK2 2 satellite 2 (subsatellite) pitch and roll angles respectively
bull at fit- tether pitch and roll angles respectively
bull If tether length
bull d = dx dy dzT- tether attachment position relative to satellite 1 (platform)
mdash
bull S = 8X 8y Sz tether flexible generalized coordinates in the longitudinal x
inplane transverse y and out-of-plane transverse z directions respectively
The attitude angles and are measured with respect to the Local Ve r t i c a l -
Loca l Horizontal ( L V L H ) frame A schematic diagram illustrating these variable
is presented in Figure 4-1 The system is taken to be in a nominal circular orbit at
an altitude of 289 km with an orbital period of 903 minutes
54
Satellite 1 (Platform) Y a w
Figure 4-1 Schematic diagram showing the generalized coordinates used to deshyscribe the system dynamics
55
43 Stationkeeping Profile
To facilitate comparison of the simulation results a reference case based on
the three-body STSS system with zero offset (d = 0 r c m 3 = 0) and a tether length
of It = 20 km is considered first (Figure 4-2) The system is subjected to an ini t ial
disturbance in pitch and roll of 2deg and 1deg respectively to all the three bodies
In addition an ini t ia l longitudinal deflection of 14 m from the tethers unstretched
position is given together with a transverse inplane and out-of-plane deflection of 1
m at the tethers mid-length From the attitude response given in Figure 4-2(a) it is
apparent that the three bodies oscillate about their respective equilibrium positions
Since coupling between the individual rigid-body dynamics is absent (zero offset) the
corresponding librational frequencies are unaffected Satellite 1 (Platform) displays
amplitude modulation arising from the products of inertia The result is a slight
increase of amplitude in the pitch direction and a significantly larger decrease in
amplitude in the roll angle
Figure 4-2(b) clearly shows coupling of the tethers attitude motion with its
flexibility dynamics The tether vibrates in the axial direction about its equilibrium
position (at approximately 138 m) with two vibration frequencies namely 012 Hz
and 31 x 1 0 - 4 Hz The former is the first longitudinal flexible mode which is dissishy
pated through the structural damping while the latter arises from the coupling with
the pitch motion of the tether Similarly in the transverse direction there are two
frequencies of oscillations arising from the flexible mode and its coupling wi th the
attitude motion As expected there is no apparent dissipation in the transverse flexshy
ible motion This is attributed to the weak coupling between the longitudinal and
transverse modes of vibration since the transverse strain in only a second order effect
relative to the longitudinal strain where the dissipation mechanism is in effect Thus
there is very litt le transfer of energy between the transverse and longitudinal modes
resulting in a very long decay time for the transverse vibrations
56
STSS configuration 3-Body a1(0)=a2(0)=a(0)=2deg (31(0)=p2(0)=p(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=dz(0)=0 Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
0 1 2 3 4 5 0 1 2 3 4 5 Time (Orbits) Time (Orbits)
Satellite 2 Pitch Angle Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-2 Stationkeeping dynamics of the three-body STSS configuration withshyout offset (a) attitude response
57
STSS configuration 3-Body a1(0)=cx2(0)=cx(0)=2deg P1(0)=p2(0)=p(0)=1deg 5x(0)=14m8y(0)=82(0)=1m dx(0)=dy(0)=dz(0)=0 Stationkeeping lt=20km
Tether Longitudinal Vibration 16| 1 1 1 1 1 r-
Time (Orbits) Tether In-Plane Transverse Vibration
imdash 1 mdash 1 mdash 1 mdash 1 mdashr
r I I 1 1 1 I 1 1 1 1 I 1 bull ^ bull bull I I J 0 1 2 3 4 5
Time (Orbits) Tether Out-of-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-2 Stationkeeping dynamics of the three-body STSS configuration withshyout offset (b) vibration response
58
Introduction of the attachment point offset significantly alters the response of
the attitude motion for the rigid bodies W i t h a i m offset along the local vertical
at the platform (dx = 1 m) and fcm^ = 100 m coupling between the tether
and end-bodies is established by providing a lever arm from which the tether is able
to exert a torque that affects the rotational motion of the rigid end-bodies (Figure
4-3) Bo th the platform and subsatellite now oscillate at the same frequency as the
tether whose motion remains unaltered In the case of the platform there is also an
amplitude modulation due to its non-zero products of inertia as explained before
However the elastic vibration response of the tether remains essentially unaffected
by the coupling
Providing an offset along the local horizontal direction (dy = 1 m) results in
a more dramatic effect on the pitch and roll response of the platform as shown in
Figure 4-4(a) Now the platform oscillates about its new equilibrium position of
-90deg The roll motion is also significantly disturbed resulting in an increase to over
10deg in amplitude On the other hand the rigid body dynamics of the tether and the
subsatellite as well as the tether flexible motion remain the same (Figure 4-4b)
Figure 4-5 presents the response when the offset of 1 m at the platform end is
along the z direction ie normal to the orbital plane The result is a larger amplitude
pitch oscillation about the reference equilibrium position as shown in Figure 4-5(a)
while the roll equilibrium is now shifted to approximately 90deg Note there is little
change in the attitude motion of the tether and end-satellites however there is a
noticeable change in the out-of-plane transverse vibration of the tether (Figure 4-5b)
which may be due to the large amplitude platform roll dynamics
Final ly by setting a i m offset in all the three direction simultaneously the
equil ibrium position of the platform in the pitch and roll angle is altered by approxishy
mately 30deg (Figure 4-6a) However the response of the system follows essentially the
59
STSS configuration 3-Body a1(0)=cx2(0)=cxt(0)=2deg p1(0)=(32(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=1m dy(0)=dz(0)=0 Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
Time (Orbits) Time (Orbits) Tether Pitch Angle Tether Roll Angle
L bull I I bull bull lt bull bull bull raquo bull bull bull bull -I h I I I I i H 0 1 2 3 4 5 0 1 2 3 4 5
Time (Orbits) Time (Orbits) Satellite 2 Pitch Angle Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-3 Stationkeeping dynamics of the three-body STSS configuration with offset along the local vertical (a) attitude response
60
STSS configuration 3-Body a 1(0)=a 2(0)=a t(0)=2deg (31(0)=p2(0)=pt(0)=1deg 5x(0)=14m5y(0)=5z(0)=1m dx(0)=1mdy(0)=dz(0)=0 Stationkeeping lt=20km
Tether Longitudinal Vibration 1 6 i mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash mdash i mdash i mdash mdash mdash mdash mdash i mdash mdash lt -
Time (Orbits)
Tether In-Plane Transverse Vibration
y bull i i i i i i i i i i i i i i i i i 0 1 2 3 4 5
Time (Orbits)
Tether Out-of-Plane Transverse Vibration r 1 1 1 1 1 1 1 1 1 1 1 1 1 1 q
h i I i i I i i i i I i i i i 3
0 1 2 3 4 5 (b) Time (Orbits)
Figure 4-3 Stationkeeping dynamics of the three-body STSS configuration with offset along the local vertical (b) vibration response
61
STSS configuration 3-Body a1(0)=cx2(0)=o(0)=2deg (31(0)=p2(0)=p(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dz(0)=0 dy(0)=1m Stationkeeping lt=20km
Satellite 1 Pitch Angle i 1 i
Satellite 1 Roll Angle
1 2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 4 Time (Orbits)
Tether Roll Angle
CD
edT
1 2 3 Time (Orbits)
Satellite 2 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 2 Roll Angle
CD
C O
(a)
Figure 4-
1 2 3 4 Time (Orbits)
1 2 3 4 Time (Orbits)
4 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal (a) attitude response
62
STSS configuration 3-Body a1(0)=cx2(0)=ot(0)=2deg p1(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dz(0)=0 dy(0)=1m Stationkeeping lt=20km
Tether Longitudinal Vibration 16| 1 gt 1 bull 1 1 1
0 1 2 3 4 5
Time (Orbits) Tether In-Plane Transverse Vibration
i mdash 1 mdash 1 mdash 1 mdash 1 mdash i mdash mdash lt mdash 1 mdash 1 mdash i mdash mdash 1 mdash 1 mdash
V I I bull bull bull I i i i I i i 1 1 3
0 1 2 3 4 5 Time (Orbits)
Tether Out-of-Plane Transverse Vibration r 1 1 1 1 1 1 1 1 1 1 1 gt bull | q
V i 1 1 1 1 1 mdash 1 mdash 3
0 1 2 3 4 5 (b) Time (Orbits)
Figure 4-4 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal (b) vibration response
63
STSS configuration 3-Body a1(0)=o2(0)=at(0)=2deg (31(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=0 dz(0)=1m Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 4 Time (Orbits)
Tether Roll Angle
1 2 3 4 Time (Orbits)
Satellite 2 Pitch Angle
2 3 4 Time (Orbits)
Satellite 2 Roll Angle
(a)
Figure 4-
1 2 3 4 Time (Orbits)
2 3 4 Time (Orbits)
5 Stationkeeping dynamics of the three-body STSS configuration with offset along the orbit normal (a) attitude response
64
STSS configuration 3-Body a1(0)=cx2(0)=ot(0)=2deg P1(0)=p2(0)=(3(0)=1deg 8x(0)=14m8y(0)=82(0)=1m dx(0)=dy(0)=0 d2(0)=1m Stationkeeping lt=20km
Tether Longitudinal Vibration 1 6 i mdash r mdash i mdash mdash i mdash i mdash mdash i mdash i mdash mdash i mdash mdash i mdash 1 mdash 1 mdash i mdash mdash lt mdash lt mdash lt mdash r
Time (Orbits) Tether In-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-5 Stationkeeping dynamics of the three-body STSS configuration with offset along the orbit normal (b) vibration response
65
same trend with only minor perturbations in the transverse elastic vibratory response
of the tether (Figure 4-6b)
44 Tether Deployment
Deployment of the tether from an ini t ial length of 200 m to 20 km is explored
next Here the deployment length profile is critical to ensure the systems stability
It is not desirable to deploy the tether too quickly since that can render the tether
slack Hence the deployment strategy detailed in Section 216 is adopted The tether
is deployed over 35 orbits with the sinusoidal acceleration and deceleration occurring
over a 4 km length ( A ^ 2 ) with the remaining deployment at a constant velocity of
V0 = 146 ms
Initially the tether is stretched only slightly S = 1304 x 1 0 _ 3 0 0 m
with al l the other states of the system ie pitch roll and transverse displacements
remaining zero The response for the zero offset case is presented in Figure 4-7 For
the platform there is no longer any coupling with the tether hence it oscillates about
the equilibrium orientation determined by its inertia matrix However there is st i l l
coupling between the subsatellite and the tether since rcm^ mdash 100 m resulting
in complete domination of the subsatellite dynamics by the tether Consequently
the pitch motion ini t ial ly grows by almost 50deg but eventually subsides as the tether
deployment rate decreases It is interesting to note that the out-of-plane motion is not
affected by the deployment This is because the Coriolis force which is responsible
for the tether pitch motion does not have a component in the z direction (as it does
in the y direction) since it is always perpendicular to the orbital rotation (Q) and
the deployment rate (It) ie Q x It
Similarly there is an increase in the amplitude of the transverse elastic oscilshy
lations of the tether again due to the Coriolis effect (Figure 4-7b) Moreover as the
66
STSS configuration 3-Body a1(0)=a2(0)=cxt(0)=2deg P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m5y(0)=52(0)=1m dx(0)=dy(0)=d2(0)=1m Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
i i i i i i i i i i i 0 1 2 3 4 5 0 1 2 3 4 5
Time (Orbits) Time (Orbits) Tether Pitch Angle Tether Roll Angle
l i i i i -I h i i i i I i i i i l- i i i i l i i i i I i i- i i H 0 1 2 3 4 5 0 1 2 3 4 5
Time (Orbits) Time (Orbits) Satellite 2 Pitch Angle Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-6 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal local vertical and orbit normal (a) attitude response
67
STSS configuration 3-Body a1(0)=cc2(0)=a(0)=2deg p1(0)=p2(0)=(3t(0)=1deg 5x(0)=14m6y(0)=52(0)=1m dx(0)=dy(0)=d2(0)=1m Stationkeeping lt=20km
Tether Longitudinal Vibration 1 6 1 mdash i mdash i mdash i mdash mdash i mdash 1 mdash 1 mdash mdash mdash i mdash mdash 1 mdash 1 mdash 1 mdash i mdash 1 mdash 1 mdash 1 mdash 1 mdash r
0 1 2 3 4 5
Time (Orbits) Tether In-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-6 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal local vertical and orbit normal (b) vibration response
68
STSS configuration 3-Body a(0)=a2(0)=ot(0)=0 P1(0)=p2(0)=pt(0)=0 8X(0)=1304 x 103m 8y(0)=8z(0)=0 dx(0)=dy(0)=dz(0)=0 Deployment
l=02km to 20km in 35 Orbits
Tether Length Profile
E
Satellite 1 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 1 Roll Angle mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash [ mdash i mdash i mdash i mdash i mdash | mdash
1 2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 4 Time (Orbits)
Tether Roll Angle
CD
oo ax
i imdashimdashimdashImdashimdashimdashimdashr-
_ 1 _ L
1 2 3 4 Time (Orbits)
Satellite 2 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 2 Roll Angle
-50
(a) 1 2 3 4
Time (Orbits) 1 2 3 4
Time (Orbits)
Figure 4-7 Deployment dynamics of the three-body STSS configuration without offset (a) attitude response
69
STSS configuration 3-Body a1(0)=a2(0)=at(0)=0 p1(0)=p2(0)=(3t(0)=0 8X(0)=1304x 103m 8y(0)=52(0)=0 dx(0)=dy(0)=d2(0)=0 Deployment
l=02km to 20km in 35 Orbits
Tether Longitudinal Vibration
to
-02
(b)
Figure 4-7
2 3 Time (Orbits)
Tether In-Plane Transverse Vibration
0 1 1 bull 1 1
1 2 3 4 Time (Orbits)
Tether Out-of-Plane Transverse Vibration
5
1 1 1 1 1 i i | i i | ^
_ i i i i L -I
1 2 3 4 5 Time (Orbits)
Deployment dynamics of the three-body STSS configuration without offset (b) vibration response
70
tether elongates its longitudinal static equilibrium position also changes due to an
increase in the gravity-gradient tether tension It may be pointed out that as in the
case of attitude motion there are no out-of-plane tether vibrations induced
When the same deployment conditions are applied to the case of 1 m offset in
the x direction (Figure 4-8) coupling effects are re-introduced The platform pitch
exceeds -100deg during the peak deployment acceleration However as the tether pitch
motion subsides the platform pitch response returns to a smaller amplitude oscillation
about its nominal equilibrium On the other hand the platform roll motion grows
to over 20deg in amplitude in the terminal stages of deployment Longitudinal and
transverse vibrations of the tether remain virtually unchanged from the zero offset
case except for minute out-of-plane perturbations due to the platform librations in
roll
45 Tether Retrieval
The system response for the case of the tether retrieval from 20 km to 200 m
with an offset of 1 m in the x and z directions at the platform end is presented in
Figure 4-9 Here the Coriolis force induced during retrieval renders the tether unstable
(Figure 4-9a) Consequently the pitch motion of the platform is also destabilized
through coupling The z offset provides an additional moment arm that shifts the
roll equilibrium to 35deg however this degree of freedom is not destabilized From
Figure 4-9(b) it is apparent that the transverse mode of the tether vibration is also
disturbed producing a high frequency response in the final stages of retrieval This
disturbance is responsible for the slight increase in the roll motion for both the tether
and subsatellite
46 Five-Body Tethered System
A five-body chain link system is considered next To facilitate comparison with
71
STSS configuration 3-Body a(0)=a2(0)=at(0)=0 P1(0)=P2(0)=P(0)=0 5X(0)=1304 x 103m 8y(0)=8z(0)=0 dx(0)=1mdy(0)=d2(0)=0 Deployment
lt=02km to 20km in 35 Orbits
Tether Length Profile
E
Satellite 1 Pitch Angle
1 2 3 Time (Orbits)
Satellite 1 Roll Angle
bull bull bull i bull bull bull bull i
1 2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 Time (Orbits)
Tether Roll Angle
i i i i J I I i _
1 2 3 4 Time (Orbits)
Satellite 2 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 2 Roll Angle 3E-3h
D )
T J ^ 0 E 0 eg
CQ
-3E-3
(a) 1 2 3 4
Time (Orbits) 1 2 3 4
Time (Orbits)
Figure 4-8 Deployment dynamics of the three-body STSS configuration with offset along the local vertical (a) attitude response
7 2
STSS configuration 3-Body a1(0)=a2(0)=at(0)=0 p1(0)=P2(0)=(3t(0)=0 Sx(0)=1304 x 103m 5y(0)=5z(0)=0 dx(0)=1mdy(0)=dz(0)=0 Deployment
l=02km to 20km in 35 Orbits
Tether Longitudinal Vibration
Time (Orbits) Tether In-Plane Transverse Vibration
n 1 1 1 f
I I I _j I I I I I 1 1 1 1 1 2 3 4 5
Time (Orbits) Tether Out-of-Plane Transverse Vibration
J 1 I I 1 I I I I I I I I L _ _ I I d 1 2 3 4 5
Time (Orbits)
Deployment dynamics of the three-body STSS configuration with offset along the local vertical (b) vibration response
(b)
Figure 4-8
73
STSS configuration 3-Body a1(0)=a2(0)=at(0)=0 p(0)=P2(0)=(3t(0)=0 Sx(0)=14m8y(0)=82(0)=0 dy(0)=0dx(0)=d2(0)=1m Retrieval
l=20km to 02km in 35 Orbits
Satellite 1 Pitch Angle
(a) Time (Orbits)
Figure 4-9 Retrieval dynamics of the along the local vertical ar
Tether Length Profile
i i i i i i i i i i I 0 1 2 3
Time (Orbits) Satellite 1 Roll Angle
Time (Orbits) Tether Roll Angle
b_ i i I i I i i L _ J
0 1 2 3 Time (Orbits)
Satellite 2 Roll Angle _ mdash | r i i | 1 i i imdashmdashj i 1 r mdash i mdash
04- -
0 1 2 3 Time (Orbits)
three-body STSS configuration with offset d orbit normal (a) attitude response
74
STSS configuration 3-Body a1(0)=a2(0)=ocl(0)=0 (31(0)=f32(0)=pt(0)=0 8x(0)=14m8y(0)=6z(0)=0 dy(0)=0dx(0)=dz(0)=1m Retrieval
l=20km to 02km in 35 Orbits
Tether Longitudinal Vibration
E 10 to
Time (Orbits) Tether In-Plane Transverse Vibration
1 2 Time (Orbits)
Tether Out-of-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-9 Retrieval dynamics of the three-body STSS configuration with offset along the local vertical and orbit normal (b) vibration response
75
the three-body system dynamics studied earlier the chain is extended through the
addition of two bodies a tether and a subsatellite with the same physical properties
as before (Section 42) Thus the five-body system consists of a platform (satellite 1)
tether 1 subsatellite 1 (satellite 2) tether 2 and subsatellite 2 (satellite 3) as shown
in Figure 4-10 The offset of tether 1 at the platform-end is denoted by d2 while that
of tether 2 to subsatellite 1 as 04 In this numerical example fcm^ = fcm^ = 0 ie
tethers 1 and 2 are attached to the centre of mass of subsatellites 1 and 2 respectively
The system response for the case where the tether attachment points coincide
wi th the centres of mass of the rigid bodies ie zero offset (d2 = d mdash 0) is presented
in Figure 4-11 Here ogtt and at2 represent the pitch motion of tethers 1 and 2
respectively whereas 03 is the pitch motions of satellite 3 A similar convention
is adopted for the rol l angle 0 A s before the zero offset eliminates the coupling
between the tethers and satellites such that they are now free to librate about their
static equil ibrium positions However their is s t i l l mutual coupling between the two
tethers This coupling is present regardless of the offset position since each tether is
capable of transferring a force to the other The coupling is clearly evident in the pitch
response of the two tethers (Figure 4 - l l a ) O n the other hand the roll motion appears
uncoupled as the in i t ia l conditions in rol l for the two tethers are identical Thus the
motion is in phase and there is no transfer of energy through coupling A s expected
during the elastic response the tethers vibrate about their static equil ibrium positions
and mutually interact (Figure 4 - l l b ) However due to the variation of tension along
the tether (x direction) they do not have the same longitudinal equilibrium Note
relatively large amplitude transverse vibrations are present particularly for tether 1
suggesting strong coupling effects with the longitudinal oscillations
76
Figure 4-10 Schematic diagram of the five-body system used in the numerical example
77
STSS configuration 5-Body a1(0)=(x2(0)=at1(0)=2deg a3(0)=a12(0)=25deg p1(0)=r32(0)=p3(0)=pt1(0)=pt2(0)=r 82(0)=84(0)=505)05m d2(0)=d4(0)=000 Stationkeeping lt1=lu=10km Pitch Roll
Satellite 1 Pitch Angle Satellite 1 Roll Angle
o 1 2 Time (Orbits)
Tether 1 Pitch and Roll Angle
0 1 2 Time (Orbits)
Tether 2 Pitch and Roll Angle
Time (Orbits) Satellite 2 Pitch and Roll Angle
Time (Orbits) Satellite 3 Pitch and Roll Angle
Time (Orbits) Time (Orbits)
Figure 4-11 Stationkeeping dynamics of the five-body STSS configuration withshyout offset (a) attitude response
78
STSS configuration 5-Body a1(0)=a2(0)=oct1(0)=2deg a3(0)=at2(0)=25o
p1(0)=p2(0)=(33(0)=pt1(0)=pt2(0)=10
62(0)=54(0)=50505m d2(0)=d4(0)=000 Stationkeeping lt1=1 =1 Okm
Tether 1 Longitudinal Vibration Tether 2 Longitudinal Vibration i i 1 i 1 r 1 1 1 1
(b) Time (Orbits) Time (Orbits)
Figure 4-11 Stationkeeping dynamics of the five-body STSS configuration withshyout offset (b) vibration response
79
The system dynamics during deployment of both the tethers in the double-
pendulum configuration is presented in Figure 4-12 Deployment of each tether takes
place from 200 m to 20 km in 35 orbits The pitch motion of the system is similar
to that observed during the three-body case The Coriolis effect causes the tethers
to pitch through a large angle ( laquo 50deg ) which in turn disturbs the platform and
subsatellite due to the presence of offset along the local vertical On the other hand
the roll response for both the tethers is damped to zero as the tether deploys in
confirmation with the principle of conservation of angular momentum whereas the
platform response in roll increases to around 20deg Subsatellite 2 remains virtually
unaffected by the other links since the tether is attached to its centre of mass thus
eliminating coupling Finally the flexibility response of the two tethers presented in
Figure 4-12(b) shows similarity with the three-body case
47 BICEPS Configuration
The mission profile of the Bl-static Canadian Experiment on Plasmas in Space
(BICEPS) is simulated next It is presently under consideration by the Canadian
Space Agency To be launched by the Black Brant 2000500 rocket it would inshy
volve interesting maneuvers of the payload before it acquires the final operational
configuration As shown in Figure 1-3 at launch the payload (two satellites) with
an undeployed tether is spinning about the orbit normal (phase 1) The internal
dampers next change the motion to a flat spin (phase 2) which in turn provides
momentum for deployment (phase 3) When fully deployed the one kilometre long
tether will connect two identical satellites carrying instrumentation video cameras
and transmitters as payload
The system parameters for the BICEPS configuration are summarized below [59 0 0 1
bull I = I2 = 0 366 0 kg bull m 2 (satellite inertias) [ 0 0 392_
bull mi = 7772 = 200 kg (mass of the satellites)
80
STSS configuration 5-Body a(0)=cx2(0)=al1(0)=2deg a3(0)=c (0)=25o
P1(0)=P2(0)=p3(0)=f3t1(0)=Pt2(0)=r 52(0)=84(0)=1304x 10300m d2(0)=d4(0)=1G0m Deployment lt1=le=02km to 20km in 35 Orbits
Pitch Roll Satellite 1 Pitch Angle
Time (Orbits) Satellite 2 Pitch and Roll Angle T 1mdash1 1 1 I I I I | I I I 1 I | I I lt~r
Tether Length Profile i 1 1 1 1 i 1 1 1 1 i 1 1 1 1 i
i i i i i 0 1 2 3 4 5
Time (Orbits) Satellite 1 Roll Angle
0 1 2 3 4 5 Time (Orbits)
Tether 1 amp 2 Roll Angle L i | I I 1 I 1 1 -I
-Ih I- bull I I i i i i I i i i i 1
0 1 2 3 4 5 Time (Orbits)
Satellite 3 Pitch and Roll Angle i- i 1 1 i 1 1 1 1 i 1 1 1 1 i 1 1 1 1 i
L i i i i i i 1 i i mdash i mdash i I i imdashimdashLJ
0 1 2 3 4 5 Time (Orbits)
Figure 4-12 Deployment dynamics of the five-body S T S S configuration with offshyset along the local vertical (a) attitude response
81
STSS configuration 5-Body a1(0)=a2(0)=oct1(0)=2deg a3(0)=at2(0)=25deg P1(0)=P2(0)=P3(0)=Pt1(0)=pt2(0)=1o
82(0)=54(0)=1304 x 10-300m d2(0)=d4(0)=100m Deployment lt1=l2=02km to 20km in 35 Orbits
Tether 1 Longitudinal Vibration Tether 2 Longitudinal Vibration
Time (Orbits) Time (Orbits) Tether 1 Z Tranverse Vibration Tether 2 Z Tranverse Vibration
Figure 4-12 Deployment dynamics of the five-body STSS configuration with offshyset along the local vertical (b) vibration response
82
bull EtAt = 61645 N (tether stiffness)
bull pt = 30 k g k m (tether density)
bull It = 1 km (tether length)
bull rid = 1-0 (tether structural damping coefficient)
The system is in a circular orbit at a 289 km altitude Offset of the tether
attachment point to the satellites at both ends is taken to be 078 m in the x
direction The response of the system in the stationkeeping phase for a prescribed set
of in i t ia l conditions is given by Figure 4-13 As in the case of the STSS configuration
there is strong coupling between the tether and the satellites rigid body dynamics
causing the latter to follow the attitude of the tether However because of the smaller
inertias of the payloads there are noticeable high frequency modulations arising from
the tether flexibility Response of the tether in the elastic degrees of freedom is
presented in Figure 4-13(b) Note the longitudinal vibrations decay quite rapidly
(lt 05 orbits) due to the structural damping however it has vir tual ly no effect on
the transverse oscillations
The unique mission requirement of B I C E P S is the proposed use of its ini t ia l
angular momentum in the cartwheeling mode to aid in the deployment of the tethered
system The maneuver is considered next The response of the system during this
maneuver wi th an ini t ia l cartwheeling rate of 5 deg s is illustrated in Figure 4-14(a) The
in i t ia l tether length is taken to be 10 m and is deployed to 1 km There is an in i t ia l
increase in the pitch motion of the tether however due to the conservation of angular
momentum the cartwheeling rate decreases proportionally to the square of the change
in tether deployment rate The result is a rapid drop in the cartwheeling rate unti l
the system stops rotating entirely and simply oscillates about its new equilibrium
point Consequently through coupling the end-bodies follow the same trend The
roll response also subsides once the cartwheeling motion ceases However
83
BICEPS configuration 3-Body a1(0)=o2(0)=at(0)=2deg P1(0)=P2(0)=pt(0)=1deg 8X(0)=001 m 6y(0)=52(0)=01 m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping l=1 km
Satellite 1 Pitch Angle Satellite 1 Roll Angle - i 1 1 1 1 i - i q r 1 1 -gt
I- I i t i i i i I H 0 1 2 0 1 2
Time (Orbits) Time (Orbits) Tether Pitch Angle Tether Roll Angle
F mdash 1 mdash 1 mdash mdash 1 mdash i mdash 1 mdash 1 mdash mdash 1 mdash = i F mdash 1 mdash mdash mdash 1 mdash i mdash 1 mdash 1 mdash 1 mdash 1 mdash q
(a) Time (Orbits) Time (Orbits)
Figure 4-13 Stationkeeping dynamics of the three-body BICEPS configuration with offset along the local vertical (a) attitude response
84
BICEPS configuration 3-Body a1(0)=o2(0)=at(0)=2deg Pl(0)=p2(0)=(3t(0)=1deg 8X(0)=001 m 8y(0)=82(0)=01 m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping lt=1km
Tether Longitudinal Vibration 1E-2[
laquo = 5 E - 3 |
000 002
Time (Orbits) Tether In-Plane Transverse Vibration
Time (Orbits) Tether Out-of-Plane Transverse Vibration
Time (Orbits)
Figure 4-13 Stationkeeping dynamics of the three-body BICEPS configuration with offset along the local vertical (b) vibration response
85
BICEPS configuration 3-Body a1(0)=o2(0)=at(0)=2deg a1(0)=a2(0)=at(0)=57s Pl(0)=P2(0)=(3t(0)=1deg 8X(0)=115 x 103m 8y(0)=82(0)=0 dx(0)=078m dy(0)=d2(0)=0 Cartwheeling Deployment
lt=10m to 1km in 1 Orbit
Tether Length Profile
E
2000
cu T3
2000
co CD
2000
C D CD
T J
Satellite 1 Pitch Angle
1 2 Time (Orbits)
Satellite 1 Roll Angle
cn CD
33 cdl
Time (Orbits) Tether Pitch Angle
Time (Orbits) Tether Roll Angle
Time (Orbits) Satellite 2 Pitch Angle
1 2 Time (Orbits)
Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-14 Cartwheeling dynamics with deployment for the three-body BICEPS with offset along the local vertical (a) attitude response
86
BICEPS configuration 3-Body a1(0)=a2(0)=at(0)=2deg a (0)=a2(0)=a(0)=57s P1(0)=P2(0)=f3(0)=1deg 8X(0)=115x103m 8y(0)=8z(0)=0 dx(0)=078m dy(0)=d2(0)=0 Cartwheeling Deployment
l=1 Om to 1 km in 1 Orbit
5E-3h
co
0E0
025
000 E to
-025
002
-002
(b)
Figure 4-14
Tether Longitudinal Vibration
Time (Orbits) Tether In-Plane Transverse Vibration
Time (Orbits) Tether Out-of-Plane Transverse Vibration
Time (Orbits)
Cartwheeling dynamics with deployment for the three-body BICEPS with offset along the local vertical (b) vibration response
87
as the deployment maneuver is completed the satellites continue to exhibit a small
amplitude low frequency roll response with small period perturbations induced by
the tether elastic oscillations superposed on it (Figure 4-14b)
48 OEDIPUS Spinning Configuration
The mission entitled Observation of Electrified Distributions of Ionospheric
Plasmas - A Unique Strategy ( O E D I P U S ) also consists of a 1 k m tethered system with
two similar end-masses It was first flown in a suborbital flight in 1989 ( O E D I P U S A )
followed by a recent study in November 1995 ( O E D I P U S - C ) in which the University
of Br i t i sh Columbia was one of the participants Dynamically O E D I P U S represents a
unique system never encountered before Launched by a Black Brant rocket developed
at Br is to l Aerospace L t d the system ie the end-bodies together wi th the tether
spins about the longitudinal axis of the tether to achieve stabilized alignment wi th
Earth s magnetic field
The response of the system undergoing a spin rate of j = l deg s is presented
in Figure 4-15 using the same parameters as those outlined for the B I C E P S conshy
figuration in Section 47 Again there is strong coupling between the three bodies
(two satellites and tether) due to the nonzero offset The spin motion introduces an
additional frequency component in the tethers elastic response which is transferred
to the l ibrational motion of the satellites A s the spin rate increases to 1 0 deg s (Figure
4-16) the amplitude and frequency of the perturbations also increase however the
general character of the response remains essentially the same
88
OEDIPUS configuration 3-Body a1(0)=a2(0)=at(0)=2deg i(0)^[2(0H(0)=Ys p1(0)=p2(0)=p(0)=r 8X(0)=12 x 103m 8y(0)=82(0)=01m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping l=1 km
Satellite 1 Pitch Angle Satellite 1 Roll Angle mdash i 1 1 1 a r 1 1 1 r
(a) Time (Orbits) Time (Orbits)
Figure 4-15 Spin dynamics (7 = ldegs) of the three-body OEDIPUS configuration with offset along the local vertical (a) attitude response
89
OEDIPUS configuration 3-Body a1(0)=a2(0)=cct(0)=2deg Y1(0)=Y2(0)=Yt(0)=l7s (31(0)=p2(0)=pt(0)=1deg 8X(0)=12 x 103m 8y(0)=5z(0)=01 m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping lt=1 km
1E-2
to
OEO
E 00
to
Tether Longitudinal Vibration
005
Time (Orbits) Tether In-Plane Transverse Vibration
Time (Orbits) Tether Out-of-Plane Transverse Vibration
Time (Orbits)
Figure 4-15 Spin dynamics (7 = l deg s ) of the three-body O E D I P U S configuration with offset along the local vertical (b) vibration response
90
OEDIPUS configuration 3-Body a1(0)=a2(0)=a(0)=2deg Yi(0)=Y2(0)=rt(0)=107s 81(0)=p2(0)=pt(0)=1deg 8X(0)=12 x 103m 8y(0)=82(0)=01m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping lt=1km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-16 Spin dynamics (7= 10degs) of the three-body OEDIPUS configura tion with offset along the local vertical (a) attitude response
91
OEDIPUS configuration 3-Body a1(0)=a2(0)=ot(0)=2deg Yi(0)H(0)4(0)laquo107s P1(0)=P2(0)=pt(0)=1deg 8X(0)=12 x 103m 8y(0)=82(0)=01m dx(0)=078m dy(0)=dz(0)=0 Stationkeeping lt=1 km
Tether Longitudinal Vibration
Figure 4-16 Spin dynamics (7 = 10degs) of the three-body OEDIPUS configurashytion with offset along the local vertical (b) vibration response
92
5 ATTITUDE AND VIBRATION CONTROL
51 Att itude Control
511 Preliminary remarks
The instability in the pitch and roll motions during the retrieval of the tether
the large amplitude librations during its deployment and the marginal stability during
the stationkeeping phase suggest that some form of active control is necessary to
satisfy the mission requirements This section focuses on the design of an attitude
controller with the objective to regulate the librational dynamics of the system
As discussed in Chapter 1 a number of methods are available to accomplish
this objective This includes a wide variety of linear and nonlinear control strategies
which can be applied in conjuction with tension thrusters offset momentum-wheels
etc and their hybrid combinations One may consider a finite number of Linear T ime
Invariant (LTI) controllers scheduled discretely over different system configurations
ie gain scheduling[43] This would be relevant during deployment and retrieval of
the tether where the configuration is changing with time A n alternative may be
to use a Linear T ime Varying ( L T V ) model in conjuction with an adaptive control
scheme where on-line parametric identification may be used to advantage[56] The
options are vir tually limitless
O f course the choice of control algorithms is governed by several important
factors effective for time-varying configurations computationally efficient for realshy
time implementation and simple in character Here the thrustermomentum-wheel
system in conjuction with the Feedback Linearization Technique ( F L T ) is chosen as
it accounts for the complete nonlinear dynamics of the system and promises to have
good overall performance over a wide range of tether lengths It is well suited for
93
highly time-varying systems whose dynamics can be modelled accurately as in the
present case
The proposed control method utilizes the thrusters located on each rigid satelshy
lite excluding the first one (platform) to regulate the pitch and rol l motion of the
tether to which it is attached ie the thrusters located on satellite 2 (link 3) regshy
ulates the attitude motion of tether 1 (link 2) O n the other hand the l ibrational
motion of the rigid bodies (platform and subsatellite) is controlled using a set of three
momentum wheels placed mutually perpendicular to each other
The F L T method is based on the transformation of the nonlinear time-varying
governing equations into a L T I system using a nonlinear time-varying feedback Deshy
tailed mathematical background to the method and associated design procedure are
discussed by several authors[435758] The resulting L T I system can be regulated
using any of the numerous linear control algorithms available in the literature In the
present study a simple P D controller is adopted arid is found to have good perforshy
mance
The choice of an F L T control scheme satisfies one of the criteria mentioned
earlier namely valid over a wide range of tether lengths The question of compushy
tational efficiency of the method wi l l have to be addressed In order to implement
this controller in real-time the computation of the systems inverse dynamics must
be executed quickly Hence a simpler model that performs well is desirable Here
the model based on the rigid system with nonlinear equations of motion is chosen
Of course its validity is assessed using the original nonlinear system that accounts
for the tether flexibility
512 Controller design using Feedback Linearization Technique
The control model used is based on the rigid system with the governing equa-
94
tions of motion given by
Mrqr + fr = Quu (51)
where the left hand side represents inertia and other forces while the right hand side
is the generalized external force due to thrusters and momentum-wheels and is given
by Eq(2109)
Lett ing
~fr = fr~ QuU (52)
and substituting in Eq(296)
ir = s(frMrdc) (53)
Expanding Eq(53) it can be shown that
qr = S(fr Mrdc) - S(QU Mr6)u (5-4)
mdash = Fr - Qru
where S(QU Mr0) is the column matrix given by
S(QuMr6) = [s(Qu(l)Mr0) 5 (Q u ( 2 ) M r | 0 ) S(QU nu) M r | 0 ) ] (55)
and Qu-i) is the i ^ column of Qu Extract ing only the controlled equations from
Eq(54) ie the attitude equations one has
Ire = Frc ~~ Qrcu
(56)
mdash vrci
where vrc is the new control input required to regulate the decoupled linear system
A t this point a simple P D controller can be applied ie
Vrc = qrcd +KvQrcd-Qrc) + Kp(Qrcd~ Qrc)- (57)
Eq(57) represents the secondary controller with Eq(54) as the primary controller
Kp and Kv are the proportional and derivative gain matrices with qrcdi Qrcd
an(^ qrcd
95
as the desired acceleration velocity and position vector for the attitude angles of each
controlled body respectively Solving for u from Eq(56) leads to
u Qrc Frc ~ vrcj bull (58)
5 13 Simulation results
The F L T controller is implemented on the three-body STSS tethered system
The choice of proportional and derivative matrix gains is based on a desired settling
time of ts mdash 05 orbit and a damping factor of = 07 for both the pitch and roll
actuators in each body Given O and ts it can be shown that
-In ) 0 5 V T ^ r (59)
and hence kn bull mdash10
V (510)
where kp and kVj are the diagonal elements of the gain matrices Kp and Kv in
Eq(57) respectively[59]
Note as each body-fixed frame is referred directly to the inertial frame FQ
the nominal pitch equilibrium angle is zero only when measured from the L V L H frame
(OJJ mdash 0) However it is 6 when referred to FQ Hence the desired position vector
qrC(l is set equal to 0(t) ie the time-varying orbital angle for the pitch motion
and zero for the roll and yaw rotations such that
qrcd 37Vxl (511)
96
The desired velocity and acceleration are then given as
e U 3 N x l (512)
E K 3 7 V x l (513)
When the system is in a circular orbit qrc^ = 0
Figure 5-1 presents the controlled response of the STSS system defined in
Chapter 4 in the stationkeeping phase with offset d2 = 111T m and If = 20
km A s mentioned earlier the F L T controller is based on the nonlinear rigid model
Note the pitch and roll angles of the rigid bodies as well as of the tether are now
attenuated in less than 1 orbit (Figure 5-la) Consequently the longitudinal vibration
response of the tether is free of coupling arising from the librational modes of the
tether This leaves only the vibration modes which are eventually damped through
structural damping The pitch and roll dynamics of the platform require relatively
higher control moments Ma and Mpi respectively (Figure 5-lb) since they have to
overcome the extra moments generated by the offset of the tether attachment point
Furthermore the platform demands an additional moment to maintain its orientation
at zero pitch and roll angle since it is not the nominal equilibrium position The nonshy
zero static equilibrium of the platform arises due to its products of inertia On the
other hand the tether pitch and roll dynamics require only a small ini t ia l control
force of about plusmn 1 N which eventually diminishes to zero once the system is
97
and
o 0
Qrcj mdash
o o
0 0
Oiit) 0 0
1
N
N
1deg I o
STSS configuration 3-Body a1(0)=a2(0)=at(0)=2lt
P1(0)=p2(0)=pt(0)=1 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=dz(0)=1m Controlled Stationkeeping lt=20km
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
05 10 15 Time (Orbits)
Sat 2 Pitch and Roll Angles
20
CO CD
T J
cdT
cT
- mdash mdash 1 1 1 1 1 I - 1
Pitch -
A Roll -
A -
1 bull bull
-
1 Time (Orbits)
Tether Y Tranverse Vibration
1 2 Time (Orbits)
Tether Z Tranverse Vibration
gt to N
to
Time (Orbits) Time (Orbits) Longitudinal Vibration
145h
(a) Time (Orbits)
Figure 5-1 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local horizontal local vertical and orbit normal (a) attitude and vibration response
98
1
STSS configuration 3-Body a1(0)=ct2(0)=cct(0)=20
P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=d2(0)=1m Controlled Stationkeeping lt=20km
Sat 1 Pitch Control Moment Sat 1 Roll Control Moment
Figure 5-1 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local horizontal local vertical and orbit normal (b) control actuator time histories
99
stabilized Similarly for the end-body a very small control torque is required to
attenuate the pitch and roll response
If the complete nonlinear flexible dynamics model is used in the feedback loop
the response performance was found to be virtually identical with minor differences
as shown in Figure 5-2 For example now the pitch and roll response of the platform
is exactly damped to zero with no steady-state error as against the minor almost
imperceptible deviation for the case of the controller basedon the rigid model (Figure
5-la) In addition the rigid controller introduces additional damping in the transverse
mode of vibration where none is present when the full flexible controller is used
This is due to a high frequency component in the at motion that slowly decays the
transverse motion through coupling
As expected the full nonlinear flexible controller which now accounts the
elastic degrees of freedom in the model introduces larger fluctuations in the control
requirement for each actuator except at the subsatellite which is not coupled to the
tether since f c m 3 = 0 The high frequency variations in the pitch and roll control
moments at the platform are due to longitudinal oscillations of the tether and the
associated changes in the tension Despite neglecting the flexible terms the overall
controlled performance of the system remains quite good Hence the feedback of the
flexible motion is not considered in subsequent analysis
When the tether offset at the platform is restricted to only 1 m along the x
direction a similar response is obtained for the system However in this case the
steady-state error in the platforms rigid body motion is much smaller (Figure 5-3a)
In addition from Figure 5-3(b) the platforms control requirement is significantly
reduced
Figure 5-4 presents the controlled response of the STSS deploying a tether
100
STSS configuration 3-Body a1(0)=a2(0)=oct(0)=2deg p1(0)=P2(0)=p(0)=1deg 8x(0)=14m5y(0)=82(0)=1m dx(0)=dy(0)=dz(0)=1m Controlled Stationkeeping l=20km
Tether Pitch and Roll Angles
c n
T J
cdT
Sat 1 Pitch and Roll Angles 1 r
00 05 10 15 Time (Orbits)
Sat 2 Pitch and Roll Angles
c n CD
cdT a
o T J
CQ
- - I 1 1 1 I 1 1 1 1
P i t c h
R o l l
bull
1 -
0 1 2 Time (Orbits)
Tether Y Tranverse Vibration
0 1 2 Time (Orbits)
Tether Z Tranverse Vibration
E gt
10 N
CO
Time (Orbits) Time (Orbits) Longitudinal Vibration
145h
(a) Time (Orbits)
Figure 5-2 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear flexible FLT controller with offset along the local horizontal local vertical and orbit normal (a) attitude and vibration response
101
STSS configuration 3-Body a1(0)=a2(0)=cxt(0)=2deg 31(0)=p2(0)=p(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=d2(0)=1m Controlled Stationkeeping lt=20km
r i i i t i mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash mdash J 0 1 2 0 1 2
Time (Orbits) Time (Orbits) Tether Pitch Control Thrust Tether Roll Control Thruster
(b) Time (Orbits) Time (Orbits)
Figure 5-2 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear flexible FLT controller with offset along the local horizontal local vertical and orbit normal (b) control actuator time histories
102
STSS configuration 3-Body a1(0)=o2(0)=a(0)=2deg P1(0)=p2(0)=(3t(0)=1deg 8x(0)=14m5y(0)=6z(0)=1m dx(0)=1mdy(0)=dz(0)=0 Controlled Stationkeeping lt=20km
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
05 10 15 Time (Orbits)
Sat 2 Pitch and Roll Angles
CD
eg CQ
0 1 2 Time (Orbits)
Tether Y Tranverse Vibration
1 2 Time (Orbits)
Tether Z Tranverse Vibration
gt CO
Time (Orbits) Time (Orbits) Longitudinal Vibration
145h
(a) Time (Orbits)
Figure 5-3 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local vertical (a) attitude and vibration response
103
STSS configuration 3-Body a1(0)=o2(0)=al(0)=2o
P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=1mdy(0)=dz(0)=0 Controlled Stationkeeping lt=20km
Sat 1 Pitch Control Moment 11 1 1 r
Sat 1 Roll Control Moment
0 1 2 Time (Orbits)
Tether Pitch Control Thrust
0 1 2 Time (Orbits)
Sat 2 Pitch Control Moment
0 1 2 Time (Orbits)
Tether Roll Control Thruster
-07 o 1
Time (Orbits Sat 2 Roll Control Moment
1E-5
0E0F
Time (Orbits)
Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local vertical (b) control actuator time histories
104
STSS configuration 3-Body a1(0)=a2(0)=a(0)=2deg Pl(0)=p2(0)=pt(0)=1deg 6X(0)=1304 x 103m 8y(0)=82(0)=0 dx(0)=1mdy(0)=d2(0)=0 Controlled Deployment
l=02km to 20km in 35 Orbits
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
o 1 2 Time (Orbits)
Tether Y Tranverse Vibration
0 1 2 Time (Orbits)
Sat 2 Pitch and Roll Angles
CD CD
CM CO
Time (Orbits) Tether Z Tranverse Vibration
to
150
100
1 2 3 0 1 2
Time (Orbits) Time (Orbits) Longitudinal Vibration
to
(a)
50
2 3 Time (Orbits)
Figure 5 - 4 Deployment dynamics of the three-body STSS using the nonlinear rigid FLT controller with offset along the local vertical (a) attitude and vibration response
105
STSS configuration 3-Body a1(0)=oc2(0)=al(0)=20
P1(0)=p2(0)=pt(0)=1deg 8X(0)=1304 x 103m 5y(0)=52(0)=0 dx(0)=1mdy(0)=dz(0)=0 Controlled Deployment
l=02km to 20km in 35 Orbits
Tether Length Profile i 1
2 0 h
E 2L
Sat 1 Pitch Control Moment t mdash r mdash i mdash i mdash | mdash i mdash i mdash i mdash i mdash | mdash r mdash i mdash i mdash i mdash | mdash r mdash
0 5
1 2 3
Time (Orbits) Sat 1 Roll Control Moment
i i t | t mdash t mdash i mdash [ i i mdash r -
1 2
Time (Orbits) Tether Pitch Control Thrust
1 2 3 Time (Orbits)
Tether Roll Control Thruster
2 E - 5
1 2 3
Time (Orbits) Sat 2 Pitch Control Moment
1 2
Time (Orbits) Sat 2 Roll Control Moment
1 E - 5
E
O E O
Time (Orbits)
Deployment dynamics of the three-body STSS using the nonlinear rigid FLT controller with offset along the local vertical (b) control actuator time histories
106
from 02 km to 20 km in 35 orbits with a i m offset along the tether length A s before
the systems attitude motion is well regulated by the controller However the control
cost increases significantly for the pitch motion of the platform and tether due to the
Coriolis force induced by the deployment maneuver (Figure 5-4b) Initially there is a
sinusoidal increase in the pitch control requirement for both the tether and platform
as the tether accelerates to its constant velocity VQ Then the control requirement
remains constant for the platform at about 60 N m as opposed to the tether where the
thrust demand increases linearly Finally when the tether decelerates the actuators
control input reduces sinusoidally back to zero The rest of the control inputs remain
essentially the same as those in the stationkeeping case In fact the tether pitch
control moment is significantly less since the tether is short during the ini t ia l stages
of control However the inplane thruster requirement Tat acts as a disturbance and
causes 6y to nearly double from the uncontrolled value (Figure 4-8a) O n the other
hand the tether has no out-of-plane deflection
Final ly the effectiveness of the F L T controller during the crit ical maneuver of
retrieval from 20 km to 02 k m in 35 orbits is assessed in Figure 5-5 A s in the earlier
cases the system response in pitch and roll is acceptable however the controller is
unable to suppress the high frequency elastic oscillations induced in the tether by the
retrieval O f course this is expected as there is no active control of the elastic degrees
of freedom However the control of the tether vibrations is discussed in detail in
Section 52 The pitch control requirement follows a similar trend in magnitude as
in the case of deployment with only minor differences due to the addition of offset in
the out-of-plane direction ( d 2 = 1 0 1 ^ m) This is also responsible for the higher
roll moment requirement for the platform control (Figure 5-5b)
107
STSS configuration 3-Body a1(0)=X2(0)=a(0)=2deg P1(0)=p2(0)=pt(0)=1deg 5x(0)=14m8y(0)=52(0)=0 dx(0)=1mdy(0)=dz(0)=0 Controlled Retrieval
l=20km to 02km in 35 Orbits
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
0 1 2 Time (Orbits)
Sat 2 Pitch and Roll Angles
CO CD
cdT
CO CD
T J CM
CO
1 2 Time (Orbits)
Tether Y Tranverse Vibration
Time (Orbits) Tether Z Tranverse Vibration
150
100
1 2 3 0 1 2 Time (Orbits) Time (Orbits)
Longitudinal Vibration
CO
(a)
x 50
2 3 Time (Orbits)
Figure 5 -5 Retrieval dynamics of the three-body STSS using the non-linear rigid FLT controller with offset along the local vertical and orbit normal (a) attitude and vibration response
108
STSS configuration 3-Body a1(0)=ct2(0)=at(0)=2deg p1(0)=p2(0)=(3(0)=1deg 5x(0)=14m5y(0)=82(0)=0 dx(0)=1mdy(0)=dz(0)=0 Controlled Retrieval
l=20km to 02km in 35 Orbits
Tether Length Profile
E
1 2 Time (Orbits)
Tether Pitch Control Thrust
2E-5
1 2 Time (Orbits)
Sat 2 Pitch Control Moment
OEOft
1 2 3 Time (Orbits)
Sat 1 Roll Control Moment
1 2 3 Time (Orbits)
Tether Roll Control Thruster
1 2 Time (Orbits)
Sat 2 Roll Control Moment
1E-5 E z
2
0E0
(b) Time (Orbits) Time (Orbits)
Figure 5-5 Retrieval dynamics of the three-body STSS using the non-linear rigid FLT controller with offset along the local vertical and orbit normal (b) control actuator time histories
109
52 Control of Tethers Elastic Vibrations
521 Prel iminary remarks
The requirement of a precisely controlled micro-gravity environment as well
as the accurate release of satellites into their final orbit demands additional control
of the tethers vibratory motion in addition to its attitude regulation To that end
an active vibration suppression strategy is designed and implemented in this section
The strategy adopted is based on offset control ie time dependent variation
of the tethers attachment point at the platform (satellite 1) A l l the three degrees
of freedom of the offset motion are used to control both the longitudinal as well as
the inplane and out-of-plane transverse modes of vibration In practice the offset
controller can be implemented through the motion of a dedicated manipulator or a
robotic arm supporting the tether which in turn is supported by the platform The
focus here is on the control of elastic deformations during stationkeeping (ie fully
deployed tether fixed length) as it represents the phase when the mission objectives
are carried out
This section begins with the linearization of the systems equations of motion
for the reduced three body stationkeeping case This is followed by the design of
the optimal control algorithm Linear Quadratic Gaussian-Loop Transfer Recovery
( L Q G L T R ) based on the reduced model and its implementation on the full nonlinear
model Final ly the systems response in the presence of offset control is presented
which tends to substantiate its effectiveness
522 System linearization and state-space realization
The design of the offset controller begins with the linearization of the equations
of motion about the systems equilibrium position Linearization of extremely lengthy
(even in matrix form) highly nonlinear nonautonomous and coupled equations of
110
motion presents a challenging problem This is further complicated by the fact the
pitch angle ot is not referred to the L V L H frame Hence the equilibrium pitch angle
is not a constant but is equal to the instantaneous orbital angle 9(t) Two methods
are available to resolve this problem
One may use the non-stationary equations of motion in their present form and
derive a controller based on the Linear T ime Varying ( L T V ) system Alternatively
one can design a Linear T ime Invariant (LTI) controller based on a new set of reduced
governing equations representing the motion of the three body tethered system A l shy
though several studies pertaining to L T V systems have been reported[5660] the latter
approach is chosen because of its simplicity Furthermore the L T I controller design
is carried out completely off-line and thus the procedure is computationally more
efficient
The reduced model is derived using the Lagrangian approach with the genershy
alized coordinates given by
where lty and are the pitch and roll angles of link i relative to the L V L H frame
respectively 5X 8y and 5Z are the generalized coordinates associated with the flexible
tether deformations in the longitudinal inplane and out-of-plane transverse direcshy
tions respectively Only the first mode of vibration is considered in the analysis
The nonlinear nonautonomous and coupled equations of motion for the tethshy
ered system can now be written as
Qred = [ltXlPlltX2P2fixSy5za302gt] (514)
M redQred + fr mdash 0gt (515)
where Mred and frec[ are the reduced mass matrix and forcing term of the system
111
respectively They are functions of qred and qred in addition to the time varying offset
position d2 and its time derivatives d2 and d2 A detailed derivation of Eq(515)
is given in Appendix II A n additional consequence of referring the pitch motion to
a local frame is that the new reduced equations are now independent of the orbital
angle 9 and under the further assumption of the system negotiating a circular orbit
91 remains constant
The nonlinear reduced equations can now be linearized about their static equishy
l ibr ium position For all the generalized coordinates the equilibrium position is zero
wi th the exception of 5X which has a non-zero equilibrium 5XQ The offset position
d2) is linearized about its ini t ial position d2^ Linearizing Eq(515) and recasting
into matr ix form gives
Msqred + CsQred + KsQred + Md^2 + Cdd2 + Kdd2 + fs = 0 (516)
where Ms Cs Ks Md Cd Kd and s are constant matrices Mul t ip ly ing throughout
by Ms
1 gives
qr = -Ms 1Csqr - Ms
lKsqr
- Ms~lCdd22 - Ms-Kdd2 - Ms~Mdd2 - Mg~1 fs r-1 (517)
Defining ud = d2 and v = qTd^T Eq(517) can be rewritten as
Pu t t ing
v =
+
-M~lCs -M~Cd -1
0
-Mg~lMd
Id
0 v + -M~lKs -M~lKd
0 0
-M-lfs
0
MCv + MKv + Mlud +Fs
- 4 ) - ( gt the L T I equations of motion can be recast into the state-space form as
(518)
(519)
x =Ax + Bud + Fd (520)
112
where
A = MC MK Id12 o
24x24
and
Let
B = ^ U 5 R 2 4 X 3 -
Fd = I FJ ) e f t 2 4 1
mdash _ i _ -5
(521)
(522)
(523)
(524)
where x is the perturbation vector from the constant equilibrium state vector xeq
Substituting Eq(524) into Eq(520) gives the linear perturbation state equation
modelling the reduced tethered system as
bullJ x mdash x mdash Ax + Bud + (Fd + Axeq) (525)
It can be shown that for ideal control[60] Fd + Axeq = 0 thus giving the familiar
state-space equation
S=Ax + Bud (526)
The selection of the output vector completes the state space realization of
the system The output vector consists of the longitudinal deformation from the
equil ibrium position SXQ of the tether at xt = h the slope of the tether due to the
transverse deformation at xt = 0 and the offset position d2 from its in i t ia l position
d2Q Thus the output vector y is given by
K(o)sx
V
dx ~ dXQ dy ~ dyQ
d z - d z0 J
Jy
dx da XQ vo
d z - dZQ ) dy dyQ
(527)
113
where lt ( 0 ) = ddxi[$vxi)]Xi=o and lt ( 0 ) = ddxi[4gtw(xi)]Xi=0
523 Linear Quadratic Gaussian control with Loop Transfer Recovery
W i t h the linear state space model defined (Eqs526527) the design of the
controller can commence The algorithm chosen is the Linear Quadratic Gaussian
( L Q G ) estimator based optimal controller[6061] The L Q G is a widely used optimal
controller with its theoretical background well developped by many authors over the
last 25 years It involves of the design of the Kalman-Bucy Fi l ter ( K B F ) which
provides an estimate of the states x and a Linear Quadratic Regulator ( L Q R ) which
is separately designed assuming all the states x are known Both the L Q R and K B F
designs independently have good robustness properties ie retain good performance
when disturbances due to model uncertainty are included However the combined
L Q R and K B F designs ie the L Q G design may have poor stability margins in the
presence of model uncertainties This l imitat ion has led to the development of an L Q G
design procedure that improves the performance of the compensator by recovering
the full state feedback robustness properties at the plant input or output (Figure 5-6)
This procedure is known as the Linear Quadratic Guassian-Loop Transfer Recovery
( L Q G L T R ) control algorithm A detailed development of its theory is given in
Ref[61] and hence is not repeated here for conciseness
The design of the L Q G L T R controller involves the repeated solution of comshy
plicated matrix equations too tedious to be executed by hand Fortunenately the enshy
tire algorithm is available in the Robust Control Toolbox[62] of the popular software
package M A T L A B The input matrices required for the function are the following
(i) state space matrices ABC and D (D = 0 for this system)
(ii) state and measurement noise covariance matrices E and 0 respectively
(iii) state and input weighting matrix Q and R respectively
114
u
MODEL UNCERTAINTY
SYSTEM PLANT
y
u LQR CONTROLLER
X KBF ESTIMATOR
u
y
LQG COMPENSATOR
PLANT
X = f v ( X t )
COMPENSATOR
x = A k x - B k y
ud
= C k f
y
Figure 5-6 Block diagram for the LQGLTR estimator based compensator
115
A s mentionned earlier the main objective of this offset controller is to regu-mdash
late the tether vibration described by the 5 equations However from the response
of the uncontrolled system presented in Chapter 4 it is clear that there is a large
difference between the magnitude of the librational and vibrational frequencies This
separation of frequencies allow for the separate design of the vibration and attitude mdash mdash
controllers Thus only the flexible subsystem composed of the S and d equations is
required in the offset controller design Similarly there is also a wide separation of
frequencies between the longitudinal and transverse modes of vibrations permitt ing mdash
the decoupling of the flexible subsystem into a set of longitudinal (Sx and dx) and
transverse (5y Sz dy and dz) subsystems The appended offset system d must also
be included since it acts as the control actuator However it is important to note
that the inplane and out-of-plane transverse modes can not be decoupled because
their oscillation frequencies are of the same order The two flexible subsystems are
summarized below
(i) Longitudinal Vibra t ion Subsystem
This is defined by u mdash Auxu + Buudu
(528)
where xu = SxdX25xdX2T u d u = dX2 and from Eq(527)
0 0 $u(lt) 0 0 0 0 1
0 0 1 0 0 0 0 1
(529)
Au and Bu are the rows and columns of A and B respectively and correspond to the
components of xu
The L Q R state weighting matrix Qu is taken as
Qu =
bull1 0 0 o -0 1 0 0 0 0 1 0
0 0 0 10
(530)
116
and the input weighting matrix Ru = 1^ The state noise covariance matr ix is
selected as 1 0 0 0 0 1 0 0 0 0 4 0
LO 0 0 1
while the measurement noise covariance matrix is taken as
(531)
i o 0 15
(532)
Given the above mentionned matrices the design of the L Q G L T R compenshy
sator is computed using M A T L A B with the following 2 commands
(i) kfu = l q r c ( ^ ( i C bdquo d i a g m x ( E u e u ) )
(ii) [AkugtBkugtCkugtDku =^ryAuBuCuDukfuQuRur)
where r is a scaler
The first command is used to compute the Ka lman filter gain matrix kfu Once
kfu is known the second command is invoked returning the state space representation
of the L Q G L T R dynamic compensator to
(533) xu mdash Akuxu Bkuyu
where xu is the state estimate vector oixu The function l t ry performs loop transfer
recovery at the system output ie the return ratio at the output approaches that
of the K B F loop given by Cu(sld - Au)Kfu[61 This is achieved by choosing a
sufficently large scaler value r in l t ry such that the singular values of the return
ratio approach those of the target design For the longitudinal controller design
r mdash 5 x 10 5 Figure 5-7(a) compares the singular values of the recovered compensator
design and the non-recovered L Q G compensator design with respect to the target
Unfortunetaly perfect recovery is not possible especially at higher frequencies
117
100
co
gt CO
-50
-100 10
Longitudinal Compensator Singular Values i I I I N I I mdash i mdash i i i i n i i mdash i mdash i i i i inimdash imdashi i i i i i i i mdash i mdash i i M i I I
Target r = 5x10 5
r = 0 (LQG)
i i i i I I i ii i i i 1 1 1 1 i i i i i 1 1 1 1 ii i I I i i i i 11111 -3 10 -2
(a)
101 10deg Frequency (rads)
101 10
Transverse Compensator Singular Values
100
50
X3
gt 0
CO
-50
bull100 10
-imdashi i 111II i 111 ni 1mdashi I I I I I I I 1mdashi I I I N I I 1mdashi i M 111
Target r = 50 r = 0 (LQG)
m i i i i i i n i i i i I I I I I I I 1mdashi I I I I I I I 1mdashi i 111 n -3 10 -2
(b)
101 10deg Frequency (rads)
101 10
Figure 5-7 Singular values for the LQG and LQGLTR compensator compared to target return ratio (a) longitudinal design (b) transverse design
118
because the system is non-minimal ie it has transmission zeros with positive real
parts[63]
(ii) Transverse Vibra t ion Subsystem
Here Xy mdash AyXy + ByU(lv
Dv = Cvxv
with xv = 6y 8Z dV2 dZ2 Sy 6Z dV2 dZ2T] udv = dy2dZ2T and
(534)
Cy
0 0 0 0 (0) 0 0 0 0 0 0 0 0 0 0 0 0
0 0 0 0
0 0 0 0 0 0 0 0 0 0 0 0
0 0 0
h 0 0 0
0 0 0 lt ( 0 ) o 0
0 1 0 0 0 1
0 0 0
2 r
h 0 0
0 0 1 0 0 1
(535)
Av and Bv are the rows and columns of the A and B corresponding to the components
O f Xy
The state L Q R weighting matrix Qv is given as
Qv
10 0 0 0 0 0 0 0 0 10 0 0 0 0 0 0 0 0 10 4 0 0 0 0 0 0 0 0 10 4 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1
(536)
mdash
Ry =w The state noise covariance matrix is taken
1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 9 x 10 4 0
0 0 0 0 0 0 0 9 x 10 4
(537)
119
while the measurement noise covariance matrix is
1 0 0 0 0 1 0 0 0 0 5 x 10 4 0
LO 0 0 5 x 10 4
(538)
As in the development of the longitudinal controller the transverse offset conshy
troller can be designed using the same two commands (lqrc and l try) wi th the input
matrices corresponding to the transverse subsystem with r = 50 The result is the
transverse compensator system given by
(539)
where xv is the state estimates of xv The singular values of the target transfer
function is compared with those achieved by the L Q G and L Q G L T R in Figure 5-
7(b) It is apparent that the recovery is not as good as that for the longitudinal case
again due to the non-minimal system Moreover the low value of r indicates that
only a lit t le recovery in the transverse subsystem is possible This suggests that the
robustness of the L Q G design is almost the maximum that can be achieved under
these conditions
Denoting f y = f j f j r xf = x^x^T and ud = udu ^ T the longishy
tudinal compensator can be combined to give
xf =
$d =
Aku
0
0 Akv
Cku o
Cu 0
Bku
0 B Xf
Xf = Ckxf
0
kv
Hf = Akxf - Bkyf
(540)
Vu Vv 0 Cv
Xf = CfXf
Defining a permutation matrix Pf such that Xj = PfX where X is the state vector
of the original nonlinear system given in Eq(32) the compensator and full nonlinear
120
system equations can be combined as
X
(541)
Ckxf
A block diagram representation of Eq(541) is shown in Figure 5-6
524 Simulation results
The dynamical response for the stationkeeping STSS with a tether length of 20
km and ini t ia l offset position of 1 m in each direction is presented in Figure 5-8 when
both the F L T attitude controller and the L Q G L T R offset controller are activated As
expected the offset controller is successful in quickly damping the elastic vibrations
in the longitudinal and transverse direction (Figure 5-8b) However from Figure 5-
8(a) it is clear that the presence of offset control requires a larger control moment
to regulate the attitude of the platform This is due to the additional torque created
by the larger moment arm around 2 m and 125 m in the inplane and out-of-plane
directions respectively introduced by the offset control In addition the control
moment for the platform is modulated by the tethers transverse vibration through
offset coupling However this does not significantly affect the l ibrational motion of
the tether whose thruster force remains relatively unchanged from the uncontrolled
vibration case
Final ly it can be concluded that the tether elastic vibration suppression
though offset control in conjunction with thruster and momentum-wheel attitude
control presents a viable strategy for regulating the dynamics tethered satellite sysshy
tems
121
STSS configuration 3-Body a1(0)=a2(0)=at(0)=2deg P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m6y(0)=52(0)=1m dx(0)=dy(0)=dz(0)=1m Stationkeeping I = 20km LQGLTR Vibration Control
Sat 1 Pitch and Roll Angles Tether Pitch and Roll Angles mdash 1 1 1 1 1 1 I 1 1 lt 1 ~ ^ 1 ^ l _
Time (Orbits) Time (Orbits) Sat 1 Roll Control Moment Tether Roll Control Thrust
(a) Time (Orbits) Time (Orbits)
Figure 5-8 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT attitude controller and LQGLTR offset vibration controller (a) attitude and libration controller reshysponse
122
STSS configuration 3-Body a1(0)=a2(0)=a(0)=2deg p1(0)=(32(0)=Q(0)=1deg ox(0)=14mSy(0)=5z(0)=1m dx(0)=dy(0)=dz(0)=1m Stationkeeping lt = 20km LQGLTR Vibration Control
Tether X Offset Position
Time (Orbits) Tether Y Tranverse Vibration 20r
Time (Orbits) Tether Y Offset Position
t o
1 2 Time (Orbits)
Tether Z Tranverse Vibration Time (Orbits)
Tether Z Offset Position
(b)
Figure 5-8
Time (Orbits) Time (Orbits)
Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT attitude controller and LQGLTR offset vibration controller (b) vibration and offset response
123
6 C O N C L U D I N G R E M A R K S
61 Summary of Results
The thesis has developed a rather general dynamics formulation for a multi-
body tethered system undergoing three-dimensional motion The system is composed
of multiple rigid bodies connected in a chain configuration by long flexible tethers
The tethers which are free to undergo libration as well as elastic vibrations in three
dimensions are also capable of deployment retrieval and constant length stationkeepshy
ing modes of operation Two types of actuators are located on the rigid satellites
thrusters and momentum-wheels
The governing equations of motion are developed using a new Order(N) apshy
proach that factorizes the mass matrix of the system such that it can be inverted
efficiently The derivation of the differential equations is generalized to account for
an arbitrary number of rigid bodies The equations were then coded in FORTRAN
for their numerical integration with the aid of a symbolic manipulation package that
algebraically evaluated the integrals involving the modes shapes functions used to dis-
cretize the flexible tether motion The simulation program was then used to assess the
uncontrolled dynamical behaviour of the system under the influence of several system
parameters including offset at the tether attachment point stationkeeping deployshy
ment and retrieval of the tether The study covered the three-body and five-body
geometries recently flown OEDIPUS system and the proposed BICEPS configurashy
tion It represents innovation at every phase a general three dimensional formulation
for multibody tethered systems an order-N algorithm for efficient computation and
application to systems of contemporary interest
Two types of controllers one for the attitude motion and the other for the
124
flexible vibratory motion are developed using the thrusters and momentum-wheels for
the former and the variable offset position for the latter These controllers are used to
regulate the motion of the system under various disturbances The attitude controller
is developed using the nonlinear Feedback Linearization Technique ( F L T ) and is based
on the nonlinear but rigid model of the system O n the other hand the separation
of the longitudinal and transverse frequencies from those of the attitude response
allows for the development of a linear optimal offset controller using the robust Linear
Quadratic Gaussian-Loop Transfer Recovery ( L Q G L T R ) method The effectiveness
of the two controllers is assessed through their subsequent application to the original
nonlinear flexible model
More important original contributions of the thesis which have not been reshy
ported in the literature include the following
(i) the model accounts for the motion of a multibody chain-type system undershy
going librational and in the case of tethers elastic vibrational motion in all
the three directions
(ii) the dynamics formulation is based on a recursive O(N) Lagrangian algorithm
that efficiently computes the systems generalized acceleration vector
(iii) the development of an attitude controller based on the Feedback Linearizashy
tion Technique for a multibody system using thrusters and momentum-wheels
located on each rigid body
(iv) the design of a three degree of freedom offset controller to regulate elastic v i shy
brations of the tether using the Linear Quadratic Gaussian and Loop Transfer
Recovery method ( L Q G L T R )
(v) substantiation of the formulation and control strategies through the applicashy
tion to a wide variety of systems thus demonstrating its versatility
125
The emphasis throughout has been on the development of a methodology to
study a large class of tethered systems efficiently It was not intended to compile
an extensive amount of data concerning the dynamical behaviour through a planned
variation of system parameters O f course a designer can easily employ the user-
friendly program to acquire such information Rather the objective was to establish
trends based on the parameters which are likely to have more significant effect on
the system dynamics both uncontrolled as well as controlled Based on the results
obtained the following general remarks can be made
(i) The presence of the platform products of inertia modulates the attitude reshy
sponse of the platform and gives rise to non-zero equilibrium pitch and roll
angles
(ii) Offset of the tether attachment point significantly affects the equil ibrium orishy
entation of the system as well as its dynamics through coupling W i t h a relshy
atively small subsatellite the tether dynamics dominate the system response
with high frequency elastic vibrations modulating the librational motion On
the other hand the effect of the platform dynamics on the tether response is
negligible As can be expected the platform dynamics is significantly affected
by the offset along the local horizontal and the orbit normal
(iii) For a three-body system deployment of the tether can destabilize the platform
in pitch as in the case of nonzero offset However the roll motion remains
undisturbed by the Coriolis force Moreover deployment can also render the
tether slack if it proceeds beyond a critical speed
(iv) Uncontrolled retrieval of the tether is always unstable in pitch It also leads
to high frequency elastic vibrations in the tether
(v) The five-body tethered system exhibits dynamical characteristics similar to
those observed for the three-body case As can be expected now there are
additional coupling effects due to two extra bodies a rigid subsatellite and a
126
flexible tether
(vi) Cartwheeling motion of the B I C E P S configuration can also be used to adshy
vantage in the deployment of the tether However exceeding a crit ical ini t ia l
cartwheeling rate can result in a large tension in the tether which eventually
causes the satellite to bounce back rendering the tether slack
(vii) Spinning the tether about its nominal length as in the case of O E D I P U S
introduces high frequency transverse vibrations in the tether which in turn
affect the dynamics of the satellites
(viii) The F L T based controller is quite successful in regulating the attitude moshy
tion of both the tether and rigid satellites in a short period of time during
stationkeeping deployment as well as retrieval phases
(ix) The offset controller is successful in suppressing the elastic vibrations of the
tether wi thin the allowable l imits of the offset mechanism ( plusmn 2 0 m)
62 Recommendations for Future Study
A s can be expected any scientific inquiry is merely a prelude to further efforts
and insight Several possible avenues exist for further investigation in this field A
few related to the study presented here are indicated below
(i) inclusion of environmental effects particularly the free molecular reaction
forces as well as interaction with Earths magnetic field and solar radiation
(ii) addition of flexible booms attached to the rigid satellites providing a mechashy
nism for energy dissipation
(iii) validation of the new three dimensional offset control strategies using ground
based experiments
(iv) animation of the simulation results to allow visual interpretation of the system
dynamics
127
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132
A P P E N D I X I TENSOR R E P R E S E N T A T I O N OF T H E
EQUATIONS OF M O T I O N
11 Prel iminary Remarks
The derivation of the equations of motion for the multibody tethered system
involves the product of several matrices as well as the derivative of matrices and
vectors with respect to other vectors In order to concisely code these relationships
in FORTRAN the matrix equations of motion are expressed in tensor notation
Tensor mathematics in general is an extremely powerful tool which can be
used to solve many complex problems in physics and engineering[6465] However
only a few preliminary results of Cartesian tensor analysis are required here They
are summarized below
12 Mathematical Background
The representation of matrices and vectors in tensor notation simply involves
the use of indices referring to the specific elements of the matrix entity For example
v = vk (k = lN) (Il)
A = Aij (i = lNj = lM)
where vk is the kth element of vector v and Aj is the element on the t h row and
3th column of matrix A It is clear from Eq(Il) that exactly one index is required to
completely define an entire vector whereas two independent indices are required for
matrices For this reason vectors and matrices are known as first-order and second-
order tensors respectively Scaler variables are known as zeroth-order tensors since
in this case an index is not required
133
Matrix operations are expressed in tensor notation similar to the way they are
programmed using a computer language such as FORTRAN or C They are expressed
in a summation notation known as Einstein notation For example the product of a
matrix with a vector is given as
w = Av (12)
or in tensor notation W i = ^2^2Aij(vkSjk)
4-1 ltL3gt j
where 8jk is the Kronecker delta Dropping the summation symbol the matrix-vector
product can be expressed compactly as
Wi = AijVj = Aimvm (14)
Here j is a dummy index since it appears twice in a single term and hence can
be replaced by any other index Note that since the resulting product is a vector
only one index is required namely i that appears exactly once on both sides of the
equation Similarly
E = aAB + bCD (15)
or Ej mdash aAimBmj + bCimDmj
~ aBmjAim I bDmjCim (16)
where A B C D and E are second-order tensors (matrices) and a and b are zeroth-
order tensors (scalers) In addition once an expression is in tensor form it can be
treated similar to a scaler and the terms can be rearranged
The transpose of a matrix is also easily described using tensors One simply
switches the position of the indices For example let w = ATv then in tensor-
notation
Wi = AjiVj (17)
134
The real power of tensor notation is in its unambiguous expression of terms
containing partial derivatives of scalers vectors or matrices with respect to other
vectors For example the Jacobian matrix of y = f(x) can be easily expressed in
tensor notation as
di-dx-j-1 ( L 8 )
mdash which is a second-order tensor as expected If f(x) = Ax where A is constant then
f - a )
according to the current definition If C = A(x)B(x) then the time derivative of C
is given by
C = Ax)Bx) + A(x)Bx) (110)
or in tensor notation
Cij = AimBmj + AimBmj (111)
= XhiAimfcBmj + AimBmj^k)
which is more clear than its equivalent vector form Note that A^^ and Bmjk are
third-order tensors that can be readily handled in F O R T R A N
13 Forcing Function
The forcing function of the equations of motion as given by Eq(260) is
- u 1-SdM- dqtdPe 9 f i m F(q qt) = Mq - -q mdashq + mdashmdash - Qd (260)
This expression can be converted into tensor form to facilitate its implementation
into F O R T R A N source code
Consider the first term Mq From Eq(289) the mass matrix M is given by
T M mdash Rv MtRv which in tensor form is
Mtj = RZiM^R^j (112)
135
Taking the time derivative of M gives
Mij = qsRv
ni^MtnrnRv
mj + Rv
niMtnTn^Rv
mj + Rv
niMtnmRv
mjs) (113)
The second term on the right hand side in Eq(260) is given by
-JPdMu 1 bdquo T x
2Q ~cWq = 2Qs s r k Q r ^ ^
Expanding Msrk leads to
M src = RnskMtnmRmr + RnsMtnmkRmr + RnsMtnmRmrk- (L15)
Finally the potential energy term in Eq(260) is also expanded into tensor
form to give dPe = dqt dPe
dq dq dqt = dq^QP^
and since ^ = Rp(q)q then
Now inserting Eqs(I13-I17) into Eq(260) and rearranging the tensor form of the
forcing term can how be stated as
Fk(q^q^)=(RP
sk + R P n J q n ) ^ - - Q d k
+ QsQr | (^Rlks ~ 2 M t n m R m r
+ (^RnkMtnms ~ 2~RnsMtnmgtk^ R m r
(^RnkRmrs ~ ~2RnsRmrk^j
where Fk(q q t) represents the kth component of F
(118)
136
A P P E N D I X II R E D U C E D EQUATIONS OF M O T I O N
II 1 Prel iminary Remarks
The reduced model used for the design of the vibration controller is represented
by two rigid end-bodies capable of three-dimensional attitude motion interconnected
with a fixed length tether The flexible tether is discretized using the assumed-
mode method with only the fundamental mode of vibration considered to represent
longitudinal as well as in-plane and out-of-plane transverse deformations In this
model tether structural damping is neglected Finally the system is restricted to a
nominal circular orbit
II 2 Derivation of the Lagrangian Equations of Mot ion
Derivation of the equations of motion for the reduced system follows a similar
procedure to that for the full case presented in Chapter 2 However in this model the
straightforward derivation of the energy expressions for the entire system is undershy
taken ignoring the Order(N) approach discussed previously Furthermore in order
to make the final governing equations stationary the attitude motion is now referred
to the local LVLH frame This is accomplished by simply defining the following new
attitude angles
ci = on + 6
Pi = Pi (HI)
7t = 0
where and are the pitch and roll angles respectively Note that spin is not
considered in the reduced model hence 7J = 0 Substituting Eq(IIl) into Eq(214)
137
gives the new expression for the rotation matrix as
T- CpiSa^ Cai+61 S 0 S a + e i
-S 0 c Pi
(II2)
W i t h the new rotation matrix defined the inertial velocity of the elemental
mass drrii o n the ^ link can now be expressed using Eq(215) as
(215)
Since the tether length is assumed to be fixed T j f j and Tj$jlt^ of Eq(215) are
identically zero Also defining
(II3) Vi = I Pi i
Rdm- f deg r the reduced system is given by
where
and
Rdm =Di + Pi(gi)m + T^iSi
Pi(9i)m= [Tai9i T0gi T^] fj
(II4)
(II5)
(II6)
D1 = D D l D S v
52 = 31 + P(d2)m +T[d2
D3 = D2 + l2 + dX2Pii)m + T2i6x
Note that $ j for i = 2 is defined in Section 212 whereas for i = 1 and 3 it is
the null matrix Since only one mode is used for each flexible degree of freedom
5i = [6Xi6yi6Zi]T for i = 2
The kinetic energy of the system can now be written as
i=l Z i = l Jmi RdmRdmdmi (UJ)
138
Expanding Kampi using Eq(II4)
KH = miDD + 2DiP
i j gidmi)fji + 2D T ltMtrade^
+tf J P^(9i)Pi9i)d^l + 2mT j PFmTi^dm^ (II8)
where the integrals are evaluated using the procedure similar to that described in
Chapter 2 The gravitational and strain energy expressions are given by Eq(247)
and Eq(251) respectively using the newly defined rotation matrix T in place of
TV
Substituting the kinetic and potential energy expressions in
d ( 8Ke d(Ke - Pe) i bull i Q ^ 0 (II9)
d t dqred d(ired
with
Qred = [ algt l gt a 2 2 A 2 A 2 lt ^ 2 Q 3 3 ] (5-14)
and d2 regarded as a time-varying parameter the nonlinear non-autonomous equashy
tions of motion for the reduced system can finally be expressed as
MredQred + fred = reg- (5-15)
139
22 Kinetics and System Energy 22
221 Kinet ic energy 22
222 Simplification for rigid links 23
223 Gravitat ional potential energy 25
224 Strain energy 26
225 Tether energy dissipation 27
23 O(N) Form of the Equations of Mot ion 28
231 Lagrange equations of motion 28
232 Generalized coordinates and position transformation 29
233 Velocity transformations 31
234 Cyl indrical coordinate modification 33
235 Mass matrix inversion 34
236 Specification of the offset position 35
24 Generalized Control Forces 36
241 Preliminary remarks 36
242 Generalized thruster forces 37
243 Generalized momentum gyro torques 39
3 C O M P U T E R I M P L E M E N T A T I O N 42
31 Preliminary Remarks 42
32 Numerical Implementation 43
321 Integration routine 43
322 Program structure 43
33 Verification of the Code 46
iv
331 Energy conservation 4 6
332 Comparison with available data 50
4 D Y N A M I C S I M U L A T I O N 53
41 Prel iminary Remarks 53
42 Parameter and Response Variable Definitions 53
43 Stationkeeping Profile 56
44 Tether Deployment 66
45 Tether Retrieval 71
46 Five-Body Tethered System 71
47 B I C E P S Configuration 80
48 O E D I P U S Spinning Configuration 88
5 A T T I T U D E A N D V I B R A T I O N C O N T R O L 93
51 Att i tude Control 93
511 Prel iminary remarks 93
512 Controller design using Feedback Linearization Technique 94
513 Simulation results 96
52 Control of Tethers Elastic Vibrations 110
521 Prel iminary remarks 110
522 System linearization and state-space realization 1 1 0
523 Linear Quadratic Gaussian control
with Loop Transfer Recovery 114
524 Simulation results 121
v
6 C O N C L U D I N G R E M A R K S 1 2 4
61 Summary of Results 124
62 Recommendations for Future Study 127
B I B L I O G R A P H Y 128
A P P E N D I C E S
I T E N S O R R E P R E S E N T A T I O N O F T H E E Q U A T I O N S O F M O T I O N 133
11 Prel iminary Remarks 133
12 Mathematical Background 133
13 Forcing Function 135
II R E D U C E D E Q U A T I O N S O F M O T I O N 137
II 1 Preliminary Remarks 137
II2 Derivation of the Lagrangian Equations of Mot ion 137
v i
LIST OF SYMBOLS
A j intermediate length over which the deploymentretrieval profile
for the itfl l ink is sinusoidal
A Lagrange multipliers
EUQU longitudinal state and measurement noise covariance matrices
respectively
matrix containing mode shape functions of the t h flexible link
pitch roll and yaw angles of the i 1 link
time-varying modal coordinate for the i 1 flexible link
link strain and stress respectively
damping factor of the t h attitude actuator
set of attitude angles (fjj = ci
structural damping coefficient for the t h l ink EjJEj
true anomaly
Earths gravitational constant GMejBe
density of the ith link
fundamental frequency of the link longitudinal tether vibrashy
tion
A tethers cross-sectional area
AfBfCjDf state-space representation of flexible subsystem
Dj inertial position vector of frame Fj
Dj magnitude of Dj mdash mdash
Df) transformation matrix relating Dj and Ds
D r i inplane radial distance of the first l ink
DSi Drv6i DZ1 T for i = 1 and DX Dy poundgt 2 T for 1
m
Pi
w 0 i
v i i
Dx- transformation matrix relating D j and Ds-
DXi Dyi Dz- Cartesian components of D
DZl out-of-plane position component of first link
E Youngs elastic modulus for the ith l ink
Ej^ contribution from structural damping to the augmented complex
modulus E
F systems conservative force vector
FQ inertial reference frame
Fi t h link body-fixed reference frame
n x n identity matrix
alternate form of inertia matrix for the i 1 l ink
K A Z $i(k + dXi+1)
Kei t h link kinetic energy
M(q t) systems coupled mass matrix
Mma Mma Mmy- control moments in the pitch roll and yaw directions respec-
tively for the t h link
Mr fr rigid mass matrix and force vector respectively
Mred fred- Qred mass matrix force vector and generalized coordinate vector for
the reduced model respectively
Mt block diagonal decoupled mass matrix
symmetric decoupled mass matrix
N total number of links
0N) order-N
Pi(9i) column matrix [T^giTpg^T^gi) mdash
Q nonconservative generalized force vector
Qu actuator coupling matrix or longitudinal L Q R state weighting
matrix
v i i i
Rdm- inertial position of the t h link mass element drrii
RP transformation matrix relating qt and tf
Ru longitudinal L Q R input weighting matrix
RvRn Rd transformation matrices relating qt and q
S(f Mdc) generalized acceleration vector of coupled system q for the full
nonlinear flexible system
th uiirv l u w u u u maniA
th
T j i l ink rotation matrix
TtaTto control thrust in the pitch and roll direction for the im l ink
respectively
rQi maximum deploymentretrieval velocity of the t h l ink
Vei l ink strain energy
Vgi link gravitational potential energy
Rayleigh dissipation function arising from structural damping
in the ith link
di position vector to the frame F from the frame F_i bull
dc desired offset acceleration vector (i 1 l ink offset position)
drrii infinitesimal mass element of the t h l ink
dx^dy^dZi Cartesian components of d along the local vertical local horishy
zontal and orbit normal directions respectively
F-Q mdash
gi r igid and flexible position vectors of drrii fj + ltEraquolt5
ijk unit vectors 1 0 0 r 010-^ and 00 l r respectively
li length of the t h l ink
raj mass of the t h link
nfj total number of flexible modes for the i 1 l ink nXi + nyi + nZi
nq total number of generalized coordinates per link nfj + 7
nqq systems total number of generalized coordinates Nnq
ix
number of flexible modes in the longitudinal inplane and out-
of-plane transverse directions respectively for the i ^ link
number of attitude control actuators
- $ T
set of generalized coordinates for the itfl l ink which accounts for
interactions with adjacent links
qv---QtNT
set of coordinates for the independent t h link (not connected to
adjacent links)
rigid position of dm in the frame Fj
position of centre of mass of the itfl l ink relative to the frame Fj
i + $DJi
desired settling time of the j 1 attitude actuator
actuator force vector for entire system
flexible deformation of the link along the Xj yi and Zj direcshy
tions respectively
control input for flexible subsystem
Cartesian components of fj
actual and estimated state of flexible subsystem respectively
output vector of flexible subsystem
LIST OF FIGURES
1-1 A schematic diagram of the space platform based N-body tethered
satellite system 2
1-2 Associated forces for the dumbbell satellite configuration 3
1- 3 Some applications of the tethered satellite systems (a) multiple
communication satellites at a fixed geostationary location
(b) retrieval maneuver (c) proposed B I C E P S configuration 6
2- 1 Vector components of the i 1 and (i- l ) 1 chain links 15
2-2 Vector components in cylindrical orbital coordinates 20
2-3 Inertial position of subsatellite thruster forces 38
2- 4 Coupled force representation of a momentum-wheel on a rigid body 40
3- 1 Flowchart showing the computer program structure 45
3-2 Kinet ic and potentialenergy transfer for the three-body platform based
tethered satellite system (a) variation of kinetic and potential energy
(b) percent change in total energy of system 48
3-3 Kinet ic and potential energy transfer for the five-body platform based
tethered satellite system (a) variation of kinetic and potential energy
(b) percent change in total energy of system 49
3-4 Simulation results for the platform based three-body tethered
system originally presented in Ref[43] 51
3- 5 Simulation results for the platform based three-body tethered
system obtained using the present computer program 52
4- 1 Schematic diagram showing the generalized coordinates used to
describe the system dynamics 55
4-2 Stationkeeping dynamics of the three-body STSS configuration
without offset (a) attitude response (b) vibration response 57
4-3 Stationkeeping dynamics of the three-body STSS configuration
xi
with offset along the local vertical
(a) attitude response (b) vibration response 60
4-4 Stationkeeping dynamics of the three-body STSS configuration
with offset along the local horizontal
(a) attitude response (b) vibration response 62
4-5 Stationkeeping dynamics of the three-body STSS configuration
with offset along the orbit normal
(a) attitude response (b) vibration response 64
4-6 Stationkeeping dynamics of the three-body STSS configuration
wi th offset along the local horizontal local vertical and orbit normal
(a) attitude response (b) vibration response 67
4-7 Deployment dynamics of the three-body STSS configuration
without offset
(a) attitude response (b) vibration response 69
4-8 Deployment dynamics of the three-body STSS configuration
with offset along the local vertical
(a) attitude response (b) vibration response 72
4-9 Retrieval dynamics of the three-body STSS configuration
with offset along the local vertical and orbit normal
(a) attitude response (b) vibration response 74
4-10 Schematic diagram of the five-body system used in the numerical example 77
4-11 Stationkeeping dynamics of the five-body STSS configuration
without offset
(a) attitude response (b) vibration response 78
4-12 Deployment dynamics of the five-body STSS configuration
wi th offset along the local vertical
(a) attitude response (b) vibration response 81
4-13 Stationkeeping dynamics of the three-body B I C E P S configuration
x i i
with offset along the local vertical
(a) attitude response (b) vibration response 84
4-14 Cartwheeling dynamics with deployment for the three-body B I C E P S
with offset along the local vertical
(a) attitude response (b) vibration response 86
4-15 Spin dynamics (7 = l deg s ) of the three-body O E D I P U S configuration
with offset along the local vertical
(a) attitude response (b) vibration response 89
4- 16 Spin dynamics (7 = 10deg s ) of the three-body O E D I P U S configuration
wi th offset along the local vertical
(a) attitude response (b) vibration response 91
5- 1 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear rigid F L T controller with offset along the
local horizontal local vertical and orbit normal
(a) attitude and vibration response (b) control actuator time histories 98
5-2 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear flexible F L T controller with offset along the
local horizontal local vertical and orbit normal
(a) attitude and vibration response (b) control actuator time histories 101
5-3 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear rigid F L T controller with offset along
the local vertical
(a) attitude and vibration response (b) control actuator time histories 103
5-4 Deployment dynamics of the three-body STSS using the nonlinear
rigid F L T controller with offset along the local vertical
(a) attitude and vibration response (b) control actuator time histories 105
5-5 Retrieval dynamics of the three-body STSS using the non-linear
rigid F L T controller with offset along the local vertical and orbit normal
(a) attitude and vibration response (b) control actuator time histories 108
x i i i
5-6 Block diagram for the L Q G L T R estimator based compensator 115
5-7 Singular values for the L Q G and L Q G L T R compensator compared to target
return ratio (a) longitudinal design (b) transverse design 118
5-8 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear rigid F L T attitude controller and L Q G L T R
offset vibration controller (a) attitude and libration controller response
(b) vibration and offset response 122
xiv
A C K N O W L E D G E M E N T
Firs t and foremost I would like to express the most genuine gratitude to
my supervisor Prof V i n o d J M o d i whose unending guidance and encouragement
proved most invaluable during the course of my studies
I am also deeply indebted to Dr Satyabrata Pradhan for his limitless patience
and technical advice from which this work would not have been possible I would also
like to extend a word of appreciation towards Prof A r u n K Mis ra ( M c G i l l University)
for ini t iat ing my interest in space dynamics and control
Also I would like to thank my collegues Dr Sandeep Munshi M r Mathieu
Caron and M r Yuan Chen as well as Dr Mae Seto M r Gary L i m and M r Mark
Chu for their words of advice and helpful suggestions that heavily contributed to all
aspects of my studies
Final ly I would like to thank all my family and friends here in Vancouver and
back home in Montreal whose unending support made the past two years the most
memorable of my life
The research project was supported by the Natural Sciences and Engineering
Research Council s P G S - A Scholarship held by the author as well as by N S E R C grant
A-2181 held by Prof V J M o d i
xv
1 INTRODUCTION
11 Prel iminary Remarks
W i t h the ever changing demands of the worlds population one often wonders
about the commitment to the space program On the other hand the importance
of worldwide communications global environmental monitoring as well as long-term
habitation in space have demanded more ambitious and versatile satellites systems
It is doubtful that the Russian scientist Tsiolkovsky[l] when he first proposed the
uti l izat ion of the Earths gravity-gradient environment in the last century would have
envisioned the current and proposed applications of tethered satellite systems
A tethered satellite system consists of two or more subsatellites connected
to each other by long thin cables or tethers which are in tension (Figure 1-1) The
subsatellite which can also include the shuttle and space station are in orbit together
The first proposed use of tethers in space was associated with the rescue of stranded
astronauts by throwing a buoy on a tether and reeling it to the rescue vehicle
Prel iminary studies of such systems led to the discovery of the inherent instability
during the tether retrieval[2] Nevertheless the bir th of tethered systems occured
during the Gemini X I and X I I missions[3] in 1966 when a short tether (10-30m)
was used to generate the first artificial gravity environment in space (000015g) by
cartwheeling (spinning) the system about its center of mass
The mission also demonstrated an important use of tethers for gravity gradient
stabilization The force analysis of a simplified model illustrates this point Consider
the dumbbell satellite configuration orbiting about the centre of Ear th as shown in
Figure 1-2 Satellite 1 and satellite 2 are connected to each other by a tether with
1
Space Station (Satellite 1 )
2
Satellite N
Figure 1-1 A schematic diagram of the space platform based N-body tethered satellite system
2
Figure 1-2 Associated forces for the dumbbell satellite configuration
3
the whole system free to librate about the systems centre of mass C M The two
major forces acting on each body are due to the centrifugal and gravitational effects
Since body 1 is further away from the earth its gravitational force Fg^y is smaller
than that of body 2 ie Fgi lt Fg^ However its centrifugal force FCl is greater
then the centrifugal force experienced by the second satellite (FCl gt FC2) Moreover
for body 1 Fg^ lt Fc^ in contrast to Fg2 gt FC2 for body 2 Thus resolving the force
components along the tether there is an evident resultant tension force Ff present
in the tether Similarly adding the normal components of the two forces vectorially
results in a force FR which restores the system to its stable equil ibrium along the
local vertical These tension and gravity-gradient restoring forces are the two most
important features of tethered systems
Several milestone missions have flown in the recent past The U S A J a p a n
project T P E (Tethered Payload Experiment) [4] was one of the first to be launched
using a sounding rocket to conduct environmental studies The results of the T P E
provided support for the N A S A I t a l y shuttle based tethered satellite system referred
to as the TSS-1 mission[5] This experiment aimed to study the electrodynamic effects
of a shuttle borne satellite system with a conductive tether connection where an
electric current was induced in the tether as it swept through the Earth s magnetic
field Unfortunetely the two attempts in August 1992 and February 1996 resulted
in only partial success The former suffered from a spool failure resulting in a tether
deployment of only 256m of a planned 20km During the latter attempt a break
in the tethers insulation resulted in arcing leading to tether rupture Nevertheless
the information gathered by these attempts st i l l provided the engineers and scientists
invaluable information on the dynamic behaviour of this complex system
Canada also paved the way with novel designs of tethered system aimed
primari ly at the study of the ionosphere and Northern Lights (Aurora Borealis)
4
The O E D I P U S A and C (Observation of Electrified Distributions in the Ionospheric
P lasma - a Unique Strategy) missions[6] launched from a sounding rocket in January
1989 and November 1995 respectively provided insight into the complex dynamical
behaviour of two comparable mass satellites connected by a 1 km long spinning
tether
Final ly two experiments were conducted by N A S A called S E D S I and II
(Small Expendable Deployable System) [7] which hold the current record of 20 km
long tether In each of these missions a small probe (26 kg) was deployed from the
second stage of a D E L T A II rocket thus succesfully demonstrating the feasibility of
long tether deployment Retrieval of the tether which is significantly more difficult
has not yet been achieved
Several proposed applications are also currently under study These include the
study of earths upper atmosphere using probes lowered from the shuttle in a low earth
orbit ( L E O ) to the deployment and retrieval of satellites for servicing Even more
promising application concerns the generation of power for the proposed International
Space Station using conductive tethers The Canadian Space Agency (CSA) has also
proposed a dual-mass satellite system B I C E P S (Bl-static Canadian Experiment on
Plasmas in Space) to make simultaneous measurements at different locations in the
environment for correlation studies[8] A unique feature of the B I C E P S mission is
the deployment of the tether aided by the angular momentum of the system which
is ini t ia l ly cartwheeling about its orbit normal (Figure 1-3)
In addition to the three-body configuration multi-tethered systems have been
proposed for monitoring Earths environment in the global project monitored by
N A S A entitled Mission to Planet Ear th ( M T P E ) The proposed mission aims to
study pollution control through the understanding of the dynamic interactions be-
5
Retrieval of Satellite for Servicing
lt mdash
lt Satellite
Figure 1-3 Some applications of the tethered satellite systems (a) multiple comshymunication satellites at a fixed geostationary location (b) retrieval maneuver (c) proposed BICEPS configuration
6
tween the Earths atmosphere biosphere hydrosphere and cryosphere This wi l l be
accomplished with multiple tethered instrumentation payloads simultaneously soundshy
ing at different altitudes for spatial correlations Such mult ibody systems are considshy
ered to be the next stage in the evolution of tethered satellites Micro-gravity payload
modules suspended from the proposed space station for long-term experiments as well
as communications satellites with increased line-of-sight capability represent just two
of the numerous possible applications under consideration
12 Brief Review of the Relevant Literature
The design of multi-payload systems wi l l require extensive dynamic analysis
and parametric study before the final mission configuration can be chosen This
wi l l include a review of the fundamental dynamics of the simple two-body tethered
systems and how its dynamics can be extended to those of multi-body tethered system
in a chain or more generally a tree topology
121 Multibody 0(N) formulation
Many studies of two-body systems have been reported One of the earliest
contribution is by Rupp[9] who was the first to study planar dynamics of the simshy
plified Shuttle based Tethered Satellite System (STSS) His pioneering work led to
the discovery of pitch oscillation growth during the retrieval phase Later Baker[10]
advanced the investigation to the third dimension in addition to adding atmospheric
effects to the system A more complete dynamical model was later developed and
analyzed by M o d i and Misra[ l l 12] which included deployment and tether flexibility
A comprehensive survey of important developments and milestones in the area have
been compiled and reviewed by Mis ra and Modi[13] The major conclusions based on
the literature may be summarized as follows
bull the stationkeeping phase is stable
7
bull deployment can be unstable if the rate exceeds a crit ical speed
bull retrieval of the tether is inherently unstable
bull transverse vibrations can grow due to the Coriolis force induced during deshy
ployment and retrieval
Misra Amier and Modi[14] were one of the first to extend these results to the
three-satellite double pendulum case This simple model which includes a variable
length tether was sufficient to uncover the multiple equilibrium configurations of
the system Later the work was extended to include out-of-plane motion and a
preliminary reel-rate control law to regulate the librational motion[15] Kesmir i and
Misra[16] developed a general formulation for N-body tethered systems based on the
Lagrangian principle where three-dimensional motion and flexibility are accounted for
However from their work it is clear that as the number of payload or bodies increases
the computational cost to solve the forward dynamics also increases dramatically
Tradit ional methods of inverting the mass matrix M in the equation
ML+F = Q
to solve for the acceleration vector proved to be computationally costly being of the
order for practical simulation algorithms O(N^) refers to a mult ibody formulashy
tion algorithm whose computational cost is proportional to the cube of the number
of bodies N used It is therefore clear that more efficient solution strategies have to
be considered
Many improved algorithms have been proposed to reduce the number of comshy
putational steps required for solving for the acceleration vector Rosenthal[17] proshy
posed a recursive formulation based on the triangularization of the equations of moshy
tion to obtain an 0(N2) formulation He also demonstrates the use of a symbolic
manipulation program to reduce the final number of computations A recursive L a -
8
grangian formulation proposed by Book[18] also points out the relative inefficiency of
conventional formulations and proceeds to derive an algorithm which is also 0(N2)
A recursive formulation is one where the equations of motion are calculated in order
from one end of the system to the other It usually involves one or more passes along
the links to calculate the acceleration of each link (forward pass) and the constraint
forces (backward or inverse pass) Non-recursive algorithms express the equations of
motion for each link independent of the other Although the coupling terms st i l l need
to be included the algorithm is in general more amenable to parallel computing
A st i l l more efficient dynamics formulation is an O(N) algorithm Here the
computational effort is directly proportional to the number of bodies or links used
Several types of such algorithms have been developed over the years The one based
on the Newton-Euler approach has recursive equations and is used extensively in the
field of multi-joint robotics[1920] It should be emphasized that efficient computation
of the forward and inverse dynamics is imperative if any on-line (real-time) control
of the joint is to be achieved Hollerbach[21] proposed a recursive O(N) formulation
based on Lagranges equations of motion However his derivation was primari ly foshy
cused on the inverse dynamics and did not improve the computational efficiency of
the forward dynamics (acceleration vector) Other recursive algorithms include methshy
ods based on the principle of vir tual work[22] and Kanes equations of motion[23]
Keat[24] proposed a method based on a velocity transformation that eliminated the
appearance of constraint forces and was recursive in nature O n the other hand
K u r d i l a Menon and Sunkel[25] proposed a non-recursive algorithm based on the
Range-Space method which employs element-by-element approach used in modern
finite-element procedures Authors also demonstrated the potential for parallel comshy
putation of their non-recursive formulation The introduction of a Spatial Operator
Factorization[262728] which utilizes an analogy between mult ibody robot dynamics
and linear filtering and smoothing theory to efficiently invert the systems mass ma-
9
t r ix is another approach to a recursive algorithm Their results were further extended
to include the case where joints follow user-specified profiles[29] More recently Prad-
han et al [30] proposed a non-recursive Lagrangian algorithm for factorizing the mass
matrix leading to an O(N) formulation of the forward dynamics of the system A n
efficient O(N) algorithm where the number of bodies N in the system varies on-line
has been developed by Banerjee[31]
122 Issues of tether modelling
The importance of including tether flexibility has been demonstrated by several
researchers in the past However there are several methods available for modelling
tether flexibility Each of these methods has relative advantages and limitations
wi th regards to the computational efficiency A continuum model where flexible
motion is discretized using the assumed mode method has been proposed[3233] and
succesfully implemented In the majority of cases even one flexible mode is sufficient
to capture the dynamical behaviour of the tether[34] K i m and Vadali[35] proposed a
so-called bead model or lumped-mass approach that discretizes the tether using point
masses along its length However in order to accurately portray the motion of the
tether a significantly high number of beads are needed thus increasing the number of
computation steps Consequently this has led to the development of a hybrid between
the last two approaches[16] ie interpolation between the beads using a continuum
approach thus requiring less number of beads Finally the tether flexibility can also
be modelled using a finite-element approach[36] which is more suitable for transient
response analysis In the present study the continuum approach is adopted due to
its simplicity and proven reliability in accurately conveying the vibratory response
In the past the modal integrals arising from the discretization process have
been evaluted numerically In general this leads to unnecessary effort by the comshy
puter Today these integrals can be evaluated analytically using a symbolic in-
10
tegration package eg Maple V Subsequently they can be coded in F O R T R A N
directly by the package which could also result in a significant reduction in debugging
time More importantly there is considerable improvement in the computational effishy
ciency [16] especially during deployment and retrieval where the modal integral must
be evaluated at each time-step
123 Attitude and vibration control
In view of the conclusions arrived at by some of the researchers mentioned
above it is clear that an appropriate control strategy is needed to regulate the dyshy
namics of the system ie attitude and tether vibration First the attitude motion of
the entire system must be controlled This of course would be essential in the case of
a satellite system intended for scientific experiments such as the micro-gravity facilishy
ties aboard the proposed Internation Space Station Vibra t ion of the flexible members
wi l l have to be checked if they affect the integrity of the on-board instrumentation
Thruster as well as momentum-wheel approach[34] have been considered to
regulate the rigid-body motion of the end-satellite as well as the swinging motion
of the tether The procedure is particularly attractive due to its effectiveness over a
wide range of tether lengths as well as ease of implementation Other methods inshy
clude tension and length rate control which regulate the tethers tension and nominal
unstretched length respectively[915] It is usually implemented at the deployment
spool of the tether More recently an offset strategy involving time dependent moshy
tion of the tether attachment point to the platform has been proposed[3738] It
overcomes the problem of plume impingement created by the thruster control and
the ineffectiveness of tension control at shorter lengths However the effectiveness of
such a controller can become limited with an exceedingly long tether due to a pracshy
tical l imi t on the permissible offset motion In response to these two control issues a
hybrid thrusteroffset scheme has been proposed to combine the best features of the
11
two methods[39]
In addition to attitude control these schemes can be used to attenuate the
flexible response of the tether Tension and length rate control[12] as well as thruster
based algorithms[3340] have been proposed to this end M o d i et al [41] have sucshy
cessfully demonstrated the effectiveness of an offset strategy to damp undesired v i shy
brations Passive energy dissipative devices eg viscous dampers are also another
viable solution to the problem
The development of various control laws to implement the above mentioned
strategies has also recieved much attention A n eigenstructure assignment in conducshy
tion wi th an offset controller for vibration attenuation and momemtum wheels for
platform libration control has been developed[42] in addition to the controller design
from a graph theoretic approach[41] Also non-linear feedback methods such as the
Feedback Linearization Technique ( F L T ) that are more suitable for controlling highly
nonlinear non-autonomous coupled systems have also been considered[43]
It is important to point out that several linear controllers including the classhy
sic state feedback Linear Quadratic Regulator ( L Q R ) have received considerable
attention[3944] Moreover robust methods such as the Linear Quadratic Guas-
s ian Loop Transfer Recovery ( L Q G L T R ) method have also been developed and imshy
plemented on tethered systems[43] A more complete review of these algorithms as
well as others applied to tethered system has been presented by Pradhan[43]
13 Scope of the Investigation
The objective of the thesis is to develop and implement a versatile as well as
computationally efficient formulation algorithm applicable to a large class of tethered
satellite systems The distinctive features of the model can be summarized as follows
12
bull TV-body 0(N) tethered satellite system in a chain-type topology
bull the system is free to negotiate a general Keplerian orbit and permitted to
undergo three dimensional inplane and out-of-plane l ibrtational motion
bull r igid bodies constituting the system are free to execute general rotational
motion
bull three dimensional flexibility present in the tether which is discretized using
the continuum assumed-mode method with an arbitrary number of flexible
modes
bull energy dissipation through structural damping is included
bull capability to model various mission configurations and maneuvers including
the tether spin about an arbitrary axis
bull user-defined time dependent deployment and retrieval profiles for the tether
as well as the tether attachment point (offset)
bull attitude control algorithm for the tether and rigid bodies using thrusters and
momentum-wheels based on the Feedback Linearization Technique
bull the algorithm for suppression of tether vibration using an active offset (tether
attachment point) control strategy based on the optimal linear control law
( L Q G L T - R )
To begin with in Chapter 2 kinematics and kinetics of the system are derived
using the O(N) formulation methodology developed by Pradhan et al [30] Chapshy
ter 3 discusses issues related to the development of the simulation program and its
validation This is followed by a detailed parametric study of several mission proshy
files simulated as particular cases of the versatile formulation in Chapter 4 Next
the attitude and vibration controllers are designed and their effectiveness assessed in
Chapter 5 The thesis ends with concluding remarks and suggestions for future work
13
2 F O R M U L A T I O N OF T H E P R O B L E M
21 Kinematics
211 Preliminary definitions and the i ^ position vector
The derivation of the equations of motion begins with the definition of the
inertial position of an arbitrary link of the mult ibody system Let the i ^ link of the
system be free to translate and rotate in 3-D space From Figure 2-1 the position
vector Rdm^ to the mass element drrn on the i ^ link wi th reference to the inertial
frame FQ can be written as
Here D represents the inertial position of the t h body fixed frame Fj relative to FQ
ri = [xiyiZi]T is the rigid position vector of drrii wi th respect to F J and f(fi) is
the flexible deformation at fj also with respect to the frame Fj Bo th these vectors
are relative to Fj Note in the present model tethers (i even) are considered flexible
and X corresponds to the nominal position along the unstretched tether while a and
ZJ are by definition equal to zero On the other hand the rigid bodies referred to as
satellites (i odd) including the platform are taken to be rigid ie f(fj) = 0 T j is
the rotation matrix used to express body-fixed vectors with reference to the inertial
frame
212 Tether flexibility discretization
The tether flexibility is discretized with an arbitrary but finite number of
Di + Ti(i + f(i)) (21)
14
15
modes in each direction using the assumed-mode method as
Ui nvi
wi I I i = 1
nzi bull
(22)
where nXi nyi and n 2 ^ are the number of modes used in the X m and Z directions
respectively For the ] t h longitudinal mode the admissible mode shape function is
taken as 2 j - l
Hi(xiii)=[j) (23)
where l is the i ^ tether length[32] In the case of inplane and out-of-plane transverse
deflections the admissible functions are
ampyi(xili) = ampZixuk) = v ^ s i n JKXj
k (24)
where y2 is added as a normalizing factor In this case both the longitudinal and
transverse mode shapes satisfy the geometric boundary conditions for a simple string
in axial and transverse vibration
Eq(22) is recast into a compact matrix form as
ff(i) = $i(xili)5i(t) (25)
where $i(x l) is the matrix of the tether mode shapes defined as
0 0
^i(xiJi) =
^ X bull bull bull ^XA
0
0 0
0
G 5R 3 x n f
(26)
and nfj = nx- + nyi + nZi Note $J(XJJ) is a function of the spatial coordinate x
and not of yi or ZJ However it is also a function of the tether length parameter li
16
which is time-varying during deployment and retrieval For this reason care must be
exercised when differentiating $J(XJJ) with respect to time such that
d d d k) = Jtregi(Xi li) = ^ 7 $ t ( ^ i k)Xi + gf$i(xigt li)k- (2-7)
Since x represents the nominal rate of change of the position of the elemental mass
drrii it is equal to j (deploymentretrieval rate) Let t ing
( 3 d
dx~ + dT)^Xili^ ( 2 8 )
one arrives at
$i(xili) = $Di(xili)ii (29)
The time varying generalized coordinate vector 5 in Eq(25) is composed of
the longitudinal and transverse components ie
ltM) G s f t n f i x l (210)
where 5X^ 5y- and Sz^ are nx- x 1 ny x 1 and nz- x 1 size vectors and correspond to
$ X and $ Z respectively
213 Rotation angles and transformations
The matrix T j in Eq(21) represents a rotation transformation from the body
fixed frame Fj to the inertial frame FQ ie FQ = T j F j It is defined by the 3-2-1
ordered sequence of elementary rotations[46]
1 P i tch F[ = Cf ( a j )F 0
2 R o l l F = C f (3j)F = C7f (3j)Cf ( a j )F 0
3 Yaw F = C7( 7 i )F = C7(7l)cf^)Cf(ai)FQ = Q F 0
17
where
c f (A) =
-Sai Cai
0 0
and
C(7i ) =
0 -s0i] = 0 1 0
0 cpi
i 0 0 = 0
0
(211)
(212)
(213)
Here Eq(211) corresponds to a rotation of ctj (pitch) about the inertial axis ZQ (orbit
normal) of frame FQ resulting in a new intermediate frame F It is then followed by
a roll rotation fa about the axis y of F- giving a second intermediate frame F-
Final ly a spin rotation 7J about the axis z of F- is given resulting in the body
fixed frame Fj for the i lt l link Since all the three rotation matrices are orthogonal
it follows that FQ = C~1Fi = Cj F and hence
C CPi Cai ~ CH Sai + SrYi SPi Sai 5 7 i Sampi + Cli SPi Cci CPi Sai C1i Cai + 5 7 i SPi ~ S1 Ca + C 7 bull Sp Sai
(214) SliCPi CliCPi
Here Cx and Sx are abbreviations for cos(x) and sin(x) respectively
214 Inertial velocity of the ith link
mdash Lett ing pj = fj - f $j(5j and differentiating Eq(21) with respect to time gives
m- D j + T j f j + T^Si + T j pound + T j $ j 5 j (215)
Each of the terms appearing in Eq(215) must be addressed individually Since
fj represents the rigid body motion within the link it only has meaning for the
deployable tether and since yi = Zj = 0 for al l tether links (i even) fj = Ifi where 1
is the unit vector 1 0 0T For the case of rigid links (i odd) fj = 0
18
Evaluating the derivative of each angle in the rotation matrix gives
T^i9i = Taiai9i + Tppigi + T^fagi (216)
where Tx = J j I V Collecting the scaler angular velocity terms and defining the set
Vi = deg^gt A) 7i^ Eq(216) can be rewritten as
T i f t = Pi9i)fji (217)
where Pi(gi) is a column matrix defined by
Pi(9i)Vi = [Taigi^p^iT^gilffi (218)
Inserting Eq(29) and Eq(217) into Eq(215) and defining Sj = 2 + leads to
215 Cylindrical orbital coordinates
The Di term in Eq(219) describes the orbital motion of the t h l ink and
is composed of the three Cartesian coordinates DXi A ^ DZi However over the
cycle of one orbit DXi and Dyi vary dramatically by around the order of Earths
radius This must be avoided since large variations in the coordinates can cause
severe truncation errors during their numerical integration For this reason it is
more convenient to express the Cartesian components in terms of more stationary
variables This is readily accomplished using cylindrical coordinates
However it is only necessary to transform the first l inks orbital components
ie Lgti since only D is a generalized coordinates From Figure 2-2 it is apparent
(219)
that
(220)
19
Figure 2-2 Vector components in cylindrical orbital coordinates
20
Eq(220) can be rewritten in matrix form as
cos(0i) 0 0 s i n ( 0 i ) 0 0
0 0 1
Dri
h (221)
= DTlDSv
where Dri is the inplane radial distance 9 the true anomaly and DZl the out-of-
plane distance normal to the orbital plane The total derivative of D wi th respect
to time gives in column matrix form
1 1 1 1 J (222)
= D D l D 3 v
where [cos(^i) -Dnsm(8i) 0 ]
(223)
For all remaining links ie i = 2 N
cos(0i) - A - j S i n ^ i ) 0 sin(^i) D r i c o s ( 0 i ) 0
0 0 1
DTl = D D i = I d 3 i = 2N (224)
where is the n x n identity matrix and
2N (225)
216 Tether deployment and retrieval profile
The deployment of the tether is a critical maneuver for this class of systems
Cr i t i ca l deployment rates would cause the system to librate out of control Normally
a smooth sinusoidal profile is chosen (S-curve) However one would also like a long
tether extending to ten twenty or even a hundred kilometers in a reasonable time
say 3-4 orbits This is usually difficult or impossible to accomplish solely wi th one
S-curve profile Often a composite S-curve-Steady-S-curve profile is adopted Thus
21
the deployment scheme for the t h tether can be summarized as
k = Tr l - cos (^r(ti - to-)) 1 ltUlt h 2 r V A V
k = Vov tiilttiltt2i
k = IT j 1 - C deg S lkitl ^ ^ lt i lt ^
(226)
where to- tt are the ini t ia l and final times of the deployment or retrieval maneuver
respectively A t j = t - tQ mdash tt mdash t2- and the steady deployment velocity VQ is
calculated based on the continuity of l and l at the specified intermediate times t
and to For the case of
and
Ai = i(ti) - ^(t 0 -) = - ^ ( t 2 i ) (227)
2Ali + li(tf) -h(tQ)
V = lUli l^lL (228)
1 tf-tQ v y
A t = (229)
22 Kinetics and System Energy
221 Kinetic energy
The kinetic energy of the ith link is given by
Kei = lJm Rdrnfi^Ami (230)
Setting Eq(219) in a matrix-vector form gives the relation
4 ^ ^ PM) T i $ 4 TiSi]i[t (231)
22
where
m
k J
x l (232)
and nq = -nfj + 7 is the number of generalized coordinates in each link Inserting
Eq(231) into Eq(230) and integrating the kinetic energy for the t h l ink can be
rewritten as
1 T
(233)
where Mti is the links symmetric mass matrix
Mt =
m i D D T D D DDTPifgidmi) DDTTi J ^dm D n T Sidm rn
sym fP(9i)Pi(9i)dmi J P[g^T^idm J P ( pound ) T ^ m sym sym J ^f^idm f ^f^dmi sym sym sym J sfsidmi
pound Sfttiqxnq
(234)
and mi is the mass of the Ith link The kinetic energy for the entire system ie N
bodies can now be stated as
1 bull T bull 1 bull T Ke = J2 MtA = o laquo Mtqt 2 ^ -i
i=l (235)
where qt = qj^^ bull bull -qfNT a n d Mt is a block diagonal decoupled mass matrix
of the system with Mf- on its diagonal
(236) Mt =
0 0
0 Mt2
0
0 0
MH
0 0 0
0 0 0 bull MtN
222 Simplification for rigid links
The mass matrix given by Eq(234) simplifies considerably for the case of rigid
23
links and is given as
Mti
m D D l
T D D l DD^Piiffidmi) 0
sym 0
0
fP(fi)Pi(i)dmi 0 0 0
0 Id nfj
0 m~
i = 1 3 5 N
(237)
Moreover when mult iplying the matrices in the (11) block term of Eq(237) one
gets
D D l
1 D D l = 1 0 0 0 D n
2 0 (238) 0 0 1
which is a diagonal matrix However more importantly i t is independent of 6 and
hence is always non-singular ensuring that is always invertible
In the (12) block term of M ^ it is apparent that f fidmi is simply the
definition of the centre of mass of the t h rigid link and hence given by
To drrii mdash miff (239)
where fcmi is the position vector of the centre of mass of link i For the case of the
(22) block term a little more attention is required Expanding the integrand using
Eq(218) where ltj = fj gives
fj Ta Tafi fj T Q Tpfi f- T Q T 7 i f j
sym sym
Now since each of the terms of Eq(240) is a scaler or a 1 x 1 matrix it is therefore
equal to its trace ie sum of its diagonal elements thus
tr(ffTai
TTaifi) tr(ffTai
TTpfi) tr(ff Ta^T^fi)
fjT^Tpn rfTp^fi (240)
P(rj)Pj(fi) = syrri trifjTpTTpfi) t r ^ T ^ T ^
sym sym tr(ff T 7 T r 7 f j) J
(241)
Using two properties of the trace function[47] namely
tr(ABC) = tr(BCA) = trCAB) (242)
24
and
j tr(A) =tr(^J A^j (243)
where A B and C are real matrices it becomes quite clear that the (22) block term
of Eq(237) simplifies to
P(fi)Pii)dmi =
tr(Tai
TTaiImi) tTTai
TTpImi) t r (T Q r T 7 7 m ) trTaTTpImi) tr(Tp
TT7Jmi) sym sym sym
(244)
where Im- is an alternate form of the inertia tensor of the t h link which when
expanded is
Imlaquo = I nfi dm bull i sym
sym sym
J x^dmi J Xydmi XZdm sym J yi2dmi f yiZidm
sym J z^dmi J
xVi -L VH sym ijxxi + lyy IZZJ)2
(245)
Note this is in terms of moments of inertia and products of inertia of the t f l rigid
link A similar simplification can be made for the flexible case where T is replaced
with Qi resulting in a new expression for Im
223 Gravitational potential energy
The gravitational potential energy of the ith link due to the central force law
can be written as
V9i f dm f
-A mdash = -A bullgttradelti RdmA
dm
inn Lgti + T j ^ | (246)
where z = 3986 x 105 km 3s 2 (Earths gravitational constant) Expanding binomially
and retaining up to third order terms gives
Vgi ~ ~ D T 2D [J gfgidrrii + 23f T j gidm mdashDT
D2 i l I I Df Tt I grffdmiTiDi
(2-lt 7)
25
where D = D is the magnitude of the inertial position vector to the frame F The
integrals in the above expression are functions of the flexible mode shapes and can be
evaluated through symbolic manipulation using an existing package such as Maple
V In turn Maple V can translate the evaluated integrals into F O R T R A N and store
the code in a file Furthermore for the case of rigid bodies it can be shown quite
readily that
j gjgidrrii = j rjfidrrn = [IXXi + Iyy + IZZi)2 (248)
Similarly the other integrals in Eq(234) can be integrated symbolically
224 Strain energy
When deriving the elastic strain energy of a simple string in tension the
assumptions of high tension and low amplitude of transverse vibrations are generally
valid Consequently the first order approximation of the strain-displacement relation
is generally acceptable However for orbiting tethers in a weak gravitational-gradient
field neglecting higher order terms can result in poorly modelled tether dynamics
Geometric nonlinearities are responsible for the foreshortening effects present in the
tether These cannot be neglected because they account for the heaving motion along
the longitudinal axis due to lateral vibrations The magnitude of the longitudinal
oscillations diminish as the tether becomes shorter[35]
W i t h this as background the tether strain which is derived from the theory
of elastic vibrations[3248] is given by
(249)
where Uj V and W are the flexible deformations along the Xj and Z direction
respectively and are obtained from Eq(25) The square bracketed term in Eq(249)
represents the geometric nonlinearity and is accounted for in the analysis
26
The total strain energy of a flexible link is given by
Ve = I f CiEidVl (250)
Substituting the stress-strain relationship 0 = E(poundi the strain energy is now given
by
Vei=l-EiAi J l l d x h (251)
where E^ is the tethers Youngs modulus and A is the cross-sectional area The
tether is assumed to be uniform thus EA which is the flexural stiffness of the
tether is assumed to be constant Eq(251) is evaluated symbolically for arbitrary
number of modes
225 Tether energy dissipation
The evaluation of the energy dissipation due to tether deformations remains
a problem not well understood even today Here this complex phenomenon is repshy
resented through a simplified structural damping model[32] In addition the system
exhibits hysteresis which also must be considered This is accomplished using an
augmented complex Youngs modulus
EX^Ei+jEi (252)
where Ej^ is the contribution from the structural damping and j = The
augmented stress-strain relation is now given as
a = Epoundi = Ei(l+jrldi)ei (253)
where 77^ = EjE^ ltC 1 is defined as the structural damping coefficient determined
experimentally [49]
27
If poundi is a harmonic function with frequency UJQ- then
jet = (254) wo-
Substituting into Eq(253) and rearranging the terms gives
e + I mdash I poundi = o-i + ad (255)
where ad and a are the stress with and without structural damping respectively
Now the Rayleigh dissipation function[50] for the ith tether can be expressed as
(ii) Jo
_ 1 EjAiVdi rU (2-56^ 2 w 0 Jo
The strain rate ii is the time derivative of Eq(249) The generalized external force
due to damping can now be written as Qd = Q^ bull bull bull QdNT where
dWd
^ = - i p = 2 4 ( 2 5 7 )
= 0 i = 13 N
23 0(N) Form of the Equations of Mot ion
231 Lagrange equations of motion
W i t h the kinetic energy expression defined and the potential energy of the
whole system given by Pe = ]Cn=l( 5i+ ej) the equations of motion can be obtained
quite readily using the Lagrangian principle
where q is the set of generalized coordinates to be defined in a later section
Substituting Eqs(235247251 and 257) into Eq(258) leads to the familiar
matr ix form of the non-linear non-autonomous coupled equations of motion for the
28
system
Mq t)t+ F(q q t) = Q(q q t) (259)
where M(q t) is the nonlinear symmetric mass matrix and F(q q t) is the forcing
function which can be written in matrix form as
Q(ltfgt Qi 0 is the vector of non-conservative generalized forces including control inshy
puts acting on the system A detailed expansion of Eq(260) in tensor notation is
developed in Appendix I
Numerical solution of Eq(259) in terms of q would require inversion of the
mass matrix However M is a full matrix of size nqq x nqq where nqq = Nnq is the
total number of generalized coordinates Therefore direct inversion of M would lead
to a large number of computation steps of the order nqq^ or higher The objective
here is to develop an order-N O(N) algorithm for obtaining M _ 1 that minimizes
the computational effort This is accomplished through the following transformations
ini t ia l ly developed for the stationkeeping case by Pradhan et al [30]
232 Generalized coordinates and position transformation
During the derivation of the energy expressions for the system the focus was
on the decoupled system ie each link was considered independent of the others
Thus each links energy expression is also uncoupled However the constraints forces
between two links must be incorporated in the final equations of motion
Let
e Unqxl (261) m
be the set of coupled coordinates of the t h l ink such that q = qT q2 bull bull bull 0T i s the
29
systems generalized coordinates Let qt- be the set of auxiliary decoupled coordinates
defined by Eq(232) such that qt = qj^ qj^ Qtj^1 bull The only difference between
qtj and q is the presence of D against d is the position of F wi th respect to
the inertial frame FQ whereas d is defined as the offset position of F relative to and
projected on the frame It can be expressed as d = dXidyidZi^ For the
special case of link 1 D = d
From Figure 2-1 can written as
Di = A - l + T i _ i J i _ i + di) + T V i S i - i f t - i + dXi)8i- (262)
Denoting KAj = $j(Zj + dx- ) Eq(262) can be rewritten in summation form as
i ampi = H ( T i - 1 ^ + T i - l K A i - l - l + T j - l ^ - l ) bull ( 2 - 6 3 )
3 = 1
Introducing the index substitution k mdash j mdash 1 into Eq(263)
i-1 D = ^2 (Tfc_ilt4 + TkKAk6k + Tkilk) + T^di (264)
k=l
since d = D TQ = 1^ and IQ DQ SQ are null vectors Recasting Eq(264) in
matr ix form leads to i-1
Qt Jfe=l
where
and
For the entire system
T fc-1 0 TkKAk
0 0 0 0 0 0 0 0 0 0 0 0 -
r T i - i 0 0 0 0 0 0 0 0 0 G
0 0 0 1
Qt = RpQ
G R n lt x l
(265)
(266)
G Unq x l (267)
(268)
30
where
RP
R 1 0 0 R R 2 0 R RP
2 R 3
0 0 0
R RP R PN _ 1 RN J
(269)
Here RP is the lower block triangular matrix which relates qt and q
233 Velocity transformations
The two sets of velocity coordinates qti and qi are related by the following
transformations
(i) First Transformation
Differentiating Eq(262) with respect to time gives the inertial velocity of the
t h l ink as
D = A - l + Ti_iK + + K A i _ i ^ _ i
+ T j _ i ^ + Ti-ii + T i - i K A i - i ^ + T j - i K A ^ i ^ i (270)
(271) (
Defining K A D = dddXi+1KAi K A L = d d K A and hi = di+1 +lfi +KA^i
results in K A ^ i = K A D j i o ^ + K A L i _ i i _ i
= KADi_iiTdi + K A L j _ i ^ _ i
Inserting Eq(271) into Eq(270) and rearranging gives
= A _ + T i ^ d g + K A D ^ x ^ i ^ ^ + Pi-^hi-iH-i
+ T i _ i K A i _ i 5 _ i + T i_ i2 + K A L j _ i 5 j _ i 4 _ i (272)
Using Eq(272) to define recursively - D j _ i and applying the index substitution
c = j mdash 1 as in the case of the position transformation it follows that
Di = ^ T ^ i K D ^ ^ + Pk(hk)ffk + T f c K A f c ^ + TfcKLfc4 Jfc=l (273)
+ T j _ i K D j _ i lt i ii
31
where K D j _ i = + K A D ^ i ^ - i r 7 and K L j = i + K A L j 5 j Finally by following a
procedure similar to that used for the position transformation it can be shown that
for the entire system
t = Rvil (2-74)
where Rv is given by the lower block diagonal matrix
Here
Rv
Rf 0 0 R R$ 0 R Rl RJ
0 0 0
T3V nN-l
R bullN
e Mnqqxl
Ri =
T i _ i K D i _ i TiKAi T j K L j 0 0 0 0 bullo 0 0 0 0 0 0 0
e K n lt x l
RS
T i-1 0 0 o-0 0 0 0 0 0
0 0 0 1
G Mnq x l
(275)
(276)
(277)
(ii) Second Transformation
The second transformation is simply Eq(272) set into matrix form This leads
to the expression for qt as
(278)
where Id Pihi) T j K A j T j K L j 0 0 0 0 0 0
0 0
0 0
0 0
G $lnq x l
Thus for the entire system
(279)
pound = Rnjt + Rdq
= Vdnqq - R nrlRdi (280)
32
where
and
- 0 0 0 RI 0 0
Rn = 0 0
0 0
Rf 0 0 0 rgtd K2 0
Rd = 0 0 Rl
0 0 0
RN-1 0
G 5R n w x l
0 0 0 e 5 R N ^ X L
(281)
(282)
234 Cylindrical coordinate modification
Because of the use of cylindrical coorfmates for the first body it is necessary
to slightly modify the above matrices to account for this coordinate system For the
first link d mdash D = D ^ D S L thus
Qti = 01 =gt DDTigi (283)
where
D D T i
Eq(269) now takes the form
DTL 0 0 0 d 3 0 0 0 ID
0 0 nfj 0
RP =
P f D T T i R2 0 P f D T T i Rp
2 R3
iJfDTTi i^
x l
0 0 0
R PN-I RN
(284)
(285)
In the velocity transformation a similar modification is necessary since D =
DD^DS^ Mak ing the substitution it can be shown that Eq(276) and Eq(279) now
33
takes the form for i = 1
T)V _ pn _ r t j mdash rt^ mdash 0 0 0 0
0 0 0 0 0 0 0 0
(286)
The rest of the matrices for the remaining links ie i = 2 N are unchanged
235 Mass matrix inversion
Returning to the expression for kinetic energy substitution of Eq(274) into
Eq(235) gives the first of two expressions for kinetic energy as
Ke = -q RvT MtRv (287)
Introduction of Eq(280) into Eq(235) results in the second expression for Ke as
1X 2q (hnqq-RnrlRd Mt (idnqq - R nrlRd (288)
Thus there are two factorizations for the systems coupled mass matrix given as
M = RV MtRv
M (hnqq - RnrlRd Mt ( i d n q q - R n r 1 R d
(289)
(290)
Inverting Eq(290)
r-1 _ ( rgtd~^ M 1 = (Ra) l d n q q - Rn)Mf1 (Rd)~l(Idnqq - Rn) (291)
Since both Rd and Mt are block diagonal matrices their inverse is simply the inverse
34
of each block on the diagonal ie
Mr1
h 0 0
0 0
0 0
0 0 0
0
0
0
tN
(292)
Each block has the dimension nq x nq and is independent of the number of bodies
in the system thus the inversion of M is only linearly dependent on N ie an O(N)
operation
236 Specification of the offset position
The derivation of the equations of motion using the O(N) formulation requires
that the offset coordinate d be treated as a generalized coordinate However this
offset may be constrained or controlled to follow a prescribed motion dictated either by
the designer or the controller This is achieved through the introduction of Lagrange
multipliers[51] To begin with the multipliers are assigned to all the d equations
Thus letting f = F - Q Eq(259) becomes
Mq + f = P C A (293)
where A G K 3 - 7 V x l is the vector of Lagrange multipliers and Pc G s R n 9 x 3 A r is the
permutation matrix assigning the appropriate Aj to its corresponding d equation
Inverting M and pre-multiplying both sides by PcT gives
pcTM-pc
= ( dc + PdegTM-^f) (294)
where dc is the desired offset acceleration vector Note both dc and PcTM lf are
known Thus the solution for A has the form
A pcTM-lpc - l dc + PC1 M (295)
35
Now substituting Eq(295) into Eq(293) and rearranging the terms gives
q = S(f Ml) = -M~lf + M~lPc
which is the new constrained vector equation of motion with the offset specified by
dc Note that pre-multiplying M~l by P c T provides the rows of M _ 1 corresponding
to the d equations Similarly post-multiplying by Pc leads the columns of the matrix
corresponding to the d equations
24 Generalized Control Forces
241 Preliminary remarks
The treatment of non-conservative external forces acting on the system is
considered next In reality the system is subjected to a wide variety of environmental
forces including atmospheric drag solar radiation pressure Earth s magnetic field as
well as dissipative devices to mention a few However in the present model only
the effect of active control thrusters and momemtum-wheels is considered These
generalized forces wi l l be used to control the attitude motion of the system
In the derivation of the equations of motion using the Lagrangian procedure
the contribution of each force must be properly distributed among all the generalized
coordinates This is acheived using the relation
nu ^5 Qk=E Pern ^ (2-97)
m=l deg Q k
where Qk and qk are the kth elements of Q and ltf respectively[51] Fem is the
mth external force acting on the system at the location Rm from the inertial frame
and nu is the total number of actuators Derivation of the generalized forces due to
thrusters and momentum wheels is given in the next two subsections
pcTM-lpc (dc + PcTM- Vj (2-96)
36
242 Generalized thruster forces
One set of thrusters is placed on each subsatellite ie al l the rigid bodies
except the first (platform) as shown in Figure 2-3 They are capable of firing in the
three orthogonal directions From the schematic diagram the inertial position of the
thruster located on body 3 is given as
Defining
R3 = di + Tid2 + T2(hi + KA2lt52) + T 3 f c m 3
= d + dfY + T 3 f c m 3
df = Tjdj+i + Tj+1lj+ii + KAj+i5j+i)
the thruster position for body 5 is given as
(298)
(299)
R5 = di+ dfx + d 3 + T5fcm
Thus in general the inertial position for the thruster on the t h body can be given
(2100)
as i - 2
k=l i1 cm (2101)
Let the column matrix
[Q i iT 9 R-K=rR dqj lA n i
WjRi A R i (2102)
where i = 3 5 N and j mdash 1 2 i The vector derivative of R wi th respect to
the scaler components qk is stored in the kth column of Thus for the case of
three bodies Eq(297) for thrusters becomes
1 Q
Q3
3 -1
T 3 T t 2 (2103)
37
Satellite 3
mdash gt -raquo cm
D1=d1
mdash gt
a mdash gt
o
Figure 2-3 Inertial position of subsatellite thruster forces
38
where ft = 0TtaTt^T is the thrust vector for the t h l ink (tether i-1) Here
Ttn and Tta are the thrust forces along the inplane m and out-of-plane Z directions i Pi
respectively For the case of 5-bodies
Q
~Qgt1
0
0
( T ^ 2 ) (2104)
The result for seven or more bodies follows the pattern established by Eq(2104)
243 Generalized momentum gyro torques
When deriving the generalized moments arising from the momemtum-wheels
on each rigid body (satellite) including the platform it is easier to view the torques as
coupled forces From Figure 2-4 it is apparent that for the ith l ink the generalized
forces constituting the couple are
(2105)
where
Fei = Faij acting at ei = eXi
Fe2 = -Fpk acting at e 2 = C a ^ l
F e 3 = F7ik acting at e 3 = eyj
Expanding Eq(2105) it becomes clear that
3
Qk = ^ F e i
i=l
d
dqk (2106)
which is independent of Ri- Thus the moments Mma = 2FaieXi Mm^ = 2FpeXi
and Mmi = 2FlieVi Transforming the coordinates to the body fixed frame using
the rotation matrix the generalized force due to the momentum-wheels on link i can
39
Figure 2-4 Coupled force representation of a momentum-wheel on a rigid body
40
be written as
d d d Qi = Td bull ^ T i t M m a - Tik bull mdashT^Mmp + T ^ bull T J j M ^
QiMmi
where Mm = Mma M m Mmi T
lt3 T J ^ T - T i f c bull ^ T i pound T ^ - ^ T z j
(2107)
(2108)
Note combining the thruster forces and momentum-wheel torques a compact
expression for the generalized force vector for 5 bodies can be written as
Q
Q T 3 0 Q T 5 0
0 0 0
0 lt33T3 Qyen Q5
3T5 0
0 0 0 Q4T5 0
0 0 0 Q5
5T5
M M I
Ti 2 m 3
Quu (2109)
The pattern of Eq(2109) is retained for the case of N links
41
3 C O M P U T E R I M P L E M E N T A T I O N
31 Prel iminary Remarks
The governing equations of motion for the TV-body tethered system were deshy
rived in the last chapter using an efficient O(N) factorization method These equashy
tions are highly nonlinear nonautonomous and coupled In order to predict and
appreciate the character of the system response under a wide variety of disturbances
it is necessary to solve this system of differential equations Although the closed-
form solutions to various simplified models can be obtained using well-known analytic
methods the complete nonlinear response of the coupled system can only be obtained
numerically
However the numerical solution of these complicated equations is not a straightshy
forward task To begin with the equations are stiff[52] be they have attitude and
vibration response frequencies separated by about an order of magnitude or more
which makes their numerical integration suceptible to accuracy errors i f a properly
designed integration routine is not used
Secondly the factorization algorithm that ensures efficient inversion of the
mass matrix as well as the time-varying offset and tether length significantly increase
the size of the resulting simulation code to well over 10000 lines This raises computer
memory issues which must be addressed in order to properly design a single program
that is capable of simulating a wide range of configurations and mission profiles It
is further desired that the program be modular in character to accomodate design
variations with ease and efficiency
This chapter begins with a discussion on issues of numerical and symbolic in-
42
tegration This is followed by an introduction to the program structure Final ly the
validity of the computer program is established using two methods namely the conshy
servation of total energy and comparison of the response results for simple particular
cases reported in the literature
32 Numerical Implementation
321 Integration routine
A computer program coded in F O R T R A N was written to integrate the equashy
tions of motion of the system The widely available routine used to integrate the difshy
ferential equations D G E A R [ 5 3 ] is based on Gears predictor-corrector method[54]
It is well suited to stiff differential equations as i t provides automatic step-size adjustshy
ment and error-checking capabilities at each iteration cycle These features provide
superior performance over traditional 4^ order Runge-Kut ta methods
Like most other routines the method requires that the differential equations be
transformed into a first order state space representation This is easily accomplished
by letting
thus
X - ^ ^ f ^ ) =ltbullbull) (3 -2 )
where S(fMdc) is defined by Eq(296) Eq(32) represents the new set of 2nqq
first order differential equations of motion
322 Program structure
A flow chart representing the computer programs structure is presented in
Figure 3-1 The M A I N program is responsible for calling the integration routine It
also coordinates the flow of information throughout the program The first subroutine
43
called is INIT which initializes the system variables such as the constant parameters
(mass density inertia etc) and the ini t ial conditions I C (attitude angles orbital
motion tether elastic deformation etc) from user-defined data files Furthermore all
the necessary parameters required by the integration subroutine D G E A R (step-size
tolerance) are provided by INIT The results of the simulation ie time history of the
generalized coordinates are then passed on to the O U T P U T subroutine which sorts
the information into different output files to be subsequently read in by the auxiliary
plott ing package
The majority of the simulation running time is spent in the F C N subroutine
which is called repeatedly by D G E A R F C N generates the fv(Xi) vector defined in
Eq(32) It is composed of several subroutines which are responsible for the comshy
putation of all the time-varying matrices vectors and scaler variables derived in
Chapter 2 that comprise the governing equations of motion for the entire system
These subroutines have been carefully constructed to ensure maximum efficiency in
computational performance as well as memory allocation
Two important aspects in the programming of F C N should be outlined at this
point As mentioned earlier some of the integrals defined in Chapter 2 are notably
more difficult to evaluate manually Traditionally they have been integrated numershy
ically on-line (during the simulation) In some cases particularly stationkeeping the
computational cost is quite reasonable since the modal integrals need only be evalushy
ated once However for the case of deployment or retrieval where the mode shape
functions vary they have to be integrated at each time-step This has considerable
repercussions on the total simulation time Thus in this program the integrals are
evaluated symbolically using M A P L E V[55] Once generated they are translated into
F O R T R A N source code and stored in I N C L U D E files to be read by the compiler
Al though these files can be lengthy (1000 lines or more) for a large number of flexible
44
INIT
DGEAR PARAM
IC amp SYS PARAM
MAIN
lt
DGEAR
lt
FCN
OUTPUT
STATE ACTUATOR FORCES
MOMENTS amp THRUST FORCES
VIBRATION
r OFFS DYNA
ET MICS
Figure 3-1 Flowchart showing the computer program structure
45
modes they st i l l require less time to compute compared to their on-line evaluation
For the case of deployment the performance improvement was found be as high as
10 times for certain cases
Secondly the O(N) algorithm presented in the last section involves matrices
that contained mostly zero terms (RP Rv and Mplusmn) The sparse structure of these mashy
trices can be used advantageously to improve the computational efficiency by avoiding
multiplications of zero elements The result is a quick evaluation of fv(Xt) Moreshy
over there is a considerable saving of computer memory as zero elements are no
longer stored
Final ly the C O N T R O L subroutine which is part of F C N is designed to calshy
culate the appropriate control forces torques and offset dynamics which are required
to stabilize the librational ( A T T I T U D E ) and vibrational ( V I B R A T I O N ) motion of
the system
33 Verification of the Code
The simulation program was checked for its validity using two methods The
first verified the conservation of total energy in the absence of dissipation The second
approach was a direct comparison of the planar response generated by the simulation
program with those available in the literature A s numerical results for the flexible
three-dimensional A^-body model are not available one is forced to be content with
the validation using a few simplified cases
331 Energy conservation
The configuration considered here is that of a 3-body tethered system with
the following parameters for the platform subsatellite and tether
46
h =
1091430 -8135 328108 -8135 8646050 27116 328108 27116 8286760
kg bull m 2 platform inertia
kg bull m 2 end-satellite inertia 200 0 0
0 400 0 0 0 400
m = 90000 kg mass of the space station platform
m2 = 500 kg mass of the end-satellite
EfAt = 61645 N tether elastic stiffness
pt mdash 49 kg km tether linear density
The tether length is taken to remain constant at 10 km (stationkeeping case)
A n in i t ia l disturbance in pitch and roll of 2deg and 1deg respectively is given to both
the rigid bodies and the flexible tether In addition the tether is ini t ia l ly deflected
by 45 m in the longitudinal direction at its end and 05 m in both the inplane and
out-of-plane directions in the first mode The tether attachment point offset at the
platform (satellite 1) end is 1 m in all the three directions ie d2 = 11 lTm The
attachment point at the subsatellite (satellite 2) end is 1 m in the x direction ie
fcm3 = 10 0T
The variation in the kinetic and potential energies is presented for the above-
mentioned case in Figure 3-2(a) As seen from the plot there is a continuous mutual
exchange between the kinetic and potential energies however the variation in the
total energy remains zero (Figure 3-2b) It should be noted that after 1 orbit the
variation of both kinetic and potential energy does not return to 0 This is due to
the attitude and elastic motion of the system which shifts the centre of mass of the
tethered system in a non-Keplerian orbit
Similar results are presented for the 5-body case ie a double pendulum (Figshy
ure 3-3) Here the system is composed of a platform and two subsatellites connected
47
STSS configuration 3-Body a 1(0)=a 2(0)=a t(0)=2deg P1(0)=p2(0)=p(0)=1deg 5x(0)=45m 8y(0)=52(0)=05m dx(0)=dy(0)=d2(0)=1m Stationkeeping lt=10km
Variation of Kinetic and Potential Energy
(a) Time (Orbits) Percent Variation in Total Energy
OOEOh
-50E-12
-10E-11 I i i
0 1 2 (b) Time (Orbits)
Figure 3-2 Kinetic and potential energy transfer for the three-body platform based tethered satellite system (a) variation of kinetic and potential energy (b) percent change in total energy of system
48
STSS configuration 5-Body a 1(0)=a 2(0)=a t 1(0)=2deg a 3(0)=a t 2(0)=25deg P1(0)=p2(0)=p3(0)=pt1(0)=pt2(0)=1lt
5x(0)=5m 8y(0)=5z(0)=05m dx(0)=dy(0)=dz(0)=0 Stationkeeping 1 =12=1 Okm
(a)
20E6
10E6r-
00E0
-10E6
-20E6
Variation of Kinetic and Potential Energy
-
1 I I I 1 I I
I I
AP e ~_
i i 1 i i i i
00E0
UJ -50E-12 U J
lt
Time (Orbits) Percent Variation in Total Energy
-10E-11h
(b) Time (Orbits)
Figure 3 -3 Kinet ic and potential energy transfer for the five-body platform based tethered satellite system (a) variation of kinetic and potential enshyergy (b) percent change in total energy of system
49
in sequence by two tethers The two subsatellites have the same mass and inertia
ie 7713 = 7 7 i 2 a n d I3 = I21 a n d are separated by identical tether segments each 10
k m in length The platform has the same properties as before however there is no
tether attachment point offset present in the system (d = fcmi = 0)
In al l the cases considered energy was found to be conserved However it
should be noted that for the cases where the tether structural damping deployment
retrieval attitude and vibration control are present energy is no longer conserved
since al l these features add or subtract energy from the system
332 Comparison with available data
The alternative approach used to validate the numerical program involves
comparison of system responses with those presented in the literature A sample case
considered here is the planar system whose parameters are the same as those given
in Section 331 A n ini t ia l pitch disturbance of 2deg is imparted to both the platform
and the tether together with an ini t ial tether deformation of 08 m and 001 m in
the longitudinal and transverse directions respectively in the same manner as before
The tether length is held constant at 5 km and it is assumed to be connected to the
centre of mass of the rigid platform and subsatellite The response time histories from
Ref[43] are compared with those obtained using the present program in Figures 3-4
and 3-5 Here ap and at represent the platform and tether pitch respectively and B i
C i are the tethers longitudinal and transverse generalized coordinates respectively
A comparison of Figures 3-4 and 3-5 clearly shows a remarkable correlation
between the two sets of results in both the amplitude and frequency The same
trend persisted even with several other comparisons Thus a considerable level of
confidence is provided in the simulation program as a tool to explore the dynamics
and control of flexible multibody tethered systems
50
a p(0) = 2deg 0(0) = 2C
6(0) = 08 m
0(0) = 001 m
Stationkeeping L = 5 km
D p y = 0 D p 2 = 0
a oo
c -ooo -0 01
Figure 3 - 4 Simulation results for the platform based three-body tethered system originally presented in Ref[43]
51
STSS configuration 3-Body a1(0)=a2(0)=at(0)=2deg p1(0)=(32(0)=(3t(0)=0 8x(0)=08m 5y(0)=001m 5Z(0)=0 dx(0)=dy(0)=dz(0)=0 Stationkeeping l=5km
Satellite 1 Pitch Angle Tether Pitch Angle
i -i i- i i i i i i d
1 2 0 1 2 Time (Orbits) Time (Orbits)
Longitudinal Vibration
^MAAAAAAAWVWWWWVVVW
V bull i bull i I i J 00 05 10 15 20
Time (Orbits) Transverse Vibration
Time (Orbits)
Figure 3-5 Simulation results for the platform based three-body tethered system obtained using the present computer program
52
4 D Y N A M I C SIMULATION
41 Prel iminary Remarks
Understanding the dynamics of a system is critical to its design and develshy
opment for engineering application Due to obvious limitations imposed by flight
tests and simulation of environmental effects in ground based facilities space based
systems are routinely designed through the use of numerical models This Chapter
studies the dynamical response of two different tethered systems during deployment
retrieval and stationkeeping phases In the first case the Space platform based Tethshy
ered Satellite System (STSS) which consists of a large mass (platform) connected to
a relatively smaller mass (subsatellite) with a long flexible tether is considered The
other system is the O E D I P U S B I C E P S configuration involving two comparable mass
satellites interconnected by a flexible tether In addition the mission requirement of
spin about an arbitrary axis the tether axis for O E D I P U S - A C and cartwheeling or
spin about the orbit normal for the proposed B I C E P S mission is accounted for
A s mentioned in Chapter 2 the tether flexibility is modelled using the asshy
sumed mode discretization method Although the simulation program can account
for an arbitrary number of vibrational modes in each direction only the first mode
is considered in this study as it accounts for most of the strain energy[34] and hence
dominates the vibratory motion The parametric study considers single as well as
double pendulum type systems with offset of the tether attachment points The
systems stability is also discussed
42 Parameter and Response Variable Definitions
The system parameters used in the simulation unless otherwise stated are
53
selected as follows for the STSS configuration
1091430 -8135 328108 -8135 8646050 27116 328108 27116 8286760
h =
h = 200 0 0
0 400 0 0 0 400
bull m = 90000 kg (mass of the space platform)
kg bull m 2 (platform inertia) 760 J
kg bull m 2 (subsatellite inertia)
bull 7772 = 500 kg (mass of the subsatellite)
bull EtAt = 61645 N (tether stiffness)
bull Pt = 49 k g k m (tether density)
bull Vd mdash 05 (tether structural damping coefficient)
bull fcm^ = l 0 0 r m (tether attachment point at the subsatellite)
The response variables are defined as follows
bull ai0 satellite 1 (platform) pitch and roll angles respectively
bull CK2 2 satellite 2 (subsatellite) pitch and roll angles respectively
bull at fit- tether pitch and roll angles respectively
bull If tether length
bull d = dx dy dzT- tether attachment position relative to satellite 1 (platform)
mdash
bull S = 8X 8y Sz tether flexible generalized coordinates in the longitudinal x
inplane transverse y and out-of-plane transverse z directions respectively
The attitude angles and are measured with respect to the Local Ve r t i c a l -
Loca l Horizontal ( L V L H ) frame A schematic diagram illustrating these variable
is presented in Figure 4-1 The system is taken to be in a nominal circular orbit at
an altitude of 289 km with an orbital period of 903 minutes
54
Satellite 1 (Platform) Y a w
Figure 4-1 Schematic diagram showing the generalized coordinates used to deshyscribe the system dynamics
55
43 Stationkeeping Profile
To facilitate comparison of the simulation results a reference case based on
the three-body STSS system with zero offset (d = 0 r c m 3 = 0) and a tether length
of It = 20 km is considered first (Figure 4-2) The system is subjected to an ini t ial
disturbance in pitch and roll of 2deg and 1deg respectively to all the three bodies
In addition an ini t ia l longitudinal deflection of 14 m from the tethers unstretched
position is given together with a transverse inplane and out-of-plane deflection of 1
m at the tethers mid-length From the attitude response given in Figure 4-2(a) it is
apparent that the three bodies oscillate about their respective equilibrium positions
Since coupling between the individual rigid-body dynamics is absent (zero offset) the
corresponding librational frequencies are unaffected Satellite 1 (Platform) displays
amplitude modulation arising from the products of inertia The result is a slight
increase of amplitude in the pitch direction and a significantly larger decrease in
amplitude in the roll angle
Figure 4-2(b) clearly shows coupling of the tethers attitude motion with its
flexibility dynamics The tether vibrates in the axial direction about its equilibrium
position (at approximately 138 m) with two vibration frequencies namely 012 Hz
and 31 x 1 0 - 4 Hz The former is the first longitudinal flexible mode which is dissishy
pated through the structural damping while the latter arises from the coupling with
the pitch motion of the tether Similarly in the transverse direction there are two
frequencies of oscillations arising from the flexible mode and its coupling wi th the
attitude motion As expected there is no apparent dissipation in the transverse flexshy
ible motion This is attributed to the weak coupling between the longitudinal and
transverse modes of vibration since the transverse strain in only a second order effect
relative to the longitudinal strain where the dissipation mechanism is in effect Thus
there is very litt le transfer of energy between the transverse and longitudinal modes
resulting in a very long decay time for the transverse vibrations
56
STSS configuration 3-Body a1(0)=a2(0)=a(0)=2deg (31(0)=p2(0)=p(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=dz(0)=0 Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
0 1 2 3 4 5 0 1 2 3 4 5 Time (Orbits) Time (Orbits)
Satellite 2 Pitch Angle Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-2 Stationkeeping dynamics of the three-body STSS configuration withshyout offset (a) attitude response
57
STSS configuration 3-Body a1(0)=cx2(0)=cx(0)=2deg P1(0)=p2(0)=p(0)=1deg 5x(0)=14m8y(0)=82(0)=1m dx(0)=dy(0)=dz(0)=0 Stationkeeping lt=20km
Tether Longitudinal Vibration 16| 1 1 1 1 1 r-
Time (Orbits) Tether In-Plane Transverse Vibration
imdash 1 mdash 1 mdash 1 mdash 1 mdashr
r I I 1 1 1 I 1 1 1 1 I 1 bull ^ bull bull I I J 0 1 2 3 4 5
Time (Orbits) Tether Out-of-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-2 Stationkeeping dynamics of the three-body STSS configuration withshyout offset (b) vibration response
58
Introduction of the attachment point offset significantly alters the response of
the attitude motion for the rigid bodies W i t h a i m offset along the local vertical
at the platform (dx = 1 m) and fcm^ = 100 m coupling between the tether
and end-bodies is established by providing a lever arm from which the tether is able
to exert a torque that affects the rotational motion of the rigid end-bodies (Figure
4-3) Bo th the platform and subsatellite now oscillate at the same frequency as the
tether whose motion remains unaltered In the case of the platform there is also an
amplitude modulation due to its non-zero products of inertia as explained before
However the elastic vibration response of the tether remains essentially unaffected
by the coupling
Providing an offset along the local horizontal direction (dy = 1 m) results in
a more dramatic effect on the pitch and roll response of the platform as shown in
Figure 4-4(a) Now the platform oscillates about its new equilibrium position of
-90deg The roll motion is also significantly disturbed resulting in an increase to over
10deg in amplitude On the other hand the rigid body dynamics of the tether and the
subsatellite as well as the tether flexible motion remain the same (Figure 4-4b)
Figure 4-5 presents the response when the offset of 1 m at the platform end is
along the z direction ie normal to the orbital plane The result is a larger amplitude
pitch oscillation about the reference equilibrium position as shown in Figure 4-5(a)
while the roll equilibrium is now shifted to approximately 90deg Note there is little
change in the attitude motion of the tether and end-satellites however there is a
noticeable change in the out-of-plane transverse vibration of the tether (Figure 4-5b)
which may be due to the large amplitude platform roll dynamics
Final ly by setting a i m offset in all the three direction simultaneously the
equil ibrium position of the platform in the pitch and roll angle is altered by approxishy
mately 30deg (Figure 4-6a) However the response of the system follows essentially the
59
STSS configuration 3-Body a1(0)=cx2(0)=cxt(0)=2deg p1(0)=(32(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=1m dy(0)=dz(0)=0 Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
Time (Orbits) Time (Orbits) Tether Pitch Angle Tether Roll Angle
L bull I I bull bull lt bull bull bull raquo bull bull bull bull -I h I I I I i H 0 1 2 3 4 5 0 1 2 3 4 5
Time (Orbits) Time (Orbits) Satellite 2 Pitch Angle Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-3 Stationkeeping dynamics of the three-body STSS configuration with offset along the local vertical (a) attitude response
60
STSS configuration 3-Body a 1(0)=a 2(0)=a t(0)=2deg (31(0)=p2(0)=pt(0)=1deg 5x(0)=14m5y(0)=5z(0)=1m dx(0)=1mdy(0)=dz(0)=0 Stationkeeping lt=20km
Tether Longitudinal Vibration 1 6 i mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash mdash i mdash i mdash mdash mdash mdash mdash i mdash mdash lt -
Time (Orbits)
Tether In-Plane Transverse Vibration
y bull i i i i i i i i i i i i i i i i i 0 1 2 3 4 5
Time (Orbits)
Tether Out-of-Plane Transverse Vibration r 1 1 1 1 1 1 1 1 1 1 1 1 1 1 q
h i I i i I i i i i I i i i i 3
0 1 2 3 4 5 (b) Time (Orbits)
Figure 4-3 Stationkeeping dynamics of the three-body STSS configuration with offset along the local vertical (b) vibration response
61
STSS configuration 3-Body a1(0)=cx2(0)=o(0)=2deg (31(0)=p2(0)=p(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dz(0)=0 dy(0)=1m Stationkeeping lt=20km
Satellite 1 Pitch Angle i 1 i
Satellite 1 Roll Angle
1 2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 4 Time (Orbits)
Tether Roll Angle
CD
edT
1 2 3 Time (Orbits)
Satellite 2 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 2 Roll Angle
CD
C O
(a)
Figure 4-
1 2 3 4 Time (Orbits)
1 2 3 4 Time (Orbits)
4 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal (a) attitude response
62
STSS configuration 3-Body a1(0)=cx2(0)=ot(0)=2deg p1(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dz(0)=0 dy(0)=1m Stationkeeping lt=20km
Tether Longitudinal Vibration 16| 1 gt 1 bull 1 1 1
0 1 2 3 4 5
Time (Orbits) Tether In-Plane Transverse Vibration
i mdash 1 mdash 1 mdash 1 mdash 1 mdash i mdash mdash lt mdash 1 mdash 1 mdash i mdash mdash 1 mdash 1 mdash
V I I bull bull bull I i i i I i i 1 1 3
0 1 2 3 4 5 Time (Orbits)
Tether Out-of-Plane Transverse Vibration r 1 1 1 1 1 1 1 1 1 1 1 gt bull | q
V i 1 1 1 1 1 mdash 1 mdash 3
0 1 2 3 4 5 (b) Time (Orbits)
Figure 4-4 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal (b) vibration response
63
STSS configuration 3-Body a1(0)=o2(0)=at(0)=2deg (31(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=0 dz(0)=1m Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 4 Time (Orbits)
Tether Roll Angle
1 2 3 4 Time (Orbits)
Satellite 2 Pitch Angle
2 3 4 Time (Orbits)
Satellite 2 Roll Angle
(a)
Figure 4-
1 2 3 4 Time (Orbits)
2 3 4 Time (Orbits)
5 Stationkeeping dynamics of the three-body STSS configuration with offset along the orbit normal (a) attitude response
64
STSS configuration 3-Body a1(0)=cx2(0)=ot(0)=2deg P1(0)=p2(0)=(3(0)=1deg 8x(0)=14m8y(0)=82(0)=1m dx(0)=dy(0)=0 d2(0)=1m Stationkeeping lt=20km
Tether Longitudinal Vibration 1 6 i mdash r mdash i mdash mdash i mdash i mdash mdash i mdash i mdash mdash i mdash mdash i mdash 1 mdash 1 mdash i mdash mdash lt mdash lt mdash lt mdash r
Time (Orbits) Tether In-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-5 Stationkeeping dynamics of the three-body STSS configuration with offset along the orbit normal (b) vibration response
65
same trend with only minor perturbations in the transverse elastic vibratory response
of the tether (Figure 4-6b)
44 Tether Deployment
Deployment of the tether from an ini t ial length of 200 m to 20 km is explored
next Here the deployment length profile is critical to ensure the systems stability
It is not desirable to deploy the tether too quickly since that can render the tether
slack Hence the deployment strategy detailed in Section 216 is adopted The tether
is deployed over 35 orbits with the sinusoidal acceleration and deceleration occurring
over a 4 km length ( A ^ 2 ) with the remaining deployment at a constant velocity of
V0 = 146 ms
Initially the tether is stretched only slightly S = 1304 x 1 0 _ 3 0 0 m
with al l the other states of the system ie pitch roll and transverse displacements
remaining zero The response for the zero offset case is presented in Figure 4-7 For
the platform there is no longer any coupling with the tether hence it oscillates about
the equilibrium orientation determined by its inertia matrix However there is st i l l
coupling between the subsatellite and the tether since rcm^ mdash 100 m resulting
in complete domination of the subsatellite dynamics by the tether Consequently
the pitch motion ini t ial ly grows by almost 50deg but eventually subsides as the tether
deployment rate decreases It is interesting to note that the out-of-plane motion is not
affected by the deployment This is because the Coriolis force which is responsible
for the tether pitch motion does not have a component in the z direction (as it does
in the y direction) since it is always perpendicular to the orbital rotation (Q) and
the deployment rate (It) ie Q x It
Similarly there is an increase in the amplitude of the transverse elastic oscilshy
lations of the tether again due to the Coriolis effect (Figure 4-7b) Moreover as the
66
STSS configuration 3-Body a1(0)=a2(0)=cxt(0)=2deg P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m5y(0)=52(0)=1m dx(0)=dy(0)=d2(0)=1m Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
i i i i i i i i i i i 0 1 2 3 4 5 0 1 2 3 4 5
Time (Orbits) Time (Orbits) Tether Pitch Angle Tether Roll Angle
l i i i i -I h i i i i I i i i i l- i i i i l i i i i I i i- i i H 0 1 2 3 4 5 0 1 2 3 4 5
Time (Orbits) Time (Orbits) Satellite 2 Pitch Angle Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-6 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal local vertical and orbit normal (a) attitude response
67
STSS configuration 3-Body a1(0)=cc2(0)=a(0)=2deg p1(0)=p2(0)=(3t(0)=1deg 5x(0)=14m6y(0)=52(0)=1m dx(0)=dy(0)=d2(0)=1m Stationkeeping lt=20km
Tether Longitudinal Vibration 1 6 1 mdash i mdash i mdash i mdash mdash i mdash 1 mdash 1 mdash mdash mdash i mdash mdash 1 mdash 1 mdash 1 mdash i mdash 1 mdash 1 mdash 1 mdash 1 mdash r
0 1 2 3 4 5
Time (Orbits) Tether In-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-6 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal local vertical and orbit normal (b) vibration response
68
STSS configuration 3-Body a(0)=a2(0)=ot(0)=0 P1(0)=p2(0)=pt(0)=0 8X(0)=1304 x 103m 8y(0)=8z(0)=0 dx(0)=dy(0)=dz(0)=0 Deployment
l=02km to 20km in 35 Orbits
Tether Length Profile
E
Satellite 1 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 1 Roll Angle mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash [ mdash i mdash i mdash i mdash i mdash | mdash
1 2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 4 Time (Orbits)
Tether Roll Angle
CD
oo ax
i imdashimdashimdashImdashimdashimdashimdashr-
_ 1 _ L
1 2 3 4 Time (Orbits)
Satellite 2 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 2 Roll Angle
-50
(a) 1 2 3 4
Time (Orbits) 1 2 3 4
Time (Orbits)
Figure 4-7 Deployment dynamics of the three-body STSS configuration without offset (a) attitude response
69
STSS configuration 3-Body a1(0)=a2(0)=at(0)=0 p1(0)=p2(0)=(3t(0)=0 8X(0)=1304x 103m 8y(0)=52(0)=0 dx(0)=dy(0)=d2(0)=0 Deployment
l=02km to 20km in 35 Orbits
Tether Longitudinal Vibration
to
-02
(b)
Figure 4-7
2 3 Time (Orbits)
Tether In-Plane Transverse Vibration
0 1 1 bull 1 1
1 2 3 4 Time (Orbits)
Tether Out-of-Plane Transverse Vibration
5
1 1 1 1 1 i i | i i | ^
_ i i i i L -I
1 2 3 4 5 Time (Orbits)
Deployment dynamics of the three-body STSS configuration without offset (b) vibration response
70
tether elongates its longitudinal static equilibrium position also changes due to an
increase in the gravity-gradient tether tension It may be pointed out that as in the
case of attitude motion there are no out-of-plane tether vibrations induced
When the same deployment conditions are applied to the case of 1 m offset in
the x direction (Figure 4-8) coupling effects are re-introduced The platform pitch
exceeds -100deg during the peak deployment acceleration However as the tether pitch
motion subsides the platform pitch response returns to a smaller amplitude oscillation
about its nominal equilibrium On the other hand the platform roll motion grows
to over 20deg in amplitude in the terminal stages of deployment Longitudinal and
transverse vibrations of the tether remain virtually unchanged from the zero offset
case except for minute out-of-plane perturbations due to the platform librations in
roll
45 Tether Retrieval
The system response for the case of the tether retrieval from 20 km to 200 m
with an offset of 1 m in the x and z directions at the platform end is presented in
Figure 4-9 Here the Coriolis force induced during retrieval renders the tether unstable
(Figure 4-9a) Consequently the pitch motion of the platform is also destabilized
through coupling The z offset provides an additional moment arm that shifts the
roll equilibrium to 35deg however this degree of freedom is not destabilized From
Figure 4-9(b) it is apparent that the transverse mode of the tether vibration is also
disturbed producing a high frequency response in the final stages of retrieval This
disturbance is responsible for the slight increase in the roll motion for both the tether
and subsatellite
46 Five-Body Tethered System
A five-body chain link system is considered next To facilitate comparison with
71
STSS configuration 3-Body a(0)=a2(0)=at(0)=0 P1(0)=P2(0)=P(0)=0 5X(0)=1304 x 103m 8y(0)=8z(0)=0 dx(0)=1mdy(0)=d2(0)=0 Deployment
lt=02km to 20km in 35 Orbits
Tether Length Profile
E
Satellite 1 Pitch Angle
1 2 3 Time (Orbits)
Satellite 1 Roll Angle
bull bull bull i bull bull bull bull i
1 2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 Time (Orbits)
Tether Roll Angle
i i i i J I I i _
1 2 3 4 Time (Orbits)
Satellite 2 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 2 Roll Angle 3E-3h
D )
T J ^ 0 E 0 eg
CQ
-3E-3
(a) 1 2 3 4
Time (Orbits) 1 2 3 4
Time (Orbits)
Figure 4-8 Deployment dynamics of the three-body STSS configuration with offset along the local vertical (a) attitude response
7 2
STSS configuration 3-Body a1(0)=a2(0)=at(0)=0 p1(0)=P2(0)=(3t(0)=0 Sx(0)=1304 x 103m 5y(0)=5z(0)=0 dx(0)=1mdy(0)=dz(0)=0 Deployment
l=02km to 20km in 35 Orbits
Tether Longitudinal Vibration
Time (Orbits) Tether In-Plane Transverse Vibration
n 1 1 1 f
I I I _j I I I I I 1 1 1 1 1 2 3 4 5
Time (Orbits) Tether Out-of-Plane Transverse Vibration
J 1 I I 1 I I I I I I I I L _ _ I I d 1 2 3 4 5
Time (Orbits)
Deployment dynamics of the three-body STSS configuration with offset along the local vertical (b) vibration response
(b)
Figure 4-8
73
STSS configuration 3-Body a1(0)=a2(0)=at(0)=0 p(0)=P2(0)=(3t(0)=0 Sx(0)=14m8y(0)=82(0)=0 dy(0)=0dx(0)=d2(0)=1m Retrieval
l=20km to 02km in 35 Orbits
Satellite 1 Pitch Angle
(a) Time (Orbits)
Figure 4-9 Retrieval dynamics of the along the local vertical ar
Tether Length Profile
i i i i i i i i i i I 0 1 2 3
Time (Orbits) Satellite 1 Roll Angle
Time (Orbits) Tether Roll Angle
b_ i i I i I i i L _ J
0 1 2 3 Time (Orbits)
Satellite 2 Roll Angle _ mdash | r i i | 1 i i imdashmdashj i 1 r mdash i mdash
04- -
0 1 2 3 Time (Orbits)
three-body STSS configuration with offset d orbit normal (a) attitude response
74
STSS configuration 3-Body a1(0)=a2(0)=ocl(0)=0 (31(0)=f32(0)=pt(0)=0 8x(0)=14m8y(0)=6z(0)=0 dy(0)=0dx(0)=dz(0)=1m Retrieval
l=20km to 02km in 35 Orbits
Tether Longitudinal Vibration
E 10 to
Time (Orbits) Tether In-Plane Transverse Vibration
1 2 Time (Orbits)
Tether Out-of-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-9 Retrieval dynamics of the three-body STSS configuration with offset along the local vertical and orbit normal (b) vibration response
75
the three-body system dynamics studied earlier the chain is extended through the
addition of two bodies a tether and a subsatellite with the same physical properties
as before (Section 42) Thus the five-body system consists of a platform (satellite 1)
tether 1 subsatellite 1 (satellite 2) tether 2 and subsatellite 2 (satellite 3) as shown
in Figure 4-10 The offset of tether 1 at the platform-end is denoted by d2 while that
of tether 2 to subsatellite 1 as 04 In this numerical example fcm^ = fcm^ = 0 ie
tethers 1 and 2 are attached to the centre of mass of subsatellites 1 and 2 respectively
The system response for the case where the tether attachment points coincide
wi th the centres of mass of the rigid bodies ie zero offset (d2 = d mdash 0) is presented
in Figure 4-11 Here ogtt and at2 represent the pitch motion of tethers 1 and 2
respectively whereas 03 is the pitch motions of satellite 3 A similar convention
is adopted for the rol l angle 0 A s before the zero offset eliminates the coupling
between the tethers and satellites such that they are now free to librate about their
static equil ibrium positions However their is s t i l l mutual coupling between the two
tethers This coupling is present regardless of the offset position since each tether is
capable of transferring a force to the other The coupling is clearly evident in the pitch
response of the two tethers (Figure 4 - l l a ) O n the other hand the roll motion appears
uncoupled as the in i t ia l conditions in rol l for the two tethers are identical Thus the
motion is in phase and there is no transfer of energy through coupling A s expected
during the elastic response the tethers vibrate about their static equil ibrium positions
and mutually interact (Figure 4 - l l b ) However due to the variation of tension along
the tether (x direction) they do not have the same longitudinal equilibrium Note
relatively large amplitude transverse vibrations are present particularly for tether 1
suggesting strong coupling effects with the longitudinal oscillations
76
Figure 4-10 Schematic diagram of the five-body system used in the numerical example
77
STSS configuration 5-Body a1(0)=(x2(0)=at1(0)=2deg a3(0)=a12(0)=25deg p1(0)=r32(0)=p3(0)=pt1(0)=pt2(0)=r 82(0)=84(0)=505)05m d2(0)=d4(0)=000 Stationkeeping lt1=lu=10km Pitch Roll
Satellite 1 Pitch Angle Satellite 1 Roll Angle
o 1 2 Time (Orbits)
Tether 1 Pitch and Roll Angle
0 1 2 Time (Orbits)
Tether 2 Pitch and Roll Angle
Time (Orbits) Satellite 2 Pitch and Roll Angle
Time (Orbits) Satellite 3 Pitch and Roll Angle
Time (Orbits) Time (Orbits)
Figure 4-11 Stationkeeping dynamics of the five-body STSS configuration withshyout offset (a) attitude response
78
STSS configuration 5-Body a1(0)=a2(0)=oct1(0)=2deg a3(0)=at2(0)=25o
p1(0)=p2(0)=(33(0)=pt1(0)=pt2(0)=10
62(0)=54(0)=50505m d2(0)=d4(0)=000 Stationkeeping lt1=1 =1 Okm
Tether 1 Longitudinal Vibration Tether 2 Longitudinal Vibration i i 1 i 1 r 1 1 1 1
(b) Time (Orbits) Time (Orbits)
Figure 4-11 Stationkeeping dynamics of the five-body STSS configuration withshyout offset (b) vibration response
79
The system dynamics during deployment of both the tethers in the double-
pendulum configuration is presented in Figure 4-12 Deployment of each tether takes
place from 200 m to 20 km in 35 orbits The pitch motion of the system is similar
to that observed during the three-body case The Coriolis effect causes the tethers
to pitch through a large angle ( laquo 50deg ) which in turn disturbs the platform and
subsatellite due to the presence of offset along the local vertical On the other hand
the roll response for both the tethers is damped to zero as the tether deploys in
confirmation with the principle of conservation of angular momentum whereas the
platform response in roll increases to around 20deg Subsatellite 2 remains virtually
unaffected by the other links since the tether is attached to its centre of mass thus
eliminating coupling Finally the flexibility response of the two tethers presented in
Figure 4-12(b) shows similarity with the three-body case
47 BICEPS Configuration
The mission profile of the Bl-static Canadian Experiment on Plasmas in Space
(BICEPS) is simulated next It is presently under consideration by the Canadian
Space Agency To be launched by the Black Brant 2000500 rocket it would inshy
volve interesting maneuvers of the payload before it acquires the final operational
configuration As shown in Figure 1-3 at launch the payload (two satellites) with
an undeployed tether is spinning about the orbit normal (phase 1) The internal
dampers next change the motion to a flat spin (phase 2) which in turn provides
momentum for deployment (phase 3) When fully deployed the one kilometre long
tether will connect two identical satellites carrying instrumentation video cameras
and transmitters as payload
The system parameters for the BICEPS configuration are summarized below [59 0 0 1
bull I = I2 = 0 366 0 kg bull m 2 (satellite inertias) [ 0 0 392_
bull mi = 7772 = 200 kg (mass of the satellites)
80
STSS configuration 5-Body a(0)=cx2(0)=al1(0)=2deg a3(0)=c (0)=25o
P1(0)=P2(0)=p3(0)=f3t1(0)=Pt2(0)=r 52(0)=84(0)=1304x 10300m d2(0)=d4(0)=1G0m Deployment lt1=le=02km to 20km in 35 Orbits
Pitch Roll Satellite 1 Pitch Angle
Time (Orbits) Satellite 2 Pitch and Roll Angle T 1mdash1 1 1 I I I I | I I I 1 I | I I lt~r
Tether Length Profile i 1 1 1 1 i 1 1 1 1 i 1 1 1 1 i
i i i i i 0 1 2 3 4 5
Time (Orbits) Satellite 1 Roll Angle
0 1 2 3 4 5 Time (Orbits)
Tether 1 amp 2 Roll Angle L i | I I 1 I 1 1 -I
-Ih I- bull I I i i i i I i i i i 1
0 1 2 3 4 5 Time (Orbits)
Satellite 3 Pitch and Roll Angle i- i 1 1 i 1 1 1 1 i 1 1 1 1 i 1 1 1 1 i
L i i i i i i 1 i i mdash i mdash i I i imdashimdashLJ
0 1 2 3 4 5 Time (Orbits)
Figure 4-12 Deployment dynamics of the five-body S T S S configuration with offshyset along the local vertical (a) attitude response
81
STSS configuration 5-Body a1(0)=a2(0)=oct1(0)=2deg a3(0)=at2(0)=25deg P1(0)=P2(0)=P3(0)=Pt1(0)=pt2(0)=1o
82(0)=54(0)=1304 x 10-300m d2(0)=d4(0)=100m Deployment lt1=l2=02km to 20km in 35 Orbits
Tether 1 Longitudinal Vibration Tether 2 Longitudinal Vibration
Time (Orbits) Time (Orbits) Tether 1 Z Tranverse Vibration Tether 2 Z Tranverse Vibration
Figure 4-12 Deployment dynamics of the five-body STSS configuration with offshyset along the local vertical (b) vibration response
82
bull EtAt = 61645 N (tether stiffness)
bull pt = 30 k g k m (tether density)
bull It = 1 km (tether length)
bull rid = 1-0 (tether structural damping coefficient)
The system is in a circular orbit at a 289 km altitude Offset of the tether
attachment point to the satellites at both ends is taken to be 078 m in the x
direction The response of the system in the stationkeeping phase for a prescribed set
of in i t ia l conditions is given by Figure 4-13 As in the case of the STSS configuration
there is strong coupling between the tether and the satellites rigid body dynamics
causing the latter to follow the attitude of the tether However because of the smaller
inertias of the payloads there are noticeable high frequency modulations arising from
the tether flexibility Response of the tether in the elastic degrees of freedom is
presented in Figure 4-13(b) Note the longitudinal vibrations decay quite rapidly
(lt 05 orbits) due to the structural damping however it has vir tual ly no effect on
the transverse oscillations
The unique mission requirement of B I C E P S is the proposed use of its ini t ia l
angular momentum in the cartwheeling mode to aid in the deployment of the tethered
system The maneuver is considered next The response of the system during this
maneuver wi th an ini t ia l cartwheeling rate of 5 deg s is illustrated in Figure 4-14(a) The
in i t ia l tether length is taken to be 10 m and is deployed to 1 km There is an in i t ia l
increase in the pitch motion of the tether however due to the conservation of angular
momentum the cartwheeling rate decreases proportionally to the square of the change
in tether deployment rate The result is a rapid drop in the cartwheeling rate unti l
the system stops rotating entirely and simply oscillates about its new equilibrium
point Consequently through coupling the end-bodies follow the same trend The
roll response also subsides once the cartwheeling motion ceases However
83
BICEPS configuration 3-Body a1(0)=o2(0)=at(0)=2deg P1(0)=P2(0)=pt(0)=1deg 8X(0)=001 m 6y(0)=52(0)=01 m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping l=1 km
Satellite 1 Pitch Angle Satellite 1 Roll Angle - i 1 1 1 1 i - i q r 1 1 -gt
I- I i t i i i i I H 0 1 2 0 1 2
Time (Orbits) Time (Orbits) Tether Pitch Angle Tether Roll Angle
F mdash 1 mdash 1 mdash mdash 1 mdash i mdash 1 mdash 1 mdash mdash 1 mdash = i F mdash 1 mdash mdash mdash 1 mdash i mdash 1 mdash 1 mdash 1 mdash 1 mdash q
(a) Time (Orbits) Time (Orbits)
Figure 4-13 Stationkeeping dynamics of the three-body BICEPS configuration with offset along the local vertical (a) attitude response
84
BICEPS configuration 3-Body a1(0)=o2(0)=at(0)=2deg Pl(0)=p2(0)=(3t(0)=1deg 8X(0)=001 m 8y(0)=82(0)=01 m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping lt=1km
Tether Longitudinal Vibration 1E-2[
laquo = 5 E - 3 |
000 002
Time (Orbits) Tether In-Plane Transverse Vibration
Time (Orbits) Tether Out-of-Plane Transverse Vibration
Time (Orbits)
Figure 4-13 Stationkeeping dynamics of the three-body BICEPS configuration with offset along the local vertical (b) vibration response
85
BICEPS configuration 3-Body a1(0)=o2(0)=at(0)=2deg a1(0)=a2(0)=at(0)=57s Pl(0)=P2(0)=(3t(0)=1deg 8X(0)=115 x 103m 8y(0)=82(0)=0 dx(0)=078m dy(0)=d2(0)=0 Cartwheeling Deployment
lt=10m to 1km in 1 Orbit
Tether Length Profile
E
2000
cu T3
2000
co CD
2000
C D CD
T J
Satellite 1 Pitch Angle
1 2 Time (Orbits)
Satellite 1 Roll Angle
cn CD
33 cdl
Time (Orbits) Tether Pitch Angle
Time (Orbits) Tether Roll Angle
Time (Orbits) Satellite 2 Pitch Angle
1 2 Time (Orbits)
Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-14 Cartwheeling dynamics with deployment for the three-body BICEPS with offset along the local vertical (a) attitude response
86
BICEPS configuration 3-Body a1(0)=a2(0)=at(0)=2deg a (0)=a2(0)=a(0)=57s P1(0)=P2(0)=f3(0)=1deg 8X(0)=115x103m 8y(0)=8z(0)=0 dx(0)=078m dy(0)=d2(0)=0 Cartwheeling Deployment
l=1 Om to 1 km in 1 Orbit
5E-3h
co
0E0
025
000 E to
-025
002
-002
(b)
Figure 4-14
Tether Longitudinal Vibration
Time (Orbits) Tether In-Plane Transverse Vibration
Time (Orbits) Tether Out-of-Plane Transverse Vibration
Time (Orbits)
Cartwheeling dynamics with deployment for the three-body BICEPS with offset along the local vertical (b) vibration response
87
as the deployment maneuver is completed the satellites continue to exhibit a small
amplitude low frequency roll response with small period perturbations induced by
the tether elastic oscillations superposed on it (Figure 4-14b)
48 OEDIPUS Spinning Configuration
The mission entitled Observation of Electrified Distributions of Ionospheric
Plasmas - A Unique Strategy ( O E D I P U S ) also consists of a 1 k m tethered system with
two similar end-masses It was first flown in a suborbital flight in 1989 ( O E D I P U S A )
followed by a recent study in November 1995 ( O E D I P U S - C ) in which the University
of Br i t i sh Columbia was one of the participants Dynamically O E D I P U S represents a
unique system never encountered before Launched by a Black Brant rocket developed
at Br is to l Aerospace L t d the system ie the end-bodies together wi th the tether
spins about the longitudinal axis of the tether to achieve stabilized alignment wi th
Earth s magnetic field
The response of the system undergoing a spin rate of j = l deg s is presented
in Figure 4-15 using the same parameters as those outlined for the B I C E P S conshy
figuration in Section 47 Again there is strong coupling between the three bodies
(two satellites and tether) due to the nonzero offset The spin motion introduces an
additional frequency component in the tethers elastic response which is transferred
to the l ibrational motion of the satellites A s the spin rate increases to 1 0 deg s (Figure
4-16) the amplitude and frequency of the perturbations also increase however the
general character of the response remains essentially the same
88
OEDIPUS configuration 3-Body a1(0)=a2(0)=at(0)=2deg i(0)^[2(0H(0)=Ys p1(0)=p2(0)=p(0)=r 8X(0)=12 x 103m 8y(0)=82(0)=01m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping l=1 km
Satellite 1 Pitch Angle Satellite 1 Roll Angle mdash i 1 1 1 a r 1 1 1 r
(a) Time (Orbits) Time (Orbits)
Figure 4-15 Spin dynamics (7 = ldegs) of the three-body OEDIPUS configuration with offset along the local vertical (a) attitude response
89
OEDIPUS configuration 3-Body a1(0)=a2(0)=cct(0)=2deg Y1(0)=Y2(0)=Yt(0)=l7s (31(0)=p2(0)=pt(0)=1deg 8X(0)=12 x 103m 8y(0)=5z(0)=01 m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping lt=1 km
1E-2
to
OEO
E 00
to
Tether Longitudinal Vibration
005
Time (Orbits) Tether In-Plane Transverse Vibration
Time (Orbits) Tether Out-of-Plane Transverse Vibration
Time (Orbits)
Figure 4-15 Spin dynamics (7 = l deg s ) of the three-body O E D I P U S configuration with offset along the local vertical (b) vibration response
90
OEDIPUS configuration 3-Body a1(0)=a2(0)=a(0)=2deg Yi(0)=Y2(0)=rt(0)=107s 81(0)=p2(0)=pt(0)=1deg 8X(0)=12 x 103m 8y(0)=82(0)=01m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping lt=1km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-16 Spin dynamics (7= 10degs) of the three-body OEDIPUS configura tion with offset along the local vertical (a) attitude response
91
OEDIPUS configuration 3-Body a1(0)=a2(0)=ot(0)=2deg Yi(0)H(0)4(0)laquo107s P1(0)=P2(0)=pt(0)=1deg 8X(0)=12 x 103m 8y(0)=82(0)=01m dx(0)=078m dy(0)=dz(0)=0 Stationkeeping lt=1 km
Tether Longitudinal Vibration
Figure 4-16 Spin dynamics (7 = 10degs) of the three-body OEDIPUS configurashytion with offset along the local vertical (b) vibration response
92
5 ATTITUDE AND VIBRATION CONTROL
51 Att itude Control
511 Preliminary remarks
The instability in the pitch and roll motions during the retrieval of the tether
the large amplitude librations during its deployment and the marginal stability during
the stationkeeping phase suggest that some form of active control is necessary to
satisfy the mission requirements This section focuses on the design of an attitude
controller with the objective to regulate the librational dynamics of the system
As discussed in Chapter 1 a number of methods are available to accomplish
this objective This includes a wide variety of linear and nonlinear control strategies
which can be applied in conjuction with tension thrusters offset momentum-wheels
etc and their hybrid combinations One may consider a finite number of Linear T ime
Invariant (LTI) controllers scheduled discretely over different system configurations
ie gain scheduling[43] This would be relevant during deployment and retrieval of
the tether where the configuration is changing with time A n alternative may be
to use a Linear T ime Varying ( L T V ) model in conjuction with an adaptive control
scheme where on-line parametric identification may be used to advantage[56] The
options are vir tually limitless
O f course the choice of control algorithms is governed by several important
factors effective for time-varying configurations computationally efficient for realshy
time implementation and simple in character Here the thrustermomentum-wheel
system in conjuction with the Feedback Linearization Technique ( F L T ) is chosen as
it accounts for the complete nonlinear dynamics of the system and promises to have
good overall performance over a wide range of tether lengths It is well suited for
93
highly time-varying systems whose dynamics can be modelled accurately as in the
present case
The proposed control method utilizes the thrusters located on each rigid satelshy
lite excluding the first one (platform) to regulate the pitch and rol l motion of the
tether to which it is attached ie the thrusters located on satellite 2 (link 3) regshy
ulates the attitude motion of tether 1 (link 2) O n the other hand the l ibrational
motion of the rigid bodies (platform and subsatellite) is controlled using a set of three
momentum wheels placed mutually perpendicular to each other
The F L T method is based on the transformation of the nonlinear time-varying
governing equations into a L T I system using a nonlinear time-varying feedback Deshy
tailed mathematical background to the method and associated design procedure are
discussed by several authors[435758] The resulting L T I system can be regulated
using any of the numerous linear control algorithms available in the literature In the
present study a simple P D controller is adopted arid is found to have good perforshy
mance
The choice of an F L T control scheme satisfies one of the criteria mentioned
earlier namely valid over a wide range of tether lengths The question of compushy
tational efficiency of the method wi l l have to be addressed In order to implement
this controller in real-time the computation of the systems inverse dynamics must
be executed quickly Hence a simpler model that performs well is desirable Here
the model based on the rigid system with nonlinear equations of motion is chosen
Of course its validity is assessed using the original nonlinear system that accounts
for the tether flexibility
512 Controller design using Feedback Linearization Technique
The control model used is based on the rigid system with the governing equa-
94
tions of motion given by
Mrqr + fr = Quu (51)
where the left hand side represents inertia and other forces while the right hand side
is the generalized external force due to thrusters and momentum-wheels and is given
by Eq(2109)
Lett ing
~fr = fr~ QuU (52)
and substituting in Eq(296)
ir = s(frMrdc) (53)
Expanding Eq(53) it can be shown that
qr = S(fr Mrdc) - S(QU Mr6)u (5-4)
mdash = Fr - Qru
where S(QU Mr0) is the column matrix given by
S(QuMr6) = [s(Qu(l)Mr0) 5 (Q u ( 2 ) M r | 0 ) S(QU nu) M r | 0 ) ] (55)
and Qu-i) is the i ^ column of Qu Extract ing only the controlled equations from
Eq(54) ie the attitude equations one has
Ire = Frc ~~ Qrcu
(56)
mdash vrci
where vrc is the new control input required to regulate the decoupled linear system
A t this point a simple P D controller can be applied ie
Vrc = qrcd +KvQrcd-Qrc) + Kp(Qrcd~ Qrc)- (57)
Eq(57) represents the secondary controller with Eq(54) as the primary controller
Kp and Kv are the proportional and derivative gain matrices with qrcdi Qrcd
an(^ qrcd
95
as the desired acceleration velocity and position vector for the attitude angles of each
controlled body respectively Solving for u from Eq(56) leads to
u Qrc Frc ~ vrcj bull (58)
5 13 Simulation results
The F L T controller is implemented on the three-body STSS tethered system
The choice of proportional and derivative matrix gains is based on a desired settling
time of ts mdash 05 orbit and a damping factor of = 07 for both the pitch and roll
actuators in each body Given O and ts it can be shown that
-In ) 0 5 V T ^ r (59)
and hence kn bull mdash10
V (510)
where kp and kVj are the diagonal elements of the gain matrices Kp and Kv in
Eq(57) respectively[59]
Note as each body-fixed frame is referred directly to the inertial frame FQ
the nominal pitch equilibrium angle is zero only when measured from the L V L H frame
(OJJ mdash 0) However it is 6 when referred to FQ Hence the desired position vector
qrC(l is set equal to 0(t) ie the time-varying orbital angle for the pitch motion
and zero for the roll and yaw rotations such that
qrcd 37Vxl (511)
96
The desired velocity and acceleration are then given as
e U 3 N x l (512)
E K 3 7 V x l (513)
When the system is in a circular orbit qrc^ = 0
Figure 5-1 presents the controlled response of the STSS system defined in
Chapter 4 in the stationkeeping phase with offset d2 = 111T m and If = 20
km A s mentioned earlier the F L T controller is based on the nonlinear rigid model
Note the pitch and roll angles of the rigid bodies as well as of the tether are now
attenuated in less than 1 orbit (Figure 5-la) Consequently the longitudinal vibration
response of the tether is free of coupling arising from the librational modes of the
tether This leaves only the vibration modes which are eventually damped through
structural damping The pitch and roll dynamics of the platform require relatively
higher control moments Ma and Mpi respectively (Figure 5-lb) since they have to
overcome the extra moments generated by the offset of the tether attachment point
Furthermore the platform demands an additional moment to maintain its orientation
at zero pitch and roll angle since it is not the nominal equilibrium position The nonshy
zero static equilibrium of the platform arises due to its products of inertia On the
other hand the tether pitch and roll dynamics require only a small ini t ia l control
force of about plusmn 1 N which eventually diminishes to zero once the system is
97
and
o 0
Qrcj mdash
o o
0 0
Oiit) 0 0
1
N
N
1deg I o
STSS configuration 3-Body a1(0)=a2(0)=at(0)=2lt
P1(0)=p2(0)=pt(0)=1 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=dz(0)=1m Controlled Stationkeeping lt=20km
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
05 10 15 Time (Orbits)
Sat 2 Pitch and Roll Angles
20
CO CD
T J
cdT
cT
- mdash mdash 1 1 1 1 1 I - 1
Pitch -
A Roll -
A -
1 bull bull
-
1 Time (Orbits)
Tether Y Tranverse Vibration
1 2 Time (Orbits)
Tether Z Tranverse Vibration
gt to N
to
Time (Orbits) Time (Orbits) Longitudinal Vibration
145h
(a) Time (Orbits)
Figure 5-1 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local horizontal local vertical and orbit normal (a) attitude and vibration response
98
1
STSS configuration 3-Body a1(0)=ct2(0)=cct(0)=20
P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=d2(0)=1m Controlled Stationkeeping lt=20km
Sat 1 Pitch Control Moment Sat 1 Roll Control Moment
Figure 5-1 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local horizontal local vertical and orbit normal (b) control actuator time histories
99
stabilized Similarly for the end-body a very small control torque is required to
attenuate the pitch and roll response
If the complete nonlinear flexible dynamics model is used in the feedback loop
the response performance was found to be virtually identical with minor differences
as shown in Figure 5-2 For example now the pitch and roll response of the platform
is exactly damped to zero with no steady-state error as against the minor almost
imperceptible deviation for the case of the controller basedon the rigid model (Figure
5-la) In addition the rigid controller introduces additional damping in the transverse
mode of vibration where none is present when the full flexible controller is used
This is due to a high frequency component in the at motion that slowly decays the
transverse motion through coupling
As expected the full nonlinear flexible controller which now accounts the
elastic degrees of freedom in the model introduces larger fluctuations in the control
requirement for each actuator except at the subsatellite which is not coupled to the
tether since f c m 3 = 0 The high frequency variations in the pitch and roll control
moments at the platform are due to longitudinal oscillations of the tether and the
associated changes in the tension Despite neglecting the flexible terms the overall
controlled performance of the system remains quite good Hence the feedback of the
flexible motion is not considered in subsequent analysis
When the tether offset at the platform is restricted to only 1 m along the x
direction a similar response is obtained for the system However in this case the
steady-state error in the platforms rigid body motion is much smaller (Figure 5-3a)
In addition from Figure 5-3(b) the platforms control requirement is significantly
reduced
Figure 5-4 presents the controlled response of the STSS deploying a tether
100
STSS configuration 3-Body a1(0)=a2(0)=oct(0)=2deg p1(0)=P2(0)=p(0)=1deg 8x(0)=14m5y(0)=82(0)=1m dx(0)=dy(0)=dz(0)=1m Controlled Stationkeeping l=20km
Tether Pitch and Roll Angles
c n
T J
cdT
Sat 1 Pitch and Roll Angles 1 r
00 05 10 15 Time (Orbits)
Sat 2 Pitch and Roll Angles
c n CD
cdT a
o T J
CQ
- - I 1 1 1 I 1 1 1 1
P i t c h
R o l l
bull
1 -
0 1 2 Time (Orbits)
Tether Y Tranverse Vibration
0 1 2 Time (Orbits)
Tether Z Tranverse Vibration
E gt
10 N
CO
Time (Orbits) Time (Orbits) Longitudinal Vibration
145h
(a) Time (Orbits)
Figure 5-2 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear flexible FLT controller with offset along the local horizontal local vertical and orbit normal (a) attitude and vibration response
101
STSS configuration 3-Body a1(0)=a2(0)=cxt(0)=2deg 31(0)=p2(0)=p(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=d2(0)=1m Controlled Stationkeeping lt=20km
r i i i t i mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash mdash J 0 1 2 0 1 2
Time (Orbits) Time (Orbits) Tether Pitch Control Thrust Tether Roll Control Thruster
(b) Time (Orbits) Time (Orbits)
Figure 5-2 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear flexible FLT controller with offset along the local horizontal local vertical and orbit normal (b) control actuator time histories
102
STSS configuration 3-Body a1(0)=o2(0)=a(0)=2deg P1(0)=p2(0)=(3t(0)=1deg 8x(0)=14m5y(0)=6z(0)=1m dx(0)=1mdy(0)=dz(0)=0 Controlled Stationkeeping lt=20km
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
05 10 15 Time (Orbits)
Sat 2 Pitch and Roll Angles
CD
eg CQ
0 1 2 Time (Orbits)
Tether Y Tranverse Vibration
1 2 Time (Orbits)
Tether Z Tranverse Vibration
gt CO
Time (Orbits) Time (Orbits) Longitudinal Vibration
145h
(a) Time (Orbits)
Figure 5-3 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local vertical (a) attitude and vibration response
103
STSS configuration 3-Body a1(0)=o2(0)=al(0)=2o
P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=1mdy(0)=dz(0)=0 Controlled Stationkeeping lt=20km
Sat 1 Pitch Control Moment 11 1 1 r
Sat 1 Roll Control Moment
0 1 2 Time (Orbits)
Tether Pitch Control Thrust
0 1 2 Time (Orbits)
Sat 2 Pitch Control Moment
0 1 2 Time (Orbits)
Tether Roll Control Thruster
-07 o 1
Time (Orbits Sat 2 Roll Control Moment
1E-5
0E0F
Time (Orbits)
Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local vertical (b) control actuator time histories
104
STSS configuration 3-Body a1(0)=a2(0)=a(0)=2deg Pl(0)=p2(0)=pt(0)=1deg 6X(0)=1304 x 103m 8y(0)=82(0)=0 dx(0)=1mdy(0)=d2(0)=0 Controlled Deployment
l=02km to 20km in 35 Orbits
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
o 1 2 Time (Orbits)
Tether Y Tranverse Vibration
0 1 2 Time (Orbits)
Sat 2 Pitch and Roll Angles
CD CD
CM CO
Time (Orbits) Tether Z Tranverse Vibration
to
150
100
1 2 3 0 1 2
Time (Orbits) Time (Orbits) Longitudinal Vibration
to
(a)
50
2 3 Time (Orbits)
Figure 5 - 4 Deployment dynamics of the three-body STSS using the nonlinear rigid FLT controller with offset along the local vertical (a) attitude and vibration response
105
STSS configuration 3-Body a1(0)=oc2(0)=al(0)=20
P1(0)=p2(0)=pt(0)=1deg 8X(0)=1304 x 103m 5y(0)=52(0)=0 dx(0)=1mdy(0)=dz(0)=0 Controlled Deployment
l=02km to 20km in 35 Orbits
Tether Length Profile i 1
2 0 h
E 2L
Sat 1 Pitch Control Moment t mdash r mdash i mdash i mdash | mdash i mdash i mdash i mdash i mdash | mdash r mdash i mdash i mdash i mdash | mdash r mdash
0 5
1 2 3
Time (Orbits) Sat 1 Roll Control Moment
i i t | t mdash t mdash i mdash [ i i mdash r -
1 2
Time (Orbits) Tether Pitch Control Thrust
1 2 3 Time (Orbits)
Tether Roll Control Thruster
2 E - 5
1 2 3
Time (Orbits) Sat 2 Pitch Control Moment
1 2
Time (Orbits) Sat 2 Roll Control Moment
1 E - 5
E
O E O
Time (Orbits)
Deployment dynamics of the three-body STSS using the nonlinear rigid FLT controller with offset along the local vertical (b) control actuator time histories
106
from 02 km to 20 km in 35 orbits with a i m offset along the tether length A s before
the systems attitude motion is well regulated by the controller However the control
cost increases significantly for the pitch motion of the platform and tether due to the
Coriolis force induced by the deployment maneuver (Figure 5-4b) Initially there is a
sinusoidal increase in the pitch control requirement for both the tether and platform
as the tether accelerates to its constant velocity VQ Then the control requirement
remains constant for the platform at about 60 N m as opposed to the tether where the
thrust demand increases linearly Finally when the tether decelerates the actuators
control input reduces sinusoidally back to zero The rest of the control inputs remain
essentially the same as those in the stationkeeping case In fact the tether pitch
control moment is significantly less since the tether is short during the ini t ia l stages
of control However the inplane thruster requirement Tat acts as a disturbance and
causes 6y to nearly double from the uncontrolled value (Figure 4-8a) O n the other
hand the tether has no out-of-plane deflection
Final ly the effectiveness of the F L T controller during the crit ical maneuver of
retrieval from 20 km to 02 k m in 35 orbits is assessed in Figure 5-5 A s in the earlier
cases the system response in pitch and roll is acceptable however the controller is
unable to suppress the high frequency elastic oscillations induced in the tether by the
retrieval O f course this is expected as there is no active control of the elastic degrees
of freedom However the control of the tether vibrations is discussed in detail in
Section 52 The pitch control requirement follows a similar trend in magnitude as
in the case of deployment with only minor differences due to the addition of offset in
the out-of-plane direction ( d 2 = 1 0 1 ^ m) This is also responsible for the higher
roll moment requirement for the platform control (Figure 5-5b)
107
STSS configuration 3-Body a1(0)=X2(0)=a(0)=2deg P1(0)=p2(0)=pt(0)=1deg 5x(0)=14m8y(0)=52(0)=0 dx(0)=1mdy(0)=dz(0)=0 Controlled Retrieval
l=20km to 02km in 35 Orbits
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
0 1 2 Time (Orbits)
Sat 2 Pitch and Roll Angles
CO CD
cdT
CO CD
T J CM
CO
1 2 Time (Orbits)
Tether Y Tranverse Vibration
Time (Orbits) Tether Z Tranverse Vibration
150
100
1 2 3 0 1 2 Time (Orbits) Time (Orbits)
Longitudinal Vibration
CO
(a)
x 50
2 3 Time (Orbits)
Figure 5 -5 Retrieval dynamics of the three-body STSS using the non-linear rigid FLT controller with offset along the local vertical and orbit normal (a) attitude and vibration response
108
STSS configuration 3-Body a1(0)=ct2(0)=at(0)=2deg p1(0)=p2(0)=(3(0)=1deg 5x(0)=14m5y(0)=82(0)=0 dx(0)=1mdy(0)=dz(0)=0 Controlled Retrieval
l=20km to 02km in 35 Orbits
Tether Length Profile
E
1 2 Time (Orbits)
Tether Pitch Control Thrust
2E-5
1 2 Time (Orbits)
Sat 2 Pitch Control Moment
OEOft
1 2 3 Time (Orbits)
Sat 1 Roll Control Moment
1 2 3 Time (Orbits)
Tether Roll Control Thruster
1 2 Time (Orbits)
Sat 2 Roll Control Moment
1E-5 E z
2
0E0
(b) Time (Orbits) Time (Orbits)
Figure 5-5 Retrieval dynamics of the three-body STSS using the non-linear rigid FLT controller with offset along the local vertical and orbit normal (b) control actuator time histories
109
52 Control of Tethers Elastic Vibrations
521 Prel iminary remarks
The requirement of a precisely controlled micro-gravity environment as well
as the accurate release of satellites into their final orbit demands additional control
of the tethers vibratory motion in addition to its attitude regulation To that end
an active vibration suppression strategy is designed and implemented in this section
The strategy adopted is based on offset control ie time dependent variation
of the tethers attachment point at the platform (satellite 1) A l l the three degrees
of freedom of the offset motion are used to control both the longitudinal as well as
the inplane and out-of-plane transverse modes of vibration In practice the offset
controller can be implemented through the motion of a dedicated manipulator or a
robotic arm supporting the tether which in turn is supported by the platform The
focus here is on the control of elastic deformations during stationkeeping (ie fully
deployed tether fixed length) as it represents the phase when the mission objectives
are carried out
This section begins with the linearization of the systems equations of motion
for the reduced three body stationkeeping case This is followed by the design of
the optimal control algorithm Linear Quadratic Gaussian-Loop Transfer Recovery
( L Q G L T R ) based on the reduced model and its implementation on the full nonlinear
model Final ly the systems response in the presence of offset control is presented
which tends to substantiate its effectiveness
522 System linearization and state-space realization
The design of the offset controller begins with the linearization of the equations
of motion about the systems equilibrium position Linearization of extremely lengthy
(even in matrix form) highly nonlinear nonautonomous and coupled equations of
110
motion presents a challenging problem This is further complicated by the fact the
pitch angle ot is not referred to the L V L H frame Hence the equilibrium pitch angle
is not a constant but is equal to the instantaneous orbital angle 9(t) Two methods
are available to resolve this problem
One may use the non-stationary equations of motion in their present form and
derive a controller based on the Linear T ime Varying ( L T V ) system Alternatively
one can design a Linear T ime Invariant (LTI) controller based on a new set of reduced
governing equations representing the motion of the three body tethered system A l shy
though several studies pertaining to L T V systems have been reported[5660] the latter
approach is chosen because of its simplicity Furthermore the L T I controller design
is carried out completely off-line and thus the procedure is computationally more
efficient
The reduced model is derived using the Lagrangian approach with the genershy
alized coordinates given by
where lty and are the pitch and roll angles of link i relative to the L V L H frame
respectively 5X 8y and 5Z are the generalized coordinates associated with the flexible
tether deformations in the longitudinal inplane and out-of-plane transverse direcshy
tions respectively Only the first mode of vibration is considered in the analysis
The nonlinear nonautonomous and coupled equations of motion for the tethshy
ered system can now be written as
Qred = [ltXlPlltX2P2fixSy5za302gt] (514)
M redQred + fr mdash 0gt (515)
where Mred and frec[ are the reduced mass matrix and forcing term of the system
111
respectively They are functions of qred and qred in addition to the time varying offset
position d2 and its time derivatives d2 and d2 A detailed derivation of Eq(515)
is given in Appendix II A n additional consequence of referring the pitch motion to
a local frame is that the new reduced equations are now independent of the orbital
angle 9 and under the further assumption of the system negotiating a circular orbit
91 remains constant
The nonlinear reduced equations can now be linearized about their static equishy
l ibr ium position For all the generalized coordinates the equilibrium position is zero
wi th the exception of 5X which has a non-zero equilibrium 5XQ The offset position
d2) is linearized about its ini t ial position d2^ Linearizing Eq(515) and recasting
into matr ix form gives
Msqred + CsQred + KsQred + Md^2 + Cdd2 + Kdd2 + fs = 0 (516)
where Ms Cs Ks Md Cd Kd and s are constant matrices Mul t ip ly ing throughout
by Ms
1 gives
qr = -Ms 1Csqr - Ms
lKsqr
- Ms~lCdd22 - Ms-Kdd2 - Ms~Mdd2 - Mg~1 fs r-1 (517)
Defining ud = d2 and v = qTd^T Eq(517) can be rewritten as
Pu t t ing
v =
+
-M~lCs -M~Cd -1
0
-Mg~lMd
Id
0 v + -M~lKs -M~lKd
0 0
-M-lfs
0
MCv + MKv + Mlud +Fs
- 4 ) - ( gt the L T I equations of motion can be recast into the state-space form as
(518)
(519)
x =Ax + Bud + Fd (520)
112
where
A = MC MK Id12 o
24x24
and
Let
B = ^ U 5 R 2 4 X 3 -
Fd = I FJ ) e f t 2 4 1
mdash _ i _ -5
(521)
(522)
(523)
(524)
where x is the perturbation vector from the constant equilibrium state vector xeq
Substituting Eq(524) into Eq(520) gives the linear perturbation state equation
modelling the reduced tethered system as
bullJ x mdash x mdash Ax + Bud + (Fd + Axeq) (525)
It can be shown that for ideal control[60] Fd + Axeq = 0 thus giving the familiar
state-space equation
S=Ax + Bud (526)
The selection of the output vector completes the state space realization of
the system The output vector consists of the longitudinal deformation from the
equil ibrium position SXQ of the tether at xt = h the slope of the tether due to the
transverse deformation at xt = 0 and the offset position d2 from its in i t ia l position
d2Q Thus the output vector y is given by
K(o)sx
V
dx ~ dXQ dy ~ dyQ
d z - d z0 J
Jy
dx da XQ vo
d z - dZQ ) dy dyQ
(527)
113
where lt ( 0 ) = ddxi[$vxi)]Xi=o and lt ( 0 ) = ddxi[4gtw(xi)]Xi=0
523 Linear Quadratic Gaussian control with Loop Transfer Recovery
W i t h the linear state space model defined (Eqs526527) the design of the
controller can commence The algorithm chosen is the Linear Quadratic Gaussian
( L Q G ) estimator based optimal controller[6061] The L Q G is a widely used optimal
controller with its theoretical background well developped by many authors over the
last 25 years It involves of the design of the Kalman-Bucy Fi l ter ( K B F ) which
provides an estimate of the states x and a Linear Quadratic Regulator ( L Q R ) which
is separately designed assuming all the states x are known Both the L Q R and K B F
designs independently have good robustness properties ie retain good performance
when disturbances due to model uncertainty are included However the combined
L Q R and K B F designs ie the L Q G design may have poor stability margins in the
presence of model uncertainties This l imitat ion has led to the development of an L Q G
design procedure that improves the performance of the compensator by recovering
the full state feedback robustness properties at the plant input or output (Figure 5-6)
This procedure is known as the Linear Quadratic Guassian-Loop Transfer Recovery
( L Q G L T R ) control algorithm A detailed development of its theory is given in
Ref[61] and hence is not repeated here for conciseness
The design of the L Q G L T R controller involves the repeated solution of comshy
plicated matrix equations too tedious to be executed by hand Fortunenately the enshy
tire algorithm is available in the Robust Control Toolbox[62] of the popular software
package M A T L A B The input matrices required for the function are the following
(i) state space matrices ABC and D (D = 0 for this system)
(ii) state and measurement noise covariance matrices E and 0 respectively
(iii) state and input weighting matrix Q and R respectively
114
u
MODEL UNCERTAINTY
SYSTEM PLANT
y
u LQR CONTROLLER
X KBF ESTIMATOR
u
y
LQG COMPENSATOR
PLANT
X = f v ( X t )
COMPENSATOR
x = A k x - B k y
ud
= C k f
y
Figure 5-6 Block diagram for the LQGLTR estimator based compensator
115
A s mentionned earlier the main objective of this offset controller is to regu-mdash
late the tether vibration described by the 5 equations However from the response
of the uncontrolled system presented in Chapter 4 it is clear that there is a large
difference between the magnitude of the librational and vibrational frequencies This
separation of frequencies allow for the separate design of the vibration and attitude mdash mdash
controllers Thus only the flexible subsystem composed of the S and d equations is
required in the offset controller design Similarly there is also a wide separation of
frequencies between the longitudinal and transverse modes of vibrations permitt ing mdash
the decoupling of the flexible subsystem into a set of longitudinal (Sx and dx) and
transverse (5y Sz dy and dz) subsystems The appended offset system d must also
be included since it acts as the control actuator However it is important to note
that the inplane and out-of-plane transverse modes can not be decoupled because
their oscillation frequencies are of the same order The two flexible subsystems are
summarized below
(i) Longitudinal Vibra t ion Subsystem
This is defined by u mdash Auxu + Buudu
(528)
where xu = SxdX25xdX2T u d u = dX2 and from Eq(527)
0 0 $u(lt) 0 0 0 0 1
0 0 1 0 0 0 0 1
(529)
Au and Bu are the rows and columns of A and B respectively and correspond to the
components of xu
The L Q R state weighting matrix Qu is taken as
Qu =
bull1 0 0 o -0 1 0 0 0 0 1 0
0 0 0 10
(530)
116
and the input weighting matrix Ru = 1^ The state noise covariance matr ix is
selected as 1 0 0 0 0 1 0 0 0 0 4 0
LO 0 0 1
while the measurement noise covariance matrix is taken as
(531)
i o 0 15
(532)
Given the above mentionned matrices the design of the L Q G L T R compenshy
sator is computed using M A T L A B with the following 2 commands
(i) kfu = l q r c ( ^ ( i C bdquo d i a g m x ( E u e u ) )
(ii) [AkugtBkugtCkugtDku =^ryAuBuCuDukfuQuRur)
where r is a scaler
The first command is used to compute the Ka lman filter gain matrix kfu Once
kfu is known the second command is invoked returning the state space representation
of the L Q G L T R dynamic compensator to
(533) xu mdash Akuxu Bkuyu
where xu is the state estimate vector oixu The function l t ry performs loop transfer
recovery at the system output ie the return ratio at the output approaches that
of the K B F loop given by Cu(sld - Au)Kfu[61 This is achieved by choosing a
sufficently large scaler value r in l t ry such that the singular values of the return
ratio approach those of the target design For the longitudinal controller design
r mdash 5 x 10 5 Figure 5-7(a) compares the singular values of the recovered compensator
design and the non-recovered L Q G compensator design with respect to the target
Unfortunetaly perfect recovery is not possible especially at higher frequencies
117
100
co
gt CO
-50
-100 10
Longitudinal Compensator Singular Values i I I I N I I mdash i mdash i i i i n i i mdash i mdash i i i i inimdash imdashi i i i i i i i mdash i mdash i i M i I I
Target r = 5x10 5
r = 0 (LQG)
i i i i I I i ii i i i 1 1 1 1 i i i i i 1 1 1 1 ii i I I i i i i 11111 -3 10 -2
(a)
101 10deg Frequency (rads)
101 10
Transverse Compensator Singular Values
100
50
X3
gt 0
CO
-50
bull100 10
-imdashi i 111II i 111 ni 1mdashi I I I I I I I 1mdashi I I I N I I 1mdashi i M 111
Target r = 50 r = 0 (LQG)
m i i i i i i n i i i i I I I I I I I 1mdashi I I I I I I I 1mdashi i 111 n -3 10 -2
(b)
101 10deg Frequency (rads)
101 10
Figure 5-7 Singular values for the LQG and LQGLTR compensator compared to target return ratio (a) longitudinal design (b) transverse design
118
because the system is non-minimal ie it has transmission zeros with positive real
parts[63]
(ii) Transverse Vibra t ion Subsystem
Here Xy mdash AyXy + ByU(lv
Dv = Cvxv
with xv = 6y 8Z dV2 dZ2 Sy 6Z dV2 dZ2T] udv = dy2dZ2T and
(534)
Cy
0 0 0 0 (0) 0 0 0 0 0 0 0 0 0 0 0 0
0 0 0 0
0 0 0 0 0 0 0 0 0 0 0 0
0 0 0
h 0 0 0
0 0 0 lt ( 0 ) o 0
0 1 0 0 0 1
0 0 0
2 r
h 0 0
0 0 1 0 0 1
(535)
Av and Bv are the rows and columns of the A and B corresponding to the components
O f Xy
The state L Q R weighting matrix Qv is given as
Qv
10 0 0 0 0 0 0 0 0 10 0 0 0 0 0 0 0 0 10 4 0 0 0 0 0 0 0 0 10 4 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1
(536)
mdash
Ry =w The state noise covariance matrix is taken
1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 9 x 10 4 0
0 0 0 0 0 0 0 9 x 10 4
(537)
119
while the measurement noise covariance matrix is
1 0 0 0 0 1 0 0 0 0 5 x 10 4 0
LO 0 0 5 x 10 4
(538)
As in the development of the longitudinal controller the transverse offset conshy
troller can be designed using the same two commands (lqrc and l try) wi th the input
matrices corresponding to the transverse subsystem with r = 50 The result is the
transverse compensator system given by
(539)
where xv is the state estimates of xv The singular values of the target transfer
function is compared with those achieved by the L Q G and L Q G L T R in Figure 5-
7(b) It is apparent that the recovery is not as good as that for the longitudinal case
again due to the non-minimal system Moreover the low value of r indicates that
only a lit t le recovery in the transverse subsystem is possible This suggests that the
robustness of the L Q G design is almost the maximum that can be achieved under
these conditions
Denoting f y = f j f j r xf = x^x^T and ud = udu ^ T the longishy
tudinal compensator can be combined to give
xf =
$d =
Aku
0
0 Akv
Cku o
Cu 0
Bku
0 B Xf
Xf = Ckxf
0
kv
Hf = Akxf - Bkyf
(540)
Vu Vv 0 Cv
Xf = CfXf
Defining a permutation matrix Pf such that Xj = PfX where X is the state vector
of the original nonlinear system given in Eq(32) the compensator and full nonlinear
120
system equations can be combined as
X
(541)
Ckxf
A block diagram representation of Eq(541) is shown in Figure 5-6
524 Simulation results
The dynamical response for the stationkeeping STSS with a tether length of 20
km and ini t ia l offset position of 1 m in each direction is presented in Figure 5-8 when
both the F L T attitude controller and the L Q G L T R offset controller are activated As
expected the offset controller is successful in quickly damping the elastic vibrations
in the longitudinal and transverse direction (Figure 5-8b) However from Figure 5-
8(a) it is clear that the presence of offset control requires a larger control moment
to regulate the attitude of the platform This is due to the additional torque created
by the larger moment arm around 2 m and 125 m in the inplane and out-of-plane
directions respectively introduced by the offset control In addition the control
moment for the platform is modulated by the tethers transverse vibration through
offset coupling However this does not significantly affect the l ibrational motion of
the tether whose thruster force remains relatively unchanged from the uncontrolled
vibration case
Final ly it can be concluded that the tether elastic vibration suppression
though offset control in conjunction with thruster and momentum-wheel attitude
control presents a viable strategy for regulating the dynamics tethered satellite sysshy
tems
121
STSS configuration 3-Body a1(0)=a2(0)=at(0)=2deg P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m6y(0)=52(0)=1m dx(0)=dy(0)=dz(0)=1m Stationkeeping I = 20km LQGLTR Vibration Control
Sat 1 Pitch and Roll Angles Tether Pitch and Roll Angles mdash 1 1 1 1 1 1 I 1 1 lt 1 ~ ^ 1 ^ l _
Time (Orbits) Time (Orbits) Sat 1 Roll Control Moment Tether Roll Control Thrust
(a) Time (Orbits) Time (Orbits)
Figure 5-8 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT attitude controller and LQGLTR offset vibration controller (a) attitude and libration controller reshysponse
122
STSS configuration 3-Body a1(0)=a2(0)=a(0)=2deg p1(0)=(32(0)=Q(0)=1deg ox(0)=14mSy(0)=5z(0)=1m dx(0)=dy(0)=dz(0)=1m Stationkeeping lt = 20km LQGLTR Vibration Control
Tether X Offset Position
Time (Orbits) Tether Y Tranverse Vibration 20r
Time (Orbits) Tether Y Offset Position
t o
1 2 Time (Orbits)
Tether Z Tranverse Vibration Time (Orbits)
Tether Z Offset Position
(b)
Figure 5-8
Time (Orbits) Time (Orbits)
Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT attitude controller and LQGLTR offset vibration controller (b) vibration and offset response
123
6 C O N C L U D I N G R E M A R K S
61 Summary of Results
The thesis has developed a rather general dynamics formulation for a multi-
body tethered system undergoing three-dimensional motion The system is composed
of multiple rigid bodies connected in a chain configuration by long flexible tethers
The tethers which are free to undergo libration as well as elastic vibrations in three
dimensions are also capable of deployment retrieval and constant length stationkeepshy
ing modes of operation Two types of actuators are located on the rigid satellites
thrusters and momentum-wheels
The governing equations of motion are developed using a new Order(N) apshy
proach that factorizes the mass matrix of the system such that it can be inverted
efficiently The derivation of the differential equations is generalized to account for
an arbitrary number of rigid bodies The equations were then coded in FORTRAN
for their numerical integration with the aid of a symbolic manipulation package that
algebraically evaluated the integrals involving the modes shapes functions used to dis-
cretize the flexible tether motion The simulation program was then used to assess the
uncontrolled dynamical behaviour of the system under the influence of several system
parameters including offset at the tether attachment point stationkeeping deployshy
ment and retrieval of the tether The study covered the three-body and five-body
geometries recently flown OEDIPUS system and the proposed BICEPS configurashy
tion It represents innovation at every phase a general three dimensional formulation
for multibody tethered systems an order-N algorithm for efficient computation and
application to systems of contemporary interest
Two types of controllers one for the attitude motion and the other for the
124
flexible vibratory motion are developed using the thrusters and momentum-wheels for
the former and the variable offset position for the latter These controllers are used to
regulate the motion of the system under various disturbances The attitude controller
is developed using the nonlinear Feedback Linearization Technique ( F L T ) and is based
on the nonlinear but rigid model of the system O n the other hand the separation
of the longitudinal and transverse frequencies from those of the attitude response
allows for the development of a linear optimal offset controller using the robust Linear
Quadratic Gaussian-Loop Transfer Recovery ( L Q G L T R ) method The effectiveness
of the two controllers is assessed through their subsequent application to the original
nonlinear flexible model
More important original contributions of the thesis which have not been reshy
ported in the literature include the following
(i) the model accounts for the motion of a multibody chain-type system undershy
going librational and in the case of tethers elastic vibrational motion in all
the three directions
(ii) the dynamics formulation is based on a recursive O(N) Lagrangian algorithm
that efficiently computes the systems generalized acceleration vector
(iii) the development of an attitude controller based on the Feedback Linearizashy
tion Technique for a multibody system using thrusters and momentum-wheels
located on each rigid body
(iv) the design of a three degree of freedom offset controller to regulate elastic v i shy
brations of the tether using the Linear Quadratic Gaussian and Loop Transfer
Recovery method ( L Q G L T R )
(v) substantiation of the formulation and control strategies through the applicashy
tion to a wide variety of systems thus demonstrating its versatility
125
The emphasis throughout has been on the development of a methodology to
study a large class of tethered systems efficiently It was not intended to compile
an extensive amount of data concerning the dynamical behaviour through a planned
variation of system parameters O f course a designer can easily employ the user-
friendly program to acquire such information Rather the objective was to establish
trends based on the parameters which are likely to have more significant effect on
the system dynamics both uncontrolled as well as controlled Based on the results
obtained the following general remarks can be made
(i) The presence of the platform products of inertia modulates the attitude reshy
sponse of the platform and gives rise to non-zero equilibrium pitch and roll
angles
(ii) Offset of the tether attachment point significantly affects the equil ibrium orishy
entation of the system as well as its dynamics through coupling W i t h a relshy
atively small subsatellite the tether dynamics dominate the system response
with high frequency elastic vibrations modulating the librational motion On
the other hand the effect of the platform dynamics on the tether response is
negligible As can be expected the platform dynamics is significantly affected
by the offset along the local horizontal and the orbit normal
(iii) For a three-body system deployment of the tether can destabilize the platform
in pitch as in the case of nonzero offset However the roll motion remains
undisturbed by the Coriolis force Moreover deployment can also render the
tether slack if it proceeds beyond a critical speed
(iv) Uncontrolled retrieval of the tether is always unstable in pitch It also leads
to high frequency elastic vibrations in the tether
(v) The five-body tethered system exhibits dynamical characteristics similar to
those observed for the three-body case As can be expected now there are
additional coupling effects due to two extra bodies a rigid subsatellite and a
126
flexible tether
(vi) Cartwheeling motion of the B I C E P S configuration can also be used to adshy
vantage in the deployment of the tether However exceeding a crit ical ini t ia l
cartwheeling rate can result in a large tension in the tether which eventually
causes the satellite to bounce back rendering the tether slack
(vii) Spinning the tether about its nominal length as in the case of O E D I P U S
introduces high frequency transverse vibrations in the tether which in turn
affect the dynamics of the satellites
(viii) The F L T based controller is quite successful in regulating the attitude moshy
tion of both the tether and rigid satellites in a short period of time during
stationkeeping deployment as well as retrieval phases
(ix) The offset controller is successful in suppressing the elastic vibrations of the
tether wi thin the allowable l imits of the offset mechanism ( plusmn 2 0 m)
62 Recommendations for Future Study
A s can be expected any scientific inquiry is merely a prelude to further efforts
and insight Several possible avenues exist for further investigation in this field A
few related to the study presented here are indicated below
(i) inclusion of environmental effects particularly the free molecular reaction
forces as well as interaction with Earths magnetic field and solar radiation
(ii) addition of flexible booms attached to the rigid satellites providing a mechashy
nism for energy dissipation
(iii) validation of the new three dimensional offset control strategies using ground
based experiments
(iv) animation of the simulation results to allow visual interpretation of the system
dynamics
127
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[30] Pradhan S M o d i V J and Misra A K A n Order N Dynamics Formulation of Flexible Mul t ibody Systems in Tree Topology - The Lagrangian Approach AIAAAAS Astrodynamics Specialist Conference San Diego C A U S A July 1996 Paper No 96-3624 pp 480-490
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[34] Pradhan S M o d i V J and Misra A K On the Inverse Control of the Tethered Satellite Systems The Journal of the Astronautical Sciences A A S Publications Office San Diego California U S A V o l 43 No 2 Apr i l - June 1995 pp 179-193
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[38] Ruying F and Bainum P M Dynamics and Control of a Space Platform with a Tethered Subsatellite Journal of Guidance Control and Dynamics V o l 11 No 4 July-August 1988 pp 377-381
[39] M o d i V J Lakshmanan P K and Misra A K Offset Control of Tethered Satellite Systems Analysis and Experimental Verification Acta Astronautica V o l 21 No 5 1990 pp 283-294
[40] X u D M Misra A K and M o d i V J Thruster Augmented Act ive Cin t ro l of a Tethered Satellite System During its Retrieval Journal of Guidance Control and Dynamics V o l 9 No 6 November-December 1986 pp 663-672
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[43] Pradhan S Planar Dynamics and Control of Tethered Satellite Systems P h D Thesis Department of Mechanical Engineering The University of Br i t i sh Columbia December 1994
[44] Bainum P M and Kumar V K Optimal-Control of the Shuttle-Tethered System Acta Astronautica V o l 7 No 12 1980 pp 1333-1348
[46] Hughes P C Spacecraft Att i tude Dynamics John Wi ley amp Sons New York U S A 1992 Chapter 2 pp 7-38
[47] Stengel R F Optimal Control and Estimation Dover Publications Inc New York U S A 1994 p 26
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[50] Meirovitch L Elements of Vibration Analysis M c G r a w - H i l l Inc New York U S A 1986 pp 255-256
[51] Greenwood D T Principles of Dynamics Prentice Ha l l Inc Englewood Cliffs New Jersey U S A 1965 pp 252-275
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131
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132
A P P E N D I X I TENSOR R E P R E S E N T A T I O N OF T H E
EQUATIONS OF M O T I O N
11 Prel iminary Remarks
The derivation of the equations of motion for the multibody tethered system
involves the product of several matrices as well as the derivative of matrices and
vectors with respect to other vectors In order to concisely code these relationships
in FORTRAN the matrix equations of motion are expressed in tensor notation
Tensor mathematics in general is an extremely powerful tool which can be
used to solve many complex problems in physics and engineering[6465] However
only a few preliminary results of Cartesian tensor analysis are required here They
are summarized below
12 Mathematical Background
The representation of matrices and vectors in tensor notation simply involves
the use of indices referring to the specific elements of the matrix entity For example
v = vk (k = lN) (Il)
A = Aij (i = lNj = lM)
where vk is the kth element of vector v and Aj is the element on the t h row and
3th column of matrix A It is clear from Eq(Il) that exactly one index is required to
completely define an entire vector whereas two independent indices are required for
matrices For this reason vectors and matrices are known as first-order and second-
order tensors respectively Scaler variables are known as zeroth-order tensors since
in this case an index is not required
133
Matrix operations are expressed in tensor notation similar to the way they are
programmed using a computer language such as FORTRAN or C They are expressed
in a summation notation known as Einstein notation For example the product of a
matrix with a vector is given as
w = Av (12)
or in tensor notation W i = ^2^2Aij(vkSjk)
4-1 ltL3gt j
where 8jk is the Kronecker delta Dropping the summation symbol the matrix-vector
product can be expressed compactly as
Wi = AijVj = Aimvm (14)
Here j is a dummy index since it appears twice in a single term and hence can
be replaced by any other index Note that since the resulting product is a vector
only one index is required namely i that appears exactly once on both sides of the
equation Similarly
E = aAB + bCD (15)
or Ej mdash aAimBmj + bCimDmj
~ aBmjAim I bDmjCim (16)
where A B C D and E are second-order tensors (matrices) and a and b are zeroth-
order tensors (scalers) In addition once an expression is in tensor form it can be
treated similar to a scaler and the terms can be rearranged
The transpose of a matrix is also easily described using tensors One simply
switches the position of the indices For example let w = ATv then in tensor-
notation
Wi = AjiVj (17)
134
The real power of tensor notation is in its unambiguous expression of terms
containing partial derivatives of scalers vectors or matrices with respect to other
vectors For example the Jacobian matrix of y = f(x) can be easily expressed in
tensor notation as
di-dx-j-1 ( L 8 )
mdash which is a second-order tensor as expected If f(x) = Ax where A is constant then
f - a )
according to the current definition If C = A(x)B(x) then the time derivative of C
is given by
C = Ax)Bx) + A(x)Bx) (110)
or in tensor notation
Cij = AimBmj + AimBmj (111)
= XhiAimfcBmj + AimBmj^k)
which is more clear than its equivalent vector form Note that A^^ and Bmjk are
third-order tensors that can be readily handled in F O R T R A N
13 Forcing Function
The forcing function of the equations of motion as given by Eq(260) is
- u 1-SdM- dqtdPe 9 f i m F(q qt) = Mq - -q mdashq + mdashmdash - Qd (260)
This expression can be converted into tensor form to facilitate its implementation
into F O R T R A N source code
Consider the first term Mq From Eq(289) the mass matrix M is given by
T M mdash Rv MtRv which in tensor form is
Mtj = RZiM^R^j (112)
135
Taking the time derivative of M gives
Mij = qsRv
ni^MtnrnRv
mj + Rv
niMtnTn^Rv
mj + Rv
niMtnmRv
mjs) (113)
The second term on the right hand side in Eq(260) is given by
-JPdMu 1 bdquo T x
2Q ~cWq = 2Qs s r k Q r ^ ^
Expanding Msrk leads to
M src = RnskMtnmRmr + RnsMtnmkRmr + RnsMtnmRmrk- (L15)
Finally the potential energy term in Eq(260) is also expanded into tensor
form to give dPe = dqt dPe
dq dq dqt = dq^QP^
and since ^ = Rp(q)q then
Now inserting Eqs(I13-I17) into Eq(260) and rearranging the tensor form of the
forcing term can how be stated as
Fk(q^q^)=(RP
sk + R P n J q n ) ^ - - Q d k
+ QsQr | (^Rlks ~ 2 M t n m R m r
+ (^RnkMtnms ~ 2~RnsMtnmgtk^ R m r
(^RnkRmrs ~ ~2RnsRmrk^j
where Fk(q q t) represents the kth component of F
(118)
136
A P P E N D I X II R E D U C E D EQUATIONS OF M O T I O N
II 1 Prel iminary Remarks
The reduced model used for the design of the vibration controller is represented
by two rigid end-bodies capable of three-dimensional attitude motion interconnected
with a fixed length tether The flexible tether is discretized using the assumed-
mode method with only the fundamental mode of vibration considered to represent
longitudinal as well as in-plane and out-of-plane transverse deformations In this
model tether structural damping is neglected Finally the system is restricted to a
nominal circular orbit
II 2 Derivation of the Lagrangian Equations of Mot ion
Derivation of the equations of motion for the reduced system follows a similar
procedure to that for the full case presented in Chapter 2 However in this model the
straightforward derivation of the energy expressions for the entire system is undershy
taken ignoring the Order(N) approach discussed previously Furthermore in order
to make the final governing equations stationary the attitude motion is now referred
to the local LVLH frame This is accomplished by simply defining the following new
attitude angles
ci = on + 6
Pi = Pi (HI)
7t = 0
where and are the pitch and roll angles respectively Note that spin is not
considered in the reduced model hence 7J = 0 Substituting Eq(IIl) into Eq(214)
137
gives the new expression for the rotation matrix as
T- CpiSa^ Cai+61 S 0 S a + e i
-S 0 c Pi
(II2)
W i t h the new rotation matrix defined the inertial velocity of the elemental
mass drrii o n the ^ link can now be expressed using Eq(215) as
(215)
Since the tether length is assumed to be fixed T j f j and Tj$jlt^ of Eq(215) are
identically zero Also defining
(II3) Vi = I Pi i
Rdm- f deg r the reduced system is given by
where
and
Rdm =Di + Pi(gi)m + T^iSi
Pi(9i)m= [Tai9i T0gi T^] fj
(II4)
(II5)
(II6)
D1 = D D l D S v
52 = 31 + P(d2)m +T[d2
D3 = D2 + l2 + dX2Pii)m + T2i6x
Note that $ j for i = 2 is defined in Section 212 whereas for i = 1 and 3 it is
the null matrix Since only one mode is used for each flexible degree of freedom
5i = [6Xi6yi6Zi]T for i = 2
The kinetic energy of the system can now be written as
i=l Z i = l Jmi RdmRdmdmi (UJ)
138
Expanding Kampi using Eq(II4)
KH = miDD + 2DiP
i j gidmi)fji + 2D T ltMtrade^
+tf J P^(9i)Pi9i)d^l + 2mT j PFmTi^dm^ (II8)
where the integrals are evaluated using the procedure similar to that described in
Chapter 2 The gravitational and strain energy expressions are given by Eq(247)
and Eq(251) respectively using the newly defined rotation matrix T in place of
TV
Substituting the kinetic and potential energy expressions in
d ( 8Ke d(Ke - Pe) i bull i Q ^ 0 (II9)
d t dqred d(ired
with
Qred = [ algt l gt a 2 2 A 2 A 2 lt ^ 2 Q 3 3 ] (5-14)
and d2 regarded as a time-varying parameter the nonlinear non-autonomous equashy
tions of motion for the reduced system can finally be expressed as
MredQred + fred = reg- (5-15)
139
331 Energy conservation 4 6
332 Comparison with available data 50
4 D Y N A M I C S I M U L A T I O N 53
41 Prel iminary Remarks 53
42 Parameter and Response Variable Definitions 53
43 Stationkeeping Profile 56
44 Tether Deployment 66
45 Tether Retrieval 71
46 Five-Body Tethered System 71
47 B I C E P S Configuration 80
48 O E D I P U S Spinning Configuration 88
5 A T T I T U D E A N D V I B R A T I O N C O N T R O L 93
51 Att i tude Control 93
511 Prel iminary remarks 93
512 Controller design using Feedback Linearization Technique 94
513 Simulation results 96
52 Control of Tethers Elastic Vibrations 110
521 Prel iminary remarks 110
522 System linearization and state-space realization 1 1 0
523 Linear Quadratic Gaussian control
with Loop Transfer Recovery 114
524 Simulation results 121
v
6 C O N C L U D I N G R E M A R K S 1 2 4
61 Summary of Results 124
62 Recommendations for Future Study 127
B I B L I O G R A P H Y 128
A P P E N D I C E S
I T E N S O R R E P R E S E N T A T I O N O F T H E E Q U A T I O N S O F M O T I O N 133
11 Prel iminary Remarks 133
12 Mathematical Background 133
13 Forcing Function 135
II R E D U C E D E Q U A T I O N S O F M O T I O N 137
II 1 Preliminary Remarks 137
II2 Derivation of the Lagrangian Equations of Mot ion 137
v i
LIST OF SYMBOLS
A j intermediate length over which the deploymentretrieval profile
for the itfl l ink is sinusoidal
A Lagrange multipliers
EUQU longitudinal state and measurement noise covariance matrices
respectively
matrix containing mode shape functions of the t h flexible link
pitch roll and yaw angles of the i 1 link
time-varying modal coordinate for the i 1 flexible link
link strain and stress respectively
damping factor of the t h attitude actuator
set of attitude angles (fjj = ci
structural damping coefficient for the t h l ink EjJEj
true anomaly
Earths gravitational constant GMejBe
density of the ith link
fundamental frequency of the link longitudinal tether vibrashy
tion
A tethers cross-sectional area
AfBfCjDf state-space representation of flexible subsystem
Dj inertial position vector of frame Fj
Dj magnitude of Dj mdash mdash
Df) transformation matrix relating Dj and Ds
D r i inplane radial distance of the first l ink
DSi Drv6i DZ1 T for i = 1 and DX Dy poundgt 2 T for 1
m
Pi
w 0 i
v i i
Dx- transformation matrix relating D j and Ds-
DXi Dyi Dz- Cartesian components of D
DZl out-of-plane position component of first link
E Youngs elastic modulus for the ith l ink
Ej^ contribution from structural damping to the augmented complex
modulus E
F systems conservative force vector
FQ inertial reference frame
Fi t h link body-fixed reference frame
n x n identity matrix
alternate form of inertia matrix for the i 1 l ink
K A Z $i(k + dXi+1)
Kei t h link kinetic energy
M(q t) systems coupled mass matrix
Mma Mma Mmy- control moments in the pitch roll and yaw directions respec-
tively for the t h link
Mr fr rigid mass matrix and force vector respectively
Mred fred- Qred mass matrix force vector and generalized coordinate vector for
the reduced model respectively
Mt block diagonal decoupled mass matrix
symmetric decoupled mass matrix
N total number of links
0N) order-N
Pi(9i) column matrix [T^giTpg^T^gi) mdash
Q nonconservative generalized force vector
Qu actuator coupling matrix or longitudinal L Q R state weighting
matrix
v i i i
Rdm- inertial position of the t h link mass element drrii
RP transformation matrix relating qt and tf
Ru longitudinal L Q R input weighting matrix
RvRn Rd transformation matrices relating qt and q
S(f Mdc) generalized acceleration vector of coupled system q for the full
nonlinear flexible system
th uiirv l u w u u u maniA
th
T j i l ink rotation matrix
TtaTto control thrust in the pitch and roll direction for the im l ink
respectively
rQi maximum deploymentretrieval velocity of the t h l ink
Vei l ink strain energy
Vgi link gravitational potential energy
Rayleigh dissipation function arising from structural damping
in the ith link
di position vector to the frame F from the frame F_i bull
dc desired offset acceleration vector (i 1 l ink offset position)
drrii infinitesimal mass element of the t h l ink
dx^dy^dZi Cartesian components of d along the local vertical local horishy
zontal and orbit normal directions respectively
F-Q mdash
gi r igid and flexible position vectors of drrii fj + ltEraquolt5
ijk unit vectors 1 0 0 r 010-^ and 00 l r respectively
li length of the t h l ink
raj mass of the t h link
nfj total number of flexible modes for the i 1 l ink nXi + nyi + nZi
nq total number of generalized coordinates per link nfj + 7
nqq systems total number of generalized coordinates Nnq
ix
number of flexible modes in the longitudinal inplane and out-
of-plane transverse directions respectively for the i ^ link
number of attitude control actuators
- $ T
set of generalized coordinates for the itfl l ink which accounts for
interactions with adjacent links
qv---QtNT
set of coordinates for the independent t h link (not connected to
adjacent links)
rigid position of dm in the frame Fj
position of centre of mass of the itfl l ink relative to the frame Fj
i + $DJi
desired settling time of the j 1 attitude actuator
actuator force vector for entire system
flexible deformation of the link along the Xj yi and Zj direcshy
tions respectively
control input for flexible subsystem
Cartesian components of fj
actual and estimated state of flexible subsystem respectively
output vector of flexible subsystem
LIST OF FIGURES
1-1 A schematic diagram of the space platform based N-body tethered
satellite system 2
1-2 Associated forces for the dumbbell satellite configuration 3
1- 3 Some applications of the tethered satellite systems (a) multiple
communication satellites at a fixed geostationary location
(b) retrieval maneuver (c) proposed B I C E P S configuration 6
2- 1 Vector components of the i 1 and (i- l ) 1 chain links 15
2-2 Vector components in cylindrical orbital coordinates 20
2-3 Inertial position of subsatellite thruster forces 38
2- 4 Coupled force representation of a momentum-wheel on a rigid body 40
3- 1 Flowchart showing the computer program structure 45
3-2 Kinet ic and potentialenergy transfer for the three-body platform based
tethered satellite system (a) variation of kinetic and potential energy
(b) percent change in total energy of system 48
3-3 Kinet ic and potential energy transfer for the five-body platform based
tethered satellite system (a) variation of kinetic and potential energy
(b) percent change in total energy of system 49
3-4 Simulation results for the platform based three-body tethered
system originally presented in Ref[43] 51
3- 5 Simulation results for the platform based three-body tethered
system obtained using the present computer program 52
4- 1 Schematic diagram showing the generalized coordinates used to
describe the system dynamics 55
4-2 Stationkeeping dynamics of the three-body STSS configuration
without offset (a) attitude response (b) vibration response 57
4-3 Stationkeeping dynamics of the three-body STSS configuration
xi
with offset along the local vertical
(a) attitude response (b) vibration response 60
4-4 Stationkeeping dynamics of the three-body STSS configuration
with offset along the local horizontal
(a) attitude response (b) vibration response 62
4-5 Stationkeeping dynamics of the three-body STSS configuration
with offset along the orbit normal
(a) attitude response (b) vibration response 64
4-6 Stationkeeping dynamics of the three-body STSS configuration
wi th offset along the local horizontal local vertical and orbit normal
(a) attitude response (b) vibration response 67
4-7 Deployment dynamics of the three-body STSS configuration
without offset
(a) attitude response (b) vibration response 69
4-8 Deployment dynamics of the three-body STSS configuration
with offset along the local vertical
(a) attitude response (b) vibration response 72
4-9 Retrieval dynamics of the three-body STSS configuration
with offset along the local vertical and orbit normal
(a) attitude response (b) vibration response 74
4-10 Schematic diagram of the five-body system used in the numerical example 77
4-11 Stationkeeping dynamics of the five-body STSS configuration
without offset
(a) attitude response (b) vibration response 78
4-12 Deployment dynamics of the five-body STSS configuration
wi th offset along the local vertical
(a) attitude response (b) vibration response 81
4-13 Stationkeeping dynamics of the three-body B I C E P S configuration
x i i
with offset along the local vertical
(a) attitude response (b) vibration response 84
4-14 Cartwheeling dynamics with deployment for the three-body B I C E P S
with offset along the local vertical
(a) attitude response (b) vibration response 86
4-15 Spin dynamics (7 = l deg s ) of the three-body O E D I P U S configuration
with offset along the local vertical
(a) attitude response (b) vibration response 89
4- 16 Spin dynamics (7 = 10deg s ) of the three-body O E D I P U S configuration
wi th offset along the local vertical
(a) attitude response (b) vibration response 91
5- 1 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear rigid F L T controller with offset along the
local horizontal local vertical and orbit normal
(a) attitude and vibration response (b) control actuator time histories 98
5-2 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear flexible F L T controller with offset along the
local horizontal local vertical and orbit normal
(a) attitude and vibration response (b) control actuator time histories 101
5-3 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear rigid F L T controller with offset along
the local vertical
(a) attitude and vibration response (b) control actuator time histories 103
5-4 Deployment dynamics of the three-body STSS using the nonlinear
rigid F L T controller with offset along the local vertical
(a) attitude and vibration response (b) control actuator time histories 105
5-5 Retrieval dynamics of the three-body STSS using the non-linear
rigid F L T controller with offset along the local vertical and orbit normal
(a) attitude and vibration response (b) control actuator time histories 108
x i i i
5-6 Block diagram for the L Q G L T R estimator based compensator 115
5-7 Singular values for the L Q G and L Q G L T R compensator compared to target
return ratio (a) longitudinal design (b) transverse design 118
5-8 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear rigid F L T attitude controller and L Q G L T R
offset vibration controller (a) attitude and libration controller response
(b) vibration and offset response 122
xiv
A C K N O W L E D G E M E N T
Firs t and foremost I would like to express the most genuine gratitude to
my supervisor Prof V i n o d J M o d i whose unending guidance and encouragement
proved most invaluable during the course of my studies
I am also deeply indebted to Dr Satyabrata Pradhan for his limitless patience
and technical advice from which this work would not have been possible I would also
like to extend a word of appreciation towards Prof A r u n K Mis ra ( M c G i l l University)
for ini t iat ing my interest in space dynamics and control
Also I would like to thank my collegues Dr Sandeep Munshi M r Mathieu
Caron and M r Yuan Chen as well as Dr Mae Seto M r Gary L i m and M r Mark
Chu for their words of advice and helpful suggestions that heavily contributed to all
aspects of my studies
Final ly I would like to thank all my family and friends here in Vancouver and
back home in Montreal whose unending support made the past two years the most
memorable of my life
The research project was supported by the Natural Sciences and Engineering
Research Council s P G S - A Scholarship held by the author as well as by N S E R C grant
A-2181 held by Prof V J M o d i
xv
1 INTRODUCTION
11 Prel iminary Remarks
W i t h the ever changing demands of the worlds population one often wonders
about the commitment to the space program On the other hand the importance
of worldwide communications global environmental monitoring as well as long-term
habitation in space have demanded more ambitious and versatile satellites systems
It is doubtful that the Russian scientist Tsiolkovsky[l] when he first proposed the
uti l izat ion of the Earths gravity-gradient environment in the last century would have
envisioned the current and proposed applications of tethered satellite systems
A tethered satellite system consists of two or more subsatellites connected
to each other by long thin cables or tethers which are in tension (Figure 1-1) The
subsatellite which can also include the shuttle and space station are in orbit together
The first proposed use of tethers in space was associated with the rescue of stranded
astronauts by throwing a buoy on a tether and reeling it to the rescue vehicle
Prel iminary studies of such systems led to the discovery of the inherent instability
during the tether retrieval[2] Nevertheless the bir th of tethered systems occured
during the Gemini X I and X I I missions[3] in 1966 when a short tether (10-30m)
was used to generate the first artificial gravity environment in space (000015g) by
cartwheeling (spinning) the system about its center of mass
The mission also demonstrated an important use of tethers for gravity gradient
stabilization The force analysis of a simplified model illustrates this point Consider
the dumbbell satellite configuration orbiting about the centre of Ear th as shown in
Figure 1-2 Satellite 1 and satellite 2 are connected to each other by a tether with
1
Space Station (Satellite 1 )
2
Satellite N
Figure 1-1 A schematic diagram of the space platform based N-body tethered satellite system
2
Figure 1-2 Associated forces for the dumbbell satellite configuration
3
the whole system free to librate about the systems centre of mass C M The two
major forces acting on each body are due to the centrifugal and gravitational effects
Since body 1 is further away from the earth its gravitational force Fg^y is smaller
than that of body 2 ie Fgi lt Fg^ However its centrifugal force FCl is greater
then the centrifugal force experienced by the second satellite (FCl gt FC2) Moreover
for body 1 Fg^ lt Fc^ in contrast to Fg2 gt FC2 for body 2 Thus resolving the force
components along the tether there is an evident resultant tension force Ff present
in the tether Similarly adding the normal components of the two forces vectorially
results in a force FR which restores the system to its stable equil ibrium along the
local vertical These tension and gravity-gradient restoring forces are the two most
important features of tethered systems
Several milestone missions have flown in the recent past The U S A J a p a n
project T P E (Tethered Payload Experiment) [4] was one of the first to be launched
using a sounding rocket to conduct environmental studies The results of the T P E
provided support for the N A S A I t a l y shuttle based tethered satellite system referred
to as the TSS-1 mission[5] This experiment aimed to study the electrodynamic effects
of a shuttle borne satellite system with a conductive tether connection where an
electric current was induced in the tether as it swept through the Earth s magnetic
field Unfortunetely the two attempts in August 1992 and February 1996 resulted
in only partial success The former suffered from a spool failure resulting in a tether
deployment of only 256m of a planned 20km During the latter attempt a break
in the tethers insulation resulted in arcing leading to tether rupture Nevertheless
the information gathered by these attempts st i l l provided the engineers and scientists
invaluable information on the dynamic behaviour of this complex system
Canada also paved the way with novel designs of tethered system aimed
primari ly at the study of the ionosphere and Northern Lights (Aurora Borealis)
4
The O E D I P U S A and C (Observation of Electrified Distributions in the Ionospheric
P lasma - a Unique Strategy) missions[6] launched from a sounding rocket in January
1989 and November 1995 respectively provided insight into the complex dynamical
behaviour of two comparable mass satellites connected by a 1 km long spinning
tether
Final ly two experiments were conducted by N A S A called S E D S I and II
(Small Expendable Deployable System) [7] which hold the current record of 20 km
long tether In each of these missions a small probe (26 kg) was deployed from the
second stage of a D E L T A II rocket thus succesfully demonstrating the feasibility of
long tether deployment Retrieval of the tether which is significantly more difficult
has not yet been achieved
Several proposed applications are also currently under study These include the
study of earths upper atmosphere using probes lowered from the shuttle in a low earth
orbit ( L E O ) to the deployment and retrieval of satellites for servicing Even more
promising application concerns the generation of power for the proposed International
Space Station using conductive tethers The Canadian Space Agency (CSA) has also
proposed a dual-mass satellite system B I C E P S (Bl-static Canadian Experiment on
Plasmas in Space) to make simultaneous measurements at different locations in the
environment for correlation studies[8] A unique feature of the B I C E P S mission is
the deployment of the tether aided by the angular momentum of the system which
is ini t ia l ly cartwheeling about its orbit normal (Figure 1-3)
In addition to the three-body configuration multi-tethered systems have been
proposed for monitoring Earths environment in the global project monitored by
N A S A entitled Mission to Planet Ear th ( M T P E ) The proposed mission aims to
study pollution control through the understanding of the dynamic interactions be-
5
Retrieval of Satellite for Servicing
lt mdash
lt Satellite
Figure 1-3 Some applications of the tethered satellite systems (a) multiple comshymunication satellites at a fixed geostationary location (b) retrieval maneuver (c) proposed BICEPS configuration
6
tween the Earths atmosphere biosphere hydrosphere and cryosphere This wi l l be
accomplished with multiple tethered instrumentation payloads simultaneously soundshy
ing at different altitudes for spatial correlations Such mult ibody systems are considshy
ered to be the next stage in the evolution of tethered satellites Micro-gravity payload
modules suspended from the proposed space station for long-term experiments as well
as communications satellites with increased line-of-sight capability represent just two
of the numerous possible applications under consideration
12 Brief Review of the Relevant Literature
The design of multi-payload systems wi l l require extensive dynamic analysis
and parametric study before the final mission configuration can be chosen This
wi l l include a review of the fundamental dynamics of the simple two-body tethered
systems and how its dynamics can be extended to those of multi-body tethered system
in a chain or more generally a tree topology
121 Multibody 0(N) formulation
Many studies of two-body systems have been reported One of the earliest
contribution is by Rupp[9] who was the first to study planar dynamics of the simshy
plified Shuttle based Tethered Satellite System (STSS) His pioneering work led to
the discovery of pitch oscillation growth during the retrieval phase Later Baker[10]
advanced the investigation to the third dimension in addition to adding atmospheric
effects to the system A more complete dynamical model was later developed and
analyzed by M o d i and Misra[ l l 12] which included deployment and tether flexibility
A comprehensive survey of important developments and milestones in the area have
been compiled and reviewed by Mis ra and Modi[13] The major conclusions based on
the literature may be summarized as follows
bull the stationkeeping phase is stable
7
bull deployment can be unstable if the rate exceeds a crit ical speed
bull retrieval of the tether is inherently unstable
bull transverse vibrations can grow due to the Coriolis force induced during deshy
ployment and retrieval
Misra Amier and Modi[14] were one of the first to extend these results to the
three-satellite double pendulum case This simple model which includes a variable
length tether was sufficient to uncover the multiple equilibrium configurations of
the system Later the work was extended to include out-of-plane motion and a
preliminary reel-rate control law to regulate the librational motion[15] Kesmir i and
Misra[16] developed a general formulation for N-body tethered systems based on the
Lagrangian principle where three-dimensional motion and flexibility are accounted for
However from their work it is clear that as the number of payload or bodies increases
the computational cost to solve the forward dynamics also increases dramatically
Tradit ional methods of inverting the mass matrix M in the equation
ML+F = Q
to solve for the acceleration vector proved to be computationally costly being of the
order for practical simulation algorithms O(N^) refers to a mult ibody formulashy
tion algorithm whose computational cost is proportional to the cube of the number
of bodies N used It is therefore clear that more efficient solution strategies have to
be considered
Many improved algorithms have been proposed to reduce the number of comshy
putational steps required for solving for the acceleration vector Rosenthal[17] proshy
posed a recursive formulation based on the triangularization of the equations of moshy
tion to obtain an 0(N2) formulation He also demonstrates the use of a symbolic
manipulation program to reduce the final number of computations A recursive L a -
8
grangian formulation proposed by Book[18] also points out the relative inefficiency of
conventional formulations and proceeds to derive an algorithm which is also 0(N2)
A recursive formulation is one where the equations of motion are calculated in order
from one end of the system to the other It usually involves one or more passes along
the links to calculate the acceleration of each link (forward pass) and the constraint
forces (backward or inverse pass) Non-recursive algorithms express the equations of
motion for each link independent of the other Although the coupling terms st i l l need
to be included the algorithm is in general more amenable to parallel computing
A st i l l more efficient dynamics formulation is an O(N) algorithm Here the
computational effort is directly proportional to the number of bodies or links used
Several types of such algorithms have been developed over the years The one based
on the Newton-Euler approach has recursive equations and is used extensively in the
field of multi-joint robotics[1920] It should be emphasized that efficient computation
of the forward and inverse dynamics is imperative if any on-line (real-time) control
of the joint is to be achieved Hollerbach[21] proposed a recursive O(N) formulation
based on Lagranges equations of motion However his derivation was primari ly foshy
cused on the inverse dynamics and did not improve the computational efficiency of
the forward dynamics (acceleration vector) Other recursive algorithms include methshy
ods based on the principle of vir tual work[22] and Kanes equations of motion[23]
Keat[24] proposed a method based on a velocity transformation that eliminated the
appearance of constraint forces and was recursive in nature O n the other hand
K u r d i l a Menon and Sunkel[25] proposed a non-recursive algorithm based on the
Range-Space method which employs element-by-element approach used in modern
finite-element procedures Authors also demonstrated the potential for parallel comshy
putation of their non-recursive formulation The introduction of a Spatial Operator
Factorization[262728] which utilizes an analogy between mult ibody robot dynamics
and linear filtering and smoothing theory to efficiently invert the systems mass ma-
9
t r ix is another approach to a recursive algorithm Their results were further extended
to include the case where joints follow user-specified profiles[29] More recently Prad-
han et al [30] proposed a non-recursive Lagrangian algorithm for factorizing the mass
matrix leading to an O(N) formulation of the forward dynamics of the system A n
efficient O(N) algorithm where the number of bodies N in the system varies on-line
has been developed by Banerjee[31]
122 Issues of tether modelling
The importance of including tether flexibility has been demonstrated by several
researchers in the past However there are several methods available for modelling
tether flexibility Each of these methods has relative advantages and limitations
wi th regards to the computational efficiency A continuum model where flexible
motion is discretized using the assumed mode method has been proposed[3233] and
succesfully implemented In the majority of cases even one flexible mode is sufficient
to capture the dynamical behaviour of the tether[34] K i m and Vadali[35] proposed a
so-called bead model or lumped-mass approach that discretizes the tether using point
masses along its length However in order to accurately portray the motion of the
tether a significantly high number of beads are needed thus increasing the number of
computation steps Consequently this has led to the development of a hybrid between
the last two approaches[16] ie interpolation between the beads using a continuum
approach thus requiring less number of beads Finally the tether flexibility can also
be modelled using a finite-element approach[36] which is more suitable for transient
response analysis In the present study the continuum approach is adopted due to
its simplicity and proven reliability in accurately conveying the vibratory response
In the past the modal integrals arising from the discretization process have
been evaluted numerically In general this leads to unnecessary effort by the comshy
puter Today these integrals can be evaluated analytically using a symbolic in-
10
tegration package eg Maple V Subsequently they can be coded in F O R T R A N
directly by the package which could also result in a significant reduction in debugging
time More importantly there is considerable improvement in the computational effishy
ciency [16] especially during deployment and retrieval where the modal integral must
be evaluated at each time-step
123 Attitude and vibration control
In view of the conclusions arrived at by some of the researchers mentioned
above it is clear that an appropriate control strategy is needed to regulate the dyshy
namics of the system ie attitude and tether vibration First the attitude motion of
the entire system must be controlled This of course would be essential in the case of
a satellite system intended for scientific experiments such as the micro-gravity facilishy
ties aboard the proposed Internation Space Station Vibra t ion of the flexible members
wi l l have to be checked if they affect the integrity of the on-board instrumentation
Thruster as well as momentum-wheel approach[34] have been considered to
regulate the rigid-body motion of the end-satellite as well as the swinging motion
of the tether The procedure is particularly attractive due to its effectiveness over a
wide range of tether lengths as well as ease of implementation Other methods inshy
clude tension and length rate control which regulate the tethers tension and nominal
unstretched length respectively[915] It is usually implemented at the deployment
spool of the tether More recently an offset strategy involving time dependent moshy
tion of the tether attachment point to the platform has been proposed[3738] It
overcomes the problem of plume impingement created by the thruster control and
the ineffectiveness of tension control at shorter lengths However the effectiveness of
such a controller can become limited with an exceedingly long tether due to a pracshy
tical l imi t on the permissible offset motion In response to these two control issues a
hybrid thrusteroffset scheme has been proposed to combine the best features of the
11
two methods[39]
In addition to attitude control these schemes can be used to attenuate the
flexible response of the tether Tension and length rate control[12] as well as thruster
based algorithms[3340] have been proposed to this end M o d i et al [41] have sucshy
cessfully demonstrated the effectiveness of an offset strategy to damp undesired v i shy
brations Passive energy dissipative devices eg viscous dampers are also another
viable solution to the problem
The development of various control laws to implement the above mentioned
strategies has also recieved much attention A n eigenstructure assignment in conducshy
tion wi th an offset controller for vibration attenuation and momemtum wheels for
platform libration control has been developed[42] in addition to the controller design
from a graph theoretic approach[41] Also non-linear feedback methods such as the
Feedback Linearization Technique ( F L T ) that are more suitable for controlling highly
nonlinear non-autonomous coupled systems have also been considered[43]
It is important to point out that several linear controllers including the classhy
sic state feedback Linear Quadratic Regulator ( L Q R ) have received considerable
attention[3944] Moreover robust methods such as the Linear Quadratic Guas-
s ian Loop Transfer Recovery ( L Q G L T R ) method have also been developed and imshy
plemented on tethered systems[43] A more complete review of these algorithms as
well as others applied to tethered system has been presented by Pradhan[43]
13 Scope of the Investigation
The objective of the thesis is to develop and implement a versatile as well as
computationally efficient formulation algorithm applicable to a large class of tethered
satellite systems The distinctive features of the model can be summarized as follows
12
bull TV-body 0(N) tethered satellite system in a chain-type topology
bull the system is free to negotiate a general Keplerian orbit and permitted to
undergo three dimensional inplane and out-of-plane l ibrtational motion
bull r igid bodies constituting the system are free to execute general rotational
motion
bull three dimensional flexibility present in the tether which is discretized using
the continuum assumed-mode method with an arbitrary number of flexible
modes
bull energy dissipation through structural damping is included
bull capability to model various mission configurations and maneuvers including
the tether spin about an arbitrary axis
bull user-defined time dependent deployment and retrieval profiles for the tether
as well as the tether attachment point (offset)
bull attitude control algorithm for the tether and rigid bodies using thrusters and
momentum-wheels based on the Feedback Linearization Technique
bull the algorithm for suppression of tether vibration using an active offset (tether
attachment point) control strategy based on the optimal linear control law
( L Q G L T - R )
To begin with in Chapter 2 kinematics and kinetics of the system are derived
using the O(N) formulation methodology developed by Pradhan et al [30] Chapshy
ter 3 discusses issues related to the development of the simulation program and its
validation This is followed by a detailed parametric study of several mission proshy
files simulated as particular cases of the versatile formulation in Chapter 4 Next
the attitude and vibration controllers are designed and their effectiveness assessed in
Chapter 5 The thesis ends with concluding remarks and suggestions for future work
13
2 F O R M U L A T I O N OF T H E P R O B L E M
21 Kinematics
211 Preliminary definitions and the i ^ position vector
The derivation of the equations of motion begins with the definition of the
inertial position of an arbitrary link of the mult ibody system Let the i ^ link of the
system be free to translate and rotate in 3-D space From Figure 2-1 the position
vector Rdm^ to the mass element drrn on the i ^ link wi th reference to the inertial
frame FQ can be written as
Here D represents the inertial position of the t h body fixed frame Fj relative to FQ
ri = [xiyiZi]T is the rigid position vector of drrii wi th respect to F J and f(fi) is
the flexible deformation at fj also with respect to the frame Fj Bo th these vectors
are relative to Fj Note in the present model tethers (i even) are considered flexible
and X corresponds to the nominal position along the unstretched tether while a and
ZJ are by definition equal to zero On the other hand the rigid bodies referred to as
satellites (i odd) including the platform are taken to be rigid ie f(fj) = 0 T j is
the rotation matrix used to express body-fixed vectors with reference to the inertial
frame
212 Tether flexibility discretization
The tether flexibility is discretized with an arbitrary but finite number of
Di + Ti(i + f(i)) (21)
14
15
modes in each direction using the assumed-mode method as
Ui nvi
wi I I i = 1
nzi bull
(22)
where nXi nyi and n 2 ^ are the number of modes used in the X m and Z directions
respectively For the ] t h longitudinal mode the admissible mode shape function is
taken as 2 j - l
Hi(xiii)=[j) (23)
where l is the i ^ tether length[32] In the case of inplane and out-of-plane transverse
deflections the admissible functions are
ampyi(xili) = ampZixuk) = v ^ s i n JKXj
k (24)
where y2 is added as a normalizing factor In this case both the longitudinal and
transverse mode shapes satisfy the geometric boundary conditions for a simple string
in axial and transverse vibration
Eq(22) is recast into a compact matrix form as
ff(i) = $i(xili)5i(t) (25)
where $i(x l) is the matrix of the tether mode shapes defined as
0 0
^i(xiJi) =
^ X bull bull bull ^XA
0
0 0
0
G 5R 3 x n f
(26)
and nfj = nx- + nyi + nZi Note $J(XJJ) is a function of the spatial coordinate x
and not of yi or ZJ However it is also a function of the tether length parameter li
16
which is time-varying during deployment and retrieval For this reason care must be
exercised when differentiating $J(XJJ) with respect to time such that
d d d k) = Jtregi(Xi li) = ^ 7 $ t ( ^ i k)Xi + gf$i(xigt li)k- (2-7)
Since x represents the nominal rate of change of the position of the elemental mass
drrii it is equal to j (deploymentretrieval rate) Let t ing
( 3 d
dx~ + dT)^Xili^ ( 2 8 )
one arrives at
$i(xili) = $Di(xili)ii (29)
The time varying generalized coordinate vector 5 in Eq(25) is composed of
the longitudinal and transverse components ie
ltM) G s f t n f i x l (210)
where 5X^ 5y- and Sz^ are nx- x 1 ny x 1 and nz- x 1 size vectors and correspond to
$ X and $ Z respectively
213 Rotation angles and transformations
The matrix T j in Eq(21) represents a rotation transformation from the body
fixed frame Fj to the inertial frame FQ ie FQ = T j F j It is defined by the 3-2-1
ordered sequence of elementary rotations[46]
1 P i tch F[ = Cf ( a j )F 0
2 R o l l F = C f (3j)F = C7f (3j)Cf ( a j )F 0
3 Yaw F = C7( 7 i )F = C7(7l)cf^)Cf(ai)FQ = Q F 0
17
where
c f (A) =
-Sai Cai
0 0
and
C(7i ) =
0 -s0i] = 0 1 0
0 cpi
i 0 0 = 0
0
(211)
(212)
(213)
Here Eq(211) corresponds to a rotation of ctj (pitch) about the inertial axis ZQ (orbit
normal) of frame FQ resulting in a new intermediate frame F It is then followed by
a roll rotation fa about the axis y of F- giving a second intermediate frame F-
Final ly a spin rotation 7J about the axis z of F- is given resulting in the body
fixed frame Fj for the i lt l link Since all the three rotation matrices are orthogonal
it follows that FQ = C~1Fi = Cj F and hence
C CPi Cai ~ CH Sai + SrYi SPi Sai 5 7 i Sampi + Cli SPi Cci CPi Sai C1i Cai + 5 7 i SPi ~ S1 Ca + C 7 bull Sp Sai
(214) SliCPi CliCPi
Here Cx and Sx are abbreviations for cos(x) and sin(x) respectively
214 Inertial velocity of the ith link
mdash Lett ing pj = fj - f $j(5j and differentiating Eq(21) with respect to time gives
m- D j + T j f j + T^Si + T j pound + T j $ j 5 j (215)
Each of the terms appearing in Eq(215) must be addressed individually Since
fj represents the rigid body motion within the link it only has meaning for the
deployable tether and since yi = Zj = 0 for al l tether links (i even) fj = Ifi where 1
is the unit vector 1 0 0T For the case of rigid links (i odd) fj = 0
18
Evaluating the derivative of each angle in the rotation matrix gives
T^i9i = Taiai9i + Tppigi + T^fagi (216)
where Tx = J j I V Collecting the scaler angular velocity terms and defining the set
Vi = deg^gt A) 7i^ Eq(216) can be rewritten as
T i f t = Pi9i)fji (217)
where Pi(gi) is a column matrix defined by
Pi(9i)Vi = [Taigi^p^iT^gilffi (218)
Inserting Eq(29) and Eq(217) into Eq(215) and defining Sj = 2 + leads to
215 Cylindrical orbital coordinates
The Di term in Eq(219) describes the orbital motion of the t h l ink and
is composed of the three Cartesian coordinates DXi A ^ DZi However over the
cycle of one orbit DXi and Dyi vary dramatically by around the order of Earths
radius This must be avoided since large variations in the coordinates can cause
severe truncation errors during their numerical integration For this reason it is
more convenient to express the Cartesian components in terms of more stationary
variables This is readily accomplished using cylindrical coordinates
However it is only necessary to transform the first l inks orbital components
ie Lgti since only D is a generalized coordinates From Figure 2-2 it is apparent
(219)
that
(220)
19
Figure 2-2 Vector components in cylindrical orbital coordinates
20
Eq(220) can be rewritten in matrix form as
cos(0i) 0 0 s i n ( 0 i ) 0 0
0 0 1
Dri
h (221)
= DTlDSv
where Dri is the inplane radial distance 9 the true anomaly and DZl the out-of-
plane distance normal to the orbital plane The total derivative of D wi th respect
to time gives in column matrix form
1 1 1 1 J (222)
= D D l D 3 v
where [cos(^i) -Dnsm(8i) 0 ]
(223)
For all remaining links ie i = 2 N
cos(0i) - A - j S i n ^ i ) 0 sin(^i) D r i c o s ( 0 i ) 0
0 0 1
DTl = D D i = I d 3 i = 2N (224)
where is the n x n identity matrix and
2N (225)
216 Tether deployment and retrieval profile
The deployment of the tether is a critical maneuver for this class of systems
Cr i t i ca l deployment rates would cause the system to librate out of control Normally
a smooth sinusoidal profile is chosen (S-curve) However one would also like a long
tether extending to ten twenty or even a hundred kilometers in a reasonable time
say 3-4 orbits This is usually difficult or impossible to accomplish solely wi th one
S-curve profile Often a composite S-curve-Steady-S-curve profile is adopted Thus
21
the deployment scheme for the t h tether can be summarized as
k = Tr l - cos (^r(ti - to-)) 1 ltUlt h 2 r V A V
k = Vov tiilttiltt2i
k = IT j 1 - C deg S lkitl ^ ^ lt i lt ^
(226)
where to- tt are the ini t ia l and final times of the deployment or retrieval maneuver
respectively A t j = t - tQ mdash tt mdash t2- and the steady deployment velocity VQ is
calculated based on the continuity of l and l at the specified intermediate times t
and to For the case of
and
Ai = i(ti) - ^(t 0 -) = - ^ ( t 2 i ) (227)
2Ali + li(tf) -h(tQ)
V = lUli l^lL (228)
1 tf-tQ v y
A t = (229)
22 Kinetics and System Energy
221 Kinetic energy
The kinetic energy of the ith link is given by
Kei = lJm Rdrnfi^Ami (230)
Setting Eq(219) in a matrix-vector form gives the relation
4 ^ ^ PM) T i $ 4 TiSi]i[t (231)
22
where
m
k J
x l (232)
and nq = -nfj + 7 is the number of generalized coordinates in each link Inserting
Eq(231) into Eq(230) and integrating the kinetic energy for the t h l ink can be
rewritten as
1 T
(233)
where Mti is the links symmetric mass matrix
Mt =
m i D D T D D DDTPifgidmi) DDTTi J ^dm D n T Sidm rn
sym fP(9i)Pi(9i)dmi J P[g^T^idm J P ( pound ) T ^ m sym sym J ^f^idm f ^f^dmi sym sym sym J sfsidmi
pound Sfttiqxnq
(234)
and mi is the mass of the Ith link The kinetic energy for the entire system ie N
bodies can now be stated as
1 bull T bull 1 bull T Ke = J2 MtA = o laquo Mtqt 2 ^ -i
i=l (235)
where qt = qj^^ bull bull -qfNT a n d Mt is a block diagonal decoupled mass matrix
of the system with Mf- on its diagonal
(236) Mt =
0 0
0 Mt2
0
0 0
MH
0 0 0
0 0 0 bull MtN
222 Simplification for rigid links
The mass matrix given by Eq(234) simplifies considerably for the case of rigid
23
links and is given as
Mti
m D D l
T D D l DD^Piiffidmi) 0
sym 0
0
fP(fi)Pi(i)dmi 0 0 0
0 Id nfj
0 m~
i = 1 3 5 N
(237)
Moreover when mult iplying the matrices in the (11) block term of Eq(237) one
gets
D D l
1 D D l = 1 0 0 0 D n
2 0 (238) 0 0 1
which is a diagonal matrix However more importantly i t is independent of 6 and
hence is always non-singular ensuring that is always invertible
In the (12) block term of M ^ it is apparent that f fidmi is simply the
definition of the centre of mass of the t h rigid link and hence given by
To drrii mdash miff (239)
where fcmi is the position vector of the centre of mass of link i For the case of the
(22) block term a little more attention is required Expanding the integrand using
Eq(218) where ltj = fj gives
fj Ta Tafi fj T Q Tpfi f- T Q T 7 i f j
sym sym
Now since each of the terms of Eq(240) is a scaler or a 1 x 1 matrix it is therefore
equal to its trace ie sum of its diagonal elements thus
tr(ffTai
TTaifi) tr(ffTai
TTpfi) tr(ff Ta^T^fi)
fjT^Tpn rfTp^fi (240)
P(rj)Pj(fi) = syrri trifjTpTTpfi) t r ^ T ^ T ^
sym sym tr(ff T 7 T r 7 f j) J
(241)
Using two properties of the trace function[47] namely
tr(ABC) = tr(BCA) = trCAB) (242)
24
and
j tr(A) =tr(^J A^j (243)
where A B and C are real matrices it becomes quite clear that the (22) block term
of Eq(237) simplifies to
P(fi)Pii)dmi =
tr(Tai
TTaiImi) tTTai
TTpImi) t r (T Q r T 7 7 m ) trTaTTpImi) tr(Tp
TT7Jmi) sym sym sym
(244)
where Im- is an alternate form of the inertia tensor of the t h link which when
expanded is
Imlaquo = I nfi dm bull i sym
sym sym
J x^dmi J Xydmi XZdm sym J yi2dmi f yiZidm
sym J z^dmi J
xVi -L VH sym ijxxi + lyy IZZJ)2
(245)
Note this is in terms of moments of inertia and products of inertia of the t f l rigid
link A similar simplification can be made for the flexible case where T is replaced
with Qi resulting in a new expression for Im
223 Gravitational potential energy
The gravitational potential energy of the ith link due to the central force law
can be written as
V9i f dm f
-A mdash = -A bullgttradelti RdmA
dm
inn Lgti + T j ^ | (246)
where z = 3986 x 105 km 3s 2 (Earths gravitational constant) Expanding binomially
and retaining up to third order terms gives
Vgi ~ ~ D T 2D [J gfgidrrii + 23f T j gidm mdashDT
D2 i l I I Df Tt I grffdmiTiDi
(2-lt 7)
25
where D = D is the magnitude of the inertial position vector to the frame F The
integrals in the above expression are functions of the flexible mode shapes and can be
evaluated through symbolic manipulation using an existing package such as Maple
V In turn Maple V can translate the evaluated integrals into F O R T R A N and store
the code in a file Furthermore for the case of rigid bodies it can be shown quite
readily that
j gjgidrrii = j rjfidrrn = [IXXi + Iyy + IZZi)2 (248)
Similarly the other integrals in Eq(234) can be integrated symbolically
224 Strain energy
When deriving the elastic strain energy of a simple string in tension the
assumptions of high tension and low amplitude of transverse vibrations are generally
valid Consequently the first order approximation of the strain-displacement relation
is generally acceptable However for orbiting tethers in a weak gravitational-gradient
field neglecting higher order terms can result in poorly modelled tether dynamics
Geometric nonlinearities are responsible for the foreshortening effects present in the
tether These cannot be neglected because they account for the heaving motion along
the longitudinal axis due to lateral vibrations The magnitude of the longitudinal
oscillations diminish as the tether becomes shorter[35]
W i t h this as background the tether strain which is derived from the theory
of elastic vibrations[3248] is given by
(249)
where Uj V and W are the flexible deformations along the Xj and Z direction
respectively and are obtained from Eq(25) The square bracketed term in Eq(249)
represents the geometric nonlinearity and is accounted for in the analysis
26
The total strain energy of a flexible link is given by
Ve = I f CiEidVl (250)
Substituting the stress-strain relationship 0 = E(poundi the strain energy is now given
by
Vei=l-EiAi J l l d x h (251)
where E^ is the tethers Youngs modulus and A is the cross-sectional area The
tether is assumed to be uniform thus EA which is the flexural stiffness of the
tether is assumed to be constant Eq(251) is evaluated symbolically for arbitrary
number of modes
225 Tether energy dissipation
The evaluation of the energy dissipation due to tether deformations remains
a problem not well understood even today Here this complex phenomenon is repshy
resented through a simplified structural damping model[32] In addition the system
exhibits hysteresis which also must be considered This is accomplished using an
augmented complex Youngs modulus
EX^Ei+jEi (252)
where Ej^ is the contribution from the structural damping and j = The
augmented stress-strain relation is now given as
a = Epoundi = Ei(l+jrldi)ei (253)
where 77^ = EjE^ ltC 1 is defined as the structural damping coefficient determined
experimentally [49]
27
If poundi is a harmonic function with frequency UJQ- then
jet = (254) wo-
Substituting into Eq(253) and rearranging the terms gives
e + I mdash I poundi = o-i + ad (255)
where ad and a are the stress with and without structural damping respectively
Now the Rayleigh dissipation function[50] for the ith tether can be expressed as
(ii) Jo
_ 1 EjAiVdi rU (2-56^ 2 w 0 Jo
The strain rate ii is the time derivative of Eq(249) The generalized external force
due to damping can now be written as Qd = Q^ bull bull bull QdNT where
dWd
^ = - i p = 2 4 ( 2 5 7 )
= 0 i = 13 N
23 0(N) Form of the Equations of Mot ion
231 Lagrange equations of motion
W i t h the kinetic energy expression defined and the potential energy of the
whole system given by Pe = ]Cn=l( 5i+ ej) the equations of motion can be obtained
quite readily using the Lagrangian principle
where q is the set of generalized coordinates to be defined in a later section
Substituting Eqs(235247251 and 257) into Eq(258) leads to the familiar
matr ix form of the non-linear non-autonomous coupled equations of motion for the
28
system
Mq t)t+ F(q q t) = Q(q q t) (259)
where M(q t) is the nonlinear symmetric mass matrix and F(q q t) is the forcing
function which can be written in matrix form as
Q(ltfgt Qi 0 is the vector of non-conservative generalized forces including control inshy
puts acting on the system A detailed expansion of Eq(260) in tensor notation is
developed in Appendix I
Numerical solution of Eq(259) in terms of q would require inversion of the
mass matrix However M is a full matrix of size nqq x nqq where nqq = Nnq is the
total number of generalized coordinates Therefore direct inversion of M would lead
to a large number of computation steps of the order nqq^ or higher The objective
here is to develop an order-N O(N) algorithm for obtaining M _ 1 that minimizes
the computational effort This is accomplished through the following transformations
ini t ia l ly developed for the stationkeeping case by Pradhan et al [30]
232 Generalized coordinates and position transformation
During the derivation of the energy expressions for the system the focus was
on the decoupled system ie each link was considered independent of the others
Thus each links energy expression is also uncoupled However the constraints forces
between two links must be incorporated in the final equations of motion
Let
e Unqxl (261) m
be the set of coupled coordinates of the t h l ink such that q = qT q2 bull bull bull 0T i s the
29
systems generalized coordinates Let qt- be the set of auxiliary decoupled coordinates
defined by Eq(232) such that qt = qj^ qj^ Qtj^1 bull The only difference between
qtj and q is the presence of D against d is the position of F wi th respect to
the inertial frame FQ whereas d is defined as the offset position of F relative to and
projected on the frame It can be expressed as d = dXidyidZi^ For the
special case of link 1 D = d
From Figure 2-1 can written as
Di = A - l + T i _ i J i _ i + di) + T V i S i - i f t - i + dXi)8i- (262)
Denoting KAj = $j(Zj + dx- ) Eq(262) can be rewritten in summation form as
i ampi = H ( T i - 1 ^ + T i - l K A i - l - l + T j - l ^ - l ) bull ( 2 - 6 3 )
3 = 1
Introducing the index substitution k mdash j mdash 1 into Eq(263)
i-1 D = ^2 (Tfc_ilt4 + TkKAk6k + Tkilk) + T^di (264)
k=l
since d = D TQ = 1^ and IQ DQ SQ are null vectors Recasting Eq(264) in
matr ix form leads to i-1
Qt Jfe=l
where
and
For the entire system
T fc-1 0 TkKAk
0 0 0 0 0 0 0 0 0 0 0 0 -
r T i - i 0 0 0 0 0 0 0 0 0 G
0 0 0 1
Qt = RpQ
G R n lt x l
(265)
(266)
G Unq x l (267)
(268)
30
where
RP
R 1 0 0 R R 2 0 R RP
2 R 3
0 0 0
R RP R PN _ 1 RN J
(269)
Here RP is the lower block triangular matrix which relates qt and q
233 Velocity transformations
The two sets of velocity coordinates qti and qi are related by the following
transformations
(i) First Transformation
Differentiating Eq(262) with respect to time gives the inertial velocity of the
t h l ink as
D = A - l + Ti_iK + + K A i _ i ^ _ i
+ T j _ i ^ + Ti-ii + T i - i K A i - i ^ + T j - i K A ^ i ^ i (270)
(271) (
Defining K A D = dddXi+1KAi K A L = d d K A and hi = di+1 +lfi +KA^i
results in K A ^ i = K A D j i o ^ + K A L i _ i i _ i
= KADi_iiTdi + K A L j _ i ^ _ i
Inserting Eq(271) into Eq(270) and rearranging gives
= A _ + T i ^ d g + K A D ^ x ^ i ^ ^ + Pi-^hi-iH-i
+ T i _ i K A i _ i 5 _ i + T i_ i2 + K A L j _ i 5 j _ i 4 _ i (272)
Using Eq(272) to define recursively - D j _ i and applying the index substitution
c = j mdash 1 as in the case of the position transformation it follows that
Di = ^ T ^ i K D ^ ^ + Pk(hk)ffk + T f c K A f c ^ + TfcKLfc4 Jfc=l (273)
+ T j _ i K D j _ i lt i ii
31
where K D j _ i = + K A D ^ i ^ - i r 7 and K L j = i + K A L j 5 j Finally by following a
procedure similar to that used for the position transformation it can be shown that
for the entire system
t = Rvil (2-74)
where Rv is given by the lower block diagonal matrix
Here
Rv
Rf 0 0 R R$ 0 R Rl RJ
0 0 0
T3V nN-l
R bullN
e Mnqqxl
Ri =
T i _ i K D i _ i TiKAi T j K L j 0 0 0 0 bullo 0 0 0 0 0 0 0
e K n lt x l
RS
T i-1 0 0 o-0 0 0 0 0 0
0 0 0 1
G Mnq x l
(275)
(276)
(277)
(ii) Second Transformation
The second transformation is simply Eq(272) set into matrix form This leads
to the expression for qt as
(278)
where Id Pihi) T j K A j T j K L j 0 0 0 0 0 0
0 0
0 0
0 0
G $lnq x l
Thus for the entire system
(279)
pound = Rnjt + Rdq
= Vdnqq - R nrlRdi (280)
32
where
and
- 0 0 0 RI 0 0
Rn = 0 0
0 0
Rf 0 0 0 rgtd K2 0
Rd = 0 0 Rl
0 0 0
RN-1 0
G 5R n w x l
0 0 0 e 5 R N ^ X L
(281)
(282)
234 Cylindrical coordinate modification
Because of the use of cylindrical coorfmates for the first body it is necessary
to slightly modify the above matrices to account for this coordinate system For the
first link d mdash D = D ^ D S L thus
Qti = 01 =gt DDTigi (283)
where
D D T i
Eq(269) now takes the form
DTL 0 0 0 d 3 0 0 0 ID
0 0 nfj 0
RP =
P f D T T i R2 0 P f D T T i Rp
2 R3
iJfDTTi i^
x l
0 0 0
R PN-I RN
(284)
(285)
In the velocity transformation a similar modification is necessary since D =
DD^DS^ Mak ing the substitution it can be shown that Eq(276) and Eq(279) now
33
takes the form for i = 1
T)V _ pn _ r t j mdash rt^ mdash 0 0 0 0
0 0 0 0 0 0 0 0
(286)
The rest of the matrices for the remaining links ie i = 2 N are unchanged
235 Mass matrix inversion
Returning to the expression for kinetic energy substitution of Eq(274) into
Eq(235) gives the first of two expressions for kinetic energy as
Ke = -q RvT MtRv (287)
Introduction of Eq(280) into Eq(235) results in the second expression for Ke as
1X 2q (hnqq-RnrlRd Mt (idnqq - R nrlRd (288)
Thus there are two factorizations for the systems coupled mass matrix given as
M = RV MtRv
M (hnqq - RnrlRd Mt ( i d n q q - R n r 1 R d
(289)
(290)
Inverting Eq(290)
r-1 _ ( rgtd~^ M 1 = (Ra) l d n q q - Rn)Mf1 (Rd)~l(Idnqq - Rn) (291)
Since both Rd and Mt are block diagonal matrices their inverse is simply the inverse
34
of each block on the diagonal ie
Mr1
h 0 0
0 0
0 0
0 0 0
0
0
0
tN
(292)
Each block has the dimension nq x nq and is independent of the number of bodies
in the system thus the inversion of M is only linearly dependent on N ie an O(N)
operation
236 Specification of the offset position
The derivation of the equations of motion using the O(N) formulation requires
that the offset coordinate d be treated as a generalized coordinate However this
offset may be constrained or controlled to follow a prescribed motion dictated either by
the designer or the controller This is achieved through the introduction of Lagrange
multipliers[51] To begin with the multipliers are assigned to all the d equations
Thus letting f = F - Q Eq(259) becomes
Mq + f = P C A (293)
where A G K 3 - 7 V x l is the vector of Lagrange multipliers and Pc G s R n 9 x 3 A r is the
permutation matrix assigning the appropriate Aj to its corresponding d equation
Inverting M and pre-multiplying both sides by PcT gives
pcTM-pc
= ( dc + PdegTM-^f) (294)
where dc is the desired offset acceleration vector Note both dc and PcTM lf are
known Thus the solution for A has the form
A pcTM-lpc - l dc + PC1 M (295)
35
Now substituting Eq(295) into Eq(293) and rearranging the terms gives
q = S(f Ml) = -M~lf + M~lPc
which is the new constrained vector equation of motion with the offset specified by
dc Note that pre-multiplying M~l by P c T provides the rows of M _ 1 corresponding
to the d equations Similarly post-multiplying by Pc leads the columns of the matrix
corresponding to the d equations
24 Generalized Control Forces
241 Preliminary remarks
The treatment of non-conservative external forces acting on the system is
considered next In reality the system is subjected to a wide variety of environmental
forces including atmospheric drag solar radiation pressure Earth s magnetic field as
well as dissipative devices to mention a few However in the present model only
the effect of active control thrusters and momemtum-wheels is considered These
generalized forces wi l l be used to control the attitude motion of the system
In the derivation of the equations of motion using the Lagrangian procedure
the contribution of each force must be properly distributed among all the generalized
coordinates This is acheived using the relation
nu ^5 Qk=E Pern ^ (2-97)
m=l deg Q k
where Qk and qk are the kth elements of Q and ltf respectively[51] Fem is the
mth external force acting on the system at the location Rm from the inertial frame
and nu is the total number of actuators Derivation of the generalized forces due to
thrusters and momentum wheels is given in the next two subsections
pcTM-lpc (dc + PcTM- Vj (2-96)
36
242 Generalized thruster forces
One set of thrusters is placed on each subsatellite ie al l the rigid bodies
except the first (platform) as shown in Figure 2-3 They are capable of firing in the
three orthogonal directions From the schematic diagram the inertial position of the
thruster located on body 3 is given as
Defining
R3 = di + Tid2 + T2(hi + KA2lt52) + T 3 f c m 3
= d + dfY + T 3 f c m 3
df = Tjdj+i + Tj+1lj+ii + KAj+i5j+i)
the thruster position for body 5 is given as
(298)
(299)
R5 = di+ dfx + d 3 + T5fcm
Thus in general the inertial position for the thruster on the t h body can be given
(2100)
as i - 2
k=l i1 cm (2101)
Let the column matrix
[Q i iT 9 R-K=rR dqj lA n i
WjRi A R i (2102)
where i = 3 5 N and j mdash 1 2 i The vector derivative of R wi th respect to
the scaler components qk is stored in the kth column of Thus for the case of
three bodies Eq(297) for thrusters becomes
1 Q
Q3
3 -1
T 3 T t 2 (2103)
37
Satellite 3
mdash gt -raquo cm
D1=d1
mdash gt
a mdash gt
o
Figure 2-3 Inertial position of subsatellite thruster forces
38
where ft = 0TtaTt^T is the thrust vector for the t h l ink (tether i-1) Here
Ttn and Tta are the thrust forces along the inplane m and out-of-plane Z directions i Pi
respectively For the case of 5-bodies
Q
~Qgt1
0
0
( T ^ 2 ) (2104)
The result for seven or more bodies follows the pattern established by Eq(2104)
243 Generalized momentum gyro torques
When deriving the generalized moments arising from the momemtum-wheels
on each rigid body (satellite) including the platform it is easier to view the torques as
coupled forces From Figure 2-4 it is apparent that for the ith l ink the generalized
forces constituting the couple are
(2105)
where
Fei = Faij acting at ei = eXi
Fe2 = -Fpk acting at e 2 = C a ^ l
F e 3 = F7ik acting at e 3 = eyj
Expanding Eq(2105) it becomes clear that
3
Qk = ^ F e i
i=l
d
dqk (2106)
which is independent of Ri- Thus the moments Mma = 2FaieXi Mm^ = 2FpeXi
and Mmi = 2FlieVi Transforming the coordinates to the body fixed frame using
the rotation matrix the generalized force due to the momentum-wheels on link i can
39
Figure 2-4 Coupled force representation of a momentum-wheel on a rigid body
40
be written as
d d d Qi = Td bull ^ T i t M m a - Tik bull mdashT^Mmp + T ^ bull T J j M ^
QiMmi
where Mm = Mma M m Mmi T
lt3 T J ^ T - T i f c bull ^ T i pound T ^ - ^ T z j
(2107)
(2108)
Note combining the thruster forces and momentum-wheel torques a compact
expression for the generalized force vector for 5 bodies can be written as
Q
Q T 3 0 Q T 5 0
0 0 0
0 lt33T3 Qyen Q5
3T5 0
0 0 0 Q4T5 0
0 0 0 Q5
5T5
M M I
Ti 2 m 3
Quu (2109)
The pattern of Eq(2109) is retained for the case of N links
41
3 C O M P U T E R I M P L E M E N T A T I O N
31 Prel iminary Remarks
The governing equations of motion for the TV-body tethered system were deshy
rived in the last chapter using an efficient O(N) factorization method These equashy
tions are highly nonlinear nonautonomous and coupled In order to predict and
appreciate the character of the system response under a wide variety of disturbances
it is necessary to solve this system of differential equations Although the closed-
form solutions to various simplified models can be obtained using well-known analytic
methods the complete nonlinear response of the coupled system can only be obtained
numerically
However the numerical solution of these complicated equations is not a straightshy
forward task To begin with the equations are stiff[52] be they have attitude and
vibration response frequencies separated by about an order of magnitude or more
which makes their numerical integration suceptible to accuracy errors i f a properly
designed integration routine is not used
Secondly the factorization algorithm that ensures efficient inversion of the
mass matrix as well as the time-varying offset and tether length significantly increase
the size of the resulting simulation code to well over 10000 lines This raises computer
memory issues which must be addressed in order to properly design a single program
that is capable of simulating a wide range of configurations and mission profiles It
is further desired that the program be modular in character to accomodate design
variations with ease and efficiency
This chapter begins with a discussion on issues of numerical and symbolic in-
42
tegration This is followed by an introduction to the program structure Final ly the
validity of the computer program is established using two methods namely the conshy
servation of total energy and comparison of the response results for simple particular
cases reported in the literature
32 Numerical Implementation
321 Integration routine
A computer program coded in F O R T R A N was written to integrate the equashy
tions of motion of the system The widely available routine used to integrate the difshy
ferential equations D G E A R [ 5 3 ] is based on Gears predictor-corrector method[54]
It is well suited to stiff differential equations as i t provides automatic step-size adjustshy
ment and error-checking capabilities at each iteration cycle These features provide
superior performance over traditional 4^ order Runge-Kut ta methods
Like most other routines the method requires that the differential equations be
transformed into a first order state space representation This is easily accomplished
by letting
thus
X - ^ ^ f ^ ) =ltbullbull) (3 -2 )
where S(fMdc) is defined by Eq(296) Eq(32) represents the new set of 2nqq
first order differential equations of motion
322 Program structure
A flow chart representing the computer programs structure is presented in
Figure 3-1 The M A I N program is responsible for calling the integration routine It
also coordinates the flow of information throughout the program The first subroutine
43
called is INIT which initializes the system variables such as the constant parameters
(mass density inertia etc) and the ini t ial conditions I C (attitude angles orbital
motion tether elastic deformation etc) from user-defined data files Furthermore all
the necessary parameters required by the integration subroutine D G E A R (step-size
tolerance) are provided by INIT The results of the simulation ie time history of the
generalized coordinates are then passed on to the O U T P U T subroutine which sorts
the information into different output files to be subsequently read in by the auxiliary
plott ing package
The majority of the simulation running time is spent in the F C N subroutine
which is called repeatedly by D G E A R F C N generates the fv(Xi) vector defined in
Eq(32) It is composed of several subroutines which are responsible for the comshy
putation of all the time-varying matrices vectors and scaler variables derived in
Chapter 2 that comprise the governing equations of motion for the entire system
These subroutines have been carefully constructed to ensure maximum efficiency in
computational performance as well as memory allocation
Two important aspects in the programming of F C N should be outlined at this
point As mentioned earlier some of the integrals defined in Chapter 2 are notably
more difficult to evaluate manually Traditionally they have been integrated numershy
ically on-line (during the simulation) In some cases particularly stationkeeping the
computational cost is quite reasonable since the modal integrals need only be evalushy
ated once However for the case of deployment or retrieval where the mode shape
functions vary they have to be integrated at each time-step This has considerable
repercussions on the total simulation time Thus in this program the integrals are
evaluated symbolically using M A P L E V[55] Once generated they are translated into
F O R T R A N source code and stored in I N C L U D E files to be read by the compiler
Al though these files can be lengthy (1000 lines or more) for a large number of flexible
44
INIT
DGEAR PARAM
IC amp SYS PARAM
MAIN
lt
DGEAR
lt
FCN
OUTPUT
STATE ACTUATOR FORCES
MOMENTS amp THRUST FORCES
VIBRATION
r OFFS DYNA
ET MICS
Figure 3-1 Flowchart showing the computer program structure
45
modes they st i l l require less time to compute compared to their on-line evaluation
For the case of deployment the performance improvement was found be as high as
10 times for certain cases
Secondly the O(N) algorithm presented in the last section involves matrices
that contained mostly zero terms (RP Rv and Mplusmn) The sparse structure of these mashy
trices can be used advantageously to improve the computational efficiency by avoiding
multiplications of zero elements The result is a quick evaluation of fv(Xt) Moreshy
over there is a considerable saving of computer memory as zero elements are no
longer stored
Final ly the C O N T R O L subroutine which is part of F C N is designed to calshy
culate the appropriate control forces torques and offset dynamics which are required
to stabilize the librational ( A T T I T U D E ) and vibrational ( V I B R A T I O N ) motion of
the system
33 Verification of the Code
The simulation program was checked for its validity using two methods The
first verified the conservation of total energy in the absence of dissipation The second
approach was a direct comparison of the planar response generated by the simulation
program with those available in the literature A s numerical results for the flexible
three-dimensional A^-body model are not available one is forced to be content with
the validation using a few simplified cases
331 Energy conservation
The configuration considered here is that of a 3-body tethered system with
the following parameters for the platform subsatellite and tether
46
h =
1091430 -8135 328108 -8135 8646050 27116 328108 27116 8286760
kg bull m 2 platform inertia
kg bull m 2 end-satellite inertia 200 0 0
0 400 0 0 0 400
m = 90000 kg mass of the space station platform
m2 = 500 kg mass of the end-satellite
EfAt = 61645 N tether elastic stiffness
pt mdash 49 kg km tether linear density
The tether length is taken to remain constant at 10 km (stationkeeping case)
A n in i t ia l disturbance in pitch and roll of 2deg and 1deg respectively is given to both
the rigid bodies and the flexible tether In addition the tether is ini t ia l ly deflected
by 45 m in the longitudinal direction at its end and 05 m in both the inplane and
out-of-plane directions in the first mode The tether attachment point offset at the
platform (satellite 1) end is 1 m in all the three directions ie d2 = 11 lTm The
attachment point at the subsatellite (satellite 2) end is 1 m in the x direction ie
fcm3 = 10 0T
The variation in the kinetic and potential energies is presented for the above-
mentioned case in Figure 3-2(a) As seen from the plot there is a continuous mutual
exchange between the kinetic and potential energies however the variation in the
total energy remains zero (Figure 3-2b) It should be noted that after 1 orbit the
variation of both kinetic and potential energy does not return to 0 This is due to
the attitude and elastic motion of the system which shifts the centre of mass of the
tethered system in a non-Keplerian orbit
Similar results are presented for the 5-body case ie a double pendulum (Figshy
ure 3-3) Here the system is composed of a platform and two subsatellites connected
47
STSS configuration 3-Body a 1(0)=a 2(0)=a t(0)=2deg P1(0)=p2(0)=p(0)=1deg 5x(0)=45m 8y(0)=52(0)=05m dx(0)=dy(0)=d2(0)=1m Stationkeeping lt=10km
Variation of Kinetic and Potential Energy
(a) Time (Orbits) Percent Variation in Total Energy
OOEOh
-50E-12
-10E-11 I i i
0 1 2 (b) Time (Orbits)
Figure 3-2 Kinetic and potential energy transfer for the three-body platform based tethered satellite system (a) variation of kinetic and potential energy (b) percent change in total energy of system
48
STSS configuration 5-Body a 1(0)=a 2(0)=a t 1(0)=2deg a 3(0)=a t 2(0)=25deg P1(0)=p2(0)=p3(0)=pt1(0)=pt2(0)=1lt
5x(0)=5m 8y(0)=5z(0)=05m dx(0)=dy(0)=dz(0)=0 Stationkeeping 1 =12=1 Okm
(a)
20E6
10E6r-
00E0
-10E6
-20E6
Variation of Kinetic and Potential Energy
-
1 I I I 1 I I
I I
AP e ~_
i i 1 i i i i
00E0
UJ -50E-12 U J
lt
Time (Orbits) Percent Variation in Total Energy
-10E-11h
(b) Time (Orbits)
Figure 3 -3 Kinet ic and potential energy transfer for the five-body platform based tethered satellite system (a) variation of kinetic and potential enshyergy (b) percent change in total energy of system
49
in sequence by two tethers The two subsatellites have the same mass and inertia
ie 7713 = 7 7 i 2 a n d I3 = I21 a n d are separated by identical tether segments each 10
k m in length The platform has the same properties as before however there is no
tether attachment point offset present in the system (d = fcmi = 0)
In al l the cases considered energy was found to be conserved However it
should be noted that for the cases where the tether structural damping deployment
retrieval attitude and vibration control are present energy is no longer conserved
since al l these features add or subtract energy from the system
332 Comparison with available data
The alternative approach used to validate the numerical program involves
comparison of system responses with those presented in the literature A sample case
considered here is the planar system whose parameters are the same as those given
in Section 331 A n ini t ia l pitch disturbance of 2deg is imparted to both the platform
and the tether together with an ini t ial tether deformation of 08 m and 001 m in
the longitudinal and transverse directions respectively in the same manner as before
The tether length is held constant at 5 km and it is assumed to be connected to the
centre of mass of the rigid platform and subsatellite The response time histories from
Ref[43] are compared with those obtained using the present program in Figures 3-4
and 3-5 Here ap and at represent the platform and tether pitch respectively and B i
C i are the tethers longitudinal and transverse generalized coordinates respectively
A comparison of Figures 3-4 and 3-5 clearly shows a remarkable correlation
between the two sets of results in both the amplitude and frequency The same
trend persisted even with several other comparisons Thus a considerable level of
confidence is provided in the simulation program as a tool to explore the dynamics
and control of flexible multibody tethered systems
50
a p(0) = 2deg 0(0) = 2C
6(0) = 08 m
0(0) = 001 m
Stationkeeping L = 5 km
D p y = 0 D p 2 = 0
a oo
c -ooo -0 01
Figure 3 - 4 Simulation results for the platform based three-body tethered system originally presented in Ref[43]
51
STSS configuration 3-Body a1(0)=a2(0)=at(0)=2deg p1(0)=(32(0)=(3t(0)=0 8x(0)=08m 5y(0)=001m 5Z(0)=0 dx(0)=dy(0)=dz(0)=0 Stationkeeping l=5km
Satellite 1 Pitch Angle Tether Pitch Angle
i -i i- i i i i i i d
1 2 0 1 2 Time (Orbits) Time (Orbits)
Longitudinal Vibration
^MAAAAAAAWVWWWWVVVW
V bull i bull i I i J 00 05 10 15 20
Time (Orbits) Transverse Vibration
Time (Orbits)
Figure 3-5 Simulation results for the platform based three-body tethered system obtained using the present computer program
52
4 D Y N A M I C SIMULATION
41 Prel iminary Remarks
Understanding the dynamics of a system is critical to its design and develshy
opment for engineering application Due to obvious limitations imposed by flight
tests and simulation of environmental effects in ground based facilities space based
systems are routinely designed through the use of numerical models This Chapter
studies the dynamical response of two different tethered systems during deployment
retrieval and stationkeeping phases In the first case the Space platform based Tethshy
ered Satellite System (STSS) which consists of a large mass (platform) connected to
a relatively smaller mass (subsatellite) with a long flexible tether is considered The
other system is the O E D I P U S B I C E P S configuration involving two comparable mass
satellites interconnected by a flexible tether In addition the mission requirement of
spin about an arbitrary axis the tether axis for O E D I P U S - A C and cartwheeling or
spin about the orbit normal for the proposed B I C E P S mission is accounted for
A s mentioned in Chapter 2 the tether flexibility is modelled using the asshy
sumed mode discretization method Although the simulation program can account
for an arbitrary number of vibrational modes in each direction only the first mode
is considered in this study as it accounts for most of the strain energy[34] and hence
dominates the vibratory motion The parametric study considers single as well as
double pendulum type systems with offset of the tether attachment points The
systems stability is also discussed
42 Parameter and Response Variable Definitions
The system parameters used in the simulation unless otherwise stated are
53
selected as follows for the STSS configuration
1091430 -8135 328108 -8135 8646050 27116 328108 27116 8286760
h =
h = 200 0 0
0 400 0 0 0 400
bull m = 90000 kg (mass of the space platform)
kg bull m 2 (platform inertia) 760 J
kg bull m 2 (subsatellite inertia)
bull 7772 = 500 kg (mass of the subsatellite)
bull EtAt = 61645 N (tether stiffness)
bull Pt = 49 k g k m (tether density)
bull Vd mdash 05 (tether structural damping coefficient)
bull fcm^ = l 0 0 r m (tether attachment point at the subsatellite)
The response variables are defined as follows
bull ai0 satellite 1 (platform) pitch and roll angles respectively
bull CK2 2 satellite 2 (subsatellite) pitch and roll angles respectively
bull at fit- tether pitch and roll angles respectively
bull If tether length
bull d = dx dy dzT- tether attachment position relative to satellite 1 (platform)
mdash
bull S = 8X 8y Sz tether flexible generalized coordinates in the longitudinal x
inplane transverse y and out-of-plane transverse z directions respectively
The attitude angles and are measured with respect to the Local Ve r t i c a l -
Loca l Horizontal ( L V L H ) frame A schematic diagram illustrating these variable
is presented in Figure 4-1 The system is taken to be in a nominal circular orbit at
an altitude of 289 km with an orbital period of 903 minutes
54
Satellite 1 (Platform) Y a w
Figure 4-1 Schematic diagram showing the generalized coordinates used to deshyscribe the system dynamics
55
43 Stationkeeping Profile
To facilitate comparison of the simulation results a reference case based on
the three-body STSS system with zero offset (d = 0 r c m 3 = 0) and a tether length
of It = 20 km is considered first (Figure 4-2) The system is subjected to an ini t ial
disturbance in pitch and roll of 2deg and 1deg respectively to all the three bodies
In addition an ini t ia l longitudinal deflection of 14 m from the tethers unstretched
position is given together with a transverse inplane and out-of-plane deflection of 1
m at the tethers mid-length From the attitude response given in Figure 4-2(a) it is
apparent that the three bodies oscillate about their respective equilibrium positions
Since coupling between the individual rigid-body dynamics is absent (zero offset) the
corresponding librational frequencies are unaffected Satellite 1 (Platform) displays
amplitude modulation arising from the products of inertia The result is a slight
increase of amplitude in the pitch direction and a significantly larger decrease in
amplitude in the roll angle
Figure 4-2(b) clearly shows coupling of the tethers attitude motion with its
flexibility dynamics The tether vibrates in the axial direction about its equilibrium
position (at approximately 138 m) with two vibration frequencies namely 012 Hz
and 31 x 1 0 - 4 Hz The former is the first longitudinal flexible mode which is dissishy
pated through the structural damping while the latter arises from the coupling with
the pitch motion of the tether Similarly in the transverse direction there are two
frequencies of oscillations arising from the flexible mode and its coupling wi th the
attitude motion As expected there is no apparent dissipation in the transverse flexshy
ible motion This is attributed to the weak coupling between the longitudinal and
transverse modes of vibration since the transverse strain in only a second order effect
relative to the longitudinal strain where the dissipation mechanism is in effect Thus
there is very litt le transfer of energy between the transverse and longitudinal modes
resulting in a very long decay time for the transverse vibrations
56
STSS configuration 3-Body a1(0)=a2(0)=a(0)=2deg (31(0)=p2(0)=p(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=dz(0)=0 Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
0 1 2 3 4 5 0 1 2 3 4 5 Time (Orbits) Time (Orbits)
Satellite 2 Pitch Angle Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-2 Stationkeeping dynamics of the three-body STSS configuration withshyout offset (a) attitude response
57
STSS configuration 3-Body a1(0)=cx2(0)=cx(0)=2deg P1(0)=p2(0)=p(0)=1deg 5x(0)=14m8y(0)=82(0)=1m dx(0)=dy(0)=dz(0)=0 Stationkeeping lt=20km
Tether Longitudinal Vibration 16| 1 1 1 1 1 r-
Time (Orbits) Tether In-Plane Transverse Vibration
imdash 1 mdash 1 mdash 1 mdash 1 mdashr
r I I 1 1 1 I 1 1 1 1 I 1 bull ^ bull bull I I J 0 1 2 3 4 5
Time (Orbits) Tether Out-of-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-2 Stationkeeping dynamics of the three-body STSS configuration withshyout offset (b) vibration response
58
Introduction of the attachment point offset significantly alters the response of
the attitude motion for the rigid bodies W i t h a i m offset along the local vertical
at the platform (dx = 1 m) and fcm^ = 100 m coupling between the tether
and end-bodies is established by providing a lever arm from which the tether is able
to exert a torque that affects the rotational motion of the rigid end-bodies (Figure
4-3) Bo th the platform and subsatellite now oscillate at the same frequency as the
tether whose motion remains unaltered In the case of the platform there is also an
amplitude modulation due to its non-zero products of inertia as explained before
However the elastic vibration response of the tether remains essentially unaffected
by the coupling
Providing an offset along the local horizontal direction (dy = 1 m) results in
a more dramatic effect on the pitch and roll response of the platform as shown in
Figure 4-4(a) Now the platform oscillates about its new equilibrium position of
-90deg The roll motion is also significantly disturbed resulting in an increase to over
10deg in amplitude On the other hand the rigid body dynamics of the tether and the
subsatellite as well as the tether flexible motion remain the same (Figure 4-4b)
Figure 4-5 presents the response when the offset of 1 m at the platform end is
along the z direction ie normal to the orbital plane The result is a larger amplitude
pitch oscillation about the reference equilibrium position as shown in Figure 4-5(a)
while the roll equilibrium is now shifted to approximately 90deg Note there is little
change in the attitude motion of the tether and end-satellites however there is a
noticeable change in the out-of-plane transverse vibration of the tether (Figure 4-5b)
which may be due to the large amplitude platform roll dynamics
Final ly by setting a i m offset in all the three direction simultaneously the
equil ibrium position of the platform in the pitch and roll angle is altered by approxishy
mately 30deg (Figure 4-6a) However the response of the system follows essentially the
59
STSS configuration 3-Body a1(0)=cx2(0)=cxt(0)=2deg p1(0)=(32(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=1m dy(0)=dz(0)=0 Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
Time (Orbits) Time (Orbits) Tether Pitch Angle Tether Roll Angle
L bull I I bull bull lt bull bull bull raquo bull bull bull bull -I h I I I I i H 0 1 2 3 4 5 0 1 2 3 4 5
Time (Orbits) Time (Orbits) Satellite 2 Pitch Angle Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-3 Stationkeeping dynamics of the three-body STSS configuration with offset along the local vertical (a) attitude response
60
STSS configuration 3-Body a 1(0)=a 2(0)=a t(0)=2deg (31(0)=p2(0)=pt(0)=1deg 5x(0)=14m5y(0)=5z(0)=1m dx(0)=1mdy(0)=dz(0)=0 Stationkeeping lt=20km
Tether Longitudinal Vibration 1 6 i mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash mdash i mdash i mdash mdash mdash mdash mdash i mdash mdash lt -
Time (Orbits)
Tether In-Plane Transverse Vibration
y bull i i i i i i i i i i i i i i i i i 0 1 2 3 4 5
Time (Orbits)
Tether Out-of-Plane Transverse Vibration r 1 1 1 1 1 1 1 1 1 1 1 1 1 1 q
h i I i i I i i i i I i i i i 3
0 1 2 3 4 5 (b) Time (Orbits)
Figure 4-3 Stationkeeping dynamics of the three-body STSS configuration with offset along the local vertical (b) vibration response
61
STSS configuration 3-Body a1(0)=cx2(0)=o(0)=2deg (31(0)=p2(0)=p(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dz(0)=0 dy(0)=1m Stationkeeping lt=20km
Satellite 1 Pitch Angle i 1 i
Satellite 1 Roll Angle
1 2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 4 Time (Orbits)
Tether Roll Angle
CD
edT
1 2 3 Time (Orbits)
Satellite 2 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 2 Roll Angle
CD
C O
(a)
Figure 4-
1 2 3 4 Time (Orbits)
1 2 3 4 Time (Orbits)
4 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal (a) attitude response
62
STSS configuration 3-Body a1(0)=cx2(0)=ot(0)=2deg p1(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dz(0)=0 dy(0)=1m Stationkeeping lt=20km
Tether Longitudinal Vibration 16| 1 gt 1 bull 1 1 1
0 1 2 3 4 5
Time (Orbits) Tether In-Plane Transverse Vibration
i mdash 1 mdash 1 mdash 1 mdash 1 mdash i mdash mdash lt mdash 1 mdash 1 mdash i mdash mdash 1 mdash 1 mdash
V I I bull bull bull I i i i I i i 1 1 3
0 1 2 3 4 5 Time (Orbits)
Tether Out-of-Plane Transverse Vibration r 1 1 1 1 1 1 1 1 1 1 1 gt bull | q
V i 1 1 1 1 1 mdash 1 mdash 3
0 1 2 3 4 5 (b) Time (Orbits)
Figure 4-4 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal (b) vibration response
63
STSS configuration 3-Body a1(0)=o2(0)=at(0)=2deg (31(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=0 dz(0)=1m Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 4 Time (Orbits)
Tether Roll Angle
1 2 3 4 Time (Orbits)
Satellite 2 Pitch Angle
2 3 4 Time (Orbits)
Satellite 2 Roll Angle
(a)
Figure 4-
1 2 3 4 Time (Orbits)
2 3 4 Time (Orbits)
5 Stationkeeping dynamics of the three-body STSS configuration with offset along the orbit normal (a) attitude response
64
STSS configuration 3-Body a1(0)=cx2(0)=ot(0)=2deg P1(0)=p2(0)=(3(0)=1deg 8x(0)=14m8y(0)=82(0)=1m dx(0)=dy(0)=0 d2(0)=1m Stationkeeping lt=20km
Tether Longitudinal Vibration 1 6 i mdash r mdash i mdash mdash i mdash i mdash mdash i mdash i mdash mdash i mdash mdash i mdash 1 mdash 1 mdash i mdash mdash lt mdash lt mdash lt mdash r
Time (Orbits) Tether In-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-5 Stationkeeping dynamics of the three-body STSS configuration with offset along the orbit normal (b) vibration response
65
same trend with only minor perturbations in the transverse elastic vibratory response
of the tether (Figure 4-6b)
44 Tether Deployment
Deployment of the tether from an ini t ial length of 200 m to 20 km is explored
next Here the deployment length profile is critical to ensure the systems stability
It is not desirable to deploy the tether too quickly since that can render the tether
slack Hence the deployment strategy detailed in Section 216 is adopted The tether
is deployed over 35 orbits with the sinusoidal acceleration and deceleration occurring
over a 4 km length ( A ^ 2 ) with the remaining deployment at a constant velocity of
V0 = 146 ms
Initially the tether is stretched only slightly S = 1304 x 1 0 _ 3 0 0 m
with al l the other states of the system ie pitch roll and transverse displacements
remaining zero The response for the zero offset case is presented in Figure 4-7 For
the platform there is no longer any coupling with the tether hence it oscillates about
the equilibrium orientation determined by its inertia matrix However there is st i l l
coupling between the subsatellite and the tether since rcm^ mdash 100 m resulting
in complete domination of the subsatellite dynamics by the tether Consequently
the pitch motion ini t ial ly grows by almost 50deg but eventually subsides as the tether
deployment rate decreases It is interesting to note that the out-of-plane motion is not
affected by the deployment This is because the Coriolis force which is responsible
for the tether pitch motion does not have a component in the z direction (as it does
in the y direction) since it is always perpendicular to the orbital rotation (Q) and
the deployment rate (It) ie Q x It
Similarly there is an increase in the amplitude of the transverse elastic oscilshy
lations of the tether again due to the Coriolis effect (Figure 4-7b) Moreover as the
66
STSS configuration 3-Body a1(0)=a2(0)=cxt(0)=2deg P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m5y(0)=52(0)=1m dx(0)=dy(0)=d2(0)=1m Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
i i i i i i i i i i i 0 1 2 3 4 5 0 1 2 3 4 5
Time (Orbits) Time (Orbits) Tether Pitch Angle Tether Roll Angle
l i i i i -I h i i i i I i i i i l- i i i i l i i i i I i i- i i H 0 1 2 3 4 5 0 1 2 3 4 5
Time (Orbits) Time (Orbits) Satellite 2 Pitch Angle Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-6 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal local vertical and orbit normal (a) attitude response
67
STSS configuration 3-Body a1(0)=cc2(0)=a(0)=2deg p1(0)=p2(0)=(3t(0)=1deg 5x(0)=14m6y(0)=52(0)=1m dx(0)=dy(0)=d2(0)=1m Stationkeeping lt=20km
Tether Longitudinal Vibration 1 6 1 mdash i mdash i mdash i mdash mdash i mdash 1 mdash 1 mdash mdash mdash i mdash mdash 1 mdash 1 mdash 1 mdash i mdash 1 mdash 1 mdash 1 mdash 1 mdash r
0 1 2 3 4 5
Time (Orbits) Tether In-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-6 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal local vertical and orbit normal (b) vibration response
68
STSS configuration 3-Body a(0)=a2(0)=ot(0)=0 P1(0)=p2(0)=pt(0)=0 8X(0)=1304 x 103m 8y(0)=8z(0)=0 dx(0)=dy(0)=dz(0)=0 Deployment
l=02km to 20km in 35 Orbits
Tether Length Profile
E
Satellite 1 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 1 Roll Angle mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash [ mdash i mdash i mdash i mdash i mdash | mdash
1 2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 4 Time (Orbits)
Tether Roll Angle
CD
oo ax
i imdashimdashimdashImdashimdashimdashimdashr-
_ 1 _ L
1 2 3 4 Time (Orbits)
Satellite 2 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 2 Roll Angle
-50
(a) 1 2 3 4
Time (Orbits) 1 2 3 4
Time (Orbits)
Figure 4-7 Deployment dynamics of the three-body STSS configuration without offset (a) attitude response
69
STSS configuration 3-Body a1(0)=a2(0)=at(0)=0 p1(0)=p2(0)=(3t(0)=0 8X(0)=1304x 103m 8y(0)=52(0)=0 dx(0)=dy(0)=d2(0)=0 Deployment
l=02km to 20km in 35 Orbits
Tether Longitudinal Vibration
to
-02
(b)
Figure 4-7
2 3 Time (Orbits)
Tether In-Plane Transverse Vibration
0 1 1 bull 1 1
1 2 3 4 Time (Orbits)
Tether Out-of-Plane Transverse Vibration
5
1 1 1 1 1 i i | i i | ^
_ i i i i L -I
1 2 3 4 5 Time (Orbits)
Deployment dynamics of the three-body STSS configuration without offset (b) vibration response
70
tether elongates its longitudinal static equilibrium position also changes due to an
increase in the gravity-gradient tether tension It may be pointed out that as in the
case of attitude motion there are no out-of-plane tether vibrations induced
When the same deployment conditions are applied to the case of 1 m offset in
the x direction (Figure 4-8) coupling effects are re-introduced The platform pitch
exceeds -100deg during the peak deployment acceleration However as the tether pitch
motion subsides the platform pitch response returns to a smaller amplitude oscillation
about its nominal equilibrium On the other hand the platform roll motion grows
to over 20deg in amplitude in the terminal stages of deployment Longitudinal and
transverse vibrations of the tether remain virtually unchanged from the zero offset
case except for minute out-of-plane perturbations due to the platform librations in
roll
45 Tether Retrieval
The system response for the case of the tether retrieval from 20 km to 200 m
with an offset of 1 m in the x and z directions at the platform end is presented in
Figure 4-9 Here the Coriolis force induced during retrieval renders the tether unstable
(Figure 4-9a) Consequently the pitch motion of the platform is also destabilized
through coupling The z offset provides an additional moment arm that shifts the
roll equilibrium to 35deg however this degree of freedom is not destabilized From
Figure 4-9(b) it is apparent that the transverse mode of the tether vibration is also
disturbed producing a high frequency response in the final stages of retrieval This
disturbance is responsible for the slight increase in the roll motion for both the tether
and subsatellite
46 Five-Body Tethered System
A five-body chain link system is considered next To facilitate comparison with
71
STSS configuration 3-Body a(0)=a2(0)=at(0)=0 P1(0)=P2(0)=P(0)=0 5X(0)=1304 x 103m 8y(0)=8z(0)=0 dx(0)=1mdy(0)=d2(0)=0 Deployment
lt=02km to 20km in 35 Orbits
Tether Length Profile
E
Satellite 1 Pitch Angle
1 2 3 Time (Orbits)
Satellite 1 Roll Angle
bull bull bull i bull bull bull bull i
1 2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 Time (Orbits)
Tether Roll Angle
i i i i J I I i _
1 2 3 4 Time (Orbits)
Satellite 2 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 2 Roll Angle 3E-3h
D )
T J ^ 0 E 0 eg
CQ
-3E-3
(a) 1 2 3 4
Time (Orbits) 1 2 3 4
Time (Orbits)
Figure 4-8 Deployment dynamics of the three-body STSS configuration with offset along the local vertical (a) attitude response
7 2
STSS configuration 3-Body a1(0)=a2(0)=at(0)=0 p1(0)=P2(0)=(3t(0)=0 Sx(0)=1304 x 103m 5y(0)=5z(0)=0 dx(0)=1mdy(0)=dz(0)=0 Deployment
l=02km to 20km in 35 Orbits
Tether Longitudinal Vibration
Time (Orbits) Tether In-Plane Transverse Vibration
n 1 1 1 f
I I I _j I I I I I 1 1 1 1 1 2 3 4 5
Time (Orbits) Tether Out-of-Plane Transverse Vibration
J 1 I I 1 I I I I I I I I L _ _ I I d 1 2 3 4 5
Time (Orbits)
Deployment dynamics of the three-body STSS configuration with offset along the local vertical (b) vibration response
(b)
Figure 4-8
73
STSS configuration 3-Body a1(0)=a2(0)=at(0)=0 p(0)=P2(0)=(3t(0)=0 Sx(0)=14m8y(0)=82(0)=0 dy(0)=0dx(0)=d2(0)=1m Retrieval
l=20km to 02km in 35 Orbits
Satellite 1 Pitch Angle
(a) Time (Orbits)
Figure 4-9 Retrieval dynamics of the along the local vertical ar
Tether Length Profile
i i i i i i i i i i I 0 1 2 3
Time (Orbits) Satellite 1 Roll Angle
Time (Orbits) Tether Roll Angle
b_ i i I i I i i L _ J
0 1 2 3 Time (Orbits)
Satellite 2 Roll Angle _ mdash | r i i | 1 i i imdashmdashj i 1 r mdash i mdash
04- -
0 1 2 3 Time (Orbits)
three-body STSS configuration with offset d orbit normal (a) attitude response
74
STSS configuration 3-Body a1(0)=a2(0)=ocl(0)=0 (31(0)=f32(0)=pt(0)=0 8x(0)=14m8y(0)=6z(0)=0 dy(0)=0dx(0)=dz(0)=1m Retrieval
l=20km to 02km in 35 Orbits
Tether Longitudinal Vibration
E 10 to
Time (Orbits) Tether In-Plane Transverse Vibration
1 2 Time (Orbits)
Tether Out-of-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-9 Retrieval dynamics of the three-body STSS configuration with offset along the local vertical and orbit normal (b) vibration response
75
the three-body system dynamics studied earlier the chain is extended through the
addition of two bodies a tether and a subsatellite with the same physical properties
as before (Section 42) Thus the five-body system consists of a platform (satellite 1)
tether 1 subsatellite 1 (satellite 2) tether 2 and subsatellite 2 (satellite 3) as shown
in Figure 4-10 The offset of tether 1 at the platform-end is denoted by d2 while that
of tether 2 to subsatellite 1 as 04 In this numerical example fcm^ = fcm^ = 0 ie
tethers 1 and 2 are attached to the centre of mass of subsatellites 1 and 2 respectively
The system response for the case where the tether attachment points coincide
wi th the centres of mass of the rigid bodies ie zero offset (d2 = d mdash 0) is presented
in Figure 4-11 Here ogtt and at2 represent the pitch motion of tethers 1 and 2
respectively whereas 03 is the pitch motions of satellite 3 A similar convention
is adopted for the rol l angle 0 A s before the zero offset eliminates the coupling
between the tethers and satellites such that they are now free to librate about their
static equil ibrium positions However their is s t i l l mutual coupling between the two
tethers This coupling is present regardless of the offset position since each tether is
capable of transferring a force to the other The coupling is clearly evident in the pitch
response of the two tethers (Figure 4 - l l a ) O n the other hand the roll motion appears
uncoupled as the in i t ia l conditions in rol l for the two tethers are identical Thus the
motion is in phase and there is no transfer of energy through coupling A s expected
during the elastic response the tethers vibrate about their static equil ibrium positions
and mutually interact (Figure 4 - l l b ) However due to the variation of tension along
the tether (x direction) they do not have the same longitudinal equilibrium Note
relatively large amplitude transverse vibrations are present particularly for tether 1
suggesting strong coupling effects with the longitudinal oscillations
76
Figure 4-10 Schematic diagram of the five-body system used in the numerical example
77
STSS configuration 5-Body a1(0)=(x2(0)=at1(0)=2deg a3(0)=a12(0)=25deg p1(0)=r32(0)=p3(0)=pt1(0)=pt2(0)=r 82(0)=84(0)=505)05m d2(0)=d4(0)=000 Stationkeeping lt1=lu=10km Pitch Roll
Satellite 1 Pitch Angle Satellite 1 Roll Angle
o 1 2 Time (Orbits)
Tether 1 Pitch and Roll Angle
0 1 2 Time (Orbits)
Tether 2 Pitch and Roll Angle
Time (Orbits) Satellite 2 Pitch and Roll Angle
Time (Orbits) Satellite 3 Pitch and Roll Angle
Time (Orbits) Time (Orbits)
Figure 4-11 Stationkeeping dynamics of the five-body STSS configuration withshyout offset (a) attitude response
78
STSS configuration 5-Body a1(0)=a2(0)=oct1(0)=2deg a3(0)=at2(0)=25o
p1(0)=p2(0)=(33(0)=pt1(0)=pt2(0)=10
62(0)=54(0)=50505m d2(0)=d4(0)=000 Stationkeeping lt1=1 =1 Okm
Tether 1 Longitudinal Vibration Tether 2 Longitudinal Vibration i i 1 i 1 r 1 1 1 1
(b) Time (Orbits) Time (Orbits)
Figure 4-11 Stationkeeping dynamics of the five-body STSS configuration withshyout offset (b) vibration response
79
The system dynamics during deployment of both the tethers in the double-
pendulum configuration is presented in Figure 4-12 Deployment of each tether takes
place from 200 m to 20 km in 35 orbits The pitch motion of the system is similar
to that observed during the three-body case The Coriolis effect causes the tethers
to pitch through a large angle ( laquo 50deg ) which in turn disturbs the platform and
subsatellite due to the presence of offset along the local vertical On the other hand
the roll response for both the tethers is damped to zero as the tether deploys in
confirmation with the principle of conservation of angular momentum whereas the
platform response in roll increases to around 20deg Subsatellite 2 remains virtually
unaffected by the other links since the tether is attached to its centre of mass thus
eliminating coupling Finally the flexibility response of the two tethers presented in
Figure 4-12(b) shows similarity with the three-body case
47 BICEPS Configuration
The mission profile of the Bl-static Canadian Experiment on Plasmas in Space
(BICEPS) is simulated next It is presently under consideration by the Canadian
Space Agency To be launched by the Black Brant 2000500 rocket it would inshy
volve interesting maneuvers of the payload before it acquires the final operational
configuration As shown in Figure 1-3 at launch the payload (two satellites) with
an undeployed tether is spinning about the orbit normal (phase 1) The internal
dampers next change the motion to a flat spin (phase 2) which in turn provides
momentum for deployment (phase 3) When fully deployed the one kilometre long
tether will connect two identical satellites carrying instrumentation video cameras
and transmitters as payload
The system parameters for the BICEPS configuration are summarized below [59 0 0 1
bull I = I2 = 0 366 0 kg bull m 2 (satellite inertias) [ 0 0 392_
bull mi = 7772 = 200 kg (mass of the satellites)
80
STSS configuration 5-Body a(0)=cx2(0)=al1(0)=2deg a3(0)=c (0)=25o
P1(0)=P2(0)=p3(0)=f3t1(0)=Pt2(0)=r 52(0)=84(0)=1304x 10300m d2(0)=d4(0)=1G0m Deployment lt1=le=02km to 20km in 35 Orbits
Pitch Roll Satellite 1 Pitch Angle
Time (Orbits) Satellite 2 Pitch and Roll Angle T 1mdash1 1 1 I I I I | I I I 1 I | I I lt~r
Tether Length Profile i 1 1 1 1 i 1 1 1 1 i 1 1 1 1 i
i i i i i 0 1 2 3 4 5
Time (Orbits) Satellite 1 Roll Angle
0 1 2 3 4 5 Time (Orbits)
Tether 1 amp 2 Roll Angle L i | I I 1 I 1 1 -I
-Ih I- bull I I i i i i I i i i i 1
0 1 2 3 4 5 Time (Orbits)
Satellite 3 Pitch and Roll Angle i- i 1 1 i 1 1 1 1 i 1 1 1 1 i 1 1 1 1 i
L i i i i i i 1 i i mdash i mdash i I i imdashimdashLJ
0 1 2 3 4 5 Time (Orbits)
Figure 4-12 Deployment dynamics of the five-body S T S S configuration with offshyset along the local vertical (a) attitude response
81
STSS configuration 5-Body a1(0)=a2(0)=oct1(0)=2deg a3(0)=at2(0)=25deg P1(0)=P2(0)=P3(0)=Pt1(0)=pt2(0)=1o
82(0)=54(0)=1304 x 10-300m d2(0)=d4(0)=100m Deployment lt1=l2=02km to 20km in 35 Orbits
Tether 1 Longitudinal Vibration Tether 2 Longitudinal Vibration
Time (Orbits) Time (Orbits) Tether 1 Z Tranverse Vibration Tether 2 Z Tranverse Vibration
Figure 4-12 Deployment dynamics of the five-body STSS configuration with offshyset along the local vertical (b) vibration response
82
bull EtAt = 61645 N (tether stiffness)
bull pt = 30 k g k m (tether density)
bull It = 1 km (tether length)
bull rid = 1-0 (tether structural damping coefficient)
The system is in a circular orbit at a 289 km altitude Offset of the tether
attachment point to the satellites at both ends is taken to be 078 m in the x
direction The response of the system in the stationkeeping phase for a prescribed set
of in i t ia l conditions is given by Figure 4-13 As in the case of the STSS configuration
there is strong coupling between the tether and the satellites rigid body dynamics
causing the latter to follow the attitude of the tether However because of the smaller
inertias of the payloads there are noticeable high frequency modulations arising from
the tether flexibility Response of the tether in the elastic degrees of freedom is
presented in Figure 4-13(b) Note the longitudinal vibrations decay quite rapidly
(lt 05 orbits) due to the structural damping however it has vir tual ly no effect on
the transverse oscillations
The unique mission requirement of B I C E P S is the proposed use of its ini t ia l
angular momentum in the cartwheeling mode to aid in the deployment of the tethered
system The maneuver is considered next The response of the system during this
maneuver wi th an ini t ia l cartwheeling rate of 5 deg s is illustrated in Figure 4-14(a) The
in i t ia l tether length is taken to be 10 m and is deployed to 1 km There is an in i t ia l
increase in the pitch motion of the tether however due to the conservation of angular
momentum the cartwheeling rate decreases proportionally to the square of the change
in tether deployment rate The result is a rapid drop in the cartwheeling rate unti l
the system stops rotating entirely and simply oscillates about its new equilibrium
point Consequently through coupling the end-bodies follow the same trend The
roll response also subsides once the cartwheeling motion ceases However
83
BICEPS configuration 3-Body a1(0)=o2(0)=at(0)=2deg P1(0)=P2(0)=pt(0)=1deg 8X(0)=001 m 6y(0)=52(0)=01 m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping l=1 km
Satellite 1 Pitch Angle Satellite 1 Roll Angle - i 1 1 1 1 i - i q r 1 1 -gt
I- I i t i i i i I H 0 1 2 0 1 2
Time (Orbits) Time (Orbits) Tether Pitch Angle Tether Roll Angle
F mdash 1 mdash 1 mdash mdash 1 mdash i mdash 1 mdash 1 mdash mdash 1 mdash = i F mdash 1 mdash mdash mdash 1 mdash i mdash 1 mdash 1 mdash 1 mdash 1 mdash q
(a) Time (Orbits) Time (Orbits)
Figure 4-13 Stationkeeping dynamics of the three-body BICEPS configuration with offset along the local vertical (a) attitude response
84
BICEPS configuration 3-Body a1(0)=o2(0)=at(0)=2deg Pl(0)=p2(0)=(3t(0)=1deg 8X(0)=001 m 8y(0)=82(0)=01 m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping lt=1km
Tether Longitudinal Vibration 1E-2[
laquo = 5 E - 3 |
000 002
Time (Orbits) Tether In-Plane Transverse Vibration
Time (Orbits) Tether Out-of-Plane Transverse Vibration
Time (Orbits)
Figure 4-13 Stationkeeping dynamics of the three-body BICEPS configuration with offset along the local vertical (b) vibration response
85
BICEPS configuration 3-Body a1(0)=o2(0)=at(0)=2deg a1(0)=a2(0)=at(0)=57s Pl(0)=P2(0)=(3t(0)=1deg 8X(0)=115 x 103m 8y(0)=82(0)=0 dx(0)=078m dy(0)=d2(0)=0 Cartwheeling Deployment
lt=10m to 1km in 1 Orbit
Tether Length Profile
E
2000
cu T3
2000
co CD
2000
C D CD
T J
Satellite 1 Pitch Angle
1 2 Time (Orbits)
Satellite 1 Roll Angle
cn CD
33 cdl
Time (Orbits) Tether Pitch Angle
Time (Orbits) Tether Roll Angle
Time (Orbits) Satellite 2 Pitch Angle
1 2 Time (Orbits)
Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-14 Cartwheeling dynamics with deployment for the three-body BICEPS with offset along the local vertical (a) attitude response
86
BICEPS configuration 3-Body a1(0)=a2(0)=at(0)=2deg a (0)=a2(0)=a(0)=57s P1(0)=P2(0)=f3(0)=1deg 8X(0)=115x103m 8y(0)=8z(0)=0 dx(0)=078m dy(0)=d2(0)=0 Cartwheeling Deployment
l=1 Om to 1 km in 1 Orbit
5E-3h
co
0E0
025
000 E to
-025
002
-002
(b)
Figure 4-14
Tether Longitudinal Vibration
Time (Orbits) Tether In-Plane Transverse Vibration
Time (Orbits) Tether Out-of-Plane Transverse Vibration
Time (Orbits)
Cartwheeling dynamics with deployment for the three-body BICEPS with offset along the local vertical (b) vibration response
87
as the deployment maneuver is completed the satellites continue to exhibit a small
amplitude low frequency roll response with small period perturbations induced by
the tether elastic oscillations superposed on it (Figure 4-14b)
48 OEDIPUS Spinning Configuration
The mission entitled Observation of Electrified Distributions of Ionospheric
Plasmas - A Unique Strategy ( O E D I P U S ) also consists of a 1 k m tethered system with
two similar end-masses It was first flown in a suborbital flight in 1989 ( O E D I P U S A )
followed by a recent study in November 1995 ( O E D I P U S - C ) in which the University
of Br i t i sh Columbia was one of the participants Dynamically O E D I P U S represents a
unique system never encountered before Launched by a Black Brant rocket developed
at Br is to l Aerospace L t d the system ie the end-bodies together wi th the tether
spins about the longitudinal axis of the tether to achieve stabilized alignment wi th
Earth s magnetic field
The response of the system undergoing a spin rate of j = l deg s is presented
in Figure 4-15 using the same parameters as those outlined for the B I C E P S conshy
figuration in Section 47 Again there is strong coupling between the three bodies
(two satellites and tether) due to the nonzero offset The spin motion introduces an
additional frequency component in the tethers elastic response which is transferred
to the l ibrational motion of the satellites A s the spin rate increases to 1 0 deg s (Figure
4-16) the amplitude and frequency of the perturbations also increase however the
general character of the response remains essentially the same
88
OEDIPUS configuration 3-Body a1(0)=a2(0)=at(0)=2deg i(0)^[2(0H(0)=Ys p1(0)=p2(0)=p(0)=r 8X(0)=12 x 103m 8y(0)=82(0)=01m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping l=1 km
Satellite 1 Pitch Angle Satellite 1 Roll Angle mdash i 1 1 1 a r 1 1 1 r
(a) Time (Orbits) Time (Orbits)
Figure 4-15 Spin dynamics (7 = ldegs) of the three-body OEDIPUS configuration with offset along the local vertical (a) attitude response
89
OEDIPUS configuration 3-Body a1(0)=a2(0)=cct(0)=2deg Y1(0)=Y2(0)=Yt(0)=l7s (31(0)=p2(0)=pt(0)=1deg 8X(0)=12 x 103m 8y(0)=5z(0)=01 m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping lt=1 km
1E-2
to
OEO
E 00
to
Tether Longitudinal Vibration
005
Time (Orbits) Tether In-Plane Transverse Vibration
Time (Orbits) Tether Out-of-Plane Transverse Vibration
Time (Orbits)
Figure 4-15 Spin dynamics (7 = l deg s ) of the three-body O E D I P U S configuration with offset along the local vertical (b) vibration response
90
OEDIPUS configuration 3-Body a1(0)=a2(0)=a(0)=2deg Yi(0)=Y2(0)=rt(0)=107s 81(0)=p2(0)=pt(0)=1deg 8X(0)=12 x 103m 8y(0)=82(0)=01m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping lt=1km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-16 Spin dynamics (7= 10degs) of the three-body OEDIPUS configura tion with offset along the local vertical (a) attitude response
91
OEDIPUS configuration 3-Body a1(0)=a2(0)=ot(0)=2deg Yi(0)H(0)4(0)laquo107s P1(0)=P2(0)=pt(0)=1deg 8X(0)=12 x 103m 8y(0)=82(0)=01m dx(0)=078m dy(0)=dz(0)=0 Stationkeeping lt=1 km
Tether Longitudinal Vibration
Figure 4-16 Spin dynamics (7 = 10degs) of the three-body OEDIPUS configurashytion with offset along the local vertical (b) vibration response
92
5 ATTITUDE AND VIBRATION CONTROL
51 Att itude Control
511 Preliminary remarks
The instability in the pitch and roll motions during the retrieval of the tether
the large amplitude librations during its deployment and the marginal stability during
the stationkeeping phase suggest that some form of active control is necessary to
satisfy the mission requirements This section focuses on the design of an attitude
controller with the objective to regulate the librational dynamics of the system
As discussed in Chapter 1 a number of methods are available to accomplish
this objective This includes a wide variety of linear and nonlinear control strategies
which can be applied in conjuction with tension thrusters offset momentum-wheels
etc and their hybrid combinations One may consider a finite number of Linear T ime
Invariant (LTI) controllers scheduled discretely over different system configurations
ie gain scheduling[43] This would be relevant during deployment and retrieval of
the tether where the configuration is changing with time A n alternative may be
to use a Linear T ime Varying ( L T V ) model in conjuction with an adaptive control
scheme where on-line parametric identification may be used to advantage[56] The
options are vir tually limitless
O f course the choice of control algorithms is governed by several important
factors effective for time-varying configurations computationally efficient for realshy
time implementation and simple in character Here the thrustermomentum-wheel
system in conjuction with the Feedback Linearization Technique ( F L T ) is chosen as
it accounts for the complete nonlinear dynamics of the system and promises to have
good overall performance over a wide range of tether lengths It is well suited for
93
highly time-varying systems whose dynamics can be modelled accurately as in the
present case
The proposed control method utilizes the thrusters located on each rigid satelshy
lite excluding the first one (platform) to regulate the pitch and rol l motion of the
tether to which it is attached ie the thrusters located on satellite 2 (link 3) regshy
ulates the attitude motion of tether 1 (link 2) O n the other hand the l ibrational
motion of the rigid bodies (platform and subsatellite) is controlled using a set of three
momentum wheels placed mutually perpendicular to each other
The F L T method is based on the transformation of the nonlinear time-varying
governing equations into a L T I system using a nonlinear time-varying feedback Deshy
tailed mathematical background to the method and associated design procedure are
discussed by several authors[435758] The resulting L T I system can be regulated
using any of the numerous linear control algorithms available in the literature In the
present study a simple P D controller is adopted arid is found to have good perforshy
mance
The choice of an F L T control scheme satisfies one of the criteria mentioned
earlier namely valid over a wide range of tether lengths The question of compushy
tational efficiency of the method wi l l have to be addressed In order to implement
this controller in real-time the computation of the systems inverse dynamics must
be executed quickly Hence a simpler model that performs well is desirable Here
the model based on the rigid system with nonlinear equations of motion is chosen
Of course its validity is assessed using the original nonlinear system that accounts
for the tether flexibility
512 Controller design using Feedback Linearization Technique
The control model used is based on the rigid system with the governing equa-
94
tions of motion given by
Mrqr + fr = Quu (51)
where the left hand side represents inertia and other forces while the right hand side
is the generalized external force due to thrusters and momentum-wheels and is given
by Eq(2109)
Lett ing
~fr = fr~ QuU (52)
and substituting in Eq(296)
ir = s(frMrdc) (53)
Expanding Eq(53) it can be shown that
qr = S(fr Mrdc) - S(QU Mr6)u (5-4)
mdash = Fr - Qru
where S(QU Mr0) is the column matrix given by
S(QuMr6) = [s(Qu(l)Mr0) 5 (Q u ( 2 ) M r | 0 ) S(QU nu) M r | 0 ) ] (55)
and Qu-i) is the i ^ column of Qu Extract ing only the controlled equations from
Eq(54) ie the attitude equations one has
Ire = Frc ~~ Qrcu
(56)
mdash vrci
where vrc is the new control input required to regulate the decoupled linear system
A t this point a simple P D controller can be applied ie
Vrc = qrcd +KvQrcd-Qrc) + Kp(Qrcd~ Qrc)- (57)
Eq(57) represents the secondary controller with Eq(54) as the primary controller
Kp and Kv are the proportional and derivative gain matrices with qrcdi Qrcd
an(^ qrcd
95
as the desired acceleration velocity and position vector for the attitude angles of each
controlled body respectively Solving for u from Eq(56) leads to
u Qrc Frc ~ vrcj bull (58)
5 13 Simulation results
The F L T controller is implemented on the three-body STSS tethered system
The choice of proportional and derivative matrix gains is based on a desired settling
time of ts mdash 05 orbit and a damping factor of = 07 for both the pitch and roll
actuators in each body Given O and ts it can be shown that
-In ) 0 5 V T ^ r (59)
and hence kn bull mdash10
V (510)
where kp and kVj are the diagonal elements of the gain matrices Kp and Kv in
Eq(57) respectively[59]
Note as each body-fixed frame is referred directly to the inertial frame FQ
the nominal pitch equilibrium angle is zero only when measured from the L V L H frame
(OJJ mdash 0) However it is 6 when referred to FQ Hence the desired position vector
qrC(l is set equal to 0(t) ie the time-varying orbital angle for the pitch motion
and zero for the roll and yaw rotations such that
qrcd 37Vxl (511)
96
The desired velocity and acceleration are then given as
e U 3 N x l (512)
E K 3 7 V x l (513)
When the system is in a circular orbit qrc^ = 0
Figure 5-1 presents the controlled response of the STSS system defined in
Chapter 4 in the stationkeeping phase with offset d2 = 111T m and If = 20
km A s mentioned earlier the F L T controller is based on the nonlinear rigid model
Note the pitch and roll angles of the rigid bodies as well as of the tether are now
attenuated in less than 1 orbit (Figure 5-la) Consequently the longitudinal vibration
response of the tether is free of coupling arising from the librational modes of the
tether This leaves only the vibration modes which are eventually damped through
structural damping The pitch and roll dynamics of the platform require relatively
higher control moments Ma and Mpi respectively (Figure 5-lb) since they have to
overcome the extra moments generated by the offset of the tether attachment point
Furthermore the platform demands an additional moment to maintain its orientation
at zero pitch and roll angle since it is not the nominal equilibrium position The nonshy
zero static equilibrium of the platform arises due to its products of inertia On the
other hand the tether pitch and roll dynamics require only a small ini t ia l control
force of about plusmn 1 N which eventually diminishes to zero once the system is
97
and
o 0
Qrcj mdash
o o
0 0
Oiit) 0 0
1
N
N
1deg I o
STSS configuration 3-Body a1(0)=a2(0)=at(0)=2lt
P1(0)=p2(0)=pt(0)=1 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=dz(0)=1m Controlled Stationkeeping lt=20km
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
05 10 15 Time (Orbits)
Sat 2 Pitch and Roll Angles
20
CO CD
T J
cdT
cT
- mdash mdash 1 1 1 1 1 I - 1
Pitch -
A Roll -
A -
1 bull bull
-
1 Time (Orbits)
Tether Y Tranverse Vibration
1 2 Time (Orbits)
Tether Z Tranverse Vibration
gt to N
to
Time (Orbits) Time (Orbits) Longitudinal Vibration
145h
(a) Time (Orbits)
Figure 5-1 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local horizontal local vertical and orbit normal (a) attitude and vibration response
98
1
STSS configuration 3-Body a1(0)=ct2(0)=cct(0)=20
P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=d2(0)=1m Controlled Stationkeeping lt=20km
Sat 1 Pitch Control Moment Sat 1 Roll Control Moment
Figure 5-1 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local horizontal local vertical and orbit normal (b) control actuator time histories
99
stabilized Similarly for the end-body a very small control torque is required to
attenuate the pitch and roll response
If the complete nonlinear flexible dynamics model is used in the feedback loop
the response performance was found to be virtually identical with minor differences
as shown in Figure 5-2 For example now the pitch and roll response of the platform
is exactly damped to zero with no steady-state error as against the minor almost
imperceptible deviation for the case of the controller basedon the rigid model (Figure
5-la) In addition the rigid controller introduces additional damping in the transverse
mode of vibration where none is present when the full flexible controller is used
This is due to a high frequency component in the at motion that slowly decays the
transverse motion through coupling
As expected the full nonlinear flexible controller which now accounts the
elastic degrees of freedom in the model introduces larger fluctuations in the control
requirement for each actuator except at the subsatellite which is not coupled to the
tether since f c m 3 = 0 The high frequency variations in the pitch and roll control
moments at the platform are due to longitudinal oscillations of the tether and the
associated changes in the tension Despite neglecting the flexible terms the overall
controlled performance of the system remains quite good Hence the feedback of the
flexible motion is not considered in subsequent analysis
When the tether offset at the platform is restricted to only 1 m along the x
direction a similar response is obtained for the system However in this case the
steady-state error in the platforms rigid body motion is much smaller (Figure 5-3a)
In addition from Figure 5-3(b) the platforms control requirement is significantly
reduced
Figure 5-4 presents the controlled response of the STSS deploying a tether
100
STSS configuration 3-Body a1(0)=a2(0)=oct(0)=2deg p1(0)=P2(0)=p(0)=1deg 8x(0)=14m5y(0)=82(0)=1m dx(0)=dy(0)=dz(0)=1m Controlled Stationkeeping l=20km
Tether Pitch and Roll Angles
c n
T J
cdT
Sat 1 Pitch and Roll Angles 1 r
00 05 10 15 Time (Orbits)
Sat 2 Pitch and Roll Angles
c n CD
cdT a
o T J
CQ
- - I 1 1 1 I 1 1 1 1
P i t c h
R o l l
bull
1 -
0 1 2 Time (Orbits)
Tether Y Tranverse Vibration
0 1 2 Time (Orbits)
Tether Z Tranverse Vibration
E gt
10 N
CO
Time (Orbits) Time (Orbits) Longitudinal Vibration
145h
(a) Time (Orbits)
Figure 5-2 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear flexible FLT controller with offset along the local horizontal local vertical and orbit normal (a) attitude and vibration response
101
STSS configuration 3-Body a1(0)=a2(0)=cxt(0)=2deg 31(0)=p2(0)=p(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=d2(0)=1m Controlled Stationkeeping lt=20km
r i i i t i mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash mdash J 0 1 2 0 1 2
Time (Orbits) Time (Orbits) Tether Pitch Control Thrust Tether Roll Control Thruster
(b) Time (Orbits) Time (Orbits)
Figure 5-2 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear flexible FLT controller with offset along the local horizontal local vertical and orbit normal (b) control actuator time histories
102
STSS configuration 3-Body a1(0)=o2(0)=a(0)=2deg P1(0)=p2(0)=(3t(0)=1deg 8x(0)=14m5y(0)=6z(0)=1m dx(0)=1mdy(0)=dz(0)=0 Controlled Stationkeeping lt=20km
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
05 10 15 Time (Orbits)
Sat 2 Pitch and Roll Angles
CD
eg CQ
0 1 2 Time (Orbits)
Tether Y Tranverse Vibration
1 2 Time (Orbits)
Tether Z Tranverse Vibration
gt CO
Time (Orbits) Time (Orbits) Longitudinal Vibration
145h
(a) Time (Orbits)
Figure 5-3 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local vertical (a) attitude and vibration response
103
STSS configuration 3-Body a1(0)=o2(0)=al(0)=2o
P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=1mdy(0)=dz(0)=0 Controlled Stationkeeping lt=20km
Sat 1 Pitch Control Moment 11 1 1 r
Sat 1 Roll Control Moment
0 1 2 Time (Orbits)
Tether Pitch Control Thrust
0 1 2 Time (Orbits)
Sat 2 Pitch Control Moment
0 1 2 Time (Orbits)
Tether Roll Control Thruster
-07 o 1
Time (Orbits Sat 2 Roll Control Moment
1E-5
0E0F
Time (Orbits)
Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local vertical (b) control actuator time histories
104
STSS configuration 3-Body a1(0)=a2(0)=a(0)=2deg Pl(0)=p2(0)=pt(0)=1deg 6X(0)=1304 x 103m 8y(0)=82(0)=0 dx(0)=1mdy(0)=d2(0)=0 Controlled Deployment
l=02km to 20km in 35 Orbits
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
o 1 2 Time (Orbits)
Tether Y Tranverse Vibration
0 1 2 Time (Orbits)
Sat 2 Pitch and Roll Angles
CD CD
CM CO
Time (Orbits) Tether Z Tranverse Vibration
to
150
100
1 2 3 0 1 2
Time (Orbits) Time (Orbits) Longitudinal Vibration
to
(a)
50
2 3 Time (Orbits)
Figure 5 - 4 Deployment dynamics of the three-body STSS using the nonlinear rigid FLT controller with offset along the local vertical (a) attitude and vibration response
105
STSS configuration 3-Body a1(0)=oc2(0)=al(0)=20
P1(0)=p2(0)=pt(0)=1deg 8X(0)=1304 x 103m 5y(0)=52(0)=0 dx(0)=1mdy(0)=dz(0)=0 Controlled Deployment
l=02km to 20km in 35 Orbits
Tether Length Profile i 1
2 0 h
E 2L
Sat 1 Pitch Control Moment t mdash r mdash i mdash i mdash | mdash i mdash i mdash i mdash i mdash | mdash r mdash i mdash i mdash i mdash | mdash r mdash
0 5
1 2 3
Time (Orbits) Sat 1 Roll Control Moment
i i t | t mdash t mdash i mdash [ i i mdash r -
1 2
Time (Orbits) Tether Pitch Control Thrust
1 2 3 Time (Orbits)
Tether Roll Control Thruster
2 E - 5
1 2 3
Time (Orbits) Sat 2 Pitch Control Moment
1 2
Time (Orbits) Sat 2 Roll Control Moment
1 E - 5
E
O E O
Time (Orbits)
Deployment dynamics of the three-body STSS using the nonlinear rigid FLT controller with offset along the local vertical (b) control actuator time histories
106
from 02 km to 20 km in 35 orbits with a i m offset along the tether length A s before
the systems attitude motion is well regulated by the controller However the control
cost increases significantly for the pitch motion of the platform and tether due to the
Coriolis force induced by the deployment maneuver (Figure 5-4b) Initially there is a
sinusoidal increase in the pitch control requirement for both the tether and platform
as the tether accelerates to its constant velocity VQ Then the control requirement
remains constant for the platform at about 60 N m as opposed to the tether where the
thrust demand increases linearly Finally when the tether decelerates the actuators
control input reduces sinusoidally back to zero The rest of the control inputs remain
essentially the same as those in the stationkeeping case In fact the tether pitch
control moment is significantly less since the tether is short during the ini t ia l stages
of control However the inplane thruster requirement Tat acts as a disturbance and
causes 6y to nearly double from the uncontrolled value (Figure 4-8a) O n the other
hand the tether has no out-of-plane deflection
Final ly the effectiveness of the F L T controller during the crit ical maneuver of
retrieval from 20 km to 02 k m in 35 orbits is assessed in Figure 5-5 A s in the earlier
cases the system response in pitch and roll is acceptable however the controller is
unable to suppress the high frequency elastic oscillations induced in the tether by the
retrieval O f course this is expected as there is no active control of the elastic degrees
of freedom However the control of the tether vibrations is discussed in detail in
Section 52 The pitch control requirement follows a similar trend in magnitude as
in the case of deployment with only minor differences due to the addition of offset in
the out-of-plane direction ( d 2 = 1 0 1 ^ m) This is also responsible for the higher
roll moment requirement for the platform control (Figure 5-5b)
107
STSS configuration 3-Body a1(0)=X2(0)=a(0)=2deg P1(0)=p2(0)=pt(0)=1deg 5x(0)=14m8y(0)=52(0)=0 dx(0)=1mdy(0)=dz(0)=0 Controlled Retrieval
l=20km to 02km in 35 Orbits
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
0 1 2 Time (Orbits)
Sat 2 Pitch and Roll Angles
CO CD
cdT
CO CD
T J CM
CO
1 2 Time (Orbits)
Tether Y Tranverse Vibration
Time (Orbits) Tether Z Tranverse Vibration
150
100
1 2 3 0 1 2 Time (Orbits) Time (Orbits)
Longitudinal Vibration
CO
(a)
x 50
2 3 Time (Orbits)
Figure 5 -5 Retrieval dynamics of the three-body STSS using the non-linear rigid FLT controller with offset along the local vertical and orbit normal (a) attitude and vibration response
108
STSS configuration 3-Body a1(0)=ct2(0)=at(0)=2deg p1(0)=p2(0)=(3(0)=1deg 5x(0)=14m5y(0)=82(0)=0 dx(0)=1mdy(0)=dz(0)=0 Controlled Retrieval
l=20km to 02km in 35 Orbits
Tether Length Profile
E
1 2 Time (Orbits)
Tether Pitch Control Thrust
2E-5
1 2 Time (Orbits)
Sat 2 Pitch Control Moment
OEOft
1 2 3 Time (Orbits)
Sat 1 Roll Control Moment
1 2 3 Time (Orbits)
Tether Roll Control Thruster
1 2 Time (Orbits)
Sat 2 Roll Control Moment
1E-5 E z
2
0E0
(b) Time (Orbits) Time (Orbits)
Figure 5-5 Retrieval dynamics of the three-body STSS using the non-linear rigid FLT controller with offset along the local vertical and orbit normal (b) control actuator time histories
109
52 Control of Tethers Elastic Vibrations
521 Prel iminary remarks
The requirement of a precisely controlled micro-gravity environment as well
as the accurate release of satellites into their final orbit demands additional control
of the tethers vibratory motion in addition to its attitude regulation To that end
an active vibration suppression strategy is designed and implemented in this section
The strategy adopted is based on offset control ie time dependent variation
of the tethers attachment point at the platform (satellite 1) A l l the three degrees
of freedom of the offset motion are used to control both the longitudinal as well as
the inplane and out-of-plane transverse modes of vibration In practice the offset
controller can be implemented through the motion of a dedicated manipulator or a
robotic arm supporting the tether which in turn is supported by the platform The
focus here is on the control of elastic deformations during stationkeeping (ie fully
deployed tether fixed length) as it represents the phase when the mission objectives
are carried out
This section begins with the linearization of the systems equations of motion
for the reduced three body stationkeeping case This is followed by the design of
the optimal control algorithm Linear Quadratic Gaussian-Loop Transfer Recovery
( L Q G L T R ) based on the reduced model and its implementation on the full nonlinear
model Final ly the systems response in the presence of offset control is presented
which tends to substantiate its effectiveness
522 System linearization and state-space realization
The design of the offset controller begins with the linearization of the equations
of motion about the systems equilibrium position Linearization of extremely lengthy
(even in matrix form) highly nonlinear nonautonomous and coupled equations of
110
motion presents a challenging problem This is further complicated by the fact the
pitch angle ot is not referred to the L V L H frame Hence the equilibrium pitch angle
is not a constant but is equal to the instantaneous orbital angle 9(t) Two methods
are available to resolve this problem
One may use the non-stationary equations of motion in their present form and
derive a controller based on the Linear T ime Varying ( L T V ) system Alternatively
one can design a Linear T ime Invariant (LTI) controller based on a new set of reduced
governing equations representing the motion of the three body tethered system A l shy
though several studies pertaining to L T V systems have been reported[5660] the latter
approach is chosen because of its simplicity Furthermore the L T I controller design
is carried out completely off-line and thus the procedure is computationally more
efficient
The reduced model is derived using the Lagrangian approach with the genershy
alized coordinates given by
where lty and are the pitch and roll angles of link i relative to the L V L H frame
respectively 5X 8y and 5Z are the generalized coordinates associated with the flexible
tether deformations in the longitudinal inplane and out-of-plane transverse direcshy
tions respectively Only the first mode of vibration is considered in the analysis
The nonlinear nonautonomous and coupled equations of motion for the tethshy
ered system can now be written as
Qred = [ltXlPlltX2P2fixSy5za302gt] (514)
M redQred + fr mdash 0gt (515)
where Mred and frec[ are the reduced mass matrix and forcing term of the system
111
respectively They are functions of qred and qred in addition to the time varying offset
position d2 and its time derivatives d2 and d2 A detailed derivation of Eq(515)
is given in Appendix II A n additional consequence of referring the pitch motion to
a local frame is that the new reduced equations are now independent of the orbital
angle 9 and under the further assumption of the system negotiating a circular orbit
91 remains constant
The nonlinear reduced equations can now be linearized about their static equishy
l ibr ium position For all the generalized coordinates the equilibrium position is zero
wi th the exception of 5X which has a non-zero equilibrium 5XQ The offset position
d2) is linearized about its ini t ial position d2^ Linearizing Eq(515) and recasting
into matr ix form gives
Msqred + CsQred + KsQred + Md^2 + Cdd2 + Kdd2 + fs = 0 (516)
where Ms Cs Ks Md Cd Kd and s are constant matrices Mul t ip ly ing throughout
by Ms
1 gives
qr = -Ms 1Csqr - Ms
lKsqr
- Ms~lCdd22 - Ms-Kdd2 - Ms~Mdd2 - Mg~1 fs r-1 (517)
Defining ud = d2 and v = qTd^T Eq(517) can be rewritten as
Pu t t ing
v =
+
-M~lCs -M~Cd -1
0
-Mg~lMd
Id
0 v + -M~lKs -M~lKd
0 0
-M-lfs
0
MCv + MKv + Mlud +Fs
- 4 ) - ( gt the L T I equations of motion can be recast into the state-space form as
(518)
(519)
x =Ax + Bud + Fd (520)
112
where
A = MC MK Id12 o
24x24
and
Let
B = ^ U 5 R 2 4 X 3 -
Fd = I FJ ) e f t 2 4 1
mdash _ i _ -5
(521)
(522)
(523)
(524)
where x is the perturbation vector from the constant equilibrium state vector xeq
Substituting Eq(524) into Eq(520) gives the linear perturbation state equation
modelling the reduced tethered system as
bullJ x mdash x mdash Ax + Bud + (Fd + Axeq) (525)
It can be shown that for ideal control[60] Fd + Axeq = 0 thus giving the familiar
state-space equation
S=Ax + Bud (526)
The selection of the output vector completes the state space realization of
the system The output vector consists of the longitudinal deformation from the
equil ibrium position SXQ of the tether at xt = h the slope of the tether due to the
transverse deformation at xt = 0 and the offset position d2 from its in i t ia l position
d2Q Thus the output vector y is given by
K(o)sx
V
dx ~ dXQ dy ~ dyQ
d z - d z0 J
Jy
dx da XQ vo
d z - dZQ ) dy dyQ
(527)
113
where lt ( 0 ) = ddxi[$vxi)]Xi=o and lt ( 0 ) = ddxi[4gtw(xi)]Xi=0
523 Linear Quadratic Gaussian control with Loop Transfer Recovery
W i t h the linear state space model defined (Eqs526527) the design of the
controller can commence The algorithm chosen is the Linear Quadratic Gaussian
( L Q G ) estimator based optimal controller[6061] The L Q G is a widely used optimal
controller with its theoretical background well developped by many authors over the
last 25 years It involves of the design of the Kalman-Bucy Fi l ter ( K B F ) which
provides an estimate of the states x and a Linear Quadratic Regulator ( L Q R ) which
is separately designed assuming all the states x are known Both the L Q R and K B F
designs independently have good robustness properties ie retain good performance
when disturbances due to model uncertainty are included However the combined
L Q R and K B F designs ie the L Q G design may have poor stability margins in the
presence of model uncertainties This l imitat ion has led to the development of an L Q G
design procedure that improves the performance of the compensator by recovering
the full state feedback robustness properties at the plant input or output (Figure 5-6)
This procedure is known as the Linear Quadratic Guassian-Loop Transfer Recovery
( L Q G L T R ) control algorithm A detailed development of its theory is given in
Ref[61] and hence is not repeated here for conciseness
The design of the L Q G L T R controller involves the repeated solution of comshy
plicated matrix equations too tedious to be executed by hand Fortunenately the enshy
tire algorithm is available in the Robust Control Toolbox[62] of the popular software
package M A T L A B The input matrices required for the function are the following
(i) state space matrices ABC and D (D = 0 for this system)
(ii) state and measurement noise covariance matrices E and 0 respectively
(iii) state and input weighting matrix Q and R respectively
114
u
MODEL UNCERTAINTY
SYSTEM PLANT
y
u LQR CONTROLLER
X KBF ESTIMATOR
u
y
LQG COMPENSATOR
PLANT
X = f v ( X t )
COMPENSATOR
x = A k x - B k y
ud
= C k f
y
Figure 5-6 Block diagram for the LQGLTR estimator based compensator
115
A s mentionned earlier the main objective of this offset controller is to regu-mdash
late the tether vibration described by the 5 equations However from the response
of the uncontrolled system presented in Chapter 4 it is clear that there is a large
difference between the magnitude of the librational and vibrational frequencies This
separation of frequencies allow for the separate design of the vibration and attitude mdash mdash
controllers Thus only the flexible subsystem composed of the S and d equations is
required in the offset controller design Similarly there is also a wide separation of
frequencies between the longitudinal and transverse modes of vibrations permitt ing mdash
the decoupling of the flexible subsystem into a set of longitudinal (Sx and dx) and
transverse (5y Sz dy and dz) subsystems The appended offset system d must also
be included since it acts as the control actuator However it is important to note
that the inplane and out-of-plane transverse modes can not be decoupled because
their oscillation frequencies are of the same order The two flexible subsystems are
summarized below
(i) Longitudinal Vibra t ion Subsystem
This is defined by u mdash Auxu + Buudu
(528)
where xu = SxdX25xdX2T u d u = dX2 and from Eq(527)
0 0 $u(lt) 0 0 0 0 1
0 0 1 0 0 0 0 1
(529)
Au and Bu are the rows and columns of A and B respectively and correspond to the
components of xu
The L Q R state weighting matrix Qu is taken as
Qu =
bull1 0 0 o -0 1 0 0 0 0 1 0
0 0 0 10
(530)
116
and the input weighting matrix Ru = 1^ The state noise covariance matr ix is
selected as 1 0 0 0 0 1 0 0 0 0 4 0
LO 0 0 1
while the measurement noise covariance matrix is taken as
(531)
i o 0 15
(532)
Given the above mentionned matrices the design of the L Q G L T R compenshy
sator is computed using M A T L A B with the following 2 commands
(i) kfu = l q r c ( ^ ( i C bdquo d i a g m x ( E u e u ) )
(ii) [AkugtBkugtCkugtDku =^ryAuBuCuDukfuQuRur)
where r is a scaler
The first command is used to compute the Ka lman filter gain matrix kfu Once
kfu is known the second command is invoked returning the state space representation
of the L Q G L T R dynamic compensator to
(533) xu mdash Akuxu Bkuyu
where xu is the state estimate vector oixu The function l t ry performs loop transfer
recovery at the system output ie the return ratio at the output approaches that
of the K B F loop given by Cu(sld - Au)Kfu[61 This is achieved by choosing a
sufficently large scaler value r in l t ry such that the singular values of the return
ratio approach those of the target design For the longitudinal controller design
r mdash 5 x 10 5 Figure 5-7(a) compares the singular values of the recovered compensator
design and the non-recovered L Q G compensator design with respect to the target
Unfortunetaly perfect recovery is not possible especially at higher frequencies
117
100
co
gt CO
-50
-100 10
Longitudinal Compensator Singular Values i I I I N I I mdash i mdash i i i i n i i mdash i mdash i i i i inimdash imdashi i i i i i i i mdash i mdash i i M i I I
Target r = 5x10 5
r = 0 (LQG)
i i i i I I i ii i i i 1 1 1 1 i i i i i 1 1 1 1 ii i I I i i i i 11111 -3 10 -2
(a)
101 10deg Frequency (rads)
101 10
Transverse Compensator Singular Values
100
50
X3
gt 0
CO
-50
bull100 10
-imdashi i 111II i 111 ni 1mdashi I I I I I I I 1mdashi I I I N I I 1mdashi i M 111
Target r = 50 r = 0 (LQG)
m i i i i i i n i i i i I I I I I I I 1mdashi I I I I I I I 1mdashi i 111 n -3 10 -2
(b)
101 10deg Frequency (rads)
101 10
Figure 5-7 Singular values for the LQG and LQGLTR compensator compared to target return ratio (a) longitudinal design (b) transverse design
118
because the system is non-minimal ie it has transmission zeros with positive real
parts[63]
(ii) Transverse Vibra t ion Subsystem
Here Xy mdash AyXy + ByU(lv
Dv = Cvxv
with xv = 6y 8Z dV2 dZ2 Sy 6Z dV2 dZ2T] udv = dy2dZ2T and
(534)
Cy
0 0 0 0 (0) 0 0 0 0 0 0 0 0 0 0 0 0
0 0 0 0
0 0 0 0 0 0 0 0 0 0 0 0
0 0 0
h 0 0 0
0 0 0 lt ( 0 ) o 0
0 1 0 0 0 1
0 0 0
2 r
h 0 0
0 0 1 0 0 1
(535)
Av and Bv are the rows and columns of the A and B corresponding to the components
O f Xy
The state L Q R weighting matrix Qv is given as
Qv
10 0 0 0 0 0 0 0 0 10 0 0 0 0 0 0 0 0 10 4 0 0 0 0 0 0 0 0 10 4 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1
(536)
mdash
Ry =w The state noise covariance matrix is taken
1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 9 x 10 4 0
0 0 0 0 0 0 0 9 x 10 4
(537)
119
while the measurement noise covariance matrix is
1 0 0 0 0 1 0 0 0 0 5 x 10 4 0
LO 0 0 5 x 10 4
(538)
As in the development of the longitudinal controller the transverse offset conshy
troller can be designed using the same two commands (lqrc and l try) wi th the input
matrices corresponding to the transverse subsystem with r = 50 The result is the
transverse compensator system given by
(539)
where xv is the state estimates of xv The singular values of the target transfer
function is compared with those achieved by the L Q G and L Q G L T R in Figure 5-
7(b) It is apparent that the recovery is not as good as that for the longitudinal case
again due to the non-minimal system Moreover the low value of r indicates that
only a lit t le recovery in the transverse subsystem is possible This suggests that the
robustness of the L Q G design is almost the maximum that can be achieved under
these conditions
Denoting f y = f j f j r xf = x^x^T and ud = udu ^ T the longishy
tudinal compensator can be combined to give
xf =
$d =
Aku
0
0 Akv
Cku o
Cu 0
Bku
0 B Xf
Xf = Ckxf
0
kv
Hf = Akxf - Bkyf
(540)
Vu Vv 0 Cv
Xf = CfXf
Defining a permutation matrix Pf such that Xj = PfX where X is the state vector
of the original nonlinear system given in Eq(32) the compensator and full nonlinear
120
system equations can be combined as
X
(541)
Ckxf
A block diagram representation of Eq(541) is shown in Figure 5-6
524 Simulation results
The dynamical response for the stationkeeping STSS with a tether length of 20
km and ini t ia l offset position of 1 m in each direction is presented in Figure 5-8 when
both the F L T attitude controller and the L Q G L T R offset controller are activated As
expected the offset controller is successful in quickly damping the elastic vibrations
in the longitudinal and transverse direction (Figure 5-8b) However from Figure 5-
8(a) it is clear that the presence of offset control requires a larger control moment
to regulate the attitude of the platform This is due to the additional torque created
by the larger moment arm around 2 m and 125 m in the inplane and out-of-plane
directions respectively introduced by the offset control In addition the control
moment for the platform is modulated by the tethers transverse vibration through
offset coupling However this does not significantly affect the l ibrational motion of
the tether whose thruster force remains relatively unchanged from the uncontrolled
vibration case
Final ly it can be concluded that the tether elastic vibration suppression
though offset control in conjunction with thruster and momentum-wheel attitude
control presents a viable strategy for regulating the dynamics tethered satellite sysshy
tems
121
STSS configuration 3-Body a1(0)=a2(0)=at(0)=2deg P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m6y(0)=52(0)=1m dx(0)=dy(0)=dz(0)=1m Stationkeeping I = 20km LQGLTR Vibration Control
Sat 1 Pitch and Roll Angles Tether Pitch and Roll Angles mdash 1 1 1 1 1 1 I 1 1 lt 1 ~ ^ 1 ^ l _
Time (Orbits) Time (Orbits) Sat 1 Roll Control Moment Tether Roll Control Thrust
(a) Time (Orbits) Time (Orbits)
Figure 5-8 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT attitude controller and LQGLTR offset vibration controller (a) attitude and libration controller reshysponse
122
STSS configuration 3-Body a1(0)=a2(0)=a(0)=2deg p1(0)=(32(0)=Q(0)=1deg ox(0)=14mSy(0)=5z(0)=1m dx(0)=dy(0)=dz(0)=1m Stationkeeping lt = 20km LQGLTR Vibration Control
Tether X Offset Position
Time (Orbits) Tether Y Tranverse Vibration 20r
Time (Orbits) Tether Y Offset Position
t o
1 2 Time (Orbits)
Tether Z Tranverse Vibration Time (Orbits)
Tether Z Offset Position
(b)
Figure 5-8
Time (Orbits) Time (Orbits)
Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT attitude controller and LQGLTR offset vibration controller (b) vibration and offset response
123
6 C O N C L U D I N G R E M A R K S
61 Summary of Results
The thesis has developed a rather general dynamics formulation for a multi-
body tethered system undergoing three-dimensional motion The system is composed
of multiple rigid bodies connected in a chain configuration by long flexible tethers
The tethers which are free to undergo libration as well as elastic vibrations in three
dimensions are also capable of deployment retrieval and constant length stationkeepshy
ing modes of operation Two types of actuators are located on the rigid satellites
thrusters and momentum-wheels
The governing equations of motion are developed using a new Order(N) apshy
proach that factorizes the mass matrix of the system such that it can be inverted
efficiently The derivation of the differential equations is generalized to account for
an arbitrary number of rigid bodies The equations were then coded in FORTRAN
for their numerical integration with the aid of a symbolic manipulation package that
algebraically evaluated the integrals involving the modes shapes functions used to dis-
cretize the flexible tether motion The simulation program was then used to assess the
uncontrolled dynamical behaviour of the system under the influence of several system
parameters including offset at the tether attachment point stationkeeping deployshy
ment and retrieval of the tether The study covered the three-body and five-body
geometries recently flown OEDIPUS system and the proposed BICEPS configurashy
tion It represents innovation at every phase a general three dimensional formulation
for multibody tethered systems an order-N algorithm for efficient computation and
application to systems of contemporary interest
Two types of controllers one for the attitude motion and the other for the
124
flexible vibratory motion are developed using the thrusters and momentum-wheels for
the former and the variable offset position for the latter These controllers are used to
regulate the motion of the system under various disturbances The attitude controller
is developed using the nonlinear Feedback Linearization Technique ( F L T ) and is based
on the nonlinear but rigid model of the system O n the other hand the separation
of the longitudinal and transverse frequencies from those of the attitude response
allows for the development of a linear optimal offset controller using the robust Linear
Quadratic Gaussian-Loop Transfer Recovery ( L Q G L T R ) method The effectiveness
of the two controllers is assessed through their subsequent application to the original
nonlinear flexible model
More important original contributions of the thesis which have not been reshy
ported in the literature include the following
(i) the model accounts for the motion of a multibody chain-type system undershy
going librational and in the case of tethers elastic vibrational motion in all
the three directions
(ii) the dynamics formulation is based on a recursive O(N) Lagrangian algorithm
that efficiently computes the systems generalized acceleration vector
(iii) the development of an attitude controller based on the Feedback Linearizashy
tion Technique for a multibody system using thrusters and momentum-wheels
located on each rigid body
(iv) the design of a three degree of freedom offset controller to regulate elastic v i shy
brations of the tether using the Linear Quadratic Gaussian and Loop Transfer
Recovery method ( L Q G L T R )
(v) substantiation of the formulation and control strategies through the applicashy
tion to a wide variety of systems thus demonstrating its versatility
125
The emphasis throughout has been on the development of a methodology to
study a large class of tethered systems efficiently It was not intended to compile
an extensive amount of data concerning the dynamical behaviour through a planned
variation of system parameters O f course a designer can easily employ the user-
friendly program to acquire such information Rather the objective was to establish
trends based on the parameters which are likely to have more significant effect on
the system dynamics both uncontrolled as well as controlled Based on the results
obtained the following general remarks can be made
(i) The presence of the platform products of inertia modulates the attitude reshy
sponse of the platform and gives rise to non-zero equilibrium pitch and roll
angles
(ii) Offset of the tether attachment point significantly affects the equil ibrium orishy
entation of the system as well as its dynamics through coupling W i t h a relshy
atively small subsatellite the tether dynamics dominate the system response
with high frequency elastic vibrations modulating the librational motion On
the other hand the effect of the platform dynamics on the tether response is
negligible As can be expected the platform dynamics is significantly affected
by the offset along the local horizontal and the orbit normal
(iii) For a three-body system deployment of the tether can destabilize the platform
in pitch as in the case of nonzero offset However the roll motion remains
undisturbed by the Coriolis force Moreover deployment can also render the
tether slack if it proceeds beyond a critical speed
(iv) Uncontrolled retrieval of the tether is always unstable in pitch It also leads
to high frequency elastic vibrations in the tether
(v) The five-body tethered system exhibits dynamical characteristics similar to
those observed for the three-body case As can be expected now there are
additional coupling effects due to two extra bodies a rigid subsatellite and a
126
flexible tether
(vi) Cartwheeling motion of the B I C E P S configuration can also be used to adshy
vantage in the deployment of the tether However exceeding a crit ical ini t ia l
cartwheeling rate can result in a large tension in the tether which eventually
causes the satellite to bounce back rendering the tether slack
(vii) Spinning the tether about its nominal length as in the case of O E D I P U S
introduces high frequency transverse vibrations in the tether which in turn
affect the dynamics of the satellites
(viii) The F L T based controller is quite successful in regulating the attitude moshy
tion of both the tether and rigid satellites in a short period of time during
stationkeeping deployment as well as retrieval phases
(ix) The offset controller is successful in suppressing the elastic vibrations of the
tether wi thin the allowable l imits of the offset mechanism ( plusmn 2 0 m)
62 Recommendations for Future Study
A s can be expected any scientific inquiry is merely a prelude to further efforts
and insight Several possible avenues exist for further investigation in this field A
few related to the study presented here are indicated below
(i) inclusion of environmental effects particularly the free molecular reaction
forces as well as interaction with Earths magnetic field and solar radiation
(ii) addition of flexible booms attached to the rigid satellites providing a mechashy
nism for energy dissipation
(iii) validation of the new three dimensional offset control strategies using ground
based experiments
(iv) animation of the simulation results to allow visual interpretation of the system
dynamics
127
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[38] Ruying F and Bainum P M Dynamics and Control of a Space Platform with a Tethered Subsatellite Journal of Guidance Control and Dynamics V o l 11 No 4 July-August 1988 pp 377-381
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6 1995 pp 373-384
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[43] Pradhan S Planar Dynamics and Control of Tethered Satellite Systems P h D Thesis Department of Mechanical Engineering The University of Br i t i sh Columbia December 1994
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[46] Hughes P C Spacecraft Att i tude Dynamics John Wi ley amp Sons New York U S A 1992 Chapter 2 pp 7-38
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[50] Meirovitch L Elements of Vibration Analysis M c G r a w - H i l l Inc New York U S A 1986 pp 255-256
[51] Greenwood D T Principles of Dynamics Prentice Ha l l Inc Englewood Cliffs New Jersey U S A 1965 pp 252-275
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132
A P P E N D I X I TENSOR R E P R E S E N T A T I O N OF T H E
EQUATIONS OF M O T I O N
11 Prel iminary Remarks
The derivation of the equations of motion for the multibody tethered system
involves the product of several matrices as well as the derivative of matrices and
vectors with respect to other vectors In order to concisely code these relationships
in FORTRAN the matrix equations of motion are expressed in tensor notation
Tensor mathematics in general is an extremely powerful tool which can be
used to solve many complex problems in physics and engineering[6465] However
only a few preliminary results of Cartesian tensor analysis are required here They
are summarized below
12 Mathematical Background
The representation of matrices and vectors in tensor notation simply involves
the use of indices referring to the specific elements of the matrix entity For example
v = vk (k = lN) (Il)
A = Aij (i = lNj = lM)
where vk is the kth element of vector v and Aj is the element on the t h row and
3th column of matrix A It is clear from Eq(Il) that exactly one index is required to
completely define an entire vector whereas two independent indices are required for
matrices For this reason vectors and matrices are known as first-order and second-
order tensors respectively Scaler variables are known as zeroth-order tensors since
in this case an index is not required
133
Matrix operations are expressed in tensor notation similar to the way they are
programmed using a computer language such as FORTRAN or C They are expressed
in a summation notation known as Einstein notation For example the product of a
matrix with a vector is given as
w = Av (12)
or in tensor notation W i = ^2^2Aij(vkSjk)
4-1 ltL3gt j
where 8jk is the Kronecker delta Dropping the summation symbol the matrix-vector
product can be expressed compactly as
Wi = AijVj = Aimvm (14)
Here j is a dummy index since it appears twice in a single term and hence can
be replaced by any other index Note that since the resulting product is a vector
only one index is required namely i that appears exactly once on both sides of the
equation Similarly
E = aAB + bCD (15)
or Ej mdash aAimBmj + bCimDmj
~ aBmjAim I bDmjCim (16)
where A B C D and E are second-order tensors (matrices) and a and b are zeroth-
order tensors (scalers) In addition once an expression is in tensor form it can be
treated similar to a scaler and the terms can be rearranged
The transpose of a matrix is also easily described using tensors One simply
switches the position of the indices For example let w = ATv then in tensor-
notation
Wi = AjiVj (17)
134
The real power of tensor notation is in its unambiguous expression of terms
containing partial derivatives of scalers vectors or matrices with respect to other
vectors For example the Jacobian matrix of y = f(x) can be easily expressed in
tensor notation as
di-dx-j-1 ( L 8 )
mdash which is a second-order tensor as expected If f(x) = Ax where A is constant then
f - a )
according to the current definition If C = A(x)B(x) then the time derivative of C
is given by
C = Ax)Bx) + A(x)Bx) (110)
or in tensor notation
Cij = AimBmj + AimBmj (111)
= XhiAimfcBmj + AimBmj^k)
which is more clear than its equivalent vector form Note that A^^ and Bmjk are
third-order tensors that can be readily handled in F O R T R A N
13 Forcing Function
The forcing function of the equations of motion as given by Eq(260) is
- u 1-SdM- dqtdPe 9 f i m F(q qt) = Mq - -q mdashq + mdashmdash - Qd (260)
This expression can be converted into tensor form to facilitate its implementation
into F O R T R A N source code
Consider the first term Mq From Eq(289) the mass matrix M is given by
T M mdash Rv MtRv which in tensor form is
Mtj = RZiM^R^j (112)
135
Taking the time derivative of M gives
Mij = qsRv
ni^MtnrnRv
mj + Rv
niMtnTn^Rv
mj + Rv
niMtnmRv
mjs) (113)
The second term on the right hand side in Eq(260) is given by
-JPdMu 1 bdquo T x
2Q ~cWq = 2Qs s r k Q r ^ ^
Expanding Msrk leads to
M src = RnskMtnmRmr + RnsMtnmkRmr + RnsMtnmRmrk- (L15)
Finally the potential energy term in Eq(260) is also expanded into tensor
form to give dPe = dqt dPe
dq dq dqt = dq^QP^
and since ^ = Rp(q)q then
Now inserting Eqs(I13-I17) into Eq(260) and rearranging the tensor form of the
forcing term can how be stated as
Fk(q^q^)=(RP
sk + R P n J q n ) ^ - - Q d k
+ QsQr | (^Rlks ~ 2 M t n m R m r
+ (^RnkMtnms ~ 2~RnsMtnmgtk^ R m r
(^RnkRmrs ~ ~2RnsRmrk^j
where Fk(q q t) represents the kth component of F
(118)
136
A P P E N D I X II R E D U C E D EQUATIONS OF M O T I O N
II 1 Prel iminary Remarks
The reduced model used for the design of the vibration controller is represented
by two rigid end-bodies capable of three-dimensional attitude motion interconnected
with a fixed length tether The flexible tether is discretized using the assumed-
mode method with only the fundamental mode of vibration considered to represent
longitudinal as well as in-plane and out-of-plane transverse deformations In this
model tether structural damping is neglected Finally the system is restricted to a
nominal circular orbit
II 2 Derivation of the Lagrangian Equations of Mot ion
Derivation of the equations of motion for the reduced system follows a similar
procedure to that for the full case presented in Chapter 2 However in this model the
straightforward derivation of the energy expressions for the entire system is undershy
taken ignoring the Order(N) approach discussed previously Furthermore in order
to make the final governing equations stationary the attitude motion is now referred
to the local LVLH frame This is accomplished by simply defining the following new
attitude angles
ci = on + 6
Pi = Pi (HI)
7t = 0
where and are the pitch and roll angles respectively Note that spin is not
considered in the reduced model hence 7J = 0 Substituting Eq(IIl) into Eq(214)
137
gives the new expression for the rotation matrix as
T- CpiSa^ Cai+61 S 0 S a + e i
-S 0 c Pi
(II2)
W i t h the new rotation matrix defined the inertial velocity of the elemental
mass drrii o n the ^ link can now be expressed using Eq(215) as
(215)
Since the tether length is assumed to be fixed T j f j and Tj$jlt^ of Eq(215) are
identically zero Also defining
(II3) Vi = I Pi i
Rdm- f deg r the reduced system is given by
where
and
Rdm =Di + Pi(gi)m + T^iSi
Pi(9i)m= [Tai9i T0gi T^] fj
(II4)
(II5)
(II6)
D1 = D D l D S v
52 = 31 + P(d2)m +T[d2
D3 = D2 + l2 + dX2Pii)m + T2i6x
Note that $ j for i = 2 is defined in Section 212 whereas for i = 1 and 3 it is
the null matrix Since only one mode is used for each flexible degree of freedom
5i = [6Xi6yi6Zi]T for i = 2
The kinetic energy of the system can now be written as
i=l Z i = l Jmi RdmRdmdmi (UJ)
138
Expanding Kampi using Eq(II4)
KH = miDD + 2DiP
i j gidmi)fji + 2D T ltMtrade^
+tf J P^(9i)Pi9i)d^l + 2mT j PFmTi^dm^ (II8)
where the integrals are evaluated using the procedure similar to that described in
Chapter 2 The gravitational and strain energy expressions are given by Eq(247)
and Eq(251) respectively using the newly defined rotation matrix T in place of
TV
Substituting the kinetic and potential energy expressions in
d ( 8Ke d(Ke - Pe) i bull i Q ^ 0 (II9)
d t dqred d(ired
with
Qred = [ algt l gt a 2 2 A 2 A 2 lt ^ 2 Q 3 3 ] (5-14)
and d2 regarded as a time-varying parameter the nonlinear non-autonomous equashy
tions of motion for the reduced system can finally be expressed as
MredQred + fred = reg- (5-15)
139
6 C O N C L U D I N G R E M A R K S 1 2 4
61 Summary of Results 124
62 Recommendations for Future Study 127
B I B L I O G R A P H Y 128
A P P E N D I C E S
I T E N S O R R E P R E S E N T A T I O N O F T H E E Q U A T I O N S O F M O T I O N 133
11 Prel iminary Remarks 133
12 Mathematical Background 133
13 Forcing Function 135
II R E D U C E D E Q U A T I O N S O F M O T I O N 137
II 1 Preliminary Remarks 137
II2 Derivation of the Lagrangian Equations of Mot ion 137
v i
LIST OF SYMBOLS
A j intermediate length over which the deploymentretrieval profile
for the itfl l ink is sinusoidal
A Lagrange multipliers
EUQU longitudinal state and measurement noise covariance matrices
respectively
matrix containing mode shape functions of the t h flexible link
pitch roll and yaw angles of the i 1 link
time-varying modal coordinate for the i 1 flexible link
link strain and stress respectively
damping factor of the t h attitude actuator
set of attitude angles (fjj = ci
structural damping coefficient for the t h l ink EjJEj
true anomaly
Earths gravitational constant GMejBe
density of the ith link
fundamental frequency of the link longitudinal tether vibrashy
tion
A tethers cross-sectional area
AfBfCjDf state-space representation of flexible subsystem
Dj inertial position vector of frame Fj
Dj magnitude of Dj mdash mdash
Df) transformation matrix relating Dj and Ds
D r i inplane radial distance of the first l ink
DSi Drv6i DZ1 T for i = 1 and DX Dy poundgt 2 T for 1
m
Pi
w 0 i
v i i
Dx- transformation matrix relating D j and Ds-
DXi Dyi Dz- Cartesian components of D
DZl out-of-plane position component of first link
E Youngs elastic modulus for the ith l ink
Ej^ contribution from structural damping to the augmented complex
modulus E
F systems conservative force vector
FQ inertial reference frame
Fi t h link body-fixed reference frame
n x n identity matrix
alternate form of inertia matrix for the i 1 l ink
K A Z $i(k + dXi+1)
Kei t h link kinetic energy
M(q t) systems coupled mass matrix
Mma Mma Mmy- control moments in the pitch roll and yaw directions respec-
tively for the t h link
Mr fr rigid mass matrix and force vector respectively
Mred fred- Qred mass matrix force vector and generalized coordinate vector for
the reduced model respectively
Mt block diagonal decoupled mass matrix
symmetric decoupled mass matrix
N total number of links
0N) order-N
Pi(9i) column matrix [T^giTpg^T^gi) mdash
Q nonconservative generalized force vector
Qu actuator coupling matrix or longitudinal L Q R state weighting
matrix
v i i i
Rdm- inertial position of the t h link mass element drrii
RP transformation matrix relating qt and tf
Ru longitudinal L Q R input weighting matrix
RvRn Rd transformation matrices relating qt and q
S(f Mdc) generalized acceleration vector of coupled system q for the full
nonlinear flexible system
th uiirv l u w u u u maniA
th
T j i l ink rotation matrix
TtaTto control thrust in the pitch and roll direction for the im l ink
respectively
rQi maximum deploymentretrieval velocity of the t h l ink
Vei l ink strain energy
Vgi link gravitational potential energy
Rayleigh dissipation function arising from structural damping
in the ith link
di position vector to the frame F from the frame F_i bull
dc desired offset acceleration vector (i 1 l ink offset position)
drrii infinitesimal mass element of the t h l ink
dx^dy^dZi Cartesian components of d along the local vertical local horishy
zontal and orbit normal directions respectively
F-Q mdash
gi r igid and flexible position vectors of drrii fj + ltEraquolt5
ijk unit vectors 1 0 0 r 010-^ and 00 l r respectively
li length of the t h l ink
raj mass of the t h link
nfj total number of flexible modes for the i 1 l ink nXi + nyi + nZi
nq total number of generalized coordinates per link nfj + 7
nqq systems total number of generalized coordinates Nnq
ix
number of flexible modes in the longitudinal inplane and out-
of-plane transverse directions respectively for the i ^ link
number of attitude control actuators
- $ T
set of generalized coordinates for the itfl l ink which accounts for
interactions with adjacent links
qv---QtNT
set of coordinates for the independent t h link (not connected to
adjacent links)
rigid position of dm in the frame Fj
position of centre of mass of the itfl l ink relative to the frame Fj
i + $DJi
desired settling time of the j 1 attitude actuator
actuator force vector for entire system
flexible deformation of the link along the Xj yi and Zj direcshy
tions respectively
control input for flexible subsystem
Cartesian components of fj
actual and estimated state of flexible subsystem respectively
output vector of flexible subsystem
LIST OF FIGURES
1-1 A schematic diagram of the space platform based N-body tethered
satellite system 2
1-2 Associated forces for the dumbbell satellite configuration 3
1- 3 Some applications of the tethered satellite systems (a) multiple
communication satellites at a fixed geostationary location
(b) retrieval maneuver (c) proposed B I C E P S configuration 6
2- 1 Vector components of the i 1 and (i- l ) 1 chain links 15
2-2 Vector components in cylindrical orbital coordinates 20
2-3 Inertial position of subsatellite thruster forces 38
2- 4 Coupled force representation of a momentum-wheel on a rigid body 40
3- 1 Flowchart showing the computer program structure 45
3-2 Kinet ic and potentialenergy transfer for the three-body platform based
tethered satellite system (a) variation of kinetic and potential energy
(b) percent change in total energy of system 48
3-3 Kinet ic and potential energy transfer for the five-body platform based
tethered satellite system (a) variation of kinetic and potential energy
(b) percent change in total energy of system 49
3-4 Simulation results for the platform based three-body tethered
system originally presented in Ref[43] 51
3- 5 Simulation results for the platform based three-body tethered
system obtained using the present computer program 52
4- 1 Schematic diagram showing the generalized coordinates used to
describe the system dynamics 55
4-2 Stationkeeping dynamics of the three-body STSS configuration
without offset (a) attitude response (b) vibration response 57
4-3 Stationkeeping dynamics of the three-body STSS configuration
xi
with offset along the local vertical
(a) attitude response (b) vibration response 60
4-4 Stationkeeping dynamics of the three-body STSS configuration
with offset along the local horizontal
(a) attitude response (b) vibration response 62
4-5 Stationkeeping dynamics of the three-body STSS configuration
with offset along the orbit normal
(a) attitude response (b) vibration response 64
4-6 Stationkeeping dynamics of the three-body STSS configuration
wi th offset along the local horizontal local vertical and orbit normal
(a) attitude response (b) vibration response 67
4-7 Deployment dynamics of the three-body STSS configuration
without offset
(a) attitude response (b) vibration response 69
4-8 Deployment dynamics of the three-body STSS configuration
with offset along the local vertical
(a) attitude response (b) vibration response 72
4-9 Retrieval dynamics of the three-body STSS configuration
with offset along the local vertical and orbit normal
(a) attitude response (b) vibration response 74
4-10 Schematic diagram of the five-body system used in the numerical example 77
4-11 Stationkeeping dynamics of the five-body STSS configuration
without offset
(a) attitude response (b) vibration response 78
4-12 Deployment dynamics of the five-body STSS configuration
wi th offset along the local vertical
(a) attitude response (b) vibration response 81
4-13 Stationkeeping dynamics of the three-body B I C E P S configuration
x i i
with offset along the local vertical
(a) attitude response (b) vibration response 84
4-14 Cartwheeling dynamics with deployment for the three-body B I C E P S
with offset along the local vertical
(a) attitude response (b) vibration response 86
4-15 Spin dynamics (7 = l deg s ) of the three-body O E D I P U S configuration
with offset along the local vertical
(a) attitude response (b) vibration response 89
4- 16 Spin dynamics (7 = 10deg s ) of the three-body O E D I P U S configuration
wi th offset along the local vertical
(a) attitude response (b) vibration response 91
5- 1 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear rigid F L T controller with offset along the
local horizontal local vertical and orbit normal
(a) attitude and vibration response (b) control actuator time histories 98
5-2 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear flexible F L T controller with offset along the
local horizontal local vertical and orbit normal
(a) attitude and vibration response (b) control actuator time histories 101
5-3 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear rigid F L T controller with offset along
the local vertical
(a) attitude and vibration response (b) control actuator time histories 103
5-4 Deployment dynamics of the three-body STSS using the nonlinear
rigid F L T controller with offset along the local vertical
(a) attitude and vibration response (b) control actuator time histories 105
5-5 Retrieval dynamics of the three-body STSS using the non-linear
rigid F L T controller with offset along the local vertical and orbit normal
(a) attitude and vibration response (b) control actuator time histories 108
x i i i
5-6 Block diagram for the L Q G L T R estimator based compensator 115
5-7 Singular values for the L Q G and L Q G L T R compensator compared to target
return ratio (a) longitudinal design (b) transverse design 118
5-8 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear rigid F L T attitude controller and L Q G L T R
offset vibration controller (a) attitude and libration controller response
(b) vibration and offset response 122
xiv
A C K N O W L E D G E M E N T
Firs t and foremost I would like to express the most genuine gratitude to
my supervisor Prof V i n o d J M o d i whose unending guidance and encouragement
proved most invaluable during the course of my studies
I am also deeply indebted to Dr Satyabrata Pradhan for his limitless patience
and technical advice from which this work would not have been possible I would also
like to extend a word of appreciation towards Prof A r u n K Mis ra ( M c G i l l University)
for ini t iat ing my interest in space dynamics and control
Also I would like to thank my collegues Dr Sandeep Munshi M r Mathieu
Caron and M r Yuan Chen as well as Dr Mae Seto M r Gary L i m and M r Mark
Chu for their words of advice and helpful suggestions that heavily contributed to all
aspects of my studies
Final ly I would like to thank all my family and friends here in Vancouver and
back home in Montreal whose unending support made the past two years the most
memorable of my life
The research project was supported by the Natural Sciences and Engineering
Research Council s P G S - A Scholarship held by the author as well as by N S E R C grant
A-2181 held by Prof V J M o d i
xv
1 INTRODUCTION
11 Prel iminary Remarks
W i t h the ever changing demands of the worlds population one often wonders
about the commitment to the space program On the other hand the importance
of worldwide communications global environmental monitoring as well as long-term
habitation in space have demanded more ambitious and versatile satellites systems
It is doubtful that the Russian scientist Tsiolkovsky[l] when he first proposed the
uti l izat ion of the Earths gravity-gradient environment in the last century would have
envisioned the current and proposed applications of tethered satellite systems
A tethered satellite system consists of two or more subsatellites connected
to each other by long thin cables or tethers which are in tension (Figure 1-1) The
subsatellite which can also include the shuttle and space station are in orbit together
The first proposed use of tethers in space was associated with the rescue of stranded
astronauts by throwing a buoy on a tether and reeling it to the rescue vehicle
Prel iminary studies of such systems led to the discovery of the inherent instability
during the tether retrieval[2] Nevertheless the bir th of tethered systems occured
during the Gemini X I and X I I missions[3] in 1966 when a short tether (10-30m)
was used to generate the first artificial gravity environment in space (000015g) by
cartwheeling (spinning) the system about its center of mass
The mission also demonstrated an important use of tethers for gravity gradient
stabilization The force analysis of a simplified model illustrates this point Consider
the dumbbell satellite configuration orbiting about the centre of Ear th as shown in
Figure 1-2 Satellite 1 and satellite 2 are connected to each other by a tether with
1
Space Station (Satellite 1 )
2
Satellite N
Figure 1-1 A schematic diagram of the space platform based N-body tethered satellite system
2
Figure 1-2 Associated forces for the dumbbell satellite configuration
3
the whole system free to librate about the systems centre of mass C M The two
major forces acting on each body are due to the centrifugal and gravitational effects
Since body 1 is further away from the earth its gravitational force Fg^y is smaller
than that of body 2 ie Fgi lt Fg^ However its centrifugal force FCl is greater
then the centrifugal force experienced by the second satellite (FCl gt FC2) Moreover
for body 1 Fg^ lt Fc^ in contrast to Fg2 gt FC2 for body 2 Thus resolving the force
components along the tether there is an evident resultant tension force Ff present
in the tether Similarly adding the normal components of the two forces vectorially
results in a force FR which restores the system to its stable equil ibrium along the
local vertical These tension and gravity-gradient restoring forces are the two most
important features of tethered systems
Several milestone missions have flown in the recent past The U S A J a p a n
project T P E (Tethered Payload Experiment) [4] was one of the first to be launched
using a sounding rocket to conduct environmental studies The results of the T P E
provided support for the N A S A I t a l y shuttle based tethered satellite system referred
to as the TSS-1 mission[5] This experiment aimed to study the electrodynamic effects
of a shuttle borne satellite system with a conductive tether connection where an
electric current was induced in the tether as it swept through the Earth s magnetic
field Unfortunetely the two attempts in August 1992 and February 1996 resulted
in only partial success The former suffered from a spool failure resulting in a tether
deployment of only 256m of a planned 20km During the latter attempt a break
in the tethers insulation resulted in arcing leading to tether rupture Nevertheless
the information gathered by these attempts st i l l provided the engineers and scientists
invaluable information on the dynamic behaviour of this complex system
Canada also paved the way with novel designs of tethered system aimed
primari ly at the study of the ionosphere and Northern Lights (Aurora Borealis)
4
The O E D I P U S A and C (Observation of Electrified Distributions in the Ionospheric
P lasma - a Unique Strategy) missions[6] launched from a sounding rocket in January
1989 and November 1995 respectively provided insight into the complex dynamical
behaviour of two comparable mass satellites connected by a 1 km long spinning
tether
Final ly two experiments were conducted by N A S A called S E D S I and II
(Small Expendable Deployable System) [7] which hold the current record of 20 km
long tether In each of these missions a small probe (26 kg) was deployed from the
second stage of a D E L T A II rocket thus succesfully demonstrating the feasibility of
long tether deployment Retrieval of the tether which is significantly more difficult
has not yet been achieved
Several proposed applications are also currently under study These include the
study of earths upper atmosphere using probes lowered from the shuttle in a low earth
orbit ( L E O ) to the deployment and retrieval of satellites for servicing Even more
promising application concerns the generation of power for the proposed International
Space Station using conductive tethers The Canadian Space Agency (CSA) has also
proposed a dual-mass satellite system B I C E P S (Bl-static Canadian Experiment on
Plasmas in Space) to make simultaneous measurements at different locations in the
environment for correlation studies[8] A unique feature of the B I C E P S mission is
the deployment of the tether aided by the angular momentum of the system which
is ini t ia l ly cartwheeling about its orbit normal (Figure 1-3)
In addition to the three-body configuration multi-tethered systems have been
proposed for monitoring Earths environment in the global project monitored by
N A S A entitled Mission to Planet Ear th ( M T P E ) The proposed mission aims to
study pollution control through the understanding of the dynamic interactions be-
5
Retrieval of Satellite for Servicing
lt mdash
lt Satellite
Figure 1-3 Some applications of the tethered satellite systems (a) multiple comshymunication satellites at a fixed geostationary location (b) retrieval maneuver (c) proposed BICEPS configuration
6
tween the Earths atmosphere biosphere hydrosphere and cryosphere This wi l l be
accomplished with multiple tethered instrumentation payloads simultaneously soundshy
ing at different altitudes for spatial correlations Such mult ibody systems are considshy
ered to be the next stage in the evolution of tethered satellites Micro-gravity payload
modules suspended from the proposed space station for long-term experiments as well
as communications satellites with increased line-of-sight capability represent just two
of the numerous possible applications under consideration
12 Brief Review of the Relevant Literature
The design of multi-payload systems wi l l require extensive dynamic analysis
and parametric study before the final mission configuration can be chosen This
wi l l include a review of the fundamental dynamics of the simple two-body tethered
systems and how its dynamics can be extended to those of multi-body tethered system
in a chain or more generally a tree topology
121 Multibody 0(N) formulation
Many studies of two-body systems have been reported One of the earliest
contribution is by Rupp[9] who was the first to study planar dynamics of the simshy
plified Shuttle based Tethered Satellite System (STSS) His pioneering work led to
the discovery of pitch oscillation growth during the retrieval phase Later Baker[10]
advanced the investigation to the third dimension in addition to adding atmospheric
effects to the system A more complete dynamical model was later developed and
analyzed by M o d i and Misra[ l l 12] which included deployment and tether flexibility
A comprehensive survey of important developments and milestones in the area have
been compiled and reviewed by Mis ra and Modi[13] The major conclusions based on
the literature may be summarized as follows
bull the stationkeeping phase is stable
7
bull deployment can be unstable if the rate exceeds a crit ical speed
bull retrieval of the tether is inherently unstable
bull transverse vibrations can grow due to the Coriolis force induced during deshy
ployment and retrieval
Misra Amier and Modi[14] were one of the first to extend these results to the
three-satellite double pendulum case This simple model which includes a variable
length tether was sufficient to uncover the multiple equilibrium configurations of
the system Later the work was extended to include out-of-plane motion and a
preliminary reel-rate control law to regulate the librational motion[15] Kesmir i and
Misra[16] developed a general formulation for N-body tethered systems based on the
Lagrangian principle where three-dimensional motion and flexibility are accounted for
However from their work it is clear that as the number of payload or bodies increases
the computational cost to solve the forward dynamics also increases dramatically
Tradit ional methods of inverting the mass matrix M in the equation
ML+F = Q
to solve for the acceleration vector proved to be computationally costly being of the
order for practical simulation algorithms O(N^) refers to a mult ibody formulashy
tion algorithm whose computational cost is proportional to the cube of the number
of bodies N used It is therefore clear that more efficient solution strategies have to
be considered
Many improved algorithms have been proposed to reduce the number of comshy
putational steps required for solving for the acceleration vector Rosenthal[17] proshy
posed a recursive formulation based on the triangularization of the equations of moshy
tion to obtain an 0(N2) formulation He also demonstrates the use of a symbolic
manipulation program to reduce the final number of computations A recursive L a -
8
grangian formulation proposed by Book[18] also points out the relative inefficiency of
conventional formulations and proceeds to derive an algorithm which is also 0(N2)
A recursive formulation is one where the equations of motion are calculated in order
from one end of the system to the other It usually involves one or more passes along
the links to calculate the acceleration of each link (forward pass) and the constraint
forces (backward or inverse pass) Non-recursive algorithms express the equations of
motion for each link independent of the other Although the coupling terms st i l l need
to be included the algorithm is in general more amenable to parallel computing
A st i l l more efficient dynamics formulation is an O(N) algorithm Here the
computational effort is directly proportional to the number of bodies or links used
Several types of such algorithms have been developed over the years The one based
on the Newton-Euler approach has recursive equations and is used extensively in the
field of multi-joint robotics[1920] It should be emphasized that efficient computation
of the forward and inverse dynamics is imperative if any on-line (real-time) control
of the joint is to be achieved Hollerbach[21] proposed a recursive O(N) formulation
based on Lagranges equations of motion However his derivation was primari ly foshy
cused on the inverse dynamics and did not improve the computational efficiency of
the forward dynamics (acceleration vector) Other recursive algorithms include methshy
ods based on the principle of vir tual work[22] and Kanes equations of motion[23]
Keat[24] proposed a method based on a velocity transformation that eliminated the
appearance of constraint forces and was recursive in nature O n the other hand
K u r d i l a Menon and Sunkel[25] proposed a non-recursive algorithm based on the
Range-Space method which employs element-by-element approach used in modern
finite-element procedures Authors also demonstrated the potential for parallel comshy
putation of their non-recursive formulation The introduction of a Spatial Operator
Factorization[262728] which utilizes an analogy between mult ibody robot dynamics
and linear filtering and smoothing theory to efficiently invert the systems mass ma-
9
t r ix is another approach to a recursive algorithm Their results were further extended
to include the case where joints follow user-specified profiles[29] More recently Prad-
han et al [30] proposed a non-recursive Lagrangian algorithm for factorizing the mass
matrix leading to an O(N) formulation of the forward dynamics of the system A n
efficient O(N) algorithm where the number of bodies N in the system varies on-line
has been developed by Banerjee[31]
122 Issues of tether modelling
The importance of including tether flexibility has been demonstrated by several
researchers in the past However there are several methods available for modelling
tether flexibility Each of these methods has relative advantages and limitations
wi th regards to the computational efficiency A continuum model where flexible
motion is discretized using the assumed mode method has been proposed[3233] and
succesfully implemented In the majority of cases even one flexible mode is sufficient
to capture the dynamical behaviour of the tether[34] K i m and Vadali[35] proposed a
so-called bead model or lumped-mass approach that discretizes the tether using point
masses along its length However in order to accurately portray the motion of the
tether a significantly high number of beads are needed thus increasing the number of
computation steps Consequently this has led to the development of a hybrid between
the last two approaches[16] ie interpolation between the beads using a continuum
approach thus requiring less number of beads Finally the tether flexibility can also
be modelled using a finite-element approach[36] which is more suitable for transient
response analysis In the present study the continuum approach is adopted due to
its simplicity and proven reliability in accurately conveying the vibratory response
In the past the modal integrals arising from the discretization process have
been evaluted numerically In general this leads to unnecessary effort by the comshy
puter Today these integrals can be evaluated analytically using a symbolic in-
10
tegration package eg Maple V Subsequently they can be coded in F O R T R A N
directly by the package which could also result in a significant reduction in debugging
time More importantly there is considerable improvement in the computational effishy
ciency [16] especially during deployment and retrieval where the modal integral must
be evaluated at each time-step
123 Attitude and vibration control
In view of the conclusions arrived at by some of the researchers mentioned
above it is clear that an appropriate control strategy is needed to regulate the dyshy
namics of the system ie attitude and tether vibration First the attitude motion of
the entire system must be controlled This of course would be essential in the case of
a satellite system intended for scientific experiments such as the micro-gravity facilishy
ties aboard the proposed Internation Space Station Vibra t ion of the flexible members
wi l l have to be checked if they affect the integrity of the on-board instrumentation
Thruster as well as momentum-wheel approach[34] have been considered to
regulate the rigid-body motion of the end-satellite as well as the swinging motion
of the tether The procedure is particularly attractive due to its effectiveness over a
wide range of tether lengths as well as ease of implementation Other methods inshy
clude tension and length rate control which regulate the tethers tension and nominal
unstretched length respectively[915] It is usually implemented at the deployment
spool of the tether More recently an offset strategy involving time dependent moshy
tion of the tether attachment point to the platform has been proposed[3738] It
overcomes the problem of plume impingement created by the thruster control and
the ineffectiveness of tension control at shorter lengths However the effectiveness of
such a controller can become limited with an exceedingly long tether due to a pracshy
tical l imi t on the permissible offset motion In response to these two control issues a
hybrid thrusteroffset scheme has been proposed to combine the best features of the
11
two methods[39]
In addition to attitude control these schemes can be used to attenuate the
flexible response of the tether Tension and length rate control[12] as well as thruster
based algorithms[3340] have been proposed to this end M o d i et al [41] have sucshy
cessfully demonstrated the effectiveness of an offset strategy to damp undesired v i shy
brations Passive energy dissipative devices eg viscous dampers are also another
viable solution to the problem
The development of various control laws to implement the above mentioned
strategies has also recieved much attention A n eigenstructure assignment in conducshy
tion wi th an offset controller for vibration attenuation and momemtum wheels for
platform libration control has been developed[42] in addition to the controller design
from a graph theoretic approach[41] Also non-linear feedback methods such as the
Feedback Linearization Technique ( F L T ) that are more suitable for controlling highly
nonlinear non-autonomous coupled systems have also been considered[43]
It is important to point out that several linear controllers including the classhy
sic state feedback Linear Quadratic Regulator ( L Q R ) have received considerable
attention[3944] Moreover robust methods such as the Linear Quadratic Guas-
s ian Loop Transfer Recovery ( L Q G L T R ) method have also been developed and imshy
plemented on tethered systems[43] A more complete review of these algorithms as
well as others applied to tethered system has been presented by Pradhan[43]
13 Scope of the Investigation
The objective of the thesis is to develop and implement a versatile as well as
computationally efficient formulation algorithm applicable to a large class of tethered
satellite systems The distinctive features of the model can be summarized as follows
12
bull TV-body 0(N) tethered satellite system in a chain-type topology
bull the system is free to negotiate a general Keplerian orbit and permitted to
undergo three dimensional inplane and out-of-plane l ibrtational motion
bull r igid bodies constituting the system are free to execute general rotational
motion
bull three dimensional flexibility present in the tether which is discretized using
the continuum assumed-mode method with an arbitrary number of flexible
modes
bull energy dissipation through structural damping is included
bull capability to model various mission configurations and maneuvers including
the tether spin about an arbitrary axis
bull user-defined time dependent deployment and retrieval profiles for the tether
as well as the tether attachment point (offset)
bull attitude control algorithm for the tether and rigid bodies using thrusters and
momentum-wheels based on the Feedback Linearization Technique
bull the algorithm for suppression of tether vibration using an active offset (tether
attachment point) control strategy based on the optimal linear control law
( L Q G L T - R )
To begin with in Chapter 2 kinematics and kinetics of the system are derived
using the O(N) formulation methodology developed by Pradhan et al [30] Chapshy
ter 3 discusses issues related to the development of the simulation program and its
validation This is followed by a detailed parametric study of several mission proshy
files simulated as particular cases of the versatile formulation in Chapter 4 Next
the attitude and vibration controllers are designed and their effectiveness assessed in
Chapter 5 The thesis ends with concluding remarks and suggestions for future work
13
2 F O R M U L A T I O N OF T H E P R O B L E M
21 Kinematics
211 Preliminary definitions and the i ^ position vector
The derivation of the equations of motion begins with the definition of the
inertial position of an arbitrary link of the mult ibody system Let the i ^ link of the
system be free to translate and rotate in 3-D space From Figure 2-1 the position
vector Rdm^ to the mass element drrn on the i ^ link wi th reference to the inertial
frame FQ can be written as
Here D represents the inertial position of the t h body fixed frame Fj relative to FQ
ri = [xiyiZi]T is the rigid position vector of drrii wi th respect to F J and f(fi) is
the flexible deformation at fj also with respect to the frame Fj Bo th these vectors
are relative to Fj Note in the present model tethers (i even) are considered flexible
and X corresponds to the nominal position along the unstretched tether while a and
ZJ are by definition equal to zero On the other hand the rigid bodies referred to as
satellites (i odd) including the platform are taken to be rigid ie f(fj) = 0 T j is
the rotation matrix used to express body-fixed vectors with reference to the inertial
frame
212 Tether flexibility discretization
The tether flexibility is discretized with an arbitrary but finite number of
Di + Ti(i + f(i)) (21)
14
15
modes in each direction using the assumed-mode method as
Ui nvi
wi I I i = 1
nzi bull
(22)
where nXi nyi and n 2 ^ are the number of modes used in the X m and Z directions
respectively For the ] t h longitudinal mode the admissible mode shape function is
taken as 2 j - l
Hi(xiii)=[j) (23)
where l is the i ^ tether length[32] In the case of inplane and out-of-plane transverse
deflections the admissible functions are
ampyi(xili) = ampZixuk) = v ^ s i n JKXj
k (24)
where y2 is added as a normalizing factor In this case both the longitudinal and
transverse mode shapes satisfy the geometric boundary conditions for a simple string
in axial and transverse vibration
Eq(22) is recast into a compact matrix form as
ff(i) = $i(xili)5i(t) (25)
where $i(x l) is the matrix of the tether mode shapes defined as
0 0
^i(xiJi) =
^ X bull bull bull ^XA
0
0 0
0
G 5R 3 x n f
(26)
and nfj = nx- + nyi + nZi Note $J(XJJ) is a function of the spatial coordinate x
and not of yi or ZJ However it is also a function of the tether length parameter li
16
which is time-varying during deployment and retrieval For this reason care must be
exercised when differentiating $J(XJJ) with respect to time such that
d d d k) = Jtregi(Xi li) = ^ 7 $ t ( ^ i k)Xi + gf$i(xigt li)k- (2-7)
Since x represents the nominal rate of change of the position of the elemental mass
drrii it is equal to j (deploymentretrieval rate) Let t ing
( 3 d
dx~ + dT)^Xili^ ( 2 8 )
one arrives at
$i(xili) = $Di(xili)ii (29)
The time varying generalized coordinate vector 5 in Eq(25) is composed of
the longitudinal and transverse components ie
ltM) G s f t n f i x l (210)
where 5X^ 5y- and Sz^ are nx- x 1 ny x 1 and nz- x 1 size vectors and correspond to
$ X and $ Z respectively
213 Rotation angles and transformations
The matrix T j in Eq(21) represents a rotation transformation from the body
fixed frame Fj to the inertial frame FQ ie FQ = T j F j It is defined by the 3-2-1
ordered sequence of elementary rotations[46]
1 P i tch F[ = Cf ( a j )F 0
2 R o l l F = C f (3j)F = C7f (3j)Cf ( a j )F 0
3 Yaw F = C7( 7 i )F = C7(7l)cf^)Cf(ai)FQ = Q F 0
17
where
c f (A) =
-Sai Cai
0 0
and
C(7i ) =
0 -s0i] = 0 1 0
0 cpi
i 0 0 = 0
0
(211)
(212)
(213)
Here Eq(211) corresponds to a rotation of ctj (pitch) about the inertial axis ZQ (orbit
normal) of frame FQ resulting in a new intermediate frame F It is then followed by
a roll rotation fa about the axis y of F- giving a second intermediate frame F-
Final ly a spin rotation 7J about the axis z of F- is given resulting in the body
fixed frame Fj for the i lt l link Since all the three rotation matrices are orthogonal
it follows that FQ = C~1Fi = Cj F and hence
C CPi Cai ~ CH Sai + SrYi SPi Sai 5 7 i Sampi + Cli SPi Cci CPi Sai C1i Cai + 5 7 i SPi ~ S1 Ca + C 7 bull Sp Sai
(214) SliCPi CliCPi
Here Cx and Sx are abbreviations for cos(x) and sin(x) respectively
214 Inertial velocity of the ith link
mdash Lett ing pj = fj - f $j(5j and differentiating Eq(21) with respect to time gives
m- D j + T j f j + T^Si + T j pound + T j $ j 5 j (215)
Each of the terms appearing in Eq(215) must be addressed individually Since
fj represents the rigid body motion within the link it only has meaning for the
deployable tether and since yi = Zj = 0 for al l tether links (i even) fj = Ifi where 1
is the unit vector 1 0 0T For the case of rigid links (i odd) fj = 0
18
Evaluating the derivative of each angle in the rotation matrix gives
T^i9i = Taiai9i + Tppigi + T^fagi (216)
where Tx = J j I V Collecting the scaler angular velocity terms and defining the set
Vi = deg^gt A) 7i^ Eq(216) can be rewritten as
T i f t = Pi9i)fji (217)
where Pi(gi) is a column matrix defined by
Pi(9i)Vi = [Taigi^p^iT^gilffi (218)
Inserting Eq(29) and Eq(217) into Eq(215) and defining Sj = 2 + leads to
215 Cylindrical orbital coordinates
The Di term in Eq(219) describes the orbital motion of the t h l ink and
is composed of the three Cartesian coordinates DXi A ^ DZi However over the
cycle of one orbit DXi and Dyi vary dramatically by around the order of Earths
radius This must be avoided since large variations in the coordinates can cause
severe truncation errors during their numerical integration For this reason it is
more convenient to express the Cartesian components in terms of more stationary
variables This is readily accomplished using cylindrical coordinates
However it is only necessary to transform the first l inks orbital components
ie Lgti since only D is a generalized coordinates From Figure 2-2 it is apparent
(219)
that
(220)
19
Figure 2-2 Vector components in cylindrical orbital coordinates
20
Eq(220) can be rewritten in matrix form as
cos(0i) 0 0 s i n ( 0 i ) 0 0
0 0 1
Dri
h (221)
= DTlDSv
where Dri is the inplane radial distance 9 the true anomaly and DZl the out-of-
plane distance normal to the orbital plane The total derivative of D wi th respect
to time gives in column matrix form
1 1 1 1 J (222)
= D D l D 3 v
where [cos(^i) -Dnsm(8i) 0 ]
(223)
For all remaining links ie i = 2 N
cos(0i) - A - j S i n ^ i ) 0 sin(^i) D r i c o s ( 0 i ) 0
0 0 1
DTl = D D i = I d 3 i = 2N (224)
where is the n x n identity matrix and
2N (225)
216 Tether deployment and retrieval profile
The deployment of the tether is a critical maneuver for this class of systems
Cr i t i ca l deployment rates would cause the system to librate out of control Normally
a smooth sinusoidal profile is chosen (S-curve) However one would also like a long
tether extending to ten twenty or even a hundred kilometers in a reasonable time
say 3-4 orbits This is usually difficult or impossible to accomplish solely wi th one
S-curve profile Often a composite S-curve-Steady-S-curve profile is adopted Thus
21
the deployment scheme for the t h tether can be summarized as
k = Tr l - cos (^r(ti - to-)) 1 ltUlt h 2 r V A V
k = Vov tiilttiltt2i
k = IT j 1 - C deg S lkitl ^ ^ lt i lt ^
(226)
where to- tt are the ini t ia l and final times of the deployment or retrieval maneuver
respectively A t j = t - tQ mdash tt mdash t2- and the steady deployment velocity VQ is
calculated based on the continuity of l and l at the specified intermediate times t
and to For the case of
and
Ai = i(ti) - ^(t 0 -) = - ^ ( t 2 i ) (227)
2Ali + li(tf) -h(tQ)
V = lUli l^lL (228)
1 tf-tQ v y
A t = (229)
22 Kinetics and System Energy
221 Kinetic energy
The kinetic energy of the ith link is given by
Kei = lJm Rdrnfi^Ami (230)
Setting Eq(219) in a matrix-vector form gives the relation
4 ^ ^ PM) T i $ 4 TiSi]i[t (231)
22
where
m
k J
x l (232)
and nq = -nfj + 7 is the number of generalized coordinates in each link Inserting
Eq(231) into Eq(230) and integrating the kinetic energy for the t h l ink can be
rewritten as
1 T
(233)
where Mti is the links symmetric mass matrix
Mt =
m i D D T D D DDTPifgidmi) DDTTi J ^dm D n T Sidm rn
sym fP(9i)Pi(9i)dmi J P[g^T^idm J P ( pound ) T ^ m sym sym J ^f^idm f ^f^dmi sym sym sym J sfsidmi
pound Sfttiqxnq
(234)
and mi is the mass of the Ith link The kinetic energy for the entire system ie N
bodies can now be stated as
1 bull T bull 1 bull T Ke = J2 MtA = o laquo Mtqt 2 ^ -i
i=l (235)
where qt = qj^^ bull bull -qfNT a n d Mt is a block diagonal decoupled mass matrix
of the system with Mf- on its diagonal
(236) Mt =
0 0
0 Mt2
0
0 0
MH
0 0 0
0 0 0 bull MtN
222 Simplification for rigid links
The mass matrix given by Eq(234) simplifies considerably for the case of rigid
23
links and is given as
Mti
m D D l
T D D l DD^Piiffidmi) 0
sym 0
0
fP(fi)Pi(i)dmi 0 0 0
0 Id nfj
0 m~
i = 1 3 5 N
(237)
Moreover when mult iplying the matrices in the (11) block term of Eq(237) one
gets
D D l
1 D D l = 1 0 0 0 D n
2 0 (238) 0 0 1
which is a diagonal matrix However more importantly i t is independent of 6 and
hence is always non-singular ensuring that is always invertible
In the (12) block term of M ^ it is apparent that f fidmi is simply the
definition of the centre of mass of the t h rigid link and hence given by
To drrii mdash miff (239)
where fcmi is the position vector of the centre of mass of link i For the case of the
(22) block term a little more attention is required Expanding the integrand using
Eq(218) where ltj = fj gives
fj Ta Tafi fj T Q Tpfi f- T Q T 7 i f j
sym sym
Now since each of the terms of Eq(240) is a scaler or a 1 x 1 matrix it is therefore
equal to its trace ie sum of its diagonal elements thus
tr(ffTai
TTaifi) tr(ffTai
TTpfi) tr(ff Ta^T^fi)
fjT^Tpn rfTp^fi (240)
P(rj)Pj(fi) = syrri trifjTpTTpfi) t r ^ T ^ T ^
sym sym tr(ff T 7 T r 7 f j) J
(241)
Using two properties of the trace function[47] namely
tr(ABC) = tr(BCA) = trCAB) (242)
24
and
j tr(A) =tr(^J A^j (243)
where A B and C are real matrices it becomes quite clear that the (22) block term
of Eq(237) simplifies to
P(fi)Pii)dmi =
tr(Tai
TTaiImi) tTTai
TTpImi) t r (T Q r T 7 7 m ) trTaTTpImi) tr(Tp
TT7Jmi) sym sym sym
(244)
where Im- is an alternate form of the inertia tensor of the t h link which when
expanded is
Imlaquo = I nfi dm bull i sym
sym sym
J x^dmi J Xydmi XZdm sym J yi2dmi f yiZidm
sym J z^dmi J
xVi -L VH sym ijxxi + lyy IZZJ)2
(245)
Note this is in terms of moments of inertia and products of inertia of the t f l rigid
link A similar simplification can be made for the flexible case where T is replaced
with Qi resulting in a new expression for Im
223 Gravitational potential energy
The gravitational potential energy of the ith link due to the central force law
can be written as
V9i f dm f
-A mdash = -A bullgttradelti RdmA
dm
inn Lgti + T j ^ | (246)
where z = 3986 x 105 km 3s 2 (Earths gravitational constant) Expanding binomially
and retaining up to third order terms gives
Vgi ~ ~ D T 2D [J gfgidrrii + 23f T j gidm mdashDT
D2 i l I I Df Tt I grffdmiTiDi
(2-lt 7)
25
where D = D is the magnitude of the inertial position vector to the frame F The
integrals in the above expression are functions of the flexible mode shapes and can be
evaluated through symbolic manipulation using an existing package such as Maple
V In turn Maple V can translate the evaluated integrals into F O R T R A N and store
the code in a file Furthermore for the case of rigid bodies it can be shown quite
readily that
j gjgidrrii = j rjfidrrn = [IXXi + Iyy + IZZi)2 (248)
Similarly the other integrals in Eq(234) can be integrated symbolically
224 Strain energy
When deriving the elastic strain energy of a simple string in tension the
assumptions of high tension and low amplitude of transverse vibrations are generally
valid Consequently the first order approximation of the strain-displacement relation
is generally acceptable However for orbiting tethers in a weak gravitational-gradient
field neglecting higher order terms can result in poorly modelled tether dynamics
Geometric nonlinearities are responsible for the foreshortening effects present in the
tether These cannot be neglected because they account for the heaving motion along
the longitudinal axis due to lateral vibrations The magnitude of the longitudinal
oscillations diminish as the tether becomes shorter[35]
W i t h this as background the tether strain which is derived from the theory
of elastic vibrations[3248] is given by
(249)
where Uj V and W are the flexible deformations along the Xj and Z direction
respectively and are obtained from Eq(25) The square bracketed term in Eq(249)
represents the geometric nonlinearity and is accounted for in the analysis
26
The total strain energy of a flexible link is given by
Ve = I f CiEidVl (250)
Substituting the stress-strain relationship 0 = E(poundi the strain energy is now given
by
Vei=l-EiAi J l l d x h (251)
where E^ is the tethers Youngs modulus and A is the cross-sectional area The
tether is assumed to be uniform thus EA which is the flexural stiffness of the
tether is assumed to be constant Eq(251) is evaluated symbolically for arbitrary
number of modes
225 Tether energy dissipation
The evaluation of the energy dissipation due to tether deformations remains
a problem not well understood even today Here this complex phenomenon is repshy
resented through a simplified structural damping model[32] In addition the system
exhibits hysteresis which also must be considered This is accomplished using an
augmented complex Youngs modulus
EX^Ei+jEi (252)
where Ej^ is the contribution from the structural damping and j = The
augmented stress-strain relation is now given as
a = Epoundi = Ei(l+jrldi)ei (253)
where 77^ = EjE^ ltC 1 is defined as the structural damping coefficient determined
experimentally [49]
27
If poundi is a harmonic function with frequency UJQ- then
jet = (254) wo-
Substituting into Eq(253) and rearranging the terms gives
e + I mdash I poundi = o-i + ad (255)
where ad and a are the stress with and without structural damping respectively
Now the Rayleigh dissipation function[50] for the ith tether can be expressed as
(ii) Jo
_ 1 EjAiVdi rU (2-56^ 2 w 0 Jo
The strain rate ii is the time derivative of Eq(249) The generalized external force
due to damping can now be written as Qd = Q^ bull bull bull QdNT where
dWd
^ = - i p = 2 4 ( 2 5 7 )
= 0 i = 13 N
23 0(N) Form of the Equations of Mot ion
231 Lagrange equations of motion
W i t h the kinetic energy expression defined and the potential energy of the
whole system given by Pe = ]Cn=l( 5i+ ej) the equations of motion can be obtained
quite readily using the Lagrangian principle
where q is the set of generalized coordinates to be defined in a later section
Substituting Eqs(235247251 and 257) into Eq(258) leads to the familiar
matr ix form of the non-linear non-autonomous coupled equations of motion for the
28
system
Mq t)t+ F(q q t) = Q(q q t) (259)
where M(q t) is the nonlinear symmetric mass matrix and F(q q t) is the forcing
function which can be written in matrix form as
Q(ltfgt Qi 0 is the vector of non-conservative generalized forces including control inshy
puts acting on the system A detailed expansion of Eq(260) in tensor notation is
developed in Appendix I
Numerical solution of Eq(259) in terms of q would require inversion of the
mass matrix However M is a full matrix of size nqq x nqq where nqq = Nnq is the
total number of generalized coordinates Therefore direct inversion of M would lead
to a large number of computation steps of the order nqq^ or higher The objective
here is to develop an order-N O(N) algorithm for obtaining M _ 1 that minimizes
the computational effort This is accomplished through the following transformations
ini t ia l ly developed for the stationkeeping case by Pradhan et al [30]
232 Generalized coordinates and position transformation
During the derivation of the energy expressions for the system the focus was
on the decoupled system ie each link was considered independent of the others
Thus each links energy expression is also uncoupled However the constraints forces
between two links must be incorporated in the final equations of motion
Let
e Unqxl (261) m
be the set of coupled coordinates of the t h l ink such that q = qT q2 bull bull bull 0T i s the
29
systems generalized coordinates Let qt- be the set of auxiliary decoupled coordinates
defined by Eq(232) such that qt = qj^ qj^ Qtj^1 bull The only difference between
qtj and q is the presence of D against d is the position of F wi th respect to
the inertial frame FQ whereas d is defined as the offset position of F relative to and
projected on the frame It can be expressed as d = dXidyidZi^ For the
special case of link 1 D = d
From Figure 2-1 can written as
Di = A - l + T i _ i J i _ i + di) + T V i S i - i f t - i + dXi)8i- (262)
Denoting KAj = $j(Zj + dx- ) Eq(262) can be rewritten in summation form as
i ampi = H ( T i - 1 ^ + T i - l K A i - l - l + T j - l ^ - l ) bull ( 2 - 6 3 )
3 = 1
Introducing the index substitution k mdash j mdash 1 into Eq(263)
i-1 D = ^2 (Tfc_ilt4 + TkKAk6k + Tkilk) + T^di (264)
k=l
since d = D TQ = 1^ and IQ DQ SQ are null vectors Recasting Eq(264) in
matr ix form leads to i-1
Qt Jfe=l
where
and
For the entire system
T fc-1 0 TkKAk
0 0 0 0 0 0 0 0 0 0 0 0 -
r T i - i 0 0 0 0 0 0 0 0 0 G
0 0 0 1
Qt = RpQ
G R n lt x l
(265)
(266)
G Unq x l (267)
(268)
30
where
RP
R 1 0 0 R R 2 0 R RP
2 R 3
0 0 0
R RP R PN _ 1 RN J
(269)
Here RP is the lower block triangular matrix which relates qt and q
233 Velocity transformations
The two sets of velocity coordinates qti and qi are related by the following
transformations
(i) First Transformation
Differentiating Eq(262) with respect to time gives the inertial velocity of the
t h l ink as
D = A - l + Ti_iK + + K A i _ i ^ _ i
+ T j _ i ^ + Ti-ii + T i - i K A i - i ^ + T j - i K A ^ i ^ i (270)
(271) (
Defining K A D = dddXi+1KAi K A L = d d K A and hi = di+1 +lfi +KA^i
results in K A ^ i = K A D j i o ^ + K A L i _ i i _ i
= KADi_iiTdi + K A L j _ i ^ _ i
Inserting Eq(271) into Eq(270) and rearranging gives
= A _ + T i ^ d g + K A D ^ x ^ i ^ ^ + Pi-^hi-iH-i
+ T i _ i K A i _ i 5 _ i + T i_ i2 + K A L j _ i 5 j _ i 4 _ i (272)
Using Eq(272) to define recursively - D j _ i and applying the index substitution
c = j mdash 1 as in the case of the position transformation it follows that
Di = ^ T ^ i K D ^ ^ + Pk(hk)ffk + T f c K A f c ^ + TfcKLfc4 Jfc=l (273)
+ T j _ i K D j _ i lt i ii
31
where K D j _ i = + K A D ^ i ^ - i r 7 and K L j = i + K A L j 5 j Finally by following a
procedure similar to that used for the position transformation it can be shown that
for the entire system
t = Rvil (2-74)
where Rv is given by the lower block diagonal matrix
Here
Rv
Rf 0 0 R R$ 0 R Rl RJ
0 0 0
T3V nN-l
R bullN
e Mnqqxl
Ri =
T i _ i K D i _ i TiKAi T j K L j 0 0 0 0 bullo 0 0 0 0 0 0 0
e K n lt x l
RS
T i-1 0 0 o-0 0 0 0 0 0
0 0 0 1
G Mnq x l
(275)
(276)
(277)
(ii) Second Transformation
The second transformation is simply Eq(272) set into matrix form This leads
to the expression for qt as
(278)
where Id Pihi) T j K A j T j K L j 0 0 0 0 0 0
0 0
0 0
0 0
G $lnq x l
Thus for the entire system
(279)
pound = Rnjt + Rdq
= Vdnqq - R nrlRdi (280)
32
where
and
- 0 0 0 RI 0 0
Rn = 0 0
0 0
Rf 0 0 0 rgtd K2 0
Rd = 0 0 Rl
0 0 0
RN-1 0
G 5R n w x l
0 0 0 e 5 R N ^ X L
(281)
(282)
234 Cylindrical coordinate modification
Because of the use of cylindrical coorfmates for the first body it is necessary
to slightly modify the above matrices to account for this coordinate system For the
first link d mdash D = D ^ D S L thus
Qti = 01 =gt DDTigi (283)
where
D D T i
Eq(269) now takes the form
DTL 0 0 0 d 3 0 0 0 ID
0 0 nfj 0
RP =
P f D T T i R2 0 P f D T T i Rp
2 R3
iJfDTTi i^
x l
0 0 0
R PN-I RN
(284)
(285)
In the velocity transformation a similar modification is necessary since D =
DD^DS^ Mak ing the substitution it can be shown that Eq(276) and Eq(279) now
33
takes the form for i = 1
T)V _ pn _ r t j mdash rt^ mdash 0 0 0 0
0 0 0 0 0 0 0 0
(286)
The rest of the matrices for the remaining links ie i = 2 N are unchanged
235 Mass matrix inversion
Returning to the expression for kinetic energy substitution of Eq(274) into
Eq(235) gives the first of two expressions for kinetic energy as
Ke = -q RvT MtRv (287)
Introduction of Eq(280) into Eq(235) results in the second expression for Ke as
1X 2q (hnqq-RnrlRd Mt (idnqq - R nrlRd (288)
Thus there are two factorizations for the systems coupled mass matrix given as
M = RV MtRv
M (hnqq - RnrlRd Mt ( i d n q q - R n r 1 R d
(289)
(290)
Inverting Eq(290)
r-1 _ ( rgtd~^ M 1 = (Ra) l d n q q - Rn)Mf1 (Rd)~l(Idnqq - Rn) (291)
Since both Rd and Mt are block diagonal matrices their inverse is simply the inverse
34
of each block on the diagonal ie
Mr1
h 0 0
0 0
0 0
0 0 0
0
0
0
tN
(292)
Each block has the dimension nq x nq and is independent of the number of bodies
in the system thus the inversion of M is only linearly dependent on N ie an O(N)
operation
236 Specification of the offset position
The derivation of the equations of motion using the O(N) formulation requires
that the offset coordinate d be treated as a generalized coordinate However this
offset may be constrained or controlled to follow a prescribed motion dictated either by
the designer or the controller This is achieved through the introduction of Lagrange
multipliers[51] To begin with the multipliers are assigned to all the d equations
Thus letting f = F - Q Eq(259) becomes
Mq + f = P C A (293)
where A G K 3 - 7 V x l is the vector of Lagrange multipliers and Pc G s R n 9 x 3 A r is the
permutation matrix assigning the appropriate Aj to its corresponding d equation
Inverting M and pre-multiplying both sides by PcT gives
pcTM-pc
= ( dc + PdegTM-^f) (294)
where dc is the desired offset acceleration vector Note both dc and PcTM lf are
known Thus the solution for A has the form
A pcTM-lpc - l dc + PC1 M (295)
35
Now substituting Eq(295) into Eq(293) and rearranging the terms gives
q = S(f Ml) = -M~lf + M~lPc
which is the new constrained vector equation of motion with the offset specified by
dc Note that pre-multiplying M~l by P c T provides the rows of M _ 1 corresponding
to the d equations Similarly post-multiplying by Pc leads the columns of the matrix
corresponding to the d equations
24 Generalized Control Forces
241 Preliminary remarks
The treatment of non-conservative external forces acting on the system is
considered next In reality the system is subjected to a wide variety of environmental
forces including atmospheric drag solar radiation pressure Earth s magnetic field as
well as dissipative devices to mention a few However in the present model only
the effect of active control thrusters and momemtum-wheels is considered These
generalized forces wi l l be used to control the attitude motion of the system
In the derivation of the equations of motion using the Lagrangian procedure
the contribution of each force must be properly distributed among all the generalized
coordinates This is acheived using the relation
nu ^5 Qk=E Pern ^ (2-97)
m=l deg Q k
where Qk and qk are the kth elements of Q and ltf respectively[51] Fem is the
mth external force acting on the system at the location Rm from the inertial frame
and nu is the total number of actuators Derivation of the generalized forces due to
thrusters and momentum wheels is given in the next two subsections
pcTM-lpc (dc + PcTM- Vj (2-96)
36
242 Generalized thruster forces
One set of thrusters is placed on each subsatellite ie al l the rigid bodies
except the first (platform) as shown in Figure 2-3 They are capable of firing in the
three orthogonal directions From the schematic diagram the inertial position of the
thruster located on body 3 is given as
Defining
R3 = di + Tid2 + T2(hi + KA2lt52) + T 3 f c m 3
= d + dfY + T 3 f c m 3
df = Tjdj+i + Tj+1lj+ii + KAj+i5j+i)
the thruster position for body 5 is given as
(298)
(299)
R5 = di+ dfx + d 3 + T5fcm
Thus in general the inertial position for the thruster on the t h body can be given
(2100)
as i - 2
k=l i1 cm (2101)
Let the column matrix
[Q i iT 9 R-K=rR dqj lA n i
WjRi A R i (2102)
where i = 3 5 N and j mdash 1 2 i The vector derivative of R wi th respect to
the scaler components qk is stored in the kth column of Thus for the case of
three bodies Eq(297) for thrusters becomes
1 Q
Q3
3 -1
T 3 T t 2 (2103)
37
Satellite 3
mdash gt -raquo cm
D1=d1
mdash gt
a mdash gt
o
Figure 2-3 Inertial position of subsatellite thruster forces
38
where ft = 0TtaTt^T is the thrust vector for the t h l ink (tether i-1) Here
Ttn and Tta are the thrust forces along the inplane m and out-of-plane Z directions i Pi
respectively For the case of 5-bodies
Q
~Qgt1
0
0
( T ^ 2 ) (2104)
The result for seven or more bodies follows the pattern established by Eq(2104)
243 Generalized momentum gyro torques
When deriving the generalized moments arising from the momemtum-wheels
on each rigid body (satellite) including the platform it is easier to view the torques as
coupled forces From Figure 2-4 it is apparent that for the ith l ink the generalized
forces constituting the couple are
(2105)
where
Fei = Faij acting at ei = eXi
Fe2 = -Fpk acting at e 2 = C a ^ l
F e 3 = F7ik acting at e 3 = eyj
Expanding Eq(2105) it becomes clear that
3
Qk = ^ F e i
i=l
d
dqk (2106)
which is independent of Ri- Thus the moments Mma = 2FaieXi Mm^ = 2FpeXi
and Mmi = 2FlieVi Transforming the coordinates to the body fixed frame using
the rotation matrix the generalized force due to the momentum-wheels on link i can
39
Figure 2-4 Coupled force representation of a momentum-wheel on a rigid body
40
be written as
d d d Qi = Td bull ^ T i t M m a - Tik bull mdashT^Mmp + T ^ bull T J j M ^
QiMmi
where Mm = Mma M m Mmi T
lt3 T J ^ T - T i f c bull ^ T i pound T ^ - ^ T z j
(2107)
(2108)
Note combining the thruster forces and momentum-wheel torques a compact
expression for the generalized force vector for 5 bodies can be written as
Q
Q T 3 0 Q T 5 0
0 0 0
0 lt33T3 Qyen Q5
3T5 0
0 0 0 Q4T5 0
0 0 0 Q5
5T5
M M I
Ti 2 m 3
Quu (2109)
The pattern of Eq(2109) is retained for the case of N links
41
3 C O M P U T E R I M P L E M E N T A T I O N
31 Prel iminary Remarks
The governing equations of motion for the TV-body tethered system were deshy
rived in the last chapter using an efficient O(N) factorization method These equashy
tions are highly nonlinear nonautonomous and coupled In order to predict and
appreciate the character of the system response under a wide variety of disturbances
it is necessary to solve this system of differential equations Although the closed-
form solutions to various simplified models can be obtained using well-known analytic
methods the complete nonlinear response of the coupled system can only be obtained
numerically
However the numerical solution of these complicated equations is not a straightshy
forward task To begin with the equations are stiff[52] be they have attitude and
vibration response frequencies separated by about an order of magnitude or more
which makes their numerical integration suceptible to accuracy errors i f a properly
designed integration routine is not used
Secondly the factorization algorithm that ensures efficient inversion of the
mass matrix as well as the time-varying offset and tether length significantly increase
the size of the resulting simulation code to well over 10000 lines This raises computer
memory issues which must be addressed in order to properly design a single program
that is capable of simulating a wide range of configurations and mission profiles It
is further desired that the program be modular in character to accomodate design
variations with ease and efficiency
This chapter begins with a discussion on issues of numerical and symbolic in-
42
tegration This is followed by an introduction to the program structure Final ly the
validity of the computer program is established using two methods namely the conshy
servation of total energy and comparison of the response results for simple particular
cases reported in the literature
32 Numerical Implementation
321 Integration routine
A computer program coded in F O R T R A N was written to integrate the equashy
tions of motion of the system The widely available routine used to integrate the difshy
ferential equations D G E A R [ 5 3 ] is based on Gears predictor-corrector method[54]
It is well suited to stiff differential equations as i t provides automatic step-size adjustshy
ment and error-checking capabilities at each iteration cycle These features provide
superior performance over traditional 4^ order Runge-Kut ta methods
Like most other routines the method requires that the differential equations be
transformed into a first order state space representation This is easily accomplished
by letting
thus
X - ^ ^ f ^ ) =ltbullbull) (3 -2 )
where S(fMdc) is defined by Eq(296) Eq(32) represents the new set of 2nqq
first order differential equations of motion
322 Program structure
A flow chart representing the computer programs structure is presented in
Figure 3-1 The M A I N program is responsible for calling the integration routine It
also coordinates the flow of information throughout the program The first subroutine
43
called is INIT which initializes the system variables such as the constant parameters
(mass density inertia etc) and the ini t ial conditions I C (attitude angles orbital
motion tether elastic deformation etc) from user-defined data files Furthermore all
the necessary parameters required by the integration subroutine D G E A R (step-size
tolerance) are provided by INIT The results of the simulation ie time history of the
generalized coordinates are then passed on to the O U T P U T subroutine which sorts
the information into different output files to be subsequently read in by the auxiliary
plott ing package
The majority of the simulation running time is spent in the F C N subroutine
which is called repeatedly by D G E A R F C N generates the fv(Xi) vector defined in
Eq(32) It is composed of several subroutines which are responsible for the comshy
putation of all the time-varying matrices vectors and scaler variables derived in
Chapter 2 that comprise the governing equations of motion for the entire system
These subroutines have been carefully constructed to ensure maximum efficiency in
computational performance as well as memory allocation
Two important aspects in the programming of F C N should be outlined at this
point As mentioned earlier some of the integrals defined in Chapter 2 are notably
more difficult to evaluate manually Traditionally they have been integrated numershy
ically on-line (during the simulation) In some cases particularly stationkeeping the
computational cost is quite reasonable since the modal integrals need only be evalushy
ated once However for the case of deployment or retrieval where the mode shape
functions vary they have to be integrated at each time-step This has considerable
repercussions on the total simulation time Thus in this program the integrals are
evaluated symbolically using M A P L E V[55] Once generated they are translated into
F O R T R A N source code and stored in I N C L U D E files to be read by the compiler
Al though these files can be lengthy (1000 lines or more) for a large number of flexible
44
INIT
DGEAR PARAM
IC amp SYS PARAM
MAIN
lt
DGEAR
lt
FCN
OUTPUT
STATE ACTUATOR FORCES
MOMENTS amp THRUST FORCES
VIBRATION
r OFFS DYNA
ET MICS
Figure 3-1 Flowchart showing the computer program structure
45
modes they st i l l require less time to compute compared to their on-line evaluation
For the case of deployment the performance improvement was found be as high as
10 times for certain cases
Secondly the O(N) algorithm presented in the last section involves matrices
that contained mostly zero terms (RP Rv and Mplusmn) The sparse structure of these mashy
trices can be used advantageously to improve the computational efficiency by avoiding
multiplications of zero elements The result is a quick evaluation of fv(Xt) Moreshy
over there is a considerable saving of computer memory as zero elements are no
longer stored
Final ly the C O N T R O L subroutine which is part of F C N is designed to calshy
culate the appropriate control forces torques and offset dynamics which are required
to stabilize the librational ( A T T I T U D E ) and vibrational ( V I B R A T I O N ) motion of
the system
33 Verification of the Code
The simulation program was checked for its validity using two methods The
first verified the conservation of total energy in the absence of dissipation The second
approach was a direct comparison of the planar response generated by the simulation
program with those available in the literature A s numerical results for the flexible
three-dimensional A^-body model are not available one is forced to be content with
the validation using a few simplified cases
331 Energy conservation
The configuration considered here is that of a 3-body tethered system with
the following parameters for the platform subsatellite and tether
46
h =
1091430 -8135 328108 -8135 8646050 27116 328108 27116 8286760
kg bull m 2 platform inertia
kg bull m 2 end-satellite inertia 200 0 0
0 400 0 0 0 400
m = 90000 kg mass of the space station platform
m2 = 500 kg mass of the end-satellite
EfAt = 61645 N tether elastic stiffness
pt mdash 49 kg km tether linear density
The tether length is taken to remain constant at 10 km (stationkeeping case)
A n in i t ia l disturbance in pitch and roll of 2deg and 1deg respectively is given to both
the rigid bodies and the flexible tether In addition the tether is ini t ia l ly deflected
by 45 m in the longitudinal direction at its end and 05 m in both the inplane and
out-of-plane directions in the first mode The tether attachment point offset at the
platform (satellite 1) end is 1 m in all the three directions ie d2 = 11 lTm The
attachment point at the subsatellite (satellite 2) end is 1 m in the x direction ie
fcm3 = 10 0T
The variation in the kinetic and potential energies is presented for the above-
mentioned case in Figure 3-2(a) As seen from the plot there is a continuous mutual
exchange between the kinetic and potential energies however the variation in the
total energy remains zero (Figure 3-2b) It should be noted that after 1 orbit the
variation of both kinetic and potential energy does not return to 0 This is due to
the attitude and elastic motion of the system which shifts the centre of mass of the
tethered system in a non-Keplerian orbit
Similar results are presented for the 5-body case ie a double pendulum (Figshy
ure 3-3) Here the system is composed of a platform and two subsatellites connected
47
STSS configuration 3-Body a 1(0)=a 2(0)=a t(0)=2deg P1(0)=p2(0)=p(0)=1deg 5x(0)=45m 8y(0)=52(0)=05m dx(0)=dy(0)=d2(0)=1m Stationkeeping lt=10km
Variation of Kinetic and Potential Energy
(a) Time (Orbits) Percent Variation in Total Energy
OOEOh
-50E-12
-10E-11 I i i
0 1 2 (b) Time (Orbits)
Figure 3-2 Kinetic and potential energy transfer for the three-body platform based tethered satellite system (a) variation of kinetic and potential energy (b) percent change in total energy of system
48
STSS configuration 5-Body a 1(0)=a 2(0)=a t 1(0)=2deg a 3(0)=a t 2(0)=25deg P1(0)=p2(0)=p3(0)=pt1(0)=pt2(0)=1lt
5x(0)=5m 8y(0)=5z(0)=05m dx(0)=dy(0)=dz(0)=0 Stationkeeping 1 =12=1 Okm
(a)
20E6
10E6r-
00E0
-10E6
-20E6
Variation of Kinetic and Potential Energy
-
1 I I I 1 I I
I I
AP e ~_
i i 1 i i i i
00E0
UJ -50E-12 U J
lt
Time (Orbits) Percent Variation in Total Energy
-10E-11h
(b) Time (Orbits)
Figure 3 -3 Kinet ic and potential energy transfer for the five-body platform based tethered satellite system (a) variation of kinetic and potential enshyergy (b) percent change in total energy of system
49
in sequence by two tethers The two subsatellites have the same mass and inertia
ie 7713 = 7 7 i 2 a n d I3 = I21 a n d are separated by identical tether segments each 10
k m in length The platform has the same properties as before however there is no
tether attachment point offset present in the system (d = fcmi = 0)
In al l the cases considered energy was found to be conserved However it
should be noted that for the cases where the tether structural damping deployment
retrieval attitude and vibration control are present energy is no longer conserved
since al l these features add or subtract energy from the system
332 Comparison with available data
The alternative approach used to validate the numerical program involves
comparison of system responses with those presented in the literature A sample case
considered here is the planar system whose parameters are the same as those given
in Section 331 A n ini t ia l pitch disturbance of 2deg is imparted to both the platform
and the tether together with an ini t ial tether deformation of 08 m and 001 m in
the longitudinal and transverse directions respectively in the same manner as before
The tether length is held constant at 5 km and it is assumed to be connected to the
centre of mass of the rigid platform and subsatellite The response time histories from
Ref[43] are compared with those obtained using the present program in Figures 3-4
and 3-5 Here ap and at represent the platform and tether pitch respectively and B i
C i are the tethers longitudinal and transverse generalized coordinates respectively
A comparison of Figures 3-4 and 3-5 clearly shows a remarkable correlation
between the two sets of results in both the amplitude and frequency The same
trend persisted even with several other comparisons Thus a considerable level of
confidence is provided in the simulation program as a tool to explore the dynamics
and control of flexible multibody tethered systems
50
a p(0) = 2deg 0(0) = 2C
6(0) = 08 m
0(0) = 001 m
Stationkeeping L = 5 km
D p y = 0 D p 2 = 0
a oo
c -ooo -0 01
Figure 3 - 4 Simulation results for the platform based three-body tethered system originally presented in Ref[43]
51
STSS configuration 3-Body a1(0)=a2(0)=at(0)=2deg p1(0)=(32(0)=(3t(0)=0 8x(0)=08m 5y(0)=001m 5Z(0)=0 dx(0)=dy(0)=dz(0)=0 Stationkeeping l=5km
Satellite 1 Pitch Angle Tether Pitch Angle
i -i i- i i i i i i d
1 2 0 1 2 Time (Orbits) Time (Orbits)
Longitudinal Vibration
^MAAAAAAAWVWWWWVVVW
V bull i bull i I i J 00 05 10 15 20
Time (Orbits) Transverse Vibration
Time (Orbits)
Figure 3-5 Simulation results for the platform based three-body tethered system obtained using the present computer program
52
4 D Y N A M I C SIMULATION
41 Prel iminary Remarks
Understanding the dynamics of a system is critical to its design and develshy
opment for engineering application Due to obvious limitations imposed by flight
tests and simulation of environmental effects in ground based facilities space based
systems are routinely designed through the use of numerical models This Chapter
studies the dynamical response of two different tethered systems during deployment
retrieval and stationkeeping phases In the first case the Space platform based Tethshy
ered Satellite System (STSS) which consists of a large mass (platform) connected to
a relatively smaller mass (subsatellite) with a long flexible tether is considered The
other system is the O E D I P U S B I C E P S configuration involving two comparable mass
satellites interconnected by a flexible tether In addition the mission requirement of
spin about an arbitrary axis the tether axis for O E D I P U S - A C and cartwheeling or
spin about the orbit normal for the proposed B I C E P S mission is accounted for
A s mentioned in Chapter 2 the tether flexibility is modelled using the asshy
sumed mode discretization method Although the simulation program can account
for an arbitrary number of vibrational modes in each direction only the first mode
is considered in this study as it accounts for most of the strain energy[34] and hence
dominates the vibratory motion The parametric study considers single as well as
double pendulum type systems with offset of the tether attachment points The
systems stability is also discussed
42 Parameter and Response Variable Definitions
The system parameters used in the simulation unless otherwise stated are
53
selected as follows for the STSS configuration
1091430 -8135 328108 -8135 8646050 27116 328108 27116 8286760
h =
h = 200 0 0
0 400 0 0 0 400
bull m = 90000 kg (mass of the space platform)
kg bull m 2 (platform inertia) 760 J
kg bull m 2 (subsatellite inertia)
bull 7772 = 500 kg (mass of the subsatellite)
bull EtAt = 61645 N (tether stiffness)
bull Pt = 49 k g k m (tether density)
bull Vd mdash 05 (tether structural damping coefficient)
bull fcm^ = l 0 0 r m (tether attachment point at the subsatellite)
The response variables are defined as follows
bull ai0 satellite 1 (platform) pitch and roll angles respectively
bull CK2 2 satellite 2 (subsatellite) pitch and roll angles respectively
bull at fit- tether pitch and roll angles respectively
bull If tether length
bull d = dx dy dzT- tether attachment position relative to satellite 1 (platform)
mdash
bull S = 8X 8y Sz tether flexible generalized coordinates in the longitudinal x
inplane transverse y and out-of-plane transverse z directions respectively
The attitude angles and are measured with respect to the Local Ve r t i c a l -
Loca l Horizontal ( L V L H ) frame A schematic diagram illustrating these variable
is presented in Figure 4-1 The system is taken to be in a nominal circular orbit at
an altitude of 289 km with an orbital period of 903 minutes
54
Satellite 1 (Platform) Y a w
Figure 4-1 Schematic diagram showing the generalized coordinates used to deshyscribe the system dynamics
55
43 Stationkeeping Profile
To facilitate comparison of the simulation results a reference case based on
the three-body STSS system with zero offset (d = 0 r c m 3 = 0) and a tether length
of It = 20 km is considered first (Figure 4-2) The system is subjected to an ini t ial
disturbance in pitch and roll of 2deg and 1deg respectively to all the three bodies
In addition an ini t ia l longitudinal deflection of 14 m from the tethers unstretched
position is given together with a transverse inplane and out-of-plane deflection of 1
m at the tethers mid-length From the attitude response given in Figure 4-2(a) it is
apparent that the three bodies oscillate about their respective equilibrium positions
Since coupling between the individual rigid-body dynamics is absent (zero offset) the
corresponding librational frequencies are unaffected Satellite 1 (Platform) displays
amplitude modulation arising from the products of inertia The result is a slight
increase of amplitude in the pitch direction and a significantly larger decrease in
amplitude in the roll angle
Figure 4-2(b) clearly shows coupling of the tethers attitude motion with its
flexibility dynamics The tether vibrates in the axial direction about its equilibrium
position (at approximately 138 m) with two vibration frequencies namely 012 Hz
and 31 x 1 0 - 4 Hz The former is the first longitudinal flexible mode which is dissishy
pated through the structural damping while the latter arises from the coupling with
the pitch motion of the tether Similarly in the transverse direction there are two
frequencies of oscillations arising from the flexible mode and its coupling wi th the
attitude motion As expected there is no apparent dissipation in the transverse flexshy
ible motion This is attributed to the weak coupling between the longitudinal and
transverse modes of vibration since the transverse strain in only a second order effect
relative to the longitudinal strain where the dissipation mechanism is in effect Thus
there is very litt le transfer of energy between the transverse and longitudinal modes
resulting in a very long decay time for the transverse vibrations
56
STSS configuration 3-Body a1(0)=a2(0)=a(0)=2deg (31(0)=p2(0)=p(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=dz(0)=0 Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
0 1 2 3 4 5 0 1 2 3 4 5 Time (Orbits) Time (Orbits)
Satellite 2 Pitch Angle Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-2 Stationkeeping dynamics of the three-body STSS configuration withshyout offset (a) attitude response
57
STSS configuration 3-Body a1(0)=cx2(0)=cx(0)=2deg P1(0)=p2(0)=p(0)=1deg 5x(0)=14m8y(0)=82(0)=1m dx(0)=dy(0)=dz(0)=0 Stationkeeping lt=20km
Tether Longitudinal Vibration 16| 1 1 1 1 1 r-
Time (Orbits) Tether In-Plane Transverse Vibration
imdash 1 mdash 1 mdash 1 mdash 1 mdashr
r I I 1 1 1 I 1 1 1 1 I 1 bull ^ bull bull I I J 0 1 2 3 4 5
Time (Orbits) Tether Out-of-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-2 Stationkeeping dynamics of the three-body STSS configuration withshyout offset (b) vibration response
58
Introduction of the attachment point offset significantly alters the response of
the attitude motion for the rigid bodies W i t h a i m offset along the local vertical
at the platform (dx = 1 m) and fcm^ = 100 m coupling between the tether
and end-bodies is established by providing a lever arm from which the tether is able
to exert a torque that affects the rotational motion of the rigid end-bodies (Figure
4-3) Bo th the platform and subsatellite now oscillate at the same frequency as the
tether whose motion remains unaltered In the case of the platform there is also an
amplitude modulation due to its non-zero products of inertia as explained before
However the elastic vibration response of the tether remains essentially unaffected
by the coupling
Providing an offset along the local horizontal direction (dy = 1 m) results in
a more dramatic effect on the pitch and roll response of the platform as shown in
Figure 4-4(a) Now the platform oscillates about its new equilibrium position of
-90deg The roll motion is also significantly disturbed resulting in an increase to over
10deg in amplitude On the other hand the rigid body dynamics of the tether and the
subsatellite as well as the tether flexible motion remain the same (Figure 4-4b)
Figure 4-5 presents the response when the offset of 1 m at the platform end is
along the z direction ie normal to the orbital plane The result is a larger amplitude
pitch oscillation about the reference equilibrium position as shown in Figure 4-5(a)
while the roll equilibrium is now shifted to approximately 90deg Note there is little
change in the attitude motion of the tether and end-satellites however there is a
noticeable change in the out-of-plane transverse vibration of the tether (Figure 4-5b)
which may be due to the large amplitude platform roll dynamics
Final ly by setting a i m offset in all the three direction simultaneously the
equil ibrium position of the platform in the pitch and roll angle is altered by approxishy
mately 30deg (Figure 4-6a) However the response of the system follows essentially the
59
STSS configuration 3-Body a1(0)=cx2(0)=cxt(0)=2deg p1(0)=(32(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=1m dy(0)=dz(0)=0 Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
Time (Orbits) Time (Orbits) Tether Pitch Angle Tether Roll Angle
L bull I I bull bull lt bull bull bull raquo bull bull bull bull -I h I I I I i H 0 1 2 3 4 5 0 1 2 3 4 5
Time (Orbits) Time (Orbits) Satellite 2 Pitch Angle Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-3 Stationkeeping dynamics of the three-body STSS configuration with offset along the local vertical (a) attitude response
60
STSS configuration 3-Body a 1(0)=a 2(0)=a t(0)=2deg (31(0)=p2(0)=pt(0)=1deg 5x(0)=14m5y(0)=5z(0)=1m dx(0)=1mdy(0)=dz(0)=0 Stationkeeping lt=20km
Tether Longitudinal Vibration 1 6 i mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash mdash i mdash i mdash mdash mdash mdash mdash i mdash mdash lt -
Time (Orbits)
Tether In-Plane Transverse Vibration
y bull i i i i i i i i i i i i i i i i i 0 1 2 3 4 5
Time (Orbits)
Tether Out-of-Plane Transverse Vibration r 1 1 1 1 1 1 1 1 1 1 1 1 1 1 q
h i I i i I i i i i I i i i i 3
0 1 2 3 4 5 (b) Time (Orbits)
Figure 4-3 Stationkeeping dynamics of the three-body STSS configuration with offset along the local vertical (b) vibration response
61
STSS configuration 3-Body a1(0)=cx2(0)=o(0)=2deg (31(0)=p2(0)=p(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dz(0)=0 dy(0)=1m Stationkeeping lt=20km
Satellite 1 Pitch Angle i 1 i
Satellite 1 Roll Angle
1 2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 4 Time (Orbits)
Tether Roll Angle
CD
edT
1 2 3 Time (Orbits)
Satellite 2 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 2 Roll Angle
CD
C O
(a)
Figure 4-
1 2 3 4 Time (Orbits)
1 2 3 4 Time (Orbits)
4 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal (a) attitude response
62
STSS configuration 3-Body a1(0)=cx2(0)=ot(0)=2deg p1(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dz(0)=0 dy(0)=1m Stationkeeping lt=20km
Tether Longitudinal Vibration 16| 1 gt 1 bull 1 1 1
0 1 2 3 4 5
Time (Orbits) Tether In-Plane Transverse Vibration
i mdash 1 mdash 1 mdash 1 mdash 1 mdash i mdash mdash lt mdash 1 mdash 1 mdash i mdash mdash 1 mdash 1 mdash
V I I bull bull bull I i i i I i i 1 1 3
0 1 2 3 4 5 Time (Orbits)
Tether Out-of-Plane Transverse Vibration r 1 1 1 1 1 1 1 1 1 1 1 gt bull | q
V i 1 1 1 1 1 mdash 1 mdash 3
0 1 2 3 4 5 (b) Time (Orbits)
Figure 4-4 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal (b) vibration response
63
STSS configuration 3-Body a1(0)=o2(0)=at(0)=2deg (31(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=0 dz(0)=1m Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 4 Time (Orbits)
Tether Roll Angle
1 2 3 4 Time (Orbits)
Satellite 2 Pitch Angle
2 3 4 Time (Orbits)
Satellite 2 Roll Angle
(a)
Figure 4-
1 2 3 4 Time (Orbits)
2 3 4 Time (Orbits)
5 Stationkeeping dynamics of the three-body STSS configuration with offset along the orbit normal (a) attitude response
64
STSS configuration 3-Body a1(0)=cx2(0)=ot(0)=2deg P1(0)=p2(0)=(3(0)=1deg 8x(0)=14m8y(0)=82(0)=1m dx(0)=dy(0)=0 d2(0)=1m Stationkeeping lt=20km
Tether Longitudinal Vibration 1 6 i mdash r mdash i mdash mdash i mdash i mdash mdash i mdash i mdash mdash i mdash mdash i mdash 1 mdash 1 mdash i mdash mdash lt mdash lt mdash lt mdash r
Time (Orbits) Tether In-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-5 Stationkeeping dynamics of the three-body STSS configuration with offset along the orbit normal (b) vibration response
65
same trend with only minor perturbations in the transverse elastic vibratory response
of the tether (Figure 4-6b)
44 Tether Deployment
Deployment of the tether from an ini t ial length of 200 m to 20 km is explored
next Here the deployment length profile is critical to ensure the systems stability
It is not desirable to deploy the tether too quickly since that can render the tether
slack Hence the deployment strategy detailed in Section 216 is adopted The tether
is deployed over 35 orbits with the sinusoidal acceleration and deceleration occurring
over a 4 km length ( A ^ 2 ) with the remaining deployment at a constant velocity of
V0 = 146 ms
Initially the tether is stretched only slightly S = 1304 x 1 0 _ 3 0 0 m
with al l the other states of the system ie pitch roll and transverse displacements
remaining zero The response for the zero offset case is presented in Figure 4-7 For
the platform there is no longer any coupling with the tether hence it oscillates about
the equilibrium orientation determined by its inertia matrix However there is st i l l
coupling between the subsatellite and the tether since rcm^ mdash 100 m resulting
in complete domination of the subsatellite dynamics by the tether Consequently
the pitch motion ini t ial ly grows by almost 50deg but eventually subsides as the tether
deployment rate decreases It is interesting to note that the out-of-plane motion is not
affected by the deployment This is because the Coriolis force which is responsible
for the tether pitch motion does not have a component in the z direction (as it does
in the y direction) since it is always perpendicular to the orbital rotation (Q) and
the deployment rate (It) ie Q x It
Similarly there is an increase in the amplitude of the transverse elastic oscilshy
lations of the tether again due to the Coriolis effect (Figure 4-7b) Moreover as the
66
STSS configuration 3-Body a1(0)=a2(0)=cxt(0)=2deg P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m5y(0)=52(0)=1m dx(0)=dy(0)=d2(0)=1m Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
i i i i i i i i i i i 0 1 2 3 4 5 0 1 2 3 4 5
Time (Orbits) Time (Orbits) Tether Pitch Angle Tether Roll Angle
l i i i i -I h i i i i I i i i i l- i i i i l i i i i I i i- i i H 0 1 2 3 4 5 0 1 2 3 4 5
Time (Orbits) Time (Orbits) Satellite 2 Pitch Angle Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-6 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal local vertical and orbit normal (a) attitude response
67
STSS configuration 3-Body a1(0)=cc2(0)=a(0)=2deg p1(0)=p2(0)=(3t(0)=1deg 5x(0)=14m6y(0)=52(0)=1m dx(0)=dy(0)=d2(0)=1m Stationkeeping lt=20km
Tether Longitudinal Vibration 1 6 1 mdash i mdash i mdash i mdash mdash i mdash 1 mdash 1 mdash mdash mdash i mdash mdash 1 mdash 1 mdash 1 mdash i mdash 1 mdash 1 mdash 1 mdash 1 mdash r
0 1 2 3 4 5
Time (Orbits) Tether In-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-6 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal local vertical and orbit normal (b) vibration response
68
STSS configuration 3-Body a(0)=a2(0)=ot(0)=0 P1(0)=p2(0)=pt(0)=0 8X(0)=1304 x 103m 8y(0)=8z(0)=0 dx(0)=dy(0)=dz(0)=0 Deployment
l=02km to 20km in 35 Orbits
Tether Length Profile
E
Satellite 1 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 1 Roll Angle mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash [ mdash i mdash i mdash i mdash i mdash | mdash
1 2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 4 Time (Orbits)
Tether Roll Angle
CD
oo ax
i imdashimdashimdashImdashimdashimdashimdashr-
_ 1 _ L
1 2 3 4 Time (Orbits)
Satellite 2 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 2 Roll Angle
-50
(a) 1 2 3 4
Time (Orbits) 1 2 3 4
Time (Orbits)
Figure 4-7 Deployment dynamics of the three-body STSS configuration without offset (a) attitude response
69
STSS configuration 3-Body a1(0)=a2(0)=at(0)=0 p1(0)=p2(0)=(3t(0)=0 8X(0)=1304x 103m 8y(0)=52(0)=0 dx(0)=dy(0)=d2(0)=0 Deployment
l=02km to 20km in 35 Orbits
Tether Longitudinal Vibration
to
-02
(b)
Figure 4-7
2 3 Time (Orbits)
Tether In-Plane Transverse Vibration
0 1 1 bull 1 1
1 2 3 4 Time (Orbits)
Tether Out-of-Plane Transverse Vibration
5
1 1 1 1 1 i i | i i | ^
_ i i i i L -I
1 2 3 4 5 Time (Orbits)
Deployment dynamics of the three-body STSS configuration without offset (b) vibration response
70
tether elongates its longitudinal static equilibrium position also changes due to an
increase in the gravity-gradient tether tension It may be pointed out that as in the
case of attitude motion there are no out-of-plane tether vibrations induced
When the same deployment conditions are applied to the case of 1 m offset in
the x direction (Figure 4-8) coupling effects are re-introduced The platform pitch
exceeds -100deg during the peak deployment acceleration However as the tether pitch
motion subsides the platform pitch response returns to a smaller amplitude oscillation
about its nominal equilibrium On the other hand the platform roll motion grows
to over 20deg in amplitude in the terminal stages of deployment Longitudinal and
transverse vibrations of the tether remain virtually unchanged from the zero offset
case except for minute out-of-plane perturbations due to the platform librations in
roll
45 Tether Retrieval
The system response for the case of the tether retrieval from 20 km to 200 m
with an offset of 1 m in the x and z directions at the platform end is presented in
Figure 4-9 Here the Coriolis force induced during retrieval renders the tether unstable
(Figure 4-9a) Consequently the pitch motion of the platform is also destabilized
through coupling The z offset provides an additional moment arm that shifts the
roll equilibrium to 35deg however this degree of freedom is not destabilized From
Figure 4-9(b) it is apparent that the transverse mode of the tether vibration is also
disturbed producing a high frequency response in the final stages of retrieval This
disturbance is responsible for the slight increase in the roll motion for both the tether
and subsatellite
46 Five-Body Tethered System
A five-body chain link system is considered next To facilitate comparison with
71
STSS configuration 3-Body a(0)=a2(0)=at(0)=0 P1(0)=P2(0)=P(0)=0 5X(0)=1304 x 103m 8y(0)=8z(0)=0 dx(0)=1mdy(0)=d2(0)=0 Deployment
lt=02km to 20km in 35 Orbits
Tether Length Profile
E
Satellite 1 Pitch Angle
1 2 3 Time (Orbits)
Satellite 1 Roll Angle
bull bull bull i bull bull bull bull i
1 2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 Time (Orbits)
Tether Roll Angle
i i i i J I I i _
1 2 3 4 Time (Orbits)
Satellite 2 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 2 Roll Angle 3E-3h
D )
T J ^ 0 E 0 eg
CQ
-3E-3
(a) 1 2 3 4
Time (Orbits) 1 2 3 4
Time (Orbits)
Figure 4-8 Deployment dynamics of the three-body STSS configuration with offset along the local vertical (a) attitude response
7 2
STSS configuration 3-Body a1(0)=a2(0)=at(0)=0 p1(0)=P2(0)=(3t(0)=0 Sx(0)=1304 x 103m 5y(0)=5z(0)=0 dx(0)=1mdy(0)=dz(0)=0 Deployment
l=02km to 20km in 35 Orbits
Tether Longitudinal Vibration
Time (Orbits) Tether In-Plane Transverse Vibration
n 1 1 1 f
I I I _j I I I I I 1 1 1 1 1 2 3 4 5
Time (Orbits) Tether Out-of-Plane Transverse Vibration
J 1 I I 1 I I I I I I I I L _ _ I I d 1 2 3 4 5
Time (Orbits)
Deployment dynamics of the three-body STSS configuration with offset along the local vertical (b) vibration response
(b)
Figure 4-8
73
STSS configuration 3-Body a1(0)=a2(0)=at(0)=0 p(0)=P2(0)=(3t(0)=0 Sx(0)=14m8y(0)=82(0)=0 dy(0)=0dx(0)=d2(0)=1m Retrieval
l=20km to 02km in 35 Orbits
Satellite 1 Pitch Angle
(a) Time (Orbits)
Figure 4-9 Retrieval dynamics of the along the local vertical ar
Tether Length Profile
i i i i i i i i i i I 0 1 2 3
Time (Orbits) Satellite 1 Roll Angle
Time (Orbits) Tether Roll Angle
b_ i i I i I i i L _ J
0 1 2 3 Time (Orbits)
Satellite 2 Roll Angle _ mdash | r i i | 1 i i imdashmdashj i 1 r mdash i mdash
04- -
0 1 2 3 Time (Orbits)
three-body STSS configuration with offset d orbit normal (a) attitude response
74
STSS configuration 3-Body a1(0)=a2(0)=ocl(0)=0 (31(0)=f32(0)=pt(0)=0 8x(0)=14m8y(0)=6z(0)=0 dy(0)=0dx(0)=dz(0)=1m Retrieval
l=20km to 02km in 35 Orbits
Tether Longitudinal Vibration
E 10 to
Time (Orbits) Tether In-Plane Transverse Vibration
1 2 Time (Orbits)
Tether Out-of-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-9 Retrieval dynamics of the three-body STSS configuration with offset along the local vertical and orbit normal (b) vibration response
75
the three-body system dynamics studied earlier the chain is extended through the
addition of two bodies a tether and a subsatellite with the same physical properties
as before (Section 42) Thus the five-body system consists of a platform (satellite 1)
tether 1 subsatellite 1 (satellite 2) tether 2 and subsatellite 2 (satellite 3) as shown
in Figure 4-10 The offset of tether 1 at the platform-end is denoted by d2 while that
of tether 2 to subsatellite 1 as 04 In this numerical example fcm^ = fcm^ = 0 ie
tethers 1 and 2 are attached to the centre of mass of subsatellites 1 and 2 respectively
The system response for the case where the tether attachment points coincide
wi th the centres of mass of the rigid bodies ie zero offset (d2 = d mdash 0) is presented
in Figure 4-11 Here ogtt and at2 represent the pitch motion of tethers 1 and 2
respectively whereas 03 is the pitch motions of satellite 3 A similar convention
is adopted for the rol l angle 0 A s before the zero offset eliminates the coupling
between the tethers and satellites such that they are now free to librate about their
static equil ibrium positions However their is s t i l l mutual coupling between the two
tethers This coupling is present regardless of the offset position since each tether is
capable of transferring a force to the other The coupling is clearly evident in the pitch
response of the two tethers (Figure 4 - l l a ) O n the other hand the roll motion appears
uncoupled as the in i t ia l conditions in rol l for the two tethers are identical Thus the
motion is in phase and there is no transfer of energy through coupling A s expected
during the elastic response the tethers vibrate about their static equil ibrium positions
and mutually interact (Figure 4 - l l b ) However due to the variation of tension along
the tether (x direction) they do not have the same longitudinal equilibrium Note
relatively large amplitude transverse vibrations are present particularly for tether 1
suggesting strong coupling effects with the longitudinal oscillations
76
Figure 4-10 Schematic diagram of the five-body system used in the numerical example
77
STSS configuration 5-Body a1(0)=(x2(0)=at1(0)=2deg a3(0)=a12(0)=25deg p1(0)=r32(0)=p3(0)=pt1(0)=pt2(0)=r 82(0)=84(0)=505)05m d2(0)=d4(0)=000 Stationkeeping lt1=lu=10km Pitch Roll
Satellite 1 Pitch Angle Satellite 1 Roll Angle
o 1 2 Time (Orbits)
Tether 1 Pitch and Roll Angle
0 1 2 Time (Orbits)
Tether 2 Pitch and Roll Angle
Time (Orbits) Satellite 2 Pitch and Roll Angle
Time (Orbits) Satellite 3 Pitch and Roll Angle
Time (Orbits) Time (Orbits)
Figure 4-11 Stationkeeping dynamics of the five-body STSS configuration withshyout offset (a) attitude response
78
STSS configuration 5-Body a1(0)=a2(0)=oct1(0)=2deg a3(0)=at2(0)=25o
p1(0)=p2(0)=(33(0)=pt1(0)=pt2(0)=10
62(0)=54(0)=50505m d2(0)=d4(0)=000 Stationkeeping lt1=1 =1 Okm
Tether 1 Longitudinal Vibration Tether 2 Longitudinal Vibration i i 1 i 1 r 1 1 1 1
(b) Time (Orbits) Time (Orbits)
Figure 4-11 Stationkeeping dynamics of the five-body STSS configuration withshyout offset (b) vibration response
79
The system dynamics during deployment of both the tethers in the double-
pendulum configuration is presented in Figure 4-12 Deployment of each tether takes
place from 200 m to 20 km in 35 orbits The pitch motion of the system is similar
to that observed during the three-body case The Coriolis effect causes the tethers
to pitch through a large angle ( laquo 50deg ) which in turn disturbs the platform and
subsatellite due to the presence of offset along the local vertical On the other hand
the roll response for both the tethers is damped to zero as the tether deploys in
confirmation with the principle of conservation of angular momentum whereas the
platform response in roll increases to around 20deg Subsatellite 2 remains virtually
unaffected by the other links since the tether is attached to its centre of mass thus
eliminating coupling Finally the flexibility response of the two tethers presented in
Figure 4-12(b) shows similarity with the three-body case
47 BICEPS Configuration
The mission profile of the Bl-static Canadian Experiment on Plasmas in Space
(BICEPS) is simulated next It is presently under consideration by the Canadian
Space Agency To be launched by the Black Brant 2000500 rocket it would inshy
volve interesting maneuvers of the payload before it acquires the final operational
configuration As shown in Figure 1-3 at launch the payload (two satellites) with
an undeployed tether is spinning about the orbit normal (phase 1) The internal
dampers next change the motion to a flat spin (phase 2) which in turn provides
momentum for deployment (phase 3) When fully deployed the one kilometre long
tether will connect two identical satellites carrying instrumentation video cameras
and transmitters as payload
The system parameters for the BICEPS configuration are summarized below [59 0 0 1
bull I = I2 = 0 366 0 kg bull m 2 (satellite inertias) [ 0 0 392_
bull mi = 7772 = 200 kg (mass of the satellites)
80
STSS configuration 5-Body a(0)=cx2(0)=al1(0)=2deg a3(0)=c (0)=25o
P1(0)=P2(0)=p3(0)=f3t1(0)=Pt2(0)=r 52(0)=84(0)=1304x 10300m d2(0)=d4(0)=1G0m Deployment lt1=le=02km to 20km in 35 Orbits
Pitch Roll Satellite 1 Pitch Angle
Time (Orbits) Satellite 2 Pitch and Roll Angle T 1mdash1 1 1 I I I I | I I I 1 I | I I lt~r
Tether Length Profile i 1 1 1 1 i 1 1 1 1 i 1 1 1 1 i
i i i i i 0 1 2 3 4 5
Time (Orbits) Satellite 1 Roll Angle
0 1 2 3 4 5 Time (Orbits)
Tether 1 amp 2 Roll Angle L i | I I 1 I 1 1 -I
-Ih I- bull I I i i i i I i i i i 1
0 1 2 3 4 5 Time (Orbits)
Satellite 3 Pitch and Roll Angle i- i 1 1 i 1 1 1 1 i 1 1 1 1 i 1 1 1 1 i
L i i i i i i 1 i i mdash i mdash i I i imdashimdashLJ
0 1 2 3 4 5 Time (Orbits)
Figure 4-12 Deployment dynamics of the five-body S T S S configuration with offshyset along the local vertical (a) attitude response
81
STSS configuration 5-Body a1(0)=a2(0)=oct1(0)=2deg a3(0)=at2(0)=25deg P1(0)=P2(0)=P3(0)=Pt1(0)=pt2(0)=1o
82(0)=54(0)=1304 x 10-300m d2(0)=d4(0)=100m Deployment lt1=l2=02km to 20km in 35 Orbits
Tether 1 Longitudinal Vibration Tether 2 Longitudinal Vibration
Time (Orbits) Time (Orbits) Tether 1 Z Tranverse Vibration Tether 2 Z Tranverse Vibration
Figure 4-12 Deployment dynamics of the five-body STSS configuration with offshyset along the local vertical (b) vibration response
82
bull EtAt = 61645 N (tether stiffness)
bull pt = 30 k g k m (tether density)
bull It = 1 km (tether length)
bull rid = 1-0 (tether structural damping coefficient)
The system is in a circular orbit at a 289 km altitude Offset of the tether
attachment point to the satellites at both ends is taken to be 078 m in the x
direction The response of the system in the stationkeeping phase for a prescribed set
of in i t ia l conditions is given by Figure 4-13 As in the case of the STSS configuration
there is strong coupling between the tether and the satellites rigid body dynamics
causing the latter to follow the attitude of the tether However because of the smaller
inertias of the payloads there are noticeable high frequency modulations arising from
the tether flexibility Response of the tether in the elastic degrees of freedom is
presented in Figure 4-13(b) Note the longitudinal vibrations decay quite rapidly
(lt 05 orbits) due to the structural damping however it has vir tual ly no effect on
the transverse oscillations
The unique mission requirement of B I C E P S is the proposed use of its ini t ia l
angular momentum in the cartwheeling mode to aid in the deployment of the tethered
system The maneuver is considered next The response of the system during this
maneuver wi th an ini t ia l cartwheeling rate of 5 deg s is illustrated in Figure 4-14(a) The
in i t ia l tether length is taken to be 10 m and is deployed to 1 km There is an in i t ia l
increase in the pitch motion of the tether however due to the conservation of angular
momentum the cartwheeling rate decreases proportionally to the square of the change
in tether deployment rate The result is a rapid drop in the cartwheeling rate unti l
the system stops rotating entirely and simply oscillates about its new equilibrium
point Consequently through coupling the end-bodies follow the same trend The
roll response also subsides once the cartwheeling motion ceases However
83
BICEPS configuration 3-Body a1(0)=o2(0)=at(0)=2deg P1(0)=P2(0)=pt(0)=1deg 8X(0)=001 m 6y(0)=52(0)=01 m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping l=1 km
Satellite 1 Pitch Angle Satellite 1 Roll Angle - i 1 1 1 1 i - i q r 1 1 -gt
I- I i t i i i i I H 0 1 2 0 1 2
Time (Orbits) Time (Orbits) Tether Pitch Angle Tether Roll Angle
F mdash 1 mdash 1 mdash mdash 1 mdash i mdash 1 mdash 1 mdash mdash 1 mdash = i F mdash 1 mdash mdash mdash 1 mdash i mdash 1 mdash 1 mdash 1 mdash 1 mdash q
(a) Time (Orbits) Time (Orbits)
Figure 4-13 Stationkeeping dynamics of the three-body BICEPS configuration with offset along the local vertical (a) attitude response
84
BICEPS configuration 3-Body a1(0)=o2(0)=at(0)=2deg Pl(0)=p2(0)=(3t(0)=1deg 8X(0)=001 m 8y(0)=82(0)=01 m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping lt=1km
Tether Longitudinal Vibration 1E-2[
laquo = 5 E - 3 |
000 002
Time (Orbits) Tether In-Plane Transverse Vibration
Time (Orbits) Tether Out-of-Plane Transverse Vibration
Time (Orbits)
Figure 4-13 Stationkeeping dynamics of the three-body BICEPS configuration with offset along the local vertical (b) vibration response
85
BICEPS configuration 3-Body a1(0)=o2(0)=at(0)=2deg a1(0)=a2(0)=at(0)=57s Pl(0)=P2(0)=(3t(0)=1deg 8X(0)=115 x 103m 8y(0)=82(0)=0 dx(0)=078m dy(0)=d2(0)=0 Cartwheeling Deployment
lt=10m to 1km in 1 Orbit
Tether Length Profile
E
2000
cu T3
2000
co CD
2000
C D CD
T J
Satellite 1 Pitch Angle
1 2 Time (Orbits)
Satellite 1 Roll Angle
cn CD
33 cdl
Time (Orbits) Tether Pitch Angle
Time (Orbits) Tether Roll Angle
Time (Orbits) Satellite 2 Pitch Angle
1 2 Time (Orbits)
Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-14 Cartwheeling dynamics with deployment for the three-body BICEPS with offset along the local vertical (a) attitude response
86
BICEPS configuration 3-Body a1(0)=a2(0)=at(0)=2deg a (0)=a2(0)=a(0)=57s P1(0)=P2(0)=f3(0)=1deg 8X(0)=115x103m 8y(0)=8z(0)=0 dx(0)=078m dy(0)=d2(0)=0 Cartwheeling Deployment
l=1 Om to 1 km in 1 Orbit
5E-3h
co
0E0
025
000 E to
-025
002
-002
(b)
Figure 4-14
Tether Longitudinal Vibration
Time (Orbits) Tether In-Plane Transverse Vibration
Time (Orbits) Tether Out-of-Plane Transverse Vibration
Time (Orbits)
Cartwheeling dynamics with deployment for the three-body BICEPS with offset along the local vertical (b) vibration response
87
as the deployment maneuver is completed the satellites continue to exhibit a small
amplitude low frequency roll response with small period perturbations induced by
the tether elastic oscillations superposed on it (Figure 4-14b)
48 OEDIPUS Spinning Configuration
The mission entitled Observation of Electrified Distributions of Ionospheric
Plasmas - A Unique Strategy ( O E D I P U S ) also consists of a 1 k m tethered system with
two similar end-masses It was first flown in a suborbital flight in 1989 ( O E D I P U S A )
followed by a recent study in November 1995 ( O E D I P U S - C ) in which the University
of Br i t i sh Columbia was one of the participants Dynamically O E D I P U S represents a
unique system never encountered before Launched by a Black Brant rocket developed
at Br is to l Aerospace L t d the system ie the end-bodies together wi th the tether
spins about the longitudinal axis of the tether to achieve stabilized alignment wi th
Earth s magnetic field
The response of the system undergoing a spin rate of j = l deg s is presented
in Figure 4-15 using the same parameters as those outlined for the B I C E P S conshy
figuration in Section 47 Again there is strong coupling between the three bodies
(two satellites and tether) due to the nonzero offset The spin motion introduces an
additional frequency component in the tethers elastic response which is transferred
to the l ibrational motion of the satellites A s the spin rate increases to 1 0 deg s (Figure
4-16) the amplitude and frequency of the perturbations also increase however the
general character of the response remains essentially the same
88
OEDIPUS configuration 3-Body a1(0)=a2(0)=at(0)=2deg i(0)^[2(0H(0)=Ys p1(0)=p2(0)=p(0)=r 8X(0)=12 x 103m 8y(0)=82(0)=01m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping l=1 km
Satellite 1 Pitch Angle Satellite 1 Roll Angle mdash i 1 1 1 a r 1 1 1 r
(a) Time (Orbits) Time (Orbits)
Figure 4-15 Spin dynamics (7 = ldegs) of the three-body OEDIPUS configuration with offset along the local vertical (a) attitude response
89
OEDIPUS configuration 3-Body a1(0)=a2(0)=cct(0)=2deg Y1(0)=Y2(0)=Yt(0)=l7s (31(0)=p2(0)=pt(0)=1deg 8X(0)=12 x 103m 8y(0)=5z(0)=01 m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping lt=1 km
1E-2
to
OEO
E 00
to
Tether Longitudinal Vibration
005
Time (Orbits) Tether In-Plane Transverse Vibration
Time (Orbits) Tether Out-of-Plane Transverse Vibration
Time (Orbits)
Figure 4-15 Spin dynamics (7 = l deg s ) of the three-body O E D I P U S configuration with offset along the local vertical (b) vibration response
90
OEDIPUS configuration 3-Body a1(0)=a2(0)=a(0)=2deg Yi(0)=Y2(0)=rt(0)=107s 81(0)=p2(0)=pt(0)=1deg 8X(0)=12 x 103m 8y(0)=82(0)=01m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping lt=1km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-16 Spin dynamics (7= 10degs) of the three-body OEDIPUS configura tion with offset along the local vertical (a) attitude response
91
OEDIPUS configuration 3-Body a1(0)=a2(0)=ot(0)=2deg Yi(0)H(0)4(0)laquo107s P1(0)=P2(0)=pt(0)=1deg 8X(0)=12 x 103m 8y(0)=82(0)=01m dx(0)=078m dy(0)=dz(0)=0 Stationkeeping lt=1 km
Tether Longitudinal Vibration
Figure 4-16 Spin dynamics (7 = 10degs) of the three-body OEDIPUS configurashytion with offset along the local vertical (b) vibration response
92
5 ATTITUDE AND VIBRATION CONTROL
51 Att itude Control
511 Preliminary remarks
The instability in the pitch and roll motions during the retrieval of the tether
the large amplitude librations during its deployment and the marginal stability during
the stationkeeping phase suggest that some form of active control is necessary to
satisfy the mission requirements This section focuses on the design of an attitude
controller with the objective to regulate the librational dynamics of the system
As discussed in Chapter 1 a number of methods are available to accomplish
this objective This includes a wide variety of linear and nonlinear control strategies
which can be applied in conjuction with tension thrusters offset momentum-wheels
etc and their hybrid combinations One may consider a finite number of Linear T ime
Invariant (LTI) controllers scheduled discretely over different system configurations
ie gain scheduling[43] This would be relevant during deployment and retrieval of
the tether where the configuration is changing with time A n alternative may be
to use a Linear T ime Varying ( L T V ) model in conjuction with an adaptive control
scheme where on-line parametric identification may be used to advantage[56] The
options are vir tually limitless
O f course the choice of control algorithms is governed by several important
factors effective for time-varying configurations computationally efficient for realshy
time implementation and simple in character Here the thrustermomentum-wheel
system in conjuction with the Feedback Linearization Technique ( F L T ) is chosen as
it accounts for the complete nonlinear dynamics of the system and promises to have
good overall performance over a wide range of tether lengths It is well suited for
93
highly time-varying systems whose dynamics can be modelled accurately as in the
present case
The proposed control method utilizes the thrusters located on each rigid satelshy
lite excluding the first one (platform) to regulate the pitch and rol l motion of the
tether to which it is attached ie the thrusters located on satellite 2 (link 3) regshy
ulates the attitude motion of tether 1 (link 2) O n the other hand the l ibrational
motion of the rigid bodies (platform and subsatellite) is controlled using a set of three
momentum wheels placed mutually perpendicular to each other
The F L T method is based on the transformation of the nonlinear time-varying
governing equations into a L T I system using a nonlinear time-varying feedback Deshy
tailed mathematical background to the method and associated design procedure are
discussed by several authors[435758] The resulting L T I system can be regulated
using any of the numerous linear control algorithms available in the literature In the
present study a simple P D controller is adopted arid is found to have good perforshy
mance
The choice of an F L T control scheme satisfies one of the criteria mentioned
earlier namely valid over a wide range of tether lengths The question of compushy
tational efficiency of the method wi l l have to be addressed In order to implement
this controller in real-time the computation of the systems inverse dynamics must
be executed quickly Hence a simpler model that performs well is desirable Here
the model based on the rigid system with nonlinear equations of motion is chosen
Of course its validity is assessed using the original nonlinear system that accounts
for the tether flexibility
512 Controller design using Feedback Linearization Technique
The control model used is based on the rigid system with the governing equa-
94
tions of motion given by
Mrqr + fr = Quu (51)
where the left hand side represents inertia and other forces while the right hand side
is the generalized external force due to thrusters and momentum-wheels and is given
by Eq(2109)
Lett ing
~fr = fr~ QuU (52)
and substituting in Eq(296)
ir = s(frMrdc) (53)
Expanding Eq(53) it can be shown that
qr = S(fr Mrdc) - S(QU Mr6)u (5-4)
mdash = Fr - Qru
where S(QU Mr0) is the column matrix given by
S(QuMr6) = [s(Qu(l)Mr0) 5 (Q u ( 2 ) M r | 0 ) S(QU nu) M r | 0 ) ] (55)
and Qu-i) is the i ^ column of Qu Extract ing only the controlled equations from
Eq(54) ie the attitude equations one has
Ire = Frc ~~ Qrcu
(56)
mdash vrci
where vrc is the new control input required to regulate the decoupled linear system
A t this point a simple P D controller can be applied ie
Vrc = qrcd +KvQrcd-Qrc) + Kp(Qrcd~ Qrc)- (57)
Eq(57) represents the secondary controller with Eq(54) as the primary controller
Kp and Kv are the proportional and derivative gain matrices with qrcdi Qrcd
an(^ qrcd
95
as the desired acceleration velocity and position vector for the attitude angles of each
controlled body respectively Solving for u from Eq(56) leads to
u Qrc Frc ~ vrcj bull (58)
5 13 Simulation results
The F L T controller is implemented on the three-body STSS tethered system
The choice of proportional and derivative matrix gains is based on a desired settling
time of ts mdash 05 orbit and a damping factor of = 07 for both the pitch and roll
actuators in each body Given O and ts it can be shown that
-In ) 0 5 V T ^ r (59)
and hence kn bull mdash10
V (510)
where kp and kVj are the diagonal elements of the gain matrices Kp and Kv in
Eq(57) respectively[59]
Note as each body-fixed frame is referred directly to the inertial frame FQ
the nominal pitch equilibrium angle is zero only when measured from the L V L H frame
(OJJ mdash 0) However it is 6 when referred to FQ Hence the desired position vector
qrC(l is set equal to 0(t) ie the time-varying orbital angle for the pitch motion
and zero for the roll and yaw rotations such that
qrcd 37Vxl (511)
96
The desired velocity and acceleration are then given as
e U 3 N x l (512)
E K 3 7 V x l (513)
When the system is in a circular orbit qrc^ = 0
Figure 5-1 presents the controlled response of the STSS system defined in
Chapter 4 in the stationkeeping phase with offset d2 = 111T m and If = 20
km A s mentioned earlier the F L T controller is based on the nonlinear rigid model
Note the pitch and roll angles of the rigid bodies as well as of the tether are now
attenuated in less than 1 orbit (Figure 5-la) Consequently the longitudinal vibration
response of the tether is free of coupling arising from the librational modes of the
tether This leaves only the vibration modes which are eventually damped through
structural damping The pitch and roll dynamics of the platform require relatively
higher control moments Ma and Mpi respectively (Figure 5-lb) since they have to
overcome the extra moments generated by the offset of the tether attachment point
Furthermore the platform demands an additional moment to maintain its orientation
at zero pitch and roll angle since it is not the nominal equilibrium position The nonshy
zero static equilibrium of the platform arises due to its products of inertia On the
other hand the tether pitch and roll dynamics require only a small ini t ia l control
force of about plusmn 1 N which eventually diminishes to zero once the system is
97
and
o 0
Qrcj mdash
o o
0 0
Oiit) 0 0
1
N
N
1deg I o
STSS configuration 3-Body a1(0)=a2(0)=at(0)=2lt
P1(0)=p2(0)=pt(0)=1 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=dz(0)=1m Controlled Stationkeeping lt=20km
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
05 10 15 Time (Orbits)
Sat 2 Pitch and Roll Angles
20
CO CD
T J
cdT
cT
- mdash mdash 1 1 1 1 1 I - 1
Pitch -
A Roll -
A -
1 bull bull
-
1 Time (Orbits)
Tether Y Tranverse Vibration
1 2 Time (Orbits)
Tether Z Tranverse Vibration
gt to N
to
Time (Orbits) Time (Orbits) Longitudinal Vibration
145h
(a) Time (Orbits)
Figure 5-1 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local horizontal local vertical and orbit normal (a) attitude and vibration response
98
1
STSS configuration 3-Body a1(0)=ct2(0)=cct(0)=20
P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=d2(0)=1m Controlled Stationkeeping lt=20km
Sat 1 Pitch Control Moment Sat 1 Roll Control Moment
Figure 5-1 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local horizontal local vertical and orbit normal (b) control actuator time histories
99
stabilized Similarly for the end-body a very small control torque is required to
attenuate the pitch and roll response
If the complete nonlinear flexible dynamics model is used in the feedback loop
the response performance was found to be virtually identical with minor differences
as shown in Figure 5-2 For example now the pitch and roll response of the platform
is exactly damped to zero with no steady-state error as against the minor almost
imperceptible deviation for the case of the controller basedon the rigid model (Figure
5-la) In addition the rigid controller introduces additional damping in the transverse
mode of vibration where none is present when the full flexible controller is used
This is due to a high frequency component in the at motion that slowly decays the
transverse motion through coupling
As expected the full nonlinear flexible controller which now accounts the
elastic degrees of freedom in the model introduces larger fluctuations in the control
requirement for each actuator except at the subsatellite which is not coupled to the
tether since f c m 3 = 0 The high frequency variations in the pitch and roll control
moments at the platform are due to longitudinal oscillations of the tether and the
associated changes in the tension Despite neglecting the flexible terms the overall
controlled performance of the system remains quite good Hence the feedback of the
flexible motion is not considered in subsequent analysis
When the tether offset at the platform is restricted to only 1 m along the x
direction a similar response is obtained for the system However in this case the
steady-state error in the platforms rigid body motion is much smaller (Figure 5-3a)
In addition from Figure 5-3(b) the platforms control requirement is significantly
reduced
Figure 5-4 presents the controlled response of the STSS deploying a tether
100
STSS configuration 3-Body a1(0)=a2(0)=oct(0)=2deg p1(0)=P2(0)=p(0)=1deg 8x(0)=14m5y(0)=82(0)=1m dx(0)=dy(0)=dz(0)=1m Controlled Stationkeeping l=20km
Tether Pitch and Roll Angles
c n
T J
cdT
Sat 1 Pitch and Roll Angles 1 r
00 05 10 15 Time (Orbits)
Sat 2 Pitch and Roll Angles
c n CD
cdT a
o T J
CQ
- - I 1 1 1 I 1 1 1 1
P i t c h
R o l l
bull
1 -
0 1 2 Time (Orbits)
Tether Y Tranverse Vibration
0 1 2 Time (Orbits)
Tether Z Tranverse Vibration
E gt
10 N
CO
Time (Orbits) Time (Orbits) Longitudinal Vibration
145h
(a) Time (Orbits)
Figure 5-2 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear flexible FLT controller with offset along the local horizontal local vertical and orbit normal (a) attitude and vibration response
101
STSS configuration 3-Body a1(0)=a2(0)=cxt(0)=2deg 31(0)=p2(0)=p(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=d2(0)=1m Controlled Stationkeeping lt=20km
r i i i t i mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash mdash J 0 1 2 0 1 2
Time (Orbits) Time (Orbits) Tether Pitch Control Thrust Tether Roll Control Thruster
(b) Time (Orbits) Time (Orbits)
Figure 5-2 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear flexible FLT controller with offset along the local horizontal local vertical and orbit normal (b) control actuator time histories
102
STSS configuration 3-Body a1(0)=o2(0)=a(0)=2deg P1(0)=p2(0)=(3t(0)=1deg 8x(0)=14m5y(0)=6z(0)=1m dx(0)=1mdy(0)=dz(0)=0 Controlled Stationkeeping lt=20km
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
05 10 15 Time (Orbits)
Sat 2 Pitch and Roll Angles
CD
eg CQ
0 1 2 Time (Orbits)
Tether Y Tranverse Vibration
1 2 Time (Orbits)
Tether Z Tranverse Vibration
gt CO
Time (Orbits) Time (Orbits) Longitudinal Vibration
145h
(a) Time (Orbits)
Figure 5-3 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local vertical (a) attitude and vibration response
103
STSS configuration 3-Body a1(0)=o2(0)=al(0)=2o
P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=1mdy(0)=dz(0)=0 Controlled Stationkeeping lt=20km
Sat 1 Pitch Control Moment 11 1 1 r
Sat 1 Roll Control Moment
0 1 2 Time (Orbits)
Tether Pitch Control Thrust
0 1 2 Time (Orbits)
Sat 2 Pitch Control Moment
0 1 2 Time (Orbits)
Tether Roll Control Thruster
-07 o 1
Time (Orbits Sat 2 Roll Control Moment
1E-5
0E0F
Time (Orbits)
Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local vertical (b) control actuator time histories
104
STSS configuration 3-Body a1(0)=a2(0)=a(0)=2deg Pl(0)=p2(0)=pt(0)=1deg 6X(0)=1304 x 103m 8y(0)=82(0)=0 dx(0)=1mdy(0)=d2(0)=0 Controlled Deployment
l=02km to 20km in 35 Orbits
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
o 1 2 Time (Orbits)
Tether Y Tranverse Vibration
0 1 2 Time (Orbits)
Sat 2 Pitch and Roll Angles
CD CD
CM CO
Time (Orbits) Tether Z Tranverse Vibration
to
150
100
1 2 3 0 1 2
Time (Orbits) Time (Orbits) Longitudinal Vibration
to
(a)
50
2 3 Time (Orbits)
Figure 5 - 4 Deployment dynamics of the three-body STSS using the nonlinear rigid FLT controller with offset along the local vertical (a) attitude and vibration response
105
STSS configuration 3-Body a1(0)=oc2(0)=al(0)=20
P1(0)=p2(0)=pt(0)=1deg 8X(0)=1304 x 103m 5y(0)=52(0)=0 dx(0)=1mdy(0)=dz(0)=0 Controlled Deployment
l=02km to 20km in 35 Orbits
Tether Length Profile i 1
2 0 h
E 2L
Sat 1 Pitch Control Moment t mdash r mdash i mdash i mdash | mdash i mdash i mdash i mdash i mdash | mdash r mdash i mdash i mdash i mdash | mdash r mdash
0 5
1 2 3
Time (Orbits) Sat 1 Roll Control Moment
i i t | t mdash t mdash i mdash [ i i mdash r -
1 2
Time (Orbits) Tether Pitch Control Thrust
1 2 3 Time (Orbits)
Tether Roll Control Thruster
2 E - 5
1 2 3
Time (Orbits) Sat 2 Pitch Control Moment
1 2
Time (Orbits) Sat 2 Roll Control Moment
1 E - 5
E
O E O
Time (Orbits)
Deployment dynamics of the three-body STSS using the nonlinear rigid FLT controller with offset along the local vertical (b) control actuator time histories
106
from 02 km to 20 km in 35 orbits with a i m offset along the tether length A s before
the systems attitude motion is well regulated by the controller However the control
cost increases significantly for the pitch motion of the platform and tether due to the
Coriolis force induced by the deployment maneuver (Figure 5-4b) Initially there is a
sinusoidal increase in the pitch control requirement for both the tether and platform
as the tether accelerates to its constant velocity VQ Then the control requirement
remains constant for the platform at about 60 N m as opposed to the tether where the
thrust demand increases linearly Finally when the tether decelerates the actuators
control input reduces sinusoidally back to zero The rest of the control inputs remain
essentially the same as those in the stationkeeping case In fact the tether pitch
control moment is significantly less since the tether is short during the ini t ia l stages
of control However the inplane thruster requirement Tat acts as a disturbance and
causes 6y to nearly double from the uncontrolled value (Figure 4-8a) O n the other
hand the tether has no out-of-plane deflection
Final ly the effectiveness of the F L T controller during the crit ical maneuver of
retrieval from 20 km to 02 k m in 35 orbits is assessed in Figure 5-5 A s in the earlier
cases the system response in pitch and roll is acceptable however the controller is
unable to suppress the high frequency elastic oscillations induced in the tether by the
retrieval O f course this is expected as there is no active control of the elastic degrees
of freedom However the control of the tether vibrations is discussed in detail in
Section 52 The pitch control requirement follows a similar trend in magnitude as
in the case of deployment with only minor differences due to the addition of offset in
the out-of-plane direction ( d 2 = 1 0 1 ^ m) This is also responsible for the higher
roll moment requirement for the platform control (Figure 5-5b)
107
STSS configuration 3-Body a1(0)=X2(0)=a(0)=2deg P1(0)=p2(0)=pt(0)=1deg 5x(0)=14m8y(0)=52(0)=0 dx(0)=1mdy(0)=dz(0)=0 Controlled Retrieval
l=20km to 02km in 35 Orbits
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
0 1 2 Time (Orbits)
Sat 2 Pitch and Roll Angles
CO CD
cdT
CO CD
T J CM
CO
1 2 Time (Orbits)
Tether Y Tranverse Vibration
Time (Orbits) Tether Z Tranverse Vibration
150
100
1 2 3 0 1 2 Time (Orbits) Time (Orbits)
Longitudinal Vibration
CO
(a)
x 50
2 3 Time (Orbits)
Figure 5 -5 Retrieval dynamics of the three-body STSS using the non-linear rigid FLT controller with offset along the local vertical and orbit normal (a) attitude and vibration response
108
STSS configuration 3-Body a1(0)=ct2(0)=at(0)=2deg p1(0)=p2(0)=(3(0)=1deg 5x(0)=14m5y(0)=82(0)=0 dx(0)=1mdy(0)=dz(0)=0 Controlled Retrieval
l=20km to 02km in 35 Orbits
Tether Length Profile
E
1 2 Time (Orbits)
Tether Pitch Control Thrust
2E-5
1 2 Time (Orbits)
Sat 2 Pitch Control Moment
OEOft
1 2 3 Time (Orbits)
Sat 1 Roll Control Moment
1 2 3 Time (Orbits)
Tether Roll Control Thruster
1 2 Time (Orbits)
Sat 2 Roll Control Moment
1E-5 E z
2
0E0
(b) Time (Orbits) Time (Orbits)
Figure 5-5 Retrieval dynamics of the three-body STSS using the non-linear rigid FLT controller with offset along the local vertical and orbit normal (b) control actuator time histories
109
52 Control of Tethers Elastic Vibrations
521 Prel iminary remarks
The requirement of a precisely controlled micro-gravity environment as well
as the accurate release of satellites into their final orbit demands additional control
of the tethers vibratory motion in addition to its attitude regulation To that end
an active vibration suppression strategy is designed and implemented in this section
The strategy adopted is based on offset control ie time dependent variation
of the tethers attachment point at the platform (satellite 1) A l l the three degrees
of freedom of the offset motion are used to control both the longitudinal as well as
the inplane and out-of-plane transverse modes of vibration In practice the offset
controller can be implemented through the motion of a dedicated manipulator or a
robotic arm supporting the tether which in turn is supported by the platform The
focus here is on the control of elastic deformations during stationkeeping (ie fully
deployed tether fixed length) as it represents the phase when the mission objectives
are carried out
This section begins with the linearization of the systems equations of motion
for the reduced three body stationkeeping case This is followed by the design of
the optimal control algorithm Linear Quadratic Gaussian-Loop Transfer Recovery
( L Q G L T R ) based on the reduced model and its implementation on the full nonlinear
model Final ly the systems response in the presence of offset control is presented
which tends to substantiate its effectiveness
522 System linearization and state-space realization
The design of the offset controller begins with the linearization of the equations
of motion about the systems equilibrium position Linearization of extremely lengthy
(even in matrix form) highly nonlinear nonautonomous and coupled equations of
110
motion presents a challenging problem This is further complicated by the fact the
pitch angle ot is not referred to the L V L H frame Hence the equilibrium pitch angle
is not a constant but is equal to the instantaneous orbital angle 9(t) Two methods
are available to resolve this problem
One may use the non-stationary equations of motion in their present form and
derive a controller based on the Linear T ime Varying ( L T V ) system Alternatively
one can design a Linear T ime Invariant (LTI) controller based on a new set of reduced
governing equations representing the motion of the three body tethered system A l shy
though several studies pertaining to L T V systems have been reported[5660] the latter
approach is chosen because of its simplicity Furthermore the L T I controller design
is carried out completely off-line and thus the procedure is computationally more
efficient
The reduced model is derived using the Lagrangian approach with the genershy
alized coordinates given by
where lty and are the pitch and roll angles of link i relative to the L V L H frame
respectively 5X 8y and 5Z are the generalized coordinates associated with the flexible
tether deformations in the longitudinal inplane and out-of-plane transverse direcshy
tions respectively Only the first mode of vibration is considered in the analysis
The nonlinear nonautonomous and coupled equations of motion for the tethshy
ered system can now be written as
Qred = [ltXlPlltX2P2fixSy5za302gt] (514)
M redQred + fr mdash 0gt (515)
where Mred and frec[ are the reduced mass matrix and forcing term of the system
111
respectively They are functions of qred and qred in addition to the time varying offset
position d2 and its time derivatives d2 and d2 A detailed derivation of Eq(515)
is given in Appendix II A n additional consequence of referring the pitch motion to
a local frame is that the new reduced equations are now independent of the orbital
angle 9 and under the further assumption of the system negotiating a circular orbit
91 remains constant
The nonlinear reduced equations can now be linearized about their static equishy
l ibr ium position For all the generalized coordinates the equilibrium position is zero
wi th the exception of 5X which has a non-zero equilibrium 5XQ The offset position
d2) is linearized about its ini t ial position d2^ Linearizing Eq(515) and recasting
into matr ix form gives
Msqred + CsQred + KsQred + Md^2 + Cdd2 + Kdd2 + fs = 0 (516)
where Ms Cs Ks Md Cd Kd and s are constant matrices Mul t ip ly ing throughout
by Ms
1 gives
qr = -Ms 1Csqr - Ms
lKsqr
- Ms~lCdd22 - Ms-Kdd2 - Ms~Mdd2 - Mg~1 fs r-1 (517)
Defining ud = d2 and v = qTd^T Eq(517) can be rewritten as
Pu t t ing
v =
+
-M~lCs -M~Cd -1
0
-Mg~lMd
Id
0 v + -M~lKs -M~lKd
0 0
-M-lfs
0
MCv + MKv + Mlud +Fs
- 4 ) - ( gt the L T I equations of motion can be recast into the state-space form as
(518)
(519)
x =Ax + Bud + Fd (520)
112
where
A = MC MK Id12 o
24x24
and
Let
B = ^ U 5 R 2 4 X 3 -
Fd = I FJ ) e f t 2 4 1
mdash _ i _ -5
(521)
(522)
(523)
(524)
where x is the perturbation vector from the constant equilibrium state vector xeq
Substituting Eq(524) into Eq(520) gives the linear perturbation state equation
modelling the reduced tethered system as
bullJ x mdash x mdash Ax + Bud + (Fd + Axeq) (525)
It can be shown that for ideal control[60] Fd + Axeq = 0 thus giving the familiar
state-space equation
S=Ax + Bud (526)
The selection of the output vector completes the state space realization of
the system The output vector consists of the longitudinal deformation from the
equil ibrium position SXQ of the tether at xt = h the slope of the tether due to the
transverse deformation at xt = 0 and the offset position d2 from its in i t ia l position
d2Q Thus the output vector y is given by
K(o)sx
V
dx ~ dXQ dy ~ dyQ
d z - d z0 J
Jy
dx da XQ vo
d z - dZQ ) dy dyQ
(527)
113
where lt ( 0 ) = ddxi[$vxi)]Xi=o and lt ( 0 ) = ddxi[4gtw(xi)]Xi=0
523 Linear Quadratic Gaussian control with Loop Transfer Recovery
W i t h the linear state space model defined (Eqs526527) the design of the
controller can commence The algorithm chosen is the Linear Quadratic Gaussian
( L Q G ) estimator based optimal controller[6061] The L Q G is a widely used optimal
controller with its theoretical background well developped by many authors over the
last 25 years It involves of the design of the Kalman-Bucy Fi l ter ( K B F ) which
provides an estimate of the states x and a Linear Quadratic Regulator ( L Q R ) which
is separately designed assuming all the states x are known Both the L Q R and K B F
designs independently have good robustness properties ie retain good performance
when disturbances due to model uncertainty are included However the combined
L Q R and K B F designs ie the L Q G design may have poor stability margins in the
presence of model uncertainties This l imitat ion has led to the development of an L Q G
design procedure that improves the performance of the compensator by recovering
the full state feedback robustness properties at the plant input or output (Figure 5-6)
This procedure is known as the Linear Quadratic Guassian-Loop Transfer Recovery
( L Q G L T R ) control algorithm A detailed development of its theory is given in
Ref[61] and hence is not repeated here for conciseness
The design of the L Q G L T R controller involves the repeated solution of comshy
plicated matrix equations too tedious to be executed by hand Fortunenately the enshy
tire algorithm is available in the Robust Control Toolbox[62] of the popular software
package M A T L A B The input matrices required for the function are the following
(i) state space matrices ABC and D (D = 0 for this system)
(ii) state and measurement noise covariance matrices E and 0 respectively
(iii) state and input weighting matrix Q and R respectively
114
u
MODEL UNCERTAINTY
SYSTEM PLANT
y
u LQR CONTROLLER
X KBF ESTIMATOR
u
y
LQG COMPENSATOR
PLANT
X = f v ( X t )
COMPENSATOR
x = A k x - B k y
ud
= C k f
y
Figure 5-6 Block diagram for the LQGLTR estimator based compensator
115
A s mentionned earlier the main objective of this offset controller is to regu-mdash
late the tether vibration described by the 5 equations However from the response
of the uncontrolled system presented in Chapter 4 it is clear that there is a large
difference between the magnitude of the librational and vibrational frequencies This
separation of frequencies allow for the separate design of the vibration and attitude mdash mdash
controllers Thus only the flexible subsystem composed of the S and d equations is
required in the offset controller design Similarly there is also a wide separation of
frequencies between the longitudinal and transverse modes of vibrations permitt ing mdash
the decoupling of the flexible subsystem into a set of longitudinal (Sx and dx) and
transverse (5y Sz dy and dz) subsystems The appended offset system d must also
be included since it acts as the control actuator However it is important to note
that the inplane and out-of-plane transverse modes can not be decoupled because
their oscillation frequencies are of the same order The two flexible subsystems are
summarized below
(i) Longitudinal Vibra t ion Subsystem
This is defined by u mdash Auxu + Buudu
(528)
where xu = SxdX25xdX2T u d u = dX2 and from Eq(527)
0 0 $u(lt) 0 0 0 0 1
0 0 1 0 0 0 0 1
(529)
Au and Bu are the rows and columns of A and B respectively and correspond to the
components of xu
The L Q R state weighting matrix Qu is taken as
Qu =
bull1 0 0 o -0 1 0 0 0 0 1 0
0 0 0 10
(530)
116
and the input weighting matrix Ru = 1^ The state noise covariance matr ix is
selected as 1 0 0 0 0 1 0 0 0 0 4 0
LO 0 0 1
while the measurement noise covariance matrix is taken as
(531)
i o 0 15
(532)
Given the above mentionned matrices the design of the L Q G L T R compenshy
sator is computed using M A T L A B with the following 2 commands
(i) kfu = l q r c ( ^ ( i C bdquo d i a g m x ( E u e u ) )
(ii) [AkugtBkugtCkugtDku =^ryAuBuCuDukfuQuRur)
where r is a scaler
The first command is used to compute the Ka lman filter gain matrix kfu Once
kfu is known the second command is invoked returning the state space representation
of the L Q G L T R dynamic compensator to
(533) xu mdash Akuxu Bkuyu
where xu is the state estimate vector oixu The function l t ry performs loop transfer
recovery at the system output ie the return ratio at the output approaches that
of the K B F loop given by Cu(sld - Au)Kfu[61 This is achieved by choosing a
sufficently large scaler value r in l t ry such that the singular values of the return
ratio approach those of the target design For the longitudinal controller design
r mdash 5 x 10 5 Figure 5-7(a) compares the singular values of the recovered compensator
design and the non-recovered L Q G compensator design with respect to the target
Unfortunetaly perfect recovery is not possible especially at higher frequencies
117
100
co
gt CO
-50
-100 10
Longitudinal Compensator Singular Values i I I I N I I mdash i mdash i i i i n i i mdash i mdash i i i i inimdash imdashi i i i i i i i mdash i mdash i i M i I I
Target r = 5x10 5
r = 0 (LQG)
i i i i I I i ii i i i 1 1 1 1 i i i i i 1 1 1 1 ii i I I i i i i 11111 -3 10 -2
(a)
101 10deg Frequency (rads)
101 10
Transverse Compensator Singular Values
100
50
X3
gt 0
CO
-50
bull100 10
-imdashi i 111II i 111 ni 1mdashi I I I I I I I 1mdashi I I I N I I 1mdashi i M 111
Target r = 50 r = 0 (LQG)
m i i i i i i n i i i i I I I I I I I 1mdashi I I I I I I I 1mdashi i 111 n -3 10 -2
(b)
101 10deg Frequency (rads)
101 10
Figure 5-7 Singular values for the LQG and LQGLTR compensator compared to target return ratio (a) longitudinal design (b) transverse design
118
because the system is non-minimal ie it has transmission zeros with positive real
parts[63]
(ii) Transverse Vibra t ion Subsystem
Here Xy mdash AyXy + ByU(lv
Dv = Cvxv
with xv = 6y 8Z dV2 dZ2 Sy 6Z dV2 dZ2T] udv = dy2dZ2T and
(534)
Cy
0 0 0 0 (0) 0 0 0 0 0 0 0 0 0 0 0 0
0 0 0 0
0 0 0 0 0 0 0 0 0 0 0 0
0 0 0
h 0 0 0
0 0 0 lt ( 0 ) o 0
0 1 0 0 0 1
0 0 0
2 r
h 0 0
0 0 1 0 0 1
(535)
Av and Bv are the rows and columns of the A and B corresponding to the components
O f Xy
The state L Q R weighting matrix Qv is given as
Qv
10 0 0 0 0 0 0 0 0 10 0 0 0 0 0 0 0 0 10 4 0 0 0 0 0 0 0 0 10 4 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1
(536)
mdash
Ry =w The state noise covariance matrix is taken
1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 9 x 10 4 0
0 0 0 0 0 0 0 9 x 10 4
(537)
119
while the measurement noise covariance matrix is
1 0 0 0 0 1 0 0 0 0 5 x 10 4 0
LO 0 0 5 x 10 4
(538)
As in the development of the longitudinal controller the transverse offset conshy
troller can be designed using the same two commands (lqrc and l try) wi th the input
matrices corresponding to the transverse subsystem with r = 50 The result is the
transverse compensator system given by
(539)
where xv is the state estimates of xv The singular values of the target transfer
function is compared with those achieved by the L Q G and L Q G L T R in Figure 5-
7(b) It is apparent that the recovery is not as good as that for the longitudinal case
again due to the non-minimal system Moreover the low value of r indicates that
only a lit t le recovery in the transverse subsystem is possible This suggests that the
robustness of the L Q G design is almost the maximum that can be achieved under
these conditions
Denoting f y = f j f j r xf = x^x^T and ud = udu ^ T the longishy
tudinal compensator can be combined to give
xf =
$d =
Aku
0
0 Akv
Cku o
Cu 0
Bku
0 B Xf
Xf = Ckxf
0
kv
Hf = Akxf - Bkyf
(540)
Vu Vv 0 Cv
Xf = CfXf
Defining a permutation matrix Pf such that Xj = PfX where X is the state vector
of the original nonlinear system given in Eq(32) the compensator and full nonlinear
120
system equations can be combined as
X
(541)
Ckxf
A block diagram representation of Eq(541) is shown in Figure 5-6
524 Simulation results
The dynamical response for the stationkeeping STSS with a tether length of 20
km and ini t ia l offset position of 1 m in each direction is presented in Figure 5-8 when
both the F L T attitude controller and the L Q G L T R offset controller are activated As
expected the offset controller is successful in quickly damping the elastic vibrations
in the longitudinal and transverse direction (Figure 5-8b) However from Figure 5-
8(a) it is clear that the presence of offset control requires a larger control moment
to regulate the attitude of the platform This is due to the additional torque created
by the larger moment arm around 2 m and 125 m in the inplane and out-of-plane
directions respectively introduced by the offset control In addition the control
moment for the platform is modulated by the tethers transverse vibration through
offset coupling However this does not significantly affect the l ibrational motion of
the tether whose thruster force remains relatively unchanged from the uncontrolled
vibration case
Final ly it can be concluded that the tether elastic vibration suppression
though offset control in conjunction with thruster and momentum-wheel attitude
control presents a viable strategy for regulating the dynamics tethered satellite sysshy
tems
121
STSS configuration 3-Body a1(0)=a2(0)=at(0)=2deg P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m6y(0)=52(0)=1m dx(0)=dy(0)=dz(0)=1m Stationkeeping I = 20km LQGLTR Vibration Control
Sat 1 Pitch and Roll Angles Tether Pitch and Roll Angles mdash 1 1 1 1 1 1 I 1 1 lt 1 ~ ^ 1 ^ l _
Time (Orbits) Time (Orbits) Sat 1 Roll Control Moment Tether Roll Control Thrust
(a) Time (Orbits) Time (Orbits)
Figure 5-8 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT attitude controller and LQGLTR offset vibration controller (a) attitude and libration controller reshysponse
122
STSS configuration 3-Body a1(0)=a2(0)=a(0)=2deg p1(0)=(32(0)=Q(0)=1deg ox(0)=14mSy(0)=5z(0)=1m dx(0)=dy(0)=dz(0)=1m Stationkeeping lt = 20km LQGLTR Vibration Control
Tether X Offset Position
Time (Orbits) Tether Y Tranverse Vibration 20r
Time (Orbits) Tether Y Offset Position
t o
1 2 Time (Orbits)
Tether Z Tranverse Vibration Time (Orbits)
Tether Z Offset Position
(b)
Figure 5-8
Time (Orbits) Time (Orbits)
Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT attitude controller and LQGLTR offset vibration controller (b) vibration and offset response
123
6 C O N C L U D I N G R E M A R K S
61 Summary of Results
The thesis has developed a rather general dynamics formulation for a multi-
body tethered system undergoing three-dimensional motion The system is composed
of multiple rigid bodies connected in a chain configuration by long flexible tethers
The tethers which are free to undergo libration as well as elastic vibrations in three
dimensions are also capable of deployment retrieval and constant length stationkeepshy
ing modes of operation Two types of actuators are located on the rigid satellites
thrusters and momentum-wheels
The governing equations of motion are developed using a new Order(N) apshy
proach that factorizes the mass matrix of the system such that it can be inverted
efficiently The derivation of the differential equations is generalized to account for
an arbitrary number of rigid bodies The equations were then coded in FORTRAN
for their numerical integration with the aid of a symbolic manipulation package that
algebraically evaluated the integrals involving the modes shapes functions used to dis-
cretize the flexible tether motion The simulation program was then used to assess the
uncontrolled dynamical behaviour of the system under the influence of several system
parameters including offset at the tether attachment point stationkeeping deployshy
ment and retrieval of the tether The study covered the three-body and five-body
geometries recently flown OEDIPUS system and the proposed BICEPS configurashy
tion It represents innovation at every phase a general three dimensional formulation
for multibody tethered systems an order-N algorithm for efficient computation and
application to systems of contemporary interest
Two types of controllers one for the attitude motion and the other for the
124
flexible vibratory motion are developed using the thrusters and momentum-wheels for
the former and the variable offset position for the latter These controllers are used to
regulate the motion of the system under various disturbances The attitude controller
is developed using the nonlinear Feedback Linearization Technique ( F L T ) and is based
on the nonlinear but rigid model of the system O n the other hand the separation
of the longitudinal and transverse frequencies from those of the attitude response
allows for the development of a linear optimal offset controller using the robust Linear
Quadratic Gaussian-Loop Transfer Recovery ( L Q G L T R ) method The effectiveness
of the two controllers is assessed through their subsequent application to the original
nonlinear flexible model
More important original contributions of the thesis which have not been reshy
ported in the literature include the following
(i) the model accounts for the motion of a multibody chain-type system undershy
going librational and in the case of tethers elastic vibrational motion in all
the three directions
(ii) the dynamics formulation is based on a recursive O(N) Lagrangian algorithm
that efficiently computes the systems generalized acceleration vector
(iii) the development of an attitude controller based on the Feedback Linearizashy
tion Technique for a multibody system using thrusters and momentum-wheels
located on each rigid body
(iv) the design of a three degree of freedom offset controller to regulate elastic v i shy
brations of the tether using the Linear Quadratic Gaussian and Loop Transfer
Recovery method ( L Q G L T R )
(v) substantiation of the formulation and control strategies through the applicashy
tion to a wide variety of systems thus demonstrating its versatility
125
The emphasis throughout has been on the development of a methodology to
study a large class of tethered systems efficiently It was not intended to compile
an extensive amount of data concerning the dynamical behaviour through a planned
variation of system parameters O f course a designer can easily employ the user-
friendly program to acquire such information Rather the objective was to establish
trends based on the parameters which are likely to have more significant effect on
the system dynamics both uncontrolled as well as controlled Based on the results
obtained the following general remarks can be made
(i) The presence of the platform products of inertia modulates the attitude reshy
sponse of the platform and gives rise to non-zero equilibrium pitch and roll
angles
(ii) Offset of the tether attachment point significantly affects the equil ibrium orishy
entation of the system as well as its dynamics through coupling W i t h a relshy
atively small subsatellite the tether dynamics dominate the system response
with high frequency elastic vibrations modulating the librational motion On
the other hand the effect of the platform dynamics on the tether response is
negligible As can be expected the platform dynamics is significantly affected
by the offset along the local horizontal and the orbit normal
(iii) For a three-body system deployment of the tether can destabilize the platform
in pitch as in the case of nonzero offset However the roll motion remains
undisturbed by the Coriolis force Moreover deployment can also render the
tether slack if it proceeds beyond a critical speed
(iv) Uncontrolled retrieval of the tether is always unstable in pitch It also leads
to high frequency elastic vibrations in the tether
(v) The five-body tethered system exhibits dynamical characteristics similar to
those observed for the three-body case As can be expected now there are
additional coupling effects due to two extra bodies a rigid subsatellite and a
126
flexible tether
(vi) Cartwheeling motion of the B I C E P S configuration can also be used to adshy
vantage in the deployment of the tether However exceeding a crit ical ini t ia l
cartwheeling rate can result in a large tension in the tether which eventually
causes the satellite to bounce back rendering the tether slack
(vii) Spinning the tether about its nominal length as in the case of O E D I P U S
introduces high frequency transverse vibrations in the tether which in turn
affect the dynamics of the satellites
(viii) The F L T based controller is quite successful in regulating the attitude moshy
tion of both the tether and rigid satellites in a short period of time during
stationkeeping deployment as well as retrieval phases
(ix) The offset controller is successful in suppressing the elastic vibrations of the
tether wi thin the allowable l imits of the offset mechanism ( plusmn 2 0 m)
62 Recommendations for Future Study
A s can be expected any scientific inquiry is merely a prelude to further efforts
and insight Several possible avenues exist for further investigation in this field A
few related to the study presented here are indicated below
(i) inclusion of environmental effects particularly the free molecular reaction
forces as well as interaction with Earths magnetic field and solar radiation
(ii) addition of flexible booms attached to the rigid satellites providing a mechashy
nism for energy dissipation
(iii) validation of the new three dimensional offset control strategies using ground
based experiments
(iv) animation of the simulation results to allow visual interpretation of the system
dynamics
127
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[34] Pradhan S M o d i V J and Misra A K On the Inverse Control of the Tethered Satellite Systems The Journal of the Astronautical Sciences A A S Publications Office San Diego California U S A V o l 43 No 2 Apr i l - June 1995 pp 179-193
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[37] Pradhan S M o d i V J and Misra A K Simultaneous Control of Platform Att i tude and Tether Vibra t ion using Offset Strategy AIAAA AS Astrodynamics Specialist Conference San Diego C A U S A July 1996 Paper No 96-3570 pp 1-11
[38] Ruying F and Bainum P M Dynamics and Control of a Space Platform with a Tethered Subsatellite Journal of Guidance Control and Dynamics V o l 11 No 4 July-August 1988 pp 377-381
[39] M o d i V J Lakshmanan P K and Misra A K Offset Control of Tethered Satellite Systems Analysis and Experimental Verification Acta Astronautica V o l 21 No 5 1990 pp 283-294
[40] X u D M Misra A K and M o d i V J Thruster Augmented Act ive Cin t ro l of a Tethered Satellite System During its Retrieval Journal of Guidance Control and Dynamics V o l 9 No 6 November-December 1986 pp 663-672
[41] M o d i V J Pradhan S and Misra A K Off-Set Control of the Tethered Systems Using a Graph Theoretic Approach Acta Astronautica V o l 35 No
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[43] Pradhan S Planar Dynamics and Control of Tethered Satellite Systems P h D Thesis Department of Mechanical Engineering The University of Br i t i sh Columbia December 1994
[44] Bainum P M and Kumar V K Optimal-Control of the Shuttle-Tethered System Acta Astronautica V o l 7 No 12 1980 pp 1333-1348
[46] Hughes P C Spacecraft Att i tude Dynamics John Wi ley amp Sons New York U S A 1992 Chapter 2 pp 7-38
[47] Stengel R F Optimal Control and Estimation Dover Publications Inc New York U S A 1994 p 26
[48] Nayfeh A H and Mook D T Nonlinear Oscillation John Wi ley and Sons New York U S A 1979 pp 485-500
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[50] Meirovitch L Elements of Vibration Analysis M c G r a w - H i l l Inc New York U S A 1986 pp 255-256
[51] Greenwood D T Principles of Dynamics Prentice Ha l l Inc Englewood Cliffs New Jersey U S A 1965 pp 252-275
[52] Boyce W E and D i P r i m a R C Elemental Differential Equations and Boundshy
ary Value Problems John Wi ley amp Sons Inc New York U S A 1986 pp 424-431
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[54] Gear C W Numerical Initial Value Problems in Ordinary Differential Equashy
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[56] Tsakalis K S and Ioannou P A Linear Time-Varying Systems Control and
Adaptation Prentice-Hall Inc Englewood Cliffs New Jersey U S A 1993 pp 148-180
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[60] Athans M The Role and Use of the Stochastic Linear-Quadrat ic-Guassian Problem in Control System Design IEEE Transactions on Automatic Control V o l A C - 1 6 No 6 December 1971 pp 529-552
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[62] Chiang R Y and Safonov M G Robust Control Toolbox Users Guide The M a t h Works Inc Natick Mass U S A 1992 pp 272-274
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132
A P P E N D I X I TENSOR R E P R E S E N T A T I O N OF T H E
EQUATIONS OF M O T I O N
11 Prel iminary Remarks
The derivation of the equations of motion for the multibody tethered system
involves the product of several matrices as well as the derivative of matrices and
vectors with respect to other vectors In order to concisely code these relationships
in FORTRAN the matrix equations of motion are expressed in tensor notation
Tensor mathematics in general is an extremely powerful tool which can be
used to solve many complex problems in physics and engineering[6465] However
only a few preliminary results of Cartesian tensor analysis are required here They
are summarized below
12 Mathematical Background
The representation of matrices and vectors in tensor notation simply involves
the use of indices referring to the specific elements of the matrix entity For example
v = vk (k = lN) (Il)
A = Aij (i = lNj = lM)
where vk is the kth element of vector v and Aj is the element on the t h row and
3th column of matrix A It is clear from Eq(Il) that exactly one index is required to
completely define an entire vector whereas two independent indices are required for
matrices For this reason vectors and matrices are known as first-order and second-
order tensors respectively Scaler variables are known as zeroth-order tensors since
in this case an index is not required
133
Matrix operations are expressed in tensor notation similar to the way they are
programmed using a computer language such as FORTRAN or C They are expressed
in a summation notation known as Einstein notation For example the product of a
matrix with a vector is given as
w = Av (12)
or in tensor notation W i = ^2^2Aij(vkSjk)
4-1 ltL3gt j
where 8jk is the Kronecker delta Dropping the summation symbol the matrix-vector
product can be expressed compactly as
Wi = AijVj = Aimvm (14)
Here j is a dummy index since it appears twice in a single term and hence can
be replaced by any other index Note that since the resulting product is a vector
only one index is required namely i that appears exactly once on both sides of the
equation Similarly
E = aAB + bCD (15)
or Ej mdash aAimBmj + bCimDmj
~ aBmjAim I bDmjCim (16)
where A B C D and E are second-order tensors (matrices) and a and b are zeroth-
order tensors (scalers) In addition once an expression is in tensor form it can be
treated similar to a scaler and the terms can be rearranged
The transpose of a matrix is also easily described using tensors One simply
switches the position of the indices For example let w = ATv then in tensor-
notation
Wi = AjiVj (17)
134
The real power of tensor notation is in its unambiguous expression of terms
containing partial derivatives of scalers vectors or matrices with respect to other
vectors For example the Jacobian matrix of y = f(x) can be easily expressed in
tensor notation as
di-dx-j-1 ( L 8 )
mdash which is a second-order tensor as expected If f(x) = Ax where A is constant then
f - a )
according to the current definition If C = A(x)B(x) then the time derivative of C
is given by
C = Ax)Bx) + A(x)Bx) (110)
or in tensor notation
Cij = AimBmj + AimBmj (111)
= XhiAimfcBmj + AimBmj^k)
which is more clear than its equivalent vector form Note that A^^ and Bmjk are
third-order tensors that can be readily handled in F O R T R A N
13 Forcing Function
The forcing function of the equations of motion as given by Eq(260) is
- u 1-SdM- dqtdPe 9 f i m F(q qt) = Mq - -q mdashq + mdashmdash - Qd (260)
This expression can be converted into tensor form to facilitate its implementation
into F O R T R A N source code
Consider the first term Mq From Eq(289) the mass matrix M is given by
T M mdash Rv MtRv which in tensor form is
Mtj = RZiM^R^j (112)
135
Taking the time derivative of M gives
Mij = qsRv
ni^MtnrnRv
mj + Rv
niMtnTn^Rv
mj + Rv
niMtnmRv
mjs) (113)
The second term on the right hand side in Eq(260) is given by
-JPdMu 1 bdquo T x
2Q ~cWq = 2Qs s r k Q r ^ ^
Expanding Msrk leads to
M src = RnskMtnmRmr + RnsMtnmkRmr + RnsMtnmRmrk- (L15)
Finally the potential energy term in Eq(260) is also expanded into tensor
form to give dPe = dqt dPe
dq dq dqt = dq^QP^
and since ^ = Rp(q)q then
Now inserting Eqs(I13-I17) into Eq(260) and rearranging the tensor form of the
forcing term can how be stated as
Fk(q^q^)=(RP
sk + R P n J q n ) ^ - - Q d k
+ QsQr | (^Rlks ~ 2 M t n m R m r
+ (^RnkMtnms ~ 2~RnsMtnmgtk^ R m r
(^RnkRmrs ~ ~2RnsRmrk^j
where Fk(q q t) represents the kth component of F
(118)
136
A P P E N D I X II R E D U C E D EQUATIONS OF M O T I O N
II 1 Prel iminary Remarks
The reduced model used for the design of the vibration controller is represented
by two rigid end-bodies capable of three-dimensional attitude motion interconnected
with a fixed length tether The flexible tether is discretized using the assumed-
mode method with only the fundamental mode of vibration considered to represent
longitudinal as well as in-plane and out-of-plane transverse deformations In this
model tether structural damping is neglected Finally the system is restricted to a
nominal circular orbit
II 2 Derivation of the Lagrangian Equations of Mot ion
Derivation of the equations of motion for the reduced system follows a similar
procedure to that for the full case presented in Chapter 2 However in this model the
straightforward derivation of the energy expressions for the entire system is undershy
taken ignoring the Order(N) approach discussed previously Furthermore in order
to make the final governing equations stationary the attitude motion is now referred
to the local LVLH frame This is accomplished by simply defining the following new
attitude angles
ci = on + 6
Pi = Pi (HI)
7t = 0
where and are the pitch and roll angles respectively Note that spin is not
considered in the reduced model hence 7J = 0 Substituting Eq(IIl) into Eq(214)
137
gives the new expression for the rotation matrix as
T- CpiSa^ Cai+61 S 0 S a + e i
-S 0 c Pi
(II2)
W i t h the new rotation matrix defined the inertial velocity of the elemental
mass drrii o n the ^ link can now be expressed using Eq(215) as
(215)
Since the tether length is assumed to be fixed T j f j and Tj$jlt^ of Eq(215) are
identically zero Also defining
(II3) Vi = I Pi i
Rdm- f deg r the reduced system is given by
where
and
Rdm =Di + Pi(gi)m + T^iSi
Pi(9i)m= [Tai9i T0gi T^] fj
(II4)
(II5)
(II6)
D1 = D D l D S v
52 = 31 + P(d2)m +T[d2
D3 = D2 + l2 + dX2Pii)m + T2i6x
Note that $ j for i = 2 is defined in Section 212 whereas for i = 1 and 3 it is
the null matrix Since only one mode is used for each flexible degree of freedom
5i = [6Xi6yi6Zi]T for i = 2
The kinetic energy of the system can now be written as
i=l Z i = l Jmi RdmRdmdmi (UJ)
138
Expanding Kampi using Eq(II4)
KH = miDD + 2DiP
i j gidmi)fji + 2D T ltMtrade^
+tf J P^(9i)Pi9i)d^l + 2mT j PFmTi^dm^ (II8)
where the integrals are evaluated using the procedure similar to that described in
Chapter 2 The gravitational and strain energy expressions are given by Eq(247)
and Eq(251) respectively using the newly defined rotation matrix T in place of
TV
Substituting the kinetic and potential energy expressions in
d ( 8Ke d(Ke - Pe) i bull i Q ^ 0 (II9)
d t dqred d(ired
with
Qred = [ algt l gt a 2 2 A 2 A 2 lt ^ 2 Q 3 3 ] (5-14)
and d2 regarded as a time-varying parameter the nonlinear non-autonomous equashy
tions of motion for the reduced system can finally be expressed as
MredQred + fred = reg- (5-15)
139
LIST OF SYMBOLS
A j intermediate length over which the deploymentretrieval profile
for the itfl l ink is sinusoidal
A Lagrange multipliers
EUQU longitudinal state and measurement noise covariance matrices
respectively
matrix containing mode shape functions of the t h flexible link
pitch roll and yaw angles of the i 1 link
time-varying modal coordinate for the i 1 flexible link
link strain and stress respectively
damping factor of the t h attitude actuator
set of attitude angles (fjj = ci
structural damping coefficient for the t h l ink EjJEj
true anomaly
Earths gravitational constant GMejBe
density of the ith link
fundamental frequency of the link longitudinal tether vibrashy
tion
A tethers cross-sectional area
AfBfCjDf state-space representation of flexible subsystem
Dj inertial position vector of frame Fj
Dj magnitude of Dj mdash mdash
Df) transformation matrix relating Dj and Ds
D r i inplane radial distance of the first l ink
DSi Drv6i DZ1 T for i = 1 and DX Dy poundgt 2 T for 1
m
Pi
w 0 i
v i i
Dx- transformation matrix relating D j and Ds-
DXi Dyi Dz- Cartesian components of D
DZl out-of-plane position component of first link
E Youngs elastic modulus for the ith l ink
Ej^ contribution from structural damping to the augmented complex
modulus E
F systems conservative force vector
FQ inertial reference frame
Fi t h link body-fixed reference frame
n x n identity matrix
alternate form of inertia matrix for the i 1 l ink
K A Z $i(k + dXi+1)
Kei t h link kinetic energy
M(q t) systems coupled mass matrix
Mma Mma Mmy- control moments in the pitch roll and yaw directions respec-
tively for the t h link
Mr fr rigid mass matrix and force vector respectively
Mred fred- Qred mass matrix force vector and generalized coordinate vector for
the reduced model respectively
Mt block diagonal decoupled mass matrix
symmetric decoupled mass matrix
N total number of links
0N) order-N
Pi(9i) column matrix [T^giTpg^T^gi) mdash
Q nonconservative generalized force vector
Qu actuator coupling matrix or longitudinal L Q R state weighting
matrix
v i i i
Rdm- inertial position of the t h link mass element drrii
RP transformation matrix relating qt and tf
Ru longitudinal L Q R input weighting matrix
RvRn Rd transformation matrices relating qt and q
S(f Mdc) generalized acceleration vector of coupled system q for the full
nonlinear flexible system
th uiirv l u w u u u maniA
th
T j i l ink rotation matrix
TtaTto control thrust in the pitch and roll direction for the im l ink
respectively
rQi maximum deploymentretrieval velocity of the t h l ink
Vei l ink strain energy
Vgi link gravitational potential energy
Rayleigh dissipation function arising from structural damping
in the ith link
di position vector to the frame F from the frame F_i bull
dc desired offset acceleration vector (i 1 l ink offset position)
drrii infinitesimal mass element of the t h l ink
dx^dy^dZi Cartesian components of d along the local vertical local horishy
zontal and orbit normal directions respectively
F-Q mdash
gi r igid and flexible position vectors of drrii fj + ltEraquolt5
ijk unit vectors 1 0 0 r 010-^ and 00 l r respectively
li length of the t h l ink
raj mass of the t h link
nfj total number of flexible modes for the i 1 l ink nXi + nyi + nZi
nq total number of generalized coordinates per link nfj + 7
nqq systems total number of generalized coordinates Nnq
ix
number of flexible modes in the longitudinal inplane and out-
of-plane transverse directions respectively for the i ^ link
number of attitude control actuators
- $ T
set of generalized coordinates for the itfl l ink which accounts for
interactions with adjacent links
qv---QtNT
set of coordinates for the independent t h link (not connected to
adjacent links)
rigid position of dm in the frame Fj
position of centre of mass of the itfl l ink relative to the frame Fj
i + $DJi
desired settling time of the j 1 attitude actuator
actuator force vector for entire system
flexible deformation of the link along the Xj yi and Zj direcshy
tions respectively
control input for flexible subsystem
Cartesian components of fj
actual and estimated state of flexible subsystem respectively
output vector of flexible subsystem
LIST OF FIGURES
1-1 A schematic diagram of the space platform based N-body tethered
satellite system 2
1-2 Associated forces for the dumbbell satellite configuration 3
1- 3 Some applications of the tethered satellite systems (a) multiple
communication satellites at a fixed geostationary location
(b) retrieval maneuver (c) proposed B I C E P S configuration 6
2- 1 Vector components of the i 1 and (i- l ) 1 chain links 15
2-2 Vector components in cylindrical orbital coordinates 20
2-3 Inertial position of subsatellite thruster forces 38
2- 4 Coupled force representation of a momentum-wheel on a rigid body 40
3- 1 Flowchart showing the computer program structure 45
3-2 Kinet ic and potentialenergy transfer for the three-body platform based
tethered satellite system (a) variation of kinetic and potential energy
(b) percent change in total energy of system 48
3-3 Kinet ic and potential energy transfer for the five-body platform based
tethered satellite system (a) variation of kinetic and potential energy
(b) percent change in total energy of system 49
3-4 Simulation results for the platform based three-body tethered
system originally presented in Ref[43] 51
3- 5 Simulation results for the platform based three-body tethered
system obtained using the present computer program 52
4- 1 Schematic diagram showing the generalized coordinates used to
describe the system dynamics 55
4-2 Stationkeeping dynamics of the three-body STSS configuration
without offset (a) attitude response (b) vibration response 57
4-3 Stationkeeping dynamics of the three-body STSS configuration
xi
with offset along the local vertical
(a) attitude response (b) vibration response 60
4-4 Stationkeeping dynamics of the three-body STSS configuration
with offset along the local horizontal
(a) attitude response (b) vibration response 62
4-5 Stationkeeping dynamics of the three-body STSS configuration
with offset along the orbit normal
(a) attitude response (b) vibration response 64
4-6 Stationkeeping dynamics of the three-body STSS configuration
wi th offset along the local horizontal local vertical and orbit normal
(a) attitude response (b) vibration response 67
4-7 Deployment dynamics of the three-body STSS configuration
without offset
(a) attitude response (b) vibration response 69
4-8 Deployment dynamics of the three-body STSS configuration
with offset along the local vertical
(a) attitude response (b) vibration response 72
4-9 Retrieval dynamics of the three-body STSS configuration
with offset along the local vertical and orbit normal
(a) attitude response (b) vibration response 74
4-10 Schematic diagram of the five-body system used in the numerical example 77
4-11 Stationkeeping dynamics of the five-body STSS configuration
without offset
(a) attitude response (b) vibration response 78
4-12 Deployment dynamics of the five-body STSS configuration
wi th offset along the local vertical
(a) attitude response (b) vibration response 81
4-13 Stationkeeping dynamics of the three-body B I C E P S configuration
x i i
with offset along the local vertical
(a) attitude response (b) vibration response 84
4-14 Cartwheeling dynamics with deployment for the three-body B I C E P S
with offset along the local vertical
(a) attitude response (b) vibration response 86
4-15 Spin dynamics (7 = l deg s ) of the three-body O E D I P U S configuration
with offset along the local vertical
(a) attitude response (b) vibration response 89
4- 16 Spin dynamics (7 = 10deg s ) of the three-body O E D I P U S configuration
wi th offset along the local vertical
(a) attitude response (b) vibration response 91
5- 1 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear rigid F L T controller with offset along the
local horizontal local vertical and orbit normal
(a) attitude and vibration response (b) control actuator time histories 98
5-2 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear flexible F L T controller with offset along the
local horizontal local vertical and orbit normal
(a) attitude and vibration response (b) control actuator time histories 101
5-3 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear rigid F L T controller with offset along
the local vertical
(a) attitude and vibration response (b) control actuator time histories 103
5-4 Deployment dynamics of the three-body STSS using the nonlinear
rigid F L T controller with offset along the local vertical
(a) attitude and vibration response (b) control actuator time histories 105
5-5 Retrieval dynamics of the three-body STSS using the non-linear
rigid F L T controller with offset along the local vertical and orbit normal
(a) attitude and vibration response (b) control actuator time histories 108
x i i i
5-6 Block diagram for the L Q G L T R estimator based compensator 115
5-7 Singular values for the L Q G and L Q G L T R compensator compared to target
return ratio (a) longitudinal design (b) transverse design 118
5-8 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear rigid F L T attitude controller and L Q G L T R
offset vibration controller (a) attitude and libration controller response
(b) vibration and offset response 122
xiv
A C K N O W L E D G E M E N T
Firs t and foremost I would like to express the most genuine gratitude to
my supervisor Prof V i n o d J M o d i whose unending guidance and encouragement
proved most invaluable during the course of my studies
I am also deeply indebted to Dr Satyabrata Pradhan for his limitless patience
and technical advice from which this work would not have been possible I would also
like to extend a word of appreciation towards Prof A r u n K Mis ra ( M c G i l l University)
for ini t iat ing my interest in space dynamics and control
Also I would like to thank my collegues Dr Sandeep Munshi M r Mathieu
Caron and M r Yuan Chen as well as Dr Mae Seto M r Gary L i m and M r Mark
Chu for their words of advice and helpful suggestions that heavily contributed to all
aspects of my studies
Final ly I would like to thank all my family and friends here in Vancouver and
back home in Montreal whose unending support made the past two years the most
memorable of my life
The research project was supported by the Natural Sciences and Engineering
Research Council s P G S - A Scholarship held by the author as well as by N S E R C grant
A-2181 held by Prof V J M o d i
xv
1 INTRODUCTION
11 Prel iminary Remarks
W i t h the ever changing demands of the worlds population one often wonders
about the commitment to the space program On the other hand the importance
of worldwide communications global environmental monitoring as well as long-term
habitation in space have demanded more ambitious and versatile satellites systems
It is doubtful that the Russian scientist Tsiolkovsky[l] when he first proposed the
uti l izat ion of the Earths gravity-gradient environment in the last century would have
envisioned the current and proposed applications of tethered satellite systems
A tethered satellite system consists of two or more subsatellites connected
to each other by long thin cables or tethers which are in tension (Figure 1-1) The
subsatellite which can also include the shuttle and space station are in orbit together
The first proposed use of tethers in space was associated with the rescue of stranded
astronauts by throwing a buoy on a tether and reeling it to the rescue vehicle
Prel iminary studies of such systems led to the discovery of the inherent instability
during the tether retrieval[2] Nevertheless the bir th of tethered systems occured
during the Gemini X I and X I I missions[3] in 1966 when a short tether (10-30m)
was used to generate the first artificial gravity environment in space (000015g) by
cartwheeling (spinning) the system about its center of mass
The mission also demonstrated an important use of tethers for gravity gradient
stabilization The force analysis of a simplified model illustrates this point Consider
the dumbbell satellite configuration orbiting about the centre of Ear th as shown in
Figure 1-2 Satellite 1 and satellite 2 are connected to each other by a tether with
1
Space Station (Satellite 1 )
2
Satellite N
Figure 1-1 A schematic diagram of the space platform based N-body tethered satellite system
2
Figure 1-2 Associated forces for the dumbbell satellite configuration
3
the whole system free to librate about the systems centre of mass C M The two
major forces acting on each body are due to the centrifugal and gravitational effects
Since body 1 is further away from the earth its gravitational force Fg^y is smaller
than that of body 2 ie Fgi lt Fg^ However its centrifugal force FCl is greater
then the centrifugal force experienced by the second satellite (FCl gt FC2) Moreover
for body 1 Fg^ lt Fc^ in contrast to Fg2 gt FC2 for body 2 Thus resolving the force
components along the tether there is an evident resultant tension force Ff present
in the tether Similarly adding the normal components of the two forces vectorially
results in a force FR which restores the system to its stable equil ibrium along the
local vertical These tension and gravity-gradient restoring forces are the two most
important features of tethered systems
Several milestone missions have flown in the recent past The U S A J a p a n
project T P E (Tethered Payload Experiment) [4] was one of the first to be launched
using a sounding rocket to conduct environmental studies The results of the T P E
provided support for the N A S A I t a l y shuttle based tethered satellite system referred
to as the TSS-1 mission[5] This experiment aimed to study the electrodynamic effects
of a shuttle borne satellite system with a conductive tether connection where an
electric current was induced in the tether as it swept through the Earth s magnetic
field Unfortunetely the two attempts in August 1992 and February 1996 resulted
in only partial success The former suffered from a spool failure resulting in a tether
deployment of only 256m of a planned 20km During the latter attempt a break
in the tethers insulation resulted in arcing leading to tether rupture Nevertheless
the information gathered by these attempts st i l l provided the engineers and scientists
invaluable information on the dynamic behaviour of this complex system
Canada also paved the way with novel designs of tethered system aimed
primari ly at the study of the ionosphere and Northern Lights (Aurora Borealis)
4
The O E D I P U S A and C (Observation of Electrified Distributions in the Ionospheric
P lasma - a Unique Strategy) missions[6] launched from a sounding rocket in January
1989 and November 1995 respectively provided insight into the complex dynamical
behaviour of two comparable mass satellites connected by a 1 km long spinning
tether
Final ly two experiments were conducted by N A S A called S E D S I and II
(Small Expendable Deployable System) [7] which hold the current record of 20 km
long tether In each of these missions a small probe (26 kg) was deployed from the
second stage of a D E L T A II rocket thus succesfully demonstrating the feasibility of
long tether deployment Retrieval of the tether which is significantly more difficult
has not yet been achieved
Several proposed applications are also currently under study These include the
study of earths upper atmosphere using probes lowered from the shuttle in a low earth
orbit ( L E O ) to the deployment and retrieval of satellites for servicing Even more
promising application concerns the generation of power for the proposed International
Space Station using conductive tethers The Canadian Space Agency (CSA) has also
proposed a dual-mass satellite system B I C E P S (Bl-static Canadian Experiment on
Plasmas in Space) to make simultaneous measurements at different locations in the
environment for correlation studies[8] A unique feature of the B I C E P S mission is
the deployment of the tether aided by the angular momentum of the system which
is ini t ia l ly cartwheeling about its orbit normal (Figure 1-3)
In addition to the three-body configuration multi-tethered systems have been
proposed for monitoring Earths environment in the global project monitored by
N A S A entitled Mission to Planet Ear th ( M T P E ) The proposed mission aims to
study pollution control through the understanding of the dynamic interactions be-
5
Retrieval of Satellite for Servicing
lt mdash
lt Satellite
Figure 1-3 Some applications of the tethered satellite systems (a) multiple comshymunication satellites at a fixed geostationary location (b) retrieval maneuver (c) proposed BICEPS configuration
6
tween the Earths atmosphere biosphere hydrosphere and cryosphere This wi l l be
accomplished with multiple tethered instrumentation payloads simultaneously soundshy
ing at different altitudes for spatial correlations Such mult ibody systems are considshy
ered to be the next stage in the evolution of tethered satellites Micro-gravity payload
modules suspended from the proposed space station for long-term experiments as well
as communications satellites with increased line-of-sight capability represent just two
of the numerous possible applications under consideration
12 Brief Review of the Relevant Literature
The design of multi-payload systems wi l l require extensive dynamic analysis
and parametric study before the final mission configuration can be chosen This
wi l l include a review of the fundamental dynamics of the simple two-body tethered
systems and how its dynamics can be extended to those of multi-body tethered system
in a chain or more generally a tree topology
121 Multibody 0(N) formulation
Many studies of two-body systems have been reported One of the earliest
contribution is by Rupp[9] who was the first to study planar dynamics of the simshy
plified Shuttle based Tethered Satellite System (STSS) His pioneering work led to
the discovery of pitch oscillation growth during the retrieval phase Later Baker[10]
advanced the investigation to the third dimension in addition to adding atmospheric
effects to the system A more complete dynamical model was later developed and
analyzed by M o d i and Misra[ l l 12] which included deployment and tether flexibility
A comprehensive survey of important developments and milestones in the area have
been compiled and reviewed by Mis ra and Modi[13] The major conclusions based on
the literature may be summarized as follows
bull the stationkeeping phase is stable
7
bull deployment can be unstable if the rate exceeds a crit ical speed
bull retrieval of the tether is inherently unstable
bull transverse vibrations can grow due to the Coriolis force induced during deshy
ployment and retrieval
Misra Amier and Modi[14] were one of the first to extend these results to the
three-satellite double pendulum case This simple model which includes a variable
length tether was sufficient to uncover the multiple equilibrium configurations of
the system Later the work was extended to include out-of-plane motion and a
preliminary reel-rate control law to regulate the librational motion[15] Kesmir i and
Misra[16] developed a general formulation for N-body tethered systems based on the
Lagrangian principle where three-dimensional motion and flexibility are accounted for
However from their work it is clear that as the number of payload or bodies increases
the computational cost to solve the forward dynamics also increases dramatically
Tradit ional methods of inverting the mass matrix M in the equation
ML+F = Q
to solve for the acceleration vector proved to be computationally costly being of the
order for practical simulation algorithms O(N^) refers to a mult ibody formulashy
tion algorithm whose computational cost is proportional to the cube of the number
of bodies N used It is therefore clear that more efficient solution strategies have to
be considered
Many improved algorithms have been proposed to reduce the number of comshy
putational steps required for solving for the acceleration vector Rosenthal[17] proshy
posed a recursive formulation based on the triangularization of the equations of moshy
tion to obtain an 0(N2) formulation He also demonstrates the use of a symbolic
manipulation program to reduce the final number of computations A recursive L a -
8
grangian formulation proposed by Book[18] also points out the relative inefficiency of
conventional formulations and proceeds to derive an algorithm which is also 0(N2)
A recursive formulation is one where the equations of motion are calculated in order
from one end of the system to the other It usually involves one or more passes along
the links to calculate the acceleration of each link (forward pass) and the constraint
forces (backward or inverse pass) Non-recursive algorithms express the equations of
motion for each link independent of the other Although the coupling terms st i l l need
to be included the algorithm is in general more amenable to parallel computing
A st i l l more efficient dynamics formulation is an O(N) algorithm Here the
computational effort is directly proportional to the number of bodies or links used
Several types of such algorithms have been developed over the years The one based
on the Newton-Euler approach has recursive equations and is used extensively in the
field of multi-joint robotics[1920] It should be emphasized that efficient computation
of the forward and inverse dynamics is imperative if any on-line (real-time) control
of the joint is to be achieved Hollerbach[21] proposed a recursive O(N) formulation
based on Lagranges equations of motion However his derivation was primari ly foshy
cused on the inverse dynamics and did not improve the computational efficiency of
the forward dynamics (acceleration vector) Other recursive algorithms include methshy
ods based on the principle of vir tual work[22] and Kanes equations of motion[23]
Keat[24] proposed a method based on a velocity transformation that eliminated the
appearance of constraint forces and was recursive in nature O n the other hand
K u r d i l a Menon and Sunkel[25] proposed a non-recursive algorithm based on the
Range-Space method which employs element-by-element approach used in modern
finite-element procedures Authors also demonstrated the potential for parallel comshy
putation of their non-recursive formulation The introduction of a Spatial Operator
Factorization[262728] which utilizes an analogy between mult ibody robot dynamics
and linear filtering and smoothing theory to efficiently invert the systems mass ma-
9
t r ix is another approach to a recursive algorithm Their results were further extended
to include the case where joints follow user-specified profiles[29] More recently Prad-
han et al [30] proposed a non-recursive Lagrangian algorithm for factorizing the mass
matrix leading to an O(N) formulation of the forward dynamics of the system A n
efficient O(N) algorithm where the number of bodies N in the system varies on-line
has been developed by Banerjee[31]
122 Issues of tether modelling
The importance of including tether flexibility has been demonstrated by several
researchers in the past However there are several methods available for modelling
tether flexibility Each of these methods has relative advantages and limitations
wi th regards to the computational efficiency A continuum model where flexible
motion is discretized using the assumed mode method has been proposed[3233] and
succesfully implemented In the majority of cases even one flexible mode is sufficient
to capture the dynamical behaviour of the tether[34] K i m and Vadali[35] proposed a
so-called bead model or lumped-mass approach that discretizes the tether using point
masses along its length However in order to accurately portray the motion of the
tether a significantly high number of beads are needed thus increasing the number of
computation steps Consequently this has led to the development of a hybrid between
the last two approaches[16] ie interpolation between the beads using a continuum
approach thus requiring less number of beads Finally the tether flexibility can also
be modelled using a finite-element approach[36] which is more suitable for transient
response analysis In the present study the continuum approach is adopted due to
its simplicity and proven reliability in accurately conveying the vibratory response
In the past the modal integrals arising from the discretization process have
been evaluted numerically In general this leads to unnecessary effort by the comshy
puter Today these integrals can be evaluated analytically using a symbolic in-
10
tegration package eg Maple V Subsequently they can be coded in F O R T R A N
directly by the package which could also result in a significant reduction in debugging
time More importantly there is considerable improvement in the computational effishy
ciency [16] especially during deployment and retrieval where the modal integral must
be evaluated at each time-step
123 Attitude and vibration control
In view of the conclusions arrived at by some of the researchers mentioned
above it is clear that an appropriate control strategy is needed to regulate the dyshy
namics of the system ie attitude and tether vibration First the attitude motion of
the entire system must be controlled This of course would be essential in the case of
a satellite system intended for scientific experiments such as the micro-gravity facilishy
ties aboard the proposed Internation Space Station Vibra t ion of the flexible members
wi l l have to be checked if they affect the integrity of the on-board instrumentation
Thruster as well as momentum-wheel approach[34] have been considered to
regulate the rigid-body motion of the end-satellite as well as the swinging motion
of the tether The procedure is particularly attractive due to its effectiveness over a
wide range of tether lengths as well as ease of implementation Other methods inshy
clude tension and length rate control which regulate the tethers tension and nominal
unstretched length respectively[915] It is usually implemented at the deployment
spool of the tether More recently an offset strategy involving time dependent moshy
tion of the tether attachment point to the platform has been proposed[3738] It
overcomes the problem of plume impingement created by the thruster control and
the ineffectiveness of tension control at shorter lengths However the effectiveness of
such a controller can become limited with an exceedingly long tether due to a pracshy
tical l imi t on the permissible offset motion In response to these two control issues a
hybrid thrusteroffset scheme has been proposed to combine the best features of the
11
two methods[39]
In addition to attitude control these schemes can be used to attenuate the
flexible response of the tether Tension and length rate control[12] as well as thruster
based algorithms[3340] have been proposed to this end M o d i et al [41] have sucshy
cessfully demonstrated the effectiveness of an offset strategy to damp undesired v i shy
brations Passive energy dissipative devices eg viscous dampers are also another
viable solution to the problem
The development of various control laws to implement the above mentioned
strategies has also recieved much attention A n eigenstructure assignment in conducshy
tion wi th an offset controller for vibration attenuation and momemtum wheels for
platform libration control has been developed[42] in addition to the controller design
from a graph theoretic approach[41] Also non-linear feedback methods such as the
Feedback Linearization Technique ( F L T ) that are more suitable for controlling highly
nonlinear non-autonomous coupled systems have also been considered[43]
It is important to point out that several linear controllers including the classhy
sic state feedback Linear Quadratic Regulator ( L Q R ) have received considerable
attention[3944] Moreover robust methods such as the Linear Quadratic Guas-
s ian Loop Transfer Recovery ( L Q G L T R ) method have also been developed and imshy
plemented on tethered systems[43] A more complete review of these algorithms as
well as others applied to tethered system has been presented by Pradhan[43]
13 Scope of the Investigation
The objective of the thesis is to develop and implement a versatile as well as
computationally efficient formulation algorithm applicable to a large class of tethered
satellite systems The distinctive features of the model can be summarized as follows
12
bull TV-body 0(N) tethered satellite system in a chain-type topology
bull the system is free to negotiate a general Keplerian orbit and permitted to
undergo three dimensional inplane and out-of-plane l ibrtational motion
bull r igid bodies constituting the system are free to execute general rotational
motion
bull three dimensional flexibility present in the tether which is discretized using
the continuum assumed-mode method with an arbitrary number of flexible
modes
bull energy dissipation through structural damping is included
bull capability to model various mission configurations and maneuvers including
the tether spin about an arbitrary axis
bull user-defined time dependent deployment and retrieval profiles for the tether
as well as the tether attachment point (offset)
bull attitude control algorithm for the tether and rigid bodies using thrusters and
momentum-wheels based on the Feedback Linearization Technique
bull the algorithm for suppression of tether vibration using an active offset (tether
attachment point) control strategy based on the optimal linear control law
( L Q G L T - R )
To begin with in Chapter 2 kinematics and kinetics of the system are derived
using the O(N) formulation methodology developed by Pradhan et al [30] Chapshy
ter 3 discusses issues related to the development of the simulation program and its
validation This is followed by a detailed parametric study of several mission proshy
files simulated as particular cases of the versatile formulation in Chapter 4 Next
the attitude and vibration controllers are designed and their effectiveness assessed in
Chapter 5 The thesis ends with concluding remarks and suggestions for future work
13
2 F O R M U L A T I O N OF T H E P R O B L E M
21 Kinematics
211 Preliminary definitions and the i ^ position vector
The derivation of the equations of motion begins with the definition of the
inertial position of an arbitrary link of the mult ibody system Let the i ^ link of the
system be free to translate and rotate in 3-D space From Figure 2-1 the position
vector Rdm^ to the mass element drrn on the i ^ link wi th reference to the inertial
frame FQ can be written as
Here D represents the inertial position of the t h body fixed frame Fj relative to FQ
ri = [xiyiZi]T is the rigid position vector of drrii wi th respect to F J and f(fi) is
the flexible deformation at fj also with respect to the frame Fj Bo th these vectors
are relative to Fj Note in the present model tethers (i even) are considered flexible
and X corresponds to the nominal position along the unstretched tether while a and
ZJ are by definition equal to zero On the other hand the rigid bodies referred to as
satellites (i odd) including the platform are taken to be rigid ie f(fj) = 0 T j is
the rotation matrix used to express body-fixed vectors with reference to the inertial
frame
212 Tether flexibility discretization
The tether flexibility is discretized with an arbitrary but finite number of
Di + Ti(i + f(i)) (21)
14
15
modes in each direction using the assumed-mode method as
Ui nvi
wi I I i = 1
nzi bull
(22)
where nXi nyi and n 2 ^ are the number of modes used in the X m and Z directions
respectively For the ] t h longitudinal mode the admissible mode shape function is
taken as 2 j - l
Hi(xiii)=[j) (23)
where l is the i ^ tether length[32] In the case of inplane and out-of-plane transverse
deflections the admissible functions are
ampyi(xili) = ampZixuk) = v ^ s i n JKXj
k (24)
where y2 is added as a normalizing factor In this case both the longitudinal and
transverse mode shapes satisfy the geometric boundary conditions for a simple string
in axial and transverse vibration
Eq(22) is recast into a compact matrix form as
ff(i) = $i(xili)5i(t) (25)
where $i(x l) is the matrix of the tether mode shapes defined as
0 0
^i(xiJi) =
^ X bull bull bull ^XA
0
0 0
0
G 5R 3 x n f
(26)
and nfj = nx- + nyi + nZi Note $J(XJJ) is a function of the spatial coordinate x
and not of yi or ZJ However it is also a function of the tether length parameter li
16
which is time-varying during deployment and retrieval For this reason care must be
exercised when differentiating $J(XJJ) with respect to time such that
d d d k) = Jtregi(Xi li) = ^ 7 $ t ( ^ i k)Xi + gf$i(xigt li)k- (2-7)
Since x represents the nominal rate of change of the position of the elemental mass
drrii it is equal to j (deploymentretrieval rate) Let t ing
( 3 d
dx~ + dT)^Xili^ ( 2 8 )
one arrives at
$i(xili) = $Di(xili)ii (29)
The time varying generalized coordinate vector 5 in Eq(25) is composed of
the longitudinal and transverse components ie
ltM) G s f t n f i x l (210)
where 5X^ 5y- and Sz^ are nx- x 1 ny x 1 and nz- x 1 size vectors and correspond to
$ X and $ Z respectively
213 Rotation angles and transformations
The matrix T j in Eq(21) represents a rotation transformation from the body
fixed frame Fj to the inertial frame FQ ie FQ = T j F j It is defined by the 3-2-1
ordered sequence of elementary rotations[46]
1 P i tch F[ = Cf ( a j )F 0
2 R o l l F = C f (3j)F = C7f (3j)Cf ( a j )F 0
3 Yaw F = C7( 7 i )F = C7(7l)cf^)Cf(ai)FQ = Q F 0
17
where
c f (A) =
-Sai Cai
0 0
and
C(7i ) =
0 -s0i] = 0 1 0
0 cpi
i 0 0 = 0
0
(211)
(212)
(213)
Here Eq(211) corresponds to a rotation of ctj (pitch) about the inertial axis ZQ (orbit
normal) of frame FQ resulting in a new intermediate frame F It is then followed by
a roll rotation fa about the axis y of F- giving a second intermediate frame F-
Final ly a spin rotation 7J about the axis z of F- is given resulting in the body
fixed frame Fj for the i lt l link Since all the three rotation matrices are orthogonal
it follows that FQ = C~1Fi = Cj F and hence
C CPi Cai ~ CH Sai + SrYi SPi Sai 5 7 i Sampi + Cli SPi Cci CPi Sai C1i Cai + 5 7 i SPi ~ S1 Ca + C 7 bull Sp Sai
(214) SliCPi CliCPi
Here Cx and Sx are abbreviations for cos(x) and sin(x) respectively
214 Inertial velocity of the ith link
mdash Lett ing pj = fj - f $j(5j and differentiating Eq(21) with respect to time gives
m- D j + T j f j + T^Si + T j pound + T j $ j 5 j (215)
Each of the terms appearing in Eq(215) must be addressed individually Since
fj represents the rigid body motion within the link it only has meaning for the
deployable tether and since yi = Zj = 0 for al l tether links (i even) fj = Ifi where 1
is the unit vector 1 0 0T For the case of rigid links (i odd) fj = 0
18
Evaluating the derivative of each angle in the rotation matrix gives
T^i9i = Taiai9i + Tppigi + T^fagi (216)
where Tx = J j I V Collecting the scaler angular velocity terms and defining the set
Vi = deg^gt A) 7i^ Eq(216) can be rewritten as
T i f t = Pi9i)fji (217)
where Pi(gi) is a column matrix defined by
Pi(9i)Vi = [Taigi^p^iT^gilffi (218)
Inserting Eq(29) and Eq(217) into Eq(215) and defining Sj = 2 + leads to
215 Cylindrical orbital coordinates
The Di term in Eq(219) describes the orbital motion of the t h l ink and
is composed of the three Cartesian coordinates DXi A ^ DZi However over the
cycle of one orbit DXi and Dyi vary dramatically by around the order of Earths
radius This must be avoided since large variations in the coordinates can cause
severe truncation errors during their numerical integration For this reason it is
more convenient to express the Cartesian components in terms of more stationary
variables This is readily accomplished using cylindrical coordinates
However it is only necessary to transform the first l inks orbital components
ie Lgti since only D is a generalized coordinates From Figure 2-2 it is apparent
(219)
that
(220)
19
Figure 2-2 Vector components in cylindrical orbital coordinates
20
Eq(220) can be rewritten in matrix form as
cos(0i) 0 0 s i n ( 0 i ) 0 0
0 0 1
Dri
h (221)
= DTlDSv
where Dri is the inplane radial distance 9 the true anomaly and DZl the out-of-
plane distance normal to the orbital plane The total derivative of D wi th respect
to time gives in column matrix form
1 1 1 1 J (222)
= D D l D 3 v
where [cos(^i) -Dnsm(8i) 0 ]
(223)
For all remaining links ie i = 2 N
cos(0i) - A - j S i n ^ i ) 0 sin(^i) D r i c o s ( 0 i ) 0
0 0 1
DTl = D D i = I d 3 i = 2N (224)
where is the n x n identity matrix and
2N (225)
216 Tether deployment and retrieval profile
The deployment of the tether is a critical maneuver for this class of systems
Cr i t i ca l deployment rates would cause the system to librate out of control Normally
a smooth sinusoidal profile is chosen (S-curve) However one would also like a long
tether extending to ten twenty or even a hundred kilometers in a reasonable time
say 3-4 orbits This is usually difficult or impossible to accomplish solely wi th one
S-curve profile Often a composite S-curve-Steady-S-curve profile is adopted Thus
21
the deployment scheme for the t h tether can be summarized as
k = Tr l - cos (^r(ti - to-)) 1 ltUlt h 2 r V A V
k = Vov tiilttiltt2i
k = IT j 1 - C deg S lkitl ^ ^ lt i lt ^
(226)
where to- tt are the ini t ia l and final times of the deployment or retrieval maneuver
respectively A t j = t - tQ mdash tt mdash t2- and the steady deployment velocity VQ is
calculated based on the continuity of l and l at the specified intermediate times t
and to For the case of
and
Ai = i(ti) - ^(t 0 -) = - ^ ( t 2 i ) (227)
2Ali + li(tf) -h(tQ)
V = lUli l^lL (228)
1 tf-tQ v y
A t = (229)
22 Kinetics and System Energy
221 Kinetic energy
The kinetic energy of the ith link is given by
Kei = lJm Rdrnfi^Ami (230)
Setting Eq(219) in a matrix-vector form gives the relation
4 ^ ^ PM) T i $ 4 TiSi]i[t (231)
22
where
m
k J
x l (232)
and nq = -nfj + 7 is the number of generalized coordinates in each link Inserting
Eq(231) into Eq(230) and integrating the kinetic energy for the t h l ink can be
rewritten as
1 T
(233)
where Mti is the links symmetric mass matrix
Mt =
m i D D T D D DDTPifgidmi) DDTTi J ^dm D n T Sidm rn
sym fP(9i)Pi(9i)dmi J P[g^T^idm J P ( pound ) T ^ m sym sym J ^f^idm f ^f^dmi sym sym sym J sfsidmi
pound Sfttiqxnq
(234)
and mi is the mass of the Ith link The kinetic energy for the entire system ie N
bodies can now be stated as
1 bull T bull 1 bull T Ke = J2 MtA = o laquo Mtqt 2 ^ -i
i=l (235)
where qt = qj^^ bull bull -qfNT a n d Mt is a block diagonal decoupled mass matrix
of the system with Mf- on its diagonal
(236) Mt =
0 0
0 Mt2
0
0 0
MH
0 0 0
0 0 0 bull MtN
222 Simplification for rigid links
The mass matrix given by Eq(234) simplifies considerably for the case of rigid
23
links and is given as
Mti
m D D l
T D D l DD^Piiffidmi) 0
sym 0
0
fP(fi)Pi(i)dmi 0 0 0
0 Id nfj
0 m~
i = 1 3 5 N
(237)
Moreover when mult iplying the matrices in the (11) block term of Eq(237) one
gets
D D l
1 D D l = 1 0 0 0 D n
2 0 (238) 0 0 1
which is a diagonal matrix However more importantly i t is independent of 6 and
hence is always non-singular ensuring that is always invertible
In the (12) block term of M ^ it is apparent that f fidmi is simply the
definition of the centre of mass of the t h rigid link and hence given by
To drrii mdash miff (239)
where fcmi is the position vector of the centre of mass of link i For the case of the
(22) block term a little more attention is required Expanding the integrand using
Eq(218) where ltj = fj gives
fj Ta Tafi fj T Q Tpfi f- T Q T 7 i f j
sym sym
Now since each of the terms of Eq(240) is a scaler or a 1 x 1 matrix it is therefore
equal to its trace ie sum of its diagonal elements thus
tr(ffTai
TTaifi) tr(ffTai
TTpfi) tr(ff Ta^T^fi)
fjT^Tpn rfTp^fi (240)
P(rj)Pj(fi) = syrri trifjTpTTpfi) t r ^ T ^ T ^
sym sym tr(ff T 7 T r 7 f j) J
(241)
Using two properties of the trace function[47] namely
tr(ABC) = tr(BCA) = trCAB) (242)
24
and
j tr(A) =tr(^J A^j (243)
where A B and C are real matrices it becomes quite clear that the (22) block term
of Eq(237) simplifies to
P(fi)Pii)dmi =
tr(Tai
TTaiImi) tTTai
TTpImi) t r (T Q r T 7 7 m ) trTaTTpImi) tr(Tp
TT7Jmi) sym sym sym
(244)
where Im- is an alternate form of the inertia tensor of the t h link which when
expanded is
Imlaquo = I nfi dm bull i sym
sym sym
J x^dmi J Xydmi XZdm sym J yi2dmi f yiZidm
sym J z^dmi J
xVi -L VH sym ijxxi + lyy IZZJ)2
(245)
Note this is in terms of moments of inertia and products of inertia of the t f l rigid
link A similar simplification can be made for the flexible case where T is replaced
with Qi resulting in a new expression for Im
223 Gravitational potential energy
The gravitational potential energy of the ith link due to the central force law
can be written as
V9i f dm f
-A mdash = -A bullgttradelti RdmA
dm
inn Lgti + T j ^ | (246)
where z = 3986 x 105 km 3s 2 (Earths gravitational constant) Expanding binomially
and retaining up to third order terms gives
Vgi ~ ~ D T 2D [J gfgidrrii + 23f T j gidm mdashDT
D2 i l I I Df Tt I grffdmiTiDi
(2-lt 7)
25
where D = D is the magnitude of the inertial position vector to the frame F The
integrals in the above expression are functions of the flexible mode shapes and can be
evaluated through symbolic manipulation using an existing package such as Maple
V In turn Maple V can translate the evaluated integrals into F O R T R A N and store
the code in a file Furthermore for the case of rigid bodies it can be shown quite
readily that
j gjgidrrii = j rjfidrrn = [IXXi + Iyy + IZZi)2 (248)
Similarly the other integrals in Eq(234) can be integrated symbolically
224 Strain energy
When deriving the elastic strain energy of a simple string in tension the
assumptions of high tension and low amplitude of transverse vibrations are generally
valid Consequently the first order approximation of the strain-displacement relation
is generally acceptable However for orbiting tethers in a weak gravitational-gradient
field neglecting higher order terms can result in poorly modelled tether dynamics
Geometric nonlinearities are responsible for the foreshortening effects present in the
tether These cannot be neglected because they account for the heaving motion along
the longitudinal axis due to lateral vibrations The magnitude of the longitudinal
oscillations diminish as the tether becomes shorter[35]
W i t h this as background the tether strain which is derived from the theory
of elastic vibrations[3248] is given by
(249)
where Uj V and W are the flexible deformations along the Xj and Z direction
respectively and are obtained from Eq(25) The square bracketed term in Eq(249)
represents the geometric nonlinearity and is accounted for in the analysis
26
The total strain energy of a flexible link is given by
Ve = I f CiEidVl (250)
Substituting the stress-strain relationship 0 = E(poundi the strain energy is now given
by
Vei=l-EiAi J l l d x h (251)
where E^ is the tethers Youngs modulus and A is the cross-sectional area The
tether is assumed to be uniform thus EA which is the flexural stiffness of the
tether is assumed to be constant Eq(251) is evaluated symbolically for arbitrary
number of modes
225 Tether energy dissipation
The evaluation of the energy dissipation due to tether deformations remains
a problem not well understood even today Here this complex phenomenon is repshy
resented through a simplified structural damping model[32] In addition the system
exhibits hysteresis which also must be considered This is accomplished using an
augmented complex Youngs modulus
EX^Ei+jEi (252)
where Ej^ is the contribution from the structural damping and j = The
augmented stress-strain relation is now given as
a = Epoundi = Ei(l+jrldi)ei (253)
where 77^ = EjE^ ltC 1 is defined as the structural damping coefficient determined
experimentally [49]
27
If poundi is a harmonic function with frequency UJQ- then
jet = (254) wo-
Substituting into Eq(253) and rearranging the terms gives
e + I mdash I poundi = o-i + ad (255)
where ad and a are the stress with and without structural damping respectively
Now the Rayleigh dissipation function[50] for the ith tether can be expressed as
(ii) Jo
_ 1 EjAiVdi rU (2-56^ 2 w 0 Jo
The strain rate ii is the time derivative of Eq(249) The generalized external force
due to damping can now be written as Qd = Q^ bull bull bull QdNT where
dWd
^ = - i p = 2 4 ( 2 5 7 )
= 0 i = 13 N
23 0(N) Form of the Equations of Mot ion
231 Lagrange equations of motion
W i t h the kinetic energy expression defined and the potential energy of the
whole system given by Pe = ]Cn=l( 5i+ ej) the equations of motion can be obtained
quite readily using the Lagrangian principle
where q is the set of generalized coordinates to be defined in a later section
Substituting Eqs(235247251 and 257) into Eq(258) leads to the familiar
matr ix form of the non-linear non-autonomous coupled equations of motion for the
28
system
Mq t)t+ F(q q t) = Q(q q t) (259)
where M(q t) is the nonlinear symmetric mass matrix and F(q q t) is the forcing
function which can be written in matrix form as
Q(ltfgt Qi 0 is the vector of non-conservative generalized forces including control inshy
puts acting on the system A detailed expansion of Eq(260) in tensor notation is
developed in Appendix I
Numerical solution of Eq(259) in terms of q would require inversion of the
mass matrix However M is a full matrix of size nqq x nqq where nqq = Nnq is the
total number of generalized coordinates Therefore direct inversion of M would lead
to a large number of computation steps of the order nqq^ or higher The objective
here is to develop an order-N O(N) algorithm for obtaining M _ 1 that minimizes
the computational effort This is accomplished through the following transformations
ini t ia l ly developed for the stationkeeping case by Pradhan et al [30]
232 Generalized coordinates and position transformation
During the derivation of the energy expressions for the system the focus was
on the decoupled system ie each link was considered independent of the others
Thus each links energy expression is also uncoupled However the constraints forces
between two links must be incorporated in the final equations of motion
Let
e Unqxl (261) m
be the set of coupled coordinates of the t h l ink such that q = qT q2 bull bull bull 0T i s the
29
systems generalized coordinates Let qt- be the set of auxiliary decoupled coordinates
defined by Eq(232) such that qt = qj^ qj^ Qtj^1 bull The only difference between
qtj and q is the presence of D against d is the position of F wi th respect to
the inertial frame FQ whereas d is defined as the offset position of F relative to and
projected on the frame It can be expressed as d = dXidyidZi^ For the
special case of link 1 D = d
From Figure 2-1 can written as
Di = A - l + T i _ i J i _ i + di) + T V i S i - i f t - i + dXi)8i- (262)
Denoting KAj = $j(Zj + dx- ) Eq(262) can be rewritten in summation form as
i ampi = H ( T i - 1 ^ + T i - l K A i - l - l + T j - l ^ - l ) bull ( 2 - 6 3 )
3 = 1
Introducing the index substitution k mdash j mdash 1 into Eq(263)
i-1 D = ^2 (Tfc_ilt4 + TkKAk6k + Tkilk) + T^di (264)
k=l
since d = D TQ = 1^ and IQ DQ SQ are null vectors Recasting Eq(264) in
matr ix form leads to i-1
Qt Jfe=l
where
and
For the entire system
T fc-1 0 TkKAk
0 0 0 0 0 0 0 0 0 0 0 0 -
r T i - i 0 0 0 0 0 0 0 0 0 G
0 0 0 1
Qt = RpQ
G R n lt x l
(265)
(266)
G Unq x l (267)
(268)
30
where
RP
R 1 0 0 R R 2 0 R RP
2 R 3
0 0 0
R RP R PN _ 1 RN J
(269)
Here RP is the lower block triangular matrix which relates qt and q
233 Velocity transformations
The two sets of velocity coordinates qti and qi are related by the following
transformations
(i) First Transformation
Differentiating Eq(262) with respect to time gives the inertial velocity of the
t h l ink as
D = A - l + Ti_iK + + K A i _ i ^ _ i
+ T j _ i ^ + Ti-ii + T i - i K A i - i ^ + T j - i K A ^ i ^ i (270)
(271) (
Defining K A D = dddXi+1KAi K A L = d d K A and hi = di+1 +lfi +KA^i
results in K A ^ i = K A D j i o ^ + K A L i _ i i _ i
= KADi_iiTdi + K A L j _ i ^ _ i
Inserting Eq(271) into Eq(270) and rearranging gives
= A _ + T i ^ d g + K A D ^ x ^ i ^ ^ + Pi-^hi-iH-i
+ T i _ i K A i _ i 5 _ i + T i_ i2 + K A L j _ i 5 j _ i 4 _ i (272)
Using Eq(272) to define recursively - D j _ i and applying the index substitution
c = j mdash 1 as in the case of the position transformation it follows that
Di = ^ T ^ i K D ^ ^ + Pk(hk)ffk + T f c K A f c ^ + TfcKLfc4 Jfc=l (273)
+ T j _ i K D j _ i lt i ii
31
where K D j _ i = + K A D ^ i ^ - i r 7 and K L j = i + K A L j 5 j Finally by following a
procedure similar to that used for the position transformation it can be shown that
for the entire system
t = Rvil (2-74)
where Rv is given by the lower block diagonal matrix
Here
Rv
Rf 0 0 R R$ 0 R Rl RJ
0 0 0
T3V nN-l
R bullN
e Mnqqxl
Ri =
T i _ i K D i _ i TiKAi T j K L j 0 0 0 0 bullo 0 0 0 0 0 0 0
e K n lt x l
RS
T i-1 0 0 o-0 0 0 0 0 0
0 0 0 1
G Mnq x l
(275)
(276)
(277)
(ii) Second Transformation
The second transformation is simply Eq(272) set into matrix form This leads
to the expression for qt as
(278)
where Id Pihi) T j K A j T j K L j 0 0 0 0 0 0
0 0
0 0
0 0
G $lnq x l
Thus for the entire system
(279)
pound = Rnjt + Rdq
= Vdnqq - R nrlRdi (280)
32
where
and
- 0 0 0 RI 0 0
Rn = 0 0
0 0
Rf 0 0 0 rgtd K2 0
Rd = 0 0 Rl
0 0 0
RN-1 0
G 5R n w x l
0 0 0 e 5 R N ^ X L
(281)
(282)
234 Cylindrical coordinate modification
Because of the use of cylindrical coorfmates for the first body it is necessary
to slightly modify the above matrices to account for this coordinate system For the
first link d mdash D = D ^ D S L thus
Qti = 01 =gt DDTigi (283)
where
D D T i
Eq(269) now takes the form
DTL 0 0 0 d 3 0 0 0 ID
0 0 nfj 0
RP =
P f D T T i R2 0 P f D T T i Rp
2 R3
iJfDTTi i^
x l
0 0 0
R PN-I RN
(284)
(285)
In the velocity transformation a similar modification is necessary since D =
DD^DS^ Mak ing the substitution it can be shown that Eq(276) and Eq(279) now
33
takes the form for i = 1
T)V _ pn _ r t j mdash rt^ mdash 0 0 0 0
0 0 0 0 0 0 0 0
(286)
The rest of the matrices for the remaining links ie i = 2 N are unchanged
235 Mass matrix inversion
Returning to the expression for kinetic energy substitution of Eq(274) into
Eq(235) gives the first of two expressions for kinetic energy as
Ke = -q RvT MtRv (287)
Introduction of Eq(280) into Eq(235) results in the second expression for Ke as
1X 2q (hnqq-RnrlRd Mt (idnqq - R nrlRd (288)
Thus there are two factorizations for the systems coupled mass matrix given as
M = RV MtRv
M (hnqq - RnrlRd Mt ( i d n q q - R n r 1 R d
(289)
(290)
Inverting Eq(290)
r-1 _ ( rgtd~^ M 1 = (Ra) l d n q q - Rn)Mf1 (Rd)~l(Idnqq - Rn) (291)
Since both Rd and Mt are block diagonal matrices their inverse is simply the inverse
34
of each block on the diagonal ie
Mr1
h 0 0
0 0
0 0
0 0 0
0
0
0
tN
(292)
Each block has the dimension nq x nq and is independent of the number of bodies
in the system thus the inversion of M is only linearly dependent on N ie an O(N)
operation
236 Specification of the offset position
The derivation of the equations of motion using the O(N) formulation requires
that the offset coordinate d be treated as a generalized coordinate However this
offset may be constrained or controlled to follow a prescribed motion dictated either by
the designer or the controller This is achieved through the introduction of Lagrange
multipliers[51] To begin with the multipliers are assigned to all the d equations
Thus letting f = F - Q Eq(259) becomes
Mq + f = P C A (293)
where A G K 3 - 7 V x l is the vector of Lagrange multipliers and Pc G s R n 9 x 3 A r is the
permutation matrix assigning the appropriate Aj to its corresponding d equation
Inverting M and pre-multiplying both sides by PcT gives
pcTM-pc
= ( dc + PdegTM-^f) (294)
where dc is the desired offset acceleration vector Note both dc and PcTM lf are
known Thus the solution for A has the form
A pcTM-lpc - l dc + PC1 M (295)
35
Now substituting Eq(295) into Eq(293) and rearranging the terms gives
q = S(f Ml) = -M~lf + M~lPc
which is the new constrained vector equation of motion with the offset specified by
dc Note that pre-multiplying M~l by P c T provides the rows of M _ 1 corresponding
to the d equations Similarly post-multiplying by Pc leads the columns of the matrix
corresponding to the d equations
24 Generalized Control Forces
241 Preliminary remarks
The treatment of non-conservative external forces acting on the system is
considered next In reality the system is subjected to a wide variety of environmental
forces including atmospheric drag solar radiation pressure Earth s magnetic field as
well as dissipative devices to mention a few However in the present model only
the effect of active control thrusters and momemtum-wheels is considered These
generalized forces wi l l be used to control the attitude motion of the system
In the derivation of the equations of motion using the Lagrangian procedure
the contribution of each force must be properly distributed among all the generalized
coordinates This is acheived using the relation
nu ^5 Qk=E Pern ^ (2-97)
m=l deg Q k
where Qk and qk are the kth elements of Q and ltf respectively[51] Fem is the
mth external force acting on the system at the location Rm from the inertial frame
and nu is the total number of actuators Derivation of the generalized forces due to
thrusters and momentum wheels is given in the next two subsections
pcTM-lpc (dc + PcTM- Vj (2-96)
36
242 Generalized thruster forces
One set of thrusters is placed on each subsatellite ie al l the rigid bodies
except the first (platform) as shown in Figure 2-3 They are capable of firing in the
three orthogonal directions From the schematic diagram the inertial position of the
thruster located on body 3 is given as
Defining
R3 = di + Tid2 + T2(hi + KA2lt52) + T 3 f c m 3
= d + dfY + T 3 f c m 3
df = Tjdj+i + Tj+1lj+ii + KAj+i5j+i)
the thruster position for body 5 is given as
(298)
(299)
R5 = di+ dfx + d 3 + T5fcm
Thus in general the inertial position for the thruster on the t h body can be given
(2100)
as i - 2
k=l i1 cm (2101)
Let the column matrix
[Q i iT 9 R-K=rR dqj lA n i
WjRi A R i (2102)
where i = 3 5 N and j mdash 1 2 i The vector derivative of R wi th respect to
the scaler components qk is stored in the kth column of Thus for the case of
three bodies Eq(297) for thrusters becomes
1 Q
Q3
3 -1
T 3 T t 2 (2103)
37
Satellite 3
mdash gt -raquo cm
D1=d1
mdash gt
a mdash gt
o
Figure 2-3 Inertial position of subsatellite thruster forces
38
where ft = 0TtaTt^T is the thrust vector for the t h l ink (tether i-1) Here
Ttn and Tta are the thrust forces along the inplane m and out-of-plane Z directions i Pi
respectively For the case of 5-bodies
Q
~Qgt1
0
0
( T ^ 2 ) (2104)
The result for seven or more bodies follows the pattern established by Eq(2104)
243 Generalized momentum gyro torques
When deriving the generalized moments arising from the momemtum-wheels
on each rigid body (satellite) including the platform it is easier to view the torques as
coupled forces From Figure 2-4 it is apparent that for the ith l ink the generalized
forces constituting the couple are
(2105)
where
Fei = Faij acting at ei = eXi
Fe2 = -Fpk acting at e 2 = C a ^ l
F e 3 = F7ik acting at e 3 = eyj
Expanding Eq(2105) it becomes clear that
3
Qk = ^ F e i
i=l
d
dqk (2106)
which is independent of Ri- Thus the moments Mma = 2FaieXi Mm^ = 2FpeXi
and Mmi = 2FlieVi Transforming the coordinates to the body fixed frame using
the rotation matrix the generalized force due to the momentum-wheels on link i can
39
Figure 2-4 Coupled force representation of a momentum-wheel on a rigid body
40
be written as
d d d Qi = Td bull ^ T i t M m a - Tik bull mdashT^Mmp + T ^ bull T J j M ^
QiMmi
where Mm = Mma M m Mmi T
lt3 T J ^ T - T i f c bull ^ T i pound T ^ - ^ T z j
(2107)
(2108)
Note combining the thruster forces and momentum-wheel torques a compact
expression for the generalized force vector for 5 bodies can be written as
Q
Q T 3 0 Q T 5 0
0 0 0
0 lt33T3 Qyen Q5
3T5 0
0 0 0 Q4T5 0
0 0 0 Q5
5T5
M M I
Ti 2 m 3
Quu (2109)
The pattern of Eq(2109) is retained for the case of N links
41
3 C O M P U T E R I M P L E M E N T A T I O N
31 Prel iminary Remarks
The governing equations of motion for the TV-body tethered system were deshy
rived in the last chapter using an efficient O(N) factorization method These equashy
tions are highly nonlinear nonautonomous and coupled In order to predict and
appreciate the character of the system response under a wide variety of disturbances
it is necessary to solve this system of differential equations Although the closed-
form solutions to various simplified models can be obtained using well-known analytic
methods the complete nonlinear response of the coupled system can only be obtained
numerically
However the numerical solution of these complicated equations is not a straightshy
forward task To begin with the equations are stiff[52] be they have attitude and
vibration response frequencies separated by about an order of magnitude or more
which makes their numerical integration suceptible to accuracy errors i f a properly
designed integration routine is not used
Secondly the factorization algorithm that ensures efficient inversion of the
mass matrix as well as the time-varying offset and tether length significantly increase
the size of the resulting simulation code to well over 10000 lines This raises computer
memory issues which must be addressed in order to properly design a single program
that is capable of simulating a wide range of configurations and mission profiles It
is further desired that the program be modular in character to accomodate design
variations with ease and efficiency
This chapter begins with a discussion on issues of numerical and symbolic in-
42
tegration This is followed by an introduction to the program structure Final ly the
validity of the computer program is established using two methods namely the conshy
servation of total energy and comparison of the response results for simple particular
cases reported in the literature
32 Numerical Implementation
321 Integration routine
A computer program coded in F O R T R A N was written to integrate the equashy
tions of motion of the system The widely available routine used to integrate the difshy
ferential equations D G E A R [ 5 3 ] is based on Gears predictor-corrector method[54]
It is well suited to stiff differential equations as i t provides automatic step-size adjustshy
ment and error-checking capabilities at each iteration cycle These features provide
superior performance over traditional 4^ order Runge-Kut ta methods
Like most other routines the method requires that the differential equations be
transformed into a first order state space representation This is easily accomplished
by letting
thus
X - ^ ^ f ^ ) =ltbullbull) (3 -2 )
where S(fMdc) is defined by Eq(296) Eq(32) represents the new set of 2nqq
first order differential equations of motion
322 Program structure
A flow chart representing the computer programs structure is presented in
Figure 3-1 The M A I N program is responsible for calling the integration routine It
also coordinates the flow of information throughout the program The first subroutine
43
called is INIT which initializes the system variables such as the constant parameters
(mass density inertia etc) and the ini t ial conditions I C (attitude angles orbital
motion tether elastic deformation etc) from user-defined data files Furthermore all
the necessary parameters required by the integration subroutine D G E A R (step-size
tolerance) are provided by INIT The results of the simulation ie time history of the
generalized coordinates are then passed on to the O U T P U T subroutine which sorts
the information into different output files to be subsequently read in by the auxiliary
plott ing package
The majority of the simulation running time is spent in the F C N subroutine
which is called repeatedly by D G E A R F C N generates the fv(Xi) vector defined in
Eq(32) It is composed of several subroutines which are responsible for the comshy
putation of all the time-varying matrices vectors and scaler variables derived in
Chapter 2 that comprise the governing equations of motion for the entire system
These subroutines have been carefully constructed to ensure maximum efficiency in
computational performance as well as memory allocation
Two important aspects in the programming of F C N should be outlined at this
point As mentioned earlier some of the integrals defined in Chapter 2 are notably
more difficult to evaluate manually Traditionally they have been integrated numershy
ically on-line (during the simulation) In some cases particularly stationkeeping the
computational cost is quite reasonable since the modal integrals need only be evalushy
ated once However for the case of deployment or retrieval where the mode shape
functions vary they have to be integrated at each time-step This has considerable
repercussions on the total simulation time Thus in this program the integrals are
evaluated symbolically using M A P L E V[55] Once generated they are translated into
F O R T R A N source code and stored in I N C L U D E files to be read by the compiler
Al though these files can be lengthy (1000 lines or more) for a large number of flexible
44
INIT
DGEAR PARAM
IC amp SYS PARAM
MAIN
lt
DGEAR
lt
FCN
OUTPUT
STATE ACTUATOR FORCES
MOMENTS amp THRUST FORCES
VIBRATION
r OFFS DYNA
ET MICS
Figure 3-1 Flowchart showing the computer program structure
45
modes they st i l l require less time to compute compared to their on-line evaluation
For the case of deployment the performance improvement was found be as high as
10 times for certain cases
Secondly the O(N) algorithm presented in the last section involves matrices
that contained mostly zero terms (RP Rv and Mplusmn) The sparse structure of these mashy
trices can be used advantageously to improve the computational efficiency by avoiding
multiplications of zero elements The result is a quick evaluation of fv(Xt) Moreshy
over there is a considerable saving of computer memory as zero elements are no
longer stored
Final ly the C O N T R O L subroutine which is part of F C N is designed to calshy
culate the appropriate control forces torques and offset dynamics which are required
to stabilize the librational ( A T T I T U D E ) and vibrational ( V I B R A T I O N ) motion of
the system
33 Verification of the Code
The simulation program was checked for its validity using two methods The
first verified the conservation of total energy in the absence of dissipation The second
approach was a direct comparison of the planar response generated by the simulation
program with those available in the literature A s numerical results for the flexible
three-dimensional A^-body model are not available one is forced to be content with
the validation using a few simplified cases
331 Energy conservation
The configuration considered here is that of a 3-body tethered system with
the following parameters for the platform subsatellite and tether
46
h =
1091430 -8135 328108 -8135 8646050 27116 328108 27116 8286760
kg bull m 2 platform inertia
kg bull m 2 end-satellite inertia 200 0 0
0 400 0 0 0 400
m = 90000 kg mass of the space station platform
m2 = 500 kg mass of the end-satellite
EfAt = 61645 N tether elastic stiffness
pt mdash 49 kg km tether linear density
The tether length is taken to remain constant at 10 km (stationkeeping case)
A n in i t ia l disturbance in pitch and roll of 2deg and 1deg respectively is given to both
the rigid bodies and the flexible tether In addition the tether is ini t ia l ly deflected
by 45 m in the longitudinal direction at its end and 05 m in both the inplane and
out-of-plane directions in the first mode The tether attachment point offset at the
platform (satellite 1) end is 1 m in all the three directions ie d2 = 11 lTm The
attachment point at the subsatellite (satellite 2) end is 1 m in the x direction ie
fcm3 = 10 0T
The variation in the kinetic and potential energies is presented for the above-
mentioned case in Figure 3-2(a) As seen from the plot there is a continuous mutual
exchange between the kinetic and potential energies however the variation in the
total energy remains zero (Figure 3-2b) It should be noted that after 1 orbit the
variation of both kinetic and potential energy does not return to 0 This is due to
the attitude and elastic motion of the system which shifts the centre of mass of the
tethered system in a non-Keplerian orbit
Similar results are presented for the 5-body case ie a double pendulum (Figshy
ure 3-3) Here the system is composed of a platform and two subsatellites connected
47
STSS configuration 3-Body a 1(0)=a 2(0)=a t(0)=2deg P1(0)=p2(0)=p(0)=1deg 5x(0)=45m 8y(0)=52(0)=05m dx(0)=dy(0)=d2(0)=1m Stationkeeping lt=10km
Variation of Kinetic and Potential Energy
(a) Time (Orbits) Percent Variation in Total Energy
OOEOh
-50E-12
-10E-11 I i i
0 1 2 (b) Time (Orbits)
Figure 3-2 Kinetic and potential energy transfer for the three-body platform based tethered satellite system (a) variation of kinetic and potential energy (b) percent change in total energy of system
48
STSS configuration 5-Body a 1(0)=a 2(0)=a t 1(0)=2deg a 3(0)=a t 2(0)=25deg P1(0)=p2(0)=p3(0)=pt1(0)=pt2(0)=1lt
5x(0)=5m 8y(0)=5z(0)=05m dx(0)=dy(0)=dz(0)=0 Stationkeeping 1 =12=1 Okm
(a)
20E6
10E6r-
00E0
-10E6
-20E6
Variation of Kinetic and Potential Energy
-
1 I I I 1 I I
I I
AP e ~_
i i 1 i i i i
00E0
UJ -50E-12 U J
lt
Time (Orbits) Percent Variation in Total Energy
-10E-11h
(b) Time (Orbits)
Figure 3 -3 Kinet ic and potential energy transfer for the five-body platform based tethered satellite system (a) variation of kinetic and potential enshyergy (b) percent change in total energy of system
49
in sequence by two tethers The two subsatellites have the same mass and inertia
ie 7713 = 7 7 i 2 a n d I3 = I21 a n d are separated by identical tether segments each 10
k m in length The platform has the same properties as before however there is no
tether attachment point offset present in the system (d = fcmi = 0)
In al l the cases considered energy was found to be conserved However it
should be noted that for the cases where the tether structural damping deployment
retrieval attitude and vibration control are present energy is no longer conserved
since al l these features add or subtract energy from the system
332 Comparison with available data
The alternative approach used to validate the numerical program involves
comparison of system responses with those presented in the literature A sample case
considered here is the planar system whose parameters are the same as those given
in Section 331 A n ini t ia l pitch disturbance of 2deg is imparted to both the platform
and the tether together with an ini t ial tether deformation of 08 m and 001 m in
the longitudinal and transverse directions respectively in the same manner as before
The tether length is held constant at 5 km and it is assumed to be connected to the
centre of mass of the rigid platform and subsatellite The response time histories from
Ref[43] are compared with those obtained using the present program in Figures 3-4
and 3-5 Here ap and at represent the platform and tether pitch respectively and B i
C i are the tethers longitudinal and transverse generalized coordinates respectively
A comparison of Figures 3-4 and 3-5 clearly shows a remarkable correlation
between the two sets of results in both the amplitude and frequency The same
trend persisted even with several other comparisons Thus a considerable level of
confidence is provided in the simulation program as a tool to explore the dynamics
and control of flexible multibody tethered systems
50
a p(0) = 2deg 0(0) = 2C
6(0) = 08 m
0(0) = 001 m
Stationkeeping L = 5 km
D p y = 0 D p 2 = 0
a oo
c -ooo -0 01
Figure 3 - 4 Simulation results for the platform based three-body tethered system originally presented in Ref[43]
51
STSS configuration 3-Body a1(0)=a2(0)=at(0)=2deg p1(0)=(32(0)=(3t(0)=0 8x(0)=08m 5y(0)=001m 5Z(0)=0 dx(0)=dy(0)=dz(0)=0 Stationkeeping l=5km
Satellite 1 Pitch Angle Tether Pitch Angle
i -i i- i i i i i i d
1 2 0 1 2 Time (Orbits) Time (Orbits)
Longitudinal Vibration
^MAAAAAAAWVWWWWVVVW
V bull i bull i I i J 00 05 10 15 20
Time (Orbits) Transverse Vibration
Time (Orbits)
Figure 3-5 Simulation results for the platform based three-body tethered system obtained using the present computer program
52
4 D Y N A M I C SIMULATION
41 Prel iminary Remarks
Understanding the dynamics of a system is critical to its design and develshy
opment for engineering application Due to obvious limitations imposed by flight
tests and simulation of environmental effects in ground based facilities space based
systems are routinely designed through the use of numerical models This Chapter
studies the dynamical response of two different tethered systems during deployment
retrieval and stationkeeping phases In the first case the Space platform based Tethshy
ered Satellite System (STSS) which consists of a large mass (platform) connected to
a relatively smaller mass (subsatellite) with a long flexible tether is considered The
other system is the O E D I P U S B I C E P S configuration involving two comparable mass
satellites interconnected by a flexible tether In addition the mission requirement of
spin about an arbitrary axis the tether axis for O E D I P U S - A C and cartwheeling or
spin about the orbit normal for the proposed B I C E P S mission is accounted for
A s mentioned in Chapter 2 the tether flexibility is modelled using the asshy
sumed mode discretization method Although the simulation program can account
for an arbitrary number of vibrational modes in each direction only the first mode
is considered in this study as it accounts for most of the strain energy[34] and hence
dominates the vibratory motion The parametric study considers single as well as
double pendulum type systems with offset of the tether attachment points The
systems stability is also discussed
42 Parameter and Response Variable Definitions
The system parameters used in the simulation unless otherwise stated are
53
selected as follows for the STSS configuration
1091430 -8135 328108 -8135 8646050 27116 328108 27116 8286760
h =
h = 200 0 0
0 400 0 0 0 400
bull m = 90000 kg (mass of the space platform)
kg bull m 2 (platform inertia) 760 J
kg bull m 2 (subsatellite inertia)
bull 7772 = 500 kg (mass of the subsatellite)
bull EtAt = 61645 N (tether stiffness)
bull Pt = 49 k g k m (tether density)
bull Vd mdash 05 (tether structural damping coefficient)
bull fcm^ = l 0 0 r m (tether attachment point at the subsatellite)
The response variables are defined as follows
bull ai0 satellite 1 (platform) pitch and roll angles respectively
bull CK2 2 satellite 2 (subsatellite) pitch and roll angles respectively
bull at fit- tether pitch and roll angles respectively
bull If tether length
bull d = dx dy dzT- tether attachment position relative to satellite 1 (platform)
mdash
bull S = 8X 8y Sz tether flexible generalized coordinates in the longitudinal x
inplane transverse y and out-of-plane transverse z directions respectively
The attitude angles and are measured with respect to the Local Ve r t i c a l -
Loca l Horizontal ( L V L H ) frame A schematic diagram illustrating these variable
is presented in Figure 4-1 The system is taken to be in a nominal circular orbit at
an altitude of 289 km with an orbital period of 903 minutes
54
Satellite 1 (Platform) Y a w
Figure 4-1 Schematic diagram showing the generalized coordinates used to deshyscribe the system dynamics
55
43 Stationkeeping Profile
To facilitate comparison of the simulation results a reference case based on
the three-body STSS system with zero offset (d = 0 r c m 3 = 0) and a tether length
of It = 20 km is considered first (Figure 4-2) The system is subjected to an ini t ial
disturbance in pitch and roll of 2deg and 1deg respectively to all the three bodies
In addition an ini t ia l longitudinal deflection of 14 m from the tethers unstretched
position is given together with a transverse inplane and out-of-plane deflection of 1
m at the tethers mid-length From the attitude response given in Figure 4-2(a) it is
apparent that the three bodies oscillate about their respective equilibrium positions
Since coupling between the individual rigid-body dynamics is absent (zero offset) the
corresponding librational frequencies are unaffected Satellite 1 (Platform) displays
amplitude modulation arising from the products of inertia The result is a slight
increase of amplitude in the pitch direction and a significantly larger decrease in
amplitude in the roll angle
Figure 4-2(b) clearly shows coupling of the tethers attitude motion with its
flexibility dynamics The tether vibrates in the axial direction about its equilibrium
position (at approximately 138 m) with two vibration frequencies namely 012 Hz
and 31 x 1 0 - 4 Hz The former is the first longitudinal flexible mode which is dissishy
pated through the structural damping while the latter arises from the coupling with
the pitch motion of the tether Similarly in the transverse direction there are two
frequencies of oscillations arising from the flexible mode and its coupling wi th the
attitude motion As expected there is no apparent dissipation in the transverse flexshy
ible motion This is attributed to the weak coupling between the longitudinal and
transverse modes of vibration since the transverse strain in only a second order effect
relative to the longitudinal strain where the dissipation mechanism is in effect Thus
there is very litt le transfer of energy between the transverse and longitudinal modes
resulting in a very long decay time for the transverse vibrations
56
STSS configuration 3-Body a1(0)=a2(0)=a(0)=2deg (31(0)=p2(0)=p(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=dz(0)=0 Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
0 1 2 3 4 5 0 1 2 3 4 5 Time (Orbits) Time (Orbits)
Satellite 2 Pitch Angle Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-2 Stationkeeping dynamics of the three-body STSS configuration withshyout offset (a) attitude response
57
STSS configuration 3-Body a1(0)=cx2(0)=cx(0)=2deg P1(0)=p2(0)=p(0)=1deg 5x(0)=14m8y(0)=82(0)=1m dx(0)=dy(0)=dz(0)=0 Stationkeeping lt=20km
Tether Longitudinal Vibration 16| 1 1 1 1 1 r-
Time (Orbits) Tether In-Plane Transverse Vibration
imdash 1 mdash 1 mdash 1 mdash 1 mdashr
r I I 1 1 1 I 1 1 1 1 I 1 bull ^ bull bull I I J 0 1 2 3 4 5
Time (Orbits) Tether Out-of-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-2 Stationkeeping dynamics of the three-body STSS configuration withshyout offset (b) vibration response
58
Introduction of the attachment point offset significantly alters the response of
the attitude motion for the rigid bodies W i t h a i m offset along the local vertical
at the platform (dx = 1 m) and fcm^ = 100 m coupling between the tether
and end-bodies is established by providing a lever arm from which the tether is able
to exert a torque that affects the rotational motion of the rigid end-bodies (Figure
4-3) Bo th the platform and subsatellite now oscillate at the same frequency as the
tether whose motion remains unaltered In the case of the platform there is also an
amplitude modulation due to its non-zero products of inertia as explained before
However the elastic vibration response of the tether remains essentially unaffected
by the coupling
Providing an offset along the local horizontal direction (dy = 1 m) results in
a more dramatic effect on the pitch and roll response of the platform as shown in
Figure 4-4(a) Now the platform oscillates about its new equilibrium position of
-90deg The roll motion is also significantly disturbed resulting in an increase to over
10deg in amplitude On the other hand the rigid body dynamics of the tether and the
subsatellite as well as the tether flexible motion remain the same (Figure 4-4b)
Figure 4-5 presents the response when the offset of 1 m at the platform end is
along the z direction ie normal to the orbital plane The result is a larger amplitude
pitch oscillation about the reference equilibrium position as shown in Figure 4-5(a)
while the roll equilibrium is now shifted to approximately 90deg Note there is little
change in the attitude motion of the tether and end-satellites however there is a
noticeable change in the out-of-plane transverse vibration of the tether (Figure 4-5b)
which may be due to the large amplitude platform roll dynamics
Final ly by setting a i m offset in all the three direction simultaneously the
equil ibrium position of the platform in the pitch and roll angle is altered by approxishy
mately 30deg (Figure 4-6a) However the response of the system follows essentially the
59
STSS configuration 3-Body a1(0)=cx2(0)=cxt(0)=2deg p1(0)=(32(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=1m dy(0)=dz(0)=0 Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
Time (Orbits) Time (Orbits) Tether Pitch Angle Tether Roll Angle
L bull I I bull bull lt bull bull bull raquo bull bull bull bull -I h I I I I i H 0 1 2 3 4 5 0 1 2 3 4 5
Time (Orbits) Time (Orbits) Satellite 2 Pitch Angle Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-3 Stationkeeping dynamics of the three-body STSS configuration with offset along the local vertical (a) attitude response
60
STSS configuration 3-Body a 1(0)=a 2(0)=a t(0)=2deg (31(0)=p2(0)=pt(0)=1deg 5x(0)=14m5y(0)=5z(0)=1m dx(0)=1mdy(0)=dz(0)=0 Stationkeeping lt=20km
Tether Longitudinal Vibration 1 6 i mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash mdash i mdash i mdash mdash mdash mdash mdash i mdash mdash lt -
Time (Orbits)
Tether In-Plane Transverse Vibration
y bull i i i i i i i i i i i i i i i i i 0 1 2 3 4 5
Time (Orbits)
Tether Out-of-Plane Transverse Vibration r 1 1 1 1 1 1 1 1 1 1 1 1 1 1 q
h i I i i I i i i i I i i i i 3
0 1 2 3 4 5 (b) Time (Orbits)
Figure 4-3 Stationkeeping dynamics of the three-body STSS configuration with offset along the local vertical (b) vibration response
61
STSS configuration 3-Body a1(0)=cx2(0)=o(0)=2deg (31(0)=p2(0)=p(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dz(0)=0 dy(0)=1m Stationkeeping lt=20km
Satellite 1 Pitch Angle i 1 i
Satellite 1 Roll Angle
1 2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 4 Time (Orbits)
Tether Roll Angle
CD
edT
1 2 3 Time (Orbits)
Satellite 2 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 2 Roll Angle
CD
C O
(a)
Figure 4-
1 2 3 4 Time (Orbits)
1 2 3 4 Time (Orbits)
4 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal (a) attitude response
62
STSS configuration 3-Body a1(0)=cx2(0)=ot(0)=2deg p1(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dz(0)=0 dy(0)=1m Stationkeeping lt=20km
Tether Longitudinal Vibration 16| 1 gt 1 bull 1 1 1
0 1 2 3 4 5
Time (Orbits) Tether In-Plane Transverse Vibration
i mdash 1 mdash 1 mdash 1 mdash 1 mdash i mdash mdash lt mdash 1 mdash 1 mdash i mdash mdash 1 mdash 1 mdash
V I I bull bull bull I i i i I i i 1 1 3
0 1 2 3 4 5 Time (Orbits)
Tether Out-of-Plane Transverse Vibration r 1 1 1 1 1 1 1 1 1 1 1 gt bull | q
V i 1 1 1 1 1 mdash 1 mdash 3
0 1 2 3 4 5 (b) Time (Orbits)
Figure 4-4 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal (b) vibration response
63
STSS configuration 3-Body a1(0)=o2(0)=at(0)=2deg (31(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=0 dz(0)=1m Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 4 Time (Orbits)
Tether Roll Angle
1 2 3 4 Time (Orbits)
Satellite 2 Pitch Angle
2 3 4 Time (Orbits)
Satellite 2 Roll Angle
(a)
Figure 4-
1 2 3 4 Time (Orbits)
2 3 4 Time (Orbits)
5 Stationkeeping dynamics of the three-body STSS configuration with offset along the orbit normal (a) attitude response
64
STSS configuration 3-Body a1(0)=cx2(0)=ot(0)=2deg P1(0)=p2(0)=(3(0)=1deg 8x(0)=14m8y(0)=82(0)=1m dx(0)=dy(0)=0 d2(0)=1m Stationkeeping lt=20km
Tether Longitudinal Vibration 1 6 i mdash r mdash i mdash mdash i mdash i mdash mdash i mdash i mdash mdash i mdash mdash i mdash 1 mdash 1 mdash i mdash mdash lt mdash lt mdash lt mdash r
Time (Orbits) Tether In-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-5 Stationkeeping dynamics of the three-body STSS configuration with offset along the orbit normal (b) vibration response
65
same trend with only minor perturbations in the transverse elastic vibratory response
of the tether (Figure 4-6b)
44 Tether Deployment
Deployment of the tether from an ini t ial length of 200 m to 20 km is explored
next Here the deployment length profile is critical to ensure the systems stability
It is not desirable to deploy the tether too quickly since that can render the tether
slack Hence the deployment strategy detailed in Section 216 is adopted The tether
is deployed over 35 orbits with the sinusoidal acceleration and deceleration occurring
over a 4 km length ( A ^ 2 ) with the remaining deployment at a constant velocity of
V0 = 146 ms
Initially the tether is stretched only slightly S = 1304 x 1 0 _ 3 0 0 m
with al l the other states of the system ie pitch roll and transverse displacements
remaining zero The response for the zero offset case is presented in Figure 4-7 For
the platform there is no longer any coupling with the tether hence it oscillates about
the equilibrium orientation determined by its inertia matrix However there is st i l l
coupling between the subsatellite and the tether since rcm^ mdash 100 m resulting
in complete domination of the subsatellite dynamics by the tether Consequently
the pitch motion ini t ial ly grows by almost 50deg but eventually subsides as the tether
deployment rate decreases It is interesting to note that the out-of-plane motion is not
affected by the deployment This is because the Coriolis force which is responsible
for the tether pitch motion does not have a component in the z direction (as it does
in the y direction) since it is always perpendicular to the orbital rotation (Q) and
the deployment rate (It) ie Q x It
Similarly there is an increase in the amplitude of the transverse elastic oscilshy
lations of the tether again due to the Coriolis effect (Figure 4-7b) Moreover as the
66
STSS configuration 3-Body a1(0)=a2(0)=cxt(0)=2deg P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m5y(0)=52(0)=1m dx(0)=dy(0)=d2(0)=1m Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
i i i i i i i i i i i 0 1 2 3 4 5 0 1 2 3 4 5
Time (Orbits) Time (Orbits) Tether Pitch Angle Tether Roll Angle
l i i i i -I h i i i i I i i i i l- i i i i l i i i i I i i- i i H 0 1 2 3 4 5 0 1 2 3 4 5
Time (Orbits) Time (Orbits) Satellite 2 Pitch Angle Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-6 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal local vertical and orbit normal (a) attitude response
67
STSS configuration 3-Body a1(0)=cc2(0)=a(0)=2deg p1(0)=p2(0)=(3t(0)=1deg 5x(0)=14m6y(0)=52(0)=1m dx(0)=dy(0)=d2(0)=1m Stationkeeping lt=20km
Tether Longitudinal Vibration 1 6 1 mdash i mdash i mdash i mdash mdash i mdash 1 mdash 1 mdash mdash mdash i mdash mdash 1 mdash 1 mdash 1 mdash i mdash 1 mdash 1 mdash 1 mdash 1 mdash r
0 1 2 3 4 5
Time (Orbits) Tether In-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-6 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal local vertical and orbit normal (b) vibration response
68
STSS configuration 3-Body a(0)=a2(0)=ot(0)=0 P1(0)=p2(0)=pt(0)=0 8X(0)=1304 x 103m 8y(0)=8z(0)=0 dx(0)=dy(0)=dz(0)=0 Deployment
l=02km to 20km in 35 Orbits
Tether Length Profile
E
Satellite 1 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 1 Roll Angle mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash [ mdash i mdash i mdash i mdash i mdash | mdash
1 2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 4 Time (Orbits)
Tether Roll Angle
CD
oo ax
i imdashimdashimdashImdashimdashimdashimdashr-
_ 1 _ L
1 2 3 4 Time (Orbits)
Satellite 2 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 2 Roll Angle
-50
(a) 1 2 3 4
Time (Orbits) 1 2 3 4
Time (Orbits)
Figure 4-7 Deployment dynamics of the three-body STSS configuration without offset (a) attitude response
69
STSS configuration 3-Body a1(0)=a2(0)=at(0)=0 p1(0)=p2(0)=(3t(0)=0 8X(0)=1304x 103m 8y(0)=52(0)=0 dx(0)=dy(0)=d2(0)=0 Deployment
l=02km to 20km in 35 Orbits
Tether Longitudinal Vibration
to
-02
(b)
Figure 4-7
2 3 Time (Orbits)
Tether In-Plane Transverse Vibration
0 1 1 bull 1 1
1 2 3 4 Time (Orbits)
Tether Out-of-Plane Transverse Vibration
5
1 1 1 1 1 i i | i i | ^
_ i i i i L -I
1 2 3 4 5 Time (Orbits)
Deployment dynamics of the three-body STSS configuration without offset (b) vibration response
70
tether elongates its longitudinal static equilibrium position also changes due to an
increase in the gravity-gradient tether tension It may be pointed out that as in the
case of attitude motion there are no out-of-plane tether vibrations induced
When the same deployment conditions are applied to the case of 1 m offset in
the x direction (Figure 4-8) coupling effects are re-introduced The platform pitch
exceeds -100deg during the peak deployment acceleration However as the tether pitch
motion subsides the platform pitch response returns to a smaller amplitude oscillation
about its nominal equilibrium On the other hand the platform roll motion grows
to over 20deg in amplitude in the terminal stages of deployment Longitudinal and
transverse vibrations of the tether remain virtually unchanged from the zero offset
case except for minute out-of-plane perturbations due to the platform librations in
roll
45 Tether Retrieval
The system response for the case of the tether retrieval from 20 km to 200 m
with an offset of 1 m in the x and z directions at the platform end is presented in
Figure 4-9 Here the Coriolis force induced during retrieval renders the tether unstable
(Figure 4-9a) Consequently the pitch motion of the platform is also destabilized
through coupling The z offset provides an additional moment arm that shifts the
roll equilibrium to 35deg however this degree of freedom is not destabilized From
Figure 4-9(b) it is apparent that the transverse mode of the tether vibration is also
disturbed producing a high frequency response in the final stages of retrieval This
disturbance is responsible for the slight increase in the roll motion for both the tether
and subsatellite
46 Five-Body Tethered System
A five-body chain link system is considered next To facilitate comparison with
71
STSS configuration 3-Body a(0)=a2(0)=at(0)=0 P1(0)=P2(0)=P(0)=0 5X(0)=1304 x 103m 8y(0)=8z(0)=0 dx(0)=1mdy(0)=d2(0)=0 Deployment
lt=02km to 20km in 35 Orbits
Tether Length Profile
E
Satellite 1 Pitch Angle
1 2 3 Time (Orbits)
Satellite 1 Roll Angle
bull bull bull i bull bull bull bull i
1 2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 Time (Orbits)
Tether Roll Angle
i i i i J I I i _
1 2 3 4 Time (Orbits)
Satellite 2 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 2 Roll Angle 3E-3h
D )
T J ^ 0 E 0 eg
CQ
-3E-3
(a) 1 2 3 4
Time (Orbits) 1 2 3 4
Time (Orbits)
Figure 4-8 Deployment dynamics of the three-body STSS configuration with offset along the local vertical (a) attitude response
7 2
STSS configuration 3-Body a1(0)=a2(0)=at(0)=0 p1(0)=P2(0)=(3t(0)=0 Sx(0)=1304 x 103m 5y(0)=5z(0)=0 dx(0)=1mdy(0)=dz(0)=0 Deployment
l=02km to 20km in 35 Orbits
Tether Longitudinal Vibration
Time (Orbits) Tether In-Plane Transverse Vibration
n 1 1 1 f
I I I _j I I I I I 1 1 1 1 1 2 3 4 5
Time (Orbits) Tether Out-of-Plane Transverse Vibration
J 1 I I 1 I I I I I I I I L _ _ I I d 1 2 3 4 5
Time (Orbits)
Deployment dynamics of the three-body STSS configuration with offset along the local vertical (b) vibration response
(b)
Figure 4-8
73
STSS configuration 3-Body a1(0)=a2(0)=at(0)=0 p(0)=P2(0)=(3t(0)=0 Sx(0)=14m8y(0)=82(0)=0 dy(0)=0dx(0)=d2(0)=1m Retrieval
l=20km to 02km in 35 Orbits
Satellite 1 Pitch Angle
(a) Time (Orbits)
Figure 4-9 Retrieval dynamics of the along the local vertical ar
Tether Length Profile
i i i i i i i i i i I 0 1 2 3
Time (Orbits) Satellite 1 Roll Angle
Time (Orbits) Tether Roll Angle
b_ i i I i I i i L _ J
0 1 2 3 Time (Orbits)
Satellite 2 Roll Angle _ mdash | r i i | 1 i i imdashmdashj i 1 r mdash i mdash
04- -
0 1 2 3 Time (Orbits)
three-body STSS configuration with offset d orbit normal (a) attitude response
74
STSS configuration 3-Body a1(0)=a2(0)=ocl(0)=0 (31(0)=f32(0)=pt(0)=0 8x(0)=14m8y(0)=6z(0)=0 dy(0)=0dx(0)=dz(0)=1m Retrieval
l=20km to 02km in 35 Orbits
Tether Longitudinal Vibration
E 10 to
Time (Orbits) Tether In-Plane Transverse Vibration
1 2 Time (Orbits)
Tether Out-of-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-9 Retrieval dynamics of the three-body STSS configuration with offset along the local vertical and orbit normal (b) vibration response
75
the three-body system dynamics studied earlier the chain is extended through the
addition of two bodies a tether and a subsatellite with the same physical properties
as before (Section 42) Thus the five-body system consists of a platform (satellite 1)
tether 1 subsatellite 1 (satellite 2) tether 2 and subsatellite 2 (satellite 3) as shown
in Figure 4-10 The offset of tether 1 at the platform-end is denoted by d2 while that
of tether 2 to subsatellite 1 as 04 In this numerical example fcm^ = fcm^ = 0 ie
tethers 1 and 2 are attached to the centre of mass of subsatellites 1 and 2 respectively
The system response for the case where the tether attachment points coincide
wi th the centres of mass of the rigid bodies ie zero offset (d2 = d mdash 0) is presented
in Figure 4-11 Here ogtt and at2 represent the pitch motion of tethers 1 and 2
respectively whereas 03 is the pitch motions of satellite 3 A similar convention
is adopted for the rol l angle 0 A s before the zero offset eliminates the coupling
between the tethers and satellites such that they are now free to librate about their
static equil ibrium positions However their is s t i l l mutual coupling between the two
tethers This coupling is present regardless of the offset position since each tether is
capable of transferring a force to the other The coupling is clearly evident in the pitch
response of the two tethers (Figure 4 - l l a ) O n the other hand the roll motion appears
uncoupled as the in i t ia l conditions in rol l for the two tethers are identical Thus the
motion is in phase and there is no transfer of energy through coupling A s expected
during the elastic response the tethers vibrate about their static equil ibrium positions
and mutually interact (Figure 4 - l l b ) However due to the variation of tension along
the tether (x direction) they do not have the same longitudinal equilibrium Note
relatively large amplitude transverse vibrations are present particularly for tether 1
suggesting strong coupling effects with the longitudinal oscillations
76
Figure 4-10 Schematic diagram of the five-body system used in the numerical example
77
STSS configuration 5-Body a1(0)=(x2(0)=at1(0)=2deg a3(0)=a12(0)=25deg p1(0)=r32(0)=p3(0)=pt1(0)=pt2(0)=r 82(0)=84(0)=505)05m d2(0)=d4(0)=000 Stationkeeping lt1=lu=10km Pitch Roll
Satellite 1 Pitch Angle Satellite 1 Roll Angle
o 1 2 Time (Orbits)
Tether 1 Pitch and Roll Angle
0 1 2 Time (Orbits)
Tether 2 Pitch and Roll Angle
Time (Orbits) Satellite 2 Pitch and Roll Angle
Time (Orbits) Satellite 3 Pitch and Roll Angle
Time (Orbits) Time (Orbits)
Figure 4-11 Stationkeeping dynamics of the five-body STSS configuration withshyout offset (a) attitude response
78
STSS configuration 5-Body a1(0)=a2(0)=oct1(0)=2deg a3(0)=at2(0)=25o
p1(0)=p2(0)=(33(0)=pt1(0)=pt2(0)=10
62(0)=54(0)=50505m d2(0)=d4(0)=000 Stationkeeping lt1=1 =1 Okm
Tether 1 Longitudinal Vibration Tether 2 Longitudinal Vibration i i 1 i 1 r 1 1 1 1
(b) Time (Orbits) Time (Orbits)
Figure 4-11 Stationkeeping dynamics of the five-body STSS configuration withshyout offset (b) vibration response
79
The system dynamics during deployment of both the tethers in the double-
pendulum configuration is presented in Figure 4-12 Deployment of each tether takes
place from 200 m to 20 km in 35 orbits The pitch motion of the system is similar
to that observed during the three-body case The Coriolis effect causes the tethers
to pitch through a large angle ( laquo 50deg ) which in turn disturbs the platform and
subsatellite due to the presence of offset along the local vertical On the other hand
the roll response for both the tethers is damped to zero as the tether deploys in
confirmation with the principle of conservation of angular momentum whereas the
platform response in roll increases to around 20deg Subsatellite 2 remains virtually
unaffected by the other links since the tether is attached to its centre of mass thus
eliminating coupling Finally the flexibility response of the two tethers presented in
Figure 4-12(b) shows similarity with the three-body case
47 BICEPS Configuration
The mission profile of the Bl-static Canadian Experiment on Plasmas in Space
(BICEPS) is simulated next It is presently under consideration by the Canadian
Space Agency To be launched by the Black Brant 2000500 rocket it would inshy
volve interesting maneuvers of the payload before it acquires the final operational
configuration As shown in Figure 1-3 at launch the payload (two satellites) with
an undeployed tether is spinning about the orbit normal (phase 1) The internal
dampers next change the motion to a flat spin (phase 2) which in turn provides
momentum for deployment (phase 3) When fully deployed the one kilometre long
tether will connect two identical satellites carrying instrumentation video cameras
and transmitters as payload
The system parameters for the BICEPS configuration are summarized below [59 0 0 1
bull I = I2 = 0 366 0 kg bull m 2 (satellite inertias) [ 0 0 392_
bull mi = 7772 = 200 kg (mass of the satellites)
80
STSS configuration 5-Body a(0)=cx2(0)=al1(0)=2deg a3(0)=c (0)=25o
P1(0)=P2(0)=p3(0)=f3t1(0)=Pt2(0)=r 52(0)=84(0)=1304x 10300m d2(0)=d4(0)=1G0m Deployment lt1=le=02km to 20km in 35 Orbits
Pitch Roll Satellite 1 Pitch Angle
Time (Orbits) Satellite 2 Pitch and Roll Angle T 1mdash1 1 1 I I I I | I I I 1 I | I I lt~r
Tether Length Profile i 1 1 1 1 i 1 1 1 1 i 1 1 1 1 i
i i i i i 0 1 2 3 4 5
Time (Orbits) Satellite 1 Roll Angle
0 1 2 3 4 5 Time (Orbits)
Tether 1 amp 2 Roll Angle L i | I I 1 I 1 1 -I
-Ih I- bull I I i i i i I i i i i 1
0 1 2 3 4 5 Time (Orbits)
Satellite 3 Pitch and Roll Angle i- i 1 1 i 1 1 1 1 i 1 1 1 1 i 1 1 1 1 i
L i i i i i i 1 i i mdash i mdash i I i imdashimdashLJ
0 1 2 3 4 5 Time (Orbits)
Figure 4-12 Deployment dynamics of the five-body S T S S configuration with offshyset along the local vertical (a) attitude response
81
STSS configuration 5-Body a1(0)=a2(0)=oct1(0)=2deg a3(0)=at2(0)=25deg P1(0)=P2(0)=P3(0)=Pt1(0)=pt2(0)=1o
82(0)=54(0)=1304 x 10-300m d2(0)=d4(0)=100m Deployment lt1=l2=02km to 20km in 35 Orbits
Tether 1 Longitudinal Vibration Tether 2 Longitudinal Vibration
Time (Orbits) Time (Orbits) Tether 1 Z Tranverse Vibration Tether 2 Z Tranverse Vibration
Figure 4-12 Deployment dynamics of the five-body STSS configuration with offshyset along the local vertical (b) vibration response
82
bull EtAt = 61645 N (tether stiffness)
bull pt = 30 k g k m (tether density)
bull It = 1 km (tether length)
bull rid = 1-0 (tether structural damping coefficient)
The system is in a circular orbit at a 289 km altitude Offset of the tether
attachment point to the satellites at both ends is taken to be 078 m in the x
direction The response of the system in the stationkeeping phase for a prescribed set
of in i t ia l conditions is given by Figure 4-13 As in the case of the STSS configuration
there is strong coupling between the tether and the satellites rigid body dynamics
causing the latter to follow the attitude of the tether However because of the smaller
inertias of the payloads there are noticeable high frequency modulations arising from
the tether flexibility Response of the tether in the elastic degrees of freedom is
presented in Figure 4-13(b) Note the longitudinal vibrations decay quite rapidly
(lt 05 orbits) due to the structural damping however it has vir tual ly no effect on
the transverse oscillations
The unique mission requirement of B I C E P S is the proposed use of its ini t ia l
angular momentum in the cartwheeling mode to aid in the deployment of the tethered
system The maneuver is considered next The response of the system during this
maneuver wi th an ini t ia l cartwheeling rate of 5 deg s is illustrated in Figure 4-14(a) The
in i t ia l tether length is taken to be 10 m and is deployed to 1 km There is an in i t ia l
increase in the pitch motion of the tether however due to the conservation of angular
momentum the cartwheeling rate decreases proportionally to the square of the change
in tether deployment rate The result is a rapid drop in the cartwheeling rate unti l
the system stops rotating entirely and simply oscillates about its new equilibrium
point Consequently through coupling the end-bodies follow the same trend The
roll response also subsides once the cartwheeling motion ceases However
83
BICEPS configuration 3-Body a1(0)=o2(0)=at(0)=2deg P1(0)=P2(0)=pt(0)=1deg 8X(0)=001 m 6y(0)=52(0)=01 m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping l=1 km
Satellite 1 Pitch Angle Satellite 1 Roll Angle - i 1 1 1 1 i - i q r 1 1 -gt
I- I i t i i i i I H 0 1 2 0 1 2
Time (Orbits) Time (Orbits) Tether Pitch Angle Tether Roll Angle
F mdash 1 mdash 1 mdash mdash 1 mdash i mdash 1 mdash 1 mdash mdash 1 mdash = i F mdash 1 mdash mdash mdash 1 mdash i mdash 1 mdash 1 mdash 1 mdash 1 mdash q
(a) Time (Orbits) Time (Orbits)
Figure 4-13 Stationkeeping dynamics of the three-body BICEPS configuration with offset along the local vertical (a) attitude response
84
BICEPS configuration 3-Body a1(0)=o2(0)=at(0)=2deg Pl(0)=p2(0)=(3t(0)=1deg 8X(0)=001 m 8y(0)=82(0)=01 m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping lt=1km
Tether Longitudinal Vibration 1E-2[
laquo = 5 E - 3 |
000 002
Time (Orbits) Tether In-Plane Transverse Vibration
Time (Orbits) Tether Out-of-Plane Transverse Vibration
Time (Orbits)
Figure 4-13 Stationkeeping dynamics of the three-body BICEPS configuration with offset along the local vertical (b) vibration response
85
BICEPS configuration 3-Body a1(0)=o2(0)=at(0)=2deg a1(0)=a2(0)=at(0)=57s Pl(0)=P2(0)=(3t(0)=1deg 8X(0)=115 x 103m 8y(0)=82(0)=0 dx(0)=078m dy(0)=d2(0)=0 Cartwheeling Deployment
lt=10m to 1km in 1 Orbit
Tether Length Profile
E
2000
cu T3
2000
co CD
2000
C D CD
T J
Satellite 1 Pitch Angle
1 2 Time (Orbits)
Satellite 1 Roll Angle
cn CD
33 cdl
Time (Orbits) Tether Pitch Angle
Time (Orbits) Tether Roll Angle
Time (Orbits) Satellite 2 Pitch Angle
1 2 Time (Orbits)
Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-14 Cartwheeling dynamics with deployment for the three-body BICEPS with offset along the local vertical (a) attitude response
86
BICEPS configuration 3-Body a1(0)=a2(0)=at(0)=2deg a (0)=a2(0)=a(0)=57s P1(0)=P2(0)=f3(0)=1deg 8X(0)=115x103m 8y(0)=8z(0)=0 dx(0)=078m dy(0)=d2(0)=0 Cartwheeling Deployment
l=1 Om to 1 km in 1 Orbit
5E-3h
co
0E0
025
000 E to
-025
002
-002
(b)
Figure 4-14
Tether Longitudinal Vibration
Time (Orbits) Tether In-Plane Transverse Vibration
Time (Orbits) Tether Out-of-Plane Transverse Vibration
Time (Orbits)
Cartwheeling dynamics with deployment for the three-body BICEPS with offset along the local vertical (b) vibration response
87
as the deployment maneuver is completed the satellites continue to exhibit a small
amplitude low frequency roll response with small period perturbations induced by
the tether elastic oscillations superposed on it (Figure 4-14b)
48 OEDIPUS Spinning Configuration
The mission entitled Observation of Electrified Distributions of Ionospheric
Plasmas - A Unique Strategy ( O E D I P U S ) also consists of a 1 k m tethered system with
two similar end-masses It was first flown in a suborbital flight in 1989 ( O E D I P U S A )
followed by a recent study in November 1995 ( O E D I P U S - C ) in which the University
of Br i t i sh Columbia was one of the participants Dynamically O E D I P U S represents a
unique system never encountered before Launched by a Black Brant rocket developed
at Br is to l Aerospace L t d the system ie the end-bodies together wi th the tether
spins about the longitudinal axis of the tether to achieve stabilized alignment wi th
Earth s magnetic field
The response of the system undergoing a spin rate of j = l deg s is presented
in Figure 4-15 using the same parameters as those outlined for the B I C E P S conshy
figuration in Section 47 Again there is strong coupling between the three bodies
(two satellites and tether) due to the nonzero offset The spin motion introduces an
additional frequency component in the tethers elastic response which is transferred
to the l ibrational motion of the satellites A s the spin rate increases to 1 0 deg s (Figure
4-16) the amplitude and frequency of the perturbations also increase however the
general character of the response remains essentially the same
88
OEDIPUS configuration 3-Body a1(0)=a2(0)=at(0)=2deg i(0)^[2(0H(0)=Ys p1(0)=p2(0)=p(0)=r 8X(0)=12 x 103m 8y(0)=82(0)=01m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping l=1 km
Satellite 1 Pitch Angle Satellite 1 Roll Angle mdash i 1 1 1 a r 1 1 1 r
(a) Time (Orbits) Time (Orbits)
Figure 4-15 Spin dynamics (7 = ldegs) of the three-body OEDIPUS configuration with offset along the local vertical (a) attitude response
89
OEDIPUS configuration 3-Body a1(0)=a2(0)=cct(0)=2deg Y1(0)=Y2(0)=Yt(0)=l7s (31(0)=p2(0)=pt(0)=1deg 8X(0)=12 x 103m 8y(0)=5z(0)=01 m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping lt=1 km
1E-2
to
OEO
E 00
to
Tether Longitudinal Vibration
005
Time (Orbits) Tether In-Plane Transverse Vibration
Time (Orbits) Tether Out-of-Plane Transverse Vibration
Time (Orbits)
Figure 4-15 Spin dynamics (7 = l deg s ) of the three-body O E D I P U S configuration with offset along the local vertical (b) vibration response
90
OEDIPUS configuration 3-Body a1(0)=a2(0)=a(0)=2deg Yi(0)=Y2(0)=rt(0)=107s 81(0)=p2(0)=pt(0)=1deg 8X(0)=12 x 103m 8y(0)=82(0)=01m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping lt=1km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-16 Spin dynamics (7= 10degs) of the three-body OEDIPUS configura tion with offset along the local vertical (a) attitude response
91
OEDIPUS configuration 3-Body a1(0)=a2(0)=ot(0)=2deg Yi(0)H(0)4(0)laquo107s P1(0)=P2(0)=pt(0)=1deg 8X(0)=12 x 103m 8y(0)=82(0)=01m dx(0)=078m dy(0)=dz(0)=0 Stationkeeping lt=1 km
Tether Longitudinal Vibration
Figure 4-16 Spin dynamics (7 = 10degs) of the three-body OEDIPUS configurashytion with offset along the local vertical (b) vibration response
92
5 ATTITUDE AND VIBRATION CONTROL
51 Att itude Control
511 Preliminary remarks
The instability in the pitch and roll motions during the retrieval of the tether
the large amplitude librations during its deployment and the marginal stability during
the stationkeeping phase suggest that some form of active control is necessary to
satisfy the mission requirements This section focuses on the design of an attitude
controller with the objective to regulate the librational dynamics of the system
As discussed in Chapter 1 a number of methods are available to accomplish
this objective This includes a wide variety of linear and nonlinear control strategies
which can be applied in conjuction with tension thrusters offset momentum-wheels
etc and their hybrid combinations One may consider a finite number of Linear T ime
Invariant (LTI) controllers scheduled discretely over different system configurations
ie gain scheduling[43] This would be relevant during deployment and retrieval of
the tether where the configuration is changing with time A n alternative may be
to use a Linear T ime Varying ( L T V ) model in conjuction with an adaptive control
scheme where on-line parametric identification may be used to advantage[56] The
options are vir tually limitless
O f course the choice of control algorithms is governed by several important
factors effective for time-varying configurations computationally efficient for realshy
time implementation and simple in character Here the thrustermomentum-wheel
system in conjuction with the Feedback Linearization Technique ( F L T ) is chosen as
it accounts for the complete nonlinear dynamics of the system and promises to have
good overall performance over a wide range of tether lengths It is well suited for
93
highly time-varying systems whose dynamics can be modelled accurately as in the
present case
The proposed control method utilizes the thrusters located on each rigid satelshy
lite excluding the first one (platform) to regulate the pitch and rol l motion of the
tether to which it is attached ie the thrusters located on satellite 2 (link 3) regshy
ulates the attitude motion of tether 1 (link 2) O n the other hand the l ibrational
motion of the rigid bodies (platform and subsatellite) is controlled using a set of three
momentum wheels placed mutually perpendicular to each other
The F L T method is based on the transformation of the nonlinear time-varying
governing equations into a L T I system using a nonlinear time-varying feedback Deshy
tailed mathematical background to the method and associated design procedure are
discussed by several authors[435758] The resulting L T I system can be regulated
using any of the numerous linear control algorithms available in the literature In the
present study a simple P D controller is adopted arid is found to have good perforshy
mance
The choice of an F L T control scheme satisfies one of the criteria mentioned
earlier namely valid over a wide range of tether lengths The question of compushy
tational efficiency of the method wi l l have to be addressed In order to implement
this controller in real-time the computation of the systems inverse dynamics must
be executed quickly Hence a simpler model that performs well is desirable Here
the model based on the rigid system with nonlinear equations of motion is chosen
Of course its validity is assessed using the original nonlinear system that accounts
for the tether flexibility
512 Controller design using Feedback Linearization Technique
The control model used is based on the rigid system with the governing equa-
94
tions of motion given by
Mrqr + fr = Quu (51)
where the left hand side represents inertia and other forces while the right hand side
is the generalized external force due to thrusters and momentum-wheels and is given
by Eq(2109)
Lett ing
~fr = fr~ QuU (52)
and substituting in Eq(296)
ir = s(frMrdc) (53)
Expanding Eq(53) it can be shown that
qr = S(fr Mrdc) - S(QU Mr6)u (5-4)
mdash = Fr - Qru
where S(QU Mr0) is the column matrix given by
S(QuMr6) = [s(Qu(l)Mr0) 5 (Q u ( 2 ) M r | 0 ) S(QU nu) M r | 0 ) ] (55)
and Qu-i) is the i ^ column of Qu Extract ing only the controlled equations from
Eq(54) ie the attitude equations one has
Ire = Frc ~~ Qrcu
(56)
mdash vrci
where vrc is the new control input required to regulate the decoupled linear system
A t this point a simple P D controller can be applied ie
Vrc = qrcd +KvQrcd-Qrc) + Kp(Qrcd~ Qrc)- (57)
Eq(57) represents the secondary controller with Eq(54) as the primary controller
Kp and Kv are the proportional and derivative gain matrices with qrcdi Qrcd
an(^ qrcd
95
as the desired acceleration velocity and position vector for the attitude angles of each
controlled body respectively Solving for u from Eq(56) leads to
u Qrc Frc ~ vrcj bull (58)
5 13 Simulation results
The F L T controller is implemented on the three-body STSS tethered system
The choice of proportional and derivative matrix gains is based on a desired settling
time of ts mdash 05 orbit and a damping factor of = 07 for both the pitch and roll
actuators in each body Given O and ts it can be shown that
-In ) 0 5 V T ^ r (59)
and hence kn bull mdash10
V (510)
where kp and kVj are the diagonal elements of the gain matrices Kp and Kv in
Eq(57) respectively[59]
Note as each body-fixed frame is referred directly to the inertial frame FQ
the nominal pitch equilibrium angle is zero only when measured from the L V L H frame
(OJJ mdash 0) However it is 6 when referred to FQ Hence the desired position vector
qrC(l is set equal to 0(t) ie the time-varying orbital angle for the pitch motion
and zero for the roll and yaw rotations such that
qrcd 37Vxl (511)
96
The desired velocity and acceleration are then given as
e U 3 N x l (512)
E K 3 7 V x l (513)
When the system is in a circular orbit qrc^ = 0
Figure 5-1 presents the controlled response of the STSS system defined in
Chapter 4 in the stationkeeping phase with offset d2 = 111T m and If = 20
km A s mentioned earlier the F L T controller is based on the nonlinear rigid model
Note the pitch and roll angles of the rigid bodies as well as of the tether are now
attenuated in less than 1 orbit (Figure 5-la) Consequently the longitudinal vibration
response of the tether is free of coupling arising from the librational modes of the
tether This leaves only the vibration modes which are eventually damped through
structural damping The pitch and roll dynamics of the platform require relatively
higher control moments Ma and Mpi respectively (Figure 5-lb) since they have to
overcome the extra moments generated by the offset of the tether attachment point
Furthermore the platform demands an additional moment to maintain its orientation
at zero pitch and roll angle since it is not the nominal equilibrium position The nonshy
zero static equilibrium of the platform arises due to its products of inertia On the
other hand the tether pitch and roll dynamics require only a small ini t ia l control
force of about plusmn 1 N which eventually diminishes to zero once the system is
97
and
o 0
Qrcj mdash
o o
0 0
Oiit) 0 0
1
N
N
1deg I o
STSS configuration 3-Body a1(0)=a2(0)=at(0)=2lt
P1(0)=p2(0)=pt(0)=1 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=dz(0)=1m Controlled Stationkeeping lt=20km
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
05 10 15 Time (Orbits)
Sat 2 Pitch and Roll Angles
20
CO CD
T J
cdT
cT
- mdash mdash 1 1 1 1 1 I - 1
Pitch -
A Roll -
A -
1 bull bull
-
1 Time (Orbits)
Tether Y Tranverse Vibration
1 2 Time (Orbits)
Tether Z Tranverse Vibration
gt to N
to
Time (Orbits) Time (Orbits) Longitudinal Vibration
145h
(a) Time (Orbits)
Figure 5-1 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local horizontal local vertical and orbit normal (a) attitude and vibration response
98
1
STSS configuration 3-Body a1(0)=ct2(0)=cct(0)=20
P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=d2(0)=1m Controlled Stationkeeping lt=20km
Sat 1 Pitch Control Moment Sat 1 Roll Control Moment
Figure 5-1 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local horizontal local vertical and orbit normal (b) control actuator time histories
99
stabilized Similarly for the end-body a very small control torque is required to
attenuate the pitch and roll response
If the complete nonlinear flexible dynamics model is used in the feedback loop
the response performance was found to be virtually identical with minor differences
as shown in Figure 5-2 For example now the pitch and roll response of the platform
is exactly damped to zero with no steady-state error as against the minor almost
imperceptible deviation for the case of the controller basedon the rigid model (Figure
5-la) In addition the rigid controller introduces additional damping in the transverse
mode of vibration where none is present when the full flexible controller is used
This is due to a high frequency component in the at motion that slowly decays the
transverse motion through coupling
As expected the full nonlinear flexible controller which now accounts the
elastic degrees of freedom in the model introduces larger fluctuations in the control
requirement for each actuator except at the subsatellite which is not coupled to the
tether since f c m 3 = 0 The high frequency variations in the pitch and roll control
moments at the platform are due to longitudinal oscillations of the tether and the
associated changes in the tension Despite neglecting the flexible terms the overall
controlled performance of the system remains quite good Hence the feedback of the
flexible motion is not considered in subsequent analysis
When the tether offset at the platform is restricted to only 1 m along the x
direction a similar response is obtained for the system However in this case the
steady-state error in the platforms rigid body motion is much smaller (Figure 5-3a)
In addition from Figure 5-3(b) the platforms control requirement is significantly
reduced
Figure 5-4 presents the controlled response of the STSS deploying a tether
100
STSS configuration 3-Body a1(0)=a2(0)=oct(0)=2deg p1(0)=P2(0)=p(0)=1deg 8x(0)=14m5y(0)=82(0)=1m dx(0)=dy(0)=dz(0)=1m Controlled Stationkeeping l=20km
Tether Pitch and Roll Angles
c n
T J
cdT
Sat 1 Pitch and Roll Angles 1 r
00 05 10 15 Time (Orbits)
Sat 2 Pitch and Roll Angles
c n CD
cdT a
o T J
CQ
- - I 1 1 1 I 1 1 1 1
P i t c h
R o l l
bull
1 -
0 1 2 Time (Orbits)
Tether Y Tranverse Vibration
0 1 2 Time (Orbits)
Tether Z Tranverse Vibration
E gt
10 N
CO
Time (Orbits) Time (Orbits) Longitudinal Vibration
145h
(a) Time (Orbits)
Figure 5-2 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear flexible FLT controller with offset along the local horizontal local vertical and orbit normal (a) attitude and vibration response
101
STSS configuration 3-Body a1(0)=a2(0)=cxt(0)=2deg 31(0)=p2(0)=p(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=d2(0)=1m Controlled Stationkeeping lt=20km
r i i i t i mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash mdash J 0 1 2 0 1 2
Time (Orbits) Time (Orbits) Tether Pitch Control Thrust Tether Roll Control Thruster
(b) Time (Orbits) Time (Orbits)
Figure 5-2 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear flexible FLT controller with offset along the local horizontal local vertical and orbit normal (b) control actuator time histories
102
STSS configuration 3-Body a1(0)=o2(0)=a(0)=2deg P1(0)=p2(0)=(3t(0)=1deg 8x(0)=14m5y(0)=6z(0)=1m dx(0)=1mdy(0)=dz(0)=0 Controlled Stationkeeping lt=20km
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
05 10 15 Time (Orbits)
Sat 2 Pitch and Roll Angles
CD
eg CQ
0 1 2 Time (Orbits)
Tether Y Tranverse Vibration
1 2 Time (Orbits)
Tether Z Tranverse Vibration
gt CO
Time (Orbits) Time (Orbits) Longitudinal Vibration
145h
(a) Time (Orbits)
Figure 5-3 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local vertical (a) attitude and vibration response
103
STSS configuration 3-Body a1(0)=o2(0)=al(0)=2o
P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=1mdy(0)=dz(0)=0 Controlled Stationkeeping lt=20km
Sat 1 Pitch Control Moment 11 1 1 r
Sat 1 Roll Control Moment
0 1 2 Time (Orbits)
Tether Pitch Control Thrust
0 1 2 Time (Orbits)
Sat 2 Pitch Control Moment
0 1 2 Time (Orbits)
Tether Roll Control Thruster
-07 o 1
Time (Orbits Sat 2 Roll Control Moment
1E-5
0E0F
Time (Orbits)
Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local vertical (b) control actuator time histories
104
STSS configuration 3-Body a1(0)=a2(0)=a(0)=2deg Pl(0)=p2(0)=pt(0)=1deg 6X(0)=1304 x 103m 8y(0)=82(0)=0 dx(0)=1mdy(0)=d2(0)=0 Controlled Deployment
l=02km to 20km in 35 Orbits
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
o 1 2 Time (Orbits)
Tether Y Tranverse Vibration
0 1 2 Time (Orbits)
Sat 2 Pitch and Roll Angles
CD CD
CM CO
Time (Orbits) Tether Z Tranverse Vibration
to
150
100
1 2 3 0 1 2
Time (Orbits) Time (Orbits) Longitudinal Vibration
to
(a)
50
2 3 Time (Orbits)
Figure 5 - 4 Deployment dynamics of the three-body STSS using the nonlinear rigid FLT controller with offset along the local vertical (a) attitude and vibration response
105
STSS configuration 3-Body a1(0)=oc2(0)=al(0)=20
P1(0)=p2(0)=pt(0)=1deg 8X(0)=1304 x 103m 5y(0)=52(0)=0 dx(0)=1mdy(0)=dz(0)=0 Controlled Deployment
l=02km to 20km in 35 Orbits
Tether Length Profile i 1
2 0 h
E 2L
Sat 1 Pitch Control Moment t mdash r mdash i mdash i mdash | mdash i mdash i mdash i mdash i mdash | mdash r mdash i mdash i mdash i mdash | mdash r mdash
0 5
1 2 3
Time (Orbits) Sat 1 Roll Control Moment
i i t | t mdash t mdash i mdash [ i i mdash r -
1 2
Time (Orbits) Tether Pitch Control Thrust
1 2 3 Time (Orbits)
Tether Roll Control Thruster
2 E - 5
1 2 3
Time (Orbits) Sat 2 Pitch Control Moment
1 2
Time (Orbits) Sat 2 Roll Control Moment
1 E - 5
E
O E O
Time (Orbits)
Deployment dynamics of the three-body STSS using the nonlinear rigid FLT controller with offset along the local vertical (b) control actuator time histories
106
from 02 km to 20 km in 35 orbits with a i m offset along the tether length A s before
the systems attitude motion is well regulated by the controller However the control
cost increases significantly for the pitch motion of the platform and tether due to the
Coriolis force induced by the deployment maneuver (Figure 5-4b) Initially there is a
sinusoidal increase in the pitch control requirement for both the tether and platform
as the tether accelerates to its constant velocity VQ Then the control requirement
remains constant for the platform at about 60 N m as opposed to the tether where the
thrust demand increases linearly Finally when the tether decelerates the actuators
control input reduces sinusoidally back to zero The rest of the control inputs remain
essentially the same as those in the stationkeeping case In fact the tether pitch
control moment is significantly less since the tether is short during the ini t ia l stages
of control However the inplane thruster requirement Tat acts as a disturbance and
causes 6y to nearly double from the uncontrolled value (Figure 4-8a) O n the other
hand the tether has no out-of-plane deflection
Final ly the effectiveness of the F L T controller during the crit ical maneuver of
retrieval from 20 km to 02 k m in 35 orbits is assessed in Figure 5-5 A s in the earlier
cases the system response in pitch and roll is acceptable however the controller is
unable to suppress the high frequency elastic oscillations induced in the tether by the
retrieval O f course this is expected as there is no active control of the elastic degrees
of freedom However the control of the tether vibrations is discussed in detail in
Section 52 The pitch control requirement follows a similar trend in magnitude as
in the case of deployment with only minor differences due to the addition of offset in
the out-of-plane direction ( d 2 = 1 0 1 ^ m) This is also responsible for the higher
roll moment requirement for the platform control (Figure 5-5b)
107
STSS configuration 3-Body a1(0)=X2(0)=a(0)=2deg P1(0)=p2(0)=pt(0)=1deg 5x(0)=14m8y(0)=52(0)=0 dx(0)=1mdy(0)=dz(0)=0 Controlled Retrieval
l=20km to 02km in 35 Orbits
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
0 1 2 Time (Orbits)
Sat 2 Pitch and Roll Angles
CO CD
cdT
CO CD
T J CM
CO
1 2 Time (Orbits)
Tether Y Tranverse Vibration
Time (Orbits) Tether Z Tranverse Vibration
150
100
1 2 3 0 1 2 Time (Orbits) Time (Orbits)
Longitudinal Vibration
CO
(a)
x 50
2 3 Time (Orbits)
Figure 5 -5 Retrieval dynamics of the three-body STSS using the non-linear rigid FLT controller with offset along the local vertical and orbit normal (a) attitude and vibration response
108
STSS configuration 3-Body a1(0)=ct2(0)=at(0)=2deg p1(0)=p2(0)=(3(0)=1deg 5x(0)=14m5y(0)=82(0)=0 dx(0)=1mdy(0)=dz(0)=0 Controlled Retrieval
l=20km to 02km in 35 Orbits
Tether Length Profile
E
1 2 Time (Orbits)
Tether Pitch Control Thrust
2E-5
1 2 Time (Orbits)
Sat 2 Pitch Control Moment
OEOft
1 2 3 Time (Orbits)
Sat 1 Roll Control Moment
1 2 3 Time (Orbits)
Tether Roll Control Thruster
1 2 Time (Orbits)
Sat 2 Roll Control Moment
1E-5 E z
2
0E0
(b) Time (Orbits) Time (Orbits)
Figure 5-5 Retrieval dynamics of the three-body STSS using the non-linear rigid FLT controller with offset along the local vertical and orbit normal (b) control actuator time histories
109
52 Control of Tethers Elastic Vibrations
521 Prel iminary remarks
The requirement of a precisely controlled micro-gravity environment as well
as the accurate release of satellites into their final orbit demands additional control
of the tethers vibratory motion in addition to its attitude regulation To that end
an active vibration suppression strategy is designed and implemented in this section
The strategy adopted is based on offset control ie time dependent variation
of the tethers attachment point at the platform (satellite 1) A l l the three degrees
of freedom of the offset motion are used to control both the longitudinal as well as
the inplane and out-of-plane transverse modes of vibration In practice the offset
controller can be implemented through the motion of a dedicated manipulator or a
robotic arm supporting the tether which in turn is supported by the platform The
focus here is on the control of elastic deformations during stationkeeping (ie fully
deployed tether fixed length) as it represents the phase when the mission objectives
are carried out
This section begins with the linearization of the systems equations of motion
for the reduced three body stationkeeping case This is followed by the design of
the optimal control algorithm Linear Quadratic Gaussian-Loop Transfer Recovery
( L Q G L T R ) based on the reduced model and its implementation on the full nonlinear
model Final ly the systems response in the presence of offset control is presented
which tends to substantiate its effectiveness
522 System linearization and state-space realization
The design of the offset controller begins with the linearization of the equations
of motion about the systems equilibrium position Linearization of extremely lengthy
(even in matrix form) highly nonlinear nonautonomous and coupled equations of
110
motion presents a challenging problem This is further complicated by the fact the
pitch angle ot is not referred to the L V L H frame Hence the equilibrium pitch angle
is not a constant but is equal to the instantaneous orbital angle 9(t) Two methods
are available to resolve this problem
One may use the non-stationary equations of motion in their present form and
derive a controller based on the Linear T ime Varying ( L T V ) system Alternatively
one can design a Linear T ime Invariant (LTI) controller based on a new set of reduced
governing equations representing the motion of the three body tethered system A l shy
though several studies pertaining to L T V systems have been reported[5660] the latter
approach is chosen because of its simplicity Furthermore the L T I controller design
is carried out completely off-line and thus the procedure is computationally more
efficient
The reduced model is derived using the Lagrangian approach with the genershy
alized coordinates given by
where lty and are the pitch and roll angles of link i relative to the L V L H frame
respectively 5X 8y and 5Z are the generalized coordinates associated with the flexible
tether deformations in the longitudinal inplane and out-of-plane transverse direcshy
tions respectively Only the first mode of vibration is considered in the analysis
The nonlinear nonautonomous and coupled equations of motion for the tethshy
ered system can now be written as
Qred = [ltXlPlltX2P2fixSy5za302gt] (514)
M redQred + fr mdash 0gt (515)
where Mred and frec[ are the reduced mass matrix and forcing term of the system
111
respectively They are functions of qred and qred in addition to the time varying offset
position d2 and its time derivatives d2 and d2 A detailed derivation of Eq(515)
is given in Appendix II A n additional consequence of referring the pitch motion to
a local frame is that the new reduced equations are now independent of the orbital
angle 9 and under the further assumption of the system negotiating a circular orbit
91 remains constant
The nonlinear reduced equations can now be linearized about their static equishy
l ibr ium position For all the generalized coordinates the equilibrium position is zero
wi th the exception of 5X which has a non-zero equilibrium 5XQ The offset position
d2) is linearized about its ini t ial position d2^ Linearizing Eq(515) and recasting
into matr ix form gives
Msqred + CsQred + KsQred + Md^2 + Cdd2 + Kdd2 + fs = 0 (516)
where Ms Cs Ks Md Cd Kd and s are constant matrices Mul t ip ly ing throughout
by Ms
1 gives
qr = -Ms 1Csqr - Ms
lKsqr
- Ms~lCdd22 - Ms-Kdd2 - Ms~Mdd2 - Mg~1 fs r-1 (517)
Defining ud = d2 and v = qTd^T Eq(517) can be rewritten as
Pu t t ing
v =
+
-M~lCs -M~Cd -1
0
-Mg~lMd
Id
0 v + -M~lKs -M~lKd
0 0
-M-lfs
0
MCv + MKv + Mlud +Fs
- 4 ) - ( gt the L T I equations of motion can be recast into the state-space form as
(518)
(519)
x =Ax + Bud + Fd (520)
112
where
A = MC MK Id12 o
24x24
and
Let
B = ^ U 5 R 2 4 X 3 -
Fd = I FJ ) e f t 2 4 1
mdash _ i _ -5
(521)
(522)
(523)
(524)
where x is the perturbation vector from the constant equilibrium state vector xeq
Substituting Eq(524) into Eq(520) gives the linear perturbation state equation
modelling the reduced tethered system as
bullJ x mdash x mdash Ax + Bud + (Fd + Axeq) (525)
It can be shown that for ideal control[60] Fd + Axeq = 0 thus giving the familiar
state-space equation
S=Ax + Bud (526)
The selection of the output vector completes the state space realization of
the system The output vector consists of the longitudinal deformation from the
equil ibrium position SXQ of the tether at xt = h the slope of the tether due to the
transverse deformation at xt = 0 and the offset position d2 from its in i t ia l position
d2Q Thus the output vector y is given by
K(o)sx
V
dx ~ dXQ dy ~ dyQ
d z - d z0 J
Jy
dx da XQ vo
d z - dZQ ) dy dyQ
(527)
113
where lt ( 0 ) = ddxi[$vxi)]Xi=o and lt ( 0 ) = ddxi[4gtw(xi)]Xi=0
523 Linear Quadratic Gaussian control with Loop Transfer Recovery
W i t h the linear state space model defined (Eqs526527) the design of the
controller can commence The algorithm chosen is the Linear Quadratic Gaussian
( L Q G ) estimator based optimal controller[6061] The L Q G is a widely used optimal
controller with its theoretical background well developped by many authors over the
last 25 years It involves of the design of the Kalman-Bucy Fi l ter ( K B F ) which
provides an estimate of the states x and a Linear Quadratic Regulator ( L Q R ) which
is separately designed assuming all the states x are known Both the L Q R and K B F
designs independently have good robustness properties ie retain good performance
when disturbances due to model uncertainty are included However the combined
L Q R and K B F designs ie the L Q G design may have poor stability margins in the
presence of model uncertainties This l imitat ion has led to the development of an L Q G
design procedure that improves the performance of the compensator by recovering
the full state feedback robustness properties at the plant input or output (Figure 5-6)
This procedure is known as the Linear Quadratic Guassian-Loop Transfer Recovery
( L Q G L T R ) control algorithm A detailed development of its theory is given in
Ref[61] and hence is not repeated here for conciseness
The design of the L Q G L T R controller involves the repeated solution of comshy
plicated matrix equations too tedious to be executed by hand Fortunenately the enshy
tire algorithm is available in the Robust Control Toolbox[62] of the popular software
package M A T L A B The input matrices required for the function are the following
(i) state space matrices ABC and D (D = 0 for this system)
(ii) state and measurement noise covariance matrices E and 0 respectively
(iii) state and input weighting matrix Q and R respectively
114
u
MODEL UNCERTAINTY
SYSTEM PLANT
y
u LQR CONTROLLER
X KBF ESTIMATOR
u
y
LQG COMPENSATOR
PLANT
X = f v ( X t )
COMPENSATOR
x = A k x - B k y
ud
= C k f
y
Figure 5-6 Block diagram for the LQGLTR estimator based compensator
115
A s mentionned earlier the main objective of this offset controller is to regu-mdash
late the tether vibration described by the 5 equations However from the response
of the uncontrolled system presented in Chapter 4 it is clear that there is a large
difference between the magnitude of the librational and vibrational frequencies This
separation of frequencies allow for the separate design of the vibration and attitude mdash mdash
controllers Thus only the flexible subsystem composed of the S and d equations is
required in the offset controller design Similarly there is also a wide separation of
frequencies between the longitudinal and transverse modes of vibrations permitt ing mdash
the decoupling of the flexible subsystem into a set of longitudinal (Sx and dx) and
transverse (5y Sz dy and dz) subsystems The appended offset system d must also
be included since it acts as the control actuator However it is important to note
that the inplane and out-of-plane transverse modes can not be decoupled because
their oscillation frequencies are of the same order The two flexible subsystems are
summarized below
(i) Longitudinal Vibra t ion Subsystem
This is defined by u mdash Auxu + Buudu
(528)
where xu = SxdX25xdX2T u d u = dX2 and from Eq(527)
0 0 $u(lt) 0 0 0 0 1
0 0 1 0 0 0 0 1
(529)
Au and Bu are the rows and columns of A and B respectively and correspond to the
components of xu
The L Q R state weighting matrix Qu is taken as
Qu =
bull1 0 0 o -0 1 0 0 0 0 1 0
0 0 0 10
(530)
116
and the input weighting matrix Ru = 1^ The state noise covariance matr ix is
selected as 1 0 0 0 0 1 0 0 0 0 4 0
LO 0 0 1
while the measurement noise covariance matrix is taken as
(531)
i o 0 15
(532)
Given the above mentionned matrices the design of the L Q G L T R compenshy
sator is computed using M A T L A B with the following 2 commands
(i) kfu = l q r c ( ^ ( i C bdquo d i a g m x ( E u e u ) )
(ii) [AkugtBkugtCkugtDku =^ryAuBuCuDukfuQuRur)
where r is a scaler
The first command is used to compute the Ka lman filter gain matrix kfu Once
kfu is known the second command is invoked returning the state space representation
of the L Q G L T R dynamic compensator to
(533) xu mdash Akuxu Bkuyu
where xu is the state estimate vector oixu The function l t ry performs loop transfer
recovery at the system output ie the return ratio at the output approaches that
of the K B F loop given by Cu(sld - Au)Kfu[61 This is achieved by choosing a
sufficently large scaler value r in l t ry such that the singular values of the return
ratio approach those of the target design For the longitudinal controller design
r mdash 5 x 10 5 Figure 5-7(a) compares the singular values of the recovered compensator
design and the non-recovered L Q G compensator design with respect to the target
Unfortunetaly perfect recovery is not possible especially at higher frequencies
117
100
co
gt CO
-50
-100 10
Longitudinal Compensator Singular Values i I I I N I I mdash i mdash i i i i n i i mdash i mdash i i i i inimdash imdashi i i i i i i i mdash i mdash i i M i I I
Target r = 5x10 5
r = 0 (LQG)
i i i i I I i ii i i i 1 1 1 1 i i i i i 1 1 1 1 ii i I I i i i i 11111 -3 10 -2
(a)
101 10deg Frequency (rads)
101 10
Transverse Compensator Singular Values
100
50
X3
gt 0
CO
-50
bull100 10
-imdashi i 111II i 111 ni 1mdashi I I I I I I I 1mdashi I I I N I I 1mdashi i M 111
Target r = 50 r = 0 (LQG)
m i i i i i i n i i i i I I I I I I I 1mdashi I I I I I I I 1mdashi i 111 n -3 10 -2
(b)
101 10deg Frequency (rads)
101 10
Figure 5-7 Singular values for the LQG and LQGLTR compensator compared to target return ratio (a) longitudinal design (b) transverse design
118
because the system is non-minimal ie it has transmission zeros with positive real
parts[63]
(ii) Transverse Vibra t ion Subsystem
Here Xy mdash AyXy + ByU(lv
Dv = Cvxv
with xv = 6y 8Z dV2 dZ2 Sy 6Z dV2 dZ2T] udv = dy2dZ2T and
(534)
Cy
0 0 0 0 (0) 0 0 0 0 0 0 0 0 0 0 0 0
0 0 0 0
0 0 0 0 0 0 0 0 0 0 0 0
0 0 0
h 0 0 0
0 0 0 lt ( 0 ) o 0
0 1 0 0 0 1
0 0 0
2 r
h 0 0
0 0 1 0 0 1
(535)
Av and Bv are the rows and columns of the A and B corresponding to the components
O f Xy
The state L Q R weighting matrix Qv is given as
Qv
10 0 0 0 0 0 0 0 0 10 0 0 0 0 0 0 0 0 10 4 0 0 0 0 0 0 0 0 10 4 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1
(536)
mdash
Ry =w The state noise covariance matrix is taken
1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 9 x 10 4 0
0 0 0 0 0 0 0 9 x 10 4
(537)
119
while the measurement noise covariance matrix is
1 0 0 0 0 1 0 0 0 0 5 x 10 4 0
LO 0 0 5 x 10 4
(538)
As in the development of the longitudinal controller the transverse offset conshy
troller can be designed using the same two commands (lqrc and l try) wi th the input
matrices corresponding to the transverse subsystem with r = 50 The result is the
transverse compensator system given by
(539)
where xv is the state estimates of xv The singular values of the target transfer
function is compared with those achieved by the L Q G and L Q G L T R in Figure 5-
7(b) It is apparent that the recovery is not as good as that for the longitudinal case
again due to the non-minimal system Moreover the low value of r indicates that
only a lit t le recovery in the transverse subsystem is possible This suggests that the
robustness of the L Q G design is almost the maximum that can be achieved under
these conditions
Denoting f y = f j f j r xf = x^x^T and ud = udu ^ T the longishy
tudinal compensator can be combined to give
xf =
$d =
Aku
0
0 Akv
Cku o
Cu 0
Bku
0 B Xf
Xf = Ckxf
0
kv
Hf = Akxf - Bkyf
(540)
Vu Vv 0 Cv
Xf = CfXf
Defining a permutation matrix Pf such that Xj = PfX where X is the state vector
of the original nonlinear system given in Eq(32) the compensator and full nonlinear
120
system equations can be combined as
X
(541)
Ckxf
A block diagram representation of Eq(541) is shown in Figure 5-6
524 Simulation results
The dynamical response for the stationkeeping STSS with a tether length of 20
km and ini t ia l offset position of 1 m in each direction is presented in Figure 5-8 when
both the F L T attitude controller and the L Q G L T R offset controller are activated As
expected the offset controller is successful in quickly damping the elastic vibrations
in the longitudinal and transverse direction (Figure 5-8b) However from Figure 5-
8(a) it is clear that the presence of offset control requires a larger control moment
to regulate the attitude of the platform This is due to the additional torque created
by the larger moment arm around 2 m and 125 m in the inplane and out-of-plane
directions respectively introduced by the offset control In addition the control
moment for the platform is modulated by the tethers transverse vibration through
offset coupling However this does not significantly affect the l ibrational motion of
the tether whose thruster force remains relatively unchanged from the uncontrolled
vibration case
Final ly it can be concluded that the tether elastic vibration suppression
though offset control in conjunction with thruster and momentum-wheel attitude
control presents a viable strategy for regulating the dynamics tethered satellite sysshy
tems
121
STSS configuration 3-Body a1(0)=a2(0)=at(0)=2deg P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m6y(0)=52(0)=1m dx(0)=dy(0)=dz(0)=1m Stationkeeping I = 20km LQGLTR Vibration Control
Sat 1 Pitch and Roll Angles Tether Pitch and Roll Angles mdash 1 1 1 1 1 1 I 1 1 lt 1 ~ ^ 1 ^ l _
Time (Orbits) Time (Orbits) Sat 1 Roll Control Moment Tether Roll Control Thrust
(a) Time (Orbits) Time (Orbits)
Figure 5-8 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT attitude controller and LQGLTR offset vibration controller (a) attitude and libration controller reshysponse
122
STSS configuration 3-Body a1(0)=a2(0)=a(0)=2deg p1(0)=(32(0)=Q(0)=1deg ox(0)=14mSy(0)=5z(0)=1m dx(0)=dy(0)=dz(0)=1m Stationkeeping lt = 20km LQGLTR Vibration Control
Tether X Offset Position
Time (Orbits) Tether Y Tranverse Vibration 20r
Time (Orbits) Tether Y Offset Position
t o
1 2 Time (Orbits)
Tether Z Tranverse Vibration Time (Orbits)
Tether Z Offset Position
(b)
Figure 5-8
Time (Orbits) Time (Orbits)
Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT attitude controller and LQGLTR offset vibration controller (b) vibration and offset response
123
6 C O N C L U D I N G R E M A R K S
61 Summary of Results
The thesis has developed a rather general dynamics formulation for a multi-
body tethered system undergoing three-dimensional motion The system is composed
of multiple rigid bodies connected in a chain configuration by long flexible tethers
The tethers which are free to undergo libration as well as elastic vibrations in three
dimensions are also capable of deployment retrieval and constant length stationkeepshy
ing modes of operation Two types of actuators are located on the rigid satellites
thrusters and momentum-wheels
The governing equations of motion are developed using a new Order(N) apshy
proach that factorizes the mass matrix of the system such that it can be inverted
efficiently The derivation of the differential equations is generalized to account for
an arbitrary number of rigid bodies The equations were then coded in FORTRAN
for their numerical integration with the aid of a symbolic manipulation package that
algebraically evaluated the integrals involving the modes shapes functions used to dis-
cretize the flexible tether motion The simulation program was then used to assess the
uncontrolled dynamical behaviour of the system under the influence of several system
parameters including offset at the tether attachment point stationkeeping deployshy
ment and retrieval of the tether The study covered the three-body and five-body
geometries recently flown OEDIPUS system and the proposed BICEPS configurashy
tion It represents innovation at every phase a general three dimensional formulation
for multibody tethered systems an order-N algorithm for efficient computation and
application to systems of contemporary interest
Two types of controllers one for the attitude motion and the other for the
124
flexible vibratory motion are developed using the thrusters and momentum-wheels for
the former and the variable offset position for the latter These controllers are used to
regulate the motion of the system under various disturbances The attitude controller
is developed using the nonlinear Feedback Linearization Technique ( F L T ) and is based
on the nonlinear but rigid model of the system O n the other hand the separation
of the longitudinal and transverse frequencies from those of the attitude response
allows for the development of a linear optimal offset controller using the robust Linear
Quadratic Gaussian-Loop Transfer Recovery ( L Q G L T R ) method The effectiveness
of the two controllers is assessed through their subsequent application to the original
nonlinear flexible model
More important original contributions of the thesis which have not been reshy
ported in the literature include the following
(i) the model accounts for the motion of a multibody chain-type system undershy
going librational and in the case of tethers elastic vibrational motion in all
the three directions
(ii) the dynamics formulation is based on a recursive O(N) Lagrangian algorithm
that efficiently computes the systems generalized acceleration vector
(iii) the development of an attitude controller based on the Feedback Linearizashy
tion Technique for a multibody system using thrusters and momentum-wheels
located on each rigid body
(iv) the design of a three degree of freedom offset controller to regulate elastic v i shy
brations of the tether using the Linear Quadratic Gaussian and Loop Transfer
Recovery method ( L Q G L T R )
(v) substantiation of the formulation and control strategies through the applicashy
tion to a wide variety of systems thus demonstrating its versatility
125
The emphasis throughout has been on the development of a methodology to
study a large class of tethered systems efficiently It was not intended to compile
an extensive amount of data concerning the dynamical behaviour through a planned
variation of system parameters O f course a designer can easily employ the user-
friendly program to acquire such information Rather the objective was to establish
trends based on the parameters which are likely to have more significant effect on
the system dynamics both uncontrolled as well as controlled Based on the results
obtained the following general remarks can be made
(i) The presence of the platform products of inertia modulates the attitude reshy
sponse of the platform and gives rise to non-zero equilibrium pitch and roll
angles
(ii) Offset of the tether attachment point significantly affects the equil ibrium orishy
entation of the system as well as its dynamics through coupling W i t h a relshy
atively small subsatellite the tether dynamics dominate the system response
with high frequency elastic vibrations modulating the librational motion On
the other hand the effect of the platform dynamics on the tether response is
negligible As can be expected the platform dynamics is significantly affected
by the offset along the local horizontal and the orbit normal
(iii) For a three-body system deployment of the tether can destabilize the platform
in pitch as in the case of nonzero offset However the roll motion remains
undisturbed by the Coriolis force Moreover deployment can also render the
tether slack if it proceeds beyond a critical speed
(iv) Uncontrolled retrieval of the tether is always unstable in pitch It also leads
to high frequency elastic vibrations in the tether
(v) The five-body tethered system exhibits dynamical characteristics similar to
those observed for the three-body case As can be expected now there are
additional coupling effects due to two extra bodies a rigid subsatellite and a
126
flexible tether
(vi) Cartwheeling motion of the B I C E P S configuration can also be used to adshy
vantage in the deployment of the tether However exceeding a crit ical ini t ia l
cartwheeling rate can result in a large tension in the tether which eventually
causes the satellite to bounce back rendering the tether slack
(vii) Spinning the tether about its nominal length as in the case of O E D I P U S
introduces high frequency transverse vibrations in the tether which in turn
affect the dynamics of the satellites
(viii) The F L T based controller is quite successful in regulating the attitude moshy
tion of both the tether and rigid satellites in a short period of time during
stationkeeping deployment as well as retrieval phases
(ix) The offset controller is successful in suppressing the elastic vibrations of the
tether wi thin the allowable l imits of the offset mechanism ( plusmn 2 0 m)
62 Recommendations for Future Study
A s can be expected any scientific inquiry is merely a prelude to further efforts
and insight Several possible avenues exist for further investigation in this field A
few related to the study presented here are indicated below
(i) inclusion of environmental effects particularly the free molecular reaction
forces as well as interaction with Earths magnetic field and solar radiation
(ii) addition of flexible booms attached to the rigid satellites providing a mechashy
nism for energy dissipation
(iii) validation of the new three dimensional offset control strategies using ground
based experiments
(iv) animation of the simulation results to allow visual interpretation of the system
dynamics
127
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[38] Ruying F and Bainum P M Dynamics and Control of a Space Platform with a Tethered Subsatellite Journal of Guidance Control and Dynamics V o l 11 No 4 July-August 1988 pp 377-381
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[43] Pradhan S Planar Dynamics and Control of Tethered Satellite Systems P h D Thesis Department of Mechanical Engineering The University of Br i t i sh Columbia December 1994
[44] Bainum P M and Kumar V K Optimal-Control of the Shuttle-Tethered System Acta Astronautica V o l 7 No 12 1980 pp 1333-1348
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[48] Nayfeh A H and Mook D T Nonlinear Oscillation John Wi ley and Sons New York U S A 1979 pp 485-500
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[50] Meirovitch L Elements of Vibration Analysis M c G r a w - H i l l Inc New York U S A 1986 pp 255-256
[51] Greenwood D T Principles of Dynamics Prentice Ha l l Inc Englewood Cliffs New Jersey U S A 1965 pp 252-275
[52] Boyce W E and D i P r i m a R C Elemental Differential Equations and Boundshy
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[54] Gear C W Numerical Initial Value Problems in Ordinary Differential Equashy
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132
A P P E N D I X I TENSOR R E P R E S E N T A T I O N OF T H E
EQUATIONS OF M O T I O N
11 Prel iminary Remarks
The derivation of the equations of motion for the multibody tethered system
involves the product of several matrices as well as the derivative of matrices and
vectors with respect to other vectors In order to concisely code these relationships
in FORTRAN the matrix equations of motion are expressed in tensor notation
Tensor mathematics in general is an extremely powerful tool which can be
used to solve many complex problems in physics and engineering[6465] However
only a few preliminary results of Cartesian tensor analysis are required here They
are summarized below
12 Mathematical Background
The representation of matrices and vectors in tensor notation simply involves
the use of indices referring to the specific elements of the matrix entity For example
v = vk (k = lN) (Il)
A = Aij (i = lNj = lM)
where vk is the kth element of vector v and Aj is the element on the t h row and
3th column of matrix A It is clear from Eq(Il) that exactly one index is required to
completely define an entire vector whereas two independent indices are required for
matrices For this reason vectors and matrices are known as first-order and second-
order tensors respectively Scaler variables are known as zeroth-order tensors since
in this case an index is not required
133
Matrix operations are expressed in tensor notation similar to the way they are
programmed using a computer language such as FORTRAN or C They are expressed
in a summation notation known as Einstein notation For example the product of a
matrix with a vector is given as
w = Av (12)
or in tensor notation W i = ^2^2Aij(vkSjk)
4-1 ltL3gt j
where 8jk is the Kronecker delta Dropping the summation symbol the matrix-vector
product can be expressed compactly as
Wi = AijVj = Aimvm (14)
Here j is a dummy index since it appears twice in a single term and hence can
be replaced by any other index Note that since the resulting product is a vector
only one index is required namely i that appears exactly once on both sides of the
equation Similarly
E = aAB + bCD (15)
or Ej mdash aAimBmj + bCimDmj
~ aBmjAim I bDmjCim (16)
where A B C D and E are second-order tensors (matrices) and a and b are zeroth-
order tensors (scalers) In addition once an expression is in tensor form it can be
treated similar to a scaler and the terms can be rearranged
The transpose of a matrix is also easily described using tensors One simply
switches the position of the indices For example let w = ATv then in tensor-
notation
Wi = AjiVj (17)
134
The real power of tensor notation is in its unambiguous expression of terms
containing partial derivatives of scalers vectors or matrices with respect to other
vectors For example the Jacobian matrix of y = f(x) can be easily expressed in
tensor notation as
di-dx-j-1 ( L 8 )
mdash which is a second-order tensor as expected If f(x) = Ax where A is constant then
f - a )
according to the current definition If C = A(x)B(x) then the time derivative of C
is given by
C = Ax)Bx) + A(x)Bx) (110)
or in tensor notation
Cij = AimBmj + AimBmj (111)
= XhiAimfcBmj + AimBmj^k)
which is more clear than its equivalent vector form Note that A^^ and Bmjk are
third-order tensors that can be readily handled in F O R T R A N
13 Forcing Function
The forcing function of the equations of motion as given by Eq(260) is
- u 1-SdM- dqtdPe 9 f i m F(q qt) = Mq - -q mdashq + mdashmdash - Qd (260)
This expression can be converted into tensor form to facilitate its implementation
into F O R T R A N source code
Consider the first term Mq From Eq(289) the mass matrix M is given by
T M mdash Rv MtRv which in tensor form is
Mtj = RZiM^R^j (112)
135
Taking the time derivative of M gives
Mij = qsRv
ni^MtnrnRv
mj + Rv
niMtnTn^Rv
mj + Rv
niMtnmRv
mjs) (113)
The second term on the right hand side in Eq(260) is given by
-JPdMu 1 bdquo T x
2Q ~cWq = 2Qs s r k Q r ^ ^
Expanding Msrk leads to
M src = RnskMtnmRmr + RnsMtnmkRmr + RnsMtnmRmrk- (L15)
Finally the potential energy term in Eq(260) is also expanded into tensor
form to give dPe = dqt dPe
dq dq dqt = dq^QP^
and since ^ = Rp(q)q then
Now inserting Eqs(I13-I17) into Eq(260) and rearranging the tensor form of the
forcing term can how be stated as
Fk(q^q^)=(RP
sk + R P n J q n ) ^ - - Q d k
+ QsQr | (^Rlks ~ 2 M t n m R m r
+ (^RnkMtnms ~ 2~RnsMtnmgtk^ R m r
(^RnkRmrs ~ ~2RnsRmrk^j
where Fk(q q t) represents the kth component of F
(118)
136
A P P E N D I X II R E D U C E D EQUATIONS OF M O T I O N
II 1 Prel iminary Remarks
The reduced model used for the design of the vibration controller is represented
by two rigid end-bodies capable of three-dimensional attitude motion interconnected
with a fixed length tether The flexible tether is discretized using the assumed-
mode method with only the fundamental mode of vibration considered to represent
longitudinal as well as in-plane and out-of-plane transverse deformations In this
model tether structural damping is neglected Finally the system is restricted to a
nominal circular orbit
II 2 Derivation of the Lagrangian Equations of Mot ion
Derivation of the equations of motion for the reduced system follows a similar
procedure to that for the full case presented in Chapter 2 However in this model the
straightforward derivation of the energy expressions for the entire system is undershy
taken ignoring the Order(N) approach discussed previously Furthermore in order
to make the final governing equations stationary the attitude motion is now referred
to the local LVLH frame This is accomplished by simply defining the following new
attitude angles
ci = on + 6
Pi = Pi (HI)
7t = 0
where and are the pitch and roll angles respectively Note that spin is not
considered in the reduced model hence 7J = 0 Substituting Eq(IIl) into Eq(214)
137
gives the new expression for the rotation matrix as
T- CpiSa^ Cai+61 S 0 S a + e i
-S 0 c Pi
(II2)
W i t h the new rotation matrix defined the inertial velocity of the elemental
mass drrii o n the ^ link can now be expressed using Eq(215) as
(215)
Since the tether length is assumed to be fixed T j f j and Tj$jlt^ of Eq(215) are
identically zero Also defining
(II3) Vi = I Pi i
Rdm- f deg r the reduced system is given by
where
and
Rdm =Di + Pi(gi)m + T^iSi
Pi(9i)m= [Tai9i T0gi T^] fj
(II4)
(II5)
(II6)
D1 = D D l D S v
52 = 31 + P(d2)m +T[d2
D3 = D2 + l2 + dX2Pii)m + T2i6x
Note that $ j for i = 2 is defined in Section 212 whereas for i = 1 and 3 it is
the null matrix Since only one mode is used for each flexible degree of freedom
5i = [6Xi6yi6Zi]T for i = 2
The kinetic energy of the system can now be written as
i=l Z i = l Jmi RdmRdmdmi (UJ)
138
Expanding Kampi using Eq(II4)
KH = miDD + 2DiP
i j gidmi)fji + 2D T ltMtrade^
+tf J P^(9i)Pi9i)d^l + 2mT j PFmTi^dm^ (II8)
where the integrals are evaluated using the procedure similar to that described in
Chapter 2 The gravitational and strain energy expressions are given by Eq(247)
and Eq(251) respectively using the newly defined rotation matrix T in place of
TV
Substituting the kinetic and potential energy expressions in
d ( 8Ke d(Ke - Pe) i bull i Q ^ 0 (II9)
d t dqred d(ired
with
Qred = [ algt l gt a 2 2 A 2 A 2 lt ^ 2 Q 3 3 ] (5-14)
and d2 regarded as a time-varying parameter the nonlinear non-autonomous equashy
tions of motion for the reduced system can finally be expressed as
MredQred + fred = reg- (5-15)
139
Dx- transformation matrix relating D j and Ds-
DXi Dyi Dz- Cartesian components of D
DZl out-of-plane position component of first link
E Youngs elastic modulus for the ith l ink
Ej^ contribution from structural damping to the augmented complex
modulus E
F systems conservative force vector
FQ inertial reference frame
Fi t h link body-fixed reference frame
n x n identity matrix
alternate form of inertia matrix for the i 1 l ink
K A Z $i(k + dXi+1)
Kei t h link kinetic energy
M(q t) systems coupled mass matrix
Mma Mma Mmy- control moments in the pitch roll and yaw directions respec-
tively for the t h link
Mr fr rigid mass matrix and force vector respectively
Mred fred- Qred mass matrix force vector and generalized coordinate vector for
the reduced model respectively
Mt block diagonal decoupled mass matrix
symmetric decoupled mass matrix
N total number of links
0N) order-N
Pi(9i) column matrix [T^giTpg^T^gi) mdash
Q nonconservative generalized force vector
Qu actuator coupling matrix or longitudinal L Q R state weighting
matrix
v i i i
Rdm- inertial position of the t h link mass element drrii
RP transformation matrix relating qt and tf
Ru longitudinal L Q R input weighting matrix
RvRn Rd transformation matrices relating qt and q
S(f Mdc) generalized acceleration vector of coupled system q for the full
nonlinear flexible system
th uiirv l u w u u u maniA
th
T j i l ink rotation matrix
TtaTto control thrust in the pitch and roll direction for the im l ink
respectively
rQi maximum deploymentretrieval velocity of the t h l ink
Vei l ink strain energy
Vgi link gravitational potential energy
Rayleigh dissipation function arising from structural damping
in the ith link
di position vector to the frame F from the frame F_i bull
dc desired offset acceleration vector (i 1 l ink offset position)
drrii infinitesimal mass element of the t h l ink
dx^dy^dZi Cartesian components of d along the local vertical local horishy
zontal and orbit normal directions respectively
F-Q mdash
gi r igid and flexible position vectors of drrii fj + ltEraquolt5
ijk unit vectors 1 0 0 r 010-^ and 00 l r respectively
li length of the t h l ink
raj mass of the t h link
nfj total number of flexible modes for the i 1 l ink nXi + nyi + nZi
nq total number of generalized coordinates per link nfj + 7
nqq systems total number of generalized coordinates Nnq
ix
number of flexible modes in the longitudinal inplane and out-
of-plane transverse directions respectively for the i ^ link
number of attitude control actuators
- $ T
set of generalized coordinates for the itfl l ink which accounts for
interactions with adjacent links
qv---QtNT
set of coordinates for the independent t h link (not connected to
adjacent links)
rigid position of dm in the frame Fj
position of centre of mass of the itfl l ink relative to the frame Fj
i + $DJi
desired settling time of the j 1 attitude actuator
actuator force vector for entire system
flexible deformation of the link along the Xj yi and Zj direcshy
tions respectively
control input for flexible subsystem
Cartesian components of fj
actual and estimated state of flexible subsystem respectively
output vector of flexible subsystem
LIST OF FIGURES
1-1 A schematic diagram of the space platform based N-body tethered
satellite system 2
1-2 Associated forces for the dumbbell satellite configuration 3
1- 3 Some applications of the tethered satellite systems (a) multiple
communication satellites at a fixed geostationary location
(b) retrieval maneuver (c) proposed B I C E P S configuration 6
2- 1 Vector components of the i 1 and (i- l ) 1 chain links 15
2-2 Vector components in cylindrical orbital coordinates 20
2-3 Inertial position of subsatellite thruster forces 38
2- 4 Coupled force representation of a momentum-wheel on a rigid body 40
3- 1 Flowchart showing the computer program structure 45
3-2 Kinet ic and potentialenergy transfer for the three-body platform based
tethered satellite system (a) variation of kinetic and potential energy
(b) percent change in total energy of system 48
3-3 Kinet ic and potential energy transfer for the five-body platform based
tethered satellite system (a) variation of kinetic and potential energy
(b) percent change in total energy of system 49
3-4 Simulation results for the platform based three-body tethered
system originally presented in Ref[43] 51
3- 5 Simulation results for the platform based three-body tethered
system obtained using the present computer program 52
4- 1 Schematic diagram showing the generalized coordinates used to
describe the system dynamics 55
4-2 Stationkeeping dynamics of the three-body STSS configuration
without offset (a) attitude response (b) vibration response 57
4-3 Stationkeeping dynamics of the three-body STSS configuration
xi
with offset along the local vertical
(a) attitude response (b) vibration response 60
4-4 Stationkeeping dynamics of the three-body STSS configuration
with offset along the local horizontal
(a) attitude response (b) vibration response 62
4-5 Stationkeeping dynamics of the three-body STSS configuration
with offset along the orbit normal
(a) attitude response (b) vibration response 64
4-6 Stationkeeping dynamics of the three-body STSS configuration
wi th offset along the local horizontal local vertical and orbit normal
(a) attitude response (b) vibration response 67
4-7 Deployment dynamics of the three-body STSS configuration
without offset
(a) attitude response (b) vibration response 69
4-8 Deployment dynamics of the three-body STSS configuration
with offset along the local vertical
(a) attitude response (b) vibration response 72
4-9 Retrieval dynamics of the three-body STSS configuration
with offset along the local vertical and orbit normal
(a) attitude response (b) vibration response 74
4-10 Schematic diagram of the five-body system used in the numerical example 77
4-11 Stationkeeping dynamics of the five-body STSS configuration
without offset
(a) attitude response (b) vibration response 78
4-12 Deployment dynamics of the five-body STSS configuration
wi th offset along the local vertical
(a) attitude response (b) vibration response 81
4-13 Stationkeeping dynamics of the three-body B I C E P S configuration
x i i
with offset along the local vertical
(a) attitude response (b) vibration response 84
4-14 Cartwheeling dynamics with deployment for the three-body B I C E P S
with offset along the local vertical
(a) attitude response (b) vibration response 86
4-15 Spin dynamics (7 = l deg s ) of the three-body O E D I P U S configuration
with offset along the local vertical
(a) attitude response (b) vibration response 89
4- 16 Spin dynamics (7 = 10deg s ) of the three-body O E D I P U S configuration
wi th offset along the local vertical
(a) attitude response (b) vibration response 91
5- 1 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear rigid F L T controller with offset along the
local horizontal local vertical and orbit normal
(a) attitude and vibration response (b) control actuator time histories 98
5-2 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear flexible F L T controller with offset along the
local horizontal local vertical and orbit normal
(a) attitude and vibration response (b) control actuator time histories 101
5-3 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear rigid F L T controller with offset along
the local vertical
(a) attitude and vibration response (b) control actuator time histories 103
5-4 Deployment dynamics of the three-body STSS using the nonlinear
rigid F L T controller with offset along the local vertical
(a) attitude and vibration response (b) control actuator time histories 105
5-5 Retrieval dynamics of the three-body STSS using the non-linear
rigid F L T controller with offset along the local vertical and orbit normal
(a) attitude and vibration response (b) control actuator time histories 108
x i i i
5-6 Block diagram for the L Q G L T R estimator based compensator 115
5-7 Singular values for the L Q G and L Q G L T R compensator compared to target
return ratio (a) longitudinal design (b) transverse design 118
5-8 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear rigid F L T attitude controller and L Q G L T R
offset vibration controller (a) attitude and libration controller response
(b) vibration and offset response 122
xiv
A C K N O W L E D G E M E N T
Firs t and foremost I would like to express the most genuine gratitude to
my supervisor Prof V i n o d J M o d i whose unending guidance and encouragement
proved most invaluable during the course of my studies
I am also deeply indebted to Dr Satyabrata Pradhan for his limitless patience
and technical advice from which this work would not have been possible I would also
like to extend a word of appreciation towards Prof A r u n K Mis ra ( M c G i l l University)
for ini t iat ing my interest in space dynamics and control
Also I would like to thank my collegues Dr Sandeep Munshi M r Mathieu
Caron and M r Yuan Chen as well as Dr Mae Seto M r Gary L i m and M r Mark
Chu for their words of advice and helpful suggestions that heavily contributed to all
aspects of my studies
Final ly I would like to thank all my family and friends here in Vancouver and
back home in Montreal whose unending support made the past two years the most
memorable of my life
The research project was supported by the Natural Sciences and Engineering
Research Council s P G S - A Scholarship held by the author as well as by N S E R C grant
A-2181 held by Prof V J M o d i
xv
1 INTRODUCTION
11 Prel iminary Remarks
W i t h the ever changing demands of the worlds population one often wonders
about the commitment to the space program On the other hand the importance
of worldwide communications global environmental monitoring as well as long-term
habitation in space have demanded more ambitious and versatile satellites systems
It is doubtful that the Russian scientist Tsiolkovsky[l] when he first proposed the
uti l izat ion of the Earths gravity-gradient environment in the last century would have
envisioned the current and proposed applications of tethered satellite systems
A tethered satellite system consists of two or more subsatellites connected
to each other by long thin cables or tethers which are in tension (Figure 1-1) The
subsatellite which can also include the shuttle and space station are in orbit together
The first proposed use of tethers in space was associated with the rescue of stranded
astronauts by throwing a buoy on a tether and reeling it to the rescue vehicle
Prel iminary studies of such systems led to the discovery of the inherent instability
during the tether retrieval[2] Nevertheless the bir th of tethered systems occured
during the Gemini X I and X I I missions[3] in 1966 when a short tether (10-30m)
was used to generate the first artificial gravity environment in space (000015g) by
cartwheeling (spinning) the system about its center of mass
The mission also demonstrated an important use of tethers for gravity gradient
stabilization The force analysis of a simplified model illustrates this point Consider
the dumbbell satellite configuration orbiting about the centre of Ear th as shown in
Figure 1-2 Satellite 1 and satellite 2 are connected to each other by a tether with
1
Space Station (Satellite 1 )
2
Satellite N
Figure 1-1 A schematic diagram of the space platform based N-body tethered satellite system
2
Figure 1-2 Associated forces for the dumbbell satellite configuration
3
the whole system free to librate about the systems centre of mass C M The two
major forces acting on each body are due to the centrifugal and gravitational effects
Since body 1 is further away from the earth its gravitational force Fg^y is smaller
than that of body 2 ie Fgi lt Fg^ However its centrifugal force FCl is greater
then the centrifugal force experienced by the second satellite (FCl gt FC2) Moreover
for body 1 Fg^ lt Fc^ in contrast to Fg2 gt FC2 for body 2 Thus resolving the force
components along the tether there is an evident resultant tension force Ff present
in the tether Similarly adding the normal components of the two forces vectorially
results in a force FR which restores the system to its stable equil ibrium along the
local vertical These tension and gravity-gradient restoring forces are the two most
important features of tethered systems
Several milestone missions have flown in the recent past The U S A J a p a n
project T P E (Tethered Payload Experiment) [4] was one of the first to be launched
using a sounding rocket to conduct environmental studies The results of the T P E
provided support for the N A S A I t a l y shuttle based tethered satellite system referred
to as the TSS-1 mission[5] This experiment aimed to study the electrodynamic effects
of a shuttle borne satellite system with a conductive tether connection where an
electric current was induced in the tether as it swept through the Earth s magnetic
field Unfortunetely the two attempts in August 1992 and February 1996 resulted
in only partial success The former suffered from a spool failure resulting in a tether
deployment of only 256m of a planned 20km During the latter attempt a break
in the tethers insulation resulted in arcing leading to tether rupture Nevertheless
the information gathered by these attempts st i l l provided the engineers and scientists
invaluable information on the dynamic behaviour of this complex system
Canada also paved the way with novel designs of tethered system aimed
primari ly at the study of the ionosphere and Northern Lights (Aurora Borealis)
4
The O E D I P U S A and C (Observation of Electrified Distributions in the Ionospheric
P lasma - a Unique Strategy) missions[6] launched from a sounding rocket in January
1989 and November 1995 respectively provided insight into the complex dynamical
behaviour of two comparable mass satellites connected by a 1 km long spinning
tether
Final ly two experiments were conducted by N A S A called S E D S I and II
(Small Expendable Deployable System) [7] which hold the current record of 20 km
long tether In each of these missions a small probe (26 kg) was deployed from the
second stage of a D E L T A II rocket thus succesfully demonstrating the feasibility of
long tether deployment Retrieval of the tether which is significantly more difficult
has not yet been achieved
Several proposed applications are also currently under study These include the
study of earths upper atmosphere using probes lowered from the shuttle in a low earth
orbit ( L E O ) to the deployment and retrieval of satellites for servicing Even more
promising application concerns the generation of power for the proposed International
Space Station using conductive tethers The Canadian Space Agency (CSA) has also
proposed a dual-mass satellite system B I C E P S (Bl-static Canadian Experiment on
Plasmas in Space) to make simultaneous measurements at different locations in the
environment for correlation studies[8] A unique feature of the B I C E P S mission is
the deployment of the tether aided by the angular momentum of the system which
is ini t ia l ly cartwheeling about its orbit normal (Figure 1-3)
In addition to the three-body configuration multi-tethered systems have been
proposed for monitoring Earths environment in the global project monitored by
N A S A entitled Mission to Planet Ear th ( M T P E ) The proposed mission aims to
study pollution control through the understanding of the dynamic interactions be-
5
Retrieval of Satellite for Servicing
lt mdash
lt Satellite
Figure 1-3 Some applications of the tethered satellite systems (a) multiple comshymunication satellites at a fixed geostationary location (b) retrieval maneuver (c) proposed BICEPS configuration
6
tween the Earths atmosphere biosphere hydrosphere and cryosphere This wi l l be
accomplished with multiple tethered instrumentation payloads simultaneously soundshy
ing at different altitudes for spatial correlations Such mult ibody systems are considshy
ered to be the next stage in the evolution of tethered satellites Micro-gravity payload
modules suspended from the proposed space station for long-term experiments as well
as communications satellites with increased line-of-sight capability represent just two
of the numerous possible applications under consideration
12 Brief Review of the Relevant Literature
The design of multi-payload systems wi l l require extensive dynamic analysis
and parametric study before the final mission configuration can be chosen This
wi l l include a review of the fundamental dynamics of the simple two-body tethered
systems and how its dynamics can be extended to those of multi-body tethered system
in a chain or more generally a tree topology
121 Multibody 0(N) formulation
Many studies of two-body systems have been reported One of the earliest
contribution is by Rupp[9] who was the first to study planar dynamics of the simshy
plified Shuttle based Tethered Satellite System (STSS) His pioneering work led to
the discovery of pitch oscillation growth during the retrieval phase Later Baker[10]
advanced the investigation to the third dimension in addition to adding atmospheric
effects to the system A more complete dynamical model was later developed and
analyzed by M o d i and Misra[ l l 12] which included deployment and tether flexibility
A comprehensive survey of important developments and milestones in the area have
been compiled and reviewed by Mis ra and Modi[13] The major conclusions based on
the literature may be summarized as follows
bull the stationkeeping phase is stable
7
bull deployment can be unstable if the rate exceeds a crit ical speed
bull retrieval of the tether is inherently unstable
bull transverse vibrations can grow due to the Coriolis force induced during deshy
ployment and retrieval
Misra Amier and Modi[14] were one of the first to extend these results to the
three-satellite double pendulum case This simple model which includes a variable
length tether was sufficient to uncover the multiple equilibrium configurations of
the system Later the work was extended to include out-of-plane motion and a
preliminary reel-rate control law to regulate the librational motion[15] Kesmir i and
Misra[16] developed a general formulation for N-body tethered systems based on the
Lagrangian principle where three-dimensional motion and flexibility are accounted for
However from their work it is clear that as the number of payload or bodies increases
the computational cost to solve the forward dynamics also increases dramatically
Tradit ional methods of inverting the mass matrix M in the equation
ML+F = Q
to solve for the acceleration vector proved to be computationally costly being of the
order for practical simulation algorithms O(N^) refers to a mult ibody formulashy
tion algorithm whose computational cost is proportional to the cube of the number
of bodies N used It is therefore clear that more efficient solution strategies have to
be considered
Many improved algorithms have been proposed to reduce the number of comshy
putational steps required for solving for the acceleration vector Rosenthal[17] proshy
posed a recursive formulation based on the triangularization of the equations of moshy
tion to obtain an 0(N2) formulation He also demonstrates the use of a symbolic
manipulation program to reduce the final number of computations A recursive L a -
8
grangian formulation proposed by Book[18] also points out the relative inefficiency of
conventional formulations and proceeds to derive an algorithm which is also 0(N2)
A recursive formulation is one where the equations of motion are calculated in order
from one end of the system to the other It usually involves one or more passes along
the links to calculate the acceleration of each link (forward pass) and the constraint
forces (backward or inverse pass) Non-recursive algorithms express the equations of
motion for each link independent of the other Although the coupling terms st i l l need
to be included the algorithm is in general more amenable to parallel computing
A st i l l more efficient dynamics formulation is an O(N) algorithm Here the
computational effort is directly proportional to the number of bodies or links used
Several types of such algorithms have been developed over the years The one based
on the Newton-Euler approach has recursive equations and is used extensively in the
field of multi-joint robotics[1920] It should be emphasized that efficient computation
of the forward and inverse dynamics is imperative if any on-line (real-time) control
of the joint is to be achieved Hollerbach[21] proposed a recursive O(N) formulation
based on Lagranges equations of motion However his derivation was primari ly foshy
cused on the inverse dynamics and did not improve the computational efficiency of
the forward dynamics (acceleration vector) Other recursive algorithms include methshy
ods based on the principle of vir tual work[22] and Kanes equations of motion[23]
Keat[24] proposed a method based on a velocity transformation that eliminated the
appearance of constraint forces and was recursive in nature O n the other hand
K u r d i l a Menon and Sunkel[25] proposed a non-recursive algorithm based on the
Range-Space method which employs element-by-element approach used in modern
finite-element procedures Authors also demonstrated the potential for parallel comshy
putation of their non-recursive formulation The introduction of a Spatial Operator
Factorization[262728] which utilizes an analogy between mult ibody robot dynamics
and linear filtering and smoothing theory to efficiently invert the systems mass ma-
9
t r ix is another approach to a recursive algorithm Their results were further extended
to include the case where joints follow user-specified profiles[29] More recently Prad-
han et al [30] proposed a non-recursive Lagrangian algorithm for factorizing the mass
matrix leading to an O(N) formulation of the forward dynamics of the system A n
efficient O(N) algorithm where the number of bodies N in the system varies on-line
has been developed by Banerjee[31]
122 Issues of tether modelling
The importance of including tether flexibility has been demonstrated by several
researchers in the past However there are several methods available for modelling
tether flexibility Each of these methods has relative advantages and limitations
wi th regards to the computational efficiency A continuum model where flexible
motion is discretized using the assumed mode method has been proposed[3233] and
succesfully implemented In the majority of cases even one flexible mode is sufficient
to capture the dynamical behaviour of the tether[34] K i m and Vadali[35] proposed a
so-called bead model or lumped-mass approach that discretizes the tether using point
masses along its length However in order to accurately portray the motion of the
tether a significantly high number of beads are needed thus increasing the number of
computation steps Consequently this has led to the development of a hybrid between
the last two approaches[16] ie interpolation between the beads using a continuum
approach thus requiring less number of beads Finally the tether flexibility can also
be modelled using a finite-element approach[36] which is more suitable for transient
response analysis In the present study the continuum approach is adopted due to
its simplicity and proven reliability in accurately conveying the vibratory response
In the past the modal integrals arising from the discretization process have
been evaluted numerically In general this leads to unnecessary effort by the comshy
puter Today these integrals can be evaluated analytically using a symbolic in-
10
tegration package eg Maple V Subsequently they can be coded in F O R T R A N
directly by the package which could also result in a significant reduction in debugging
time More importantly there is considerable improvement in the computational effishy
ciency [16] especially during deployment and retrieval where the modal integral must
be evaluated at each time-step
123 Attitude and vibration control
In view of the conclusions arrived at by some of the researchers mentioned
above it is clear that an appropriate control strategy is needed to regulate the dyshy
namics of the system ie attitude and tether vibration First the attitude motion of
the entire system must be controlled This of course would be essential in the case of
a satellite system intended for scientific experiments such as the micro-gravity facilishy
ties aboard the proposed Internation Space Station Vibra t ion of the flexible members
wi l l have to be checked if they affect the integrity of the on-board instrumentation
Thruster as well as momentum-wheel approach[34] have been considered to
regulate the rigid-body motion of the end-satellite as well as the swinging motion
of the tether The procedure is particularly attractive due to its effectiveness over a
wide range of tether lengths as well as ease of implementation Other methods inshy
clude tension and length rate control which regulate the tethers tension and nominal
unstretched length respectively[915] It is usually implemented at the deployment
spool of the tether More recently an offset strategy involving time dependent moshy
tion of the tether attachment point to the platform has been proposed[3738] It
overcomes the problem of plume impingement created by the thruster control and
the ineffectiveness of tension control at shorter lengths However the effectiveness of
such a controller can become limited with an exceedingly long tether due to a pracshy
tical l imi t on the permissible offset motion In response to these two control issues a
hybrid thrusteroffset scheme has been proposed to combine the best features of the
11
two methods[39]
In addition to attitude control these schemes can be used to attenuate the
flexible response of the tether Tension and length rate control[12] as well as thruster
based algorithms[3340] have been proposed to this end M o d i et al [41] have sucshy
cessfully demonstrated the effectiveness of an offset strategy to damp undesired v i shy
brations Passive energy dissipative devices eg viscous dampers are also another
viable solution to the problem
The development of various control laws to implement the above mentioned
strategies has also recieved much attention A n eigenstructure assignment in conducshy
tion wi th an offset controller for vibration attenuation and momemtum wheels for
platform libration control has been developed[42] in addition to the controller design
from a graph theoretic approach[41] Also non-linear feedback methods such as the
Feedback Linearization Technique ( F L T ) that are more suitable for controlling highly
nonlinear non-autonomous coupled systems have also been considered[43]
It is important to point out that several linear controllers including the classhy
sic state feedback Linear Quadratic Regulator ( L Q R ) have received considerable
attention[3944] Moreover robust methods such as the Linear Quadratic Guas-
s ian Loop Transfer Recovery ( L Q G L T R ) method have also been developed and imshy
plemented on tethered systems[43] A more complete review of these algorithms as
well as others applied to tethered system has been presented by Pradhan[43]
13 Scope of the Investigation
The objective of the thesis is to develop and implement a versatile as well as
computationally efficient formulation algorithm applicable to a large class of tethered
satellite systems The distinctive features of the model can be summarized as follows
12
bull TV-body 0(N) tethered satellite system in a chain-type topology
bull the system is free to negotiate a general Keplerian orbit and permitted to
undergo three dimensional inplane and out-of-plane l ibrtational motion
bull r igid bodies constituting the system are free to execute general rotational
motion
bull three dimensional flexibility present in the tether which is discretized using
the continuum assumed-mode method with an arbitrary number of flexible
modes
bull energy dissipation through structural damping is included
bull capability to model various mission configurations and maneuvers including
the tether spin about an arbitrary axis
bull user-defined time dependent deployment and retrieval profiles for the tether
as well as the tether attachment point (offset)
bull attitude control algorithm for the tether and rigid bodies using thrusters and
momentum-wheels based on the Feedback Linearization Technique
bull the algorithm for suppression of tether vibration using an active offset (tether
attachment point) control strategy based on the optimal linear control law
( L Q G L T - R )
To begin with in Chapter 2 kinematics and kinetics of the system are derived
using the O(N) formulation methodology developed by Pradhan et al [30] Chapshy
ter 3 discusses issues related to the development of the simulation program and its
validation This is followed by a detailed parametric study of several mission proshy
files simulated as particular cases of the versatile formulation in Chapter 4 Next
the attitude and vibration controllers are designed and their effectiveness assessed in
Chapter 5 The thesis ends with concluding remarks and suggestions for future work
13
2 F O R M U L A T I O N OF T H E P R O B L E M
21 Kinematics
211 Preliminary definitions and the i ^ position vector
The derivation of the equations of motion begins with the definition of the
inertial position of an arbitrary link of the mult ibody system Let the i ^ link of the
system be free to translate and rotate in 3-D space From Figure 2-1 the position
vector Rdm^ to the mass element drrn on the i ^ link wi th reference to the inertial
frame FQ can be written as
Here D represents the inertial position of the t h body fixed frame Fj relative to FQ
ri = [xiyiZi]T is the rigid position vector of drrii wi th respect to F J and f(fi) is
the flexible deformation at fj also with respect to the frame Fj Bo th these vectors
are relative to Fj Note in the present model tethers (i even) are considered flexible
and X corresponds to the nominal position along the unstretched tether while a and
ZJ are by definition equal to zero On the other hand the rigid bodies referred to as
satellites (i odd) including the platform are taken to be rigid ie f(fj) = 0 T j is
the rotation matrix used to express body-fixed vectors with reference to the inertial
frame
212 Tether flexibility discretization
The tether flexibility is discretized with an arbitrary but finite number of
Di + Ti(i + f(i)) (21)
14
15
modes in each direction using the assumed-mode method as
Ui nvi
wi I I i = 1
nzi bull
(22)
where nXi nyi and n 2 ^ are the number of modes used in the X m and Z directions
respectively For the ] t h longitudinal mode the admissible mode shape function is
taken as 2 j - l
Hi(xiii)=[j) (23)
where l is the i ^ tether length[32] In the case of inplane and out-of-plane transverse
deflections the admissible functions are
ampyi(xili) = ampZixuk) = v ^ s i n JKXj
k (24)
where y2 is added as a normalizing factor In this case both the longitudinal and
transverse mode shapes satisfy the geometric boundary conditions for a simple string
in axial and transverse vibration
Eq(22) is recast into a compact matrix form as
ff(i) = $i(xili)5i(t) (25)
where $i(x l) is the matrix of the tether mode shapes defined as
0 0
^i(xiJi) =
^ X bull bull bull ^XA
0
0 0
0
G 5R 3 x n f
(26)
and nfj = nx- + nyi + nZi Note $J(XJJ) is a function of the spatial coordinate x
and not of yi or ZJ However it is also a function of the tether length parameter li
16
which is time-varying during deployment and retrieval For this reason care must be
exercised when differentiating $J(XJJ) with respect to time such that
d d d k) = Jtregi(Xi li) = ^ 7 $ t ( ^ i k)Xi + gf$i(xigt li)k- (2-7)
Since x represents the nominal rate of change of the position of the elemental mass
drrii it is equal to j (deploymentretrieval rate) Let t ing
( 3 d
dx~ + dT)^Xili^ ( 2 8 )
one arrives at
$i(xili) = $Di(xili)ii (29)
The time varying generalized coordinate vector 5 in Eq(25) is composed of
the longitudinal and transverse components ie
ltM) G s f t n f i x l (210)
where 5X^ 5y- and Sz^ are nx- x 1 ny x 1 and nz- x 1 size vectors and correspond to
$ X and $ Z respectively
213 Rotation angles and transformations
The matrix T j in Eq(21) represents a rotation transformation from the body
fixed frame Fj to the inertial frame FQ ie FQ = T j F j It is defined by the 3-2-1
ordered sequence of elementary rotations[46]
1 P i tch F[ = Cf ( a j )F 0
2 R o l l F = C f (3j)F = C7f (3j)Cf ( a j )F 0
3 Yaw F = C7( 7 i )F = C7(7l)cf^)Cf(ai)FQ = Q F 0
17
where
c f (A) =
-Sai Cai
0 0
and
C(7i ) =
0 -s0i] = 0 1 0
0 cpi
i 0 0 = 0
0
(211)
(212)
(213)
Here Eq(211) corresponds to a rotation of ctj (pitch) about the inertial axis ZQ (orbit
normal) of frame FQ resulting in a new intermediate frame F It is then followed by
a roll rotation fa about the axis y of F- giving a second intermediate frame F-
Final ly a spin rotation 7J about the axis z of F- is given resulting in the body
fixed frame Fj for the i lt l link Since all the three rotation matrices are orthogonal
it follows that FQ = C~1Fi = Cj F and hence
C CPi Cai ~ CH Sai + SrYi SPi Sai 5 7 i Sampi + Cli SPi Cci CPi Sai C1i Cai + 5 7 i SPi ~ S1 Ca + C 7 bull Sp Sai
(214) SliCPi CliCPi
Here Cx and Sx are abbreviations for cos(x) and sin(x) respectively
214 Inertial velocity of the ith link
mdash Lett ing pj = fj - f $j(5j and differentiating Eq(21) with respect to time gives
m- D j + T j f j + T^Si + T j pound + T j $ j 5 j (215)
Each of the terms appearing in Eq(215) must be addressed individually Since
fj represents the rigid body motion within the link it only has meaning for the
deployable tether and since yi = Zj = 0 for al l tether links (i even) fj = Ifi where 1
is the unit vector 1 0 0T For the case of rigid links (i odd) fj = 0
18
Evaluating the derivative of each angle in the rotation matrix gives
T^i9i = Taiai9i + Tppigi + T^fagi (216)
where Tx = J j I V Collecting the scaler angular velocity terms and defining the set
Vi = deg^gt A) 7i^ Eq(216) can be rewritten as
T i f t = Pi9i)fji (217)
where Pi(gi) is a column matrix defined by
Pi(9i)Vi = [Taigi^p^iT^gilffi (218)
Inserting Eq(29) and Eq(217) into Eq(215) and defining Sj = 2 + leads to
215 Cylindrical orbital coordinates
The Di term in Eq(219) describes the orbital motion of the t h l ink and
is composed of the three Cartesian coordinates DXi A ^ DZi However over the
cycle of one orbit DXi and Dyi vary dramatically by around the order of Earths
radius This must be avoided since large variations in the coordinates can cause
severe truncation errors during their numerical integration For this reason it is
more convenient to express the Cartesian components in terms of more stationary
variables This is readily accomplished using cylindrical coordinates
However it is only necessary to transform the first l inks orbital components
ie Lgti since only D is a generalized coordinates From Figure 2-2 it is apparent
(219)
that
(220)
19
Figure 2-2 Vector components in cylindrical orbital coordinates
20
Eq(220) can be rewritten in matrix form as
cos(0i) 0 0 s i n ( 0 i ) 0 0
0 0 1
Dri
h (221)
= DTlDSv
where Dri is the inplane radial distance 9 the true anomaly and DZl the out-of-
plane distance normal to the orbital plane The total derivative of D wi th respect
to time gives in column matrix form
1 1 1 1 J (222)
= D D l D 3 v
where [cos(^i) -Dnsm(8i) 0 ]
(223)
For all remaining links ie i = 2 N
cos(0i) - A - j S i n ^ i ) 0 sin(^i) D r i c o s ( 0 i ) 0
0 0 1
DTl = D D i = I d 3 i = 2N (224)
where is the n x n identity matrix and
2N (225)
216 Tether deployment and retrieval profile
The deployment of the tether is a critical maneuver for this class of systems
Cr i t i ca l deployment rates would cause the system to librate out of control Normally
a smooth sinusoidal profile is chosen (S-curve) However one would also like a long
tether extending to ten twenty or even a hundred kilometers in a reasonable time
say 3-4 orbits This is usually difficult or impossible to accomplish solely wi th one
S-curve profile Often a composite S-curve-Steady-S-curve profile is adopted Thus
21
the deployment scheme for the t h tether can be summarized as
k = Tr l - cos (^r(ti - to-)) 1 ltUlt h 2 r V A V
k = Vov tiilttiltt2i
k = IT j 1 - C deg S lkitl ^ ^ lt i lt ^
(226)
where to- tt are the ini t ia l and final times of the deployment or retrieval maneuver
respectively A t j = t - tQ mdash tt mdash t2- and the steady deployment velocity VQ is
calculated based on the continuity of l and l at the specified intermediate times t
and to For the case of
and
Ai = i(ti) - ^(t 0 -) = - ^ ( t 2 i ) (227)
2Ali + li(tf) -h(tQ)
V = lUli l^lL (228)
1 tf-tQ v y
A t = (229)
22 Kinetics and System Energy
221 Kinetic energy
The kinetic energy of the ith link is given by
Kei = lJm Rdrnfi^Ami (230)
Setting Eq(219) in a matrix-vector form gives the relation
4 ^ ^ PM) T i $ 4 TiSi]i[t (231)
22
where
m
k J
x l (232)
and nq = -nfj + 7 is the number of generalized coordinates in each link Inserting
Eq(231) into Eq(230) and integrating the kinetic energy for the t h l ink can be
rewritten as
1 T
(233)
where Mti is the links symmetric mass matrix
Mt =
m i D D T D D DDTPifgidmi) DDTTi J ^dm D n T Sidm rn
sym fP(9i)Pi(9i)dmi J P[g^T^idm J P ( pound ) T ^ m sym sym J ^f^idm f ^f^dmi sym sym sym J sfsidmi
pound Sfttiqxnq
(234)
and mi is the mass of the Ith link The kinetic energy for the entire system ie N
bodies can now be stated as
1 bull T bull 1 bull T Ke = J2 MtA = o laquo Mtqt 2 ^ -i
i=l (235)
where qt = qj^^ bull bull -qfNT a n d Mt is a block diagonal decoupled mass matrix
of the system with Mf- on its diagonal
(236) Mt =
0 0
0 Mt2
0
0 0
MH
0 0 0
0 0 0 bull MtN
222 Simplification for rigid links
The mass matrix given by Eq(234) simplifies considerably for the case of rigid
23
links and is given as
Mti
m D D l
T D D l DD^Piiffidmi) 0
sym 0
0
fP(fi)Pi(i)dmi 0 0 0
0 Id nfj
0 m~
i = 1 3 5 N
(237)
Moreover when mult iplying the matrices in the (11) block term of Eq(237) one
gets
D D l
1 D D l = 1 0 0 0 D n
2 0 (238) 0 0 1
which is a diagonal matrix However more importantly i t is independent of 6 and
hence is always non-singular ensuring that is always invertible
In the (12) block term of M ^ it is apparent that f fidmi is simply the
definition of the centre of mass of the t h rigid link and hence given by
To drrii mdash miff (239)
where fcmi is the position vector of the centre of mass of link i For the case of the
(22) block term a little more attention is required Expanding the integrand using
Eq(218) where ltj = fj gives
fj Ta Tafi fj T Q Tpfi f- T Q T 7 i f j
sym sym
Now since each of the terms of Eq(240) is a scaler or a 1 x 1 matrix it is therefore
equal to its trace ie sum of its diagonal elements thus
tr(ffTai
TTaifi) tr(ffTai
TTpfi) tr(ff Ta^T^fi)
fjT^Tpn rfTp^fi (240)
P(rj)Pj(fi) = syrri trifjTpTTpfi) t r ^ T ^ T ^
sym sym tr(ff T 7 T r 7 f j) J
(241)
Using two properties of the trace function[47] namely
tr(ABC) = tr(BCA) = trCAB) (242)
24
and
j tr(A) =tr(^J A^j (243)
where A B and C are real matrices it becomes quite clear that the (22) block term
of Eq(237) simplifies to
P(fi)Pii)dmi =
tr(Tai
TTaiImi) tTTai
TTpImi) t r (T Q r T 7 7 m ) trTaTTpImi) tr(Tp
TT7Jmi) sym sym sym
(244)
where Im- is an alternate form of the inertia tensor of the t h link which when
expanded is
Imlaquo = I nfi dm bull i sym
sym sym
J x^dmi J Xydmi XZdm sym J yi2dmi f yiZidm
sym J z^dmi J
xVi -L VH sym ijxxi + lyy IZZJ)2
(245)
Note this is in terms of moments of inertia and products of inertia of the t f l rigid
link A similar simplification can be made for the flexible case where T is replaced
with Qi resulting in a new expression for Im
223 Gravitational potential energy
The gravitational potential energy of the ith link due to the central force law
can be written as
V9i f dm f
-A mdash = -A bullgttradelti RdmA
dm
inn Lgti + T j ^ | (246)
where z = 3986 x 105 km 3s 2 (Earths gravitational constant) Expanding binomially
and retaining up to third order terms gives
Vgi ~ ~ D T 2D [J gfgidrrii + 23f T j gidm mdashDT
D2 i l I I Df Tt I grffdmiTiDi
(2-lt 7)
25
where D = D is the magnitude of the inertial position vector to the frame F The
integrals in the above expression are functions of the flexible mode shapes and can be
evaluated through symbolic manipulation using an existing package such as Maple
V In turn Maple V can translate the evaluated integrals into F O R T R A N and store
the code in a file Furthermore for the case of rigid bodies it can be shown quite
readily that
j gjgidrrii = j rjfidrrn = [IXXi + Iyy + IZZi)2 (248)
Similarly the other integrals in Eq(234) can be integrated symbolically
224 Strain energy
When deriving the elastic strain energy of a simple string in tension the
assumptions of high tension and low amplitude of transverse vibrations are generally
valid Consequently the first order approximation of the strain-displacement relation
is generally acceptable However for orbiting tethers in a weak gravitational-gradient
field neglecting higher order terms can result in poorly modelled tether dynamics
Geometric nonlinearities are responsible for the foreshortening effects present in the
tether These cannot be neglected because they account for the heaving motion along
the longitudinal axis due to lateral vibrations The magnitude of the longitudinal
oscillations diminish as the tether becomes shorter[35]
W i t h this as background the tether strain which is derived from the theory
of elastic vibrations[3248] is given by
(249)
where Uj V and W are the flexible deformations along the Xj and Z direction
respectively and are obtained from Eq(25) The square bracketed term in Eq(249)
represents the geometric nonlinearity and is accounted for in the analysis
26
The total strain energy of a flexible link is given by
Ve = I f CiEidVl (250)
Substituting the stress-strain relationship 0 = E(poundi the strain energy is now given
by
Vei=l-EiAi J l l d x h (251)
where E^ is the tethers Youngs modulus and A is the cross-sectional area The
tether is assumed to be uniform thus EA which is the flexural stiffness of the
tether is assumed to be constant Eq(251) is evaluated symbolically for arbitrary
number of modes
225 Tether energy dissipation
The evaluation of the energy dissipation due to tether deformations remains
a problem not well understood even today Here this complex phenomenon is repshy
resented through a simplified structural damping model[32] In addition the system
exhibits hysteresis which also must be considered This is accomplished using an
augmented complex Youngs modulus
EX^Ei+jEi (252)
where Ej^ is the contribution from the structural damping and j = The
augmented stress-strain relation is now given as
a = Epoundi = Ei(l+jrldi)ei (253)
where 77^ = EjE^ ltC 1 is defined as the structural damping coefficient determined
experimentally [49]
27
If poundi is a harmonic function with frequency UJQ- then
jet = (254) wo-
Substituting into Eq(253) and rearranging the terms gives
e + I mdash I poundi = o-i + ad (255)
where ad and a are the stress with and without structural damping respectively
Now the Rayleigh dissipation function[50] for the ith tether can be expressed as
(ii) Jo
_ 1 EjAiVdi rU (2-56^ 2 w 0 Jo
The strain rate ii is the time derivative of Eq(249) The generalized external force
due to damping can now be written as Qd = Q^ bull bull bull QdNT where
dWd
^ = - i p = 2 4 ( 2 5 7 )
= 0 i = 13 N
23 0(N) Form of the Equations of Mot ion
231 Lagrange equations of motion
W i t h the kinetic energy expression defined and the potential energy of the
whole system given by Pe = ]Cn=l( 5i+ ej) the equations of motion can be obtained
quite readily using the Lagrangian principle
where q is the set of generalized coordinates to be defined in a later section
Substituting Eqs(235247251 and 257) into Eq(258) leads to the familiar
matr ix form of the non-linear non-autonomous coupled equations of motion for the
28
system
Mq t)t+ F(q q t) = Q(q q t) (259)
where M(q t) is the nonlinear symmetric mass matrix and F(q q t) is the forcing
function which can be written in matrix form as
Q(ltfgt Qi 0 is the vector of non-conservative generalized forces including control inshy
puts acting on the system A detailed expansion of Eq(260) in tensor notation is
developed in Appendix I
Numerical solution of Eq(259) in terms of q would require inversion of the
mass matrix However M is a full matrix of size nqq x nqq where nqq = Nnq is the
total number of generalized coordinates Therefore direct inversion of M would lead
to a large number of computation steps of the order nqq^ or higher The objective
here is to develop an order-N O(N) algorithm for obtaining M _ 1 that minimizes
the computational effort This is accomplished through the following transformations
ini t ia l ly developed for the stationkeeping case by Pradhan et al [30]
232 Generalized coordinates and position transformation
During the derivation of the energy expressions for the system the focus was
on the decoupled system ie each link was considered independent of the others
Thus each links energy expression is also uncoupled However the constraints forces
between two links must be incorporated in the final equations of motion
Let
e Unqxl (261) m
be the set of coupled coordinates of the t h l ink such that q = qT q2 bull bull bull 0T i s the
29
systems generalized coordinates Let qt- be the set of auxiliary decoupled coordinates
defined by Eq(232) such that qt = qj^ qj^ Qtj^1 bull The only difference between
qtj and q is the presence of D against d is the position of F wi th respect to
the inertial frame FQ whereas d is defined as the offset position of F relative to and
projected on the frame It can be expressed as d = dXidyidZi^ For the
special case of link 1 D = d
From Figure 2-1 can written as
Di = A - l + T i _ i J i _ i + di) + T V i S i - i f t - i + dXi)8i- (262)
Denoting KAj = $j(Zj + dx- ) Eq(262) can be rewritten in summation form as
i ampi = H ( T i - 1 ^ + T i - l K A i - l - l + T j - l ^ - l ) bull ( 2 - 6 3 )
3 = 1
Introducing the index substitution k mdash j mdash 1 into Eq(263)
i-1 D = ^2 (Tfc_ilt4 + TkKAk6k + Tkilk) + T^di (264)
k=l
since d = D TQ = 1^ and IQ DQ SQ are null vectors Recasting Eq(264) in
matr ix form leads to i-1
Qt Jfe=l
where
and
For the entire system
T fc-1 0 TkKAk
0 0 0 0 0 0 0 0 0 0 0 0 -
r T i - i 0 0 0 0 0 0 0 0 0 G
0 0 0 1
Qt = RpQ
G R n lt x l
(265)
(266)
G Unq x l (267)
(268)
30
where
RP
R 1 0 0 R R 2 0 R RP
2 R 3
0 0 0
R RP R PN _ 1 RN J
(269)
Here RP is the lower block triangular matrix which relates qt and q
233 Velocity transformations
The two sets of velocity coordinates qti and qi are related by the following
transformations
(i) First Transformation
Differentiating Eq(262) with respect to time gives the inertial velocity of the
t h l ink as
D = A - l + Ti_iK + + K A i _ i ^ _ i
+ T j _ i ^ + Ti-ii + T i - i K A i - i ^ + T j - i K A ^ i ^ i (270)
(271) (
Defining K A D = dddXi+1KAi K A L = d d K A and hi = di+1 +lfi +KA^i
results in K A ^ i = K A D j i o ^ + K A L i _ i i _ i
= KADi_iiTdi + K A L j _ i ^ _ i
Inserting Eq(271) into Eq(270) and rearranging gives
= A _ + T i ^ d g + K A D ^ x ^ i ^ ^ + Pi-^hi-iH-i
+ T i _ i K A i _ i 5 _ i + T i_ i2 + K A L j _ i 5 j _ i 4 _ i (272)
Using Eq(272) to define recursively - D j _ i and applying the index substitution
c = j mdash 1 as in the case of the position transformation it follows that
Di = ^ T ^ i K D ^ ^ + Pk(hk)ffk + T f c K A f c ^ + TfcKLfc4 Jfc=l (273)
+ T j _ i K D j _ i lt i ii
31
where K D j _ i = + K A D ^ i ^ - i r 7 and K L j = i + K A L j 5 j Finally by following a
procedure similar to that used for the position transformation it can be shown that
for the entire system
t = Rvil (2-74)
where Rv is given by the lower block diagonal matrix
Here
Rv
Rf 0 0 R R$ 0 R Rl RJ
0 0 0
T3V nN-l
R bullN
e Mnqqxl
Ri =
T i _ i K D i _ i TiKAi T j K L j 0 0 0 0 bullo 0 0 0 0 0 0 0
e K n lt x l
RS
T i-1 0 0 o-0 0 0 0 0 0
0 0 0 1
G Mnq x l
(275)
(276)
(277)
(ii) Second Transformation
The second transformation is simply Eq(272) set into matrix form This leads
to the expression for qt as
(278)
where Id Pihi) T j K A j T j K L j 0 0 0 0 0 0
0 0
0 0
0 0
G $lnq x l
Thus for the entire system
(279)
pound = Rnjt + Rdq
= Vdnqq - R nrlRdi (280)
32
where
and
- 0 0 0 RI 0 0
Rn = 0 0
0 0
Rf 0 0 0 rgtd K2 0
Rd = 0 0 Rl
0 0 0
RN-1 0
G 5R n w x l
0 0 0 e 5 R N ^ X L
(281)
(282)
234 Cylindrical coordinate modification
Because of the use of cylindrical coorfmates for the first body it is necessary
to slightly modify the above matrices to account for this coordinate system For the
first link d mdash D = D ^ D S L thus
Qti = 01 =gt DDTigi (283)
where
D D T i
Eq(269) now takes the form
DTL 0 0 0 d 3 0 0 0 ID
0 0 nfj 0
RP =
P f D T T i R2 0 P f D T T i Rp
2 R3
iJfDTTi i^
x l
0 0 0
R PN-I RN
(284)
(285)
In the velocity transformation a similar modification is necessary since D =
DD^DS^ Mak ing the substitution it can be shown that Eq(276) and Eq(279) now
33
takes the form for i = 1
T)V _ pn _ r t j mdash rt^ mdash 0 0 0 0
0 0 0 0 0 0 0 0
(286)
The rest of the matrices for the remaining links ie i = 2 N are unchanged
235 Mass matrix inversion
Returning to the expression for kinetic energy substitution of Eq(274) into
Eq(235) gives the first of two expressions for kinetic energy as
Ke = -q RvT MtRv (287)
Introduction of Eq(280) into Eq(235) results in the second expression for Ke as
1X 2q (hnqq-RnrlRd Mt (idnqq - R nrlRd (288)
Thus there are two factorizations for the systems coupled mass matrix given as
M = RV MtRv
M (hnqq - RnrlRd Mt ( i d n q q - R n r 1 R d
(289)
(290)
Inverting Eq(290)
r-1 _ ( rgtd~^ M 1 = (Ra) l d n q q - Rn)Mf1 (Rd)~l(Idnqq - Rn) (291)
Since both Rd and Mt are block diagonal matrices their inverse is simply the inverse
34
of each block on the diagonal ie
Mr1
h 0 0
0 0
0 0
0 0 0
0
0
0
tN
(292)
Each block has the dimension nq x nq and is independent of the number of bodies
in the system thus the inversion of M is only linearly dependent on N ie an O(N)
operation
236 Specification of the offset position
The derivation of the equations of motion using the O(N) formulation requires
that the offset coordinate d be treated as a generalized coordinate However this
offset may be constrained or controlled to follow a prescribed motion dictated either by
the designer or the controller This is achieved through the introduction of Lagrange
multipliers[51] To begin with the multipliers are assigned to all the d equations
Thus letting f = F - Q Eq(259) becomes
Mq + f = P C A (293)
where A G K 3 - 7 V x l is the vector of Lagrange multipliers and Pc G s R n 9 x 3 A r is the
permutation matrix assigning the appropriate Aj to its corresponding d equation
Inverting M and pre-multiplying both sides by PcT gives
pcTM-pc
= ( dc + PdegTM-^f) (294)
where dc is the desired offset acceleration vector Note both dc and PcTM lf are
known Thus the solution for A has the form
A pcTM-lpc - l dc + PC1 M (295)
35
Now substituting Eq(295) into Eq(293) and rearranging the terms gives
q = S(f Ml) = -M~lf + M~lPc
which is the new constrained vector equation of motion with the offset specified by
dc Note that pre-multiplying M~l by P c T provides the rows of M _ 1 corresponding
to the d equations Similarly post-multiplying by Pc leads the columns of the matrix
corresponding to the d equations
24 Generalized Control Forces
241 Preliminary remarks
The treatment of non-conservative external forces acting on the system is
considered next In reality the system is subjected to a wide variety of environmental
forces including atmospheric drag solar radiation pressure Earth s magnetic field as
well as dissipative devices to mention a few However in the present model only
the effect of active control thrusters and momemtum-wheels is considered These
generalized forces wi l l be used to control the attitude motion of the system
In the derivation of the equations of motion using the Lagrangian procedure
the contribution of each force must be properly distributed among all the generalized
coordinates This is acheived using the relation
nu ^5 Qk=E Pern ^ (2-97)
m=l deg Q k
where Qk and qk are the kth elements of Q and ltf respectively[51] Fem is the
mth external force acting on the system at the location Rm from the inertial frame
and nu is the total number of actuators Derivation of the generalized forces due to
thrusters and momentum wheels is given in the next two subsections
pcTM-lpc (dc + PcTM- Vj (2-96)
36
242 Generalized thruster forces
One set of thrusters is placed on each subsatellite ie al l the rigid bodies
except the first (platform) as shown in Figure 2-3 They are capable of firing in the
three orthogonal directions From the schematic diagram the inertial position of the
thruster located on body 3 is given as
Defining
R3 = di + Tid2 + T2(hi + KA2lt52) + T 3 f c m 3
= d + dfY + T 3 f c m 3
df = Tjdj+i + Tj+1lj+ii + KAj+i5j+i)
the thruster position for body 5 is given as
(298)
(299)
R5 = di+ dfx + d 3 + T5fcm
Thus in general the inertial position for the thruster on the t h body can be given
(2100)
as i - 2
k=l i1 cm (2101)
Let the column matrix
[Q i iT 9 R-K=rR dqj lA n i
WjRi A R i (2102)
where i = 3 5 N and j mdash 1 2 i The vector derivative of R wi th respect to
the scaler components qk is stored in the kth column of Thus for the case of
three bodies Eq(297) for thrusters becomes
1 Q
Q3
3 -1
T 3 T t 2 (2103)
37
Satellite 3
mdash gt -raquo cm
D1=d1
mdash gt
a mdash gt
o
Figure 2-3 Inertial position of subsatellite thruster forces
38
where ft = 0TtaTt^T is the thrust vector for the t h l ink (tether i-1) Here
Ttn and Tta are the thrust forces along the inplane m and out-of-plane Z directions i Pi
respectively For the case of 5-bodies
Q
~Qgt1
0
0
( T ^ 2 ) (2104)
The result for seven or more bodies follows the pattern established by Eq(2104)
243 Generalized momentum gyro torques
When deriving the generalized moments arising from the momemtum-wheels
on each rigid body (satellite) including the platform it is easier to view the torques as
coupled forces From Figure 2-4 it is apparent that for the ith l ink the generalized
forces constituting the couple are
(2105)
where
Fei = Faij acting at ei = eXi
Fe2 = -Fpk acting at e 2 = C a ^ l
F e 3 = F7ik acting at e 3 = eyj
Expanding Eq(2105) it becomes clear that
3
Qk = ^ F e i
i=l
d
dqk (2106)
which is independent of Ri- Thus the moments Mma = 2FaieXi Mm^ = 2FpeXi
and Mmi = 2FlieVi Transforming the coordinates to the body fixed frame using
the rotation matrix the generalized force due to the momentum-wheels on link i can
39
Figure 2-4 Coupled force representation of a momentum-wheel on a rigid body
40
be written as
d d d Qi = Td bull ^ T i t M m a - Tik bull mdashT^Mmp + T ^ bull T J j M ^
QiMmi
where Mm = Mma M m Mmi T
lt3 T J ^ T - T i f c bull ^ T i pound T ^ - ^ T z j
(2107)
(2108)
Note combining the thruster forces and momentum-wheel torques a compact
expression for the generalized force vector for 5 bodies can be written as
Q
Q T 3 0 Q T 5 0
0 0 0
0 lt33T3 Qyen Q5
3T5 0
0 0 0 Q4T5 0
0 0 0 Q5
5T5
M M I
Ti 2 m 3
Quu (2109)
The pattern of Eq(2109) is retained for the case of N links
41
3 C O M P U T E R I M P L E M E N T A T I O N
31 Prel iminary Remarks
The governing equations of motion for the TV-body tethered system were deshy
rived in the last chapter using an efficient O(N) factorization method These equashy
tions are highly nonlinear nonautonomous and coupled In order to predict and
appreciate the character of the system response under a wide variety of disturbances
it is necessary to solve this system of differential equations Although the closed-
form solutions to various simplified models can be obtained using well-known analytic
methods the complete nonlinear response of the coupled system can only be obtained
numerically
However the numerical solution of these complicated equations is not a straightshy
forward task To begin with the equations are stiff[52] be they have attitude and
vibration response frequencies separated by about an order of magnitude or more
which makes their numerical integration suceptible to accuracy errors i f a properly
designed integration routine is not used
Secondly the factorization algorithm that ensures efficient inversion of the
mass matrix as well as the time-varying offset and tether length significantly increase
the size of the resulting simulation code to well over 10000 lines This raises computer
memory issues which must be addressed in order to properly design a single program
that is capable of simulating a wide range of configurations and mission profiles It
is further desired that the program be modular in character to accomodate design
variations with ease and efficiency
This chapter begins with a discussion on issues of numerical and symbolic in-
42
tegration This is followed by an introduction to the program structure Final ly the
validity of the computer program is established using two methods namely the conshy
servation of total energy and comparison of the response results for simple particular
cases reported in the literature
32 Numerical Implementation
321 Integration routine
A computer program coded in F O R T R A N was written to integrate the equashy
tions of motion of the system The widely available routine used to integrate the difshy
ferential equations D G E A R [ 5 3 ] is based on Gears predictor-corrector method[54]
It is well suited to stiff differential equations as i t provides automatic step-size adjustshy
ment and error-checking capabilities at each iteration cycle These features provide
superior performance over traditional 4^ order Runge-Kut ta methods
Like most other routines the method requires that the differential equations be
transformed into a first order state space representation This is easily accomplished
by letting
thus
X - ^ ^ f ^ ) =ltbullbull) (3 -2 )
where S(fMdc) is defined by Eq(296) Eq(32) represents the new set of 2nqq
first order differential equations of motion
322 Program structure
A flow chart representing the computer programs structure is presented in
Figure 3-1 The M A I N program is responsible for calling the integration routine It
also coordinates the flow of information throughout the program The first subroutine
43
called is INIT which initializes the system variables such as the constant parameters
(mass density inertia etc) and the ini t ial conditions I C (attitude angles orbital
motion tether elastic deformation etc) from user-defined data files Furthermore all
the necessary parameters required by the integration subroutine D G E A R (step-size
tolerance) are provided by INIT The results of the simulation ie time history of the
generalized coordinates are then passed on to the O U T P U T subroutine which sorts
the information into different output files to be subsequently read in by the auxiliary
plott ing package
The majority of the simulation running time is spent in the F C N subroutine
which is called repeatedly by D G E A R F C N generates the fv(Xi) vector defined in
Eq(32) It is composed of several subroutines which are responsible for the comshy
putation of all the time-varying matrices vectors and scaler variables derived in
Chapter 2 that comprise the governing equations of motion for the entire system
These subroutines have been carefully constructed to ensure maximum efficiency in
computational performance as well as memory allocation
Two important aspects in the programming of F C N should be outlined at this
point As mentioned earlier some of the integrals defined in Chapter 2 are notably
more difficult to evaluate manually Traditionally they have been integrated numershy
ically on-line (during the simulation) In some cases particularly stationkeeping the
computational cost is quite reasonable since the modal integrals need only be evalushy
ated once However for the case of deployment or retrieval where the mode shape
functions vary they have to be integrated at each time-step This has considerable
repercussions on the total simulation time Thus in this program the integrals are
evaluated symbolically using M A P L E V[55] Once generated they are translated into
F O R T R A N source code and stored in I N C L U D E files to be read by the compiler
Al though these files can be lengthy (1000 lines or more) for a large number of flexible
44
INIT
DGEAR PARAM
IC amp SYS PARAM
MAIN
lt
DGEAR
lt
FCN
OUTPUT
STATE ACTUATOR FORCES
MOMENTS amp THRUST FORCES
VIBRATION
r OFFS DYNA
ET MICS
Figure 3-1 Flowchart showing the computer program structure
45
modes they st i l l require less time to compute compared to their on-line evaluation
For the case of deployment the performance improvement was found be as high as
10 times for certain cases
Secondly the O(N) algorithm presented in the last section involves matrices
that contained mostly zero terms (RP Rv and Mplusmn) The sparse structure of these mashy
trices can be used advantageously to improve the computational efficiency by avoiding
multiplications of zero elements The result is a quick evaluation of fv(Xt) Moreshy
over there is a considerable saving of computer memory as zero elements are no
longer stored
Final ly the C O N T R O L subroutine which is part of F C N is designed to calshy
culate the appropriate control forces torques and offset dynamics which are required
to stabilize the librational ( A T T I T U D E ) and vibrational ( V I B R A T I O N ) motion of
the system
33 Verification of the Code
The simulation program was checked for its validity using two methods The
first verified the conservation of total energy in the absence of dissipation The second
approach was a direct comparison of the planar response generated by the simulation
program with those available in the literature A s numerical results for the flexible
three-dimensional A^-body model are not available one is forced to be content with
the validation using a few simplified cases
331 Energy conservation
The configuration considered here is that of a 3-body tethered system with
the following parameters for the platform subsatellite and tether
46
h =
1091430 -8135 328108 -8135 8646050 27116 328108 27116 8286760
kg bull m 2 platform inertia
kg bull m 2 end-satellite inertia 200 0 0
0 400 0 0 0 400
m = 90000 kg mass of the space station platform
m2 = 500 kg mass of the end-satellite
EfAt = 61645 N tether elastic stiffness
pt mdash 49 kg km tether linear density
The tether length is taken to remain constant at 10 km (stationkeeping case)
A n in i t ia l disturbance in pitch and roll of 2deg and 1deg respectively is given to both
the rigid bodies and the flexible tether In addition the tether is ini t ia l ly deflected
by 45 m in the longitudinal direction at its end and 05 m in both the inplane and
out-of-plane directions in the first mode The tether attachment point offset at the
platform (satellite 1) end is 1 m in all the three directions ie d2 = 11 lTm The
attachment point at the subsatellite (satellite 2) end is 1 m in the x direction ie
fcm3 = 10 0T
The variation in the kinetic and potential energies is presented for the above-
mentioned case in Figure 3-2(a) As seen from the plot there is a continuous mutual
exchange between the kinetic and potential energies however the variation in the
total energy remains zero (Figure 3-2b) It should be noted that after 1 orbit the
variation of both kinetic and potential energy does not return to 0 This is due to
the attitude and elastic motion of the system which shifts the centre of mass of the
tethered system in a non-Keplerian orbit
Similar results are presented for the 5-body case ie a double pendulum (Figshy
ure 3-3) Here the system is composed of a platform and two subsatellites connected
47
STSS configuration 3-Body a 1(0)=a 2(0)=a t(0)=2deg P1(0)=p2(0)=p(0)=1deg 5x(0)=45m 8y(0)=52(0)=05m dx(0)=dy(0)=d2(0)=1m Stationkeeping lt=10km
Variation of Kinetic and Potential Energy
(a) Time (Orbits) Percent Variation in Total Energy
OOEOh
-50E-12
-10E-11 I i i
0 1 2 (b) Time (Orbits)
Figure 3-2 Kinetic and potential energy transfer for the three-body platform based tethered satellite system (a) variation of kinetic and potential energy (b) percent change in total energy of system
48
STSS configuration 5-Body a 1(0)=a 2(0)=a t 1(0)=2deg a 3(0)=a t 2(0)=25deg P1(0)=p2(0)=p3(0)=pt1(0)=pt2(0)=1lt
5x(0)=5m 8y(0)=5z(0)=05m dx(0)=dy(0)=dz(0)=0 Stationkeeping 1 =12=1 Okm
(a)
20E6
10E6r-
00E0
-10E6
-20E6
Variation of Kinetic and Potential Energy
-
1 I I I 1 I I
I I
AP e ~_
i i 1 i i i i
00E0
UJ -50E-12 U J
lt
Time (Orbits) Percent Variation in Total Energy
-10E-11h
(b) Time (Orbits)
Figure 3 -3 Kinet ic and potential energy transfer for the five-body platform based tethered satellite system (a) variation of kinetic and potential enshyergy (b) percent change in total energy of system
49
in sequence by two tethers The two subsatellites have the same mass and inertia
ie 7713 = 7 7 i 2 a n d I3 = I21 a n d are separated by identical tether segments each 10
k m in length The platform has the same properties as before however there is no
tether attachment point offset present in the system (d = fcmi = 0)
In al l the cases considered energy was found to be conserved However it
should be noted that for the cases where the tether structural damping deployment
retrieval attitude and vibration control are present energy is no longer conserved
since al l these features add or subtract energy from the system
332 Comparison with available data
The alternative approach used to validate the numerical program involves
comparison of system responses with those presented in the literature A sample case
considered here is the planar system whose parameters are the same as those given
in Section 331 A n ini t ia l pitch disturbance of 2deg is imparted to both the platform
and the tether together with an ini t ial tether deformation of 08 m and 001 m in
the longitudinal and transverse directions respectively in the same manner as before
The tether length is held constant at 5 km and it is assumed to be connected to the
centre of mass of the rigid platform and subsatellite The response time histories from
Ref[43] are compared with those obtained using the present program in Figures 3-4
and 3-5 Here ap and at represent the platform and tether pitch respectively and B i
C i are the tethers longitudinal and transverse generalized coordinates respectively
A comparison of Figures 3-4 and 3-5 clearly shows a remarkable correlation
between the two sets of results in both the amplitude and frequency The same
trend persisted even with several other comparisons Thus a considerable level of
confidence is provided in the simulation program as a tool to explore the dynamics
and control of flexible multibody tethered systems
50
a p(0) = 2deg 0(0) = 2C
6(0) = 08 m
0(0) = 001 m
Stationkeeping L = 5 km
D p y = 0 D p 2 = 0
a oo
c -ooo -0 01
Figure 3 - 4 Simulation results for the platform based three-body tethered system originally presented in Ref[43]
51
STSS configuration 3-Body a1(0)=a2(0)=at(0)=2deg p1(0)=(32(0)=(3t(0)=0 8x(0)=08m 5y(0)=001m 5Z(0)=0 dx(0)=dy(0)=dz(0)=0 Stationkeeping l=5km
Satellite 1 Pitch Angle Tether Pitch Angle
i -i i- i i i i i i d
1 2 0 1 2 Time (Orbits) Time (Orbits)
Longitudinal Vibration
^MAAAAAAAWVWWWWVVVW
V bull i bull i I i J 00 05 10 15 20
Time (Orbits) Transverse Vibration
Time (Orbits)
Figure 3-5 Simulation results for the platform based three-body tethered system obtained using the present computer program
52
4 D Y N A M I C SIMULATION
41 Prel iminary Remarks
Understanding the dynamics of a system is critical to its design and develshy
opment for engineering application Due to obvious limitations imposed by flight
tests and simulation of environmental effects in ground based facilities space based
systems are routinely designed through the use of numerical models This Chapter
studies the dynamical response of two different tethered systems during deployment
retrieval and stationkeeping phases In the first case the Space platform based Tethshy
ered Satellite System (STSS) which consists of a large mass (platform) connected to
a relatively smaller mass (subsatellite) with a long flexible tether is considered The
other system is the O E D I P U S B I C E P S configuration involving two comparable mass
satellites interconnected by a flexible tether In addition the mission requirement of
spin about an arbitrary axis the tether axis for O E D I P U S - A C and cartwheeling or
spin about the orbit normal for the proposed B I C E P S mission is accounted for
A s mentioned in Chapter 2 the tether flexibility is modelled using the asshy
sumed mode discretization method Although the simulation program can account
for an arbitrary number of vibrational modes in each direction only the first mode
is considered in this study as it accounts for most of the strain energy[34] and hence
dominates the vibratory motion The parametric study considers single as well as
double pendulum type systems with offset of the tether attachment points The
systems stability is also discussed
42 Parameter and Response Variable Definitions
The system parameters used in the simulation unless otherwise stated are
53
selected as follows for the STSS configuration
1091430 -8135 328108 -8135 8646050 27116 328108 27116 8286760
h =
h = 200 0 0
0 400 0 0 0 400
bull m = 90000 kg (mass of the space platform)
kg bull m 2 (platform inertia) 760 J
kg bull m 2 (subsatellite inertia)
bull 7772 = 500 kg (mass of the subsatellite)
bull EtAt = 61645 N (tether stiffness)
bull Pt = 49 k g k m (tether density)
bull Vd mdash 05 (tether structural damping coefficient)
bull fcm^ = l 0 0 r m (tether attachment point at the subsatellite)
The response variables are defined as follows
bull ai0 satellite 1 (platform) pitch and roll angles respectively
bull CK2 2 satellite 2 (subsatellite) pitch and roll angles respectively
bull at fit- tether pitch and roll angles respectively
bull If tether length
bull d = dx dy dzT- tether attachment position relative to satellite 1 (platform)
mdash
bull S = 8X 8y Sz tether flexible generalized coordinates in the longitudinal x
inplane transverse y and out-of-plane transverse z directions respectively
The attitude angles and are measured with respect to the Local Ve r t i c a l -
Loca l Horizontal ( L V L H ) frame A schematic diagram illustrating these variable
is presented in Figure 4-1 The system is taken to be in a nominal circular orbit at
an altitude of 289 km with an orbital period of 903 minutes
54
Satellite 1 (Platform) Y a w
Figure 4-1 Schematic diagram showing the generalized coordinates used to deshyscribe the system dynamics
55
43 Stationkeeping Profile
To facilitate comparison of the simulation results a reference case based on
the three-body STSS system with zero offset (d = 0 r c m 3 = 0) and a tether length
of It = 20 km is considered first (Figure 4-2) The system is subjected to an ini t ial
disturbance in pitch and roll of 2deg and 1deg respectively to all the three bodies
In addition an ini t ia l longitudinal deflection of 14 m from the tethers unstretched
position is given together with a transverse inplane and out-of-plane deflection of 1
m at the tethers mid-length From the attitude response given in Figure 4-2(a) it is
apparent that the three bodies oscillate about their respective equilibrium positions
Since coupling between the individual rigid-body dynamics is absent (zero offset) the
corresponding librational frequencies are unaffected Satellite 1 (Platform) displays
amplitude modulation arising from the products of inertia The result is a slight
increase of amplitude in the pitch direction and a significantly larger decrease in
amplitude in the roll angle
Figure 4-2(b) clearly shows coupling of the tethers attitude motion with its
flexibility dynamics The tether vibrates in the axial direction about its equilibrium
position (at approximately 138 m) with two vibration frequencies namely 012 Hz
and 31 x 1 0 - 4 Hz The former is the first longitudinal flexible mode which is dissishy
pated through the structural damping while the latter arises from the coupling with
the pitch motion of the tether Similarly in the transverse direction there are two
frequencies of oscillations arising from the flexible mode and its coupling wi th the
attitude motion As expected there is no apparent dissipation in the transverse flexshy
ible motion This is attributed to the weak coupling between the longitudinal and
transverse modes of vibration since the transverse strain in only a second order effect
relative to the longitudinal strain where the dissipation mechanism is in effect Thus
there is very litt le transfer of energy between the transverse and longitudinal modes
resulting in a very long decay time for the transverse vibrations
56
STSS configuration 3-Body a1(0)=a2(0)=a(0)=2deg (31(0)=p2(0)=p(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=dz(0)=0 Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
0 1 2 3 4 5 0 1 2 3 4 5 Time (Orbits) Time (Orbits)
Satellite 2 Pitch Angle Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-2 Stationkeeping dynamics of the three-body STSS configuration withshyout offset (a) attitude response
57
STSS configuration 3-Body a1(0)=cx2(0)=cx(0)=2deg P1(0)=p2(0)=p(0)=1deg 5x(0)=14m8y(0)=82(0)=1m dx(0)=dy(0)=dz(0)=0 Stationkeeping lt=20km
Tether Longitudinal Vibration 16| 1 1 1 1 1 r-
Time (Orbits) Tether In-Plane Transverse Vibration
imdash 1 mdash 1 mdash 1 mdash 1 mdashr
r I I 1 1 1 I 1 1 1 1 I 1 bull ^ bull bull I I J 0 1 2 3 4 5
Time (Orbits) Tether Out-of-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-2 Stationkeeping dynamics of the three-body STSS configuration withshyout offset (b) vibration response
58
Introduction of the attachment point offset significantly alters the response of
the attitude motion for the rigid bodies W i t h a i m offset along the local vertical
at the platform (dx = 1 m) and fcm^ = 100 m coupling between the tether
and end-bodies is established by providing a lever arm from which the tether is able
to exert a torque that affects the rotational motion of the rigid end-bodies (Figure
4-3) Bo th the platform and subsatellite now oscillate at the same frequency as the
tether whose motion remains unaltered In the case of the platform there is also an
amplitude modulation due to its non-zero products of inertia as explained before
However the elastic vibration response of the tether remains essentially unaffected
by the coupling
Providing an offset along the local horizontal direction (dy = 1 m) results in
a more dramatic effect on the pitch and roll response of the platform as shown in
Figure 4-4(a) Now the platform oscillates about its new equilibrium position of
-90deg The roll motion is also significantly disturbed resulting in an increase to over
10deg in amplitude On the other hand the rigid body dynamics of the tether and the
subsatellite as well as the tether flexible motion remain the same (Figure 4-4b)
Figure 4-5 presents the response when the offset of 1 m at the platform end is
along the z direction ie normal to the orbital plane The result is a larger amplitude
pitch oscillation about the reference equilibrium position as shown in Figure 4-5(a)
while the roll equilibrium is now shifted to approximately 90deg Note there is little
change in the attitude motion of the tether and end-satellites however there is a
noticeable change in the out-of-plane transverse vibration of the tether (Figure 4-5b)
which may be due to the large amplitude platform roll dynamics
Final ly by setting a i m offset in all the three direction simultaneously the
equil ibrium position of the platform in the pitch and roll angle is altered by approxishy
mately 30deg (Figure 4-6a) However the response of the system follows essentially the
59
STSS configuration 3-Body a1(0)=cx2(0)=cxt(0)=2deg p1(0)=(32(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=1m dy(0)=dz(0)=0 Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
Time (Orbits) Time (Orbits) Tether Pitch Angle Tether Roll Angle
L bull I I bull bull lt bull bull bull raquo bull bull bull bull -I h I I I I i H 0 1 2 3 4 5 0 1 2 3 4 5
Time (Orbits) Time (Orbits) Satellite 2 Pitch Angle Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-3 Stationkeeping dynamics of the three-body STSS configuration with offset along the local vertical (a) attitude response
60
STSS configuration 3-Body a 1(0)=a 2(0)=a t(0)=2deg (31(0)=p2(0)=pt(0)=1deg 5x(0)=14m5y(0)=5z(0)=1m dx(0)=1mdy(0)=dz(0)=0 Stationkeeping lt=20km
Tether Longitudinal Vibration 1 6 i mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash mdash i mdash i mdash mdash mdash mdash mdash i mdash mdash lt -
Time (Orbits)
Tether In-Plane Transverse Vibration
y bull i i i i i i i i i i i i i i i i i 0 1 2 3 4 5
Time (Orbits)
Tether Out-of-Plane Transverse Vibration r 1 1 1 1 1 1 1 1 1 1 1 1 1 1 q
h i I i i I i i i i I i i i i 3
0 1 2 3 4 5 (b) Time (Orbits)
Figure 4-3 Stationkeeping dynamics of the three-body STSS configuration with offset along the local vertical (b) vibration response
61
STSS configuration 3-Body a1(0)=cx2(0)=o(0)=2deg (31(0)=p2(0)=p(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dz(0)=0 dy(0)=1m Stationkeeping lt=20km
Satellite 1 Pitch Angle i 1 i
Satellite 1 Roll Angle
1 2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 4 Time (Orbits)
Tether Roll Angle
CD
edT
1 2 3 Time (Orbits)
Satellite 2 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 2 Roll Angle
CD
C O
(a)
Figure 4-
1 2 3 4 Time (Orbits)
1 2 3 4 Time (Orbits)
4 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal (a) attitude response
62
STSS configuration 3-Body a1(0)=cx2(0)=ot(0)=2deg p1(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dz(0)=0 dy(0)=1m Stationkeeping lt=20km
Tether Longitudinal Vibration 16| 1 gt 1 bull 1 1 1
0 1 2 3 4 5
Time (Orbits) Tether In-Plane Transverse Vibration
i mdash 1 mdash 1 mdash 1 mdash 1 mdash i mdash mdash lt mdash 1 mdash 1 mdash i mdash mdash 1 mdash 1 mdash
V I I bull bull bull I i i i I i i 1 1 3
0 1 2 3 4 5 Time (Orbits)
Tether Out-of-Plane Transverse Vibration r 1 1 1 1 1 1 1 1 1 1 1 gt bull | q
V i 1 1 1 1 1 mdash 1 mdash 3
0 1 2 3 4 5 (b) Time (Orbits)
Figure 4-4 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal (b) vibration response
63
STSS configuration 3-Body a1(0)=o2(0)=at(0)=2deg (31(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=0 dz(0)=1m Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 4 Time (Orbits)
Tether Roll Angle
1 2 3 4 Time (Orbits)
Satellite 2 Pitch Angle
2 3 4 Time (Orbits)
Satellite 2 Roll Angle
(a)
Figure 4-
1 2 3 4 Time (Orbits)
2 3 4 Time (Orbits)
5 Stationkeeping dynamics of the three-body STSS configuration with offset along the orbit normal (a) attitude response
64
STSS configuration 3-Body a1(0)=cx2(0)=ot(0)=2deg P1(0)=p2(0)=(3(0)=1deg 8x(0)=14m8y(0)=82(0)=1m dx(0)=dy(0)=0 d2(0)=1m Stationkeeping lt=20km
Tether Longitudinal Vibration 1 6 i mdash r mdash i mdash mdash i mdash i mdash mdash i mdash i mdash mdash i mdash mdash i mdash 1 mdash 1 mdash i mdash mdash lt mdash lt mdash lt mdash r
Time (Orbits) Tether In-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-5 Stationkeeping dynamics of the three-body STSS configuration with offset along the orbit normal (b) vibration response
65
same trend with only minor perturbations in the transverse elastic vibratory response
of the tether (Figure 4-6b)
44 Tether Deployment
Deployment of the tether from an ini t ial length of 200 m to 20 km is explored
next Here the deployment length profile is critical to ensure the systems stability
It is not desirable to deploy the tether too quickly since that can render the tether
slack Hence the deployment strategy detailed in Section 216 is adopted The tether
is deployed over 35 orbits with the sinusoidal acceleration and deceleration occurring
over a 4 km length ( A ^ 2 ) with the remaining deployment at a constant velocity of
V0 = 146 ms
Initially the tether is stretched only slightly S = 1304 x 1 0 _ 3 0 0 m
with al l the other states of the system ie pitch roll and transverse displacements
remaining zero The response for the zero offset case is presented in Figure 4-7 For
the platform there is no longer any coupling with the tether hence it oscillates about
the equilibrium orientation determined by its inertia matrix However there is st i l l
coupling between the subsatellite and the tether since rcm^ mdash 100 m resulting
in complete domination of the subsatellite dynamics by the tether Consequently
the pitch motion ini t ial ly grows by almost 50deg but eventually subsides as the tether
deployment rate decreases It is interesting to note that the out-of-plane motion is not
affected by the deployment This is because the Coriolis force which is responsible
for the tether pitch motion does not have a component in the z direction (as it does
in the y direction) since it is always perpendicular to the orbital rotation (Q) and
the deployment rate (It) ie Q x It
Similarly there is an increase in the amplitude of the transverse elastic oscilshy
lations of the tether again due to the Coriolis effect (Figure 4-7b) Moreover as the
66
STSS configuration 3-Body a1(0)=a2(0)=cxt(0)=2deg P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m5y(0)=52(0)=1m dx(0)=dy(0)=d2(0)=1m Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
i i i i i i i i i i i 0 1 2 3 4 5 0 1 2 3 4 5
Time (Orbits) Time (Orbits) Tether Pitch Angle Tether Roll Angle
l i i i i -I h i i i i I i i i i l- i i i i l i i i i I i i- i i H 0 1 2 3 4 5 0 1 2 3 4 5
Time (Orbits) Time (Orbits) Satellite 2 Pitch Angle Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-6 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal local vertical and orbit normal (a) attitude response
67
STSS configuration 3-Body a1(0)=cc2(0)=a(0)=2deg p1(0)=p2(0)=(3t(0)=1deg 5x(0)=14m6y(0)=52(0)=1m dx(0)=dy(0)=d2(0)=1m Stationkeeping lt=20km
Tether Longitudinal Vibration 1 6 1 mdash i mdash i mdash i mdash mdash i mdash 1 mdash 1 mdash mdash mdash i mdash mdash 1 mdash 1 mdash 1 mdash i mdash 1 mdash 1 mdash 1 mdash 1 mdash r
0 1 2 3 4 5
Time (Orbits) Tether In-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-6 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal local vertical and orbit normal (b) vibration response
68
STSS configuration 3-Body a(0)=a2(0)=ot(0)=0 P1(0)=p2(0)=pt(0)=0 8X(0)=1304 x 103m 8y(0)=8z(0)=0 dx(0)=dy(0)=dz(0)=0 Deployment
l=02km to 20km in 35 Orbits
Tether Length Profile
E
Satellite 1 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 1 Roll Angle mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash [ mdash i mdash i mdash i mdash i mdash | mdash
1 2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 4 Time (Orbits)
Tether Roll Angle
CD
oo ax
i imdashimdashimdashImdashimdashimdashimdashr-
_ 1 _ L
1 2 3 4 Time (Orbits)
Satellite 2 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 2 Roll Angle
-50
(a) 1 2 3 4
Time (Orbits) 1 2 3 4
Time (Orbits)
Figure 4-7 Deployment dynamics of the three-body STSS configuration without offset (a) attitude response
69
STSS configuration 3-Body a1(0)=a2(0)=at(0)=0 p1(0)=p2(0)=(3t(0)=0 8X(0)=1304x 103m 8y(0)=52(0)=0 dx(0)=dy(0)=d2(0)=0 Deployment
l=02km to 20km in 35 Orbits
Tether Longitudinal Vibration
to
-02
(b)
Figure 4-7
2 3 Time (Orbits)
Tether In-Plane Transverse Vibration
0 1 1 bull 1 1
1 2 3 4 Time (Orbits)
Tether Out-of-Plane Transverse Vibration
5
1 1 1 1 1 i i | i i | ^
_ i i i i L -I
1 2 3 4 5 Time (Orbits)
Deployment dynamics of the three-body STSS configuration without offset (b) vibration response
70
tether elongates its longitudinal static equilibrium position also changes due to an
increase in the gravity-gradient tether tension It may be pointed out that as in the
case of attitude motion there are no out-of-plane tether vibrations induced
When the same deployment conditions are applied to the case of 1 m offset in
the x direction (Figure 4-8) coupling effects are re-introduced The platform pitch
exceeds -100deg during the peak deployment acceleration However as the tether pitch
motion subsides the platform pitch response returns to a smaller amplitude oscillation
about its nominal equilibrium On the other hand the platform roll motion grows
to over 20deg in amplitude in the terminal stages of deployment Longitudinal and
transverse vibrations of the tether remain virtually unchanged from the zero offset
case except for minute out-of-plane perturbations due to the platform librations in
roll
45 Tether Retrieval
The system response for the case of the tether retrieval from 20 km to 200 m
with an offset of 1 m in the x and z directions at the platform end is presented in
Figure 4-9 Here the Coriolis force induced during retrieval renders the tether unstable
(Figure 4-9a) Consequently the pitch motion of the platform is also destabilized
through coupling The z offset provides an additional moment arm that shifts the
roll equilibrium to 35deg however this degree of freedom is not destabilized From
Figure 4-9(b) it is apparent that the transverse mode of the tether vibration is also
disturbed producing a high frequency response in the final stages of retrieval This
disturbance is responsible for the slight increase in the roll motion for both the tether
and subsatellite
46 Five-Body Tethered System
A five-body chain link system is considered next To facilitate comparison with
71
STSS configuration 3-Body a(0)=a2(0)=at(0)=0 P1(0)=P2(0)=P(0)=0 5X(0)=1304 x 103m 8y(0)=8z(0)=0 dx(0)=1mdy(0)=d2(0)=0 Deployment
lt=02km to 20km in 35 Orbits
Tether Length Profile
E
Satellite 1 Pitch Angle
1 2 3 Time (Orbits)
Satellite 1 Roll Angle
bull bull bull i bull bull bull bull i
1 2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 Time (Orbits)
Tether Roll Angle
i i i i J I I i _
1 2 3 4 Time (Orbits)
Satellite 2 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 2 Roll Angle 3E-3h
D )
T J ^ 0 E 0 eg
CQ
-3E-3
(a) 1 2 3 4
Time (Orbits) 1 2 3 4
Time (Orbits)
Figure 4-8 Deployment dynamics of the three-body STSS configuration with offset along the local vertical (a) attitude response
7 2
STSS configuration 3-Body a1(0)=a2(0)=at(0)=0 p1(0)=P2(0)=(3t(0)=0 Sx(0)=1304 x 103m 5y(0)=5z(0)=0 dx(0)=1mdy(0)=dz(0)=0 Deployment
l=02km to 20km in 35 Orbits
Tether Longitudinal Vibration
Time (Orbits) Tether In-Plane Transverse Vibration
n 1 1 1 f
I I I _j I I I I I 1 1 1 1 1 2 3 4 5
Time (Orbits) Tether Out-of-Plane Transverse Vibration
J 1 I I 1 I I I I I I I I L _ _ I I d 1 2 3 4 5
Time (Orbits)
Deployment dynamics of the three-body STSS configuration with offset along the local vertical (b) vibration response
(b)
Figure 4-8
73
STSS configuration 3-Body a1(0)=a2(0)=at(0)=0 p(0)=P2(0)=(3t(0)=0 Sx(0)=14m8y(0)=82(0)=0 dy(0)=0dx(0)=d2(0)=1m Retrieval
l=20km to 02km in 35 Orbits
Satellite 1 Pitch Angle
(a) Time (Orbits)
Figure 4-9 Retrieval dynamics of the along the local vertical ar
Tether Length Profile
i i i i i i i i i i I 0 1 2 3
Time (Orbits) Satellite 1 Roll Angle
Time (Orbits) Tether Roll Angle
b_ i i I i I i i L _ J
0 1 2 3 Time (Orbits)
Satellite 2 Roll Angle _ mdash | r i i | 1 i i imdashmdashj i 1 r mdash i mdash
04- -
0 1 2 3 Time (Orbits)
three-body STSS configuration with offset d orbit normal (a) attitude response
74
STSS configuration 3-Body a1(0)=a2(0)=ocl(0)=0 (31(0)=f32(0)=pt(0)=0 8x(0)=14m8y(0)=6z(0)=0 dy(0)=0dx(0)=dz(0)=1m Retrieval
l=20km to 02km in 35 Orbits
Tether Longitudinal Vibration
E 10 to
Time (Orbits) Tether In-Plane Transverse Vibration
1 2 Time (Orbits)
Tether Out-of-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-9 Retrieval dynamics of the three-body STSS configuration with offset along the local vertical and orbit normal (b) vibration response
75
the three-body system dynamics studied earlier the chain is extended through the
addition of two bodies a tether and a subsatellite with the same physical properties
as before (Section 42) Thus the five-body system consists of a platform (satellite 1)
tether 1 subsatellite 1 (satellite 2) tether 2 and subsatellite 2 (satellite 3) as shown
in Figure 4-10 The offset of tether 1 at the platform-end is denoted by d2 while that
of tether 2 to subsatellite 1 as 04 In this numerical example fcm^ = fcm^ = 0 ie
tethers 1 and 2 are attached to the centre of mass of subsatellites 1 and 2 respectively
The system response for the case where the tether attachment points coincide
wi th the centres of mass of the rigid bodies ie zero offset (d2 = d mdash 0) is presented
in Figure 4-11 Here ogtt and at2 represent the pitch motion of tethers 1 and 2
respectively whereas 03 is the pitch motions of satellite 3 A similar convention
is adopted for the rol l angle 0 A s before the zero offset eliminates the coupling
between the tethers and satellites such that they are now free to librate about their
static equil ibrium positions However their is s t i l l mutual coupling between the two
tethers This coupling is present regardless of the offset position since each tether is
capable of transferring a force to the other The coupling is clearly evident in the pitch
response of the two tethers (Figure 4 - l l a ) O n the other hand the roll motion appears
uncoupled as the in i t ia l conditions in rol l for the two tethers are identical Thus the
motion is in phase and there is no transfer of energy through coupling A s expected
during the elastic response the tethers vibrate about their static equil ibrium positions
and mutually interact (Figure 4 - l l b ) However due to the variation of tension along
the tether (x direction) they do not have the same longitudinal equilibrium Note
relatively large amplitude transverse vibrations are present particularly for tether 1
suggesting strong coupling effects with the longitudinal oscillations
76
Figure 4-10 Schematic diagram of the five-body system used in the numerical example
77
STSS configuration 5-Body a1(0)=(x2(0)=at1(0)=2deg a3(0)=a12(0)=25deg p1(0)=r32(0)=p3(0)=pt1(0)=pt2(0)=r 82(0)=84(0)=505)05m d2(0)=d4(0)=000 Stationkeeping lt1=lu=10km Pitch Roll
Satellite 1 Pitch Angle Satellite 1 Roll Angle
o 1 2 Time (Orbits)
Tether 1 Pitch and Roll Angle
0 1 2 Time (Orbits)
Tether 2 Pitch and Roll Angle
Time (Orbits) Satellite 2 Pitch and Roll Angle
Time (Orbits) Satellite 3 Pitch and Roll Angle
Time (Orbits) Time (Orbits)
Figure 4-11 Stationkeeping dynamics of the five-body STSS configuration withshyout offset (a) attitude response
78
STSS configuration 5-Body a1(0)=a2(0)=oct1(0)=2deg a3(0)=at2(0)=25o
p1(0)=p2(0)=(33(0)=pt1(0)=pt2(0)=10
62(0)=54(0)=50505m d2(0)=d4(0)=000 Stationkeeping lt1=1 =1 Okm
Tether 1 Longitudinal Vibration Tether 2 Longitudinal Vibration i i 1 i 1 r 1 1 1 1
(b) Time (Orbits) Time (Orbits)
Figure 4-11 Stationkeeping dynamics of the five-body STSS configuration withshyout offset (b) vibration response
79
The system dynamics during deployment of both the tethers in the double-
pendulum configuration is presented in Figure 4-12 Deployment of each tether takes
place from 200 m to 20 km in 35 orbits The pitch motion of the system is similar
to that observed during the three-body case The Coriolis effect causes the tethers
to pitch through a large angle ( laquo 50deg ) which in turn disturbs the platform and
subsatellite due to the presence of offset along the local vertical On the other hand
the roll response for both the tethers is damped to zero as the tether deploys in
confirmation with the principle of conservation of angular momentum whereas the
platform response in roll increases to around 20deg Subsatellite 2 remains virtually
unaffected by the other links since the tether is attached to its centre of mass thus
eliminating coupling Finally the flexibility response of the two tethers presented in
Figure 4-12(b) shows similarity with the three-body case
47 BICEPS Configuration
The mission profile of the Bl-static Canadian Experiment on Plasmas in Space
(BICEPS) is simulated next It is presently under consideration by the Canadian
Space Agency To be launched by the Black Brant 2000500 rocket it would inshy
volve interesting maneuvers of the payload before it acquires the final operational
configuration As shown in Figure 1-3 at launch the payload (two satellites) with
an undeployed tether is spinning about the orbit normal (phase 1) The internal
dampers next change the motion to a flat spin (phase 2) which in turn provides
momentum for deployment (phase 3) When fully deployed the one kilometre long
tether will connect two identical satellites carrying instrumentation video cameras
and transmitters as payload
The system parameters for the BICEPS configuration are summarized below [59 0 0 1
bull I = I2 = 0 366 0 kg bull m 2 (satellite inertias) [ 0 0 392_
bull mi = 7772 = 200 kg (mass of the satellites)
80
STSS configuration 5-Body a(0)=cx2(0)=al1(0)=2deg a3(0)=c (0)=25o
P1(0)=P2(0)=p3(0)=f3t1(0)=Pt2(0)=r 52(0)=84(0)=1304x 10300m d2(0)=d4(0)=1G0m Deployment lt1=le=02km to 20km in 35 Orbits
Pitch Roll Satellite 1 Pitch Angle
Time (Orbits) Satellite 2 Pitch and Roll Angle T 1mdash1 1 1 I I I I | I I I 1 I | I I lt~r
Tether Length Profile i 1 1 1 1 i 1 1 1 1 i 1 1 1 1 i
i i i i i 0 1 2 3 4 5
Time (Orbits) Satellite 1 Roll Angle
0 1 2 3 4 5 Time (Orbits)
Tether 1 amp 2 Roll Angle L i | I I 1 I 1 1 -I
-Ih I- bull I I i i i i I i i i i 1
0 1 2 3 4 5 Time (Orbits)
Satellite 3 Pitch and Roll Angle i- i 1 1 i 1 1 1 1 i 1 1 1 1 i 1 1 1 1 i
L i i i i i i 1 i i mdash i mdash i I i imdashimdashLJ
0 1 2 3 4 5 Time (Orbits)
Figure 4-12 Deployment dynamics of the five-body S T S S configuration with offshyset along the local vertical (a) attitude response
81
STSS configuration 5-Body a1(0)=a2(0)=oct1(0)=2deg a3(0)=at2(0)=25deg P1(0)=P2(0)=P3(0)=Pt1(0)=pt2(0)=1o
82(0)=54(0)=1304 x 10-300m d2(0)=d4(0)=100m Deployment lt1=l2=02km to 20km in 35 Orbits
Tether 1 Longitudinal Vibration Tether 2 Longitudinal Vibration
Time (Orbits) Time (Orbits) Tether 1 Z Tranverse Vibration Tether 2 Z Tranverse Vibration
Figure 4-12 Deployment dynamics of the five-body STSS configuration with offshyset along the local vertical (b) vibration response
82
bull EtAt = 61645 N (tether stiffness)
bull pt = 30 k g k m (tether density)
bull It = 1 km (tether length)
bull rid = 1-0 (tether structural damping coefficient)
The system is in a circular orbit at a 289 km altitude Offset of the tether
attachment point to the satellites at both ends is taken to be 078 m in the x
direction The response of the system in the stationkeeping phase for a prescribed set
of in i t ia l conditions is given by Figure 4-13 As in the case of the STSS configuration
there is strong coupling between the tether and the satellites rigid body dynamics
causing the latter to follow the attitude of the tether However because of the smaller
inertias of the payloads there are noticeable high frequency modulations arising from
the tether flexibility Response of the tether in the elastic degrees of freedom is
presented in Figure 4-13(b) Note the longitudinal vibrations decay quite rapidly
(lt 05 orbits) due to the structural damping however it has vir tual ly no effect on
the transverse oscillations
The unique mission requirement of B I C E P S is the proposed use of its ini t ia l
angular momentum in the cartwheeling mode to aid in the deployment of the tethered
system The maneuver is considered next The response of the system during this
maneuver wi th an ini t ia l cartwheeling rate of 5 deg s is illustrated in Figure 4-14(a) The
in i t ia l tether length is taken to be 10 m and is deployed to 1 km There is an in i t ia l
increase in the pitch motion of the tether however due to the conservation of angular
momentum the cartwheeling rate decreases proportionally to the square of the change
in tether deployment rate The result is a rapid drop in the cartwheeling rate unti l
the system stops rotating entirely and simply oscillates about its new equilibrium
point Consequently through coupling the end-bodies follow the same trend The
roll response also subsides once the cartwheeling motion ceases However
83
BICEPS configuration 3-Body a1(0)=o2(0)=at(0)=2deg P1(0)=P2(0)=pt(0)=1deg 8X(0)=001 m 6y(0)=52(0)=01 m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping l=1 km
Satellite 1 Pitch Angle Satellite 1 Roll Angle - i 1 1 1 1 i - i q r 1 1 -gt
I- I i t i i i i I H 0 1 2 0 1 2
Time (Orbits) Time (Orbits) Tether Pitch Angle Tether Roll Angle
F mdash 1 mdash 1 mdash mdash 1 mdash i mdash 1 mdash 1 mdash mdash 1 mdash = i F mdash 1 mdash mdash mdash 1 mdash i mdash 1 mdash 1 mdash 1 mdash 1 mdash q
(a) Time (Orbits) Time (Orbits)
Figure 4-13 Stationkeeping dynamics of the three-body BICEPS configuration with offset along the local vertical (a) attitude response
84
BICEPS configuration 3-Body a1(0)=o2(0)=at(0)=2deg Pl(0)=p2(0)=(3t(0)=1deg 8X(0)=001 m 8y(0)=82(0)=01 m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping lt=1km
Tether Longitudinal Vibration 1E-2[
laquo = 5 E - 3 |
000 002
Time (Orbits) Tether In-Plane Transverse Vibration
Time (Orbits) Tether Out-of-Plane Transverse Vibration
Time (Orbits)
Figure 4-13 Stationkeeping dynamics of the three-body BICEPS configuration with offset along the local vertical (b) vibration response
85
BICEPS configuration 3-Body a1(0)=o2(0)=at(0)=2deg a1(0)=a2(0)=at(0)=57s Pl(0)=P2(0)=(3t(0)=1deg 8X(0)=115 x 103m 8y(0)=82(0)=0 dx(0)=078m dy(0)=d2(0)=0 Cartwheeling Deployment
lt=10m to 1km in 1 Orbit
Tether Length Profile
E
2000
cu T3
2000
co CD
2000
C D CD
T J
Satellite 1 Pitch Angle
1 2 Time (Orbits)
Satellite 1 Roll Angle
cn CD
33 cdl
Time (Orbits) Tether Pitch Angle
Time (Orbits) Tether Roll Angle
Time (Orbits) Satellite 2 Pitch Angle
1 2 Time (Orbits)
Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-14 Cartwheeling dynamics with deployment for the three-body BICEPS with offset along the local vertical (a) attitude response
86
BICEPS configuration 3-Body a1(0)=a2(0)=at(0)=2deg a (0)=a2(0)=a(0)=57s P1(0)=P2(0)=f3(0)=1deg 8X(0)=115x103m 8y(0)=8z(0)=0 dx(0)=078m dy(0)=d2(0)=0 Cartwheeling Deployment
l=1 Om to 1 km in 1 Orbit
5E-3h
co
0E0
025
000 E to
-025
002
-002
(b)
Figure 4-14
Tether Longitudinal Vibration
Time (Orbits) Tether In-Plane Transverse Vibration
Time (Orbits) Tether Out-of-Plane Transverse Vibration
Time (Orbits)
Cartwheeling dynamics with deployment for the three-body BICEPS with offset along the local vertical (b) vibration response
87
as the deployment maneuver is completed the satellites continue to exhibit a small
amplitude low frequency roll response with small period perturbations induced by
the tether elastic oscillations superposed on it (Figure 4-14b)
48 OEDIPUS Spinning Configuration
The mission entitled Observation of Electrified Distributions of Ionospheric
Plasmas - A Unique Strategy ( O E D I P U S ) also consists of a 1 k m tethered system with
two similar end-masses It was first flown in a suborbital flight in 1989 ( O E D I P U S A )
followed by a recent study in November 1995 ( O E D I P U S - C ) in which the University
of Br i t i sh Columbia was one of the participants Dynamically O E D I P U S represents a
unique system never encountered before Launched by a Black Brant rocket developed
at Br is to l Aerospace L t d the system ie the end-bodies together wi th the tether
spins about the longitudinal axis of the tether to achieve stabilized alignment wi th
Earth s magnetic field
The response of the system undergoing a spin rate of j = l deg s is presented
in Figure 4-15 using the same parameters as those outlined for the B I C E P S conshy
figuration in Section 47 Again there is strong coupling between the three bodies
(two satellites and tether) due to the nonzero offset The spin motion introduces an
additional frequency component in the tethers elastic response which is transferred
to the l ibrational motion of the satellites A s the spin rate increases to 1 0 deg s (Figure
4-16) the amplitude and frequency of the perturbations also increase however the
general character of the response remains essentially the same
88
OEDIPUS configuration 3-Body a1(0)=a2(0)=at(0)=2deg i(0)^[2(0H(0)=Ys p1(0)=p2(0)=p(0)=r 8X(0)=12 x 103m 8y(0)=82(0)=01m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping l=1 km
Satellite 1 Pitch Angle Satellite 1 Roll Angle mdash i 1 1 1 a r 1 1 1 r
(a) Time (Orbits) Time (Orbits)
Figure 4-15 Spin dynamics (7 = ldegs) of the three-body OEDIPUS configuration with offset along the local vertical (a) attitude response
89
OEDIPUS configuration 3-Body a1(0)=a2(0)=cct(0)=2deg Y1(0)=Y2(0)=Yt(0)=l7s (31(0)=p2(0)=pt(0)=1deg 8X(0)=12 x 103m 8y(0)=5z(0)=01 m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping lt=1 km
1E-2
to
OEO
E 00
to
Tether Longitudinal Vibration
005
Time (Orbits) Tether In-Plane Transverse Vibration
Time (Orbits) Tether Out-of-Plane Transverse Vibration
Time (Orbits)
Figure 4-15 Spin dynamics (7 = l deg s ) of the three-body O E D I P U S configuration with offset along the local vertical (b) vibration response
90
OEDIPUS configuration 3-Body a1(0)=a2(0)=a(0)=2deg Yi(0)=Y2(0)=rt(0)=107s 81(0)=p2(0)=pt(0)=1deg 8X(0)=12 x 103m 8y(0)=82(0)=01m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping lt=1km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-16 Spin dynamics (7= 10degs) of the three-body OEDIPUS configura tion with offset along the local vertical (a) attitude response
91
OEDIPUS configuration 3-Body a1(0)=a2(0)=ot(0)=2deg Yi(0)H(0)4(0)laquo107s P1(0)=P2(0)=pt(0)=1deg 8X(0)=12 x 103m 8y(0)=82(0)=01m dx(0)=078m dy(0)=dz(0)=0 Stationkeeping lt=1 km
Tether Longitudinal Vibration
Figure 4-16 Spin dynamics (7 = 10degs) of the three-body OEDIPUS configurashytion with offset along the local vertical (b) vibration response
92
5 ATTITUDE AND VIBRATION CONTROL
51 Att itude Control
511 Preliminary remarks
The instability in the pitch and roll motions during the retrieval of the tether
the large amplitude librations during its deployment and the marginal stability during
the stationkeeping phase suggest that some form of active control is necessary to
satisfy the mission requirements This section focuses on the design of an attitude
controller with the objective to regulate the librational dynamics of the system
As discussed in Chapter 1 a number of methods are available to accomplish
this objective This includes a wide variety of linear and nonlinear control strategies
which can be applied in conjuction with tension thrusters offset momentum-wheels
etc and their hybrid combinations One may consider a finite number of Linear T ime
Invariant (LTI) controllers scheduled discretely over different system configurations
ie gain scheduling[43] This would be relevant during deployment and retrieval of
the tether where the configuration is changing with time A n alternative may be
to use a Linear T ime Varying ( L T V ) model in conjuction with an adaptive control
scheme where on-line parametric identification may be used to advantage[56] The
options are vir tually limitless
O f course the choice of control algorithms is governed by several important
factors effective for time-varying configurations computationally efficient for realshy
time implementation and simple in character Here the thrustermomentum-wheel
system in conjuction with the Feedback Linearization Technique ( F L T ) is chosen as
it accounts for the complete nonlinear dynamics of the system and promises to have
good overall performance over a wide range of tether lengths It is well suited for
93
highly time-varying systems whose dynamics can be modelled accurately as in the
present case
The proposed control method utilizes the thrusters located on each rigid satelshy
lite excluding the first one (platform) to regulate the pitch and rol l motion of the
tether to which it is attached ie the thrusters located on satellite 2 (link 3) regshy
ulates the attitude motion of tether 1 (link 2) O n the other hand the l ibrational
motion of the rigid bodies (platform and subsatellite) is controlled using a set of three
momentum wheels placed mutually perpendicular to each other
The F L T method is based on the transformation of the nonlinear time-varying
governing equations into a L T I system using a nonlinear time-varying feedback Deshy
tailed mathematical background to the method and associated design procedure are
discussed by several authors[435758] The resulting L T I system can be regulated
using any of the numerous linear control algorithms available in the literature In the
present study a simple P D controller is adopted arid is found to have good perforshy
mance
The choice of an F L T control scheme satisfies one of the criteria mentioned
earlier namely valid over a wide range of tether lengths The question of compushy
tational efficiency of the method wi l l have to be addressed In order to implement
this controller in real-time the computation of the systems inverse dynamics must
be executed quickly Hence a simpler model that performs well is desirable Here
the model based on the rigid system with nonlinear equations of motion is chosen
Of course its validity is assessed using the original nonlinear system that accounts
for the tether flexibility
512 Controller design using Feedback Linearization Technique
The control model used is based on the rigid system with the governing equa-
94
tions of motion given by
Mrqr + fr = Quu (51)
where the left hand side represents inertia and other forces while the right hand side
is the generalized external force due to thrusters and momentum-wheels and is given
by Eq(2109)
Lett ing
~fr = fr~ QuU (52)
and substituting in Eq(296)
ir = s(frMrdc) (53)
Expanding Eq(53) it can be shown that
qr = S(fr Mrdc) - S(QU Mr6)u (5-4)
mdash = Fr - Qru
where S(QU Mr0) is the column matrix given by
S(QuMr6) = [s(Qu(l)Mr0) 5 (Q u ( 2 ) M r | 0 ) S(QU nu) M r | 0 ) ] (55)
and Qu-i) is the i ^ column of Qu Extract ing only the controlled equations from
Eq(54) ie the attitude equations one has
Ire = Frc ~~ Qrcu
(56)
mdash vrci
where vrc is the new control input required to regulate the decoupled linear system
A t this point a simple P D controller can be applied ie
Vrc = qrcd +KvQrcd-Qrc) + Kp(Qrcd~ Qrc)- (57)
Eq(57) represents the secondary controller with Eq(54) as the primary controller
Kp and Kv are the proportional and derivative gain matrices with qrcdi Qrcd
an(^ qrcd
95
as the desired acceleration velocity and position vector for the attitude angles of each
controlled body respectively Solving for u from Eq(56) leads to
u Qrc Frc ~ vrcj bull (58)
5 13 Simulation results
The F L T controller is implemented on the three-body STSS tethered system
The choice of proportional and derivative matrix gains is based on a desired settling
time of ts mdash 05 orbit and a damping factor of = 07 for both the pitch and roll
actuators in each body Given O and ts it can be shown that
-In ) 0 5 V T ^ r (59)
and hence kn bull mdash10
V (510)
where kp and kVj are the diagonal elements of the gain matrices Kp and Kv in
Eq(57) respectively[59]
Note as each body-fixed frame is referred directly to the inertial frame FQ
the nominal pitch equilibrium angle is zero only when measured from the L V L H frame
(OJJ mdash 0) However it is 6 when referred to FQ Hence the desired position vector
qrC(l is set equal to 0(t) ie the time-varying orbital angle for the pitch motion
and zero for the roll and yaw rotations such that
qrcd 37Vxl (511)
96
The desired velocity and acceleration are then given as
e U 3 N x l (512)
E K 3 7 V x l (513)
When the system is in a circular orbit qrc^ = 0
Figure 5-1 presents the controlled response of the STSS system defined in
Chapter 4 in the stationkeeping phase with offset d2 = 111T m and If = 20
km A s mentioned earlier the F L T controller is based on the nonlinear rigid model
Note the pitch and roll angles of the rigid bodies as well as of the tether are now
attenuated in less than 1 orbit (Figure 5-la) Consequently the longitudinal vibration
response of the tether is free of coupling arising from the librational modes of the
tether This leaves only the vibration modes which are eventually damped through
structural damping The pitch and roll dynamics of the platform require relatively
higher control moments Ma and Mpi respectively (Figure 5-lb) since they have to
overcome the extra moments generated by the offset of the tether attachment point
Furthermore the platform demands an additional moment to maintain its orientation
at zero pitch and roll angle since it is not the nominal equilibrium position The nonshy
zero static equilibrium of the platform arises due to its products of inertia On the
other hand the tether pitch and roll dynamics require only a small ini t ia l control
force of about plusmn 1 N which eventually diminishes to zero once the system is
97
and
o 0
Qrcj mdash
o o
0 0
Oiit) 0 0
1
N
N
1deg I o
STSS configuration 3-Body a1(0)=a2(0)=at(0)=2lt
P1(0)=p2(0)=pt(0)=1 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=dz(0)=1m Controlled Stationkeeping lt=20km
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
05 10 15 Time (Orbits)
Sat 2 Pitch and Roll Angles
20
CO CD
T J
cdT
cT
- mdash mdash 1 1 1 1 1 I - 1
Pitch -
A Roll -
A -
1 bull bull
-
1 Time (Orbits)
Tether Y Tranverse Vibration
1 2 Time (Orbits)
Tether Z Tranverse Vibration
gt to N
to
Time (Orbits) Time (Orbits) Longitudinal Vibration
145h
(a) Time (Orbits)
Figure 5-1 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local horizontal local vertical and orbit normal (a) attitude and vibration response
98
1
STSS configuration 3-Body a1(0)=ct2(0)=cct(0)=20
P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=d2(0)=1m Controlled Stationkeeping lt=20km
Sat 1 Pitch Control Moment Sat 1 Roll Control Moment
Figure 5-1 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local horizontal local vertical and orbit normal (b) control actuator time histories
99
stabilized Similarly for the end-body a very small control torque is required to
attenuate the pitch and roll response
If the complete nonlinear flexible dynamics model is used in the feedback loop
the response performance was found to be virtually identical with minor differences
as shown in Figure 5-2 For example now the pitch and roll response of the platform
is exactly damped to zero with no steady-state error as against the minor almost
imperceptible deviation for the case of the controller basedon the rigid model (Figure
5-la) In addition the rigid controller introduces additional damping in the transverse
mode of vibration where none is present when the full flexible controller is used
This is due to a high frequency component in the at motion that slowly decays the
transverse motion through coupling
As expected the full nonlinear flexible controller which now accounts the
elastic degrees of freedom in the model introduces larger fluctuations in the control
requirement for each actuator except at the subsatellite which is not coupled to the
tether since f c m 3 = 0 The high frequency variations in the pitch and roll control
moments at the platform are due to longitudinal oscillations of the tether and the
associated changes in the tension Despite neglecting the flexible terms the overall
controlled performance of the system remains quite good Hence the feedback of the
flexible motion is not considered in subsequent analysis
When the tether offset at the platform is restricted to only 1 m along the x
direction a similar response is obtained for the system However in this case the
steady-state error in the platforms rigid body motion is much smaller (Figure 5-3a)
In addition from Figure 5-3(b) the platforms control requirement is significantly
reduced
Figure 5-4 presents the controlled response of the STSS deploying a tether
100
STSS configuration 3-Body a1(0)=a2(0)=oct(0)=2deg p1(0)=P2(0)=p(0)=1deg 8x(0)=14m5y(0)=82(0)=1m dx(0)=dy(0)=dz(0)=1m Controlled Stationkeeping l=20km
Tether Pitch and Roll Angles
c n
T J
cdT
Sat 1 Pitch and Roll Angles 1 r
00 05 10 15 Time (Orbits)
Sat 2 Pitch and Roll Angles
c n CD
cdT a
o T J
CQ
- - I 1 1 1 I 1 1 1 1
P i t c h
R o l l
bull
1 -
0 1 2 Time (Orbits)
Tether Y Tranverse Vibration
0 1 2 Time (Orbits)
Tether Z Tranverse Vibration
E gt
10 N
CO
Time (Orbits) Time (Orbits) Longitudinal Vibration
145h
(a) Time (Orbits)
Figure 5-2 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear flexible FLT controller with offset along the local horizontal local vertical and orbit normal (a) attitude and vibration response
101
STSS configuration 3-Body a1(0)=a2(0)=cxt(0)=2deg 31(0)=p2(0)=p(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=d2(0)=1m Controlled Stationkeeping lt=20km
r i i i t i mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash mdash J 0 1 2 0 1 2
Time (Orbits) Time (Orbits) Tether Pitch Control Thrust Tether Roll Control Thruster
(b) Time (Orbits) Time (Orbits)
Figure 5-2 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear flexible FLT controller with offset along the local horizontal local vertical and orbit normal (b) control actuator time histories
102
STSS configuration 3-Body a1(0)=o2(0)=a(0)=2deg P1(0)=p2(0)=(3t(0)=1deg 8x(0)=14m5y(0)=6z(0)=1m dx(0)=1mdy(0)=dz(0)=0 Controlled Stationkeeping lt=20km
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
05 10 15 Time (Orbits)
Sat 2 Pitch and Roll Angles
CD
eg CQ
0 1 2 Time (Orbits)
Tether Y Tranverse Vibration
1 2 Time (Orbits)
Tether Z Tranverse Vibration
gt CO
Time (Orbits) Time (Orbits) Longitudinal Vibration
145h
(a) Time (Orbits)
Figure 5-3 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local vertical (a) attitude and vibration response
103
STSS configuration 3-Body a1(0)=o2(0)=al(0)=2o
P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=1mdy(0)=dz(0)=0 Controlled Stationkeeping lt=20km
Sat 1 Pitch Control Moment 11 1 1 r
Sat 1 Roll Control Moment
0 1 2 Time (Orbits)
Tether Pitch Control Thrust
0 1 2 Time (Orbits)
Sat 2 Pitch Control Moment
0 1 2 Time (Orbits)
Tether Roll Control Thruster
-07 o 1
Time (Orbits Sat 2 Roll Control Moment
1E-5
0E0F
Time (Orbits)
Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local vertical (b) control actuator time histories
104
STSS configuration 3-Body a1(0)=a2(0)=a(0)=2deg Pl(0)=p2(0)=pt(0)=1deg 6X(0)=1304 x 103m 8y(0)=82(0)=0 dx(0)=1mdy(0)=d2(0)=0 Controlled Deployment
l=02km to 20km in 35 Orbits
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
o 1 2 Time (Orbits)
Tether Y Tranverse Vibration
0 1 2 Time (Orbits)
Sat 2 Pitch and Roll Angles
CD CD
CM CO
Time (Orbits) Tether Z Tranverse Vibration
to
150
100
1 2 3 0 1 2
Time (Orbits) Time (Orbits) Longitudinal Vibration
to
(a)
50
2 3 Time (Orbits)
Figure 5 - 4 Deployment dynamics of the three-body STSS using the nonlinear rigid FLT controller with offset along the local vertical (a) attitude and vibration response
105
STSS configuration 3-Body a1(0)=oc2(0)=al(0)=20
P1(0)=p2(0)=pt(0)=1deg 8X(0)=1304 x 103m 5y(0)=52(0)=0 dx(0)=1mdy(0)=dz(0)=0 Controlled Deployment
l=02km to 20km in 35 Orbits
Tether Length Profile i 1
2 0 h
E 2L
Sat 1 Pitch Control Moment t mdash r mdash i mdash i mdash | mdash i mdash i mdash i mdash i mdash | mdash r mdash i mdash i mdash i mdash | mdash r mdash
0 5
1 2 3
Time (Orbits) Sat 1 Roll Control Moment
i i t | t mdash t mdash i mdash [ i i mdash r -
1 2
Time (Orbits) Tether Pitch Control Thrust
1 2 3 Time (Orbits)
Tether Roll Control Thruster
2 E - 5
1 2 3
Time (Orbits) Sat 2 Pitch Control Moment
1 2
Time (Orbits) Sat 2 Roll Control Moment
1 E - 5
E
O E O
Time (Orbits)
Deployment dynamics of the three-body STSS using the nonlinear rigid FLT controller with offset along the local vertical (b) control actuator time histories
106
from 02 km to 20 km in 35 orbits with a i m offset along the tether length A s before
the systems attitude motion is well regulated by the controller However the control
cost increases significantly for the pitch motion of the platform and tether due to the
Coriolis force induced by the deployment maneuver (Figure 5-4b) Initially there is a
sinusoidal increase in the pitch control requirement for both the tether and platform
as the tether accelerates to its constant velocity VQ Then the control requirement
remains constant for the platform at about 60 N m as opposed to the tether where the
thrust demand increases linearly Finally when the tether decelerates the actuators
control input reduces sinusoidally back to zero The rest of the control inputs remain
essentially the same as those in the stationkeeping case In fact the tether pitch
control moment is significantly less since the tether is short during the ini t ia l stages
of control However the inplane thruster requirement Tat acts as a disturbance and
causes 6y to nearly double from the uncontrolled value (Figure 4-8a) O n the other
hand the tether has no out-of-plane deflection
Final ly the effectiveness of the F L T controller during the crit ical maneuver of
retrieval from 20 km to 02 k m in 35 orbits is assessed in Figure 5-5 A s in the earlier
cases the system response in pitch and roll is acceptable however the controller is
unable to suppress the high frequency elastic oscillations induced in the tether by the
retrieval O f course this is expected as there is no active control of the elastic degrees
of freedom However the control of the tether vibrations is discussed in detail in
Section 52 The pitch control requirement follows a similar trend in magnitude as
in the case of deployment with only minor differences due to the addition of offset in
the out-of-plane direction ( d 2 = 1 0 1 ^ m) This is also responsible for the higher
roll moment requirement for the platform control (Figure 5-5b)
107
STSS configuration 3-Body a1(0)=X2(0)=a(0)=2deg P1(0)=p2(0)=pt(0)=1deg 5x(0)=14m8y(0)=52(0)=0 dx(0)=1mdy(0)=dz(0)=0 Controlled Retrieval
l=20km to 02km in 35 Orbits
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
0 1 2 Time (Orbits)
Sat 2 Pitch and Roll Angles
CO CD
cdT
CO CD
T J CM
CO
1 2 Time (Orbits)
Tether Y Tranverse Vibration
Time (Orbits) Tether Z Tranverse Vibration
150
100
1 2 3 0 1 2 Time (Orbits) Time (Orbits)
Longitudinal Vibration
CO
(a)
x 50
2 3 Time (Orbits)
Figure 5 -5 Retrieval dynamics of the three-body STSS using the non-linear rigid FLT controller with offset along the local vertical and orbit normal (a) attitude and vibration response
108
STSS configuration 3-Body a1(0)=ct2(0)=at(0)=2deg p1(0)=p2(0)=(3(0)=1deg 5x(0)=14m5y(0)=82(0)=0 dx(0)=1mdy(0)=dz(0)=0 Controlled Retrieval
l=20km to 02km in 35 Orbits
Tether Length Profile
E
1 2 Time (Orbits)
Tether Pitch Control Thrust
2E-5
1 2 Time (Orbits)
Sat 2 Pitch Control Moment
OEOft
1 2 3 Time (Orbits)
Sat 1 Roll Control Moment
1 2 3 Time (Orbits)
Tether Roll Control Thruster
1 2 Time (Orbits)
Sat 2 Roll Control Moment
1E-5 E z
2
0E0
(b) Time (Orbits) Time (Orbits)
Figure 5-5 Retrieval dynamics of the three-body STSS using the non-linear rigid FLT controller with offset along the local vertical and orbit normal (b) control actuator time histories
109
52 Control of Tethers Elastic Vibrations
521 Prel iminary remarks
The requirement of a precisely controlled micro-gravity environment as well
as the accurate release of satellites into their final orbit demands additional control
of the tethers vibratory motion in addition to its attitude regulation To that end
an active vibration suppression strategy is designed and implemented in this section
The strategy adopted is based on offset control ie time dependent variation
of the tethers attachment point at the platform (satellite 1) A l l the three degrees
of freedom of the offset motion are used to control both the longitudinal as well as
the inplane and out-of-plane transverse modes of vibration In practice the offset
controller can be implemented through the motion of a dedicated manipulator or a
robotic arm supporting the tether which in turn is supported by the platform The
focus here is on the control of elastic deformations during stationkeeping (ie fully
deployed tether fixed length) as it represents the phase when the mission objectives
are carried out
This section begins with the linearization of the systems equations of motion
for the reduced three body stationkeeping case This is followed by the design of
the optimal control algorithm Linear Quadratic Gaussian-Loop Transfer Recovery
( L Q G L T R ) based on the reduced model and its implementation on the full nonlinear
model Final ly the systems response in the presence of offset control is presented
which tends to substantiate its effectiveness
522 System linearization and state-space realization
The design of the offset controller begins with the linearization of the equations
of motion about the systems equilibrium position Linearization of extremely lengthy
(even in matrix form) highly nonlinear nonautonomous and coupled equations of
110
motion presents a challenging problem This is further complicated by the fact the
pitch angle ot is not referred to the L V L H frame Hence the equilibrium pitch angle
is not a constant but is equal to the instantaneous orbital angle 9(t) Two methods
are available to resolve this problem
One may use the non-stationary equations of motion in their present form and
derive a controller based on the Linear T ime Varying ( L T V ) system Alternatively
one can design a Linear T ime Invariant (LTI) controller based on a new set of reduced
governing equations representing the motion of the three body tethered system A l shy
though several studies pertaining to L T V systems have been reported[5660] the latter
approach is chosen because of its simplicity Furthermore the L T I controller design
is carried out completely off-line and thus the procedure is computationally more
efficient
The reduced model is derived using the Lagrangian approach with the genershy
alized coordinates given by
where lty and are the pitch and roll angles of link i relative to the L V L H frame
respectively 5X 8y and 5Z are the generalized coordinates associated with the flexible
tether deformations in the longitudinal inplane and out-of-plane transverse direcshy
tions respectively Only the first mode of vibration is considered in the analysis
The nonlinear nonautonomous and coupled equations of motion for the tethshy
ered system can now be written as
Qred = [ltXlPlltX2P2fixSy5za302gt] (514)
M redQred + fr mdash 0gt (515)
where Mred and frec[ are the reduced mass matrix and forcing term of the system
111
respectively They are functions of qred and qred in addition to the time varying offset
position d2 and its time derivatives d2 and d2 A detailed derivation of Eq(515)
is given in Appendix II A n additional consequence of referring the pitch motion to
a local frame is that the new reduced equations are now independent of the orbital
angle 9 and under the further assumption of the system negotiating a circular orbit
91 remains constant
The nonlinear reduced equations can now be linearized about their static equishy
l ibr ium position For all the generalized coordinates the equilibrium position is zero
wi th the exception of 5X which has a non-zero equilibrium 5XQ The offset position
d2) is linearized about its ini t ial position d2^ Linearizing Eq(515) and recasting
into matr ix form gives
Msqred + CsQred + KsQred + Md^2 + Cdd2 + Kdd2 + fs = 0 (516)
where Ms Cs Ks Md Cd Kd and s are constant matrices Mul t ip ly ing throughout
by Ms
1 gives
qr = -Ms 1Csqr - Ms
lKsqr
- Ms~lCdd22 - Ms-Kdd2 - Ms~Mdd2 - Mg~1 fs r-1 (517)
Defining ud = d2 and v = qTd^T Eq(517) can be rewritten as
Pu t t ing
v =
+
-M~lCs -M~Cd -1
0
-Mg~lMd
Id
0 v + -M~lKs -M~lKd
0 0
-M-lfs
0
MCv + MKv + Mlud +Fs
- 4 ) - ( gt the L T I equations of motion can be recast into the state-space form as
(518)
(519)
x =Ax + Bud + Fd (520)
112
where
A = MC MK Id12 o
24x24
and
Let
B = ^ U 5 R 2 4 X 3 -
Fd = I FJ ) e f t 2 4 1
mdash _ i _ -5
(521)
(522)
(523)
(524)
where x is the perturbation vector from the constant equilibrium state vector xeq
Substituting Eq(524) into Eq(520) gives the linear perturbation state equation
modelling the reduced tethered system as
bullJ x mdash x mdash Ax + Bud + (Fd + Axeq) (525)
It can be shown that for ideal control[60] Fd + Axeq = 0 thus giving the familiar
state-space equation
S=Ax + Bud (526)
The selection of the output vector completes the state space realization of
the system The output vector consists of the longitudinal deformation from the
equil ibrium position SXQ of the tether at xt = h the slope of the tether due to the
transverse deformation at xt = 0 and the offset position d2 from its in i t ia l position
d2Q Thus the output vector y is given by
K(o)sx
V
dx ~ dXQ dy ~ dyQ
d z - d z0 J
Jy
dx da XQ vo
d z - dZQ ) dy dyQ
(527)
113
where lt ( 0 ) = ddxi[$vxi)]Xi=o and lt ( 0 ) = ddxi[4gtw(xi)]Xi=0
523 Linear Quadratic Gaussian control with Loop Transfer Recovery
W i t h the linear state space model defined (Eqs526527) the design of the
controller can commence The algorithm chosen is the Linear Quadratic Gaussian
( L Q G ) estimator based optimal controller[6061] The L Q G is a widely used optimal
controller with its theoretical background well developped by many authors over the
last 25 years It involves of the design of the Kalman-Bucy Fi l ter ( K B F ) which
provides an estimate of the states x and a Linear Quadratic Regulator ( L Q R ) which
is separately designed assuming all the states x are known Both the L Q R and K B F
designs independently have good robustness properties ie retain good performance
when disturbances due to model uncertainty are included However the combined
L Q R and K B F designs ie the L Q G design may have poor stability margins in the
presence of model uncertainties This l imitat ion has led to the development of an L Q G
design procedure that improves the performance of the compensator by recovering
the full state feedback robustness properties at the plant input or output (Figure 5-6)
This procedure is known as the Linear Quadratic Guassian-Loop Transfer Recovery
( L Q G L T R ) control algorithm A detailed development of its theory is given in
Ref[61] and hence is not repeated here for conciseness
The design of the L Q G L T R controller involves the repeated solution of comshy
plicated matrix equations too tedious to be executed by hand Fortunenately the enshy
tire algorithm is available in the Robust Control Toolbox[62] of the popular software
package M A T L A B The input matrices required for the function are the following
(i) state space matrices ABC and D (D = 0 for this system)
(ii) state and measurement noise covariance matrices E and 0 respectively
(iii) state and input weighting matrix Q and R respectively
114
u
MODEL UNCERTAINTY
SYSTEM PLANT
y
u LQR CONTROLLER
X KBF ESTIMATOR
u
y
LQG COMPENSATOR
PLANT
X = f v ( X t )
COMPENSATOR
x = A k x - B k y
ud
= C k f
y
Figure 5-6 Block diagram for the LQGLTR estimator based compensator
115
A s mentionned earlier the main objective of this offset controller is to regu-mdash
late the tether vibration described by the 5 equations However from the response
of the uncontrolled system presented in Chapter 4 it is clear that there is a large
difference between the magnitude of the librational and vibrational frequencies This
separation of frequencies allow for the separate design of the vibration and attitude mdash mdash
controllers Thus only the flexible subsystem composed of the S and d equations is
required in the offset controller design Similarly there is also a wide separation of
frequencies between the longitudinal and transverse modes of vibrations permitt ing mdash
the decoupling of the flexible subsystem into a set of longitudinal (Sx and dx) and
transverse (5y Sz dy and dz) subsystems The appended offset system d must also
be included since it acts as the control actuator However it is important to note
that the inplane and out-of-plane transverse modes can not be decoupled because
their oscillation frequencies are of the same order The two flexible subsystems are
summarized below
(i) Longitudinal Vibra t ion Subsystem
This is defined by u mdash Auxu + Buudu
(528)
where xu = SxdX25xdX2T u d u = dX2 and from Eq(527)
0 0 $u(lt) 0 0 0 0 1
0 0 1 0 0 0 0 1
(529)
Au and Bu are the rows and columns of A and B respectively and correspond to the
components of xu
The L Q R state weighting matrix Qu is taken as
Qu =
bull1 0 0 o -0 1 0 0 0 0 1 0
0 0 0 10
(530)
116
and the input weighting matrix Ru = 1^ The state noise covariance matr ix is
selected as 1 0 0 0 0 1 0 0 0 0 4 0
LO 0 0 1
while the measurement noise covariance matrix is taken as
(531)
i o 0 15
(532)
Given the above mentionned matrices the design of the L Q G L T R compenshy
sator is computed using M A T L A B with the following 2 commands
(i) kfu = l q r c ( ^ ( i C bdquo d i a g m x ( E u e u ) )
(ii) [AkugtBkugtCkugtDku =^ryAuBuCuDukfuQuRur)
where r is a scaler
The first command is used to compute the Ka lman filter gain matrix kfu Once
kfu is known the second command is invoked returning the state space representation
of the L Q G L T R dynamic compensator to
(533) xu mdash Akuxu Bkuyu
where xu is the state estimate vector oixu The function l t ry performs loop transfer
recovery at the system output ie the return ratio at the output approaches that
of the K B F loop given by Cu(sld - Au)Kfu[61 This is achieved by choosing a
sufficently large scaler value r in l t ry such that the singular values of the return
ratio approach those of the target design For the longitudinal controller design
r mdash 5 x 10 5 Figure 5-7(a) compares the singular values of the recovered compensator
design and the non-recovered L Q G compensator design with respect to the target
Unfortunetaly perfect recovery is not possible especially at higher frequencies
117
100
co
gt CO
-50
-100 10
Longitudinal Compensator Singular Values i I I I N I I mdash i mdash i i i i n i i mdash i mdash i i i i inimdash imdashi i i i i i i i mdash i mdash i i M i I I
Target r = 5x10 5
r = 0 (LQG)
i i i i I I i ii i i i 1 1 1 1 i i i i i 1 1 1 1 ii i I I i i i i 11111 -3 10 -2
(a)
101 10deg Frequency (rads)
101 10
Transverse Compensator Singular Values
100
50
X3
gt 0
CO
-50
bull100 10
-imdashi i 111II i 111 ni 1mdashi I I I I I I I 1mdashi I I I N I I 1mdashi i M 111
Target r = 50 r = 0 (LQG)
m i i i i i i n i i i i I I I I I I I 1mdashi I I I I I I I 1mdashi i 111 n -3 10 -2
(b)
101 10deg Frequency (rads)
101 10
Figure 5-7 Singular values for the LQG and LQGLTR compensator compared to target return ratio (a) longitudinal design (b) transverse design
118
because the system is non-minimal ie it has transmission zeros with positive real
parts[63]
(ii) Transverse Vibra t ion Subsystem
Here Xy mdash AyXy + ByU(lv
Dv = Cvxv
with xv = 6y 8Z dV2 dZ2 Sy 6Z dV2 dZ2T] udv = dy2dZ2T and
(534)
Cy
0 0 0 0 (0) 0 0 0 0 0 0 0 0 0 0 0 0
0 0 0 0
0 0 0 0 0 0 0 0 0 0 0 0
0 0 0
h 0 0 0
0 0 0 lt ( 0 ) o 0
0 1 0 0 0 1
0 0 0
2 r
h 0 0
0 0 1 0 0 1
(535)
Av and Bv are the rows and columns of the A and B corresponding to the components
O f Xy
The state L Q R weighting matrix Qv is given as
Qv
10 0 0 0 0 0 0 0 0 10 0 0 0 0 0 0 0 0 10 4 0 0 0 0 0 0 0 0 10 4 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1
(536)
mdash
Ry =w The state noise covariance matrix is taken
1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 9 x 10 4 0
0 0 0 0 0 0 0 9 x 10 4
(537)
119
while the measurement noise covariance matrix is
1 0 0 0 0 1 0 0 0 0 5 x 10 4 0
LO 0 0 5 x 10 4
(538)
As in the development of the longitudinal controller the transverse offset conshy
troller can be designed using the same two commands (lqrc and l try) wi th the input
matrices corresponding to the transverse subsystem with r = 50 The result is the
transverse compensator system given by
(539)
where xv is the state estimates of xv The singular values of the target transfer
function is compared with those achieved by the L Q G and L Q G L T R in Figure 5-
7(b) It is apparent that the recovery is not as good as that for the longitudinal case
again due to the non-minimal system Moreover the low value of r indicates that
only a lit t le recovery in the transverse subsystem is possible This suggests that the
robustness of the L Q G design is almost the maximum that can be achieved under
these conditions
Denoting f y = f j f j r xf = x^x^T and ud = udu ^ T the longishy
tudinal compensator can be combined to give
xf =
$d =
Aku
0
0 Akv
Cku o
Cu 0
Bku
0 B Xf
Xf = Ckxf
0
kv
Hf = Akxf - Bkyf
(540)
Vu Vv 0 Cv
Xf = CfXf
Defining a permutation matrix Pf such that Xj = PfX where X is the state vector
of the original nonlinear system given in Eq(32) the compensator and full nonlinear
120
system equations can be combined as
X
(541)
Ckxf
A block diagram representation of Eq(541) is shown in Figure 5-6
524 Simulation results
The dynamical response for the stationkeeping STSS with a tether length of 20
km and ini t ia l offset position of 1 m in each direction is presented in Figure 5-8 when
both the F L T attitude controller and the L Q G L T R offset controller are activated As
expected the offset controller is successful in quickly damping the elastic vibrations
in the longitudinal and transverse direction (Figure 5-8b) However from Figure 5-
8(a) it is clear that the presence of offset control requires a larger control moment
to regulate the attitude of the platform This is due to the additional torque created
by the larger moment arm around 2 m and 125 m in the inplane and out-of-plane
directions respectively introduced by the offset control In addition the control
moment for the platform is modulated by the tethers transverse vibration through
offset coupling However this does not significantly affect the l ibrational motion of
the tether whose thruster force remains relatively unchanged from the uncontrolled
vibration case
Final ly it can be concluded that the tether elastic vibration suppression
though offset control in conjunction with thruster and momentum-wheel attitude
control presents a viable strategy for regulating the dynamics tethered satellite sysshy
tems
121
STSS configuration 3-Body a1(0)=a2(0)=at(0)=2deg P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m6y(0)=52(0)=1m dx(0)=dy(0)=dz(0)=1m Stationkeeping I = 20km LQGLTR Vibration Control
Sat 1 Pitch and Roll Angles Tether Pitch and Roll Angles mdash 1 1 1 1 1 1 I 1 1 lt 1 ~ ^ 1 ^ l _
Time (Orbits) Time (Orbits) Sat 1 Roll Control Moment Tether Roll Control Thrust
(a) Time (Orbits) Time (Orbits)
Figure 5-8 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT attitude controller and LQGLTR offset vibration controller (a) attitude and libration controller reshysponse
122
STSS configuration 3-Body a1(0)=a2(0)=a(0)=2deg p1(0)=(32(0)=Q(0)=1deg ox(0)=14mSy(0)=5z(0)=1m dx(0)=dy(0)=dz(0)=1m Stationkeeping lt = 20km LQGLTR Vibration Control
Tether X Offset Position
Time (Orbits) Tether Y Tranverse Vibration 20r
Time (Orbits) Tether Y Offset Position
t o
1 2 Time (Orbits)
Tether Z Tranverse Vibration Time (Orbits)
Tether Z Offset Position
(b)
Figure 5-8
Time (Orbits) Time (Orbits)
Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT attitude controller and LQGLTR offset vibration controller (b) vibration and offset response
123
6 C O N C L U D I N G R E M A R K S
61 Summary of Results
The thesis has developed a rather general dynamics formulation for a multi-
body tethered system undergoing three-dimensional motion The system is composed
of multiple rigid bodies connected in a chain configuration by long flexible tethers
The tethers which are free to undergo libration as well as elastic vibrations in three
dimensions are also capable of deployment retrieval and constant length stationkeepshy
ing modes of operation Two types of actuators are located on the rigid satellites
thrusters and momentum-wheels
The governing equations of motion are developed using a new Order(N) apshy
proach that factorizes the mass matrix of the system such that it can be inverted
efficiently The derivation of the differential equations is generalized to account for
an arbitrary number of rigid bodies The equations were then coded in FORTRAN
for their numerical integration with the aid of a symbolic manipulation package that
algebraically evaluated the integrals involving the modes shapes functions used to dis-
cretize the flexible tether motion The simulation program was then used to assess the
uncontrolled dynamical behaviour of the system under the influence of several system
parameters including offset at the tether attachment point stationkeeping deployshy
ment and retrieval of the tether The study covered the three-body and five-body
geometries recently flown OEDIPUS system and the proposed BICEPS configurashy
tion It represents innovation at every phase a general three dimensional formulation
for multibody tethered systems an order-N algorithm for efficient computation and
application to systems of contemporary interest
Two types of controllers one for the attitude motion and the other for the
124
flexible vibratory motion are developed using the thrusters and momentum-wheels for
the former and the variable offset position for the latter These controllers are used to
regulate the motion of the system under various disturbances The attitude controller
is developed using the nonlinear Feedback Linearization Technique ( F L T ) and is based
on the nonlinear but rigid model of the system O n the other hand the separation
of the longitudinal and transverse frequencies from those of the attitude response
allows for the development of a linear optimal offset controller using the robust Linear
Quadratic Gaussian-Loop Transfer Recovery ( L Q G L T R ) method The effectiveness
of the two controllers is assessed through their subsequent application to the original
nonlinear flexible model
More important original contributions of the thesis which have not been reshy
ported in the literature include the following
(i) the model accounts for the motion of a multibody chain-type system undershy
going librational and in the case of tethers elastic vibrational motion in all
the three directions
(ii) the dynamics formulation is based on a recursive O(N) Lagrangian algorithm
that efficiently computes the systems generalized acceleration vector
(iii) the development of an attitude controller based on the Feedback Linearizashy
tion Technique for a multibody system using thrusters and momentum-wheels
located on each rigid body
(iv) the design of a three degree of freedom offset controller to regulate elastic v i shy
brations of the tether using the Linear Quadratic Gaussian and Loop Transfer
Recovery method ( L Q G L T R )
(v) substantiation of the formulation and control strategies through the applicashy
tion to a wide variety of systems thus demonstrating its versatility
125
The emphasis throughout has been on the development of a methodology to
study a large class of tethered systems efficiently It was not intended to compile
an extensive amount of data concerning the dynamical behaviour through a planned
variation of system parameters O f course a designer can easily employ the user-
friendly program to acquire such information Rather the objective was to establish
trends based on the parameters which are likely to have more significant effect on
the system dynamics both uncontrolled as well as controlled Based on the results
obtained the following general remarks can be made
(i) The presence of the platform products of inertia modulates the attitude reshy
sponse of the platform and gives rise to non-zero equilibrium pitch and roll
angles
(ii) Offset of the tether attachment point significantly affects the equil ibrium orishy
entation of the system as well as its dynamics through coupling W i t h a relshy
atively small subsatellite the tether dynamics dominate the system response
with high frequency elastic vibrations modulating the librational motion On
the other hand the effect of the platform dynamics on the tether response is
negligible As can be expected the platform dynamics is significantly affected
by the offset along the local horizontal and the orbit normal
(iii) For a three-body system deployment of the tether can destabilize the platform
in pitch as in the case of nonzero offset However the roll motion remains
undisturbed by the Coriolis force Moreover deployment can also render the
tether slack if it proceeds beyond a critical speed
(iv) Uncontrolled retrieval of the tether is always unstable in pitch It also leads
to high frequency elastic vibrations in the tether
(v) The five-body tethered system exhibits dynamical characteristics similar to
those observed for the three-body case As can be expected now there are
additional coupling effects due to two extra bodies a rigid subsatellite and a
126
flexible tether
(vi) Cartwheeling motion of the B I C E P S configuration can also be used to adshy
vantage in the deployment of the tether However exceeding a crit ical ini t ia l
cartwheeling rate can result in a large tension in the tether which eventually
causes the satellite to bounce back rendering the tether slack
(vii) Spinning the tether about its nominal length as in the case of O E D I P U S
introduces high frequency transverse vibrations in the tether which in turn
affect the dynamics of the satellites
(viii) The F L T based controller is quite successful in regulating the attitude moshy
tion of both the tether and rigid satellites in a short period of time during
stationkeeping deployment as well as retrieval phases
(ix) The offset controller is successful in suppressing the elastic vibrations of the
tether wi thin the allowable l imits of the offset mechanism ( plusmn 2 0 m)
62 Recommendations for Future Study
A s can be expected any scientific inquiry is merely a prelude to further efforts
and insight Several possible avenues exist for further investigation in this field A
few related to the study presented here are indicated below
(i) inclusion of environmental effects particularly the free molecular reaction
forces as well as interaction with Earths magnetic field and solar radiation
(ii) addition of flexible booms attached to the rigid satellites providing a mechashy
nism for energy dissipation
(iii) validation of the new three dimensional offset control strategies using ground
based experiments
(iv) animation of the simulation results to allow visual interpretation of the system
dynamics
127
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[18] Book W J Recursive Lagrangian Dynamics of Flexible Manipulator Arms The International Journal of Robotics Research V o l 3 No 3 1984 pp 87-101
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[21] Hollerbach J M A Recursive Lagrangian Formulation of Manipulator Dyshynamics and a Comparative Study of Dynamics Formulation Complexity IEEE Transactions on Systems Man and Cybernetics V o l SMC-10 No 11 Novemshyber 1980 pp 730-736
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[23] Rosenthal D E A n Order n Formulation for Robotic Systems The Journal of the Astronautical Sciences V o l 38 No 4 October-December 1990 pp 511-529
[24] Keat J E Mul t ibody System Order n Dynamics Formulation Based on Velocshyity Transform Method Journal of Guidance Control and Dynamics V o l 13 No 2 March -Apr i l 1990 pp 207-212
[25] Kurd i l a A J Menon R G and Sunkel J W Nonrecursive Order N Formushylation of Mul t ibody Dynamics Journal of Guidance Control and Dynamics V o l 16 No 5 September-October 1993 pp 838-844
[26] Ja in A Unified Formulation of Dynamics for Serial R ig id Mul t i body Systems Journal of Guidance Control and Dynamics V o l 14 No 3 May-June 1991 pp 531-542
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[28] Ja in A and Rodriguez G Recursive Flexible Mul t ibody Systems Dynamics Using Spatial Operators Journal of Guidance Control and Dynamics V o l 15 No 6 November-December 1992 pp 1453-1466
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[30] Pradhan S M o d i V J and Misra A K A n Order N Dynamics Formulation of Flexible Mul t ibody Systems in Tree Topology - The Lagrangian Approach AIAAAAS Astrodynamics Specialist Conference San Diego C A U S A July 1996 Paper No 96-3624 pp 480-490
[31] Banerjee A K Order-n Formulation of Extrusion of a Beam with Large Bendshying and Rotation Journal of Guidance Control and Dynamics V o l 15 No 1 January-February 1992 pp 121-127
[32] X u D M The Dynamics and Control of the Shuttle Supported Tethered Sub-satellite System P h D Thesis Department of Mechanical Engineering M c G i l l University November 1984
[33] X u D M Misra A K and M o d i V J On Vibra t ion Control of Tethered Satellite Systems Proceedings of theNASAJPL Workshop on Applications of Distributed Systems Theory to the Control of Large Space Structures Pasadena California July 1982 pp 317-327
[34] Pradhan S M o d i V J and Misra A K On the Inverse Control of the Tethered Satellite Systems The Journal of the Astronautical Sciences A A S Publications Office San Diego California U S A V o l 43 No 2 Apr i l - June 1995 pp 179-193
[35] K i m E and Vadal i S R Modeling Issues Related to Retrieval of Flexible Tethered Satellite Systems Journal of Guidance Control and Dynamics V o l 18 No 5 September-October 1995 pp 1169-1176
[36] W u S Chang C and Housner J M Finite Element Approach for Transient Analysis of Mul t ibody Systems Journal of Guidance Control and Dynamics V o l 15 No 4 July-August 1992 pp 847-854
[37] Pradhan S M o d i V J and Misra A K Simultaneous Control of Platform Att i tude and Tether Vibra t ion using Offset Strategy AIAAA AS Astrodynamics Specialist Conference San Diego C A U S A July 1996 Paper No 96-3570 pp 1-11
[38] Ruying F and Bainum P M Dynamics and Control of a Space Platform with a Tethered Subsatellite Journal of Guidance Control and Dynamics V o l 11 No 4 July-August 1988 pp 377-381
[39] M o d i V J Lakshmanan P K and Misra A K Offset Control of Tethered Satellite Systems Analysis and Experimental Verification Acta Astronautica V o l 21 No 5 1990 pp 283-294
[40] X u D M Misra A K and M o d i V J Thruster Augmented Act ive Cin t ro l of a Tethered Satellite System During its Retrieval Journal of Guidance Control and Dynamics V o l 9 No 6 November-December 1986 pp 663-672
[41] M o d i V J Pradhan S and Misra A K Off-Set Control of the Tethered Systems Using a Graph Theoretic Approach Acta Astronautica V o l 35 No
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6 1995 pp 373-384
[42] Pradhan S M o d i V J Bhat M S and Misra A K Mat r ix Method for Eigenstructure Assignment The Mult i - Input Case with Appl ica t ion Journal of Guidance Control and Dynamics V o l 17 No 5 September-October 1994 pp 983-989
[43] Pradhan S Planar Dynamics and Control of Tethered Satellite Systems P h D Thesis Department of Mechanical Engineering The University of Br i t i sh Columbia December 1994
[44] Bainum P M and Kumar V K Optimal-Control of the Shuttle-Tethered System Acta Astronautica V o l 7 No 12 1980 pp 1333-1348
[46] Hughes P C Spacecraft Att i tude Dynamics John Wi ley amp Sons New York U S A 1992 Chapter 2 pp 7-38
[47] Stengel R F Optimal Control and Estimation Dover Publications Inc New York U S A 1994 p 26
[48] Nayfeh A H and Mook D T Nonlinear Oscillation John Wi ley and Sons New York U S A 1979 pp 485-500
[49] He X and Powell J D Tether Damping in Space Journal of Guidance
Control and Dynamics V o l 13 No 1 1990 pp 104-112
[50] Meirovitch L Elements of Vibration Analysis M c G r a w - H i l l Inc New York U S A 1986 pp 255-256
[51] Greenwood D T Principles of Dynamics Prentice Ha l l Inc Englewood Cliffs New Jersey U S A 1965 pp 252-275
[52] Boyce W E and D i P r i m a R C Elemental Differential Equations and Boundshy
ary Value Problems John Wi ley amp Sons Inc New York U S A 1986 pp 424-431
[53] IMSL Library Reference Manual Vol 1 I M S L Inc Houston Texas U S A 1980 pp D G E A R 1 - D G E A R 9
[54] Gear C W Numerical Initial Value Problems in Ordinary Differential Equashy
tions Prentice-Hall Englewood Cliffs New Jersey U S A 1971 pp 158-166
[55] Char B W Geddes K O Gonnet G H Leong B L Monagan M B and Wat t S M First Leaves A Tutorial Introduction to Maple V Springer-Verlag New York U S A 1992
[56] Tsakalis K S and Ioannou P A Linear Time-Varying Systems Control and
Adaptation Prentice-Hall Inc Englewood Cliffs New Jersey U S A 1993 pp 148-180
[57] Su R O n the Linear Equivalents of Nonlinear Systems Systems and Control
Letter V o l 2 No 1 1982 pp 48-52
[58] M o d i V J Karray F and Chan J K On the Control of a Class of Flexible Manipulators Using Feedback Linearization Approach 42nd Congress of the International Astronautical Federation October 1991 Montreal Canada Paper No IAF-91-324
131
[59] K u o B C Automatic Control Systems Prentice-Hall Inc Englewood Cliffs New Jersey U S A 1987 pp 314-326
[60] Athans M The Role and Use of the Stochastic Linear-Quadrat ic-Guassian Problem in Control System Design IEEE Transactions on Automatic Control V o l A C - 1 6 No 6 December 1971 pp 529-552
[61] Maciejowski J M Multivariable Feedback Design Addison-Wesley Publ ishing Company Wokingham England 1989 Chapters 1-5
[62] Chiang R Y and Safonov M G Robust Control Toolbox Users Guide The M a t h Works Inc Natick Mass U S A 1992 pp 272-274
[63] Stein G and Athans M The L Q G L T R Procedure for Mult ivariable Feedback Control Design IEEE Transactions on Automatic Control V o l A C - 3 2 No 2 February 1987 pp 105-114
[64] Borisenko A I and Tarapov I E Vector and Tensor Analysis with Applicashytions Dover Publications Inc New York U S A 1979
[65] Lass H Vector and Tensor Analysis M c G r a w - H i l l Book Company Inc New York U S A 1950
132
A P P E N D I X I TENSOR R E P R E S E N T A T I O N OF T H E
EQUATIONS OF M O T I O N
11 Prel iminary Remarks
The derivation of the equations of motion for the multibody tethered system
involves the product of several matrices as well as the derivative of matrices and
vectors with respect to other vectors In order to concisely code these relationships
in FORTRAN the matrix equations of motion are expressed in tensor notation
Tensor mathematics in general is an extremely powerful tool which can be
used to solve many complex problems in physics and engineering[6465] However
only a few preliminary results of Cartesian tensor analysis are required here They
are summarized below
12 Mathematical Background
The representation of matrices and vectors in tensor notation simply involves
the use of indices referring to the specific elements of the matrix entity For example
v = vk (k = lN) (Il)
A = Aij (i = lNj = lM)
where vk is the kth element of vector v and Aj is the element on the t h row and
3th column of matrix A It is clear from Eq(Il) that exactly one index is required to
completely define an entire vector whereas two independent indices are required for
matrices For this reason vectors and matrices are known as first-order and second-
order tensors respectively Scaler variables are known as zeroth-order tensors since
in this case an index is not required
133
Matrix operations are expressed in tensor notation similar to the way they are
programmed using a computer language such as FORTRAN or C They are expressed
in a summation notation known as Einstein notation For example the product of a
matrix with a vector is given as
w = Av (12)
or in tensor notation W i = ^2^2Aij(vkSjk)
4-1 ltL3gt j
where 8jk is the Kronecker delta Dropping the summation symbol the matrix-vector
product can be expressed compactly as
Wi = AijVj = Aimvm (14)
Here j is a dummy index since it appears twice in a single term and hence can
be replaced by any other index Note that since the resulting product is a vector
only one index is required namely i that appears exactly once on both sides of the
equation Similarly
E = aAB + bCD (15)
or Ej mdash aAimBmj + bCimDmj
~ aBmjAim I bDmjCim (16)
where A B C D and E are second-order tensors (matrices) and a and b are zeroth-
order tensors (scalers) In addition once an expression is in tensor form it can be
treated similar to a scaler and the terms can be rearranged
The transpose of a matrix is also easily described using tensors One simply
switches the position of the indices For example let w = ATv then in tensor-
notation
Wi = AjiVj (17)
134
The real power of tensor notation is in its unambiguous expression of terms
containing partial derivatives of scalers vectors or matrices with respect to other
vectors For example the Jacobian matrix of y = f(x) can be easily expressed in
tensor notation as
di-dx-j-1 ( L 8 )
mdash which is a second-order tensor as expected If f(x) = Ax where A is constant then
f - a )
according to the current definition If C = A(x)B(x) then the time derivative of C
is given by
C = Ax)Bx) + A(x)Bx) (110)
or in tensor notation
Cij = AimBmj + AimBmj (111)
= XhiAimfcBmj + AimBmj^k)
which is more clear than its equivalent vector form Note that A^^ and Bmjk are
third-order tensors that can be readily handled in F O R T R A N
13 Forcing Function
The forcing function of the equations of motion as given by Eq(260) is
- u 1-SdM- dqtdPe 9 f i m F(q qt) = Mq - -q mdashq + mdashmdash - Qd (260)
This expression can be converted into tensor form to facilitate its implementation
into F O R T R A N source code
Consider the first term Mq From Eq(289) the mass matrix M is given by
T M mdash Rv MtRv which in tensor form is
Mtj = RZiM^R^j (112)
135
Taking the time derivative of M gives
Mij = qsRv
ni^MtnrnRv
mj + Rv
niMtnTn^Rv
mj + Rv
niMtnmRv
mjs) (113)
The second term on the right hand side in Eq(260) is given by
-JPdMu 1 bdquo T x
2Q ~cWq = 2Qs s r k Q r ^ ^
Expanding Msrk leads to
M src = RnskMtnmRmr + RnsMtnmkRmr + RnsMtnmRmrk- (L15)
Finally the potential energy term in Eq(260) is also expanded into tensor
form to give dPe = dqt dPe
dq dq dqt = dq^QP^
and since ^ = Rp(q)q then
Now inserting Eqs(I13-I17) into Eq(260) and rearranging the tensor form of the
forcing term can how be stated as
Fk(q^q^)=(RP
sk + R P n J q n ) ^ - - Q d k
+ QsQr | (^Rlks ~ 2 M t n m R m r
+ (^RnkMtnms ~ 2~RnsMtnmgtk^ R m r
(^RnkRmrs ~ ~2RnsRmrk^j
where Fk(q q t) represents the kth component of F
(118)
136
A P P E N D I X II R E D U C E D EQUATIONS OF M O T I O N
II 1 Prel iminary Remarks
The reduced model used for the design of the vibration controller is represented
by two rigid end-bodies capable of three-dimensional attitude motion interconnected
with a fixed length tether The flexible tether is discretized using the assumed-
mode method with only the fundamental mode of vibration considered to represent
longitudinal as well as in-plane and out-of-plane transverse deformations In this
model tether structural damping is neglected Finally the system is restricted to a
nominal circular orbit
II 2 Derivation of the Lagrangian Equations of Mot ion
Derivation of the equations of motion for the reduced system follows a similar
procedure to that for the full case presented in Chapter 2 However in this model the
straightforward derivation of the energy expressions for the entire system is undershy
taken ignoring the Order(N) approach discussed previously Furthermore in order
to make the final governing equations stationary the attitude motion is now referred
to the local LVLH frame This is accomplished by simply defining the following new
attitude angles
ci = on + 6
Pi = Pi (HI)
7t = 0
where and are the pitch and roll angles respectively Note that spin is not
considered in the reduced model hence 7J = 0 Substituting Eq(IIl) into Eq(214)
137
gives the new expression for the rotation matrix as
T- CpiSa^ Cai+61 S 0 S a + e i
-S 0 c Pi
(II2)
W i t h the new rotation matrix defined the inertial velocity of the elemental
mass drrii o n the ^ link can now be expressed using Eq(215) as
(215)
Since the tether length is assumed to be fixed T j f j and Tj$jlt^ of Eq(215) are
identically zero Also defining
(II3) Vi = I Pi i
Rdm- f deg r the reduced system is given by
where
and
Rdm =Di + Pi(gi)m + T^iSi
Pi(9i)m= [Tai9i T0gi T^] fj
(II4)
(II5)
(II6)
D1 = D D l D S v
52 = 31 + P(d2)m +T[d2
D3 = D2 + l2 + dX2Pii)m + T2i6x
Note that $ j for i = 2 is defined in Section 212 whereas for i = 1 and 3 it is
the null matrix Since only one mode is used for each flexible degree of freedom
5i = [6Xi6yi6Zi]T for i = 2
The kinetic energy of the system can now be written as
i=l Z i = l Jmi RdmRdmdmi (UJ)
138
Expanding Kampi using Eq(II4)
KH = miDD + 2DiP
i j gidmi)fji + 2D T ltMtrade^
+tf J P^(9i)Pi9i)d^l + 2mT j PFmTi^dm^ (II8)
where the integrals are evaluated using the procedure similar to that described in
Chapter 2 The gravitational and strain energy expressions are given by Eq(247)
and Eq(251) respectively using the newly defined rotation matrix T in place of
TV
Substituting the kinetic and potential energy expressions in
d ( 8Ke d(Ke - Pe) i bull i Q ^ 0 (II9)
d t dqred d(ired
with
Qred = [ algt l gt a 2 2 A 2 A 2 lt ^ 2 Q 3 3 ] (5-14)
and d2 regarded as a time-varying parameter the nonlinear non-autonomous equashy
tions of motion for the reduced system can finally be expressed as
MredQred + fred = reg- (5-15)
139
Rdm- inertial position of the t h link mass element drrii
RP transformation matrix relating qt and tf
Ru longitudinal L Q R input weighting matrix
RvRn Rd transformation matrices relating qt and q
S(f Mdc) generalized acceleration vector of coupled system q for the full
nonlinear flexible system
th uiirv l u w u u u maniA
th
T j i l ink rotation matrix
TtaTto control thrust in the pitch and roll direction for the im l ink
respectively
rQi maximum deploymentretrieval velocity of the t h l ink
Vei l ink strain energy
Vgi link gravitational potential energy
Rayleigh dissipation function arising from structural damping
in the ith link
di position vector to the frame F from the frame F_i bull
dc desired offset acceleration vector (i 1 l ink offset position)
drrii infinitesimal mass element of the t h l ink
dx^dy^dZi Cartesian components of d along the local vertical local horishy
zontal and orbit normal directions respectively
F-Q mdash
gi r igid and flexible position vectors of drrii fj + ltEraquolt5
ijk unit vectors 1 0 0 r 010-^ and 00 l r respectively
li length of the t h l ink
raj mass of the t h link
nfj total number of flexible modes for the i 1 l ink nXi + nyi + nZi
nq total number of generalized coordinates per link nfj + 7
nqq systems total number of generalized coordinates Nnq
ix
number of flexible modes in the longitudinal inplane and out-
of-plane transverse directions respectively for the i ^ link
number of attitude control actuators
- $ T
set of generalized coordinates for the itfl l ink which accounts for
interactions with adjacent links
qv---QtNT
set of coordinates for the independent t h link (not connected to
adjacent links)
rigid position of dm in the frame Fj
position of centre of mass of the itfl l ink relative to the frame Fj
i + $DJi
desired settling time of the j 1 attitude actuator
actuator force vector for entire system
flexible deformation of the link along the Xj yi and Zj direcshy
tions respectively
control input for flexible subsystem
Cartesian components of fj
actual and estimated state of flexible subsystem respectively
output vector of flexible subsystem
LIST OF FIGURES
1-1 A schematic diagram of the space platform based N-body tethered
satellite system 2
1-2 Associated forces for the dumbbell satellite configuration 3
1- 3 Some applications of the tethered satellite systems (a) multiple
communication satellites at a fixed geostationary location
(b) retrieval maneuver (c) proposed B I C E P S configuration 6
2- 1 Vector components of the i 1 and (i- l ) 1 chain links 15
2-2 Vector components in cylindrical orbital coordinates 20
2-3 Inertial position of subsatellite thruster forces 38
2- 4 Coupled force representation of a momentum-wheel on a rigid body 40
3- 1 Flowchart showing the computer program structure 45
3-2 Kinet ic and potentialenergy transfer for the three-body platform based
tethered satellite system (a) variation of kinetic and potential energy
(b) percent change in total energy of system 48
3-3 Kinet ic and potential energy transfer for the five-body platform based
tethered satellite system (a) variation of kinetic and potential energy
(b) percent change in total energy of system 49
3-4 Simulation results for the platform based three-body tethered
system originally presented in Ref[43] 51
3- 5 Simulation results for the platform based three-body tethered
system obtained using the present computer program 52
4- 1 Schematic diagram showing the generalized coordinates used to
describe the system dynamics 55
4-2 Stationkeeping dynamics of the three-body STSS configuration
without offset (a) attitude response (b) vibration response 57
4-3 Stationkeeping dynamics of the three-body STSS configuration
xi
with offset along the local vertical
(a) attitude response (b) vibration response 60
4-4 Stationkeeping dynamics of the three-body STSS configuration
with offset along the local horizontal
(a) attitude response (b) vibration response 62
4-5 Stationkeeping dynamics of the three-body STSS configuration
with offset along the orbit normal
(a) attitude response (b) vibration response 64
4-6 Stationkeeping dynamics of the three-body STSS configuration
wi th offset along the local horizontal local vertical and orbit normal
(a) attitude response (b) vibration response 67
4-7 Deployment dynamics of the three-body STSS configuration
without offset
(a) attitude response (b) vibration response 69
4-8 Deployment dynamics of the three-body STSS configuration
with offset along the local vertical
(a) attitude response (b) vibration response 72
4-9 Retrieval dynamics of the three-body STSS configuration
with offset along the local vertical and orbit normal
(a) attitude response (b) vibration response 74
4-10 Schematic diagram of the five-body system used in the numerical example 77
4-11 Stationkeeping dynamics of the five-body STSS configuration
without offset
(a) attitude response (b) vibration response 78
4-12 Deployment dynamics of the five-body STSS configuration
wi th offset along the local vertical
(a) attitude response (b) vibration response 81
4-13 Stationkeeping dynamics of the three-body B I C E P S configuration
x i i
with offset along the local vertical
(a) attitude response (b) vibration response 84
4-14 Cartwheeling dynamics with deployment for the three-body B I C E P S
with offset along the local vertical
(a) attitude response (b) vibration response 86
4-15 Spin dynamics (7 = l deg s ) of the three-body O E D I P U S configuration
with offset along the local vertical
(a) attitude response (b) vibration response 89
4- 16 Spin dynamics (7 = 10deg s ) of the three-body O E D I P U S configuration
wi th offset along the local vertical
(a) attitude response (b) vibration response 91
5- 1 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear rigid F L T controller with offset along the
local horizontal local vertical and orbit normal
(a) attitude and vibration response (b) control actuator time histories 98
5-2 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear flexible F L T controller with offset along the
local horizontal local vertical and orbit normal
(a) attitude and vibration response (b) control actuator time histories 101
5-3 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear rigid F L T controller with offset along
the local vertical
(a) attitude and vibration response (b) control actuator time histories 103
5-4 Deployment dynamics of the three-body STSS using the nonlinear
rigid F L T controller with offset along the local vertical
(a) attitude and vibration response (b) control actuator time histories 105
5-5 Retrieval dynamics of the three-body STSS using the non-linear
rigid F L T controller with offset along the local vertical and orbit normal
(a) attitude and vibration response (b) control actuator time histories 108
x i i i
5-6 Block diagram for the L Q G L T R estimator based compensator 115
5-7 Singular values for the L Q G and L Q G L T R compensator compared to target
return ratio (a) longitudinal design (b) transverse design 118
5-8 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear rigid F L T attitude controller and L Q G L T R
offset vibration controller (a) attitude and libration controller response
(b) vibration and offset response 122
xiv
A C K N O W L E D G E M E N T
Firs t and foremost I would like to express the most genuine gratitude to
my supervisor Prof V i n o d J M o d i whose unending guidance and encouragement
proved most invaluable during the course of my studies
I am also deeply indebted to Dr Satyabrata Pradhan for his limitless patience
and technical advice from which this work would not have been possible I would also
like to extend a word of appreciation towards Prof A r u n K Mis ra ( M c G i l l University)
for ini t iat ing my interest in space dynamics and control
Also I would like to thank my collegues Dr Sandeep Munshi M r Mathieu
Caron and M r Yuan Chen as well as Dr Mae Seto M r Gary L i m and M r Mark
Chu for their words of advice and helpful suggestions that heavily contributed to all
aspects of my studies
Final ly I would like to thank all my family and friends here in Vancouver and
back home in Montreal whose unending support made the past two years the most
memorable of my life
The research project was supported by the Natural Sciences and Engineering
Research Council s P G S - A Scholarship held by the author as well as by N S E R C grant
A-2181 held by Prof V J M o d i
xv
1 INTRODUCTION
11 Prel iminary Remarks
W i t h the ever changing demands of the worlds population one often wonders
about the commitment to the space program On the other hand the importance
of worldwide communications global environmental monitoring as well as long-term
habitation in space have demanded more ambitious and versatile satellites systems
It is doubtful that the Russian scientist Tsiolkovsky[l] when he first proposed the
uti l izat ion of the Earths gravity-gradient environment in the last century would have
envisioned the current and proposed applications of tethered satellite systems
A tethered satellite system consists of two or more subsatellites connected
to each other by long thin cables or tethers which are in tension (Figure 1-1) The
subsatellite which can also include the shuttle and space station are in orbit together
The first proposed use of tethers in space was associated with the rescue of stranded
astronauts by throwing a buoy on a tether and reeling it to the rescue vehicle
Prel iminary studies of such systems led to the discovery of the inherent instability
during the tether retrieval[2] Nevertheless the bir th of tethered systems occured
during the Gemini X I and X I I missions[3] in 1966 when a short tether (10-30m)
was used to generate the first artificial gravity environment in space (000015g) by
cartwheeling (spinning) the system about its center of mass
The mission also demonstrated an important use of tethers for gravity gradient
stabilization The force analysis of a simplified model illustrates this point Consider
the dumbbell satellite configuration orbiting about the centre of Ear th as shown in
Figure 1-2 Satellite 1 and satellite 2 are connected to each other by a tether with
1
Space Station (Satellite 1 )
2
Satellite N
Figure 1-1 A schematic diagram of the space platform based N-body tethered satellite system
2
Figure 1-2 Associated forces for the dumbbell satellite configuration
3
the whole system free to librate about the systems centre of mass C M The two
major forces acting on each body are due to the centrifugal and gravitational effects
Since body 1 is further away from the earth its gravitational force Fg^y is smaller
than that of body 2 ie Fgi lt Fg^ However its centrifugal force FCl is greater
then the centrifugal force experienced by the second satellite (FCl gt FC2) Moreover
for body 1 Fg^ lt Fc^ in contrast to Fg2 gt FC2 for body 2 Thus resolving the force
components along the tether there is an evident resultant tension force Ff present
in the tether Similarly adding the normal components of the two forces vectorially
results in a force FR which restores the system to its stable equil ibrium along the
local vertical These tension and gravity-gradient restoring forces are the two most
important features of tethered systems
Several milestone missions have flown in the recent past The U S A J a p a n
project T P E (Tethered Payload Experiment) [4] was one of the first to be launched
using a sounding rocket to conduct environmental studies The results of the T P E
provided support for the N A S A I t a l y shuttle based tethered satellite system referred
to as the TSS-1 mission[5] This experiment aimed to study the electrodynamic effects
of a shuttle borne satellite system with a conductive tether connection where an
electric current was induced in the tether as it swept through the Earth s magnetic
field Unfortunetely the two attempts in August 1992 and February 1996 resulted
in only partial success The former suffered from a spool failure resulting in a tether
deployment of only 256m of a planned 20km During the latter attempt a break
in the tethers insulation resulted in arcing leading to tether rupture Nevertheless
the information gathered by these attempts st i l l provided the engineers and scientists
invaluable information on the dynamic behaviour of this complex system
Canada also paved the way with novel designs of tethered system aimed
primari ly at the study of the ionosphere and Northern Lights (Aurora Borealis)
4
The O E D I P U S A and C (Observation of Electrified Distributions in the Ionospheric
P lasma - a Unique Strategy) missions[6] launched from a sounding rocket in January
1989 and November 1995 respectively provided insight into the complex dynamical
behaviour of two comparable mass satellites connected by a 1 km long spinning
tether
Final ly two experiments were conducted by N A S A called S E D S I and II
(Small Expendable Deployable System) [7] which hold the current record of 20 km
long tether In each of these missions a small probe (26 kg) was deployed from the
second stage of a D E L T A II rocket thus succesfully demonstrating the feasibility of
long tether deployment Retrieval of the tether which is significantly more difficult
has not yet been achieved
Several proposed applications are also currently under study These include the
study of earths upper atmosphere using probes lowered from the shuttle in a low earth
orbit ( L E O ) to the deployment and retrieval of satellites for servicing Even more
promising application concerns the generation of power for the proposed International
Space Station using conductive tethers The Canadian Space Agency (CSA) has also
proposed a dual-mass satellite system B I C E P S (Bl-static Canadian Experiment on
Plasmas in Space) to make simultaneous measurements at different locations in the
environment for correlation studies[8] A unique feature of the B I C E P S mission is
the deployment of the tether aided by the angular momentum of the system which
is ini t ia l ly cartwheeling about its orbit normal (Figure 1-3)
In addition to the three-body configuration multi-tethered systems have been
proposed for monitoring Earths environment in the global project monitored by
N A S A entitled Mission to Planet Ear th ( M T P E ) The proposed mission aims to
study pollution control through the understanding of the dynamic interactions be-
5
Retrieval of Satellite for Servicing
lt mdash
lt Satellite
Figure 1-3 Some applications of the tethered satellite systems (a) multiple comshymunication satellites at a fixed geostationary location (b) retrieval maneuver (c) proposed BICEPS configuration
6
tween the Earths atmosphere biosphere hydrosphere and cryosphere This wi l l be
accomplished with multiple tethered instrumentation payloads simultaneously soundshy
ing at different altitudes for spatial correlations Such mult ibody systems are considshy
ered to be the next stage in the evolution of tethered satellites Micro-gravity payload
modules suspended from the proposed space station for long-term experiments as well
as communications satellites with increased line-of-sight capability represent just two
of the numerous possible applications under consideration
12 Brief Review of the Relevant Literature
The design of multi-payload systems wi l l require extensive dynamic analysis
and parametric study before the final mission configuration can be chosen This
wi l l include a review of the fundamental dynamics of the simple two-body tethered
systems and how its dynamics can be extended to those of multi-body tethered system
in a chain or more generally a tree topology
121 Multibody 0(N) formulation
Many studies of two-body systems have been reported One of the earliest
contribution is by Rupp[9] who was the first to study planar dynamics of the simshy
plified Shuttle based Tethered Satellite System (STSS) His pioneering work led to
the discovery of pitch oscillation growth during the retrieval phase Later Baker[10]
advanced the investigation to the third dimension in addition to adding atmospheric
effects to the system A more complete dynamical model was later developed and
analyzed by M o d i and Misra[ l l 12] which included deployment and tether flexibility
A comprehensive survey of important developments and milestones in the area have
been compiled and reviewed by Mis ra and Modi[13] The major conclusions based on
the literature may be summarized as follows
bull the stationkeeping phase is stable
7
bull deployment can be unstable if the rate exceeds a crit ical speed
bull retrieval of the tether is inherently unstable
bull transverse vibrations can grow due to the Coriolis force induced during deshy
ployment and retrieval
Misra Amier and Modi[14] were one of the first to extend these results to the
three-satellite double pendulum case This simple model which includes a variable
length tether was sufficient to uncover the multiple equilibrium configurations of
the system Later the work was extended to include out-of-plane motion and a
preliminary reel-rate control law to regulate the librational motion[15] Kesmir i and
Misra[16] developed a general formulation for N-body tethered systems based on the
Lagrangian principle where three-dimensional motion and flexibility are accounted for
However from their work it is clear that as the number of payload or bodies increases
the computational cost to solve the forward dynamics also increases dramatically
Tradit ional methods of inverting the mass matrix M in the equation
ML+F = Q
to solve for the acceleration vector proved to be computationally costly being of the
order for practical simulation algorithms O(N^) refers to a mult ibody formulashy
tion algorithm whose computational cost is proportional to the cube of the number
of bodies N used It is therefore clear that more efficient solution strategies have to
be considered
Many improved algorithms have been proposed to reduce the number of comshy
putational steps required for solving for the acceleration vector Rosenthal[17] proshy
posed a recursive formulation based on the triangularization of the equations of moshy
tion to obtain an 0(N2) formulation He also demonstrates the use of a symbolic
manipulation program to reduce the final number of computations A recursive L a -
8
grangian formulation proposed by Book[18] also points out the relative inefficiency of
conventional formulations and proceeds to derive an algorithm which is also 0(N2)
A recursive formulation is one where the equations of motion are calculated in order
from one end of the system to the other It usually involves one or more passes along
the links to calculate the acceleration of each link (forward pass) and the constraint
forces (backward or inverse pass) Non-recursive algorithms express the equations of
motion for each link independent of the other Although the coupling terms st i l l need
to be included the algorithm is in general more amenable to parallel computing
A st i l l more efficient dynamics formulation is an O(N) algorithm Here the
computational effort is directly proportional to the number of bodies or links used
Several types of such algorithms have been developed over the years The one based
on the Newton-Euler approach has recursive equations and is used extensively in the
field of multi-joint robotics[1920] It should be emphasized that efficient computation
of the forward and inverse dynamics is imperative if any on-line (real-time) control
of the joint is to be achieved Hollerbach[21] proposed a recursive O(N) formulation
based on Lagranges equations of motion However his derivation was primari ly foshy
cused on the inverse dynamics and did not improve the computational efficiency of
the forward dynamics (acceleration vector) Other recursive algorithms include methshy
ods based on the principle of vir tual work[22] and Kanes equations of motion[23]
Keat[24] proposed a method based on a velocity transformation that eliminated the
appearance of constraint forces and was recursive in nature O n the other hand
K u r d i l a Menon and Sunkel[25] proposed a non-recursive algorithm based on the
Range-Space method which employs element-by-element approach used in modern
finite-element procedures Authors also demonstrated the potential for parallel comshy
putation of their non-recursive formulation The introduction of a Spatial Operator
Factorization[262728] which utilizes an analogy between mult ibody robot dynamics
and linear filtering and smoothing theory to efficiently invert the systems mass ma-
9
t r ix is another approach to a recursive algorithm Their results were further extended
to include the case where joints follow user-specified profiles[29] More recently Prad-
han et al [30] proposed a non-recursive Lagrangian algorithm for factorizing the mass
matrix leading to an O(N) formulation of the forward dynamics of the system A n
efficient O(N) algorithm where the number of bodies N in the system varies on-line
has been developed by Banerjee[31]
122 Issues of tether modelling
The importance of including tether flexibility has been demonstrated by several
researchers in the past However there are several methods available for modelling
tether flexibility Each of these methods has relative advantages and limitations
wi th regards to the computational efficiency A continuum model where flexible
motion is discretized using the assumed mode method has been proposed[3233] and
succesfully implemented In the majority of cases even one flexible mode is sufficient
to capture the dynamical behaviour of the tether[34] K i m and Vadali[35] proposed a
so-called bead model or lumped-mass approach that discretizes the tether using point
masses along its length However in order to accurately portray the motion of the
tether a significantly high number of beads are needed thus increasing the number of
computation steps Consequently this has led to the development of a hybrid between
the last two approaches[16] ie interpolation between the beads using a continuum
approach thus requiring less number of beads Finally the tether flexibility can also
be modelled using a finite-element approach[36] which is more suitable for transient
response analysis In the present study the continuum approach is adopted due to
its simplicity and proven reliability in accurately conveying the vibratory response
In the past the modal integrals arising from the discretization process have
been evaluted numerically In general this leads to unnecessary effort by the comshy
puter Today these integrals can be evaluated analytically using a symbolic in-
10
tegration package eg Maple V Subsequently they can be coded in F O R T R A N
directly by the package which could also result in a significant reduction in debugging
time More importantly there is considerable improvement in the computational effishy
ciency [16] especially during deployment and retrieval where the modal integral must
be evaluated at each time-step
123 Attitude and vibration control
In view of the conclusions arrived at by some of the researchers mentioned
above it is clear that an appropriate control strategy is needed to regulate the dyshy
namics of the system ie attitude and tether vibration First the attitude motion of
the entire system must be controlled This of course would be essential in the case of
a satellite system intended for scientific experiments such as the micro-gravity facilishy
ties aboard the proposed Internation Space Station Vibra t ion of the flexible members
wi l l have to be checked if they affect the integrity of the on-board instrumentation
Thruster as well as momentum-wheel approach[34] have been considered to
regulate the rigid-body motion of the end-satellite as well as the swinging motion
of the tether The procedure is particularly attractive due to its effectiveness over a
wide range of tether lengths as well as ease of implementation Other methods inshy
clude tension and length rate control which regulate the tethers tension and nominal
unstretched length respectively[915] It is usually implemented at the deployment
spool of the tether More recently an offset strategy involving time dependent moshy
tion of the tether attachment point to the platform has been proposed[3738] It
overcomes the problem of plume impingement created by the thruster control and
the ineffectiveness of tension control at shorter lengths However the effectiveness of
such a controller can become limited with an exceedingly long tether due to a pracshy
tical l imi t on the permissible offset motion In response to these two control issues a
hybrid thrusteroffset scheme has been proposed to combine the best features of the
11
two methods[39]
In addition to attitude control these schemes can be used to attenuate the
flexible response of the tether Tension and length rate control[12] as well as thruster
based algorithms[3340] have been proposed to this end M o d i et al [41] have sucshy
cessfully demonstrated the effectiveness of an offset strategy to damp undesired v i shy
brations Passive energy dissipative devices eg viscous dampers are also another
viable solution to the problem
The development of various control laws to implement the above mentioned
strategies has also recieved much attention A n eigenstructure assignment in conducshy
tion wi th an offset controller for vibration attenuation and momemtum wheels for
platform libration control has been developed[42] in addition to the controller design
from a graph theoretic approach[41] Also non-linear feedback methods such as the
Feedback Linearization Technique ( F L T ) that are more suitable for controlling highly
nonlinear non-autonomous coupled systems have also been considered[43]
It is important to point out that several linear controllers including the classhy
sic state feedback Linear Quadratic Regulator ( L Q R ) have received considerable
attention[3944] Moreover robust methods such as the Linear Quadratic Guas-
s ian Loop Transfer Recovery ( L Q G L T R ) method have also been developed and imshy
plemented on tethered systems[43] A more complete review of these algorithms as
well as others applied to tethered system has been presented by Pradhan[43]
13 Scope of the Investigation
The objective of the thesis is to develop and implement a versatile as well as
computationally efficient formulation algorithm applicable to a large class of tethered
satellite systems The distinctive features of the model can be summarized as follows
12
bull TV-body 0(N) tethered satellite system in a chain-type topology
bull the system is free to negotiate a general Keplerian orbit and permitted to
undergo three dimensional inplane and out-of-plane l ibrtational motion
bull r igid bodies constituting the system are free to execute general rotational
motion
bull three dimensional flexibility present in the tether which is discretized using
the continuum assumed-mode method with an arbitrary number of flexible
modes
bull energy dissipation through structural damping is included
bull capability to model various mission configurations and maneuvers including
the tether spin about an arbitrary axis
bull user-defined time dependent deployment and retrieval profiles for the tether
as well as the tether attachment point (offset)
bull attitude control algorithm for the tether and rigid bodies using thrusters and
momentum-wheels based on the Feedback Linearization Technique
bull the algorithm for suppression of tether vibration using an active offset (tether
attachment point) control strategy based on the optimal linear control law
( L Q G L T - R )
To begin with in Chapter 2 kinematics and kinetics of the system are derived
using the O(N) formulation methodology developed by Pradhan et al [30] Chapshy
ter 3 discusses issues related to the development of the simulation program and its
validation This is followed by a detailed parametric study of several mission proshy
files simulated as particular cases of the versatile formulation in Chapter 4 Next
the attitude and vibration controllers are designed and their effectiveness assessed in
Chapter 5 The thesis ends with concluding remarks and suggestions for future work
13
2 F O R M U L A T I O N OF T H E P R O B L E M
21 Kinematics
211 Preliminary definitions and the i ^ position vector
The derivation of the equations of motion begins with the definition of the
inertial position of an arbitrary link of the mult ibody system Let the i ^ link of the
system be free to translate and rotate in 3-D space From Figure 2-1 the position
vector Rdm^ to the mass element drrn on the i ^ link wi th reference to the inertial
frame FQ can be written as
Here D represents the inertial position of the t h body fixed frame Fj relative to FQ
ri = [xiyiZi]T is the rigid position vector of drrii wi th respect to F J and f(fi) is
the flexible deformation at fj also with respect to the frame Fj Bo th these vectors
are relative to Fj Note in the present model tethers (i even) are considered flexible
and X corresponds to the nominal position along the unstretched tether while a and
ZJ are by definition equal to zero On the other hand the rigid bodies referred to as
satellites (i odd) including the platform are taken to be rigid ie f(fj) = 0 T j is
the rotation matrix used to express body-fixed vectors with reference to the inertial
frame
212 Tether flexibility discretization
The tether flexibility is discretized with an arbitrary but finite number of
Di + Ti(i + f(i)) (21)
14
15
modes in each direction using the assumed-mode method as
Ui nvi
wi I I i = 1
nzi bull
(22)
where nXi nyi and n 2 ^ are the number of modes used in the X m and Z directions
respectively For the ] t h longitudinal mode the admissible mode shape function is
taken as 2 j - l
Hi(xiii)=[j) (23)
where l is the i ^ tether length[32] In the case of inplane and out-of-plane transverse
deflections the admissible functions are
ampyi(xili) = ampZixuk) = v ^ s i n JKXj
k (24)
where y2 is added as a normalizing factor In this case both the longitudinal and
transverse mode shapes satisfy the geometric boundary conditions for a simple string
in axial and transverse vibration
Eq(22) is recast into a compact matrix form as
ff(i) = $i(xili)5i(t) (25)
where $i(x l) is the matrix of the tether mode shapes defined as
0 0
^i(xiJi) =
^ X bull bull bull ^XA
0
0 0
0
G 5R 3 x n f
(26)
and nfj = nx- + nyi + nZi Note $J(XJJ) is a function of the spatial coordinate x
and not of yi or ZJ However it is also a function of the tether length parameter li
16
which is time-varying during deployment and retrieval For this reason care must be
exercised when differentiating $J(XJJ) with respect to time such that
d d d k) = Jtregi(Xi li) = ^ 7 $ t ( ^ i k)Xi + gf$i(xigt li)k- (2-7)
Since x represents the nominal rate of change of the position of the elemental mass
drrii it is equal to j (deploymentretrieval rate) Let t ing
( 3 d
dx~ + dT)^Xili^ ( 2 8 )
one arrives at
$i(xili) = $Di(xili)ii (29)
The time varying generalized coordinate vector 5 in Eq(25) is composed of
the longitudinal and transverse components ie
ltM) G s f t n f i x l (210)
where 5X^ 5y- and Sz^ are nx- x 1 ny x 1 and nz- x 1 size vectors and correspond to
$ X and $ Z respectively
213 Rotation angles and transformations
The matrix T j in Eq(21) represents a rotation transformation from the body
fixed frame Fj to the inertial frame FQ ie FQ = T j F j It is defined by the 3-2-1
ordered sequence of elementary rotations[46]
1 P i tch F[ = Cf ( a j )F 0
2 R o l l F = C f (3j)F = C7f (3j)Cf ( a j )F 0
3 Yaw F = C7( 7 i )F = C7(7l)cf^)Cf(ai)FQ = Q F 0
17
where
c f (A) =
-Sai Cai
0 0
and
C(7i ) =
0 -s0i] = 0 1 0
0 cpi
i 0 0 = 0
0
(211)
(212)
(213)
Here Eq(211) corresponds to a rotation of ctj (pitch) about the inertial axis ZQ (orbit
normal) of frame FQ resulting in a new intermediate frame F It is then followed by
a roll rotation fa about the axis y of F- giving a second intermediate frame F-
Final ly a spin rotation 7J about the axis z of F- is given resulting in the body
fixed frame Fj for the i lt l link Since all the three rotation matrices are orthogonal
it follows that FQ = C~1Fi = Cj F and hence
C CPi Cai ~ CH Sai + SrYi SPi Sai 5 7 i Sampi + Cli SPi Cci CPi Sai C1i Cai + 5 7 i SPi ~ S1 Ca + C 7 bull Sp Sai
(214) SliCPi CliCPi
Here Cx and Sx are abbreviations for cos(x) and sin(x) respectively
214 Inertial velocity of the ith link
mdash Lett ing pj = fj - f $j(5j and differentiating Eq(21) with respect to time gives
m- D j + T j f j + T^Si + T j pound + T j $ j 5 j (215)
Each of the terms appearing in Eq(215) must be addressed individually Since
fj represents the rigid body motion within the link it only has meaning for the
deployable tether and since yi = Zj = 0 for al l tether links (i even) fj = Ifi where 1
is the unit vector 1 0 0T For the case of rigid links (i odd) fj = 0
18
Evaluating the derivative of each angle in the rotation matrix gives
T^i9i = Taiai9i + Tppigi + T^fagi (216)
where Tx = J j I V Collecting the scaler angular velocity terms and defining the set
Vi = deg^gt A) 7i^ Eq(216) can be rewritten as
T i f t = Pi9i)fji (217)
where Pi(gi) is a column matrix defined by
Pi(9i)Vi = [Taigi^p^iT^gilffi (218)
Inserting Eq(29) and Eq(217) into Eq(215) and defining Sj = 2 + leads to
215 Cylindrical orbital coordinates
The Di term in Eq(219) describes the orbital motion of the t h l ink and
is composed of the three Cartesian coordinates DXi A ^ DZi However over the
cycle of one orbit DXi and Dyi vary dramatically by around the order of Earths
radius This must be avoided since large variations in the coordinates can cause
severe truncation errors during their numerical integration For this reason it is
more convenient to express the Cartesian components in terms of more stationary
variables This is readily accomplished using cylindrical coordinates
However it is only necessary to transform the first l inks orbital components
ie Lgti since only D is a generalized coordinates From Figure 2-2 it is apparent
(219)
that
(220)
19
Figure 2-2 Vector components in cylindrical orbital coordinates
20
Eq(220) can be rewritten in matrix form as
cos(0i) 0 0 s i n ( 0 i ) 0 0
0 0 1
Dri
h (221)
= DTlDSv
where Dri is the inplane radial distance 9 the true anomaly and DZl the out-of-
plane distance normal to the orbital plane The total derivative of D wi th respect
to time gives in column matrix form
1 1 1 1 J (222)
= D D l D 3 v
where [cos(^i) -Dnsm(8i) 0 ]
(223)
For all remaining links ie i = 2 N
cos(0i) - A - j S i n ^ i ) 0 sin(^i) D r i c o s ( 0 i ) 0
0 0 1
DTl = D D i = I d 3 i = 2N (224)
where is the n x n identity matrix and
2N (225)
216 Tether deployment and retrieval profile
The deployment of the tether is a critical maneuver for this class of systems
Cr i t i ca l deployment rates would cause the system to librate out of control Normally
a smooth sinusoidal profile is chosen (S-curve) However one would also like a long
tether extending to ten twenty or even a hundred kilometers in a reasonable time
say 3-4 orbits This is usually difficult or impossible to accomplish solely wi th one
S-curve profile Often a composite S-curve-Steady-S-curve profile is adopted Thus
21
the deployment scheme for the t h tether can be summarized as
k = Tr l - cos (^r(ti - to-)) 1 ltUlt h 2 r V A V
k = Vov tiilttiltt2i
k = IT j 1 - C deg S lkitl ^ ^ lt i lt ^
(226)
where to- tt are the ini t ia l and final times of the deployment or retrieval maneuver
respectively A t j = t - tQ mdash tt mdash t2- and the steady deployment velocity VQ is
calculated based on the continuity of l and l at the specified intermediate times t
and to For the case of
and
Ai = i(ti) - ^(t 0 -) = - ^ ( t 2 i ) (227)
2Ali + li(tf) -h(tQ)
V = lUli l^lL (228)
1 tf-tQ v y
A t = (229)
22 Kinetics and System Energy
221 Kinetic energy
The kinetic energy of the ith link is given by
Kei = lJm Rdrnfi^Ami (230)
Setting Eq(219) in a matrix-vector form gives the relation
4 ^ ^ PM) T i $ 4 TiSi]i[t (231)
22
where
m
k J
x l (232)
and nq = -nfj + 7 is the number of generalized coordinates in each link Inserting
Eq(231) into Eq(230) and integrating the kinetic energy for the t h l ink can be
rewritten as
1 T
(233)
where Mti is the links symmetric mass matrix
Mt =
m i D D T D D DDTPifgidmi) DDTTi J ^dm D n T Sidm rn
sym fP(9i)Pi(9i)dmi J P[g^T^idm J P ( pound ) T ^ m sym sym J ^f^idm f ^f^dmi sym sym sym J sfsidmi
pound Sfttiqxnq
(234)
and mi is the mass of the Ith link The kinetic energy for the entire system ie N
bodies can now be stated as
1 bull T bull 1 bull T Ke = J2 MtA = o laquo Mtqt 2 ^ -i
i=l (235)
where qt = qj^^ bull bull -qfNT a n d Mt is a block diagonal decoupled mass matrix
of the system with Mf- on its diagonal
(236) Mt =
0 0
0 Mt2
0
0 0
MH
0 0 0
0 0 0 bull MtN
222 Simplification for rigid links
The mass matrix given by Eq(234) simplifies considerably for the case of rigid
23
links and is given as
Mti
m D D l
T D D l DD^Piiffidmi) 0
sym 0
0
fP(fi)Pi(i)dmi 0 0 0
0 Id nfj
0 m~
i = 1 3 5 N
(237)
Moreover when mult iplying the matrices in the (11) block term of Eq(237) one
gets
D D l
1 D D l = 1 0 0 0 D n
2 0 (238) 0 0 1
which is a diagonal matrix However more importantly i t is independent of 6 and
hence is always non-singular ensuring that is always invertible
In the (12) block term of M ^ it is apparent that f fidmi is simply the
definition of the centre of mass of the t h rigid link and hence given by
To drrii mdash miff (239)
where fcmi is the position vector of the centre of mass of link i For the case of the
(22) block term a little more attention is required Expanding the integrand using
Eq(218) where ltj = fj gives
fj Ta Tafi fj T Q Tpfi f- T Q T 7 i f j
sym sym
Now since each of the terms of Eq(240) is a scaler or a 1 x 1 matrix it is therefore
equal to its trace ie sum of its diagonal elements thus
tr(ffTai
TTaifi) tr(ffTai
TTpfi) tr(ff Ta^T^fi)
fjT^Tpn rfTp^fi (240)
P(rj)Pj(fi) = syrri trifjTpTTpfi) t r ^ T ^ T ^
sym sym tr(ff T 7 T r 7 f j) J
(241)
Using two properties of the trace function[47] namely
tr(ABC) = tr(BCA) = trCAB) (242)
24
and
j tr(A) =tr(^J A^j (243)
where A B and C are real matrices it becomes quite clear that the (22) block term
of Eq(237) simplifies to
P(fi)Pii)dmi =
tr(Tai
TTaiImi) tTTai
TTpImi) t r (T Q r T 7 7 m ) trTaTTpImi) tr(Tp
TT7Jmi) sym sym sym
(244)
where Im- is an alternate form of the inertia tensor of the t h link which when
expanded is
Imlaquo = I nfi dm bull i sym
sym sym
J x^dmi J Xydmi XZdm sym J yi2dmi f yiZidm
sym J z^dmi J
xVi -L VH sym ijxxi + lyy IZZJ)2
(245)
Note this is in terms of moments of inertia and products of inertia of the t f l rigid
link A similar simplification can be made for the flexible case where T is replaced
with Qi resulting in a new expression for Im
223 Gravitational potential energy
The gravitational potential energy of the ith link due to the central force law
can be written as
V9i f dm f
-A mdash = -A bullgttradelti RdmA
dm
inn Lgti + T j ^ | (246)
where z = 3986 x 105 km 3s 2 (Earths gravitational constant) Expanding binomially
and retaining up to third order terms gives
Vgi ~ ~ D T 2D [J gfgidrrii + 23f T j gidm mdashDT
D2 i l I I Df Tt I grffdmiTiDi
(2-lt 7)
25
where D = D is the magnitude of the inertial position vector to the frame F The
integrals in the above expression are functions of the flexible mode shapes and can be
evaluated through symbolic manipulation using an existing package such as Maple
V In turn Maple V can translate the evaluated integrals into F O R T R A N and store
the code in a file Furthermore for the case of rigid bodies it can be shown quite
readily that
j gjgidrrii = j rjfidrrn = [IXXi + Iyy + IZZi)2 (248)
Similarly the other integrals in Eq(234) can be integrated symbolically
224 Strain energy
When deriving the elastic strain energy of a simple string in tension the
assumptions of high tension and low amplitude of transverse vibrations are generally
valid Consequently the first order approximation of the strain-displacement relation
is generally acceptable However for orbiting tethers in a weak gravitational-gradient
field neglecting higher order terms can result in poorly modelled tether dynamics
Geometric nonlinearities are responsible for the foreshortening effects present in the
tether These cannot be neglected because they account for the heaving motion along
the longitudinal axis due to lateral vibrations The magnitude of the longitudinal
oscillations diminish as the tether becomes shorter[35]
W i t h this as background the tether strain which is derived from the theory
of elastic vibrations[3248] is given by
(249)
where Uj V and W are the flexible deformations along the Xj and Z direction
respectively and are obtained from Eq(25) The square bracketed term in Eq(249)
represents the geometric nonlinearity and is accounted for in the analysis
26
The total strain energy of a flexible link is given by
Ve = I f CiEidVl (250)
Substituting the stress-strain relationship 0 = E(poundi the strain energy is now given
by
Vei=l-EiAi J l l d x h (251)
where E^ is the tethers Youngs modulus and A is the cross-sectional area The
tether is assumed to be uniform thus EA which is the flexural stiffness of the
tether is assumed to be constant Eq(251) is evaluated symbolically for arbitrary
number of modes
225 Tether energy dissipation
The evaluation of the energy dissipation due to tether deformations remains
a problem not well understood even today Here this complex phenomenon is repshy
resented through a simplified structural damping model[32] In addition the system
exhibits hysteresis which also must be considered This is accomplished using an
augmented complex Youngs modulus
EX^Ei+jEi (252)
where Ej^ is the contribution from the structural damping and j = The
augmented stress-strain relation is now given as
a = Epoundi = Ei(l+jrldi)ei (253)
where 77^ = EjE^ ltC 1 is defined as the structural damping coefficient determined
experimentally [49]
27
If poundi is a harmonic function with frequency UJQ- then
jet = (254) wo-
Substituting into Eq(253) and rearranging the terms gives
e + I mdash I poundi = o-i + ad (255)
where ad and a are the stress with and without structural damping respectively
Now the Rayleigh dissipation function[50] for the ith tether can be expressed as
(ii) Jo
_ 1 EjAiVdi rU (2-56^ 2 w 0 Jo
The strain rate ii is the time derivative of Eq(249) The generalized external force
due to damping can now be written as Qd = Q^ bull bull bull QdNT where
dWd
^ = - i p = 2 4 ( 2 5 7 )
= 0 i = 13 N
23 0(N) Form of the Equations of Mot ion
231 Lagrange equations of motion
W i t h the kinetic energy expression defined and the potential energy of the
whole system given by Pe = ]Cn=l( 5i+ ej) the equations of motion can be obtained
quite readily using the Lagrangian principle
where q is the set of generalized coordinates to be defined in a later section
Substituting Eqs(235247251 and 257) into Eq(258) leads to the familiar
matr ix form of the non-linear non-autonomous coupled equations of motion for the
28
system
Mq t)t+ F(q q t) = Q(q q t) (259)
where M(q t) is the nonlinear symmetric mass matrix and F(q q t) is the forcing
function which can be written in matrix form as
Q(ltfgt Qi 0 is the vector of non-conservative generalized forces including control inshy
puts acting on the system A detailed expansion of Eq(260) in tensor notation is
developed in Appendix I
Numerical solution of Eq(259) in terms of q would require inversion of the
mass matrix However M is a full matrix of size nqq x nqq where nqq = Nnq is the
total number of generalized coordinates Therefore direct inversion of M would lead
to a large number of computation steps of the order nqq^ or higher The objective
here is to develop an order-N O(N) algorithm for obtaining M _ 1 that minimizes
the computational effort This is accomplished through the following transformations
ini t ia l ly developed for the stationkeeping case by Pradhan et al [30]
232 Generalized coordinates and position transformation
During the derivation of the energy expressions for the system the focus was
on the decoupled system ie each link was considered independent of the others
Thus each links energy expression is also uncoupled However the constraints forces
between two links must be incorporated in the final equations of motion
Let
e Unqxl (261) m
be the set of coupled coordinates of the t h l ink such that q = qT q2 bull bull bull 0T i s the
29
systems generalized coordinates Let qt- be the set of auxiliary decoupled coordinates
defined by Eq(232) such that qt = qj^ qj^ Qtj^1 bull The only difference between
qtj and q is the presence of D against d is the position of F wi th respect to
the inertial frame FQ whereas d is defined as the offset position of F relative to and
projected on the frame It can be expressed as d = dXidyidZi^ For the
special case of link 1 D = d
From Figure 2-1 can written as
Di = A - l + T i _ i J i _ i + di) + T V i S i - i f t - i + dXi)8i- (262)
Denoting KAj = $j(Zj + dx- ) Eq(262) can be rewritten in summation form as
i ampi = H ( T i - 1 ^ + T i - l K A i - l - l + T j - l ^ - l ) bull ( 2 - 6 3 )
3 = 1
Introducing the index substitution k mdash j mdash 1 into Eq(263)
i-1 D = ^2 (Tfc_ilt4 + TkKAk6k + Tkilk) + T^di (264)
k=l
since d = D TQ = 1^ and IQ DQ SQ are null vectors Recasting Eq(264) in
matr ix form leads to i-1
Qt Jfe=l
where
and
For the entire system
T fc-1 0 TkKAk
0 0 0 0 0 0 0 0 0 0 0 0 -
r T i - i 0 0 0 0 0 0 0 0 0 G
0 0 0 1
Qt = RpQ
G R n lt x l
(265)
(266)
G Unq x l (267)
(268)
30
where
RP
R 1 0 0 R R 2 0 R RP
2 R 3
0 0 0
R RP R PN _ 1 RN J
(269)
Here RP is the lower block triangular matrix which relates qt and q
233 Velocity transformations
The two sets of velocity coordinates qti and qi are related by the following
transformations
(i) First Transformation
Differentiating Eq(262) with respect to time gives the inertial velocity of the
t h l ink as
D = A - l + Ti_iK + + K A i _ i ^ _ i
+ T j _ i ^ + Ti-ii + T i - i K A i - i ^ + T j - i K A ^ i ^ i (270)
(271) (
Defining K A D = dddXi+1KAi K A L = d d K A and hi = di+1 +lfi +KA^i
results in K A ^ i = K A D j i o ^ + K A L i _ i i _ i
= KADi_iiTdi + K A L j _ i ^ _ i
Inserting Eq(271) into Eq(270) and rearranging gives
= A _ + T i ^ d g + K A D ^ x ^ i ^ ^ + Pi-^hi-iH-i
+ T i _ i K A i _ i 5 _ i + T i_ i2 + K A L j _ i 5 j _ i 4 _ i (272)
Using Eq(272) to define recursively - D j _ i and applying the index substitution
c = j mdash 1 as in the case of the position transformation it follows that
Di = ^ T ^ i K D ^ ^ + Pk(hk)ffk + T f c K A f c ^ + TfcKLfc4 Jfc=l (273)
+ T j _ i K D j _ i lt i ii
31
where K D j _ i = + K A D ^ i ^ - i r 7 and K L j = i + K A L j 5 j Finally by following a
procedure similar to that used for the position transformation it can be shown that
for the entire system
t = Rvil (2-74)
where Rv is given by the lower block diagonal matrix
Here
Rv
Rf 0 0 R R$ 0 R Rl RJ
0 0 0
T3V nN-l
R bullN
e Mnqqxl
Ri =
T i _ i K D i _ i TiKAi T j K L j 0 0 0 0 bullo 0 0 0 0 0 0 0
e K n lt x l
RS
T i-1 0 0 o-0 0 0 0 0 0
0 0 0 1
G Mnq x l
(275)
(276)
(277)
(ii) Second Transformation
The second transformation is simply Eq(272) set into matrix form This leads
to the expression for qt as
(278)
where Id Pihi) T j K A j T j K L j 0 0 0 0 0 0
0 0
0 0
0 0
G $lnq x l
Thus for the entire system
(279)
pound = Rnjt + Rdq
= Vdnqq - R nrlRdi (280)
32
where
and
- 0 0 0 RI 0 0
Rn = 0 0
0 0
Rf 0 0 0 rgtd K2 0
Rd = 0 0 Rl
0 0 0
RN-1 0
G 5R n w x l
0 0 0 e 5 R N ^ X L
(281)
(282)
234 Cylindrical coordinate modification
Because of the use of cylindrical coorfmates for the first body it is necessary
to slightly modify the above matrices to account for this coordinate system For the
first link d mdash D = D ^ D S L thus
Qti = 01 =gt DDTigi (283)
where
D D T i
Eq(269) now takes the form
DTL 0 0 0 d 3 0 0 0 ID
0 0 nfj 0
RP =
P f D T T i R2 0 P f D T T i Rp
2 R3
iJfDTTi i^
x l
0 0 0
R PN-I RN
(284)
(285)
In the velocity transformation a similar modification is necessary since D =
DD^DS^ Mak ing the substitution it can be shown that Eq(276) and Eq(279) now
33
takes the form for i = 1
T)V _ pn _ r t j mdash rt^ mdash 0 0 0 0
0 0 0 0 0 0 0 0
(286)
The rest of the matrices for the remaining links ie i = 2 N are unchanged
235 Mass matrix inversion
Returning to the expression for kinetic energy substitution of Eq(274) into
Eq(235) gives the first of two expressions for kinetic energy as
Ke = -q RvT MtRv (287)
Introduction of Eq(280) into Eq(235) results in the second expression for Ke as
1X 2q (hnqq-RnrlRd Mt (idnqq - R nrlRd (288)
Thus there are two factorizations for the systems coupled mass matrix given as
M = RV MtRv
M (hnqq - RnrlRd Mt ( i d n q q - R n r 1 R d
(289)
(290)
Inverting Eq(290)
r-1 _ ( rgtd~^ M 1 = (Ra) l d n q q - Rn)Mf1 (Rd)~l(Idnqq - Rn) (291)
Since both Rd and Mt are block diagonal matrices their inverse is simply the inverse
34
of each block on the diagonal ie
Mr1
h 0 0
0 0
0 0
0 0 0
0
0
0
tN
(292)
Each block has the dimension nq x nq and is independent of the number of bodies
in the system thus the inversion of M is only linearly dependent on N ie an O(N)
operation
236 Specification of the offset position
The derivation of the equations of motion using the O(N) formulation requires
that the offset coordinate d be treated as a generalized coordinate However this
offset may be constrained or controlled to follow a prescribed motion dictated either by
the designer or the controller This is achieved through the introduction of Lagrange
multipliers[51] To begin with the multipliers are assigned to all the d equations
Thus letting f = F - Q Eq(259) becomes
Mq + f = P C A (293)
where A G K 3 - 7 V x l is the vector of Lagrange multipliers and Pc G s R n 9 x 3 A r is the
permutation matrix assigning the appropriate Aj to its corresponding d equation
Inverting M and pre-multiplying both sides by PcT gives
pcTM-pc
= ( dc + PdegTM-^f) (294)
where dc is the desired offset acceleration vector Note both dc and PcTM lf are
known Thus the solution for A has the form
A pcTM-lpc - l dc + PC1 M (295)
35
Now substituting Eq(295) into Eq(293) and rearranging the terms gives
q = S(f Ml) = -M~lf + M~lPc
which is the new constrained vector equation of motion with the offset specified by
dc Note that pre-multiplying M~l by P c T provides the rows of M _ 1 corresponding
to the d equations Similarly post-multiplying by Pc leads the columns of the matrix
corresponding to the d equations
24 Generalized Control Forces
241 Preliminary remarks
The treatment of non-conservative external forces acting on the system is
considered next In reality the system is subjected to a wide variety of environmental
forces including atmospheric drag solar radiation pressure Earth s magnetic field as
well as dissipative devices to mention a few However in the present model only
the effect of active control thrusters and momemtum-wheels is considered These
generalized forces wi l l be used to control the attitude motion of the system
In the derivation of the equations of motion using the Lagrangian procedure
the contribution of each force must be properly distributed among all the generalized
coordinates This is acheived using the relation
nu ^5 Qk=E Pern ^ (2-97)
m=l deg Q k
where Qk and qk are the kth elements of Q and ltf respectively[51] Fem is the
mth external force acting on the system at the location Rm from the inertial frame
and nu is the total number of actuators Derivation of the generalized forces due to
thrusters and momentum wheels is given in the next two subsections
pcTM-lpc (dc + PcTM- Vj (2-96)
36
242 Generalized thruster forces
One set of thrusters is placed on each subsatellite ie al l the rigid bodies
except the first (platform) as shown in Figure 2-3 They are capable of firing in the
three orthogonal directions From the schematic diagram the inertial position of the
thruster located on body 3 is given as
Defining
R3 = di + Tid2 + T2(hi + KA2lt52) + T 3 f c m 3
= d + dfY + T 3 f c m 3
df = Tjdj+i + Tj+1lj+ii + KAj+i5j+i)
the thruster position for body 5 is given as
(298)
(299)
R5 = di+ dfx + d 3 + T5fcm
Thus in general the inertial position for the thruster on the t h body can be given
(2100)
as i - 2
k=l i1 cm (2101)
Let the column matrix
[Q i iT 9 R-K=rR dqj lA n i
WjRi A R i (2102)
where i = 3 5 N and j mdash 1 2 i The vector derivative of R wi th respect to
the scaler components qk is stored in the kth column of Thus for the case of
three bodies Eq(297) for thrusters becomes
1 Q
Q3
3 -1
T 3 T t 2 (2103)
37
Satellite 3
mdash gt -raquo cm
D1=d1
mdash gt
a mdash gt
o
Figure 2-3 Inertial position of subsatellite thruster forces
38
where ft = 0TtaTt^T is the thrust vector for the t h l ink (tether i-1) Here
Ttn and Tta are the thrust forces along the inplane m and out-of-plane Z directions i Pi
respectively For the case of 5-bodies
Q
~Qgt1
0
0
( T ^ 2 ) (2104)
The result for seven or more bodies follows the pattern established by Eq(2104)
243 Generalized momentum gyro torques
When deriving the generalized moments arising from the momemtum-wheels
on each rigid body (satellite) including the platform it is easier to view the torques as
coupled forces From Figure 2-4 it is apparent that for the ith l ink the generalized
forces constituting the couple are
(2105)
where
Fei = Faij acting at ei = eXi
Fe2 = -Fpk acting at e 2 = C a ^ l
F e 3 = F7ik acting at e 3 = eyj
Expanding Eq(2105) it becomes clear that
3
Qk = ^ F e i
i=l
d
dqk (2106)
which is independent of Ri- Thus the moments Mma = 2FaieXi Mm^ = 2FpeXi
and Mmi = 2FlieVi Transforming the coordinates to the body fixed frame using
the rotation matrix the generalized force due to the momentum-wheels on link i can
39
Figure 2-4 Coupled force representation of a momentum-wheel on a rigid body
40
be written as
d d d Qi = Td bull ^ T i t M m a - Tik bull mdashT^Mmp + T ^ bull T J j M ^
QiMmi
where Mm = Mma M m Mmi T
lt3 T J ^ T - T i f c bull ^ T i pound T ^ - ^ T z j
(2107)
(2108)
Note combining the thruster forces and momentum-wheel torques a compact
expression for the generalized force vector for 5 bodies can be written as
Q
Q T 3 0 Q T 5 0
0 0 0
0 lt33T3 Qyen Q5
3T5 0
0 0 0 Q4T5 0
0 0 0 Q5
5T5
M M I
Ti 2 m 3
Quu (2109)
The pattern of Eq(2109) is retained for the case of N links
41
3 C O M P U T E R I M P L E M E N T A T I O N
31 Prel iminary Remarks
The governing equations of motion for the TV-body tethered system were deshy
rived in the last chapter using an efficient O(N) factorization method These equashy
tions are highly nonlinear nonautonomous and coupled In order to predict and
appreciate the character of the system response under a wide variety of disturbances
it is necessary to solve this system of differential equations Although the closed-
form solutions to various simplified models can be obtained using well-known analytic
methods the complete nonlinear response of the coupled system can only be obtained
numerically
However the numerical solution of these complicated equations is not a straightshy
forward task To begin with the equations are stiff[52] be they have attitude and
vibration response frequencies separated by about an order of magnitude or more
which makes their numerical integration suceptible to accuracy errors i f a properly
designed integration routine is not used
Secondly the factorization algorithm that ensures efficient inversion of the
mass matrix as well as the time-varying offset and tether length significantly increase
the size of the resulting simulation code to well over 10000 lines This raises computer
memory issues which must be addressed in order to properly design a single program
that is capable of simulating a wide range of configurations and mission profiles It
is further desired that the program be modular in character to accomodate design
variations with ease and efficiency
This chapter begins with a discussion on issues of numerical and symbolic in-
42
tegration This is followed by an introduction to the program structure Final ly the
validity of the computer program is established using two methods namely the conshy
servation of total energy and comparison of the response results for simple particular
cases reported in the literature
32 Numerical Implementation
321 Integration routine
A computer program coded in F O R T R A N was written to integrate the equashy
tions of motion of the system The widely available routine used to integrate the difshy
ferential equations D G E A R [ 5 3 ] is based on Gears predictor-corrector method[54]
It is well suited to stiff differential equations as i t provides automatic step-size adjustshy
ment and error-checking capabilities at each iteration cycle These features provide
superior performance over traditional 4^ order Runge-Kut ta methods
Like most other routines the method requires that the differential equations be
transformed into a first order state space representation This is easily accomplished
by letting
thus
X - ^ ^ f ^ ) =ltbullbull) (3 -2 )
where S(fMdc) is defined by Eq(296) Eq(32) represents the new set of 2nqq
first order differential equations of motion
322 Program structure
A flow chart representing the computer programs structure is presented in
Figure 3-1 The M A I N program is responsible for calling the integration routine It
also coordinates the flow of information throughout the program The first subroutine
43
called is INIT which initializes the system variables such as the constant parameters
(mass density inertia etc) and the ini t ial conditions I C (attitude angles orbital
motion tether elastic deformation etc) from user-defined data files Furthermore all
the necessary parameters required by the integration subroutine D G E A R (step-size
tolerance) are provided by INIT The results of the simulation ie time history of the
generalized coordinates are then passed on to the O U T P U T subroutine which sorts
the information into different output files to be subsequently read in by the auxiliary
plott ing package
The majority of the simulation running time is spent in the F C N subroutine
which is called repeatedly by D G E A R F C N generates the fv(Xi) vector defined in
Eq(32) It is composed of several subroutines which are responsible for the comshy
putation of all the time-varying matrices vectors and scaler variables derived in
Chapter 2 that comprise the governing equations of motion for the entire system
These subroutines have been carefully constructed to ensure maximum efficiency in
computational performance as well as memory allocation
Two important aspects in the programming of F C N should be outlined at this
point As mentioned earlier some of the integrals defined in Chapter 2 are notably
more difficult to evaluate manually Traditionally they have been integrated numershy
ically on-line (during the simulation) In some cases particularly stationkeeping the
computational cost is quite reasonable since the modal integrals need only be evalushy
ated once However for the case of deployment or retrieval where the mode shape
functions vary they have to be integrated at each time-step This has considerable
repercussions on the total simulation time Thus in this program the integrals are
evaluated symbolically using M A P L E V[55] Once generated they are translated into
F O R T R A N source code and stored in I N C L U D E files to be read by the compiler
Al though these files can be lengthy (1000 lines or more) for a large number of flexible
44
INIT
DGEAR PARAM
IC amp SYS PARAM
MAIN
lt
DGEAR
lt
FCN
OUTPUT
STATE ACTUATOR FORCES
MOMENTS amp THRUST FORCES
VIBRATION
r OFFS DYNA
ET MICS
Figure 3-1 Flowchart showing the computer program structure
45
modes they st i l l require less time to compute compared to their on-line evaluation
For the case of deployment the performance improvement was found be as high as
10 times for certain cases
Secondly the O(N) algorithm presented in the last section involves matrices
that contained mostly zero terms (RP Rv and Mplusmn) The sparse structure of these mashy
trices can be used advantageously to improve the computational efficiency by avoiding
multiplications of zero elements The result is a quick evaluation of fv(Xt) Moreshy
over there is a considerable saving of computer memory as zero elements are no
longer stored
Final ly the C O N T R O L subroutine which is part of F C N is designed to calshy
culate the appropriate control forces torques and offset dynamics which are required
to stabilize the librational ( A T T I T U D E ) and vibrational ( V I B R A T I O N ) motion of
the system
33 Verification of the Code
The simulation program was checked for its validity using two methods The
first verified the conservation of total energy in the absence of dissipation The second
approach was a direct comparison of the planar response generated by the simulation
program with those available in the literature A s numerical results for the flexible
three-dimensional A^-body model are not available one is forced to be content with
the validation using a few simplified cases
331 Energy conservation
The configuration considered here is that of a 3-body tethered system with
the following parameters for the platform subsatellite and tether
46
h =
1091430 -8135 328108 -8135 8646050 27116 328108 27116 8286760
kg bull m 2 platform inertia
kg bull m 2 end-satellite inertia 200 0 0
0 400 0 0 0 400
m = 90000 kg mass of the space station platform
m2 = 500 kg mass of the end-satellite
EfAt = 61645 N tether elastic stiffness
pt mdash 49 kg km tether linear density
The tether length is taken to remain constant at 10 km (stationkeeping case)
A n in i t ia l disturbance in pitch and roll of 2deg and 1deg respectively is given to both
the rigid bodies and the flexible tether In addition the tether is ini t ia l ly deflected
by 45 m in the longitudinal direction at its end and 05 m in both the inplane and
out-of-plane directions in the first mode The tether attachment point offset at the
platform (satellite 1) end is 1 m in all the three directions ie d2 = 11 lTm The
attachment point at the subsatellite (satellite 2) end is 1 m in the x direction ie
fcm3 = 10 0T
The variation in the kinetic and potential energies is presented for the above-
mentioned case in Figure 3-2(a) As seen from the plot there is a continuous mutual
exchange between the kinetic and potential energies however the variation in the
total energy remains zero (Figure 3-2b) It should be noted that after 1 orbit the
variation of both kinetic and potential energy does not return to 0 This is due to
the attitude and elastic motion of the system which shifts the centre of mass of the
tethered system in a non-Keplerian orbit
Similar results are presented for the 5-body case ie a double pendulum (Figshy
ure 3-3) Here the system is composed of a platform and two subsatellites connected
47
STSS configuration 3-Body a 1(0)=a 2(0)=a t(0)=2deg P1(0)=p2(0)=p(0)=1deg 5x(0)=45m 8y(0)=52(0)=05m dx(0)=dy(0)=d2(0)=1m Stationkeeping lt=10km
Variation of Kinetic and Potential Energy
(a) Time (Orbits) Percent Variation in Total Energy
OOEOh
-50E-12
-10E-11 I i i
0 1 2 (b) Time (Orbits)
Figure 3-2 Kinetic and potential energy transfer for the three-body platform based tethered satellite system (a) variation of kinetic and potential energy (b) percent change in total energy of system
48
STSS configuration 5-Body a 1(0)=a 2(0)=a t 1(0)=2deg a 3(0)=a t 2(0)=25deg P1(0)=p2(0)=p3(0)=pt1(0)=pt2(0)=1lt
5x(0)=5m 8y(0)=5z(0)=05m dx(0)=dy(0)=dz(0)=0 Stationkeeping 1 =12=1 Okm
(a)
20E6
10E6r-
00E0
-10E6
-20E6
Variation of Kinetic and Potential Energy
-
1 I I I 1 I I
I I
AP e ~_
i i 1 i i i i
00E0
UJ -50E-12 U J
lt
Time (Orbits) Percent Variation in Total Energy
-10E-11h
(b) Time (Orbits)
Figure 3 -3 Kinet ic and potential energy transfer for the five-body platform based tethered satellite system (a) variation of kinetic and potential enshyergy (b) percent change in total energy of system
49
in sequence by two tethers The two subsatellites have the same mass and inertia
ie 7713 = 7 7 i 2 a n d I3 = I21 a n d are separated by identical tether segments each 10
k m in length The platform has the same properties as before however there is no
tether attachment point offset present in the system (d = fcmi = 0)
In al l the cases considered energy was found to be conserved However it
should be noted that for the cases where the tether structural damping deployment
retrieval attitude and vibration control are present energy is no longer conserved
since al l these features add or subtract energy from the system
332 Comparison with available data
The alternative approach used to validate the numerical program involves
comparison of system responses with those presented in the literature A sample case
considered here is the planar system whose parameters are the same as those given
in Section 331 A n ini t ia l pitch disturbance of 2deg is imparted to both the platform
and the tether together with an ini t ial tether deformation of 08 m and 001 m in
the longitudinal and transverse directions respectively in the same manner as before
The tether length is held constant at 5 km and it is assumed to be connected to the
centre of mass of the rigid platform and subsatellite The response time histories from
Ref[43] are compared with those obtained using the present program in Figures 3-4
and 3-5 Here ap and at represent the platform and tether pitch respectively and B i
C i are the tethers longitudinal and transverse generalized coordinates respectively
A comparison of Figures 3-4 and 3-5 clearly shows a remarkable correlation
between the two sets of results in both the amplitude and frequency The same
trend persisted even with several other comparisons Thus a considerable level of
confidence is provided in the simulation program as a tool to explore the dynamics
and control of flexible multibody tethered systems
50
a p(0) = 2deg 0(0) = 2C
6(0) = 08 m
0(0) = 001 m
Stationkeeping L = 5 km
D p y = 0 D p 2 = 0
a oo
c -ooo -0 01
Figure 3 - 4 Simulation results for the platform based three-body tethered system originally presented in Ref[43]
51
STSS configuration 3-Body a1(0)=a2(0)=at(0)=2deg p1(0)=(32(0)=(3t(0)=0 8x(0)=08m 5y(0)=001m 5Z(0)=0 dx(0)=dy(0)=dz(0)=0 Stationkeeping l=5km
Satellite 1 Pitch Angle Tether Pitch Angle
i -i i- i i i i i i d
1 2 0 1 2 Time (Orbits) Time (Orbits)
Longitudinal Vibration
^MAAAAAAAWVWWWWVVVW
V bull i bull i I i J 00 05 10 15 20
Time (Orbits) Transverse Vibration
Time (Orbits)
Figure 3-5 Simulation results for the platform based three-body tethered system obtained using the present computer program
52
4 D Y N A M I C SIMULATION
41 Prel iminary Remarks
Understanding the dynamics of a system is critical to its design and develshy
opment for engineering application Due to obvious limitations imposed by flight
tests and simulation of environmental effects in ground based facilities space based
systems are routinely designed through the use of numerical models This Chapter
studies the dynamical response of two different tethered systems during deployment
retrieval and stationkeeping phases In the first case the Space platform based Tethshy
ered Satellite System (STSS) which consists of a large mass (platform) connected to
a relatively smaller mass (subsatellite) with a long flexible tether is considered The
other system is the O E D I P U S B I C E P S configuration involving two comparable mass
satellites interconnected by a flexible tether In addition the mission requirement of
spin about an arbitrary axis the tether axis for O E D I P U S - A C and cartwheeling or
spin about the orbit normal for the proposed B I C E P S mission is accounted for
A s mentioned in Chapter 2 the tether flexibility is modelled using the asshy
sumed mode discretization method Although the simulation program can account
for an arbitrary number of vibrational modes in each direction only the first mode
is considered in this study as it accounts for most of the strain energy[34] and hence
dominates the vibratory motion The parametric study considers single as well as
double pendulum type systems with offset of the tether attachment points The
systems stability is also discussed
42 Parameter and Response Variable Definitions
The system parameters used in the simulation unless otherwise stated are
53
selected as follows for the STSS configuration
1091430 -8135 328108 -8135 8646050 27116 328108 27116 8286760
h =
h = 200 0 0
0 400 0 0 0 400
bull m = 90000 kg (mass of the space platform)
kg bull m 2 (platform inertia) 760 J
kg bull m 2 (subsatellite inertia)
bull 7772 = 500 kg (mass of the subsatellite)
bull EtAt = 61645 N (tether stiffness)
bull Pt = 49 k g k m (tether density)
bull Vd mdash 05 (tether structural damping coefficient)
bull fcm^ = l 0 0 r m (tether attachment point at the subsatellite)
The response variables are defined as follows
bull ai0 satellite 1 (platform) pitch and roll angles respectively
bull CK2 2 satellite 2 (subsatellite) pitch and roll angles respectively
bull at fit- tether pitch and roll angles respectively
bull If tether length
bull d = dx dy dzT- tether attachment position relative to satellite 1 (platform)
mdash
bull S = 8X 8y Sz tether flexible generalized coordinates in the longitudinal x
inplane transverse y and out-of-plane transverse z directions respectively
The attitude angles and are measured with respect to the Local Ve r t i c a l -
Loca l Horizontal ( L V L H ) frame A schematic diagram illustrating these variable
is presented in Figure 4-1 The system is taken to be in a nominal circular orbit at
an altitude of 289 km with an orbital period of 903 minutes
54
Satellite 1 (Platform) Y a w
Figure 4-1 Schematic diagram showing the generalized coordinates used to deshyscribe the system dynamics
55
43 Stationkeeping Profile
To facilitate comparison of the simulation results a reference case based on
the three-body STSS system with zero offset (d = 0 r c m 3 = 0) and a tether length
of It = 20 km is considered first (Figure 4-2) The system is subjected to an ini t ial
disturbance in pitch and roll of 2deg and 1deg respectively to all the three bodies
In addition an ini t ia l longitudinal deflection of 14 m from the tethers unstretched
position is given together with a transverse inplane and out-of-plane deflection of 1
m at the tethers mid-length From the attitude response given in Figure 4-2(a) it is
apparent that the three bodies oscillate about their respective equilibrium positions
Since coupling between the individual rigid-body dynamics is absent (zero offset) the
corresponding librational frequencies are unaffected Satellite 1 (Platform) displays
amplitude modulation arising from the products of inertia The result is a slight
increase of amplitude in the pitch direction and a significantly larger decrease in
amplitude in the roll angle
Figure 4-2(b) clearly shows coupling of the tethers attitude motion with its
flexibility dynamics The tether vibrates in the axial direction about its equilibrium
position (at approximately 138 m) with two vibration frequencies namely 012 Hz
and 31 x 1 0 - 4 Hz The former is the first longitudinal flexible mode which is dissishy
pated through the structural damping while the latter arises from the coupling with
the pitch motion of the tether Similarly in the transverse direction there are two
frequencies of oscillations arising from the flexible mode and its coupling wi th the
attitude motion As expected there is no apparent dissipation in the transverse flexshy
ible motion This is attributed to the weak coupling between the longitudinal and
transverse modes of vibration since the transverse strain in only a second order effect
relative to the longitudinal strain where the dissipation mechanism is in effect Thus
there is very litt le transfer of energy between the transverse and longitudinal modes
resulting in a very long decay time for the transverse vibrations
56
STSS configuration 3-Body a1(0)=a2(0)=a(0)=2deg (31(0)=p2(0)=p(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=dz(0)=0 Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
0 1 2 3 4 5 0 1 2 3 4 5 Time (Orbits) Time (Orbits)
Satellite 2 Pitch Angle Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-2 Stationkeeping dynamics of the three-body STSS configuration withshyout offset (a) attitude response
57
STSS configuration 3-Body a1(0)=cx2(0)=cx(0)=2deg P1(0)=p2(0)=p(0)=1deg 5x(0)=14m8y(0)=82(0)=1m dx(0)=dy(0)=dz(0)=0 Stationkeeping lt=20km
Tether Longitudinal Vibration 16| 1 1 1 1 1 r-
Time (Orbits) Tether In-Plane Transverse Vibration
imdash 1 mdash 1 mdash 1 mdash 1 mdashr
r I I 1 1 1 I 1 1 1 1 I 1 bull ^ bull bull I I J 0 1 2 3 4 5
Time (Orbits) Tether Out-of-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-2 Stationkeeping dynamics of the three-body STSS configuration withshyout offset (b) vibration response
58
Introduction of the attachment point offset significantly alters the response of
the attitude motion for the rigid bodies W i t h a i m offset along the local vertical
at the platform (dx = 1 m) and fcm^ = 100 m coupling between the tether
and end-bodies is established by providing a lever arm from which the tether is able
to exert a torque that affects the rotational motion of the rigid end-bodies (Figure
4-3) Bo th the platform and subsatellite now oscillate at the same frequency as the
tether whose motion remains unaltered In the case of the platform there is also an
amplitude modulation due to its non-zero products of inertia as explained before
However the elastic vibration response of the tether remains essentially unaffected
by the coupling
Providing an offset along the local horizontal direction (dy = 1 m) results in
a more dramatic effect on the pitch and roll response of the platform as shown in
Figure 4-4(a) Now the platform oscillates about its new equilibrium position of
-90deg The roll motion is also significantly disturbed resulting in an increase to over
10deg in amplitude On the other hand the rigid body dynamics of the tether and the
subsatellite as well as the tether flexible motion remain the same (Figure 4-4b)
Figure 4-5 presents the response when the offset of 1 m at the platform end is
along the z direction ie normal to the orbital plane The result is a larger amplitude
pitch oscillation about the reference equilibrium position as shown in Figure 4-5(a)
while the roll equilibrium is now shifted to approximately 90deg Note there is little
change in the attitude motion of the tether and end-satellites however there is a
noticeable change in the out-of-plane transverse vibration of the tether (Figure 4-5b)
which may be due to the large amplitude platform roll dynamics
Final ly by setting a i m offset in all the three direction simultaneously the
equil ibrium position of the platform in the pitch and roll angle is altered by approxishy
mately 30deg (Figure 4-6a) However the response of the system follows essentially the
59
STSS configuration 3-Body a1(0)=cx2(0)=cxt(0)=2deg p1(0)=(32(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=1m dy(0)=dz(0)=0 Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
Time (Orbits) Time (Orbits) Tether Pitch Angle Tether Roll Angle
L bull I I bull bull lt bull bull bull raquo bull bull bull bull -I h I I I I i H 0 1 2 3 4 5 0 1 2 3 4 5
Time (Orbits) Time (Orbits) Satellite 2 Pitch Angle Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-3 Stationkeeping dynamics of the three-body STSS configuration with offset along the local vertical (a) attitude response
60
STSS configuration 3-Body a 1(0)=a 2(0)=a t(0)=2deg (31(0)=p2(0)=pt(0)=1deg 5x(0)=14m5y(0)=5z(0)=1m dx(0)=1mdy(0)=dz(0)=0 Stationkeeping lt=20km
Tether Longitudinal Vibration 1 6 i mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash mdash i mdash i mdash mdash mdash mdash mdash i mdash mdash lt -
Time (Orbits)
Tether In-Plane Transverse Vibration
y bull i i i i i i i i i i i i i i i i i 0 1 2 3 4 5
Time (Orbits)
Tether Out-of-Plane Transverse Vibration r 1 1 1 1 1 1 1 1 1 1 1 1 1 1 q
h i I i i I i i i i I i i i i 3
0 1 2 3 4 5 (b) Time (Orbits)
Figure 4-3 Stationkeeping dynamics of the three-body STSS configuration with offset along the local vertical (b) vibration response
61
STSS configuration 3-Body a1(0)=cx2(0)=o(0)=2deg (31(0)=p2(0)=p(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dz(0)=0 dy(0)=1m Stationkeeping lt=20km
Satellite 1 Pitch Angle i 1 i
Satellite 1 Roll Angle
1 2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 4 Time (Orbits)
Tether Roll Angle
CD
edT
1 2 3 Time (Orbits)
Satellite 2 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 2 Roll Angle
CD
C O
(a)
Figure 4-
1 2 3 4 Time (Orbits)
1 2 3 4 Time (Orbits)
4 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal (a) attitude response
62
STSS configuration 3-Body a1(0)=cx2(0)=ot(0)=2deg p1(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dz(0)=0 dy(0)=1m Stationkeeping lt=20km
Tether Longitudinal Vibration 16| 1 gt 1 bull 1 1 1
0 1 2 3 4 5
Time (Orbits) Tether In-Plane Transverse Vibration
i mdash 1 mdash 1 mdash 1 mdash 1 mdash i mdash mdash lt mdash 1 mdash 1 mdash i mdash mdash 1 mdash 1 mdash
V I I bull bull bull I i i i I i i 1 1 3
0 1 2 3 4 5 Time (Orbits)
Tether Out-of-Plane Transverse Vibration r 1 1 1 1 1 1 1 1 1 1 1 gt bull | q
V i 1 1 1 1 1 mdash 1 mdash 3
0 1 2 3 4 5 (b) Time (Orbits)
Figure 4-4 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal (b) vibration response
63
STSS configuration 3-Body a1(0)=o2(0)=at(0)=2deg (31(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=0 dz(0)=1m Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 4 Time (Orbits)
Tether Roll Angle
1 2 3 4 Time (Orbits)
Satellite 2 Pitch Angle
2 3 4 Time (Orbits)
Satellite 2 Roll Angle
(a)
Figure 4-
1 2 3 4 Time (Orbits)
2 3 4 Time (Orbits)
5 Stationkeeping dynamics of the three-body STSS configuration with offset along the orbit normal (a) attitude response
64
STSS configuration 3-Body a1(0)=cx2(0)=ot(0)=2deg P1(0)=p2(0)=(3(0)=1deg 8x(0)=14m8y(0)=82(0)=1m dx(0)=dy(0)=0 d2(0)=1m Stationkeeping lt=20km
Tether Longitudinal Vibration 1 6 i mdash r mdash i mdash mdash i mdash i mdash mdash i mdash i mdash mdash i mdash mdash i mdash 1 mdash 1 mdash i mdash mdash lt mdash lt mdash lt mdash r
Time (Orbits) Tether In-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-5 Stationkeeping dynamics of the three-body STSS configuration with offset along the orbit normal (b) vibration response
65
same trend with only minor perturbations in the transverse elastic vibratory response
of the tether (Figure 4-6b)
44 Tether Deployment
Deployment of the tether from an ini t ial length of 200 m to 20 km is explored
next Here the deployment length profile is critical to ensure the systems stability
It is not desirable to deploy the tether too quickly since that can render the tether
slack Hence the deployment strategy detailed in Section 216 is adopted The tether
is deployed over 35 orbits with the sinusoidal acceleration and deceleration occurring
over a 4 km length ( A ^ 2 ) with the remaining deployment at a constant velocity of
V0 = 146 ms
Initially the tether is stretched only slightly S = 1304 x 1 0 _ 3 0 0 m
with al l the other states of the system ie pitch roll and transverse displacements
remaining zero The response for the zero offset case is presented in Figure 4-7 For
the platform there is no longer any coupling with the tether hence it oscillates about
the equilibrium orientation determined by its inertia matrix However there is st i l l
coupling between the subsatellite and the tether since rcm^ mdash 100 m resulting
in complete domination of the subsatellite dynamics by the tether Consequently
the pitch motion ini t ial ly grows by almost 50deg but eventually subsides as the tether
deployment rate decreases It is interesting to note that the out-of-plane motion is not
affected by the deployment This is because the Coriolis force which is responsible
for the tether pitch motion does not have a component in the z direction (as it does
in the y direction) since it is always perpendicular to the orbital rotation (Q) and
the deployment rate (It) ie Q x It
Similarly there is an increase in the amplitude of the transverse elastic oscilshy
lations of the tether again due to the Coriolis effect (Figure 4-7b) Moreover as the
66
STSS configuration 3-Body a1(0)=a2(0)=cxt(0)=2deg P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m5y(0)=52(0)=1m dx(0)=dy(0)=d2(0)=1m Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
i i i i i i i i i i i 0 1 2 3 4 5 0 1 2 3 4 5
Time (Orbits) Time (Orbits) Tether Pitch Angle Tether Roll Angle
l i i i i -I h i i i i I i i i i l- i i i i l i i i i I i i- i i H 0 1 2 3 4 5 0 1 2 3 4 5
Time (Orbits) Time (Orbits) Satellite 2 Pitch Angle Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-6 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal local vertical and orbit normal (a) attitude response
67
STSS configuration 3-Body a1(0)=cc2(0)=a(0)=2deg p1(0)=p2(0)=(3t(0)=1deg 5x(0)=14m6y(0)=52(0)=1m dx(0)=dy(0)=d2(0)=1m Stationkeeping lt=20km
Tether Longitudinal Vibration 1 6 1 mdash i mdash i mdash i mdash mdash i mdash 1 mdash 1 mdash mdash mdash i mdash mdash 1 mdash 1 mdash 1 mdash i mdash 1 mdash 1 mdash 1 mdash 1 mdash r
0 1 2 3 4 5
Time (Orbits) Tether In-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-6 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal local vertical and orbit normal (b) vibration response
68
STSS configuration 3-Body a(0)=a2(0)=ot(0)=0 P1(0)=p2(0)=pt(0)=0 8X(0)=1304 x 103m 8y(0)=8z(0)=0 dx(0)=dy(0)=dz(0)=0 Deployment
l=02km to 20km in 35 Orbits
Tether Length Profile
E
Satellite 1 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 1 Roll Angle mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash [ mdash i mdash i mdash i mdash i mdash | mdash
1 2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 4 Time (Orbits)
Tether Roll Angle
CD
oo ax
i imdashimdashimdashImdashimdashimdashimdashr-
_ 1 _ L
1 2 3 4 Time (Orbits)
Satellite 2 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 2 Roll Angle
-50
(a) 1 2 3 4
Time (Orbits) 1 2 3 4
Time (Orbits)
Figure 4-7 Deployment dynamics of the three-body STSS configuration without offset (a) attitude response
69
STSS configuration 3-Body a1(0)=a2(0)=at(0)=0 p1(0)=p2(0)=(3t(0)=0 8X(0)=1304x 103m 8y(0)=52(0)=0 dx(0)=dy(0)=d2(0)=0 Deployment
l=02km to 20km in 35 Orbits
Tether Longitudinal Vibration
to
-02
(b)
Figure 4-7
2 3 Time (Orbits)
Tether In-Plane Transverse Vibration
0 1 1 bull 1 1
1 2 3 4 Time (Orbits)
Tether Out-of-Plane Transverse Vibration
5
1 1 1 1 1 i i | i i | ^
_ i i i i L -I
1 2 3 4 5 Time (Orbits)
Deployment dynamics of the three-body STSS configuration without offset (b) vibration response
70
tether elongates its longitudinal static equilibrium position also changes due to an
increase in the gravity-gradient tether tension It may be pointed out that as in the
case of attitude motion there are no out-of-plane tether vibrations induced
When the same deployment conditions are applied to the case of 1 m offset in
the x direction (Figure 4-8) coupling effects are re-introduced The platform pitch
exceeds -100deg during the peak deployment acceleration However as the tether pitch
motion subsides the platform pitch response returns to a smaller amplitude oscillation
about its nominal equilibrium On the other hand the platform roll motion grows
to over 20deg in amplitude in the terminal stages of deployment Longitudinal and
transverse vibrations of the tether remain virtually unchanged from the zero offset
case except for minute out-of-plane perturbations due to the platform librations in
roll
45 Tether Retrieval
The system response for the case of the tether retrieval from 20 km to 200 m
with an offset of 1 m in the x and z directions at the platform end is presented in
Figure 4-9 Here the Coriolis force induced during retrieval renders the tether unstable
(Figure 4-9a) Consequently the pitch motion of the platform is also destabilized
through coupling The z offset provides an additional moment arm that shifts the
roll equilibrium to 35deg however this degree of freedom is not destabilized From
Figure 4-9(b) it is apparent that the transverse mode of the tether vibration is also
disturbed producing a high frequency response in the final stages of retrieval This
disturbance is responsible for the slight increase in the roll motion for both the tether
and subsatellite
46 Five-Body Tethered System
A five-body chain link system is considered next To facilitate comparison with
71
STSS configuration 3-Body a(0)=a2(0)=at(0)=0 P1(0)=P2(0)=P(0)=0 5X(0)=1304 x 103m 8y(0)=8z(0)=0 dx(0)=1mdy(0)=d2(0)=0 Deployment
lt=02km to 20km in 35 Orbits
Tether Length Profile
E
Satellite 1 Pitch Angle
1 2 3 Time (Orbits)
Satellite 1 Roll Angle
bull bull bull i bull bull bull bull i
1 2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 Time (Orbits)
Tether Roll Angle
i i i i J I I i _
1 2 3 4 Time (Orbits)
Satellite 2 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 2 Roll Angle 3E-3h
D )
T J ^ 0 E 0 eg
CQ
-3E-3
(a) 1 2 3 4
Time (Orbits) 1 2 3 4
Time (Orbits)
Figure 4-8 Deployment dynamics of the three-body STSS configuration with offset along the local vertical (a) attitude response
7 2
STSS configuration 3-Body a1(0)=a2(0)=at(0)=0 p1(0)=P2(0)=(3t(0)=0 Sx(0)=1304 x 103m 5y(0)=5z(0)=0 dx(0)=1mdy(0)=dz(0)=0 Deployment
l=02km to 20km in 35 Orbits
Tether Longitudinal Vibration
Time (Orbits) Tether In-Plane Transverse Vibration
n 1 1 1 f
I I I _j I I I I I 1 1 1 1 1 2 3 4 5
Time (Orbits) Tether Out-of-Plane Transverse Vibration
J 1 I I 1 I I I I I I I I L _ _ I I d 1 2 3 4 5
Time (Orbits)
Deployment dynamics of the three-body STSS configuration with offset along the local vertical (b) vibration response
(b)
Figure 4-8
73
STSS configuration 3-Body a1(0)=a2(0)=at(0)=0 p(0)=P2(0)=(3t(0)=0 Sx(0)=14m8y(0)=82(0)=0 dy(0)=0dx(0)=d2(0)=1m Retrieval
l=20km to 02km in 35 Orbits
Satellite 1 Pitch Angle
(a) Time (Orbits)
Figure 4-9 Retrieval dynamics of the along the local vertical ar
Tether Length Profile
i i i i i i i i i i I 0 1 2 3
Time (Orbits) Satellite 1 Roll Angle
Time (Orbits) Tether Roll Angle
b_ i i I i I i i L _ J
0 1 2 3 Time (Orbits)
Satellite 2 Roll Angle _ mdash | r i i | 1 i i imdashmdashj i 1 r mdash i mdash
04- -
0 1 2 3 Time (Orbits)
three-body STSS configuration with offset d orbit normal (a) attitude response
74
STSS configuration 3-Body a1(0)=a2(0)=ocl(0)=0 (31(0)=f32(0)=pt(0)=0 8x(0)=14m8y(0)=6z(0)=0 dy(0)=0dx(0)=dz(0)=1m Retrieval
l=20km to 02km in 35 Orbits
Tether Longitudinal Vibration
E 10 to
Time (Orbits) Tether In-Plane Transverse Vibration
1 2 Time (Orbits)
Tether Out-of-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-9 Retrieval dynamics of the three-body STSS configuration with offset along the local vertical and orbit normal (b) vibration response
75
the three-body system dynamics studied earlier the chain is extended through the
addition of two bodies a tether and a subsatellite with the same physical properties
as before (Section 42) Thus the five-body system consists of a platform (satellite 1)
tether 1 subsatellite 1 (satellite 2) tether 2 and subsatellite 2 (satellite 3) as shown
in Figure 4-10 The offset of tether 1 at the platform-end is denoted by d2 while that
of tether 2 to subsatellite 1 as 04 In this numerical example fcm^ = fcm^ = 0 ie
tethers 1 and 2 are attached to the centre of mass of subsatellites 1 and 2 respectively
The system response for the case where the tether attachment points coincide
wi th the centres of mass of the rigid bodies ie zero offset (d2 = d mdash 0) is presented
in Figure 4-11 Here ogtt and at2 represent the pitch motion of tethers 1 and 2
respectively whereas 03 is the pitch motions of satellite 3 A similar convention
is adopted for the rol l angle 0 A s before the zero offset eliminates the coupling
between the tethers and satellites such that they are now free to librate about their
static equil ibrium positions However their is s t i l l mutual coupling between the two
tethers This coupling is present regardless of the offset position since each tether is
capable of transferring a force to the other The coupling is clearly evident in the pitch
response of the two tethers (Figure 4 - l l a ) O n the other hand the roll motion appears
uncoupled as the in i t ia l conditions in rol l for the two tethers are identical Thus the
motion is in phase and there is no transfer of energy through coupling A s expected
during the elastic response the tethers vibrate about their static equil ibrium positions
and mutually interact (Figure 4 - l l b ) However due to the variation of tension along
the tether (x direction) they do not have the same longitudinal equilibrium Note
relatively large amplitude transverse vibrations are present particularly for tether 1
suggesting strong coupling effects with the longitudinal oscillations
76
Figure 4-10 Schematic diagram of the five-body system used in the numerical example
77
STSS configuration 5-Body a1(0)=(x2(0)=at1(0)=2deg a3(0)=a12(0)=25deg p1(0)=r32(0)=p3(0)=pt1(0)=pt2(0)=r 82(0)=84(0)=505)05m d2(0)=d4(0)=000 Stationkeeping lt1=lu=10km Pitch Roll
Satellite 1 Pitch Angle Satellite 1 Roll Angle
o 1 2 Time (Orbits)
Tether 1 Pitch and Roll Angle
0 1 2 Time (Orbits)
Tether 2 Pitch and Roll Angle
Time (Orbits) Satellite 2 Pitch and Roll Angle
Time (Orbits) Satellite 3 Pitch and Roll Angle
Time (Orbits) Time (Orbits)
Figure 4-11 Stationkeeping dynamics of the five-body STSS configuration withshyout offset (a) attitude response
78
STSS configuration 5-Body a1(0)=a2(0)=oct1(0)=2deg a3(0)=at2(0)=25o
p1(0)=p2(0)=(33(0)=pt1(0)=pt2(0)=10
62(0)=54(0)=50505m d2(0)=d4(0)=000 Stationkeeping lt1=1 =1 Okm
Tether 1 Longitudinal Vibration Tether 2 Longitudinal Vibration i i 1 i 1 r 1 1 1 1
(b) Time (Orbits) Time (Orbits)
Figure 4-11 Stationkeeping dynamics of the five-body STSS configuration withshyout offset (b) vibration response
79
The system dynamics during deployment of both the tethers in the double-
pendulum configuration is presented in Figure 4-12 Deployment of each tether takes
place from 200 m to 20 km in 35 orbits The pitch motion of the system is similar
to that observed during the three-body case The Coriolis effect causes the tethers
to pitch through a large angle ( laquo 50deg ) which in turn disturbs the platform and
subsatellite due to the presence of offset along the local vertical On the other hand
the roll response for both the tethers is damped to zero as the tether deploys in
confirmation with the principle of conservation of angular momentum whereas the
platform response in roll increases to around 20deg Subsatellite 2 remains virtually
unaffected by the other links since the tether is attached to its centre of mass thus
eliminating coupling Finally the flexibility response of the two tethers presented in
Figure 4-12(b) shows similarity with the three-body case
47 BICEPS Configuration
The mission profile of the Bl-static Canadian Experiment on Plasmas in Space
(BICEPS) is simulated next It is presently under consideration by the Canadian
Space Agency To be launched by the Black Brant 2000500 rocket it would inshy
volve interesting maneuvers of the payload before it acquires the final operational
configuration As shown in Figure 1-3 at launch the payload (two satellites) with
an undeployed tether is spinning about the orbit normal (phase 1) The internal
dampers next change the motion to a flat spin (phase 2) which in turn provides
momentum for deployment (phase 3) When fully deployed the one kilometre long
tether will connect two identical satellites carrying instrumentation video cameras
and transmitters as payload
The system parameters for the BICEPS configuration are summarized below [59 0 0 1
bull I = I2 = 0 366 0 kg bull m 2 (satellite inertias) [ 0 0 392_
bull mi = 7772 = 200 kg (mass of the satellites)
80
STSS configuration 5-Body a(0)=cx2(0)=al1(0)=2deg a3(0)=c (0)=25o
P1(0)=P2(0)=p3(0)=f3t1(0)=Pt2(0)=r 52(0)=84(0)=1304x 10300m d2(0)=d4(0)=1G0m Deployment lt1=le=02km to 20km in 35 Orbits
Pitch Roll Satellite 1 Pitch Angle
Time (Orbits) Satellite 2 Pitch and Roll Angle T 1mdash1 1 1 I I I I | I I I 1 I | I I lt~r
Tether Length Profile i 1 1 1 1 i 1 1 1 1 i 1 1 1 1 i
i i i i i 0 1 2 3 4 5
Time (Orbits) Satellite 1 Roll Angle
0 1 2 3 4 5 Time (Orbits)
Tether 1 amp 2 Roll Angle L i | I I 1 I 1 1 -I
-Ih I- bull I I i i i i I i i i i 1
0 1 2 3 4 5 Time (Orbits)
Satellite 3 Pitch and Roll Angle i- i 1 1 i 1 1 1 1 i 1 1 1 1 i 1 1 1 1 i
L i i i i i i 1 i i mdash i mdash i I i imdashimdashLJ
0 1 2 3 4 5 Time (Orbits)
Figure 4-12 Deployment dynamics of the five-body S T S S configuration with offshyset along the local vertical (a) attitude response
81
STSS configuration 5-Body a1(0)=a2(0)=oct1(0)=2deg a3(0)=at2(0)=25deg P1(0)=P2(0)=P3(0)=Pt1(0)=pt2(0)=1o
82(0)=54(0)=1304 x 10-300m d2(0)=d4(0)=100m Deployment lt1=l2=02km to 20km in 35 Orbits
Tether 1 Longitudinal Vibration Tether 2 Longitudinal Vibration
Time (Orbits) Time (Orbits) Tether 1 Z Tranverse Vibration Tether 2 Z Tranverse Vibration
Figure 4-12 Deployment dynamics of the five-body STSS configuration with offshyset along the local vertical (b) vibration response
82
bull EtAt = 61645 N (tether stiffness)
bull pt = 30 k g k m (tether density)
bull It = 1 km (tether length)
bull rid = 1-0 (tether structural damping coefficient)
The system is in a circular orbit at a 289 km altitude Offset of the tether
attachment point to the satellites at both ends is taken to be 078 m in the x
direction The response of the system in the stationkeeping phase for a prescribed set
of in i t ia l conditions is given by Figure 4-13 As in the case of the STSS configuration
there is strong coupling between the tether and the satellites rigid body dynamics
causing the latter to follow the attitude of the tether However because of the smaller
inertias of the payloads there are noticeable high frequency modulations arising from
the tether flexibility Response of the tether in the elastic degrees of freedom is
presented in Figure 4-13(b) Note the longitudinal vibrations decay quite rapidly
(lt 05 orbits) due to the structural damping however it has vir tual ly no effect on
the transverse oscillations
The unique mission requirement of B I C E P S is the proposed use of its ini t ia l
angular momentum in the cartwheeling mode to aid in the deployment of the tethered
system The maneuver is considered next The response of the system during this
maneuver wi th an ini t ia l cartwheeling rate of 5 deg s is illustrated in Figure 4-14(a) The
in i t ia l tether length is taken to be 10 m and is deployed to 1 km There is an in i t ia l
increase in the pitch motion of the tether however due to the conservation of angular
momentum the cartwheeling rate decreases proportionally to the square of the change
in tether deployment rate The result is a rapid drop in the cartwheeling rate unti l
the system stops rotating entirely and simply oscillates about its new equilibrium
point Consequently through coupling the end-bodies follow the same trend The
roll response also subsides once the cartwheeling motion ceases However
83
BICEPS configuration 3-Body a1(0)=o2(0)=at(0)=2deg P1(0)=P2(0)=pt(0)=1deg 8X(0)=001 m 6y(0)=52(0)=01 m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping l=1 km
Satellite 1 Pitch Angle Satellite 1 Roll Angle - i 1 1 1 1 i - i q r 1 1 -gt
I- I i t i i i i I H 0 1 2 0 1 2
Time (Orbits) Time (Orbits) Tether Pitch Angle Tether Roll Angle
F mdash 1 mdash 1 mdash mdash 1 mdash i mdash 1 mdash 1 mdash mdash 1 mdash = i F mdash 1 mdash mdash mdash 1 mdash i mdash 1 mdash 1 mdash 1 mdash 1 mdash q
(a) Time (Orbits) Time (Orbits)
Figure 4-13 Stationkeeping dynamics of the three-body BICEPS configuration with offset along the local vertical (a) attitude response
84
BICEPS configuration 3-Body a1(0)=o2(0)=at(0)=2deg Pl(0)=p2(0)=(3t(0)=1deg 8X(0)=001 m 8y(0)=82(0)=01 m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping lt=1km
Tether Longitudinal Vibration 1E-2[
laquo = 5 E - 3 |
000 002
Time (Orbits) Tether In-Plane Transverse Vibration
Time (Orbits) Tether Out-of-Plane Transverse Vibration
Time (Orbits)
Figure 4-13 Stationkeeping dynamics of the three-body BICEPS configuration with offset along the local vertical (b) vibration response
85
BICEPS configuration 3-Body a1(0)=o2(0)=at(0)=2deg a1(0)=a2(0)=at(0)=57s Pl(0)=P2(0)=(3t(0)=1deg 8X(0)=115 x 103m 8y(0)=82(0)=0 dx(0)=078m dy(0)=d2(0)=0 Cartwheeling Deployment
lt=10m to 1km in 1 Orbit
Tether Length Profile
E
2000
cu T3
2000
co CD
2000
C D CD
T J
Satellite 1 Pitch Angle
1 2 Time (Orbits)
Satellite 1 Roll Angle
cn CD
33 cdl
Time (Orbits) Tether Pitch Angle
Time (Orbits) Tether Roll Angle
Time (Orbits) Satellite 2 Pitch Angle
1 2 Time (Orbits)
Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-14 Cartwheeling dynamics with deployment for the three-body BICEPS with offset along the local vertical (a) attitude response
86
BICEPS configuration 3-Body a1(0)=a2(0)=at(0)=2deg a (0)=a2(0)=a(0)=57s P1(0)=P2(0)=f3(0)=1deg 8X(0)=115x103m 8y(0)=8z(0)=0 dx(0)=078m dy(0)=d2(0)=0 Cartwheeling Deployment
l=1 Om to 1 km in 1 Orbit
5E-3h
co
0E0
025
000 E to
-025
002
-002
(b)
Figure 4-14
Tether Longitudinal Vibration
Time (Orbits) Tether In-Plane Transverse Vibration
Time (Orbits) Tether Out-of-Plane Transverse Vibration
Time (Orbits)
Cartwheeling dynamics with deployment for the three-body BICEPS with offset along the local vertical (b) vibration response
87
as the deployment maneuver is completed the satellites continue to exhibit a small
amplitude low frequency roll response with small period perturbations induced by
the tether elastic oscillations superposed on it (Figure 4-14b)
48 OEDIPUS Spinning Configuration
The mission entitled Observation of Electrified Distributions of Ionospheric
Plasmas - A Unique Strategy ( O E D I P U S ) also consists of a 1 k m tethered system with
two similar end-masses It was first flown in a suborbital flight in 1989 ( O E D I P U S A )
followed by a recent study in November 1995 ( O E D I P U S - C ) in which the University
of Br i t i sh Columbia was one of the participants Dynamically O E D I P U S represents a
unique system never encountered before Launched by a Black Brant rocket developed
at Br is to l Aerospace L t d the system ie the end-bodies together wi th the tether
spins about the longitudinal axis of the tether to achieve stabilized alignment wi th
Earth s magnetic field
The response of the system undergoing a spin rate of j = l deg s is presented
in Figure 4-15 using the same parameters as those outlined for the B I C E P S conshy
figuration in Section 47 Again there is strong coupling between the three bodies
(two satellites and tether) due to the nonzero offset The spin motion introduces an
additional frequency component in the tethers elastic response which is transferred
to the l ibrational motion of the satellites A s the spin rate increases to 1 0 deg s (Figure
4-16) the amplitude and frequency of the perturbations also increase however the
general character of the response remains essentially the same
88
OEDIPUS configuration 3-Body a1(0)=a2(0)=at(0)=2deg i(0)^[2(0H(0)=Ys p1(0)=p2(0)=p(0)=r 8X(0)=12 x 103m 8y(0)=82(0)=01m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping l=1 km
Satellite 1 Pitch Angle Satellite 1 Roll Angle mdash i 1 1 1 a r 1 1 1 r
(a) Time (Orbits) Time (Orbits)
Figure 4-15 Spin dynamics (7 = ldegs) of the three-body OEDIPUS configuration with offset along the local vertical (a) attitude response
89
OEDIPUS configuration 3-Body a1(0)=a2(0)=cct(0)=2deg Y1(0)=Y2(0)=Yt(0)=l7s (31(0)=p2(0)=pt(0)=1deg 8X(0)=12 x 103m 8y(0)=5z(0)=01 m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping lt=1 km
1E-2
to
OEO
E 00
to
Tether Longitudinal Vibration
005
Time (Orbits) Tether In-Plane Transverse Vibration
Time (Orbits) Tether Out-of-Plane Transverse Vibration
Time (Orbits)
Figure 4-15 Spin dynamics (7 = l deg s ) of the three-body O E D I P U S configuration with offset along the local vertical (b) vibration response
90
OEDIPUS configuration 3-Body a1(0)=a2(0)=a(0)=2deg Yi(0)=Y2(0)=rt(0)=107s 81(0)=p2(0)=pt(0)=1deg 8X(0)=12 x 103m 8y(0)=82(0)=01m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping lt=1km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-16 Spin dynamics (7= 10degs) of the three-body OEDIPUS configura tion with offset along the local vertical (a) attitude response
91
OEDIPUS configuration 3-Body a1(0)=a2(0)=ot(0)=2deg Yi(0)H(0)4(0)laquo107s P1(0)=P2(0)=pt(0)=1deg 8X(0)=12 x 103m 8y(0)=82(0)=01m dx(0)=078m dy(0)=dz(0)=0 Stationkeeping lt=1 km
Tether Longitudinal Vibration
Figure 4-16 Spin dynamics (7 = 10degs) of the three-body OEDIPUS configurashytion with offset along the local vertical (b) vibration response
92
5 ATTITUDE AND VIBRATION CONTROL
51 Att itude Control
511 Preliminary remarks
The instability in the pitch and roll motions during the retrieval of the tether
the large amplitude librations during its deployment and the marginal stability during
the stationkeeping phase suggest that some form of active control is necessary to
satisfy the mission requirements This section focuses on the design of an attitude
controller with the objective to regulate the librational dynamics of the system
As discussed in Chapter 1 a number of methods are available to accomplish
this objective This includes a wide variety of linear and nonlinear control strategies
which can be applied in conjuction with tension thrusters offset momentum-wheels
etc and their hybrid combinations One may consider a finite number of Linear T ime
Invariant (LTI) controllers scheduled discretely over different system configurations
ie gain scheduling[43] This would be relevant during deployment and retrieval of
the tether where the configuration is changing with time A n alternative may be
to use a Linear T ime Varying ( L T V ) model in conjuction with an adaptive control
scheme where on-line parametric identification may be used to advantage[56] The
options are vir tually limitless
O f course the choice of control algorithms is governed by several important
factors effective for time-varying configurations computationally efficient for realshy
time implementation and simple in character Here the thrustermomentum-wheel
system in conjuction with the Feedback Linearization Technique ( F L T ) is chosen as
it accounts for the complete nonlinear dynamics of the system and promises to have
good overall performance over a wide range of tether lengths It is well suited for
93
highly time-varying systems whose dynamics can be modelled accurately as in the
present case
The proposed control method utilizes the thrusters located on each rigid satelshy
lite excluding the first one (platform) to regulate the pitch and rol l motion of the
tether to which it is attached ie the thrusters located on satellite 2 (link 3) regshy
ulates the attitude motion of tether 1 (link 2) O n the other hand the l ibrational
motion of the rigid bodies (platform and subsatellite) is controlled using a set of three
momentum wheels placed mutually perpendicular to each other
The F L T method is based on the transformation of the nonlinear time-varying
governing equations into a L T I system using a nonlinear time-varying feedback Deshy
tailed mathematical background to the method and associated design procedure are
discussed by several authors[435758] The resulting L T I system can be regulated
using any of the numerous linear control algorithms available in the literature In the
present study a simple P D controller is adopted arid is found to have good perforshy
mance
The choice of an F L T control scheme satisfies one of the criteria mentioned
earlier namely valid over a wide range of tether lengths The question of compushy
tational efficiency of the method wi l l have to be addressed In order to implement
this controller in real-time the computation of the systems inverse dynamics must
be executed quickly Hence a simpler model that performs well is desirable Here
the model based on the rigid system with nonlinear equations of motion is chosen
Of course its validity is assessed using the original nonlinear system that accounts
for the tether flexibility
512 Controller design using Feedback Linearization Technique
The control model used is based on the rigid system with the governing equa-
94
tions of motion given by
Mrqr + fr = Quu (51)
where the left hand side represents inertia and other forces while the right hand side
is the generalized external force due to thrusters and momentum-wheels and is given
by Eq(2109)
Lett ing
~fr = fr~ QuU (52)
and substituting in Eq(296)
ir = s(frMrdc) (53)
Expanding Eq(53) it can be shown that
qr = S(fr Mrdc) - S(QU Mr6)u (5-4)
mdash = Fr - Qru
where S(QU Mr0) is the column matrix given by
S(QuMr6) = [s(Qu(l)Mr0) 5 (Q u ( 2 ) M r | 0 ) S(QU nu) M r | 0 ) ] (55)
and Qu-i) is the i ^ column of Qu Extract ing only the controlled equations from
Eq(54) ie the attitude equations one has
Ire = Frc ~~ Qrcu
(56)
mdash vrci
where vrc is the new control input required to regulate the decoupled linear system
A t this point a simple P D controller can be applied ie
Vrc = qrcd +KvQrcd-Qrc) + Kp(Qrcd~ Qrc)- (57)
Eq(57) represents the secondary controller with Eq(54) as the primary controller
Kp and Kv are the proportional and derivative gain matrices with qrcdi Qrcd
an(^ qrcd
95
as the desired acceleration velocity and position vector for the attitude angles of each
controlled body respectively Solving for u from Eq(56) leads to
u Qrc Frc ~ vrcj bull (58)
5 13 Simulation results
The F L T controller is implemented on the three-body STSS tethered system
The choice of proportional and derivative matrix gains is based on a desired settling
time of ts mdash 05 orbit and a damping factor of = 07 for both the pitch and roll
actuators in each body Given O and ts it can be shown that
-In ) 0 5 V T ^ r (59)
and hence kn bull mdash10
V (510)
where kp and kVj are the diagonal elements of the gain matrices Kp and Kv in
Eq(57) respectively[59]
Note as each body-fixed frame is referred directly to the inertial frame FQ
the nominal pitch equilibrium angle is zero only when measured from the L V L H frame
(OJJ mdash 0) However it is 6 when referred to FQ Hence the desired position vector
qrC(l is set equal to 0(t) ie the time-varying orbital angle for the pitch motion
and zero for the roll and yaw rotations such that
qrcd 37Vxl (511)
96
The desired velocity and acceleration are then given as
e U 3 N x l (512)
E K 3 7 V x l (513)
When the system is in a circular orbit qrc^ = 0
Figure 5-1 presents the controlled response of the STSS system defined in
Chapter 4 in the stationkeeping phase with offset d2 = 111T m and If = 20
km A s mentioned earlier the F L T controller is based on the nonlinear rigid model
Note the pitch and roll angles of the rigid bodies as well as of the tether are now
attenuated in less than 1 orbit (Figure 5-la) Consequently the longitudinal vibration
response of the tether is free of coupling arising from the librational modes of the
tether This leaves only the vibration modes which are eventually damped through
structural damping The pitch and roll dynamics of the platform require relatively
higher control moments Ma and Mpi respectively (Figure 5-lb) since they have to
overcome the extra moments generated by the offset of the tether attachment point
Furthermore the platform demands an additional moment to maintain its orientation
at zero pitch and roll angle since it is not the nominal equilibrium position The nonshy
zero static equilibrium of the platform arises due to its products of inertia On the
other hand the tether pitch and roll dynamics require only a small ini t ia l control
force of about plusmn 1 N which eventually diminishes to zero once the system is
97
and
o 0
Qrcj mdash
o o
0 0
Oiit) 0 0
1
N
N
1deg I o
STSS configuration 3-Body a1(0)=a2(0)=at(0)=2lt
P1(0)=p2(0)=pt(0)=1 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=dz(0)=1m Controlled Stationkeeping lt=20km
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
05 10 15 Time (Orbits)
Sat 2 Pitch and Roll Angles
20
CO CD
T J
cdT
cT
- mdash mdash 1 1 1 1 1 I - 1
Pitch -
A Roll -
A -
1 bull bull
-
1 Time (Orbits)
Tether Y Tranverse Vibration
1 2 Time (Orbits)
Tether Z Tranverse Vibration
gt to N
to
Time (Orbits) Time (Orbits) Longitudinal Vibration
145h
(a) Time (Orbits)
Figure 5-1 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local horizontal local vertical and orbit normal (a) attitude and vibration response
98
1
STSS configuration 3-Body a1(0)=ct2(0)=cct(0)=20
P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=d2(0)=1m Controlled Stationkeeping lt=20km
Sat 1 Pitch Control Moment Sat 1 Roll Control Moment
Figure 5-1 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local horizontal local vertical and orbit normal (b) control actuator time histories
99
stabilized Similarly for the end-body a very small control torque is required to
attenuate the pitch and roll response
If the complete nonlinear flexible dynamics model is used in the feedback loop
the response performance was found to be virtually identical with minor differences
as shown in Figure 5-2 For example now the pitch and roll response of the platform
is exactly damped to zero with no steady-state error as against the minor almost
imperceptible deviation for the case of the controller basedon the rigid model (Figure
5-la) In addition the rigid controller introduces additional damping in the transverse
mode of vibration where none is present when the full flexible controller is used
This is due to a high frequency component in the at motion that slowly decays the
transverse motion through coupling
As expected the full nonlinear flexible controller which now accounts the
elastic degrees of freedom in the model introduces larger fluctuations in the control
requirement for each actuator except at the subsatellite which is not coupled to the
tether since f c m 3 = 0 The high frequency variations in the pitch and roll control
moments at the platform are due to longitudinal oscillations of the tether and the
associated changes in the tension Despite neglecting the flexible terms the overall
controlled performance of the system remains quite good Hence the feedback of the
flexible motion is not considered in subsequent analysis
When the tether offset at the platform is restricted to only 1 m along the x
direction a similar response is obtained for the system However in this case the
steady-state error in the platforms rigid body motion is much smaller (Figure 5-3a)
In addition from Figure 5-3(b) the platforms control requirement is significantly
reduced
Figure 5-4 presents the controlled response of the STSS deploying a tether
100
STSS configuration 3-Body a1(0)=a2(0)=oct(0)=2deg p1(0)=P2(0)=p(0)=1deg 8x(0)=14m5y(0)=82(0)=1m dx(0)=dy(0)=dz(0)=1m Controlled Stationkeeping l=20km
Tether Pitch and Roll Angles
c n
T J
cdT
Sat 1 Pitch and Roll Angles 1 r
00 05 10 15 Time (Orbits)
Sat 2 Pitch and Roll Angles
c n CD
cdT a
o T J
CQ
- - I 1 1 1 I 1 1 1 1
P i t c h
R o l l
bull
1 -
0 1 2 Time (Orbits)
Tether Y Tranverse Vibration
0 1 2 Time (Orbits)
Tether Z Tranverse Vibration
E gt
10 N
CO
Time (Orbits) Time (Orbits) Longitudinal Vibration
145h
(a) Time (Orbits)
Figure 5-2 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear flexible FLT controller with offset along the local horizontal local vertical and orbit normal (a) attitude and vibration response
101
STSS configuration 3-Body a1(0)=a2(0)=cxt(0)=2deg 31(0)=p2(0)=p(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=d2(0)=1m Controlled Stationkeeping lt=20km
r i i i t i mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash mdash J 0 1 2 0 1 2
Time (Orbits) Time (Orbits) Tether Pitch Control Thrust Tether Roll Control Thruster
(b) Time (Orbits) Time (Orbits)
Figure 5-2 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear flexible FLT controller with offset along the local horizontal local vertical and orbit normal (b) control actuator time histories
102
STSS configuration 3-Body a1(0)=o2(0)=a(0)=2deg P1(0)=p2(0)=(3t(0)=1deg 8x(0)=14m5y(0)=6z(0)=1m dx(0)=1mdy(0)=dz(0)=0 Controlled Stationkeeping lt=20km
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
05 10 15 Time (Orbits)
Sat 2 Pitch and Roll Angles
CD
eg CQ
0 1 2 Time (Orbits)
Tether Y Tranverse Vibration
1 2 Time (Orbits)
Tether Z Tranverse Vibration
gt CO
Time (Orbits) Time (Orbits) Longitudinal Vibration
145h
(a) Time (Orbits)
Figure 5-3 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local vertical (a) attitude and vibration response
103
STSS configuration 3-Body a1(0)=o2(0)=al(0)=2o
P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=1mdy(0)=dz(0)=0 Controlled Stationkeeping lt=20km
Sat 1 Pitch Control Moment 11 1 1 r
Sat 1 Roll Control Moment
0 1 2 Time (Orbits)
Tether Pitch Control Thrust
0 1 2 Time (Orbits)
Sat 2 Pitch Control Moment
0 1 2 Time (Orbits)
Tether Roll Control Thruster
-07 o 1
Time (Orbits Sat 2 Roll Control Moment
1E-5
0E0F
Time (Orbits)
Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local vertical (b) control actuator time histories
104
STSS configuration 3-Body a1(0)=a2(0)=a(0)=2deg Pl(0)=p2(0)=pt(0)=1deg 6X(0)=1304 x 103m 8y(0)=82(0)=0 dx(0)=1mdy(0)=d2(0)=0 Controlled Deployment
l=02km to 20km in 35 Orbits
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
o 1 2 Time (Orbits)
Tether Y Tranverse Vibration
0 1 2 Time (Orbits)
Sat 2 Pitch and Roll Angles
CD CD
CM CO
Time (Orbits) Tether Z Tranverse Vibration
to
150
100
1 2 3 0 1 2
Time (Orbits) Time (Orbits) Longitudinal Vibration
to
(a)
50
2 3 Time (Orbits)
Figure 5 - 4 Deployment dynamics of the three-body STSS using the nonlinear rigid FLT controller with offset along the local vertical (a) attitude and vibration response
105
STSS configuration 3-Body a1(0)=oc2(0)=al(0)=20
P1(0)=p2(0)=pt(0)=1deg 8X(0)=1304 x 103m 5y(0)=52(0)=0 dx(0)=1mdy(0)=dz(0)=0 Controlled Deployment
l=02km to 20km in 35 Orbits
Tether Length Profile i 1
2 0 h
E 2L
Sat 1 Pitch Control Moment t mdash r mdash i mdash i mdash | mdash i mdash i mdash i mdash i mdash | mdash r mdash i mdash i mdash i mdash | mdash r mdash
0 5
1 2 3
Time (Orbits) Sat 1 Roll Control Moment
i i t | t mdash t mdash i mdash [ i i mdash r -
1 2
Time (Orbits) Tether Pitch Control Thrust
1 2 3 Time (Orbits)
Tether Roll Control Thruster
2 E - 5
1 2 3
Time (Orbits) Sat 2 Pitch Control Moment
1 2
Time (Orbits) Sat 2 Roll Control Moment
1 E - 5
E
O E O
Time (Orbits)
Deployment dynamics of the three-body STSS using the nonlinear rigid FLT controller with offset along the local vertical (b) control actuator time histories
106
from 02 km to 20 km in 35 orbits with a i m offset along the tether length A s before
the systems attitude motion is well regulated by the controller However the control
cost increases significantly for the pitch motion of the platform and tether due to the
Coriolis force induced by the deployment maneuver (Figure 5-4b) Initially there is a
sinusoidal increase in the pitch control requirement for both the tether and platform
as the tether accelerates to its constant velocity VQ Then the control requirement
remains constant for the platform at about 60 N m as opposed to the tether where the
thrust demand increases linearly Finally when the tether decelerates the actuators
control input reduces sinusoidally back to zero The rest of the control inputs remain
essentially the same as those in the stationkeeping case In fact the tether pitch
control moment is significantly less since the tether is short during the ini t ia l stages
of control However the inplane thruster requirement Tat acts as a disturbance and
causes 6y to nearly double from the uncontrolled value (Figure 4-8a) O n the other
hand the tether has no out-of-plane deflection
Final ly the effectiveness of the F L T controller during the crit ical maneuver of
retrieval from 20 km to 02 k m in 35 orbits is assessed in Figure 5-5 A s in the earlier
cases the system response in pitch and roll is acceptable however the controller is
unable to suppress the high frequency elastic oscillations induced in the tether by the
retrieval O f course this is expected as there is no active control of the elastic degrees
of freedom However the control of the tether vibrations is discussed in detail in
Section 52 The pitch control requirement follows a similar trend in magnitude as
in the case of deployment with only minor differences due to the addition of offset in
the out-of-plane direction ( d 2 = 1 0 1 ^ m) This is also responsible for the higher
roll moment requirement for the platform control (Figure 5-5b)
107
STSS configuration 3-Body a1(0)=X2(0)=a(0)=2deg P1(0)=p2(0)=pt(0)=1deg 5x(0)=14m8y(0)=52(0)=0 dx(0)=1mdy(0)=dz(0)=0 Controlled Retrieval
l=20km to 02km in 35 Orbits
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
0 1 2 Time (Orbits)
Sat 2 Pitch and Roll Angles
CO CD
cdT
CO CD
T J CM
CO
1 2 Time (Orbits)
Tether Y Tranverse Vibration
Time (Orbits) Tether Z Tranverse Vibration
150
100
1 2 3 0 1 2 Time (Orbits) Time (Orbits)
Longitudinal Vibration
CO
(a)
x 50
2 3 Time (Orbits)
Figure 5 -5 Retrieval dynamics of the three-body STSS using the non-linear rigid FLT controller with offset along the local vertical and orbit normal (a) attitude and vibration response
108
STSS configuration 3-Body a1(0)=ct2(0)=at(0)=2deg p1(0)=p2(0)=(3(0)=1deg 5x(0)=14m5y(0)=82(0)=0 dx(0)=1mdy(0)=dz(0)=0 Controlled Retrieval
l=20km to 02km in 35 Orbits
Tether Length Profile
E
1 2 Time (Orbits)
Tether Pitch Control Thrust
2E-5
1 2 Time (Orbits)
Sat 2 Pitch Control Moment
OEOft
1 2 3 Time (Orbits)
Sat 1 Roll Control Moment
1 2 3 Time (Orbits)
Tether Roll Control Thruster
1 2 Time (Orbits)
Sat 2 Roll Control Moment
1E-5 E z
2
0E0
(b) Time (Orbits) Time (Orbits)
Figure 5-5 Retrieval dynamics of the three-body STSS using the non-linear rigid FLT controller with offset along the local vertical and orbit normal (b) control actuator time histories
109
52 Control of Tethers Elastic Vibrations
521 Prel iminary remarks
The requirement of a precisely controlled micro-gravity environment as well
as the accurate release of satellites into their final orbit demands additional control
of the tethers vibratory motion in addition to its attitude regulation To that end
an active vibration suppression strategy is designed and implemented in this section
The strategy adopted is based on offset control ie time dependent variation
of the tethers attachment point at the platform (satellite 1) A l l the three degrees
of freedom of the offset motion are used to control both the longitudinal as well as
the inplane and out-of-plane transverse modes of vibration In practice the offset
controller can be implemented through the motion of a dedicated manipulator or a
robotic arm supporting the tether which in turn is supported by the platform The
focus here is on the control of elastic deformations during stationkeeping (ie fully
deployed tether fixed length) as it represents the phase when the mission objectives
are carried out
This section begins with the linearization of the systems equations of motion
for the reduced three body stationkeeping case This is followed by the design of
the optimal control algorithm Linear Quadratic Gaussian-Loop Transfer Recovery
( L Q G L T R ) based on the reduced model and its implementation on the full nonlinear
model Final ly the systems response in the presence of offset control is presented
which tends to substantiate its effectiveness
522 System linearization and state-space realization
The design of the offset controller begins with the linearization of the equations
of motion about the systems equilibrium position Linearization of extremely lengthy
(even in matrix form) highly nonlinear nonautonomous and coupled equations of
110
motion presents a challenging problem This is further complicated by the fact the
pitch angle ot is not referred to the L V L H frame Hence the equilibrium pitch angle
is not a constant but is equal to the instantaneous orbital angle 9(t) Two methods
are available to resolve this problem
One may use the non-stationary equations of motion in their present form and
derive a controller based on the Linear T ime Varying ( L T V ) system Alternatively
one can design a Linear T ime Invariant (LTI) controller based on a new set of reduced
governing equations representing the motion of the three body tethered system A l shy
though several studies pertaining to L T V systems have been reported[5660] the latter
approach is chosen because of its simplicity Furthermore the L T I controller design
is carried out completely off-line and thus the procedure is computationally more
efficient
The reduced model is derived using the Lagrangian approach with the genershy
alized coordinates given by
where lty and are the pitch and roll angles of link i relative to the L V L H frame
respectively 5X 8y and 5Z are the generalized coordinates associated with the flexible
tether deformations in the longitudinal inplane and out-of-plane transverse direcshy
tions respectively Only the first mode of vibration is considered in the analysis
The nonlinear nonautonomous and coupled equations of motion for the tethshy
ered system can now be written as
Qred = [ltXlPlltX2P2fixSy5za302gt] (514)
M redQred + fr mdash 0gt (515)
where Mred and frec[ are the reduced mass matrix and forcing term of the system
111
respectively They are functions of qred and qred in addition to the time varying offset
position d2 and its time derivatives d2 and d2 A detailed derivation of Eq(515)
is given in Appendix II A n additional consequence of referring the pitch motion to
a local frame is that the new reduced equations are now independent of the orbital
angle 9 and under the further assumption of the system negotiating a circular orbit
91 remains constant
The nonlinear reduced equations can now be linearized about their static equishy
l ibr ium position For all the generalized coordinates the equilibrium position is zero
wi th the exception of 5X which has a non-zero equilibrium 5XQ The offset position
d2) is linearized about its ini t ial position d2^ Linearizing Eq(515) and recasting
into matr ix form gives
Msqred + CsQred + KsQred + Md^2 + Cdd2 + Kdd2 + fs = 0 (516)
where Ms Cs Ks Md Cd Kd and s are constant matrices Mul t ip ly ing throughout
by Ms
1 gives
qr = -Ms 1Csqr - Ms
lKsqr
- Ms~lCdd22 - Ms-Kdd2 - Ms~Mdd2 - Mg~1 fs r-1 (517)
Defining ud = d2 and v = qTd^T Eq(517) can be rewritten as
Pu t t ing
v =
+
-M~lCs -M~Cd -1
0
-Mg~lMd
Id
0 v + -M~lKs -M~lKd
0 0
-M-lfs
0
MCv + MKv + Mlud +Fs
- 4 ) - ( gt the L T I equations of motion can be recast into the state-space form as
(518)
(519)
x =Ax + Bud + Fd (520)
112
where
A = MC MK Id12 o
24x24
and
Let
B = ^ U 5 R 2 4 X 3 -
Fd = I FJ ) e f t 2 4 1
mdash _ i _ -5
(521)
(522)
(523)
(524)
where x is the perturbation vector from the constant equilibrium state vector xeq
Substituting Eq(524) into Eq(520) gives the linear perturbation state equation
modelling the reduced tethered system as
bullJ x mdash x mdash Ax + Bud + (Fd + Axeq) (525)
It can be shown that for ideal control[60] Fd + Axeq = 0 thus giving the familiar
state-space equation
S=Ax + Bud (526)
The selection of the output vector completes the state space realization of
the system The output vector consists of the longitudinal deformation from the
equil ibrium position SXQ of the tether at xt = h the slope of the tether due to the
transverse deformation at xt = 0 and the offset position d2 from its in i t ia l position
d2Q Thus the output vector y is given by
K(o)sx
V
dx ~ dXQ dy ~ dyQ
d z - d z0 J
Jy
dx da XQ vo
d z - dZQ ) dy dyQ
(527)
113
where lt ( 0 ) = ddxi[$vxi)]Xi=o and lt ( 0 ) = ddxi[4gtw(xi)]Xi=0
523 Linear Quadratic Gaussian control with Loop Transfer Recovery
W i t h the linear state space model defined (Eqs526527) the design of the
controller can commence The algorithm chosen is the Linear Quadratic Gaussian
( L Q G ) estimator based optimal controller[6061] The L Q G is a widely used optimal
controller with its theoretical background well developped by many authors over the
last 25 years It involves of the design of the Kalman-Bucy Fi l ter ( K B F ) which
provides an estimate of the states x and a Linear Quadratic Regulator ( L Q R ) which
is separately designed assuming all the states x are known Both the L Q R and K B F
designs independently have good robustness properties ie retain good performance
when disturbances due to model uncertainty are included However the combined
L Q R and K B F designs ie the L Q G design may have poor stability margins in the
presence of model uncertainties This l imitat ion has led to the development of an L Q G
design procedure that improves the performance of the compensator by recovering
the full state feedback robustness properties at the plant input or output (Figure 5-6)
This procedure is known as the Linear Quadratic Guassian-Loop Transfer Recovery
( L Q G L T R ) control algorithm A detailed development of its theory is given in
Ref[61] and hence is not repeated here for conciseness
The design of the L Q G L T R controller involves the repeated solution of comshy
plicated matrix equations too tedious to be executed by hand Fortunenately the enshy
tire algorithm is available in the Robust Control Toolbox[62] of the popular software
package M A T L A B The input matrices required for the function are the following
(i) state space matrices ABC and D (D = 0 for this system)
(ii) state and measurement noise covariance matrices E and 0 respectively
(iii) state and input weighting matrix Q and R respectively
114
u
MODEL UNCERTAINTY
SYSTEM PLANT
y
u LQR CONTROLLER
X KBF ESTIMATOR
u
y
LQG COMPENSATOR
PLANT
X = f v ( X t )
COMPENSATOR
x = A k x - B k y
ud
= C k f
y
Figure 5-6 Block diagram for the LQGLTR estimator based compensator
115
A s mentionned earlier the main objective of this offset controller is to regu-mdash
late the tether vibration described by the 5 equations However from the response
of the uncontrolled system presented in Chapter 4 it is clear that there is a large
difference between the magnitude of the librational and vibrational frequencies This
separation of frequencies allow for the separate design of the vibration and attitude mdash mdash
controllers Thus only the flexible subsystem composed of the S and d equations is
required in the offset controller design Similarly there is also a wide separation of
frequencies between the longitudinal and transverse modes of vibrations permitt ing mdash
the decoupling of the flexible subsystem into a set of longitudinal (Sx and dx) and
transverse (5y Sz dy and dz) subsystems The appended offset system d must also
be included since it acts as the control actuator However it is important to note
that the inplane and out-of-plane transverse modes can not be decoupled because
their oscillation frequencies are of the same order The two flexible subsystems are
summarized below
(i) Longitudinal Vibra t ion Subsystem
This is defined by u mdash Auxu + Buudu
(528)
where xu = SxdX25xdX2T u d u = dX2 and from Eq(527)
0 0 $u(lt) 0 0 0 0 1
0 0 1 0 0 0 0 1
(529)
Au and Bu are the rows and columns of A and B respectively and correspond to the
components of xu
The L Q R state weighting matrix Qu is taken as
Qu =
bull1 0 0 o -0 1 0 0 0 0 1 0
0 0 0 10
(530)
116
and the input weighting matrix Ru = 1^ The state noise covariance matr ix is
selected as 1 0 0 0 0 1 0 0 0 0 4 0
LO 0 0 1
while the measurement noise covariance matrix is taken as
(531)
i o 0 15
(532)
Given the above mentionned matrices the design of the L Q G L T R compenshy
sator is computed using M A T L A B with the following 2 commands
(i) kfu = l q r c ( ^ ( i C bdquo d i a g m x ( E u e u ) )
(ii) [AkugtBkugtCkugtDku =^ryAuBuCuDukfuQuRur)
where r is a scaler
The first command is used to compute the Ka lman filter gain matrix kfu Once
kfu is known the second command is invoked returning the state space representation
of the L Q G L T R dynamic compensator to
(533) xu mdash Akuxu Bkuyu
where xu is the state estimate vector oixu The function l t ry performs loop transfer
recovery at the system output ie the return ratio at the output approaches that
of the K B F loop given by Cu(sld - Au)Kfu[61 This is achieved by choosing a
sufficently large scaler value r in l t ry such that the singular values of the return
ratio approach those of the target design For the longitudinal controller design
r mdash 5 x 10 5 Figure 5-7(a) compares the singular values of the recovered compensator
design and the non-recovered L Q G compensator design with respect to the target
Unfortunetaly perfect recovery is not possible especially at higher frequencies
117
100
co
gt CO
-50
-100 10
Longitudinal Compensator Singular Values i I I I N I I mdash i mdash i i i i n i i mdash i mdash i i i i inimdash imdashi i i i i i i i mdash i mdash i i M i I I
Target r = 5x10 5
r = 0 (LQG)
i i i i I I i ii i i i 1 1 1 1 i i i i i 1 1 1 1 ii i I I i i i i 11111 -3 10 -2
(a)
101 10deg Frequency (rads)
101 10
Transverse Compensator Singular Values
100
50
X3
gt 0
CO
-50
bull100 10
-imdashi i 111II i 111 ni 1mdashi I I I I I I I 1mdashi I I I N I I 1mdashi i M 111
Target r = 50 r = 0 (LQG)
m i i i i i i n i i i i I I I I I I I 1mdashi I I I I I I I 1mdashi i 111 n -3 10 -2
(b)
101 10deg Frequency (rads)
101 10
Figure 5-7 Singular values for the LQG and LQGLTR compensator compared to target return ratio (a) longitudinal design (b) transverse design
118
because the system is non-minimal ie it has transmission zeros with positive real
parts[63]
(ii) Transverse Vibra t ion Subsystem
Here Xy mdash AyXy + ByU(lv
Dv = Cvxv
with xv = 6y 8Z dV2 dZ2 Sy 6Z dV2 dZ2T] udv = dy2dZ2T and
(534)
Cy
0 0 0 0 (0) 0 0 0 0 0 0 0 0 0 0 0 0
0 0 0 0
0 0 0 0 0 0 0 0 0 0 0 0
0 0 0
h 0 0 0
0 0 0 lt ( 0 ) o 0
0 1 0 0 0 1
0 0 0
2 r
h 0 0
0 0 1 0 0 1
(535)
Av and Bv are the rows and columns of the A and B corresponding to the components
O f Xy
The state L Q R weighting matrix Qv is given as
Qv
10 0 0 0 0 0 0 0 0 10 0 0 0 0 0 0 0 0 10 4 0 0 0 0 0 0 0 0 10 4 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1
(536)
mdash
Ry =w The state noise covariance matrix is taken
1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 9 x 10 4 0
0 0 0 0 0 0 0 9 x 10 4
(537)
119
while the measurement noise covariance matrix is
1 0 0 0 0 1 0 0 0 0 5 x 10 4 0
LO 0 0 5 x 10 4
(538)
As in the development of the longitudinal controller the transverse offset conshy
troller can be designed using the same two commands (lqrc and l try) wi th the input
matrices corresponding to the transverse subsystem with r = 50 The result is the
transverse compensator system given by
(539)
where xv is the state estimates of xv The singular values of the target transfer
function is compared with those achieved by the L Q G and L Q G L T R in Figure 5-
7(b) It is apparent that the recovery is not as good as that for the longitudinal case
again due to the non-minimal system Moreover the low value of r indicates that
only a lit t le recovery in the transverse subsystem is possible This suggests that the
robustness of the L Q G design is almost the maximum that can be achieved under
these conditions
Denoting f y = f j f j r xf = x^x^T and ud = udu ^ T the longishy
tudinal compensator can be combined to give
xf =
$d =
Aku
0
0 Akv
Cku o
Cu 0
Bku
0 B Xf
Xf = Ckxf
0
kv
Hf = Akxf - Bkyf
(540)
Vu Vv 0 Cv
Xf = CfXf
Defining a permutation matrix Pf such that Xj = PfX where X is the state vector
of the original nonlinear system given in Eq(32) the compensator and full nonlinear
120
system equations can be combined as
X
(541)
Ckxf
A block diagram representation of Eq(541) is shown in Figure 5-6
524 Simulation results
The dynamical response for the stationkeeping STSS with a tether length of 20
km and ini t ia l offset position of 1 m in each direction is presented in Figure 5-8 when
both the F L T attitude controller and the L Q G L T R offset controller are activated As
expected the offset controller is successful in quickly damping the elastic vibrations
in the longitudinal and transverse direction (Figure 5-8b) However from Figure 5-
8(a) it is clear that the presence of offset control requires a larger control moment
to regulate the attitude of the platform This is due to the additional torque created
by the larger moment arm around 2 m and 125 m in the inplane and out-of-plane
directions respectively introduced by the offset control In addition the control
moment for the platform is modulated by the tethers transverse vibration through
offset coupling However this does not significantly affect the l ibrational motion of
the tether whose thruster force remains relatively unchanged from the uncontrolled
vibration case
Final ly it can be concluded that the tether elastic vibration suppression
though offset control in conjunction with thruster and momentum-wheel attitude
control presents a viable strategy for regulating the dynamics tethered satellite sysshy
tems
121
STSS configuration 3-Body a1(0)=a2(0)=at(0)=2deg P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m6y(0)=52(0)=1m dx(0)=dy(0)=dz(0)=1m Stationkeeping I = 20km LQGLTR Vibration Control
Sat 1 Pitch and Roll Angles Tether Pitch and Roll Angles mdash 1 1 1 1 1 1 I 1 1 lt 1 ~ ^ 1 ^ l _
Time (Orbits) Time (Orbits) Sat 1 Roll Control Moment Tether Roll Control Thrust
(a) Time (Orbits) Time (Orbits)
Figure 5-8 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT attitude controller and LQGLTR offset vibration controller (a) attitude and libration controller reshysponse
122
STSS configuration 3-Body a1(0)=a2(0)=a(0)=2deg p1(0)=(32(0)=Q(0)=1deg ox(0)=14mSy(0)=5z(0)=1m dx(0)=dy(0)=dz(0)=1m Stationkeeping lt = 20km LQGLTR Vibration Control
Tether X Offset Position
Time (Orbits) Tether Y Tranverse Vibration 20r
Time (Orbits) Tether Y Offset Position
t o
1 2 Time (Orbits)
Tether Z Tranverse Vibration Time (Orbits)
Tether Z Offset Position
(b)
Figure 5-8
Time (Orbits) Time (Orbits)
Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT attitude controller and LQGLTR offset vibration controller (b) vibration and offset response
123
6 C O N C L U D I N G R E M A R K S
61 Summary of Results
The thesis has developed a rather general dynamics formulation for a multi-
body tethered system undergoing three-dimensional motion The system is composed
of multiple rigid bodies connected in a chain configuration by long flexible tethers
The tethers which are free to undergo libration as well as elastic vibrations in three
dimensions are also capable of deployment retrieval and constant length stationkeepshy
ing modes of operation Two types of actuators are located on the rigid satellites
thrusters and momentum-wheels
The governing equations of motion are developed using a new Order(N) apshy
proach that factorizes the mass matrix of the system such that it can be inverted
efficiently The derivation of the differential equations is generalized to account for
an arbitrary number of rigid bodies The equations were then coded in FORTRAN
for their numerical integration with the aid of a symbolic manipulation package that
algebraically evaluated the integrals involving the modes shapes functions used to dis-
cretize the flexible tether motion The simulation program was then used to assess the
uncontrolled dynamical behaviour of the system under the influence of several system
parameters including offset at the tether attachment point stationkeeping deployshy
ment and retrieval of the tether The study covered the three-body and five-body
geometries recently flown OEDIPUS system and the proposed BICEPS configurashy
tion It represents innovation at every phase a general three dimensional formulation
for multibody tethered systems an order-N algorithm for efficient computation and
application to systems of contemporary interest
Two types of controllers one for the attitude motion and the other for the
124
flexible vibratory motion are developed using the thrusters and momentum-wheels for
the former and the variable offset position for the latter These controllers are used to
regulate the motion of the system under various disturbances The attitude controller
is developed using the nonlinear Feedback Linearization Technique ( F L T ) and is based
on the nonlinear but rigid model of the system O n the other hand the separation
of the longitudinal and transverse frequencies from those of the attitude response
allows for the development of a linear optimal offset controller using the robust Linear
Quadratic Gaussian-Loop Transfer Recovery ( L Q G L T R ) method The effectiveness
of the two controllers is assessed through their subsequent application to the original
nonlinear flexible model
More important original contributions of the thesis which have not been reshy
ported in the literature include the following
(i) the model accounts for the motion of a multibody chain-type system undershy
going librational and in the case of tethers elastic vibrational motion in all
the three directions
(ii) the dynamics formulation is based on a recursive O(N) Lagrangian algorithm
that efficiently computes the systems generalized acceleration vector
(iii) the development of an attitude controller based on the Feedback Linearizashy
tion Technique for a multibody system using thrusters and momentum-wheels
located on each rigid body
(iv) the design of a three degree of freedom offset controller to regulate elastic v i shy
brations of the tether using the Linear Quadratic Gaussian and Loop Transfer
Recovery method ( L Q G L T R )
(v) substantiation of the formulation and control strategies through the applicashy
tion to a wide variety of systems thus demonstrating its versatility
125
The emphasis throughout has been on the development of a methodology to
study a large class of tethered systems efficiently It was not intended to compile
an extensive amount of data concerning the dynamical behaviour through a planned
variation of system parameters O f course a designer can easily employ the user-
friendly program to acquire such information Rather the objective was to establish
trends based on the parameters which are likely to have more significant effect on
the system dynamics both uncontrolled as well as controlled Based on the results
obtained the following general remarks can be made
(i) The presence of the platform products of inertia modulates the attitude reshy
sponse of the platform and gives rise to non-zero equilibrium pitch and roll
angles
(ii) Offset of the tether attachment point significantly affects the equil ibrium orishy
entation of the system as well as its dynamics through coupling W i t h a relshy
atively small subsatellite the tether dynamics dominate the system response
with high frequency elastic vibrations modulating the librational motion On
the other hand the effect of the platform dynamics on the tether response is
negligible As can be expected the platform dynamics is significantly affected
by the offset along the local horizontal and the orbit normal
(iii) For a three-body system deployment of the tether can destabilize the platform
in pitch as in the case of nonzero offset However the roll motion remains
undisturbed by the Coriolis force Moreover deployment can also render the
tether slack if it proceeds beyond a critical speed
(iv) Uncontrolled retrieval of the tether is always unstable in pitch It also leads
to high frequency elastic vibrations in the tether
(v) The five-body tethered system exhibits dynamical characteristics similar to
those observed for the three-body case As can be expected now there are
additional coupling effects due to two extra bodies a rigid subsatellite and a
126
flexible tether
(vi) Cartwheeling motion of the B I C E P S configuration can also be used to adshy
vantage in the deployment of the tether However exceeding a crit ical ini t ia l
cartwheeling rate can result in a large tension in the tether which eventually
causes the satellite to bounce back rendering the tether slack
(vii) Spinning the tether about its nominal length as in the case of O E D I P U S
introduces high frequency transverse vibrations in the tether which in turn
affect the dynamics of the satellites
(viii) The F L T based controller is quite successful in regulating the attitude moshy
tion of both the tether and rigid satellites in a short period of time during
stationkeeping deployment as well as retrieval phases
(ix) The offset controller is successful in suppressing the elastic vibrations of the
tether wi thin the allowable l imits of the offset mechanism ( plusmn 2 0 m)
62 Recommendations for Future Study
A s can be expected any scientific inquiry is merely a prelude to further efforts
and insight Several possible avenues exist for further investigation in this field A
few related to the study presented here are indicated below
(i) inclusion of environmental effects particularly the free molecular reaction
forces as well as interaction with Earths magnetic field and solar radiation
(ii) addition of flexible booms attached to the rigid satellites providing a mechashy
nism for energy dissipation
(iii) validation of the new three dimensional offset control strategies using ground
based experiments
(iv) animation of the simulation results to allow visual interpretation of the system
dynamics
127
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[23] Rosenthal D E A n Order n Formulation for Robotic Systems The Journal of the Astronautical Sciences V o l 38 No 4 October-December 1990 pp 511-529
[24] Keat J E Mul t ibody System Order n Dynamics Formulation Based on Velocshyity Transform Method Journal of Guidance Control and Dynamics V o l 13 No 2 March -Apr i l 1990 pp 207-212
[25] Kurd i l a A J Menon R G and Sunkel J W Nonrecursive Order N Formushylation of Mul t ibody Dynamics Journal of Guidance Control and Dynamics V o l 16 No 5 September-October 1993 pp 838-844
[26] Ja in A Unified Formulation of Dynamics for Serial R ig id Mul t i body Systems Journal of Guidance Control and Dynamics V o l 14 No 3 May-June 1991 pp 531-542
[27] Rodriguez G and Kreutz-Delgato K Spatial Operator Factorization and Inshyversion of the Manipulator Mass Mat r ix IEEE Transaction on Robotics and Automation V o l 8 No 1 February 1992 pp 65-76
[28] Ja in A and Rodriguez G Recursive Flexible Mul t ibody Systems Dynamics Using Spatial Operators Journal of Guidance Control and Dynamics V o l 15 No 6 November-December 1992 pp 1453-1466
129
[29] Jain A and Rodriguez G Recursive Dynamics Algor i thm for Mul t ibody Sysshytems with Prescribed Mot ion Journal of Guidance Control and Dynamics V o l 16 No 5 September-October 1993 pp 830-837
[30] Pradhan S M o d i V J and Misra A K A n Order N Dynamics Formulation of Flexible Mul t ibody Systems in Tree Topology - The Lagrangian Approach AIAAAAS Astrodynamics Specialist Conference San Diego C A U S A July 1996 Paper No 96-3624 pp 480-490
[31] Banerjee A K Order-n Formulation of Extrusion of a Beam with Large Bendshying and Rotation Journal of Guidance Control and Dynamics V o l 15 No 1 January-February 1992 pp 121-127
[32] X u D M The Dynamics and Control of the Shuttle Supported Tethered Sub-satellite System P h D Thesis Department of Mechanical Engineering M c G i l l University November 1984
[33] X u D M Misra A K and M o d i V J On Vibra t ion Control of Tethered Satellite Systems Proceedings of theNASAJPL Workshop on Applications of Distributed Systems Theory to the Control of Large Space Structures Pasadena California July 1982 pp 317-327
[34] Pradhan S M o d i V J and Misra A K On the Inverse Control of the Tethered Satellite Systems The Journal of the Astronautical Sciences A A S Publications Office San Diego California U S A V o l 43 No 2 Apr i l - June 1995 pp 179-193
[35] K i m E and Vadal i S R Modeling Issues Related to Retrieval of Flexible Tethered Satellite Systems Journal of Guidance Control and Dynamics V o l 18 No 5 September-October 1995 pp 1169-1176
[36] W u S Chang C and Housner J M Finite Element Approach for Transient Analysis of Mul t ibody Systems Journal of Guidance Control and Dynamics V o l 15 No 4 July-August 1992 pp 847-854
[37] Pradhan S M o d i V J and Misra A K Simultaneous Control of Platform Att i tude and Tether Vibra t ion using Offset Strategy AIAAA AS Astrodynamics Specialist Conference San Diego C A U S A July 1996 Paper No 96-3570 pp 1-11
[38] Ruying F and Bainum P M Dynamics and Control of a Space Platform with a Tethered Subsatellite Journal of Guidance Control and Dynamics V o l 11 No 4 July-August 1988 pp 377-381
[39] M o d i V J Lakshmanan P K and Misra A K Offset Control of Tethered Satellite Systems Analysis and Experimental Verification Acta Astronautica V o l 21 No 5 1990 pp 283-294
[40] X u D M Misra A K and M o d i V J Thruster Augmented Act ive Cin t ro l of a Tethered Satellite System During its Retrieval Journal of Guidance Control and Dynamics V o l 9 No 6 November-December 1986 pp 663-672
[41] M o d i V J Pradhan S and Misra A K Off-Set Control of the Tethered Systems Using a Graph Theoretic Approach Acta Astronautica V o l 35 No
130
6 1995 pp 373-384
[42] Pradhan S M o d i V J Bhat M S and Misra A K Mat r ix Method for Eigenstructure Assignment The Mult i - Input Case with Appl ica t ion Journal of Guidance Control and Dynamics V o l 17 No 5 September-October 1994 pp 983-989
[43] Pradhan S Planar Dynamics and Control of Tethered Satellite Systems P h D Thesis Department of Mechanical Engineering The University of Br i t i sh Columbia December 1994
[44] Bainum P M and Kumar V K Optimal-Control of the Shuttle-Tethered System Acta Astronautica V o l 7 No 12 1980 pp 1333-1348
[46] Hughes P C Spacecraft Att i tude Dynamics John Wi ley amp Sons New York U S A 1992 Chapter 2 pp 7-38
[47] Stengel R F Optimal Control and Estimation Dover Publications Inc New York U S A 1994 p 26
[48] Nayfeh A H and Mook D T Nonlinear Oscillation John Wi ley and Sons New York U S A 1979 pp 485-500
[49] He X and Powell J D Tether Damping in Space Journal of Guidance
Control and Dynamics V o l 13 No 1 1990 pp 104-112
[50] Meirovitch L Elements of Vibration Analysis M c G r a w - H i l l Inc New York U S A 1986 pp 255-256
[51] Greenwood D T Principles of Dynamics Prentice Ha l l Inc Englewood Cliffs New Jersey U S A 1965 pp 252-275
[52] Boyce W E and D i P r i m a R C Elemental Differential Equations and Boundshy
ary Value Problems John Wi ley amp Sons Inc New York U S A 1986 pp 424-431
[53] IMSL Library Reference Manual Vol 1 I M S L Inc Houston Texas U S A 1980 pp D G E A R 1 - D G E A R 9
[54] Gear C W Numerical Initial Value Problems in Ordinary Differential Equashy
tions Prentice-Hall Englewood Cliffs New Jersey U S A 1971 pp 158-166
[55] Char B W Geddes K O Gonnet G H Leong B L Monagan M B and Wat t S M First Leaves A Tutorial Introduction to Maple V Springer-Verlag New York U S A 1992
[56] Tsakalis K S and Ioannou P A Linear Time-Varying Systems Control and
Adaptation Prentice-Hall Inc Englewood Cliffs New Jersey U S A 1993 pp 148-180
[57] Su R O n the Linear Equivalents of Nonlinear Systems Systems and Control
Letter V o l 2 No 1 1982 pp 48-52
[58] M o d i V J Karray F and Chan J K On the Control of a Class of Flexible Manipulators Using Feedback Linearization Approach 42nd Congress of the International Astronautical Federation October 1991 Montreal Canada Paper No IAF-91-324
131
[59] K u o B C Automatic Control Systems Prentice-Hall Inc Englewood Cliffs New Jersey U S A 1987 pp 314-326
[60] Athans M The Role and Use of the Stochastic Linear-Quadrat ic-Guassian Problem in Control System Design IEEE Transactions on Automatic Control V o l A C - 1 6 No 6 December 1971 pp 529-552
[61] Maciejowski J M Multivariable Feedback Design Addison-Wesley Publ ishing Company Wokingham England 1989 Chapters 1-5
[62] Chiang R Y and Safonov M G Robust Control Toolbox Users Guide The M a t h Works Inc Natick Mass U S A 1992 pp 272-274
[63] Stein G and Athans M The L Q G L T R Procedure for Mult ivariable Feedback Control Design IEEE Transactions on Automatic Control V o l A C - 3 2 No 2 February 1987 pp 105-114
[64] Borisenko A I and Tarapov I E Vector and Tensor Analysis with Applicashytions Dover Publications Inc New York U S A 1979
[65] Lass H Vector and Tensor Analysis M c G r a w - H i l l Book Company Inc New York U S A 1950
132
A P P E N D I X I TENSOR R E P R E S E N T A T I O N OF T H E
EQUATIONS OF M O T I O N
11 Prel iminary Remarks
The derivation of the equations of motion for the multibody tethered system
involves the product of several matrices as well as the derivative of matrices and
vectors with respect to other vectors In order to concisely code these relationships
in FORTRAN the matrix equations of motion are expressed in tensor notation
Tensor mathematics in general is an extremely powerful tool which can be
used to solve many complex problems in physics and engineering[6465] However
only a few preliminary results of Cartesian tensor analysis are required here They
are summarized below
12 Mathematical Background
The representation of matrices and vectors in tensor notation simply involves
the use of indices referring to the specific elements of the matrix entity For example
v = vk (k = lN) (Il)
A = Aij (i = lNj = lM)
where vk is the kth element of vector v and Aj is the element on the t h row and
3th column of matrix A It is clear from Eq(Il) that exactly one index is required to
completely define an entire vector whereas two independent indices are required for
matrices For this reason vectors and matrices are known as first-order and second-
order tensors respectively Scaler variables are known as zeroth-order tensors since
in this case an index is not required
133
Matrix operations are expressed in tensor notation similar to the way they are
programmed using a computer language such as FORTRAN or C They are expressed
in a summation notation known as Einstein notation For example the product of a
matrix with a vector is given as
w = Av (12)
or in tensor notation W i = ^2^2Aij(vkSjk)
4-1 ltL3gt j
where 8jk is the Kronecker delta Dropping the summation symbol the matrix-vector
product can be expressed compactly as
Wi = AijVj = Aimvm (14)
Here j is a dummy index since it appears twice in a single term and hence can
be replaced by any other index Note that since the resulting product is a vector
only one index is required namely i that appears exactly once on both sides of the
equation Similarly
E = aAB + bCD (15)
or Ej mdash aAimBmj + bCimDmj
~ aBmjAim I bDmjCim (16)
where A B C D and E are second-order tensors (matrices) and a and b are zeroth-
order tensors (scalers) In addition once an expression is in tensor form it can be
treated similar to a scaler and the terms can be rearranged
The transpose of a matrix is also easily described using tensors One simply
switches the position of the indices For example let w = ATv then in tensor-
notation
Wi = AjiVj (17)
134
The real power of tensor notation is in its unambiguous expression of terms
containing partial derivatives of scalers vectors or matrices with respect to other
vectors For example the Jacobian matrix of y = f(x) can be easily expressed in
tensor notation as
di-dx-j-1 ( L 8 )
mdash which is a second-order tensor as expected If f(x) = Ax where A is constant then
f - a )
according to the current definition If C = A(x)B(x) then the time derivative of C
is given by
C = Ax)Bx) + A(x)Bx) (110)
or in tensor notation
Cij = AimBmj + AimBmj (111)
= XhiAimfcBmj + AimBmj^k)
which is more clear than its equivalent vector form Note that A^^ and Bmjk are
third-order tensors that can be readily handled in F O R T R A N
13 Forcing Function
The forcing function of the equations of motion as given by Eq(260) is
- u 1-SdM- dqtdPe 9 f i m F(q qt) = Mq - -q mdashq + mdashmdash - Qd (260)
This expression can be converted into tensor form to facilitate its implementation
into F O R T R A N source code
Consider the first term Mq From Eq(289) the mass matrix M is given by
T M mdash Rv MtRv which in tensor form is
Mtj = RZiM^R^j (112)
135
Taking the time derivative of M gives
Mij = qsRv
ni^MtnrnRv
mj + Rv
niMtnTn^Rv
mj + Rv
niMtnmRv
mjs) (113)
The second term on the right hand side in Eq(260) is given by
-JPdMu 1 bdquo T x
2Q ~cWq = 2Qs s r k Q r ^ ^
Expanding Msrk leads to
M src = RnskMtnmRmr + RnsMtnmkRmr + RnsMtnmRmrk- (L15)
Finally the potential energy term in Eq(260) is also expanded into tensor
form to give dPe = dqt dPe
dq dq dqt = dq^QP^
and since ^ = Rp(q)q then
Now inserting Eqs(I13-I17) into Eq(260) and rearranging the tensor form of the
forcing term can how be stated as
Fk(q^q^)=(RP
sk + R P n J q n ) ^ - - Q d k
+ QsQr | (^Rlks ~ 2 M t n m R m r
+ (^RnkMtnms ~ 2~RnsMtnmgtk^ R m r
(^RnkRmrs ~ ~2RnsRmrk^j
where Fk(q q t) represents the kth component of F
(118)
136
A P P E N D I X II R E D U C E D EQUATIONS OF M O T I O N
II 1 Prel iminary Remarks
The reduced model used for the design of the vibration controller is represented
by two rigid end-bodies capable of three-dimensional attitude motion interconnected
with a fixed length tether The flexible tether is discretized using the assumed-
mode method with only the fundamental mode of vibration considered to represent
longitudinal as well as in-plane and out-of-plane transverse deformations In this
model tether structural damping is neglected Finally the system is restricted to a
nominal circular orbit
II 2 Derivation of the Lagrangian Equations of Mot ion
Derivation of the equations of motion for the reduced system follows a similar
procedure to that for the full case presented in Chapter 2 However in this model the
straightforward derivation of the energy expressions for the entire system is undershy
taken ignoring the Order(N) approach discussed previously Furthermore in order
to make the final governing equations stationary the attitude motion is now referred
to the local LVLH frame This is accomplished by simply defining the following new
attitude angles
ci = on + 6
Pi = Pi (HI)
7t = 0
where and are the pitch and roll angles respectively Note that spin is not
considered in the reduced model hence 7J = 0 Substituting Eq(IIl) into Eq(214)
137
gives the new expression for the rotation matrix as
T- CpiSa^ Cai+61 S 0 S a + e i
-S 0 c Pi
(II2)
W i t h the new rotation matrix defined the inertial velocity of the elemental
mass drrii o n the ^ link can now be expressed using Eq(215) as
(215)
Since the tether length is assumed to be fixed T j f j and Tj$jlt^ of Eq(215) are
identically zero Also defining
(II3) Vi = I Pi i
Rdm- f deg r the reduced system is given by
where
and
Rdm =Di + Pi(gi)m + T^iSi
Pi(9i)m= [Tai9i T0gi T^] fj
(II4)
(II5)
(II6)
D1 = D D l D S v
52 = 31 + P(d2)m +T[d2
D3 = D2 + l2 + dX2Pii)m + T2i6x
Note that $ j for i = 2 is defined in Section 212 whereas for i = 1 and 3 it is
the null matrix Since only one mode is used for each flexible degree of freedom
5i = [6Xi6yi6Zi]T for i = 2
The kinetic energy of the system can now be written as
i=l Z i = l Jmi RdmRdmdmi (UJ)
138
Expanding Kampi using Eq(II4)
KH = miDD + 2DiP
i j gidmi)fji + 2D T ltMtrade^
+tf J P^(9i)Pi9i)d^l + 2mT j PFmTi^dm^ (II8)
where the integrals are evaluated using the procedure similar to that described in
Chapter 2 The gravitational and strain energy expressions are given by Eq(247)
and Eq(251) respectively using the newly defined rotation matrix T in place of
TV
Substituting the kinetic and potential energy expressions in
d ( 8Ke d(Ke - Pe) i bull i Q ^ 0 (II9)
d t dqred d(ired
with
Qred = [ algt l gt a 2 2 A 2 A 2 lt ^ 2 Q 3 3 ] (5-14)
and d2 regarded as a time-varying parameter the nonlinear non-autonomous equashy
tions of motion for the reduced system can finally be expressed as
MredQred + fred = reg- (5-15)
139
number of flexible modes in the longitudinal inplane and out-
of-plane transverse directions respectively for the i ^ link
number of attitude control actuators
- $ T
set of generalized coordinates for the itfl l ink which accounts for
interactions with adjacent links
qv---QtNT
set of coordinates for the independent t h link (not connected to
adjacent links)
rigid position of dm in the frame Fj
position of centre of mass of the itfl l ink relative to the frame Fj
i + $DJi
desired settling time of the j 1 attitude actuator
actuator force vector for entire system
flexible deformation of the link along the Xj yi and Zj direcshy
tions respectively
control input for flexible subsystem
Cartesian components of fj
actual and estimated state of flexible subsystem respectively
output vector of flexible subsystem
LIST OF FIGURES
1-1 A schematic diagram of the space platform based N-body tethered
satellite system 2
1-2 Associated forces for the dumbbell satellite configuration 3
1- 3 Some applications of the tethered satellite systems (a) multiple
communication satellites at a fixed geostationary location
(b) retrieval maneuver (c) proposed B I C E P S configuration 6
2- 1 Vector components of the i 1 and (i- l ) 1 chain links 15
2-2 Vector components in cylindrical orbital coordinates 20
2-3 Inertial position of subsatellite thruster forces 38
2- 4 Coupled force representation of a momentum-wheel on a rigid body 40
3- 1 Flowchart showing the computer program structure 45
3-2 Kinet ic and potentialenergy transfer for the three-body platform based
tethered satellite system (a) variation of kinetic and potential energy
(b) percent change in total energy of system 48
3-3 Kinet ic and potential energy transfer for the five-body platform based
tethered satellite system (a) variation of kinetic and potential energy
(b) percent change in total energy of system 49
3-4 Simulation results for the platform based three-body tethered
system originally presented in Ref[43] 51
3- 5 Simulation results for the platform based three-body tethered
system obtained using the present computer program 52
4- 1 Schematic diagram showing the generalized coordinates used to
describe the system dynamics 55
4-2 Stationkeeping dynamics of the three-body STSS configuration
without offset (a) attitude response (b) vibration response 57
4-3 Stationkeeping dynamics of the three-body STSS configuration
xi
with offset along the local vertical
(a) attitude response (b) vibration response 60
4-4 Stationkeeping dynamics of the three-body STSS configuration
with offset along the local horizontal
(a) attitude response (b) vibration response 62
4-5 Stationkeeping dynamics of the three-body STSS configuration
with offset along the orbit normal
(a) attitude response (b) vibration response 64
4-6 Stationkeeping dynamics of the three-body STSS configuration
wi th offset along the local horizontal local vertical and orbit normal
(a) attitude response (b) vibration response 67
4-7 Deployment dynamics of the three-body STSS configuration
without offset
(a) attitude response (b) vibration response 69
4-8 Deployment dynamics of the three-body STSS configuration
with offset along the local vertical
(a) attitude response (b) vibration response 72
4-9 Retrieval dynamics of the three-body STSS configuration
with offset along the local vertical and orbit normal
(a) attitude response (b) vibration response 74
4-10 Schematic diagram of the five-body system used in the numerical example 77
4-11 Stationkeeping dynamics of the five-body STSS configuration
without offset
(a) attitude response (b) vibration response 78
4-12 Deployment dynamics of the five-body STSS configuration
wi th offset along the local vertical
(a) attitude response (b) vibration response 81
4-13 Stationkeeping dynamics of the three-body B I C E P S configuration
x i i
with offset along the local vertical
(a) attitude response (b) vibration response 84
4-14 Cartwheeling dynamics with deployment for the three-body B I C E P S
with offset along the local vertical
(a) attitude response (b) vibration response 86
4-15 Spin dynamics (7 = l deg s ) of the three-body O E D I P U S configuration
with offset along the local vertical
(a) attitude response (b) vibration response 89
4- 16 Spin dynamics (7 = 10deg s ) of the three-body O E D I P U S configuration
wi th offset along the local vertical
(a) attitude response (b) vibration response 91
5- 1 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear rigid F L T controller with offset along the
local horizontal local vertical and orbit normal
(a) attitude and vibration response (b) control actuator time histories 98
5-2 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear flexible F L T controller with offset along the
local horizontal local vertical and orbit normal
(a) attitude and vibration response (b) control actuator time histories 101
5-3 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear rigid F L T controller with offset along
the local vertical
(a) attitude and vibration response (b) control actuator time histories 103
5-4 Deployment dynamics of the three-body STSS using the nonlinear
rigid F L T controller with offset along the local vertical
(a) attitude and vibration response (b) control actuator time histories 105
5-5 Retrieval dynamics of the three-body STSS using the non-linear
rigid F L T controller with offset along the local vertical and orbit normal
(a) attitude and vibration response (b) control actuator time histories 108
x i i i
5-6 Block diagram for the L Q G L T R estimator based compensator 115
5-7 Singular values for the L Q G and L Q G L T R compensator compared to target
return ratio (a) longitudinal design (b) transverse design 118
5-8 Controlled dynamics of the three-body STSS during stationkeeping
using the nonlinear rigid F L T attitude controller and L Q G L T R
offset vibration controller (a) attitude and libration controller response
(b) vibration and offset response 122
xiv
A C K N O W L E D G E M E N T
Firs t and foremost I would like to express the most genuine gratitude to
my supervisor Prof V i n o d J M o d i whose unending guidance and encouragement
proved most invaluable during the course of my studies
I am also deeply indebted to Dr Satyabrata Pradhan for his limitless patience
and technical advice from which this work would not have been possible I would also
like to extend a word of appreciation towards Prof A r u n K Mis ra ( M c G i l l University)
for ini t iat ing my interest in space dynamics and control
Also I would like to thank my collegues Dr Sandeep Munshi M r Mathieu
Caron and M r Yuan Chen as well as Dr Mae Seto M r Gary L i m and M r Mark
Chu for their words of advice and helpful suggestions that heavily contributed to all
aspects of my studies
Final ly I would like to thank all my family and friends here in Vancouver and
back home in Montreal whose unending support made the past two years the most
memorable of my life
The research project was supported by the Natural Sciences and Engineering
Research Council s P G S - A Scholarship held by the author as well as by N S E R C grant
A-2181 held by Prof V J M o d i
xv
1 INTRODUCTION
11 Prel iminary Remarks
W i t h the ever changing demands of the worlds population one often wonders
about the commitment to the space program On the other hand the importance
of worldwide communications global environmental monitoring as well as long-term
habitation in space have demanded more ambitious and versatile satellites systems
It is doubtful that the Russian scientist Tsiolkovsky[l] when he first proposed the
uti l izat ion of the Earths gravity-gradient environment in the last century would have
envisioned the current and proposed applications of tethered satellite systems
A tethered satellite system consists of two or more subsatellites connected
to each other by long thin cables or tethers which are in tension (Figure 1-1) The
subsatellite which can also include the shuttle and space station are in orbit together
The first proposed use of tethers in space was associated with the rescue of stranded
astronauts by throwing a buoy on a tether and reeling it to the rescue vehicle
Prel iminary studies of such systems led to the discovery of the inherent instability
during the tether retrieval[2] Nevertheless the bir th of tethered systems occured
during the Gemini X I and X I I missions[3] in 1966 when a short tether (10-30m)
was used to generate the first artificial gravity environment in space (000015g) by
cartwheeling (spinning) the system about its center of mass
The mission also demonstrated an important use of tethers for gravity gradient
stabilization The force analysis of a simplified model illustrates this point Consider
the dumbbell satellite configuration orbiting about the centre of Ear th as shown in
Figure 1-2 Satellite 1 and satellite 2 are connected to each other by a tether with
1
Space Station (Satellite 1 )
2
Satellite N
Figure 1-1 A schematic diagram of the space platform based N-body tethered satellite system
2
Figure 1-2 Associated forces for the dumbbell satellite configuration
3
the whole system free to librate about the systems centre of mass C M The two
major forces acting on each body are due to the centrifugal and gravitational effects
Since body 1 is further away from the earth its gravitational force Fg^y is smaller
than that of body 2 ie Fgi lt Fg^ However its centrifugal force FCl is greater
then the centrifugal force experienced by the second satellite (FCl gt FC2) Moreover
for body 1 Fg^ lt Fc^ in contrast to Fg2 gt FC2 for body 2 Thus resolving the force
components along the tether there is an evident resultant tension force Ff present
in the tether Similarly adding the normal components of the two forces vectorially
results in a force FR which restores the system to its stable equil ibrium along the
local vertical These tension and gravity-gradient restoring forces are the two most
important features of tethered systems
Several milestone missions have flown in the recent past The U S A J a p a n
project T P E (Tethered Payload Experiment) [4] was one of the first to be launched
using a sounding rocket to conduct environmental studies The results of the T P E
provided support for the N A S A I t a l y shuttle based tethered satellite system referred
to as the TSS-1 mission[5] This experiment aimed to study the electrodynamic effects
of a shuttle borne satellite system with a conductive tether connection where an
electric current was induced in the tether as it swept through the Earth s magnetic
field Unfortunetely the two attempts in August 1992 and February 1996 resulted
in only partial success The former suffered from a spool failure resulting in a tether
deployment of only 256m of a planned 20km During the latter attempt a break
in the tethers insulation resulted in arcing leading to tether rupture Nevertheless
the information gathered by these attempts st i l l provided the engineers and scientists
invaluable information on the dynamic behaviour of this complex system
Canada also paved the way with novel designs of tethered system aimed
primari ly at the study of the ionosphere and Northern Lights (Aurora Borealis)
4
The O E D I P U S A and C (Observation of Electrified Distributions in the Ionospheric
P lasma - a Unique Strategy) missions[6] launched from a sounding rocket in January
1989 and November 1995 respectively provided insight into the complex dynamical
behaviour of two comparable mass satellites connected by a 1 km long spinning
tether
Final ly two experiments were conducted by N A S A called S E D S I and II
(Small Expendable Deployable System) [7] which hold the current record of 20 km
long tether In each of these missions a small probe (26 kg) was deployed from the
second stage of a D E L T A II rocket thus succesfully demonstrating the feasibility of
long tether deployment Retrieval of the tether which is significantly more difficult
has not yet been achieved
Several proposed applications are also currently under study These include the
study of earths upper atmosphere using probes lowered from the shuttle in a low earth
orbit ( L E O ) to the deployment and retrieval of satellites for servicing Even more
promising application concerns the generation of power for the proposed International
Space Station using conductive tethers The Canadian Space Agency (CSA) has also
proposed a dual-mass satellite system B I C E P S (Bl-static Canadian Experiment on
Plasmas in Space) to make simultaneous measurements at different locations in the
environment for correlation studies[8] A unique feature of the B I C E P S mission is
the deployment of the tether aided by the angular momentum of the system which
is ini t ia l ly cartwheeling about its orbit normal (Figure 1-3)
In addition to the three-body configuration multi-tethered systems have been
proposed for monitoring Earths environment in the global project monitored by
N A S A entitled Mission to Planet Ear th ( M T P E ) The proposed mission aims to
study pollution control through the understanding of the dynamic interactions be-
5
Retrieval of Satellite for Servicing
lt mdash
lt Satellite
Figure 1-3 Some applications of the tethered satellite systems (a) multiple comshymunication satellites at a fixed geostationary location (b) retrieval maneuver (c) proposed BICEPS configuration
6
tween the Earths atmosphere biosphere hydrosphere and cryosphere This wi l l be
accomplished with multiple tethered instrumentation payloads simultaneously soundshy
ing at different altitudes for spatial correlations Such mult ibody systems are considshy
ered to be the next stage in the evolution of tethered satellites Micro-gravity payload
modules suspended from the proposed space station for long-term experiments as well
as communications satellites with increased line-of-sight capability represent just two
of the numerous possible applications under consideration
12 Brief Review of the Relevant Literature
The design of multi-payload systems wi l l require extensive dynamic analysis
and parametric study before the final mission configuration can be chosen This
wi l l include a review of the fundamental dynamics of the simple two-body tethered
systems and how its dynamics can be extended to those of multi-body tethered system
in a chain or more generally a tree topology
121 Multibody 0(N) formulation
Many studies of two-body systems have been reported One of the earliest
contribution is by Rupp[9] who was the first to study planar dynamics of the simshy
plified Shuttle based Tethered Satellite System (STSS) His pioneering work led to
the discovery of pitch oscillation growth during the retrieval phase Later Baker[10]
advanced the investigation to the third dimension in addition to adding atmospheric
effects to the system A more complete dynamical model was later developed and
analyzed by M o d i and Misra[ l l 12] which included deployment and tether flexibility
A comprehensive survey of important developments and milestones in the area have
been compiled and reviewed by Mis ra and Modi[13] The major conclusions based on
the literature may be summarized as follows
bull the stationkeeping phase is stable
7
bull deployment can be unstable if the rate exceeds a crit ical speed
bull retrieval of the tether is inherently unstable
bull transverse vibrations can grow due to the Coriolis force induced during deshy
ployment and retrieval
Misra Amier and Modi[14] were one of the first to extend these results to the
three-satellite double pendulum case This simple model which includes a variable
length tether was sufficient to uncover the multiple equilibrium configurations of
the system Later the work was extended to include out-of-plane motion and a
preliminary reel-rate control law to regulate the librational motion[15] Kesmir i and
Misra[16] developed a general formulation for N-body tethered systems based on the
Lagrangian principle where three-dimensional motion and flexibility are accounted for
However from their work it is clear that as the number of payload or bodies increases
the computational cost to solve the forward dynamics also increases dramatically
Tradit ional methods of inverting the mass matrix M in the equation
ML+F = Q
to solve for the acceleration vector proved to be computationally costly being of the
order for practical simulation algorithms O(N^) refers to a mult ibody formulashy
tion algorithm whose computational cost is proportional to the cube of the number
of bodies N used It is therefore clear that more efficient solution strategies have to
be considered
Many improved algorithms have been proposed to reduce the number of comshy
putational steps required for solving for the acceleration vector Rosenthal[17] proshy
posed a recursive formulation based on the triangularization of the equations of moshy
tion to obtain an 0(N2) formulation He also demonstrates the use of a symbolic
manipulation program to reduce the final number of computations A recursive L a -
8
grangian formulation proposed by Book[18] also points out the relative inefficiency of
conventional formulations and proceeds to derive an algorithm which is also 0(N2)
A recursive formulation is one where the equations of motion are calculated in order
from one end of the system to the other It usually involves one or more passes along
the links to calculate the acceleration of each link (forward pass) and the constraint
forces (backward or inverse pass) Non-recursive algorithms express the equations of
motion for each link independent of the other Although the coupling terms st i l l need
to be included the algorithm is in general more amenable to parallel computing
A st i l l more efficient dynamics formulation is an O(N) algorithm Here the
computational effort is directly proportional to the number of bodies or links used
Several types of such algorithms have been developed over the years The one based
on the Newton-Euler approach has recursive equations and is used extensively in the
field of multi-joint robotics[1920] It should be emphasized that efficient computation
of the forward and inverse dynamics is imperative if any on-line (real-time) control
of the joint is to be achieved Hollerbach[21] proposed a recursive O(N) formulation
based on Lagranges equations of motion However his derivation was primari ly foshy
cused on the inverse dynamics and did not improve the computational efficiency of
the forward dynamics (acceleration vector) Other recursive algorithms include methshy
ods based on the principle of vir tual work[22] and Kanes equations of motion[23]
Keat[24] proposed a method based on a velocity transformation that eliminated the
appearance of constraint forces and was recursive in nature O n the other hand
K u r d i l a Menon and Sunkel[25] proposed a non-recursive algorithm based on the
Range-Space method which employs element-by-element approach used in modern
finite-element procedures Authors also demonstrated the potential for parallel comshy
putation of their non-recursive formulation The introduction of a Spatial Operator
Factorization[262728] which utilizes an analogy between mult ibody robot dynamics
and linear filtering and smoothing theory to efficiently invert the systems mass ma-
9
t r ix is another approach to a recursive algorithm Their results were further extended
to include the case where joints follow user-specified profiles[29] More recently Prad-
han et al [30] proposed a non-recursive Lagrangian algorithm for factorizing the mass
matrix leading to an O(N) formulation of the forward dynamics of the system A n
efficient O(N) algorithm where the number of bodies N in the system varies on-line
has been developed by Banerjee[31]
122 Issues of tether modelling
The importance of including tether flexibility has been demonstrated by several
researchers in the past However there are several methods available for modelling
tether flexibility Each of these methods has relative advantages and limitations
wi th regards to the computational efficiency A continuum model where flexible
motion is discretized using the assumed mode method has been proposed[3233] and
succesfully implemented In the majority of cases even one flexible mode is sufficient
to capture the dynamical behaviour of the tether[34] K i m and Vadali[35] proposed a
so-called bead model or lumped-mass approach that discretizes the tether using point
masses along its length However in order to accurately portray the motion of the
tether a significantly high number of beads are needed thus increasing the number of
computation steps Consequently this has led to the development of a hybrid between
the last two approaches[16] ie interpolation between the beads using a continuum
approach thus requiring less number of beads Finally the tether flexibility can also
be modelled using a finite-element approach[36] which is more suitable for transient
response analysis In the present study the continuum approach is adopted due to
its simplicity and proven reliability in accurately conveying the vibratory response
In the past the modal integrals arising from the discretization process have
been evaluted numerically In general this leads to unnecessary effort by the comshy
puter Today these integrals can be evaluated analytically using a symbolic in-
10
tegration package eg Maple V Subsequently they can be coded in F O R T R A N
directly by the package which could also result in a significant reduction in debugging
time More importantly there is considerable improvement in the computational effishy
ciency [16] especially during deployment and retrieval where the modal integral must
be evaluated at each time-step
123 Attitude and vibration control
In view of the conclusions arrived at by some of the researchers mentioned
above it is clear that an appropriate control strategy is needed to regulate the dyshy
namics of the system ie attitude and tether vibration First the attitude motion of
the entire system must be controlled This of course would be essential in the case of
a satellite system intended for scientific experiments such as the micro-gravity facilishy
ties aboard the proposed Internation Space Station Vibra t ion of the flexible members
wi l l have to be checked if they affect the integrity of the on-board instrumentation
Thruster as well as momentum-wheel approach[34] have been considered to
regulate the rigid-body motion of the end-satellite as well as the swinging motion
of the tether The procedure is particularly attractive due to its effectiveness over a
wide range of tether lengths as well as ease of implementation Other methods inshy
clude tension and length rate control which regulate the tethers tension and nominal
unstretched length respectively[915] It is usually implemented at the deployment
spool of the tether More recently an offset strategy involving time dependent moshy
tion of the tether attachment point to the platform has been proposed[3738] It
overcomes the problem of plume impingement created by the thruster control and
the ineffectiveness of tension control at shorter lengths However the effectiveness of
such a controller can become limited with an exceedingly long tether due to a pracshy
tical l imi t on the permissible offset motion In response to these two control issues a
hybrid thrusteroffset scheme has been proposed to combine the best features of the
11
two methods[39]
In addition to attitude control these schemes can be used to attenuate the
flexible response of the tether Tension and length rate control[12] as well as thruster
based algorithms[3340] have been proposed to this end M o d i et al [41] have sucshy
cessfully demonstrated the effectiveness of an offset strategy to damp undesired v i shy
brations Passive energy dissipative devices eg viscous dampers are also another
viable solution to the problem
The development of various control laws to implement the above mentioned
strategies has also recieved much attention A n eigenstructure assignment in conducshy
tion wi th an offset controller for vibration attenuation and momemtum wheels for
platform libration control has been developed[42] in addition to the controller design
from a graph theoretic approach[41] Also non-linear feedback methods such as the
Feedback Linearization Technique ( F L T ) that are more suitable for controlling highly
nonlinear non-autonomous coupled systems have also been considered[43]
It is important to point out that several linear controllers including the classhy
sic state feedback Linear Quadratic Regulator ( L Q R ) have received considerable
attention[3944] Moreover robust methods such as the Linear Quadratic Guas-
s ian Loop Transfer Recovery ( L Q G L T R ) method have also been developed and imshy
plemented on tethered systems[43] A more complete review of these algorithms as
well as others applied to tethered system has been presented by Pradhan[43]
13 Scope of the Investigation
The objective of the thesis is to develop and implement a versatile as well as
computationally efficient formulation algorithm applicable to a large class of tethered
satellite systems The distinctive features of the model can be summarized as follows
12
bull TV-body 0(N) tethered satellite system in a chain-type topology
bull the system is free to negotiate a general Keplerian orbit and permitted to
undergo three dimensional inplane and out-of-plane l ibrtational motion
bull r igid bodies constituting the system are free to execute general rotational
motion
bull three dimensional flexibility present in the tether which is discretized using
the continuum assumed-mode method with an arbitrary number of flexible
modes
bull energy dissipation through structural damping is included
bull capability to model various mission configurations and maneuvers including
the tether spin about an arbitrary axis
bull user-defined time dependent deployment and retrieval profiles for the tether
as well as the tether attachment point (offset)
bull attitude control algorithm for the tether and rigid bodies using thrusters and
momentum-wheels based on the Feedback Linearization Technique
bull the algorithm for suppression of tether vibration using an active offset (tether
attachment point) control strategy based on the optimal linear control law
( L Q G L T - R )
To begin with in Chapter 2 kinematics and kinetics of the system are derived
using the O(N) formulation methodology developed by Pradhan et al [30] Chapshy
ter 3 discusses issues related to the development of the simulation program and its
validation This is followed by a detailed parametric study of several mission proshy
files simulated as particular cases of the versatile formulation in Chapter 4 Next
the attitude and vibration controllers are designed and their effectiveness assessed in
Chapter 5 The thesis ends with concluding remarks and suggestions for future work
13
2 F O R M U L A T I O N OF T H E P R O B L E M
21 Kinematics
211 Preliminary definitions and the i ^ position vector
The derivation of the equations of motion begins with the definition of the
inertial position of an arbitrary link of the mult ibody system Let the i ^ link of the
system be free to translate and rotate in 3-D space From Figure 2-1 the position
vector Rdm^ to the mass element drrn on the i ^ link wi th reference to the inertial
frame FQ can be written as
Here D represents the inertial position of the t h body fixed frame Fj relative to FQ
ri = [xiyiZi]T is the rigid position vector of drrii wi th respect to F J and f(fi) is
the flexible deformation at fj also with respect to the frame Fj Bo th these vectors
are relative to Fj Note in the present model tethers (i even) are considered flexible
and X corresponds to the nominal position along the unstretched tether while a and
ZJ are by definition equal to zero On the other hand the rigid bodies referred to as
satellites (i odd) including the platform are taken to be rigid ie f(fj) = 0 T j is
the rotation matrix used to express body-fixed vectors with reference to the inertial
frame
212 Tether flexibility discretization
The tether flexibility is discretized with an arbitrary but finite number of
Di + Ti(i + f(i)) (21)
14
15
modes in each direction using the assumed-mode method as
Ui nvi
wi I I i = 1
nzi bull
(22)
where nXi nyi and n 2 ^ are the number of modes used in the X m and Z directions
respectively For the ] t h longitudinal mode the admissible mode shape function is
taken as 2 j - l
Hi(xiii)=[j) (23)
where l is the i ^ tether length[32] In the case of inplane and out-of-plane transverse
deflections the admissible functions are
ampyi(xili) = ampZixuk) = v ^ s i n JKXj
k (24)
where y2 is added as a normalizing factor In this case both the longitudinal and
transverse mode shapes satisfy the geometric boundary conditions for a simple string
in axial and transverse vibration
Eq(22) is recast into a compact matrix form as
ff(i) = $i(xili)5i(t) (25)
where $i(x l) is the matrix of the tether mode shapes defined as
0 0
^i(xiJi) =
^ X bull bull bull ^XA
0
0 0
0
G 5R 3 x n f
(26)
and nfj = nx- + nyi + nZi Note $J(XJJ) is a function of the spatial coordinate x
and not of yi or ZJ However it is also a function of the tether length parameter li
16
which is time-varying during deployment and retrieval For this reason care must be
exercised when differentiating $J(XJJ) with respect to time such that
d d d k) = Jtregi(Xi li) = ^ 7 $ t ( ^ i k)Xi + gf$i(xigt li)k- (2-7)
Since x represents the nominal rate of change of the position of the elemental mass
drrii it is equal to j (deploymentretrieval rate) Let t ing
( 3 d
dx~ + dT)^Xili^ ( 2 8 )
one arrives at
$i(xili) = $Di(xili)ii (29)
The time varying generalized coordinate vector 5 in Eq(25) is composed of
the longitudinal and transverse components ie
ltM) G s f t n f i x l (210)
where 5X^ 5y- and Sz^ are nx- x 1 ny x 1 and nz- x 1 size vectors and correspond to
$ X and $ Z respectively
213 Rotation angles and transformations
The matrix T j in Eq(21) represents a rotation transformation from the body
fixed frame Fj to the inertial frame FQ ie FQ = T j F j It is defined by the 3-2-1
ordered sequence of elementary rotations[46]
1 P i tch F[ = Cf ( a j )F 0
2 R o l l F = C f (3j)F = C7f (3j)Cf ( a j )F 0
3 Yaw F = C7( 7 i )F = C7(7l)cf^)Cf(ai)FQ = Q F 0
17
where
c f (A) =
-Sai Cai
0 0
and
C(7i ) =
0 -s0i] = 0 1 0
0 cpi
i 0 0 = 0
0
(211)
(212)
(213)
Here Eq(211) corresponds to a rotation of ctj (pitch) about the inertial axis ZQ (orbit
normal) of frame FQ resulting in a new intermediate frame F It is then followed by
a roll rotation fa about the axis y of F- giving a second intermediate frame F-
Final ly a spin rotation 7J about the axis z of F- is given resulting in the body
fixed frame Fj for the i lt l link Since all the three rotation matrices are orthogonal
it follows that FQ = C~1Fi = Cj F and hence
C CPi Cai ~ CH Sai + SrYi SPi Sai 5 7 i Sampi + Cli SPi Cci CPi Sai C1i Cai + 5 7 i SPi ~ S1 Ca + C 7 bull Sp Sai
(214) SliCPi CliCPi
Here Cx and Sx are abbreviations for cos(x) and sin(x) respectively
214 Inertial velocity of the ith link
mdash Lett ing pj = fj - f $j(5j and differentiating Eq(21) with respect to time gives
m- D j + T j f j + T^Si + T j pound + T j $ j 5 j (215)
Each of the terms appearing in Eq(215) must be addressed individually Since
fj represents the rigid body motion within the link it only has meaning for the
deployable tether and since yi = Zj = 0 for al l tether links (i even) fj = Ifi where 1
is the unit vector 1 0 0T For the case of rigid links (i odd) fj = 0
18
Evaluating the derivative of each angle in the rotation matrix gives
T^i9i = Taiai9i + Tppigi + T^fagi (216)
where Tx = J j I V Collecting the scaler angular velocity terms and defining the set
Vi = deg^gt A) 7i^ Eq(216) can be rewritten as
T i f t = Pi9i)fji (217)
where Pi(gi) is a column matrix defined by
Pi(9i)Vi = [Taigi^p^iT^gilffi (218)
Inserting Eq(29) and Eq(217) into Eq(215) and defining Sj = 2 + leads to
215 Cylindrical orbital coordinates
The Di term in Eq(219) describes the orbital motion of the t h l ink and
is composed of the three Cartesian coordinates DXi A ^ DZi However over the
cycle of one orbit DXi and Dyi vary dramatically by around the order of Earths
radius This must be avoided since large variations in the coordinates can cause
severe truncation errors during their numerical integration For this reason it is
more convenient to express the Cartesian components in terms of more stationary
variables This is readily accomplished using cylindrical coordinates
However it is only necessary to transform the first l inks orbital components
ie Lgti since only D is a generalized coordinates From Figure 2-2 it is apparent
(219)
that
(220)
19
Figure 2-2 Vector components in cylindrical orbital coordinates
20
Eq(220) can be rewritten in matrix form as
cos(0i) 0 0 s i n ( 0 i ) 0 0
0 0 1
Dri
h (221)
= DTlDSv
where Dri is the inplane radial distance 9 the true anomaly and DZl the out-of-
plane distance normal to the orbital plane The total derivative of D wi th respect
to time gives in column matrix form
1 1 1 1 J (222)
= D D l D 3 v
where [cos(^i) -Dnsm(8i) 0 ]
(223)
For all remaining links ie i = 2 N
cos(0i) - A - j S i n ^ i ) 0 sin(^i) D r i c o s ( 0 i ) 0
0 0 1
DTl = D D i = I d 3 i = 2N (224)
where is the n x n identity matrix and
2N (225)
216 Tether deployment and retrieval profile
The deployment of the tether is a critical maneuver for this class of systems
Cr i t i ca l deployment rates would cause the system to librate out of control Normally
a smooth sinusoidal profile is chosen (S-curve) However one would also like a long
tether extending to ten twenty or even a hundred kilometers in a reasonable time
say 3-4 orbits This is usually difficult or impossible to accomplish solely wi th one
S-curve profile Often a composite S-curve-Steady-S-curve profile is adopted Thus
21
the deployment scheme for the t h tether can be summarized as
k = Tr l - cos (^r(ti - to-)) 1 ltUlt h 2 r V A V
k = Vov tiilttiltt2i
k = IT j 1 - C deg S lkitl ^ ^ lt i lt ^
(226)
where to- tt are the ini t ia l and final times of the deployment or retrieval maneuver
respectively A t j = t - tQ mdash tt mdash t2- and the steady deployment velocity VQ is
calculated based on the continuity of l and l at the specified intermediate times t
and to For the case of
and
Ai = i(ti) - ^(t 0 -) = - ^ ( t 2 i ) (227)
2Ali + li(tf) -h(tQ)
V = lUli l^lL (228)
1 tf-tQ v y
A t = (229)
22 Kinetics and System Energy
221 Kinetic energy
The kinetic energy of the ith link is given by
Kei = lJm Rdrnfi^Ami (230)
Setting Eq(219) in a matrix-vector form gives the relation
4 ^ ^ PM) T i $ 4 TiSi]i[t (231)
22
where
m
k J
x l (232)
and nq = -nfj + 7 is the number of generalized coordinates in each link Inserting
Eq(231) into Eq(230) and integrating the kinetic energy for the t h l ink can be
rewritten as
1 T
(233)
where Mti is the links symmetric mass matrix
Mt =
m i D D T D D DDTPifgidmi) DDTTi J ^dm D n T Sidm rn
sym fP(9i)Pi(9i)dmi J P[g^T^idm J P ( pound ) T ^ m sym sym J ^f^idm f ^f^dmi sym sym sym J sfsidmi
pound Sfttiqxnq
(234)
and mi is the mass of the Ith link The kinetic energy for the entire system ie N
bodies can now be stated as
1 bull T bull 1 bull T Ke = J2 MtA = o laquo Mtqt 2 ^ -i
i=l (235)
where qt = qj^^ bull bull -qfNT a n d Mt is a block diagonal decoupled mass matrix
of the system with Mf- on its diagonal
(236) Mt =
0 0
0 Mt2
0
0 0
MH
0 0 0
0 0 0 bull MtN
222 Simplification for rigid links
The mass matrix given by Eq(234) simplifies considerably for the case of rigid
23
links and is given as
Mti
m D D l
T D D l DD^Piiffidmi) 0
sym 0
0
fP(fi)Pi(i)dmi 0 0 0
0 Id nfj
0 m~
i = 1 3 5 N
(237)
Moreover when mult iplying the matrices in the (11) block term of Eq(237) one
gets
D D l
1 D D l = 1 0 0 0 D n
2 0 (238) 0 0 1
which is a diagonal matrix However more importantly i t is independent of 6 and
hence is always non-singular ensuring that is always invertible
In the (12) block term of M ^ it is apparent that f fidmi is simply the
definition of the centre of mass of the t h rigid link and hence given by
To drrii mdash miff (239)
where fcmi is the position vector of the centre of mass of link i For the case of the
(22) block term a little more attention is required Expanding the integrand using
Eq(218) where ltj = fj gives
fj Ta Tafi fj T Q Tpfi f- T Q T 7 i f j
sym sym
Now since each of the terms of Eq(240) is a scaler or a 1 x 1 matrix it is therefore
equal to its trace ie sum of its diagonal elements thus
tr(ffTai
TTaifi) tr(ffTai
TTpfi) tr(ff Ta^T^fi)
fjT^Tpn rfTp^fi (240)
P(rj)Pj(fi) = syrri trifjTpTTpfi) t r ^ T ^ T ^
sym sym tr(ff T 7 T r 7 f j) J
(241)
Using two properties of the trace function[47] namely
tr(ABC) = tr(BCA) = trCAB) (242)
24
and
j tr(A) =tr(^J A^j (243)
where A B and C are real matrices it becomes quite clear that the (22) block term
of Eq(237) simplifies to
P(fi)Pii)dmi =
tr(Tai
TTaiImi) tTTai
TTpImi) t r (T Q r T 7 7 m ) trTaTTpImi) tr(Tp
TT7Jmi) sym sym sym
(244)
where Im- is an alternate form of the inertia tensor of the t h link which when
expanded is
Imlaquo = I nfi dm bull i sym
sym sym
J x^dmi J Xydmi XZdm sym J yi2dmi f yiZidm
sym J z^dmi J
xVi -L VH sym ijxxi + lyy IZZJ)2
(245)
Note this is in terms of moments of inertia and products of inertia of the t f l rigid
link A similar simplification can be made for the flexible case where T is replaced
with Qi resulting in a new expression for Im
223 Gravitational potential energy
The gravitational potential energy of the ith link due to the central force law
can be written as
V9i f dm f
-A mdash = -A bullgttradelti RdmA
dm
inn Lgti + T j ^ | (246)
where z = 3986 x 105 km 3s 2 (Earths gravitational constant) Expanding binomially
and retaining up to third order terms gives
Vgi ~ ~ D T 2D [J gfgidrrii + 23f T j gidm mdashDT
D2 i l I I Df Tt I grffdmiTiDi
(2-lt 7)
25
where D = D is the magnitude of the inertial position vector to the frame F The
integrals in the above expression are functions of the flexible mode shapes and can be
evaluated through symbolic manipulation using an existing package such as Maple
V In turn Maple V can translate the evaluated integrals into F O R T R A N and store
the code in a file Furthermore for the case of rigid bodies it can be shown quite
readily that
j gjgidrrii = j rjfidrrn = [IXXi + Iyy + IZZi)2 (248)
Similarly the other integrals in Eq(234) can be integrated symbolically
224 Strain energy
When deriving the elastic strain energy of a simple string in tension the
assumptions of high tension and low amplitude of transverse vibrations are generally
valid Consequently the first order approximation of the strain-displacement relation
is generally acceptable However for orbiting tethers in a weak gravitational-gradient
field neglecting higher order terms can result in poorly modelled tether dynamics
Geometric nonlinearities are responsible for the foreshortening effects present in the
tether These cannot be neglected because they account for the heaving motion along
the longitudinal axis due to lateral vibrations The magnitude of the longitudinal
oscillations diminish as the tether becomes shorter[35]
W i t h this as background the tether strain which is derived from the theory
of elastic vibrations[3248] is given by
(249)
where Uj V and W are the flexible deformations along the Xj and Z direction
respectively and are obtained from Eq(25) The square bracketed term in Eq(249)
represents the geometric nonlinearity and is accounted for in the analysis
26
The total strain energy of a flexible link is given by
Ve = I f CiEidVl (250)
Substituting the stress-strain relationship 0 = E(poundi the strain energy is now given
by
Vei=l-EiAi J l l d x h (251)
where E^ is the tethers Youngs modulus and A is the cross-sectional area The
tether is assumed to be uniform thus EA which is the flexural stiffness of the
tether is assumed to be constant Eq(251) is evaluated symbolically for arbitrary
number of modes
225 Tether energy dissipation
The evaluation of the energy dissipation due to tether deformations remains
a problem not well understood even today Here this complex phenomenon is repshy
resented through a simplified structural damping model[32] In addition the system
exhibits hysteresis which also must be considered This is accomplished using an
augmented complex Youngs modulus
EX^Ei+jEi (252)
where Ej^ is the contribution from the structural damping and j = The
augmented stress-strain relation is now given as
a = Epoundi = Ei(l+jrldi)ei (253)
where 77^ = EjE^ ltC 1 is defined as the structural damping coefficient determined
experimentally [49]
27
If poundi is a harmonic function with frequency UJQ- then
jet = (254) wo-
Substituting into Eq(253) and rearranging the terms gives
e + I mdash I poundi = o-i + ad (255)
where ad and a are the stress with and without structural damping respectively
Now the Rayleigh dissipation function[50] for the ith tether can be expressed as
(ii) Jo
_ 1 EjAiVdi rU (2-56^ 2 w 0 Jo
The strain rate ii is the time derivative of Eq(249) The generalized external force
due to damping can now be written as Qd = Q^ bull bull bull QdNT where
dWd
^ = - i p = 2 4 ( 2 5 7 )
= 0 i = 13 N
23 0(N) Form of the Equations of Mot ion
231 Lagrange equations of motion
W i t h the kinetic energy expression defined and the potential energy of the
whole system given by Pe = ]Cn=l( 5i+ ej) the equations of motion can be obtained
quite readily using the Lagrangian principle
where q is the set of generalized coordinates to be defined in a later section
Substituting Eqs(235247251 and 257) into Eq(258) leads to the familiar
matr ix form of the non-linear non-autonomous coupled equations of motion for the
28
system
Mq t)t+ F(q q t) = Q(q q t) (259)
where M(q t) is the nonlinear symmetric mass matrix and F(q q t) is the forcing
function which can be written in matrix form as
Q(ltfgt Qi 0 is the vector of non-conservative generalized forces including control inshy
puts acting on the system A detailed expansion of Eq(260) in tensor notation is
developed in Appendix I
Numerical solution of Eq(259) in terms of q would require inversion of the
mass matrix However M is a full matrix of size nqq x nqq where nqq = Nnq is the
total number of generalized coordinates Therefore direct inversion of M would lead
to a large number of computation steps of the order nqq^ or higher The objective
here is to develop an order-N O(N) algorithm for obtaining M _ 1 that minimizes
the computational effort This is accomplished through the following transformations
ini t ia l ly developed for the stationkeeping case by Pradhan et al [30]
232 Generalized coordinates and position transformation
During the derivation of the energy expressions for the system the focus was
on the decoupled system ie each link was considered independent of the others
Thus each links energy expression is also uncoupled However the constraints forces
between two links must be incorporated in the final equations of motion
Let
e Unqxl (261) m
be the set of coupled coordinates of the t h l ink such that q = qT q2 bull bull bull 0T i s the
29
systems generalized coordinates Let qt- be the set of auxiliary decoupled coordinates
defined by Eq(232) such that qt = qj^ qj^ Qtj^1 bull The only difference between
qtj and q is the presence of D against d is the position of F wi th respect to
the inertial frame FQ whereas d is defined as the offset position of F relative to and
projected on the frame It can be expressed as d = dXidyidZi^ For the
special case of link 1 D = d
From Figure 2-1 can written as
Di = A - l + T i _ i J i _ i + di) + T V i S i - i f t - i + dXi)8i- (262)
Denoting KAj = $j(Zj + dx- ) Eq(262) can be rewritten in summation form as
i ampi = H ( T i - 1 ^ + T i - l K A i - l - l + T j - l ^ - l ) bull ( 2 - 6 3 )
3 = 1
Introducing the index substitution k mdash j mdash 1 into Eq(263)
i-1 D = ^2 (Tfc_ilt4 + TkKAk6k + Tkilk) + T^di (264)
k=l
since d = D TQ = 1^ and IQ DQ SQ are null vectors Recasting Eq(264) in
matr ix form leads to i-1
Qt Jfe=l
where
and
For the entire system
T fc-1 0 TkKAk
0 0 0 0 0 0 0 0 0 0 0 0 -
r T i - i 0 0 0 0 0 0 0 0 0 G
0 0 0 1
Qt = RpQ
G R n lt x l
(265)
(266)
G Unq x l (267)
(268)
30
where
RP
R 1 0 0 R R 2 0 R RP
2 R 3
0 0 0
R RP R PN _ 1 RN J
(269)
Here RP is the lower block triangular matrix which relates qt and q
233 Velocity transformations
The two sets of velocity coordinates qti and qi are related by the following
transformations
(i) First Transformation
Differentiating Eq(262) with respect to time gives the inertial velocity of the
t h l ink as
D = A - l + Ti_iK + + K A i _ i ^ _ i
+ T j _ i ^ + Ti-ii + T i - i K A i - i ^ + T j - i K A ^ i ^ i (270)
(271) (
Defining K A D = dddXi+1KAi K A L = d d K A and hi = di+1 +lfi +KA^i
results in K A ^ i = K A D j i o ^ + K A L i _ i i _ i
= KADi_iiTdi + K A L j _ i ^ _ i
Inserting Eq(271) into Eq(270) and rearranging gives
= A _ + T i ^ d g + K A D ^ x ^ i ^ ^ + Pi-^hi-iH-i
+ T i _ i K A i _ i 5 _ i + T i_ i2 + K A L j _ i 5 j _ i 4 _ i (272)
Using Eq(272) to define recursively - D j _ i and applying the index substitution
c = j mdash 1 as in the case of the position transformation it follows that
Di = ^ T ^ i K D ^ ^ + Pk(hk)ffk + T f c K A f c ^ + TfcKLfc4 Jfc=l (273)
+ T j _ i K D j _ i lt i ii
31
where K D j _ i = + K A D ^ i ^ - i r 7 and K L j = i + K A L j 5 j Finally by following a
procedure similar to that used for the position transformation it can be shown that
for the entire system
t = Rvil (2-74)
where Rv is given by the lower block diagonal matrix
Here
Rv
Rf 0 0 R R$ 0 R Rl RJ
0 0 0
T3V nN-l
R bullN
e Mnqqxl
Ri =
T i _ i K D i _ i TiKAi T j K L j 0 0 0 0 bullo 0 0 0 0 0 0 0
e K n lt x l
RS
T i-1 0 0 o-0 0 0 0 0 0
0 0 0 1
G Mnq x l
(275)
(276)
(277)
(ii) Second Transformation
The second transformation is simply Eq(272) set into matrix form This leads
to the expression for qt as
(278)
where Id Pihi) T j K A j T j K L j 0 0 0 0 0 0
0 0
0 0
0 0
G $lnq x l
Thus for the entire system
(279)
pound = Rnjt + Rdq
= Vdnqq - R nrlRdi (280)
32
where
and
- 0 0 0 RI 0 0
Rn = 0 0
0 0
Rf 0 0 0 rgtd K2 0
Rd = 0 0 Rl
0 0 0
RN-1 0
G 5R n w x l
0 0 0 e 5 R N ^ X L
(281)
(282)
234 Cylindrical coordinate modification
Because of the use of cylindrical coorfmates for the first body it is necessary
to slightly modify the above matrices to account for this coordinate system For the
first link d mdash D = D ^ D S L thus
Qti = 01 =gt DDTigi (283)
where
D D T i
Eq(269) now takes the form
DTL 0 0 0 d 3 0 0 0 ID
0 0 nfj 0
RP =
P f D T T i R2 0 P f D T T i Rp
2 R3
iJfDTTi i^
x l
0 0 0
R PN-I RN
(284)
(285)
In the velocity transformation a similar modification is necessary since D =
DD^DS^ Mak ing the substitution it can be shown that Eq(276) and Eq(279) now
33
takes the form for i = 1
T)V _ pn _ r t j mdash rt^ mdash 0 0 0 0
0 0 0 0 0 0 0 0
(286)
The rest of the matrices for the remaining links ie i = 2 N are unchanged
235 Mass matrix inversion
Returning to the expression for kinetic energy substitution of Eq(274) into
Eq(235) gives the first of two expressions for kinetic energy as
Ke = -q RvT MtRv (287)
Introduction of Eq(280) into Eq(235) results in the second expression for Ke as
1X 2q (hnqq-RnrlRd Mt (idnqq - R nrlRd (288)
Thus there are two factorizations for the systems coupled mass matrix given as
M = RV MtRv
M (hnqq - RnrlRd Mt ( i d n q q - R n r 1 R d
(289)
(290)
Inverting Eq(290)
r-1 _ ( rgtd~^ M 1 = (Ra) l d n q q - Rn)Mf1 (Rd)~l(Idnqq - Rn) (291)
Since both Rd and Mt are block diagonal matrices their inverse is simply the inverse
34
of each block on the diagonal ie
Mr1
h 0 0
0 0
0 0
0 0 0
0
0
0
tN
(292)
Each block has the dimension nq x nq and is independent of the number of bodies
in the system thus the inversion of M is only linearly dependent on N ie an O(N)
operation
236 Specification of the offset position
The derivation of the equations of motion using the O(N) formulation requires
that the offset coordinate d be treated as a generalized coordinate However this
offset may be constrained or controlled to follow a prescribed motion dictated either by
the designer or the controller This is achieved through the introduction of Lagrange
multipliers[51] To begin with the multipliers are assigned to all the d equations
Thus letting f = F - Q Eq(259) becomes
Mq + f = P C A (293)
where A G K 3 - 7 V x l is the vector of Lagrange multipliers and Pc G s R n 9 x 3 A r is the
permutation matrix assigning the appropriate Aj to its corresponding d equation
Inverting M and pre-multiplying both sides by PcT gives
pcTM-pc
= ( dc + PdegTM-^f) (294)
where dc is the desired offset acceleration vector Note both dc and PcTM lf are
known Thus the solution for A has the form
A pcTM-lpc - l dc + PC1 M (295)
35
Now substituting Eq(295) into Eq(293) and rearranging the terms gives
q = S(f Ml) = -M~lf + M~lPc
which is the new constrained vector equation of motion with the offset specified by
dc Note that pre-multiplying M~l by P c T provides the rows of M _ 1 corresponding
to the d equations Similarly post-multiplying by Pc leads the columns of the matrix
corresponding to the d equations
24 Generalized Control Forces
241 Preliminary remarks
The treatment of non-conservative external forces acting on the system is
considered next In reality the system is subjected to a wide variety of environmental
forces including atmospheric drag solar radiation pressure Earth s magnetic field as
well as dissipative devices to mention a few However in the present model only
the effect of active control thrusters and momemtum-wheels is considered These
generalized forces wi l l be used to control the attitude motion of the system
In the derivation of the equations of motion using the Lagrangian procedure
the contribution of each force must be properly distributed among all the generalized
coordinates This is acheived using the relation
nu ^5 Qk=E Pern ^ (2-97)
m=l deg Q k
where Qk and qk are the kth elements of Q and ltf respectively[51] Fem is the
mth external force acting on the system at the location Rm from the inertial frame
and nu is the total number of actuators Derivation of the generalized forces due to
thrusters and momentum wheels is given in the next two subsections
pcTM-lpc (dc + PcTM- Vj (2-96)
36
242 Generalized thruster forces
One set of thrusters is placed on each subsatellite ie al l the rigid bodies
except the first (platform) as shown in Figure 2-3 They are capable of firing in the
three orthogonal directions From the schematic diagram the inertial position of the
thruster located on body 3 is given as
Defining
R3 = di + Tid2 + T2(hi + KA2lt52) + T 3 f c m 3
= d + dfY + T 3 f c m 3
df = Tjdj+i + Tj+1lj+ii + KAj+i5j+i)
the thruster position for body 5 is given as
(298)
(299)
R5 = di+ dfx + d 3 + T5fcm
Thus in general the inertial position for the thruster on the t h body can be given
(2100)
as i - 2
k=l i1 cm (2101)
Let the column matrix
[Q i iT 9 R-K=rR dqj lA n i
WjRi A R i (2102)
where i = 3 5 N and j mdash 1 2 i The vector derivative of R wi th respect to
the scaler components qk is stored in the kth column of Thus for the case of
three bodies Eq(297) for thrusters becomes
1 Q
Q3
3 -1
T 3 T t 2 (2103)
37
Satellite 3
mdash gt -raquo cm
D1=d1
mdash gt
a mdash gt
o
Figure 2-3 Inertial position of subsatellite thruster forces
38
where ft = 0TtaTt^T is the thrust vector for the t h l ink (tether i-1) Here
Ttn and Tta are the thrust forces along the inplane m and out-of-plane Z directions i Pi
respectively For the case of 5-bodies
Q
~Qgt1
0
0
( T ^ 2 ) (2104)
The result for seven or more bodies follows the pattern established by Eq(2104)
243 Generalized momentum gyro torques
When deriving the generalized moments arising from the momemtum-wheels
on each rigid body (satellite) including the platform it is easier to view the torques as
coupled forces From Figure 2-4 it is apparent that for the ith l ink the generalized
forces constituting the couple are
(2105)
where
Fei = Faij acting at ei = eXi
Fe2 = -Fpk acting at e 2 = C a ^ l
F e 3 = F7ik acting at e 3 = eyj
Expanding Eq(2105) it becomes clear that
3
Qk = ^ F e i
i=l
d
dqk (2106)
which is independent of Ri- Thus the moments Mma = 2FaieXi Mm^ = 2FpeXi
and Mmi = 2FlieVi Transforming the coordinates to the body fixed frame using
the rotation matrix the generalized force due to the momentum-wheels on link i can
39
Figure 2-4 Coupled force representation of a momentum-wheel on a rigid body
40
be written as
d d d Qi = Td bull ^ T i t M m a - Tik bull mdashT^Mmp + T ^ bull T J j M ^
QiMmi
where Mm = Mma M m Mmi T
lt3 T J ^ T - T i f c bull ^ T i pound T ^ - ^ T z j
(2107)
(2108)
Note combining the thruster forces and momentum-wheel torques a compact
expression for the generalized force vector for 5 bodies can be written as
Q
Q T 3 0 Q T 5 0
0 0 0
0 lt33T3 Qyen Q5
3T5 0
0 0 0 Q4T5 0
0 0 0 Q5
5T5
M M I
Ti 2 m 3
Quu (2109)
The pattern of Eq(2109) is retained for the case of N links
41
3 C O M P U T E R I M P L E M E N T A T I O N
31 Prel iminary Remarks
The governing equations of motion for the TV-body tethered system were deshy
rived in the last chapter using an efficient O(N) factorization method These equashy
tions are highly nonlinear nonautonomous and coupled In order to predict and
appreciate the character of the system response under a wide variety of disturbances
it is necessary to solve this system of differential equations Although the closed-
form solutions to various simplified models can be obtained using well-known analytic
methods the complete nonlinear response of the coupled system can only be obtained
numerically
However the numerical solution of these complicated equations is not a straightshy
forward task To begin with the equations are stiff[52] be they have attitude and
vibration response frequencies separated by about an order of magnitude or more
which makes their numerical integration suceptible to accuracy errors i f a properly
designed integration routine is not used
Secondly the factorization algorithm that ensures efficient inversion of the
mass matrix as well as the time-varying offset and tether length significantly increase
the size of the resulting simulation code to well over 10000 lines This raises computer
memory issues which must be addressed in order to properly design a single program
that is capable of simulating a wide range of configurations and mission profiles It
is further desired that the program be modular in character to accomodate design
variations with ease and efficiency
This chapter begins with a discussion on issues of numerical and symbolic in-
42
tegration This is followed by an introduction to the program structure Final ly the
validity of the computer program is established using two methods namely the conshy
servation of total energy and comparison of the response results for simple particular
cases reported in the literature
32 Numerical Implementation
321 Integration routine
A computer program coded in F O R T R A N was written to integrate the equashy
tions of motion of the system The widely available routine used to integrate the difshy
ferential equations D G E A R [ 5 3 ] is based on Gears predictor-corrector method[54]
It is well suited to stiff differential equations as i t provides automatic step-size adjustshy
ment and error-checking capabilities at each iteration cycle These features provide
superior performance over traditional 4^ order Runge-Kut ta methods
Like most other routines the method requires that the differential equations be
transformed into a first order state space representation This is easily accomplished
by letting
thus
X - ^ ^ f ^ ) =ltbullbull) (3 -2 )
where S(fMdc) is defined by Eq(296) Eq(32) represents the new set of 2nqq
first order differential equations of motion
322 Program structure
A flow chart representing the computer programs structure is presented in
Figure 3-1 The M A I N program is responsible for calling the integration routine It
also coordinates the flow of information throughout the program The first subroutine
43
called is INIT which initializes the system variables such as the constant parameters
(mass density inertia etc) and the ini t ial conditions I C (attitude angles orbital
motion tether elastic deformation etc) from user-defined data files Furthermore all
the necessary parameters required by the integration subroutine D G E A R (step-size
tolerance) are provided by INIT The results of the simulation ie time history of the
generalized coordinates are then passed on to the O U T P U T subroutine which sorts
the information into different output files to be subsequently read in by the auxiliary
plott ing package
The majority of the simulation running time is spent in the F C N subroutine
which is called repeatedly by D G E A R F C N generates the fv(Xi) vector defined in
Eq(32) It is composed of several subroutines which are responsible for the comshy
putation of all the time-varying matrices vectors and scaler variables derived in
Chapter 2 that comprise the governing equations of motion for the entire system
These subroutines have been carefully constructed to ensure maximum efficiency in
computational performance as well as memory allocation
Two important aspects in the programming of F C N should be outlined at this
point As mentioned earlier some of the integrals defined in Chapter 2 are notably
more difficult to evaluate manually Traditionally they have been integrated numershy
ically on-line (during the simulation) In some cases particularly stationkeeping the
computational cost is quite reasonable since the modal integrals need only be evalushy
ated once However for the case of deployment or retrieval where the mode shape
functions vary they have to be integrated at each time-step This has considerable
repercussions on the total simulation time Thus in this program the integrals are
evaluated symbolically using M A P L E V[55] Once generated they are translated into
F O R T R A N source code and stored in I N C L U D E files to be read by the compiler
Al though these files can be lengthy (1000 lines or more) for a large number of flexible
44
INIT
DGEAR PARAM
IC amp SYS PARAM
MAIN
lt
DGEAR
lt
FCN
OUTPUT
STATE ACTUATOR FORCES
MOMENTS amp THRUST FORCES
VIBRATION
r OFFS DYNA
ET MICS
Figure 3-1 Flowchart showing the computer program structure
45
modes they st i l l require less time to compute compared to their on-line evaluation
For the case of deployment the performance improvement was found be as high as
10 times for certain cases
Secondly the O(N) algorithm presented in the last section involves matrices
that contained mostly zero terms (RP Rv and Mplusmn) The sparse structure of these mashy
trices can be used advantageously to improve the computational efficiency by avoiding
multiplications of zero elements The result is a quick evaluation of fv(Xt) Moreshy
over there is a considerable saving of computer memory as zero elements are no
longer stored
Final ly the C O N T R O L subroutine which is part of F C N is designed to calshy
culate the appropriate control forces torques and offset dynamics which are required
to stabilize the librational ( A T T I T U D E ) and vibrational ( V I B R A T I O N ) motion of
the system
33 Verification of the Code
The simulation program was checked for its validity using two methods The
first verified the conservation of total energy in the absence of dissipation The second
approach was a direct comparison of the planar response generated by the simulation
program with those available in the literature A s numerical results for the flexible
three-dimensional A^-body model are not available one is forced to be content with
the validation using a few simplified cases
331 Energy conservation
The configuration considered here is that of a 3-body tethered system with
the following parameters for the platform subsatellite and tether
46
h =
1091430 -8135 328108 -8135 8646050 27116 328108 27116 8286760
kg bull m 2 platform inertia
kg bull m 2 end-satellite inertia 200 0 0
0 400 0 0 0 400
m = 90000 kg mass of the space station platform
m2 = 500 kg mass of the end-satellite
EfAt = 61645 N tether elastic stiffness
pt mdash 49 kg km tether linear density
The tether length is taken to remain constant at 10 km (stationkeeping case)
A n in i t ia l disturbance in pitch and roll of 2deg and 1deg respectively is given to both
the rigid bodies and the flexible tether In addition the tether is ini t ia l ly deflected
by 45 m in the longitudinal direction at its end and 05 m in both the inplane and
out-of-plane directions in the first mode The tether attachment point offset at the
platform (satellite 1) end is 1 m in all the three directions ie d2 = 11 lTm The
attachment point at the subsatellite (satellite 2) end is 1 m in the x direction ie
fcm3 = 10 0T
The variation in the kinetic and potential energies is presented for the above-
mentioned case in Figure 3-2(a) As seen from the plot there is a continuous mutual
exchange between the kinetic and potential energies however the variation in the
total energy remains zero (Figure 3-2b) It should be noted that after 1 orbit the
variation of both kinetic and potential energy does not return to 0 This is due to
the attitude and elastic motion of the system which shifts the centre of mass of the
tethered system in a non-Keplerian orbit
Similar results are presented for the 5-body case ie a double pendulum (Figshy
ure 3-3) Here the system is composed of a platform and two subsatellites connected
47
STSS configuration 3-Body a 1(0)=a 2(0)=a t(0)=2deg P1(0)=p2(0)=p(0)=1deg 5x(0)=45m 8y(0)=52(0)=05m dx(0)=dy(0)=d2(0)=1m Stationkeeping lt=10km
Variation of Kinetic and Potential Energy
(a) Time (Orbits) Percent Variation in Total Energy
OOEOh
-50E-12
-10E-11 I i i
0 1 2 (b) Time (Orbits)
Figure 3-2 Kinetic and potential energy transfer for the three-body platform based tethered satellite system (a) variation of kinetic and potential energy (b) percent change in total energy of system
48
STSS configuration 5-Body a 1(0)=a 2(0)=a t 1(0)=2deg a 3(0)=a t 2(0)=25deg P1(0)=p2(0)=p3(0)=pt1(0)=pt2(0)=1lt
5x(0)=5m 8y(0)=5z(0)=05m dx(0)=dy(0)=dz(0)=0 Stationkeeping 1 =12=1 Okm
(a)
20E6
10E6r-
00E0
-10E6
-20E6
Variation of Kinetic and Potential Energy
-
1 I I I 1 I I
I I
AP e ~_
i i 1 i i i i
00E0
UJ -50E-12 U J
lt
Time (Orbits) Percent Variation in Total Energy
-10E-11h
(b) Time (Orbits)
Figure 3 -3 Kinet ic and potential energy transfer for the five-body platform based tethered satellite system (a) variation of kinetic and potential enshyergy (b) percent change in total energy of system
49
in sequence by two tethers The two subsatellites have the same mass and inertia
ie 7713 = 7 7 i 2 a n d I3 = I21 a n d are separated by identical tether segments each 10
k m in length The platform has the same properties as before however there is no
tether attachment point offset present in the system (d = fcmi = 0)
In al l the cases considered energy was found to be conserved However it
should be noted that for the cases where the tether structural damping deployment
retrieval attitude and vibration control are present energy is no longer conserved
since al l these features add or subtract energy from the system
332 Comparison with available data
The alternative approach used to validate the numerical program involves
comparison of system responses with those presented in the literature A sample case
considered here is the planar system whose parameters are the same as those given
in Section 331 A n ini t ia l pitch disturbance of 2deg is imparted to both the platform
and the tether together with an ini t ial tether deformation of 08 m and 001 m in
the longitudinal and transverse directions respectively in the same manner as before
The tether length is held constant at 5 km and it is assumed to be connected to the
centre of mass of the rigid platform and subsatellite The response time histories from
Ref[43] are compared with those obtained using the present program in Figures 3-4
and 3-5 Here ap and at represent the platform and tether pitch respectively and B i
C i are the tethers longitudinal and transverse generalized coordinates respectively
A comparison of Figures 3-4 and 3-5 clearly shows a remarkable correlation
between the two sets of results in both the amplitude and frequency The same
trend persisted even with several other comparisons Thus a considerable level of
confidence is provided in the simulation program as a tool to explore the dynamics
and control of flexible multibody tethered systems
50
a p(0) = 2deg 0(0) = 2C
6(0) = 08 m
0(0) = 001 m
Stationkeeping L = 5 km
D p y = 0 D p 2 = 0
a oo
c -ooo -0 01
Figure 3 - 4 Simulation results for the platform based three-body tethered system originally presented in Ref[43]
51
STSS configuration 3-Body a1(0)=a2(0)=at(0)=2deg p1(0)=(32(0)=(3t(0)=0 8x(0)=08m 5y(0)=001m 5Z(0)=0 dx(0)=dy(0)=dz(0)=0 Stationkeeping l=5km
Satellite 1 Pitch Angle Tether Pitch Angle
i -i i- i i i i i i d
1 2 0 1 2 Time (Orbits) Time (Orbits)
Longitudinal Vibration
^MAAAAAAAWVWWWWVVVW
V bull i bull i I i J 00 05 10 15 20
Time (Orbits) Transverse Vibration
Time (Orbits)
Figure 3-5 Simulation results for the platform based three-body tethered system obtained using the present computer program
52
4 D Y N A M I C SIMULATION
41 Prel iminary Remarks
Understanding the dynamics of a system is critical to its design and develshy
opment for engineering application Due to obvious limitations imposed by flight
tests and simulation of environmental effects in ground based facilities space based
systems are routinely designed through the use of numerical models This Chapter
studies the dynamical response of two different tethered systems during deployment
retrieval and stationkeeping phases In the first case the Space platform based Tethshy
ered Satellite System (STSS) which consists of a large mass (platform) connected to
a relatively smaller mass (subsatellite) with a long flexible tether is considered The
other system is the O E D I P U S B I C E P S configuration involving two comparable mass
satellites interconnected by a flexible tether In addition the mission requirement of
spin about an arbitrary axis the tether axis for O E D I P U S - A C and cartwheeling or
spin about the orbit normal for the proposed B I C E P S mission is accounted for
A s mentioned in Chapter 2 the tether flexibility is modelled using the asshy
sumed mode discretization method Although the simulation program can account
for an arbitrary number of vibrational modes in each direction only the first mode
is considered in this study as it accounts for most of the strain energy[34] and hence
dominates the vibratory motion The parametric study considers single as well as
double pendulum type systems with offset of the tether attachment points The
systems stability is also discussed
42 Parameter and Response Variable Definitions
The system parameters used in the simulation unless otherwise stated are
53
selected as follows for the STSS configuration
1091430 -8135 328108 -8135 8646050 27116 328108 27116 8286760
h =
h = 200 0 0
0 400 0 0 0 400
bull m = 90000 kg (mass of the space platform)
kg bull m 2 (platform inertia) 760 J
kg bull m 2 (subsatellite inertia)
bull 7772 = 500 kg (mass of the subsatellite)
bull EtAt = 61645 N (tether stiffness)
bull Pt = 49 k g k m (tether density)
bull Vd mdash 05 (tether structural damping coefficient)
bull fcm^ = l 0 0 r m (tether attachment point at the subsatellite)
The response variables are defined as follows
bull ai0 satellite 1 (platform) pitch and roll angles respectively
bull CK2 2 satellite 2 (subsatellite) pitch and roll angles respectively
bull at fit- tether pitch and roll angles respectively
bull If tether length
bull d = dx dy dzT- tether attachment position relative to satellite 1 (platform)
mdash
bull S = 8X 8y Sz tether flexible generalized coordinates in the longitudinal x
inplane transverse y and out-of-plane transverse z directions respectively
The attitude angles and are measured with respect to the Local Ve r t i c a l -
Loca l Horizontal ( L V L H ) frame A schematic diagram illustrating these variable
is presented in Figure 4-1 The system is taken to be in a nominal circular orbit at
an altitude of 289 km with an orbital period of 903 minutes
54
Satellite 1 (Platform) Y a w
Figure 4-1 Schematic diagram showing the generalized coordinates used to deshyscribe the system dynamics
55
43 Stationkeeping Profile
To facilitate comparison of the simulation results a reference case based on
the three-body STSS system with zero offset (d = 0 r c m 3 = 0) and a tether length
of It = 20 km is considered first (Figure 4-2) The system is subjected to an ini t ial
disturbance in pitch and roll of 2deg and 1deg respectively to all the three bodies
In addition an ini t ia l longitudinal deflection of 14 m from the tethers unstretched
position is given together with a transverse inplane and out-of-plane deflection of 1
m at the tethers mid-length From the attitude response given in Figure 4-2(a) it is
apparent that the three bodies oscillate about their respective equilibrium positions
Since coupling between the individual rigid-body dynamics is absent (zero offset) the
corresponding librational frequencies are unaffected Satellite 1 (Platform) displays
amplitude modulation arising from the products of inertia The result is a slight
increase of amplitude in the pitch direction and a significantly larger decrease in
amplitude in the roll angle
Figure 4-2(b) clearly shows coupling of the tethers attitude motion with its
flexibility dynamics The tether vibrates in the axial direction about its equilibrium
position (at approximately 138 m) with two vibration frequencies namely 012 Hz
and 31 x 1 0 - 4 Hz The former is the first longitudinal flexible mode which is dissishy
pated through the structural damping while the latter arises from the coupling with
the pitch motion of the tether Similarly in the transverse direction there are two
frequencies of oscillations arising from the flexible mode and its coupling wi th the
attitude motion As expected there is no apparent dissipation in the transverse flexshy
ible motion This is attributed to the weak coupling between the longitudinal and
transverse modes of vibration since the transverse strain in only a second order effect
relative to the longitudinal strain where the dissipation mechanism is in effect Thus
there is very litt le transfer of energy between the transverse and longitudinal modes
resulting in a very long decay time for the transverse vibrations
56
STSS configuration 3-Body a1(0)=a2(0)=a(0)=2deg (31(0)=p2(0)=p(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=dz(0)=0 Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
0 1 2 3 4 5 0 1 2 3 4 5 Time (Orbits) Time (Orbits)
Satellite 2 Pitch Angle Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-2 Stationkeeping dynamics of the three-body STSS configuration withshyout offset (a) attitude response
57
STSS configuration 3-Body a1(0)=cx2(0)=cx(0)=2deg P1(0)=p2(0)=p(0)=1deg 5x(0)=14m8y(0)=82(0)=1m dx(0)=dy(0)=dz(0)=0 Stationkeeping lt=20km
Tether Longitudinal Vibration 16| 1 1 1 1 1 r-
Time (Orbits) Tether In-Plane Transverse Vibration
imdash 1 mdash 1 mdash 1 mdash 1 mdashr
r I I 1 1 1 I 1 1 1 1 I 1 bull ^ bull bull I I J 0 1 2 3 4 5
Time (Orbits) Tether Out-of-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-2 Stationkeeping dynamics of the three-body STSS configuration withshyout offset (b) vibration response
58
Introduction of the attachment point offset significantly alters the response of
the attitude motion for the rigid bodies W i t h a i m offset along the local vertical
at the platform (dx = 1 m) and fcm^ = 100 m coupling between the tether
and end-bodies is established by providing a lever arm from which the tether is able
to exert a torque that affects the rotational motion of the rigid end-bodies (Figure
4-3) Bo th the platform and subsatellite now oscillate at the same frequency as the
tether whose motion remains unaltered In the case of the platform there is also an
amplitude modulation due to its non-zero products of inertia as explained before
However the elastic vibration response of the tether remains essentially unaffected
by the coupling
Providing an offset along the local horizontal direction (dy = 1 m) results in
a more dramatic effect on the pitch and roll response of the platform as shown in
Figure 4-4(a) Now the platform oscillates about its new equilibrium position of
-90deg The roll motion is also significantly disturbed resulting in an increase to over
10deg in amplitude On the other hand the rigid body dynamics of the tether and the
subsatellite as well as the tether flexible motion remain the same (Figure 4-4b)
Figure 4-5 presents the response when the offset of 1 m at the platform end is
along the z direction ie normal to the orbital plane The result is a larger amplitude
pitch oscillation about the reference equilibrium position as shown in Figure 4-5(a)
while the roll equilibrium is now shifted to approximately 90deg Note there is little
change in the attitude motion of the tether and end-satellites however there is a
noticeable change in the out-of-plane transverse vibration of the tether (Figure 4-5b)
which may be due to the large amplitude platform roll dynamics
Final ly by setting a i m offset in all the three direction simultaneously the
equil ibrium position of the platform in the pitch and roll angle is altered by approxishy
mately 30deg (Figure 4-6a) However the response of the system follows essentially the
59
STSS configuration 3-Body a1(0)=cx2(0)=cxt(0)=2deg p1(0)=(32(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=1m dy(0)=dz(0)=0 Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
Time (Orbits) Time (Orbits) Tether Pitch Angle Tether Roll Angle
L bull I I bull bull lt bull bull bull raquo bull bull bull bull -I h I I I I i H 0 1 2 3 4 5 0 1 2 3 4 5
Time (Orbits) Time (Orbits) Satellite 2 Pitch Angle Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-3 Stationkeeping dynamics of the three-body STSS configuration with offset along the local vertical (a) attitude response
60
STSS configuration 3-Body a 1(0)=a 2(0)=a t(0)=2deg (31(0)=p2(0)=pt(0)=1deg 5x(0)=14m5y(0)=5z(0)=1m dx(0)=1mdy(0)=dz(0)=0 Stationkeeping lt=20km
Tether Longitudinal Vibration 1 6 i mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash mdash i mdash i mdash mdash mdash mdash mdash i mdash mdash lt -
Time (Orbits)
Tether In-Plane Transverse Vibration
y bull i i i i i i i i i i i i i i i i i 0 1 2 3 4 5
Time (Orbits)
Tether Out-of-Plane Transverse Vibration r 1 1 1 1 1 1 1 1 1 1 1 1 1 1 q
h i I i i I i i i i I i i i i 3
0 1 2 3 4 5 (b) Time (Orbits)
Figure 4-3 Stationkeeping dynamics of the three-body STSS configuration with offset along the local vertical (b) vibration response
61
STSS configuration 3-Body a1(0)=cx2(0)=o(0)=2deg (31(0)=p2(0)=p(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dz(0)=0 dy(0)=1m Stationkeeping lt=20km
Satellite 1 Pitch Angle i 1 i
Satellite 1 Roll Angle
1 2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 4 Time (Orbits)
Tether Roll Angle
CD
edT
1 2 3 Time (Orbits)
Satellite 2 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 2 Roll Angle
CD
C O
(a)
Figure 4-
1 2 3 4 Time (Orbits)
1 2 3 4 Time (Orbits)
4 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal (a) attitude response
62
STSS configuration 3-Body a1(0)=cx2(0)=ot(0)=2deg p1(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dz(0)=0 dy(0)=1m Stationkeeping lt=20km
Tether Longitudinal Vibration 16| 1 gt 1 bull 1 1 1
0 1 2 3 4 5
Time (Orbits) Tether In-Plane Transverse Vibration
i mdash 1 mdash 1 mdash 1 mdash 1 mdash i mdash mdash lt mdash 1 mdash 1 mdash i mdash mdash 1 mdash 1 mdash
V I I bull bull bull I i i i I i i 1 1 3
0 1 2 3 4 5 Time (Orbits)
Tether Out-of-Plane Transverse Vibration r 1 1 1 1 1 1 1 1 1 1 1 gt bull | q
V i 1 1 1 1 1 mdash 1 mdash 3
0 1 2 3 4 5 (b) Time (Orbits)
Figure 4-4 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal (b) vibration response
63
STSS configuration 3-Body a1(0)=o2(0)=at(0)=2deg (31(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=0 dz(0)=1m Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 4 Time (Orbits)
Tether Roll Angle
1 2 3 4 Time (Orbits)
Satellite 2 Pitch Angle
2 3 4 Time (Orbits)
Satellite 2 Roll Angle
(a)
Figure 4-
1 2 3 4 Time (Orbits)
2 3 4 Time (Orbits)
5 Stationkeeping dynamics of the three-body STSS configuration with offset along the orbit normal (a) attitude response
64
STSS configuration 3-Body a1(0)=cx2(0)=ot(0)=2deg P1(0)=p2(0)=(3(0)=1deg 8x(0)=14m8y(0)=82(0)=1m dx(0)=dy(0)=0 d2(0)=1m Stationkeeping lt=20km
Tether Longitudinal Vibration 1 6 i mdash r mdash i mdash mdash i mdash i mdash mdash i mdash i mdash mdash i mdash mdash i mdash 1 mdash 1 mdash i mdash mdash lt mdash lt mdash lt mdash r
Time (Orbits) Tether In-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-5 Stationkeeping dynamics of the three-body STSS configuration with offset along the orbit normal (b) vibration response
65
same trend with only minor perturbations in the transverse elastic vibratory response
of the tether (Figure 4-6b)
44 Tether Deployment
Deployment of the tether from an ini t ial length of 200 m to 20 km is explored
next Here the deployment length profile is critical to ensure the systems stability
It is not desirable to deploy the tether too quickly since that can render the tether
slack Hence the deployment strategy detailed in Section 216 is adopted The tether
is deployed over 35 orbits with the sinusoidal acceleration and deceleration occurring
over a 4 km length ( A ^ 2 ) with the remaining deployment at a constant velocity of
V0 = 146 ms
Initially the tether is stretched only slightly S = 1304 x 1 0 _ 3 0 0 m
with al l the other states of the system ie pitch roll and transverse displacements
remaining zero The response for the zero offset case is presented in Figure 4-7 For
the platform there is no longer any coupling with the tether hence it oscillates about
the equilibrium orientation determined by its inertia matrix However there is st i l l
coupling between the subsatellite and the tether since rcm^ mdash 100 m resulting
in complete domination of the subsatellite dynamics by the tether Consequently
the pitch motion ini t ial ly grows by almost 50deg but eventually subsides as the tether
deployment rate decreases It is interesting to note that the out-of-plane motion is not
affected by the deployment This is because the Coriolis force which is responsible
for the tether pitch motion does not have a component in the z direction (as it does
in the y direction) since it is always perpendicular to the orbital rotation (Q) and
the deployment rate (It) ie Q x It
Similarly there is an increase in the amplitude of the transverse elastic oscilshy
lations of the tether again due to the Coriolis effect (Figure 4-7b) Moreover as the
66
STSS configuration 3-Body a1(0)=a2(0)=cxt(0)=2deg P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m5y(0)=52(0)=1m dx(0)=dy(0)=d2(0)=1m Stationkeeping lt=20km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
i i i i i i i i i i i 0 1 2 3 4 5 0 1 2 3 4 5
Time (Orbits) Time (Orbits) Tether Pitch Angle Tether Roll Angle
l i i i i -I h i i i i I i i i i l- i i i i l i i i i I i i- i i H 0 1 2 3 4 5 0 1 2 3 4 5
Time (Orbits) Time (Orbits) Satellite 2 Pitch Angle Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-6 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal local vertical and orbit normal (a) attitude response
67
STSS configuration 3-Body a1(0)=cc2(0)=a(0)=2deg p1(0)=p2(0)=(3t(0)=1deg 5x(0)=14m6y(0)=52(0)=1m dx(0)=dy(0)=d2(0)=1m Stationkeeping lt=20km
Tether Longitudinal Vibration 1 6 1 mdash i mdash i mdash i mdash mdash i mdash 1 mdash 1 mdash mdash mdash i mdash mdash 1 mdash 1 mdash 1 mdash i mdash 1 mdash 1 mdash 1 mdash 1 mdash r
0 1 2 3 4 5
Time (Orbits) Tether In-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-6 Stationkeeping dynamics of the three-body STSS configuration with offset along the local horizontal local vertical and orbit normal (b) vibration response
68
STSS configuration 3-Body a(0)=a2(0)=ot(0)=0 P1(0)=p2(0)=pt(0)=0 8X(0)=1304 x 103m 8y(0)=8z(0)=0 dx(0)=dy(0)=dz(0)=0 Deployment
l=02km to 20km in 35 Orbits
Tether Length Profile
E
Satellite 1 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 1 Roll Angle mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash [ mdash i mdash i mdash i mdash i mdash | mdash
1 2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 4 Time (Orbits)
Tether Roll Angle
CD
oo ax
i imdashimdashimdashImdashimdashimdashimdashr-
_ 1 _ L
1 2 3 4 Time (Orbits)
Satellite 2 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 2 Roll Angle
-50
(a) 1 2 3 4
Time (Orbits) 1 2 3 4
Time (Orbits)
Figure 4-7 Deployment dynamics of the three-body STSS configuration without offset (a) attitude response
69
STSS configuration 3-Body a1(0)=a2(0)=at(0)=0 p1(0)=p2(0)=(3t(0)=0 8X(0)=1304x 103m 8y(0)=52(0)=0 dx(0)=dy(0)=d2(0)=0 Deployment
l=02km to 20km in 35 Orbits
Tether Longitudinal Vibration
to
-02
(b)
Figure 4-7
2 3 Time (Orbits)
Tether In-Plane Transverse Vibration
0 1 1 bull 1 1
1 2 3 4 Time (Orbits)
Tether Out-of-Plane Transverse Vibration
5
1 1 1 1 1 i i | i i | ^
_ i i i i L -I
1 2 3 4 5 Time (Orbits)
Deployment dynamics of the three-body STSS configuration without offset (b) vibration response
70
tether elongates its longitudinal static equilibrium position also changes due to an
increase in the gravity-gradient tether tension It may be pointed out that as in the
case of attitude motion there are no out-of-plane tether vibrations induced
When the same deployment conditions are applied to the case of 1 m offset in
the x direction (Figure 4-8) coupling effects are re-introduced The platform pitch
exceeds -100deg during the peak deployment acceleration However as the tether pitch
motion subsides the platform pitch response returns to a smaller amplitude oscillation
about its nominal equilibrium On the other hand the platform roll motion grows
to over 20deg in amplitude in the terminal stages of deployment Longitudinal and
transverse vibrations of the tether remain virtually unchanged from the zero offset
case except for minute out-of-plane perturbations due to the platform librations in
roll
45 Tether Retrieval
The system response for the case of the tether retrieval from 20 km to 200 m
with an offset of 1 m in the x and z directions at the platform end is presented in
Figure 4-9 Here the Coriolis force induced during retrieval renders the tether unstable
(Figure 4-9a) Consequently the pitch motion of the platform is also destabilized
through coupling The z offset provides an additional moment arm that shifts the
roll equilibrium to 35deg however this degree of freedom is not destabilized From
Figure 4-9(b) it is apparent that the transverse mode of the tether vibration is also
disturbed producing a high frequency response in the final stages of retrieval This
disturbance is responsible for the slight increase in the roll motion for both the tether
and subsatellite
46 Five-Body Tethered System
A five-body chain link system is considered next To facilitate comparison with
71
STSS configuration 3-Body a(0)=a2(0)=at(0)=0 P1(0)=P2(0)=P(0)=0 5X(0)=1304 x 103m 8y(0)=8z(0)=0 dx(0)=1mdy(0)=d2(0)=0 Deployment
lt=02km to 20km in 35 Orbits
Tether Length Profile
E
Satellite 1 Pitch Angle
1 2 3 Time (Orbits)
Satellite 1 Roll Angle
bull bull bull i bull bull bull bull i
1 2 3 4 Time (Orbits)
Tether Pitch Angle
1 2 3 Time (Orbits)
Tether Roll Angle
i i i i J I I i _
1 2 3 4 Time (Orbits)
Satellite 2 Pitch Angle
1 2 3 4 Time (Orbits)
Satellite 2 Roll Angle 3E-3h
D )
T J ^ 0 E 0 eg
CQ
-3E-3
(a) 1 2 3 4
Time (Orbits) 1 2 3 4
Time (Orbits)
Figure 4-8 Deployment dynamics of the three-body STSS configuration with offset along the local vertical (a) attitude response
7 2
STSS configuration 3-Body a1(0)=a2(0)=at(0)=0 p1(0)=P2(0)=(3t(0)=0 Sx(0)=1304 x 103m 5y(0)=5z(0)=0 dx(0)=1mdy(0)=dz(0)=0 Deployment
l=02km to 20km in 35 Orbits
Tether Longitudinal Vibration
Time (Orbits) Tether In-Plane Transverse Vibration
n 1 1 1 f
I I I _j I I I I I 1 1 1 1 1 2 3 4 5
Time (Orbits) Tether Out-of-Plane Transverse Vibration
J 1 I I 1 I I I I I I I I L _ _ I I d 1 2 3 4 5
Time (Orbits)
Deployment dynamics of the three-body STSS configuration with offset along the local vertical (b) vibration response
(b)
Figure 4-8
73
STSS configuration 3-Body a1(0)=a2(0)=at(0)=0 p(0)=P2(0)=(3t(0)=0 Sx(0)=14m8y(0)=82(0)=0 dy(0)=0dx(0)=d2(0)=1m Retrieval
l=20km to 02km in 35 Orbits
Satellite 1 Pitch Angle
(a) Time (Orbits)
Figure 4-9 Retrieval dynamics of the along the local vertical ar
Tether Length Profile
i i i i i i i i i i I 0 1 2 3
Time (Orbits) Satellite 1 Roll Angle
Time (Orbits) Tether Roll Angle
b_ i i I i I i i L _ J
0 1 2 3 Time (Orbits)
Satellite 2 Roll Angle _ mdash | r i i | 1 i i imdashmdashj i 1 r mdash i mdash
04- -
0 1 2 3 Time (Orbits)
three-body STSS configuration with offset d orbit normal (a) attitude response
74
STSS configuration 3-Body a1(0)=a2(0)=ocl(0)=0 (31(0)=f32(0)=pt(0)=0 8x(0)=14m8y(0)=6z(0)=0 dy(0)=0dx(0)=dz(0)=1m Retrieval
l=20km to 02km in 35 Orbits
Tether Longitudinal Vibration
E 10 to
Time (Orbits) Tether In-Plane Transverse Vibration
1 2 Time (Orbits)
Tether Out-of-Plane Transverse Vibration
(b) Time (Orbits)
Figure 4-9 Retrieval dynamics of the three-body STSS configuration with offset along the local vertical and orbit normal (b) vibration response
75
the three-body system dynamics studied earlier the chain is extended through the
addition of two bodies a tether and a subsatellite with the same physical properties
as before (Section 42) Thus the five-body system consists of a platform (satellite 1)
tether 1 subsatellite 1 (satellite 2) tether 2 and subsatellite 2 (satellite 3) as shown
in Figure 4-10 The offset of tether 1 at the platform-end is denoted by d2 while that
of tether 2 to subsatellite 1 as 04 In this numerical example fcm^ = fcm^ = 0 ie
tethers 1 and 2 are attached to the centre of mass of subsatellites 1 and 2 respectively
The system response for the case where the tether attachment points coincide
wi th the centres of mass of the rigid bodies ie zero offset (d2 = d mdash 0) is presented
in Figure 4-11 Here ogtt and at2 represent the pitch motion of tethers 1 and 2
respectively whereas 03 is the pitch motions of satellite 3 A similar convention
is adopted for the rol l angle 0 A s before the zero offset eliminates the coupling
between the tethers and satellites such that they are now free to librate about their
static equil ibrium positions However their is s t i l l mutual coupling between the two
tethers This coupling is present regardless of the offset position since each tether is
capable of transferring a force to the other The coupling is clearly evident in the pitch
response of the two tethers (Figure 4 - l l a ) O n the other hand the roll motion appears
uncoupled as the in i t ia l conditions in rol l for the two tethers are identical Thus the
motion is in phase and there is no transfer of energy through coupling A s expected
during the elastic response the tethers vibrate about their static equil ibrium positions
and mutually interact (Figure 4 - l l b ) However due to the variation of tension along
the tether (x direction) they do not have the same longitudinal equilibrium Note
relatively large amplitude transverse vibrations are present particularly for tether 1
suggesting strong coupling effects with the longitudinal oscillations
76
Figure 4-10 Schematic diagram of the five-body system used in the numerical example
77
STSS configuration 5-Body a1(0)=(x2(0)=at1(0)=2deg a3(0)=a12(0)=25deg p1(0)=r32(0)=p3(0)=pt1(0)=pt2(0)=r 82(0)=84(0)=505)05m d2(0)=d4(0)=000 Stationkeeping lt1=lu=10km Pitch Roll
Satellite 1 Pitch Angle Satellite 1 Roll Angle
o 1 2 Time (Orbits)
Tether 1 Pitch and Roll Angle
0 1 2 Time (Orbits)
Tether 2 Pitch and Roll Angle
Time (Orbits) Satellite 2 Pitch and Roll Angle
Time (Orbits) Satellite 3 Pitch and Roll Angle
Time (Orbits) Time (Orbits)
Figure 4-11 Stationkeeping dynamics of the five-body STSS configuration withshyout offset (a) attitude response
78
STSS configuration 5-Body a1(0)=a2(0)=oct1(0)=2deg a3(0)=at2(0)=25o
p1(0)=p2(0)=(33(0)=pt1(0)=pt2(0)=10
62(0)=54(0)=50505m d2(0)=d4(0)=000 Stationkeeping lt1=1 =1 Okm
Tether 1 Longitudinal Vibration Tether 2 Longitudinal Vibration i i 1 i 1 r 1 1 1 1
(b) Time (Orbits) Time (Orbits)
Figure 4-11 Stationkeeping dynamics of the five-body STSS configuration withshyout offset (b) vibration response
79
The system dynamics during deployment of both the tethers in the double-
pendulum configuration is presented in Figure 4-12 Deployment of each tether takes
place from 200 m to 20 km in 35 orbits The pitch motion of the system is similar
to that observed during the three-body case The Coriolis effect causes the tethers
to pitch through a large angle ( laquo 50deg ) which in turn disturbs the platform and
subsatellite due to the presence of offset along the local vertical On the other hand
the roll response for both the tethers is damped to zero as the tether deploys in
confirmation with the principle of conservation of angular momentum whereas the
platform response in roll increases to around 20deg Subsatellite 2 remains virtually
unaffected by the other links since the tether is attached to its centre of mass thus
eliminating coupling Finally the flexibility response of the two tethers presented in
Figure 4-12(b) shows similarity with the three-body case
47 BICEPS Configuration
The mission profile of the Bl-static Canadian Experiment on Plasmas in Space
(BICEPS) is simulated next It is presently under consideration by the Canadian
Space Agency To be launched by the Black Brant 2000500 rocket it would inshy
volve interesting maneuvers of the payload before it acquires the final operational
configuration As shown in Figure 1-3 at launch the payload (two satellites) with
an undeployed tether is spinning about the orbit normal (phase 1) The internal
dampers next change the motion to a flat spin (phase 2) which in turn provides
momentum for deployment (phase 3) When fully deployed the one kilometre long
tether will connect two identical satellites carrying instrumentation video cameras
and transmitters as payload
The system parameters for the BICEPS configuration are summarized below [59 0 0 1
bull I = I2 = 0 366 0 kg bull m 2 (satellite inertias) [ 0 0 392_
bull mi = 7772 = 200 kg (mass of the satellites)
80
STSS configuration 5-Body a(0)=cx2(0)=al1(0)=2deg a3(0)=c (0)=25o
P1(0)=P2(0)=p3(0)=f3t1(0)=Pt2(0)=r 52(0)=84(0)=1304x 10300m d2(0)=d4(0)=1G0m Deployment lt1=le=02km to 20km in 35 Orbits
Pitch Roll Satellite 1 Pitch Angle
Time (Orbits) Satellite 2 Pitch and Roll Angle T 1mdash1 1 1 I I I I | I I I 1 I | I I lt~r
Tether Length Profile i 1 1 1 1 i 1 1 1 1 i 1 1 1 1 i
i i i i i 0 1 2 3 4 5
Time (Orbits) Satellite 1 Roll Angle
0 1 2 3 4 5 Time (Orbits)
Tether 1 amp 2 Roll Angle L i | I I 1 I 1 1 -I
-Ih I- bull I I i i i i I i i i i 1
0 1 2 3 4 5 Time (Orbits)
Satellite 3 Pitch and Roll Angle i- i 1 1 i 1 1 1 1 i 1 1 1 1 i 1 1 1 1 i
L i i i i i i 1 i i mdash i mdash i I i imdashimdashLJ
0 1 2 3 4 5 Time (Orbits)
Figure 4-12 Deployment dynamics of the five-body S T S S configuration with offshyset along the local vertical (a) attitude response
81
STSS configuration 5-Body a1(0)=a2(0)=oct1(0)=2deg a3(0)=at2(0)=25deg P1(0)=P2(0)=P3(0)=Pt1(0)=pt2(0)=1o
82(0)=54(0)=1304 x 10-300m d2(0)=d4(0)=100m Deployment lt1=l2=02km to 20km in 35 Orbits
Tether 1 Longitudinal Vibration Tether 2 Longitudinal Vibration
Time (Orbits) Time (Orbits) Tether 1 Z Tranverse Vibration Tether 2 Z Tranverse Vibration
Figure 4-12 Deployment dynamics of the five-body STSS configuration with offshyset along the local vertical (b) vibration response
82
bull EtAt = 61645 N (tether stiffness)
bull pt = 30 k g k m (tether density)
bull It = 1 km (tether length)
bull rid = 1-0 (tether structural damping coefficient)
The system is in a circular orbit at a 289 km altitude Offset of the tether
attachment point to the satellites at both ends is taken to be 078 m in the x
direction The response of the system in the stationkeeping phase for a prescribed set
of in i t ia l conditions is given by Figure 4-13 As in the case of the STSS configuration
there is strong coupling between the tether and the satellites rigid body dynamics
causing the latter to follow the attitude of the tether However because of the smaller
inertias of the payloads there are noticeable high frequency modulations arising from
the tether flexibility Response of the tether in the elastic degrees of freedom is
presented in Figure 4-13(b) Note the longitudinal vibrations decay quite rapidly
(lt 05 orbits) due to the structural damping however it has vir tual ly no effect on
the transverse oscillations
The unique mission requirement of B I C E P S is the proposed use of its ini t ia l
angular momentum in the cartwheeling mode to aid in the deployment of the tethered
system The maneuver is considered next The response of the system during this
maneuver wi th an ini t ia l cartwheeling rate of 5 deg s is illustrated in Figure 4-14(a) The
in i t ia l tether length is taken to be 10 m and is deployed to 1 km There is an in i t ia l
increase in the pitch motion of the tether however due to the conservation of angular
momentum the cartwheeling rate decreases proportionally to the square of the change
in tether deployment rate The result is a rapid drop in the cartwheeling rate unti l
the system stops rotating entirely and simply oscillates about its new equilibrium
point Consequently through coupling the end-bodies follow the same trend The
roll response also subsides once the cartwheeling motion ceases However
83
BICEPS configuration 3-Body a1(0)=o2(0)=at(0)=2deg P1(0)=P2(0)=pt(0)=1deg 8X(0)=001 m 6y(0)=52(0)=01 m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping l=1 km
Satellite 1 Pitch Angle Satellite 1 Roll Angle - i 1 1 1 1 i - i q r 1 1 -gt
I- I i t i i i i I H 0 1 2 0 1 2
Time (Orbits) Time (Orbits) Tether Pitch Angle Tether Roll Angle
F mdash 1 mdash 1 mdash mdash 1 mdash i mdash 1 mdash 1 mdash mdash 1 mdash = i F mdash 1 mdash mdash mdash 1 mdash i mdash 1 mdash 1 mdash 1 mdash 1 mdash q
(a) Time (Orbits) Time (Orbits)
Figure 4-13 Stationkeeping dynamics of the three-body BICEPS configuration with offset along the local vertical (a) attitude response
84
BICEPS configuration 3-Body a1(0)=o2(0)=at(0)=2deg Pl(0)=p2(0)=(3t(0)=1deg 8X(0)=001 m 8y(0)=82(0)=01 m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping lt=1km
Tether Longitudinal Vibration 1E-2[
laquo = 5 E - 3 |
000 002
Time (Orbits) Tether In-Plane Transverse Vibration
Time (Orbits) Tether Out-of-Plane Transverse Vibration
Time (Orbits)
Figure 4-13 Stationkeeping dynamics of the three-body BICEPS configuration with offset along the local vertical (b) vibration response
85
BICEPS configuration 3-Body a1(0)=o2(0)=at(0)=2deg a1(0)=a2(0)=at(0)=57s Pl(0)=P2(0)=(3t(0)=1deg 8X(0)=115 x 103m 8y(0)=82(0)=0 dx(0)=078m dy(0)=d2(0)=0 Cartwheeling Deployment
lt=10m to 1km in 1 Orbit
Tether Length Profile
E
2000
cu T3
2000
co CD
2000
C D CD
T J
Satellite 1 Pitch Angle
1 2 Time (Orbits)
Satellite 1 Roll Angle
cn CD
33 cdl
Time (Orbits) Tether Pitch Angle
Time (Orbits) Tether Roll Angle
Time (Orbits) Satellite 2 Pitch Angle
1 2 Time (Orbits)
Satellite 2 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-14 Cartwheeling dynamics with deployment for the three-body BICEPS with offset along the local vertical (a) attitude response
86
BICEPS configuration 3-Body a1(0)=a2(0)=at(0)=2deg a (0)=a2(0)=a(0)=57s P1(0)=P2(0)=f3(0)=1deg 8X(0)=115x103m 8y(0)=8z(0)=0 dx(0)=078m dy(0)=d2(0)=0 Cartwheeling Deployment
l=1 Om to 1 km in 1 Orbit
5E-3h
co
0E0
025
000 E to
-025
002
-002
(b)
Figure 4-14
Tether Longitudinal Vibration
Time (Orbits) Tether In-Plane Transverse Vibration
Time (Orbits) Tether Out-of-Plane Transverse Vibration
Time (Orbits)
Cartwheeling dynamics with deployment for the three-body BICEPS with offset along the local vertical (b) vibration response
87
as the deployment maneuver is completed the satellites continue to exhibit a small
amplitude low frequency roll response with small period perturbations induced by
the tether elastic oscillations superposed on it (Figure 4-14b)
48 OEDIPUS Spinning Configuration
The mission entitled Observation of Electrified Distributions of Ionospheric
Plasmas - A Unique Strategy ( O E D I P U S ) also consists of a 1 k m tethered system with
two similar end-masses It was first flown in a suborbital flight in 1989 ( O E D I P U S A )
followed by a recent study in November 1995 ( O E D I P U S - C ) in which the University
of Br i t i sh Columbia was one of the participants Dynamically O E D I P U S represents a
unique system never encountered before Launched by a Black Brant rocket developed
at Br is to l Aerospace L t d the system ie the end-bodies together wi th the tether
spins about the longitudinal axis of the tether to achieve stabilized alignment wi th
Earth s magnetic field
The response of the system undergoing a spin rate of j = l deg s is presented
in Figure 4-15 using the same parameters as those outlined for the B I C E P S conshy
figuration in Section 47 Again there is strong coupling between the three bodies
(two satellites and tether) due to the nonzero offset The spin motion introduces an
additional frequency component in the tethers elastic response which is transferred
to the l ibrational motion of the satellites A s the spin rate increases to 1 0 deg s (Figure
4-16) the amplitude and frequency of the perturbations also increase however the
general character of the response remains essentially the same
88
OEDIPUS configuration 3-Body a1(0)=a2(0)=at(0)=2deg i(0)^[2(0H(0)=Ys p1(0)=p2(0)=p(0)=r 8X(0)=12 x 103m 8y(0)=82(0)=01m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping l=1 km
Satellite 1 Pitch Angle Satellite 1 Roll Angle mdash i 1 1 1 a r 1 1 1 r
(a) Time (Orbits) Time (Orbits)
Figure 4-15 Spin dynamics (7 = ldegs) of the three-body OEDIPUS configuration with offset along the local vertical (a) attitude response
89
OEDIPUS configuration 3-Body a1(0)=a2(0)=cct(0)=2deg Y1(0)=Y2(0)=Yt(0)=l7s (31(0)=p2(0)=pt(0)=1deg 8X(0)=12 x 103m 8y(0)=5z(0)=01 m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping lt=1 km
1E-2
to
OEO
E 00
to
Tether Longitudinal Vibration
005
Time (Orbits) Tether In-Plane Transverse Vibration
Time (Orbits) Tether Out-of-Plane Transverse Vibration
Time (Orbits)
Figure 4-15 Spin dynamics (7 = l deg s ) of the three-body O E D I P U S configuration with offset along the local vertical (b) vibration response
90
OEDIPUS configuration 3-Body a1(0)=a2(0)=a(0)=2deg Yi(0)=Y2(0)=rt(0)=107s 81(0)=p2(0)=pt(0)=1deg 8X(0)=12 x 103m 8y(0)=82(0)=01m dx(0)=078m dy(0)=d2(0)=0 Stationkeeping lt=1km
Satellite 1 Pitch Angle Satellite 1 Roll Angle
(a) Time (Orbits) Time (Orbits)
Figure 4-16 Spin dynamics (7= 10degs) of the three-body OEDIPUS configura tion with offset along the local vertical (a) attitude response
91
OEDIPUS configuration 3-Body a1(0)=a2(0)=ot(0)=2deg Yi(0)H(0)4(0)laquo107s P1(0)=P2(0)=pt(0)=1deg 8X(0)=12 x 103m 8y(0)=82(0)=01m dx(0)=078m dy(0)=dz(0)=0 Stationkeeping lt=1 km
Tether Longitudinal Vibration
Figure 4-16 Spin dynamics (7 = 10degs) of the three-body OEDIPUS configurashytion with offset along the local vertical (b) vibration response
92
5 ATTITUDE AND VIBRATION CONTROL
51 Att itude Control
511 Preliminary remarks
The instability in the pitch and roll motions during the retrieval of the tether
the large amplitude librations during its deployment and the marginal stability during
the stationkeeping phase suggest that some form of active control is necessary to
satisfy the mission requirements This section focuses on the design of an attitude
controller with the objective to regulate the librational dynamics of the system
As discussed in Chapter 1 a number of methods are available to accomplish
this objective This includes a wide variety of linear and nonlinear control strategies
which can be applied in conjuction with tension thrusters offset momentum-wheels
etc and their hybrid combinations One may consider a finite number of Linear T ime
Invariant (LTI) controllers scheduled discretely over different system configurations
ie gain scheduling[43] This would be relevant during deployment and retrieval of
the tether where the configuration is changing with time A n alternative may be
to use a Linear T ime Varying ( L T V ) model in conjuction with an adaptive control
scheme where on-line parametric identification may be used to advantage[56] The
options are vir tually limitless
O f course the choice of control algorithms is governed by several important
factors effective for time-varying configurations computationally efficient for realshy
time implementation and simple in character Here the thrustermomentum-wheel
system in conjuction with the Feedback Linearization Technique ( F L T ) is chosen as
it accounts for the complete nonlinear dynamics of the system and promises to have
good overall performance over a wide range of tether lengths It is well suited for
93
highly time-varying systems whose dynamics can be modelled accurately as in the
present case
The proposed control method utilizes the thrusters located on each rigid satelshy
lite excluding the first one (platform) to regulate the pitch and rol l motion of the
tether to which it is attached ie the thrusters located on satellite 2 (link 3) regshy
ulates the attitude motion of tether 1 (link 2) O n the other hand the l ibrational
motion of the rigid bodies (platform and subsatellite) is controlled using a set of three
momentum wheels placed mutually perpendicular to each other
The F L T method is based on the transformation of the nonlinear time-varying
governing equations into a L T I system using a nonlinear time-varying feedback Deshy
tailed mathematical background to the method and associated design procedure are
discussed by several authors[435758] The resulting L T I system can be regulated
using any of the numerous linear control algorithms available in the literature In the
present study a simple P D controller is adopted arid is found to have good perforshy
mance
The choice of an F L T control scheme satisfies one of the criteria mentioned
earlier namely valid over a wide range of tether lengths The question of compushy
tational efficiency of the method wi l l have to be addressed In order to implement
this controller in real-time the computation of the systems inverse dynamics must
be executed quickly Hence a simpler model that performs well is desirable Here
the model based on the rigid system with nonlinear equations of motion is chosen
Of course its validity is assessed using the original nonlinear system that accounts
for the tether flexibility
512 Controller design using Feedback Linearization Technique
The control model used is based on the rigid system with the governing equa-
94
tions of motion given by
Mrqr + fr = Quu (51)
where the left hand side represents inertia and other forces while the right hand side
is the generalized external force due to thrusters and momentum-wheels and is given
by Eq(2109)
Lett ing
~fr = fr~ QuU (52)
and substituting in Eq(296)
ir = s(frMrdc) (53)
Expanding Eq(53) it can be shown that
qr = S(fr Mrdc) - S(QU Mr6)u (5-4)
mdash = Fr - Qru
where S(QU Mr0) is the column matrix given by
S(QuMr6) = [s(Qu(l)Mr0) 5 (Q u ( 2 ) M r | 0 ) S(QU nu) M r | 0 ) ] (55)
and Qu-i) is the i ^ column of Qu Extract ing only the controlled equations from
Eq(54) ie the attitude equations one has
Ire = Frc ~~ Qrcu
(56)
mdash vrci
where vrc is the new control input required to regulate the decoupled linear system
A t this point a simple P D controller can be applied ie
Vrc = qrcd +KvQrcd-Qrc) + Kp(Qrcd~ Qrc)- (57)
Eq(57) represents the secondary controller with Eq(54) as the primary controller
Kp and Kv are the proportional and derivative gain matrices with qrcdi Qrcd
an(^ qrcd
95
as the desired acceleration velocity and position vector for the attitude angles of each
controlled body respectively Solving for u from Eq(56) leads to
u Qrc Frc ~ vrcj bull (58)
5 13 Simulation results
The F L T controller is implemented on the three-body STSS tethered system
The choice of proportional and derivative matrix gains is based on a desired settling
time of ts mdash 05 orbit and a damping factor of = 07 for both the pitch and roll
actuators in each body Given O and ts it can be shown that
-In ) 0 5 V T ^ r (59)
and hence kn bull mdash10
V (510)
where kp and kVj are the diagonal elements of the gain matrices Kp and Kv in
Eq(57) respectively[59]
Note as each body-fixed frame is referred directly to the inertial frame FQ
the nominal pitch equilibrium angle is zero only when measured from the L V L H frame
(OJJ mdash 0) However it is 6 when referred to FQ Hence the desired position vector
qrC(l is set equal to 0(t) ie the time-varying orbital angle for the pitch motion
and zero for the roll and yaw rotations such that
qrcd 37Vxl (511)
96
The desired velocity and acceleration are then given as
e U 3 N x l (512)
E K 3 7 V x l (513)
When the system is in a circular orbit qrc^ = 0
Figure 5-1 presents the controlled response of the STSS system defined in
Chapter 4 in the stationkeeping phase with offset d2 = 111T m and If = 20
km A s mentioned earlier the F L T controller is based on the nonlinear rigid model
Note the pitch and roll angles of the rigid bodies as well as of the tether are now
attenuated in less than 1 orbit (Figure 5-la) Consequently the longitudinal vibration
response of the tether is free of coupling arising from the librational modes of the
tether This leaves only the vibration modes which are eventually damped through
structural damping The pitch and roll dynamics of the platform require relatively
higher control moments Ma and Mpi respectively (Figure 5-lb) since they have to
overcome the extra moments generated by the offset of the tether attachment point
Furthermore the platform demands an additional moment to maintain its orientation
at zero pitch and roll angle since it is not the nominal equilibrium position The nonshy
zero static equilibrium of the platform arises due to its products of inertia On the
other hand the tether pitch and roll dynamics require only a small ini t ia l control
force of about plusmn 1 N which eventually diminishes to zero once the system is
97
and
o 0
Qrcj mdash
o o
0 0
Oiit) 0 0
1
N
N
1deg I o
STSS configuration 3-Body a1(0)=a2(0)=at(0)=2lt
P1(0)=p2(0)=pt(0)=1 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=dz(0)=1m Controlled Stationkeeping lt=20km
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
05 10 15 Time (Orbits)
Sat 2 Pitch and Roll Angles
20
CO CD
T J
cdT
cT
- mdash mdash 1 1 1 1 1 I - 1
Pitch -
A Roll -
A -
1 bull bull
-
1 Time (Orbits)
Tether Y Tranverse Vibration
1 2 Time (Orbits)
Tether Z Tranverse Vibration
gt to N
to
Time (Orbits) Time (Orbits) Longitudinal Vibration
145h
(a) Time (Orbits)
Figure 5-1 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local horizontal local vertical and orbit normal (a) attitude and vibration response
98
1
STSS configuration 3-Body a1(0)=ct2(0)=cct(0)=20
P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=d2(0)=1m Controlled Stationkeeping lt=20km
Sat 1 Pitch Control Moment Sat 1 Roll Control Moment
Figure 5-1 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local horizontal local vertical and orbit normal (b) control actuator time histories
99
stabilized Similarly for the end-body a very small control torque is required to
attenuate the pitch and roll response
If the complete nonlinear flexible dynamics model is used in the feedback loop
the response performance was found to be virtually identical with minor differences
as shown in Figure 5-2 For example now the pitch and roll response of the platform
is exactly damped to zero with no steady-state error as against the minor almost
imperceptible deviation for the case of the controller basedon the rigid model (Figure
5-la) In addition the rigid controller introduces additional damping in the transverse
mode of vibration where none is present when the full flexible controller is used
This is due to a high frequency component in the at motion that slowly decays the
transverse motion through coupling
As expected the full nonlinear flexible controller which now accounts the
elastic degrees of freedom in the model introduces larger fluctuations in the control
requirement for each actuator except at the subsatellite which is not coupled to the
tether since f c m 3 = 0 The high frequency variations in the pitch and roll control
moments at the platform are due to longitudinal oscillations of the tether and the
associated changes in the tension Despite neglecting the flexible terms the overall
controlled performance of the system remains quite good Hence the feedback of the
flexible motion is not considered in subsequent analysis
When the tether offset at the platform is restricted to only 1 m along the x
direction a similar response is obtained for the system However in this case the
steady-state error in the platforms rigid body motion is much smaller (Figure 5-3a)
In addition from Figure 5-3(b) the platforms control requirement is significantly
reduced
Figure 5-4 presents the controlled response of the STSS deploying a tether
100
STSS configuration 3-Body a1(0)=a2(0)=oct(0)=2deg p1(0)=P2(0)=p(0)=1deg 8x(0)=14m5y(0)=82(0)=1m dx(0)=dy(0)=dz(0)=1m Controlled Stationkeeping l=20km
Tether Pitch and Roll Angles
c n
T J
cdT
Sat 1 Pitch and Roll Angles 1 r
00 05 10 15 Time (Orbits)
Sat 2 Pitch and Roll Angles
c n CD
cdT a
o T J
CQ
- - I 1 1 1 I 1 1 1 1
P i t c h
R o l l
bull
1 -
0 1 2 Time (Orbits)
Tether Y Tranverse Vibration
0 1 2 Time (Orbits)
Tether Z Tranverse Vibration
E gt
10 N
CO
Time (Orbits) Time (Orbits) Longitudinal Vibration
145h
(a) Time (Orbits)
Figure 5-2 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear flexible FLT controller with offset along the local horizontal local vertical and orbit normal (a) attitude and vibration response
101
STSS configuration 3-Body a1(0)=a2(0)=cxt(0)=2deg 31(0)=p2(0)=p(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=dy(0)=d2(0)=1m Controlled Stationkeeping lt=20km
r i i i t i mdash i mdash i mdash i mdash i mdash i mdash i mdash i mdash mdash J 0 1 2 0 1 2
Time (Orbits) Time (Orbits) Tether Pitch Control Thrust Tether Roll Control Thruster
(b) Time (Orbits) Time (Orbits)
Figure 5-2 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear flexible FLT controller with offset along the local horizontal local vertical and orbit normal (b) control actuator time histories
102
STSS configuration 3-Body a1(0)=o2(0)=a(0)=2deg P1(0)=p2(0)=(3t(0)=1deg 8x(0)=14m5y(0)=6z(0)=1m dx(0)=1mdy(0)=dz(0)=0 Controlled Stationkeeping lt=20km
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
05 10 15 Time (Orbits)
Sat 2 Pitch and Roll Angles
CD
eg CQ
0 1 2 Time (Orbits)
Tether Y Tranverse Vibration
1 2 Time (Orbits)
Tether Z Tranverse Vibration
gt CO
Time (Orbits) Time (Orbits) Longitudinal Vibration
145h
(a) Time (Orbits)
Figure 5-3 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local vertical (a) attitude and vibration response
103
STSS configuration 3-Body a1(0)=o2(0)=al(0)=2o
P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m8y(0)=8z(0)=1m dx(0)=1mdy(0)=dz(0)=0 Controlled Stationkeeping lt=20km
Sat 1 Pitch Control Moment 11 1 1 r
Sat 1 Roll Control Moment
0 1 2 Time (Orbits)
Tether Pitch Control Thrust
0 1 2 Time (Orbits)
Sat 2 Pitch Control Moment
0 1 2 Time (Orbits)
Tether Roll Control Thruster
-07 o 1
Time (Orbits Sat 2 Roll Control Moment
1E-5
0E0F
Time (Orbits)
Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT controller with offset along the local vertical (b) control actuator time histories
104
STSS configuration 3-Body a1(0)=a2(0)=a(0)=2deg Pl(0)=p2(0)=pt(0)=1deg 6X(0)=1304 x 103m 8y(0)=82(0)=0 dx(0)=1mdy(0)=d2(0)=0 Controlled Deployment
l=02km to 20km in 35 Orbits
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
o 1 2 Time (Orbits)
Tether Y Tranverse Vibration
0 1 2 Time (Orbits)
Sat 2 Pitch and Roll Angles
CD CD
CM CO
Time (Orbits) Tether Z Tranverse Vibration
to
150
100
1 2 3 0 1 2
Time (Orbits) Time (Orbits) Longitudinal Vibration
to
(a)
50
2 3 Time (Orbits)
Figure 5 - 4 Deployment dynamics of the three-body STSS using the nonlinear rigid FLT controller with offset along the local vertical (a) attitude and vibration response
105
STSS configuration 3-Body a1(0)=oc2(0)=al(0)=20
P1(0)=p2(0)=pt(0)=1deg 8X(0)=1304 x 103m 5y(0)=52(0)=0 dx(0)=1mdy(0)=dz(0)=0 Controlled Deployment
l=02km to 20km in 35 Orbits
Tether Length Profile i 1
2 0 h
E 2L
Sat 1 Pitch Control Moment t mdash r mdash i mdash i mdash | mdash i mdash i mdash i mdash i mdash | mdash r mdash i mdash i mdash i mdash | mdash r mdash
0 5
1 2 3
Time (Orbits) Sat 1 Roll Control Moment
i i t | t mdash t mdash i mdash [ i i mdash r -
1 2
Time (Orbits) Tether Pitch Control Thrust
1 2 3 Time (Orbits)
Tether Roll Control Thruster
2 E - 5
1 2 3
Time (Orbits) Sat 2 Pitch Control Moment
1 2
Time (Orbits) Sat 2 Roll Control Moment
1 E - 5
E
O E O
Time (Orbits)
Deployment dynamics of the three-body STSS using the nonlinear rigid FLT controller with offset along the local vertical (b) control actuator time histories
106
from 02 km to 20 km in 35 orbits with a i m offset along the tether length A s before
the systems attitude motion is well regulated by the controller However the control
cost increases significantly for the pitch motion of the platform and tether due to the
Coriolis force induced by the deployment maneuver (Figure 5-4b) Initially there is a
sinusoidal increase in the pitch control requirement for both the tether and platform
as the tether accelerates to its constant velocity VQ Then the control requirement
remains constant for the platform at about 60 N m as opposed to the tether where the
thrust demand increases linearly Finally when the tether decelerates the actuators
control input reduces sinusoidally back to zero The rest of the control inputs remain
essentially the same as those in the stationkeeping case In fact the tether pitch
control moment is significantly less since the tether is short during the ini t ia l stages
of control However the inplane thruster requirement Tat acts as a disturbance and
causes 6y to nearly double from the uncontrolled value (Figure 4-8a) O n the other
hand the tether has no out-of-plane deflection
Final ly the effectiveness of the F L T controller during the crit ical maneuver of
retrieval from 20 km to 02 k m in 35 orbits is assessed in Figure 5-5 A s in the earlier
cases the system response in pitch and roll is acceptable however the controller is
unable to suppress the high frequency elastic oscillations induced in the tether by the
retrieval O f course this is expected as there is no active control of the elastic degrees
of freedom However the control of the tether vibrations is discussed in detail in
Section 52 The pitch control requirement follows a similar trend in magnitude as
in the case of deployment with only minor differences due to the addition of offset in
the out-of-plane direction ( d 2 = 1 0 1 ^ m) This is also responsible for the higher
roll moment requirement for the platform control (Figure 5-5b)
107
STSS configuration 3-Body a1(0)=X2(0)=a(0)=2deg P1(0)=p2(0)=pt(0)=1deg 5x(0)=14m8y(0)=52(0)=0 dx(0)=1mdy(0)=dz(0)=0 Controlled Retrieval
l=20km to 02km in 35 Orbits
Tether Pitch and Roll Angles
Sat 1 Pitch and Roll Angles
0 1 2 Time (Orbits)
Sat 2 Pitch and Roll Angles
CO CD
cdT
CO CD
T J CM
CO
1 2 Time (Orbits)
Tether Y Tranverse Vibration
Time (Orbits) Tether Z Tranverse Vibration
150
100
1 2 3 0 1 2 Time (Orbits) Time (Orbits)
Longitudinal Vibration
CO
(a)
x 50
2 3 Time (Orbits)
Figure 5 -5 Retrieval dynamics of the three-body STSS using the non-linear rigid FLT controller with offset along the local vertical and orbit normal (a) attitude and vibration response
108
STSS configuration 3-Body a1(0)=ct2(0)=at(0)=2deg p1(0)=p2(0)=(3(0)=1deg 5x(0)=14m5y(0)=82(0)=0 dx(0)=1mdy(0)=dz(0)=0 Controlled Retrieval
l=20km to 02km in 35 Orbits
Tether Length Profile
E
1 2 Time (Orbits)
Tether Pitch Control Thrust
2E-5
1 2 Time (Orbits)
Sat 2 Pitch Control Moment
OEOft
1 2 3 Time (Orbits)
Sat 1 Roll Control Moment
1 2 3 Time (Orbits)
Tether Roll Control Thruster
1 2 Time (Orbits)
Sat 2 Roll Control Moment
1E-5 E z
2
0E0
(b) Time (Orbits) Time (Orbits)
Figure 5-5 Retrieval dynamics of the three-body STSS using the non-linear rigid FLT controller with offset along the local vertical and orbit normal (b) control actuator time histories
109
52 Control of Tethers Elastic Vibrations
521 Prel iminary remarks
The requirement of a precisely controlled micro-gravity environment as well
as the accurate release of satellites into their final orbit demands additional control
of the tethers vibratory motion in addition to its attitude regulation To that end
an active vibration suppression strategy is designed and implemented in this section
The strategy adopted is based on offset control ie time dependent variation
of the tethers attachment point at the platform (satellite 1) A l l the three degrees
of freedom of the offset motion are used to control both the longitudinal as well as
the inplane and out-of-plane transverse modes of vibration In practice the offset
controller can be implemented through the motion of a dedicated manipulator or a
robotic arm supporting the tether which in turn is supported by the platform The
focus here is on the control of elastic deformations during stationkeeping (ie fully
deployed tether fixed length) as it represents the phase when the mission objectives
are carried out
This section begins with the linearization of the systems equations of motion
for the reduced three body stationkeeping case This is followed by the design of
the optimal control algorithm Linear Quadratic Gaussian-Loop Transfer Recovery
( L Q G L T R ) based on the reduced model and its implementation on the full nonlinear
model Final ly the systems response in the presence of offset control is presented
which tends to substantiate its effectiveness
522 System linearization and state-space realization
The design of the offset controller begins with the linearization of the equations
of motion about the systems equilibrium position Linearization of extremely lengthy
(even in matrix form) highly nonlinear nonautonomous and coupled equations of
110
motion presents a challenging problem This is further complicated by the fact the
pitch angle ot is not referred to the L V L H frame Hence the equilibrium pitch angle
is not a constant but is equal to the instantaneous orbital angle 9(t) Two methods
are available to resolve this problem
One may use the non-stationary equations of motion in their present form and
derive a controller based on the Linear T ime Varying ( L T V ) system Alternatively
one can design a Linear T ime Invariant (LTI) controller based on a new set of reduced
governing equations representing the motion of the three body tethered system A l shy
though several studies pertaining to L T V systems have been reported[5660] the latter
approach is chosen because of its simplicity Furthermore the L T I controller design
is carried out completely off-line and thus the procedure is computationally more
efficient
The reduced model is derived using the Lagrangian approach with the genershy
alized coordinates given by
where lty and are the pitch and roll angles of link i relative to the L V L H frame
respectively 5X 8y and 5Z are the generalized coordinates associated with the flexible
tether deformations in the longitudinal inplane and out-of-plane transverse direcshy
tions respectively Only the first mode of vibration is considered in the analysis
The nonlinear nonautonomous and coupled equations of motion for the tethshy
ered system can now be written as
Qred = [ltXlPlltX2P2fixSy5za302gt] (514)
M redQred + fr mdash 0gt (515)
where Mred and frec[ are the reduced mass matrix and forcing term of the system
111
respectively They are functions of qred and qred in addition to the time varying offset
position d2 and its time derivatives d2 and d2 A detailed derivation of Eq(515)
is given in Appendix II A n additional consequence of referring the pitch motion to
a local frame is that the new reduced equations are now independent of the orbital
angle 9 and under the further assumption of the system negotiating a circular orbit
91 remains constant
The nonlinear reduced equations can now be linearized about their static equishy
l ibr ium position For all the generalized coordinates the equilibrium position is zero
wi th the exception of 5X which has a non-zero equilibrium 5XQ The offset position
d2) is linearized about its ini t ial position d2^ Linearizing Eq(515) and recasting
into matr ix form gives
Msqred + CsQred + KsQred + Md^2 + Cdd2 + Kdd2 + fs = 0 (516)
where Ms Cs Ks Md Cd Kd and s are constant matrices Mul t ip ly ing throughout
by Ms
1 gives
qr = -Ms 1Csqr - Ms
lKsqr
- Ms~lCdd22 - Ms-Kdd2 - Ms~Mdd2 - Mg~1 fs r-1 (517)
Defining ud = d2 and v = qTd^T Eq(517) can be rewritten as
Pu t t ing
v =
+
-M~lCs -M~Cd -1
0
-Mg~lMd
Id
0 v + -M~lKs -M~lKd
0 0
-M-lfs
0
MCv + MKv + Mlud +Fs
- 4 ) - ( gt the L T I equations of motion can be recast into the state-space form as
(518)
(519)
x =Ax + Bud + Fd (520)
112
where
A = MC MK Id12 o
24x24
and
Let
B = ^ U 5 R 2 4 X 3 -
Fd = I FJ ) e f t 2 4 1
mdash _ i _ -5
(521)
(522)
(523)
(524)
where x is the perturbation vector from the constant equilibrium state vector xeq
Substituting Eq(524) into Eq(520) gives the linear perturbation state equation
modelling the reduced tethered system as
bullJ x mdash x mdash Ax + Bud + (Fd + Axeq) (525)
It can be shown that for ideal control[60] Fd + Axeq = 0 thus giving the familiar
state-space equation
S=Ax + Bud (526)
The selection of the output vector completes the state space realization of
the system The output vector consists of the longitudinal deformation from the
equil ibrium position SXQ of the tether at xt = h the slope of the tether due to the
transverse deformation at xt = 0 and the offset position d2 from its in i t ia l position
d2Q Thus the output vector y is given by
K(o)sx
V
dx ~ dXQ dy ~ dyQ
d z - d z0 J
Jy
dx da XQ vo
d z - dZQ ) dy dyQ
(527)
113
where lt ( 0 ) = ddxi[$vxi)]Xi=o and lt ( 0 ) = ddxi[4gtw(xi)]Xi=0
523 Linear Quadratic Gaussian control with Loop Transfer Recovery
W i t h the linear state space model defined (Eqs526527) the design of the
controller can commence The algorithm chosen is the Linear Quadratic Gaussian
( L Q G ) estimator based optimal controller[6061] The L Q G is a widely used optimal
controller with its theoretical background well developped by many authors over the
last 25 years It involves of the design of the Kalman-Bucy Fi l ter ( K B F ) which
provides an estimate of the states x and a Linear Quadratic Regulator ( L Q R ) which
is separately designed assuming all the states x are known Both the L Q R and K B F
designs independently have good robustness properties ie retain good performance
when disturbances due to model uncertainty are included However the combined
L Q R and K B F designs ie the L Q G design may have poor stability margins in the
presence of model uncertainties This l imitat ion has led to the development of an L Q G
design procedure that improves the performance of the compensator by recovering
the full state feedback robustness properties at the plant input or output (Figure 5-6)
This procedure is known as the Linear Quadratic Guassian-Loop Transfer Recovery
( L Q G L T R ) control algorithm A detailed development of its theory is given in
Ref[61] and hence is not repeated here for conciseness
The design of the L Q G L T R controller involves the repeated solution of comshy
plicated matrix equations too tedious to be executed by hand Fortunenately the enshy
tire algorithm is available in the Robust Control Toolbox[62] of the popular software
package M A T L A B The input matrices required for the function are the following
(i) state space matrices ABC and D (D = 0 for this system)
(ii) state and measurement noise covariance matrices E and 0 respectively
(iii) state and input weighting matrix Q and R respectively
114
u
MODEL UNCERTAINTY
SYSTEM PLANT
y
u LQR CONTROLLER
X KBF ESTIMATOR
u
y
LQG COMPENSATOR
PLANT
X = f v ( X t )
COMPENSATOR
x = A k x - B k y
ud
= C k f
y
Figure 5-6 Block diagram for the LQGLTR estimator based compensator
115
A s mentionned earlier the main objective of this offset controller is to regu-mdash
late the tether vibration described by the 5 equations However from the response
of the uncontrolled system presented in Chapter 4 it is clear that there is a large
difference between the magnitude of the librational and vibrational frequencies This
separation of frequencies allow for the separate design of the vibration and attitude mdash mdash
controllers Thus only the flexible subsystem composed of the S and d equations is
required in the offset controller design Similarly there is also a wide separation of
frequencies between the longitudinal and transverse modes of vibrations permitt ing mdash
the decoupling of the flexible subsystem into a set of longitudinal (Sx and dx) and
transverse (5y Sz dy and dz) subsystems The appended offset system d must also
be included since it acts as the control actuator However it is important to note
that the inplane and out-of-plane transverse modes can not be decoupled because
their oscillation frequencies are of the same order The two flexible subsystems are
summarized below
(i) Longitudinal Vibra t ion Subsystem
This is defined by u mdash Auxu + Buudu
(528)
where xu = SxdX25xdX2T u d u = dX2 and from Eq(527)
0 0 $u(lt) 0 0 0 0 1
0 0 1 0 0 0 0 1
(529)
Au and Bu are the rows and columns of A and B respectively and correspond to the
components of xu
The L Q R state weighting matrix Qu is taken as
Qu =
bull1 0 0 o -0 1 0 0 0 0 1 0
0 0 0 10
(530)
116
and the input weighting matrix Ru = 1^ The state noise covariance matr ix is
selected as 1 0 0 0 0 1 0 0 0 0 4 0
LO 0 0 1
while the measurement noise covariance matrix is taken as
(531)
i o 0 15
(532)
Given the above mentionned matrices the design of the L Q G L T R compenshy
sator is computed using M A T L A B with the following 2 commands
(i) kfu = l q r c ( ^ ( i C bdquo d i a g m x ( E u e u ) )
(ii) [AkugtBkugtCkugtDku =^ryAuBuCuDukfuQuRur)
where r is a scaler
The first command is used to compute the Ka lman filter gain matrix kfu Once
kfu is known the second command is invoked returning the state space representation
of the L Q G L T R dynamic compensator to
(533) xu mdash Akuxu Bkuyu
where xu is the state estimate vector oixu The function l t ry performs loop transfer
recovery at the system output ie the return ratio at the output approaches that
of the K B F loop given by Cu(sld - Au)Kfu[61 This is achieved by choosing a
sufficently large scaler value r in l t ry such that the singular values of the return
ratio approach those of the target design For the longitudinal controller design
r mdash 5 x 10 5 Figure 5-7(a) compares the singular values of the recovered compensator
design and the non-recovered L Q G compensator design with respect to the target
Unfortunetaly perfect recovery is not possible especially at higher frequencies
117
100
co
gt CO
-50
-100 10
Longitudinal Compensator Singular Values i I I I N I I mdash i mdash i i i i n i i mdash i mdash i i i i inimdash imdashi i i i i i i i mdash i mdash i i M i I I
Target r = 5x10 5
r = 0 (LQG)
i i i i I I i ii i i i 1 1 1 1 i i i i i 1 1 1 1 ii i I I i i i i 11111 -3 10 -2
(a)
101 10deg Frequency (rads)
101 10
Transverse Compensator Singular Values
100
50
X3
gt 0
CO
-50
bull100 10
-imdashi i 111II i 111 ni 1mdashi I I I I I I I 1mdashi I I I N I I 1mdashi i M 111
Target r = 50 r = 0 (LQG)
m i i i i i i n i i i i I I I I I I I 1mdashi I I I I I I I 1mdashi i 111 n -3 10 -2
(b)
101 10deg Frequency (rads)
101 10
Figure 5-7 Singular values for the LQG and LQGLTR compensator compared to target return ratio (a) longitudinal design (b) transverse design
118
because the system is non-minimal ie it has transmission zeros with positive real
parts[63]
(ii) Transverse Vibra t ion Subsystem
Here Xy mdash AyXy + ByU(lv
Dv = Cvxv
with xv = 6y 8Z dV2 dZ2 Sy 6Z dV2 dZ2T] udv = dy2dZ2T and
(534)
Cy
0 0 0 0 (0) 0 0 0 0 0 0 0 0 0 0 0 0
0 0 0 0
0 0 0 0 0 0 0 0 0 0 0 0
0 0 0
h 0 0 0
0 0 0 lt ( 0 ) o 0
0 1 0 0 0 1
0 0 0
2 r
h 0 0
0 0 1 0 0 1
(535)
Av and Bv are the rows and columns of the A and B corresponding to the components
O f Xy
The state L Q R weighting matrix Qv is given as
Qv
10 0 0 0 0 0 0 0 0 10 0 0 0 0 0 0 0 0 10 4 0 0 0 0 0 0 0 0 10 4 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1
(536)
mdash
Ry =w The state noise covariance matrix is taken
1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 1 0 0 0 0 0 0 0 0 9 x 10 4 0
0 0 0 0 0 0 0 9 x 10 4
(537)
119
while the measurement noise covariance matrix is
1 0 0 0 0 1 0 0 0 0 5 x 10 4 0
LO 0 0 5 x 10 4
(538)
As in the development of the longitudinal controller the transverse offset conshy
troller can be designed using the same two commands (lqrc and l try) wi th the input
matrices corresponding to the transverse subsystem with r = 50 The result is the
transverse compensator system given by
(539)
where xv is the state estimates of xv The singular values of the target transfer
function is compared with those achieved by the L Q G and L Q G L T R in Figure 5-
7(b) It is apparent that the recovery is not as good as that for the longitudinal case
again due to the non-minimal system Moreover the low value of r indicates that
only a lit t le recovery in the transverse subsystem is possible This suggests that the
robustness of the L Q G design is almost the maximum that can be achieved under
these conditions
Denoting f y = f j f j r xf = x^x^T and ud = udu ^ T the longishy
tudinal compensator can be combined to give
xf =
$d =
Aku
0
0 Akv
Cku o
Cu 0
Bku
0 B Xf
Xf = Ckxf
0
kv
Hf = Akxf - Bkyf
(540)
Vu Vv 0 Cv
Xf = CfXf
Defining a permutation matrix Pf such that Xj = PfX where X is the state vector
of the original nonlinear system given in Eq(32) the compensator and full nonlinear
120
system equations can be combined as
X
(541)
Ckxf
A block diagram representation of Eq(541) is shown in Figure 5-6
524 Simulation results
The dynamical response for the stationkeeping STSS with a tether length of 20
km and ini t ia l offset position of 1 m in each direction is presented in Figure 5-8 when
both the F L T attitude controller and the L Q G L T R offset controller are activated As
expected the offset controller is successful in quickly damping the elastic vibrations
in the longitudinal and transverse direction (Figure 5-8b) However from Figure 5-
8(a) it is clear that the presence of offset control requires a larger control moment
to regulate the attitude of the platform This is due to the additional torque created
by the larger moment arm around 2 m and 125 m in the inplane and out-of-plane
directions respectively introduced by the offset control In addition the control
moment for the platform is modulated by the tethers transverse vibration through
offset coupling However this does not significantly affect the l ibrational motion of
the tether whose thruster force remains relatively unchanged from the uncontrolled
vibration case
Final ly it can be concluded that the tether elastic vibration suppression
though offset control in conjunction with thruster and momentum-wheel attitude
control presents a viable strategy for regulating the dynamics tethered satellite sysshy
tems
121
STSS configuration 3-Body a1(0)=a2(0)=at(0)=2deg P1(0)=p2(0)=pt(0)=1deg 8x(0)=14m6y(0)=52(0)=1m dx(0)=dy(0)=dz(0)=1m Stationkeeping I = 20km LQGLTR Vibration Control
Sat 1 Pitch and Roll Angles Tether Pitch and Roll Angles mdash 1 1 1 1 1 1 I 1 1 lt 1 ~ ^ 1 ^ l _
Time (Orbits) Time (Orbits) Sat 1 Roll Control Moment Tether Roll Control Thrust
(a) Time (Orbits) Time (Orbits)
Figure 5-8 Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT attitude controller and LQGLTR offset vibration controller (a) attitude and libration controller reshysponse
122
STSS configuration 3-Body a1(0)=a2(0)=a(0)=2deg p1(0)=(32(0)=Q(0)=1deg ox(0)=14mSy(0)=5z(0)=1m dx(0)=dy(0)=dz(0)=1m Stationkeeping lt = 20km LQGLTR Vibration Control
Tether X Offset Position
Time (Orbits) Tether Y Tranverse Vibration 20r
Time (Orbits) Tether Y Offset Position
t o
1 2 Time (Orbits)
Tether Z Tranverse Vibration Time (Orbits)
Tether Z Offset Position
(b)
Figure 5-8
Time (Orbits) Time (Orbits)
Controlled dynamics of the three-body STSS during stationkeeping using the nonlinear rigid FLT attitude controller and LQGLTR offset vibration controller (b) vibration and offset response
123
6 C O N C L U D I N G R E M A R K S
61 Summary of Results
The thesis has developed a rather general dynamics formulation for a multi-
body tethered system undergoing three-dimensional motion The system is composed
of multiple rigid bodies connected in a chain configuration by long flexible tethers
The tethers which are free to undergo libration as well as elastic vibrations in three
dimensions are also capable of deployment retrieval and constant length stationkeepshy
ing modes of operation Two types of actuators are located on the rigid satellites
thrusters and momentum-wheels
The governing equations of motion are developed using a new Order(N) apshy
proach that factorizes the mass matrix of the system such that it can be inverted
efficiently The derivation of the differential equations is generalized to account for
an arbitrary number of rigid bodies The equations were then coded in FORTRAN
for their numerical integration with the aid of a symbolic manipulation package that
algebraically evaluated the integrals involving the modes shapes functions used to dis-
cretize the flexible tether motion The simulation program was then used to assess the
uncontrolled dynamical behaviour of the system under the influence of several system
parameters including offset at the tether attachment point stationkeeping deployshy
ment and retrieval of the tether The study covered the three-body and five-body
geometries recently flown OEDIPUS system and the proposed BICEPS configurashy
tion It represents innovation at every phase a general three dimensional formulation
for multibody tethered systems an order-N algorithm for efficient computation and
application to systems of contemporary interest
Two types of controllers one for the attitude motion and the other for the
124
flexible vibratory motion are developed using the thrusters and momentum-wheels for
the former and the variable offset position for the latter These controllers are used to
regulate the motion of the system under various disturbances The attitude controller
is developed using the nonlinear Feedback Linearization Technique ( F L T ) and is based
on the nonlinear but rigid model of the system O n the other hand the separation
of the longitudinal and transverse frequencies from those of the attitude response
allows for the development of a linear optimal offset controller using the robust Linear
Quadratic Gaussian-Loop Transfer Recovery ( L Q G L T R ) method The effectiveness
of the two controllers is assessed through their subsequent application to the original
nonlinear flexible model
More important original contributions of the thesis which have not been reshy
ported in the literature include the following
(i) the model accounts for the motion of a multibody chain-type system undershy
going librational and in the case of tethers elastic vibrational motion in all
the three directions
(ii) the dynamics formulation is based on a recursive O(N) Lagrangian algorithm
that efficiently computes the systems generalized acceleration vector
(iii) the development of an attitude controller based on the Feedback Linearizashy
tion Technique for a multibody system using thrusters and momentum-wheels
located on each rigid body
(iv) the design of a three degree of freedom offset controller to regulate elastic v i shy
brations of the tether using the Linear Quadratic Gaussian and Loop Transfer
Recovery method ( L Q G L T R )
(v) substantiation of the formulation and control strategies through the applicashy
tion to a wide variety of systems thus demonstrating its versatility
125
The emphasis throughout has been on the development of a methodology to
study a large class of tethered systems efficiently It was not intended to compile
an extensive amount of data concerning the dynamical behaviour through a planned
variation of system parameters O f course a designer can easily employ the user-
friendly program to acquire such information Rather the objective was to establish
trends based on the parameters which are likely to have more significant effect on
the system dynamics both uncontrolled as well as controlled Based on the results
obtained the following general remarks can be made
(i) The presence of the platform products of inertia modulates the attitude reshy
sponse of the platform and gives rise to non-zero equilibrium pitch and roll
angles
(ii) Offset of the tether attachment point significantly affects the equil ibrium orishy
entation of the system as well as its dynamics through coupling W i t h a relshy
atively small subsatellite the tether dynamics dominate the system response
with high frequency elastic vibrations modulating the librational motion On
the other hand the effect of the platform dynamics on the tether response is
negligible As can be expected the platform dynamics is significantly affected
by the offset along the local horizontal and the orbit normal
(iii) For a three-body system deployment of the tether can destabilize the platform
in pitch as in the case of nonzero offset However the roll motion remains
undisturbed by the Coriolis force Moreover deployment can also render the
tether slack if it proceeds beyond a critical speed
(iv) Uncontrolled retrieval of the tether is always unstable in pitch It also leads
to high frequency elastic vibrations in the tether
(v) The five-body tethered system exhibits dynamical characteristics similar to
those observed for the three-body case As can be expected now there are
additional coupling effects due to two extra bodies a rigid subsatellite and a
126
flexible tether
(vi) Cartwheeling motion of the B I C E P S configuration can also be used to adshy
vantage in the deployment of the tether However exceeding a crit ical ini t ia l
cartwheeling rate can result in a large tension in the tether which eventually
causes the satellite to bounce back rendering the tether slack
(vii) Spinning the tether about its nominal length as in the case of O E D I P U S
introduces high frequency transverse vibrations in the tether which in turn
affect the dynamics of the satellites
(viii) The F L T based controller is quite successful in regulating the attitude moshy
tion of both the tether and rigid satellites in a short period of time during
stationkeeping deployment as well as retrieval phases
(ix) The offset controller is successful in suppressing the elastic vibrations of the
tether wi thin the allowable l imits of the offset mechanism ( plusmn 2 0 m)
62 Recommendations for Future Study
A s can be expected any scientific inquiry is merely a prelude to further efforts
and insight Several possible avenues exist for further investigation in this field A
few related to the study presented here are indicated below
(i) inclusion of environmental effects particularly the free molecular reaction
forces as well as interaction with Earths magnetic field and solar radiation
(ii) addition of flexible booms attached to the rigid satellites providing a mechashy
nism for energy dissipation
(iii) validation of the new three dimensional offset control strategies using ground
based experiments
(iv) animation of the simulation results to allow visual interpretation of the system
dynamics
127
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[63] Stein G and Athans M The L Q G L T R Procedure for Mult ivariable Feedback Control Design IEEE Transactions on Automatic Control V o l A C - 3 2 No 2 February 1987 pp 105-114
[64] Borisenko A I and Tarapov I E Vector and Tensor Analysis with Applicashytions Dover Publications Inc New York U S A 1979
[65] Lass H Vector and Tensor Analysis M c G r a w - H i l l Book Company Inc New York U S A 1950
132
A P P E N D I X I TENSOR R E P R E S E N T A T I O N OF T H E
EQUATIONS OF M O T I O N
11 Prel iminary Remarks
The derivation of the equations of motion for the multibody tethered system
involves the product of several matrices as well as the derivative of matrices and
vectors with respect to other vectors In order to concisely code these relationships
in FORTRAN the matrix equations of motion are expressed in tensor notation
Tensor mathematics in general is an extremely powerful tool which can be
used to solve many complex problems in physics and engineering[6465] However
only a few preliminary results of Cartesian tensor analysis are required here They
are summarized below
12 Mathematical Background
The representation of matrices and vectors in tensor notation simply involves
the use of indices referring to the specific elements of the matrix entity For example
v = vk (k = lN) (Il)
A = Aij (i = lNj = lM)
where vk is the kth element of vector v and Aj is the element on the t h row and
3th column of matrix A It is clear from Eq(Il) that exactly one index is required to
completely define an entire vector whereas two independent indices are required for
matrices For this reason vectors and matrices are known as first-order and second-
order tensors respectively Scaler variables are known as zeroth-order tensors since
in this case an index is not required
133
Matrix operations are expressed in tensor notation similar to the way they are
programmed using a computer language such as FORTRAN or C They are expressed
in a summation notation known as Einstein notation For example the product of a
matrix with a vector is given as
w = Av (12)
or in tensor notation W i = ^2^2Aij(vkSjk)
4-1 ltL3gt j
where 8jk is the Kronecker delta Dropping the summation symbol the matrix-vector
product can be expressed compactly as
Wi = AijVj = Aimvm (14)
Here j is a dummy index since it appears twice in a single term and hence can
be replaced by any other index Note that since the resulting product is a vector
only one index is required namely i that appears exactly once on both sides of the
equation Similarly
E = aAB + bCD (15)
or Ej mdash aAimBmj + bCimDmj
~ aBmjAim I bDmjCim (16)
where A B C D and E are second-order tensors (matrices) and a and b are zeroth-
order tensors (scalers) In addition once an expression is in tensor form it can be
treated similar to a scaler and the terms can be rearranged
The transpose of a matrix is also easily described using tensors One simply
switches the position of the indices For example let w = ATv then in tensor-
notation
Wi = AjiVj (17)
134
The real power of tensor notation is in its unambiguous expression of terms
containing partial derivatives of scalers vectors or matrices with respect to other
vectors For example the Jacobian matrix of y = f(x) can be easily expressed in
tensor notation as
di-dx-j-1 ( L 8 )
mdash which is a second-order tensor as expected If f(x) = Ax where A is constant then
f - a )
according to the current definition If C = A(x)B(x) then the time derivative of C
is given by
C = Ax)Bx) + A(x)Bx) (110)
or in tensor notation
Cij = AimBmj + AimBmj (111)
= XhiAimfcBmj + AimBmj^k)
which is more clear than its equivalent vector form Note that A^^ and Bmjk are
third-order tensors that can be readily handled in F O R T R A N
13 Forcing Function
The forcing function of the equations of motion as given by Eq(260) is
- u 1-SdM- dqtdPe 9 f i m F(q qt) = Mq - -q mdashq + mdashmdash - Qd (260)
This expression can be converted into tensor form to facilitate its implementation
into F O R T R A N source code
Consider the first term Mq From Eq(289) the mass matrix M is given by
T M mdash Rv MtRv which in tensor form is
Mtj = RZiM^R^j (112)
135
Taking the time derivative of M gives
Mij = qsRv
ni^MtnrnRv
mj + Rv
niMtnTn^Rv
mj + Rv
niMtnmRv
mjs) (113)
The second term on the right hand side in Eq(260) is given by
-JPdMu 1 bdquo T x
2Q ~cWq = 2Qs s r k Q r ^ ^
Expanding Msrk leads to
M src = RnskMtnmRmr + RnsMtnmkRmr + RnsMtnmRmrk- (L15)
Finally the potential energy term in Eq(260) is also expanded into tensor
form to give dPe = dqt dPe
dq dq dqt = dq^QP^
and since ^ = Rp(q)q then
Now inserting Eqs(I13-I17) into Eq(260) and rearranging the tensor form of the
forcing term can how be stated as
Fk(q^q^)=(RP
sk + R P n J q n ) ^ - - Q d k
+ QsQr | (^Rlks ~ 2 M t n m R m r
+ (^RnkMtnms ~ 2~RnsMtnmgtk^ R m r
(^RnkRmrs ~ ~2RnsRmrk^j
where Fk(q q t) represents the kth component of F
(118)
136
A P P E N D I X II R E D U C E D EQUATIONS OF M O T I O N
II 1 Prel iminary Remarks
The reduced model used for the design of the vibration controller is represented
by two rigid end-bodies capable of three-dimensional attitude motion interconnected
with a fixed length tether The flexible tether is discretized using the assumed-
mode method with only the fundamental mode of vibration considered to represent
longitudinal as well as in-plane and out-of-plane transverse deformations In this
model tether structural damping is neglected Finally the system is restricted to a
nominal circular orbit
II 2 Derivation of the Lagrangian Equations of Mot ion
Derivation of the equations of motion for the reduced system follows a similar
procedure to that for the full case presented in Chapter 2 However in this model the
straightforward derivation of the energy expressions for the entire system is undershy
taken ignoring the Order(N) approach discussed previously Furthermore in order
to make the final governing equations stationary the attitude motion is now referred
to the local LVLH frame This is accomplished by simply defining the following new
attitude angles
ci = on + 6
Pi = Pi (HI)
7t = 0
where and are the pitch and roll angles respectively Note that spin is not
considered in the reduced model hence 7J = 0 Substituting Eq(IIl) into Eq(214)
137
gives the new expression for the rotation matrix as
T- CpiSa^ Cai+61 S 0 S a + e i
-S 0 c Pi
(II2)
W i t h the new rotation matrix defined the inertial velocity of the elemental
mass drrii o n the ^ link can now be expressed using Eq(215) as
(215)
Since the tether length is assumed to be fixed T j f j and Tj$jlt^ of Eq(215) are
identically zero Also defining
(II3) Vi = I Pi i
Rdm- f deg r the reduced system is given by
where
and
Rdm =Di + Pi(gi)m + T^iSi
Pi(9i)m= [Tai9i T0gi T^] fj
(II4)
(II5)
(II6)
D1 = D D l D S v
52 = 31 + P(d2)m +T[d2
D3 = D2 + l2 + dX2Pii)m + T2i6x
Note that $ j for i = 2 is defined in Section 212 whereas for i = 1 and 3 it is
the null matrix Since only one mode is used for each flexible degree of freedom
5i = [6Xi6yi6Zi]T for i = 2
The kinetic energy of the system can now be written as
i=l Z i = l Jmi RdmRdmdmi (UJ)
138
Expanding Kampi using Eq(II4)
KH = miDD + 2DiP
i j gidmi)fji + 2D T ltMtrade^
+tf J P^(9i)Pi9i)d^l + 2mT j PFmTi^dm^ (II8)
where the integrals are evaluated using the procedure similar to that described in
Chapter 2 The gravitational and strain energy expressions are given by Eq(247)
and Eq(251) respectively using the newly defined rotation matrix T in place of
TV
Substituting the kinetic and potential energy expressions in
d ( 8Ke d(Ke - Pe) i bull i Q ^ 0 (II9)
d t dqred d(ired
with
Qred = [ algt l gt a 2 2 A 2 A 2 lt ^ 2 Q 3 3 ] (5-14)
and d2 regarded as a time-varying parameter the nonlinear non-autonomous equashy
tions of motion for the reduced system can finally be expressed as
MredQred + fred = reg- (5-15)
139