5.6 Modeling of Induction Motor under Light Load Condition
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Analysis of Switching Transient Overvoltage in the Power System of Floating Production Storage and Offloading Vessel Master of Science Thesis Haoyan Xue Department of Electrical Sustainable Energy Faculty of Electrical Engineering, Mathematics and Computer Science Delft University of Technology
Transcript of 5.6 Modeling of Induction Motor under Light Load Condition
Analysis of Switching Transient Overvoltage in the Power
System of Floating Production Storage and Offloading Vessel
Master of Science Thesis
Haoyan Xue
Department of Electrical Sustainable Energy
Faculty of Electrical Engineering, Mathematics and Computer Science
Delft University of Technology
Analysis of Switching Transient Overvoltage in the Power
System of Floating Production Storage and Offloading Vessel
Haoyan Xue
Thesis Committee:
Prof. Lou van der Sluis
Dr.Ir. Marjan Popov
Dr.Ir. Dhiradj Djairam
Mr. Mark Ringlever
Mr. Jan de Vreede
Department of Electrical Sustainable Energy
Faculty of Electrical Engineering, Mathematics and Computer Science
Delft University of Technology
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Abstract
Large transient overvoltages can be caused by the switching operation of vacuum
circuit breakers of induction motors. In order to analyze the switching transient
overvoltage and use an appropriate protective method in the power system of floating
production storage and offloading (FPSO) vessel, the accurate models of electrical
equipments are necessary. In this study, vacuum circuit breakers, generators, cables,
busbars, surge arresters and induction motors are modeled in Alternative Transient
Program-Electromagnetic Transients Program (ATP-EMTP) software. The switching
transient overvoltages of four typical induction motors under the starting, the full load
and the light load working conditions in the power system of the selected FPSO vessel
are analyzed. A suitable protection against the switching transient overvoltage is
included in this study.
Key Words: switching transient overvoltage, ATP-EMTP, FPSO, VCB, cable,
induction motors, surge arrester, protection.
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Acknowledgement
I would love to express my gratitude to all of the people who have given me a lot of
help and contributed greatly to this work.
I am very grateful to Mr. Mark Ringlever who offers me this challenging project in
SBM Schiedam. I am truly indebted and thankful for Mr. Marjan Popov and Mr. Jan
de Vreede to be my daily supervisor at TU Delft and SBM Schiedam respectively.
I would like to thank Prof. Lou van der Sluis and Mr. Dhiradj Djairam to be members
of academic committee.
Very special thanks of mine go to my colleagues: Mr. Ricardinho Pietersz and Mr.
Zoran Maric, who have provided much useful information to this study.
Last but not least, I wish to thank my parents for their encouragement and undivided
support.
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Contents
Abstract ...................................................................................................................................... i
Acknowledgement ..................................................................................................................... ii
Abbreviations ............................................................................................................................ v
Chapter 1 Introduction ............................................................................................................... 1
1.1 Background .......................................................................................................................1
1.2 Aims and Scopes ..............................................................................................................2
1.3 Outline of Thesis ..............................................................................................................3
Chapter 2 Modeling of Vacuum Circuit Breaker ...................................................................... 4
2.1 Introduction ......................................................................................................................4
2.2 Arcing Time .....................................................................................................................5
2.3 Current Chopping .............................................................................................................5
2.4 Recovery of Dielectric Strength .......................................................................................6
2.5 Quenching Capability of HF Current ...............................................................................7
2.6 Test Circuit and Results ...................................................................................................8
2.6.1 Introduction of Test Circuit ...................................................................................... 8
2.6.2 Test Results of VCB ................................................................................................. 9
2.6.3 Sensitivity Analysis of VCB .................................................................................. 14
2.7 Conclusions ....................................................................................................................16
Chapter 3 Modeling of Different Equipments in Power System ............................................. 17
3.1 Modeling of Cable in ATP-EMTP .................................................................................17
3.1.1 Introduction ............................................................................................................ 17
3.1.2 Geometry of Cable ................................................................................................. 18
3.1.3 Material properties .................................................................................................. 22
3.1.4 Sensitivity Analysis of Semiconducting Layer ...................................................... 23
3.2 Modeling of Generator and Busbar ................................................................................25
3.3 Modeling of Busbar ........................................................................................................26
3.4 Conclusions ....................................................................................................................26
Chapter 4 Surge Protection Device ......................................................................................... 28
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4.1 Introduction ....................................................................................................................28
4.2 Impulse function for Current Source ..............................................................................28
4.3 Modeling of Surge Arrester ............................................................................................31
4.4 Conclusions ....................................................................................................................37
Chapter 5 Modeling of Induction Motor ................................................................................. 38
5.1 Introduction ....................................................................................................................38
5.2 Modeling of Induction Motor under Starting Condition ................................................38
5.3 Evaluation Circuit and Simulation Results of Motor under Starting Condition .............40
5.4 Modeling of Induction Motor under Full Load Condition .............................................49
5.4.1 Optimized Parameters of Motor under Full Load Condition .................................. 49
5.5 Evaluation Circuit and Simulation Results of Motor under Full Load Condition ..........57
5.6 Modeling of Induction Motor under Light Load Condition ...........................................60
5.7 Evaluation Circuit and Simulation Results of Motor under Light Load Condition .......60
5.8 Location and Protective Effect of Surge Arrester ..........................................................62
5.9 Conclusions ....................................................................................................................64
Chapter 6 Analysis of Switching Transient Overvoltage in the Power System of A FPSO
Vessel ...................................................................................................................................... 65
6.1 Introduction ....................................................................................................................65
6.2 Results of Switching Operation of Motor under Starting Condition ................................. 67
6.3 Results of Switching Operation of Motor under Full Load Condition ...........................71
6.4 Results of Switching Operation of Motor under Light Load Condition .........................73
Chapter 7 Conclusions and Future Work ................................................................................ 76
7.1 Conclusions ....................................................................................................................76
7.2 Future Work ...................................................................................................................77
References ............................................................................................................................... 78
Appendix ................................................................................................................................. 80
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Abbreviations
- FPSO: Floating Production Storage and Offloading
- ATP-EMTP: Alternative Transient Program-Electromagnetic
Transients Program
- VCB: Vacuum Circuit Breaker
- IEC: International Eletrotechnical Commission
- IEEE: Institute of Electrical and Electronics Engineers
- TRV: Transient Recovery Voltage
- HF: High Frequency
- BIL: Basic Insulation Level
- MOV: Metal Oxide Varistor
- TACS: Transient Analysis Control System
- UM: Universal Machine
- EVA: Ethylene Vinyl Acetate
- SiC: Silicon Carbide
- EMF: Electromotive Force
- ABS: American Shipping of Bureau
- RMS: Root Mean Square
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Chapter 1 Introduction
1.1 Background
The vacuum circuit breaker (VCB) has been involved in the modern electrical power
industry since last several decades. Due to several obvious advantages such as small
size, reduced weight, less maintenance, great and reliable performance of interruption,
the VCB has become one of the most widely used circuit breakers in the medium
voltage power system. However, everything has another side. The excellent
performance of interruption and recovery of dielectric strength cannot cover the
phenomenon of switching transient overvoltage due to multiple reignitions and virtual
current chopping of VCB. This phenomenon of VCB is associated with a series of
characteristics including current chopping, multiple reignitions, voltage escalations
and virtual current chopping [1]. In this study, the prestrikes which are caused from
the closing operation of VCB are not considered. The current chopping is referred to
the event that could lead to overvoltage during interruption of capacitive and
inductive current [2]. When the vacuum arc is conducting a small current, the arc
could become unstable and disappear before the zero point of the current. Once the
current is chopped, the transient recovery voltage (TRV) appears between the contacts
of VCB. If the TRV exceeds the dielectric strength of vacuum gap, the reignition
occurs. The multiple reignitions refer to the situation where reignition and interruption
of high frequency (HF) current repeat several times [1]. If the HF current due to
multiple reignitions of one phase is flowing into the other two phases through the
electrical couplings of the load, the HF current could superimpose on the power
frequency current of the other two phases and force the power frequency current to
zero [1]. Therefore, virtual current chopping occurs and will lead to serious
overvoltage in three phases.
The 11kV electrical power system of floating production, storage and offloading
(FPSO) vessel is very compact. Thirteen induction motors ranging from 0.8MW to
10.9MW are working together to provide power to different driven machines such as
compressors and pumps for supplying continued process of oil and gas production.
Each motor is equipped with a vacuum circuit breaker or fused vacuum contactor,
depending on the rated power of the motor. Once the switching operation is performed,
the overvoltage could occur, and if the overvoltage reaches to the basic insulation
level of motor, the insulation system of motor can be deteriorated and damaged
followed by the possible failure of motor. Consequently, the total reliability of
electrical power system in the FPSO vessel is decreased. In this study, the major
purpose is to check the switching transient overvoltage for four typical induction
motors, and if the overvoltage exceeds the basic insulation level, the suitable way to
mitigate the overvoltage is also discussed in this study.
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1.2 Aims and Scopes
The main part of this work is finished in Alternative Transient
Program-Electromagnetic Transients Program (ATP-EMTP). The ATP-EMTP is a
universal program for digital simulation of transient phenomena [3]. With the help of
this software, the complex power system in the FPSO vessel can be simulated and
analyzed. In order to perform the function of ATP-EMTP, the different electrical
equipments should be modeled first. A so called stochastic model is used to model the
VCB and its statistical behavior [4, 5]. The parameters of VCB are acquired from the
Siemens medium voltage VCB. The cable is modeled by the JMarti model in
ATP-EMTP [3, 6], which requires dimensions and electrical characteristics of
materials inside the cable such as thickness of different layers and relative permittivity
of insulation. The generator is represented by an ideal voltage source behind its
subtransient impedance [7]. The busbar of 11kV switchgear is treated to be the same
as the short overhead lines due to the similar environment of installation [8]. Three
models are used to represent the induction motors, one is the motor under starting
condition and the others are the motor under full and light load conditions. The surge
arrester is chosen as the protective device to mitigate the switching transient
overvoltage. The metal oxide varistor (MOV) component is used to model the surge
arrester in ATP-EMTP [8, 9].
After finish modeling all the equipments, a simplified layout of 11kV power system in
the selected FPSO vessel is used to verify the switching transient overvoltage for four
typical induction motors under starting, full and light load conditions: the 10.2MW
main gas compressor A motor connected by 180m cable to VCB, the 10.2MW main
gas compressor B motor connected by 300m cable to VCB, the 5.5MW water
injection pump motor connected by 160m cable to VCB, and the 1.25MW refrigerant
compressor motor connected by 240m cable to VCB.
The switching transient overvoltage of four typical motors under different working
conditions is calculated and compared with the requirement of basic insulation level
(BIL) in IEC 60034-15 standard. Once the overvoltage exceeds the BIL of motor, the
surge arrester is connected to the different locations of motor to find out the most
suitable position to mitigate the overvoltage.
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1.3 Outline of Thesis
This thesis is organized by the following sequence.
Chapter 1 is the general introduction of work, including the background, the method,
and the objective.
Chapter 2 discusses the characteristics, the modeling and the sensitivity analysis of
vacuum circuit breaker.
Chapter 3 introduces the modeling of equipments of power system, including the
cable, the generator and the busbar.
Chapter 4 explains the functions of surge arrester and surge capacitor. The method to
generate different impulse waves is discussed. Then the model of surge arrester is
tested in this chapter.
Chapter 5 mainly studies the modeling of induction motor under starting, full load and
light load working conditions. A comparison of switching transient overvoltage due to
three different working conditions and arcing time is given. Then the influence and
location of surge arrester are discussed.
Chapter 6 a comparison of switching transient overvoltage for four typical motors
under different working conditions is given. The effect of surge arrester in the worst
situation is also discussed.
Chapter 7 gives an overall review of what have been achieved in this thesis. Then, the
possible suggestions and future work are proposed.
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Chapter 2 Modeling of Vacuum Circuit Breaker
2.1 Introduction
In order to do transient analysis in power system, the correct model of vacuum circuit
breaker (VCB) is very important. The IEEE database provides several published
papers about development of VCB [5, 10, 11]. However, due to the confidential
information of vendors, the limitations of experiments and the statistical behavior of
VCB, the accurate model of VCB is very difficult to obtain. Therefore, the stochastic
model of VCB is proposed by several researchers [4, 5, 8]. In this study, the VCB
model is based on the stochastic model and all the parameters are chosen by normal
distribution from the general values of Siemens medium voltage VCB.
Figure 2 - 1 Basic model of VCB
The basic structure of VCB model is shown in Figure 2-1. The Default Model of
ATP-EMTP is used to generate opening or closed signal to TACS-Controlled Type 13
Switch [3]. After each switching operation of VCB, the Default Model judges and
executes the next output signal by analyzing the current and two terminal voltages
from the Type 13 Switch. During the switching operation of VCB, the arc is inevitable.
The resistance of arc can be well contributed to the voltage of arc, but the voltage of
arc is quite small compared with the system voltage and has little effects on the
transient overvoltage. Therefore, the resistance of arc can be ignored and the VCB
model is simplified.
Before explaining the VCB model in detail, four VCB key parameters are introduced
first,
Arcing Time: The time interval between instant of contacts opening and the
power frequency current becomes zero.
Current Chopping: The premature suppression of power frequency current before
current becomes zero.
Recovery of Dielectric Strength: The increases of breakdown strength of vacuum
gap.
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Quenching Capability of High Frequency (HF) Current: After the reignition
occurs in VCB, the ability of quenching HF current at zero point.
After the contacts of VCB are opened, the dielectric strength of vacuum gap will
increase with the time. If the increase of transient recovery voltage (TRV) is faster
than the increase of dielectric strength, the reignition will occur and the Default
Model will send a closed signal to Type 13 Switch. When the changing rate of HF
current in zero point is smaller than the quenching capability of VCB, the Default
Model will give an opening signal to Type 13 Switch and the HF current will be
extinguished. If further reignitions occur, the above procedure will repeat until the
dielectric strength could withstand the TRV. This phenomenon is shown in Figure 2-2.
Figure 2 - 2 Successful interruption with multiple reignitions of VCB
2.2 Arcing Time
Arching time is a random event for the different switching operations of VCB. The
opening time could be located in any point of one electrical period. In order to model
the arcing time, the uniform distribution is used to generate random opening time in
one electrical period. Then, the defined opening time is recorded in Default Model to
calculate the arcing time [4].
2.3 Current Chopping
The phenomenon of current chopping is that the power frequency current of circuit is
suppressed before its natural zero point in a VCB. If the inductive or capacitive
current is interrupted by VCB, the arc appears instantaneously and power frequency
current is conducting through the arc. When the power frequency current of first pole
to clear reaches to the low level, where a few amperes before the natural zero point,
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the arc becomes very unstable. In the other word, the arc will disappear during
conducting a small current and the power frequency current is chopped before its
natural zero point. The current chopping is a major disadvantage of VCB, because it is
accompanied with the transient overvoltage which is put on the load side due to
oscillations [12].
The accurate value of current chopping is very complex for the VCB. The early
researchers have proposed several experiments that conclude mean values of current
chopping for the modern VCB. Based on the paper [5], the load current and contact
materials have great influence on the mean value of current chopping. Generally
speaking, the value is ranging from 3A to 12A which depends on the alloy of contacts.
In this study, the contact material is chosen as Cu/Cr (75/25) which is widely used for
modern VCB. Hence, the mean value of current chopping is 6A [5]. The statistical
characteristics of current chopping are represented by normal distribution with mean
value 6A and 15% standard deviation [8, 13]. If the absolute value of the power
frequency current is smaller than the defined value of current chopping, the Default
Model sends an opening signal to Type 13 Switch and the current is chopped
immediately.
2.4 Recovery of Dielectric Strength
In general, two different breakdown mechanisms exist in the VCB. The first one is the
cold gap breakdown. After the contacts of VCB separate, the dielectric strength of
vacuum gap increases almost linearly [4]. If the TRV exceeds the dielectric strength,
the reignition will occur. This phenomenon is named cold gap breakdown and is
dependent on the recovery of dielectric strength. Besides cold gap breakdown, the
second one is the hot gap breakdown. After reignition occurs, the pressure and the
temperature of vacuum gap increase. If the arc is extinguished, the vacuum gap may
still have metal vapors and different ions. These residual particles make the
breakdown voltage decrease [4].
In this research, only the cold gap breakdown is considered, the recovery of dielectric
strength is modeled based on the paper [14]. It shows the linear relationship between
the value of dielectric strength and the time, the typical representation is shown by the
equation,
( )openU A t t B
2 - 1
Where,
U: the value of dielectric strength;
A: the slope of linear equation;
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t: the internal time of software;
topen: the opening time of VCB;
B: the intercept of linear equation;
The value of A and B is acquired from Siemens medium voltage VCB and the general
value is shown in Table 2-1.
Table 2 - 1 Dielectric characteristics of Siemens medium voltage VCB
A (kV/ms) B (V)
20-30 0
The mean value of dielectric strength is calculated by equation 2-1. The statistical
characteristics of dielectric strength are represented by using normal distribution with
the defined mean value and 15% standard deviation [8].
2.5 Quenching Capability of HF Current
If the reignition occurs, the high frequency (HF) current is superimposed on the power
frequency current through the arc. The HF current has several zero points and the
VCB has the capability to extinguish the HF current in one of its zero points [4, 8, 13].
The changing rate of HF current at zero point determines whether the VCB can
interrupt the current successfully or not.
In this study, the quenching capability of HF current is modeled based on the paper
[14]. The value is the linear relationship between the changing rate of current and the
time. The typical equation is shown below,
( )open
diC t t D
dt
2 - 2
Where,
di/dt: the value of changing rate of current;
C: the slope of linear equation;
t: the internal time of software;
topen : the opening time of VCB;
D: the intercept of linear equation;
The value of C and D is acquired from Siemens medium voltage VCB and the general
value is shown in Table 2-2.
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Table 2 - 2 Quenching capability of HF current of Siemens medium voltage VCB
C (A/(μs)2) D (kA/μs)
0 0.5-0.7
The mean value of quenching capability is calculated by equation 2-2 and the
statistical value is calculated by using normal distribution with the defined mean value
and 15% standard deviation [15]. If the changing rate of HF current in the zero point
is smaller thandi
dt, the Default Model will send an opening signal to Type 13 Switch
and the HF current will be extinguished.
2.6 Test Circuit and Results
2.6.1 Introduction of Test Circuit
A single phase test circuit of VCB is shown in Figure 2-4 [4]. The basic structure
consists of the voltage source UN, the inductance LN and the Capacitance CN that
represent the system voltage and the source impedance. The RK and the LK simulate
the resistance and the inductance of cable respectively. The capacitance of the load
and the cable is contained by CL. An ohmic-inductive load RL and LL is connected to
the VCB by the cable. The damping resistance RS, the gap capacitance CS and the
inductance LS are also considered in the test circuit. The possible values of different
elements are shown in Table 2-3.
Figure 2 - 3 Single phase test circuit of VCB
Table 2 - 3 Parameters of test circuit
Resistance (Ω) Inductance (mH) Capacitance (nF)
RS 50 LN 5 CN 100
RK 2 LS 5×10-5 CS 0.2
RL 10000 LK 0.04 CL 10
- - LL 120 - -
In order to verify the function of VCB model, the following parameters of VCB and
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source are chosen: the arcing time is between 500μs and 600μs, the recovery of
dielectric strength is 30kV/ms, the quenching capability of HF current is 0.6kA/μs and
the RMS value of UN is 6350.9V.
2.6.2 Test Results of VCB
When a load is disconnected from the network by the VCB, in general, four possible
situations can be concluded:
The load current is successfully interrupted by the VCB and without reignition.
After multiple reignitions and interruptions of HF current, the VCB could
withstand the TRV and interrupt the current successfully.
The VCB has failed to interrupt the HF current and the power frequency current
is conducting through the arc. However, the successful interruption is extended in
the next zero point of power frequency current.
The VCB is unable to interrupt current in any position of electrical period. It
makes serious damage to equipments.
The Figure 2-4 shows the first possible result which the VCB could make in switching
operation. In this case, the arcing time is around 600μs. After the contacts are opened
at 7.52ms, the increase of TRV is always slower than the pace of dielectric strength.
Therefore, the VCB successfully withstands the TRV and interrupts the load current
effectively. The upper trace of Figure 2-4 shows the clear phenomenon of current
chopping. When the current chopping occurs in the VCB, a transient overvoltage with
several kHz oscillations appears in the load side, as shown in middle trace of Figure
2-4. That is because the energy stored in the load inductance is transferred into load
capacitance to generate such transient overvoltage [1, 12]. The peak value of the
transient overvoltage on the load side is given by the equation, [12]
2 2 2( )mL P a a
LU U i
C
2 - 3
Where,
Ua: the instantaneous value of power frequency voltage at the instant of current
chopping;
ia: the level of current chopping;
ηm: the magnetic efficiency;
L and C: the impedance of load;
Moreover, if the magnetic efficiency is assumed equal to 1, the calculated result of
UL-P by equation 2-3 is around 22.6kV and the simulated result is around 21kV. The
small deviation comes from the equation 2-3, because the losses of load are not
considered in it.
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Figure 2 - 4 Upper trace: phase current in the VCB; Middle trace: voltage on the load side; Lower
trace: voltage across the contacts of VCB
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If the arcing time is decreased from 600μs to 550μs, the second situation takes over,
as shown in Figure 2-5 and Figure 2-6.
Figure 2 - 5 Voltage across the contacts of VCB under second situation
At t=7.57ms, the contacts of VCB are opened. At this moment, the level of current
chopping is calculated by the mean value with normal distribution. After the arc has
existed around 550μs, the power frequency current reaches to the level of current
chopping and is chopped immediately, as shown in upper right trace of Figure 2-6.
The voltage across the VCB starts to have a very HF oscillation after the power
frequency current is chopped, as shown in lower left trace of Figure 2-5. The HF
oscillation is induced from the interaction of inductance of cable LK, capacitance of
gap CS, capacitance of cable and load CL. The frequency is given by, [4]
1
11.8
(2 )S LK
S L
f MHzC C
LC C
2 - 4
The first peak of oscillation is called suppression peak, as shown in lower left trace of
Figure 2-5, the value is given by, [13]
5.96chI A
2 - 5
12
( ) 2691.9KS P ch S L
S L
LU I C C V
C C
2 - 6
The above HF voltage oscillation will be damped in short time interval. The
frequency of next voltage oscillation is much lower and can be defined by, [4]
2
14.6
2 L L
f kHzL C
2 - 7
It represents the natural frequency of the load, as shown in lower trace of Figure 2-4
and upper left trace of Figure 2-5.
Figure 2 - 6 Phase current in VCB under second situation
As time goes by, the TRV catches up the dielectric strength of vacuum gap at 8.25ms,
and then the first reignition will occur, as shown in lower right trace of Figure 2-5. As
a direct result of reignition, the HF current is injected into network and two HF
oscillations can be observed, as shown in Figure 2-6. The First HF oscillation is due
to the interaction of CS and LS, as shown in lower left trace of Figure 2-5. The value of
frequency can be defined by, [4]
3
150
2 S S
f MHzL C
2 - 8
13
This HF current is quickly damped and is not interrupted in its zero point by the VCB.
The second HF current is caused by the interaction of LK and CL, as shown in lower
right trace of Figure 2-6. The value of frequency can be defined by, [4]
4
10.25
2 K L
f MHzL C
2 - 9
The quenching capability of HF current of VCB plays an important role in
interrupting the second HF current. When the rate of rise of this HF current is lower
than the quenching capability of VCB in one of its zero points, the HF current is
interrupted and the TRV appears again, as shown in upper trace of Figure 2-7. After
multiple reignitions have occurred, the TRV is built to a high level, while the
dielectric strength of vacuum gap is also increased due to the movement of contacts.
Once the dielectric strength could withstand the peak of TRV, the load current will be
successfully interrupted. Besides above two situations, the third and the fourth
situations are not observed and recognized in test circuit with different combinations
of parameters of Siemens medium voltage VCB.
Figure 2 - 7 Upper trace: zoom in the first and the second reignitions of Figure 2-5; Lower trace:
zoom in the HF current of Figure 2-6
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2.6.3 Sensitivity Analysis of VCB
As has been discussed in the previous sections, the parameters of VCB are chosen
randomly with normal distribution from information of vendor. Hence, it is very
important to examine and verify the reignition behavior of different parameters of
VCB. The test circuit is the same as shown in Figure 2-3. In this part, the first check is
performed by changing the recovery of dielectric strength while keeping the
quenching capability of HF current constant. Then, the quenching capability of HF
current is changed, while the other parameters are kept constant. Finally, forty-five
different combinations of VCB parameters are analyzed. The performance is recorded
based on the random switching operations in three different arcing time intervals.
If the recovery of dielectric strength is increased from 20kV/ms to 30kV/ms, the
numbers of reignitions are decreased, as shown in Figure 2-8. Obviously, the reason
for this phenomenon is that the VCB with higher recovery of dielectric strength can
build up the higher dielectric strength in short time to withstand the TRV.
Figure 2 - 8 Voltage across the contacts of VCB; Left trace: dielectric strength = 20kV/ms,
quenching capability = 0.5kA/μs; Right trace: dielectric strength = 30kV/ms, quenching capability
= 0.5kA/μs
As shown in Figure 2-9, the higher quenching capability of HF current, the more
voltage spikes across the contacts of VCB. The effect of spike reveals a potential risk
in that if the VCB has lower quenching capability of HF current, it may experience
more failures of interruption. If the quenching capability is decreased from 0.7kA/μs
to 0.5kA/μs, the HF current will be successfully interrupted at the tenth HF current
zero point instead of the ninth. It is because the VCB is waiting for the HF current
acquires enough damping to make the rate of rise of HF current is lower than the
quenching capability of VCB,
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Figure 2 - 9 Left column traces: voltage across the contacts of VCB and phase current in VCB,
dielectric strength = 30kV/ms, quenching capability = 0.5kA/μs; Right column traces: voltage
across the contacts of VCB and phase current in VCB, dielectric strength = 30kV/ms, quenching
capability = 0.7kA/μs
The accurate performance of VCB is difficult to simulate, the reasons are not only the
confidential information of vendor, but also the limitation of experiment. The model
of VCB in this chapter is based on the general parameters of Siemens medium voltage
VCB. Therefore, the sensitivity analysis of parameters is important to define the trend
which would lead to the serious multiple reignitions and voltage escalation in the
VCB. Before starting the analysis, the boundaries of simulations should be defined.
First of all, three arcing time intervals are chosen: 0-150μs, 150-300μs and 300-600μs.
Each opening time of VCB is generated randomly by uniform distribution. Then, as
shown in Appendix, the second boundary is about the combination of parameters,
totally forty-five combinations of parameters are simulated within three arcing time
intervals. After each simulation is finished, the peak voltage across the contacts of
VCB is recorded and the final performance of VCB with different parameters is
drawn on Figure 2-10.
As mentioned in previous sections, the VCB has risks to get multiple reignitions and
voltage escalations during switching operation. The worst case occurs in the
combination of middle part of quenching capability of HF current and minimum part
of recovery of dielectric strength. Moreover, when do the comparison of parameters in
Figure 2-10, it can be concluded that the quenching capability of HF current has less
effect on the multiple reignitions and voltage escalations. Furthermore, the shorter
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arcing time, the more serious multiple reignitions and the voltage escalations.
Consequently, the arcing time and the recovery of dielectric strength play a dominant
role in the multiple reignitions and the voltage escalations of VCB. The worst
combination of parameters will be used in the next several chapters.
Figure 2 - 10 Reignition behavior of Siemens medium voltage VCB
2.7 Conclusions
In this section, the detail characteristics and phenomena of VCB are discussed. With
the information of Siemens, a vacuum circuit breaker model that includes the current
chopping, the multiple reignitions and the voltage escalations is introduced. An
ohmic-inductive load test circuit is given to perform the sensitivity analysis with
different combinations of VCB parameters. Moreover, the result of sensitivity analysis
has shown that the arcing time and the recovery of dielectric strength play a dominant
role in the estimation of multiple reignitions and overvoltage, whereas the quenching
capability of HF current has less effect. Consequently, if the Siemens medium voltage
VCB is used to analyze the overvoltage due to switching operation in power system,
the combination of middle part of quenching capability of HF current and minimum
part of recovery of dielectric strength imply that the worst multiple reignitions and
overvoltage could occur. In order to give the most suitable protective scheme of
overvoltage, the VCB parameters that could generate the worst case will be used in
the later chapters.
17
Chapter 3 Modeling of Different Equipments in
Power System
3.1 Modeling of Cable in ATP-EMTP
3.1.1 Introduction
In order to acquire an accurate simulation result of overvoltage, adequate estimation
of cable is needed in the simulation. The purpose of this chapter is to accurately
model cable in ATP-EMTP and make cable response in a large frequency spectrum.
Several models of cable are available in ATP-EMTP, such as PI, J-Marti and Semlyen
[3]. In this study, the J-Marti model is chosen, because it could represent the
performance of cable best in a wide range of frequency [3]. The following basic
equations show the electrical characteristics of insulated cable, [16]
( ) ( ) ( )Z R j L
3 - 1
( ) ( ) ( )Y G j C
3 - 2
Where,
R: the series resistance per unit length of cable system;
L: the series inductance per unit length of cable system;
G: the shunt conductance per unit length of cable system;
C: the shunt capacitance per unit length of cable system;
The electrical frequency ω is the variable that makes the above quantities are
calculated as the function of frequency. Both Z and Y are calculated by using the
Cable Constants with information of geometry and material properties from CCI cable
datasheet and IEC 60092-354 standard. In general, the information below must be
specified, [3, 16]
Geometry of Cable System:
- Position of each cable (X-Y coordinates);
- Diameters of conductor and thickness of different layers inside the cable;
Material properties:
- Resistivity and relative permeability of conductor;
- Resistivity and relative permeability of surrounding material;
- Relative permittivity of insulating material;
18
The model of cable is achieved by ATP-EMTP which does not consider the effect of
stranded conductors, inner and outer semiconducting layers and metal screen of
insulation. Therefore, an approximate method is combined with Cable Constants to
take account of the effect of stranded conductors and semiconducting layers in the
model of cable. The CCI XAI Type O 8.7/15kV 1×300mm2 medium voltage single
core power cable of marine and offshore application is modeled as an example in this
chapter and will be used in the later studies.
3.1.2 Geometry of Cable
The typical CCI XAI single core cable consists of conductors, different layers, metal
screen, sheath and armor, as shown in Figure 3-1. The conductor is made of tinned
flexible compact copper, and it is a kind of stranded conductors. The insulation is one
of the most important parts of cable. It is used to prevent the connection between the
conductor and the sheath. Therefore, it should have the capacity to withstand the
electrical field of cable for both steady and transient states. Hard ethylene propylene
rubber (HEPR) is the insulation of CCI single core cable, and it has characteristics
which is similar to cross the linked polyethylene (XLPE). The semi-conducting layers
are placed between the conductor and the insulation, and also between the insulation
and the metal screen to prevent the partial discharge. The major purpose of metal
screen is to avoid the electrical field outside the cable, provide the return path for
charging current and conduct the fault current to earth. The metal screen is the
individual copper tape. The material of inner and outer sheath of cable is SHF1
compound. Based on the IEC 60092-359 standard, the SHF1 compound is the halogen
free thermoplastic. This kind of material has very good performance for situations
with low smoke emissions, low generation of corrosive and toxic gases and low fire
propagation characteristics. The copper wire braid armor is placed between the inner
sheath and the outer sheath to provide mechanical protection to the cable.
The geometry of cable is based on the information of CCI cable datasheets. The
geometric data of cable is shown in Table 3-1.
Table 3 - 1 Basic geometric data of cable
Construction
(mm2)
Conductor
Diameter
(mm)
Diameter Over
insulation
(mm)
Under Armour
Diameter
(mm)
Overall
Diameter
(mm)
1×300 21.7 33.0 37.0 44.0
The information above is the basic characteristics of cable configuration. The real
structure of cable is very complicated and some compulsory parameters of simulation
should be calculated further.
19
Figure 3 - 1 Cross sectional structure of cable
1r : Outer Radius of Core
(= Inner Radius of Conductor Semiconducting Layer)
2r : Outer Radius of Conductor Semiconducting Layer
(= Inner Radius of Insulation)
3r : Outer Radius of Insulation
(= Inner Radius of Insulation Semiconducting Layer)
4r : Outer Radius of Insulation Semiconducting Layer
(= Inner Radius of Metal Screen)
5r : Outer Radius of Metal Screen
(= Inner Radius of Inner Sheath)
6r : Outer Radius of Inner Sheath
(= Inner Radius of Armor)
7r : Outer Radius of Armor
20
(= Inner Radius of Outer Sheath)
8r : Outer Radius of Outer Sheath
(= Overall Radius of Cable)
1T : Thickness of Conductor Semiconducting Layer
2T : Thickness of Insulation
3T : Thickness of Insulation Semiconducting Layer
4T : Thickness of Metal Screen
5T : Thickness of Inner Sheath
6T : Thickness of Amor
Based on the IEC 60502.2-1998 standard, the nominal thickness of hard ethylene
propylene rubber (HEPR) insulation of 1×300mm2 8.7/15kV cable is 4.5mm. Hence,
the T1 and r2 can be calculated by the equations,
1 3 2 1
16.5 4.5 10.85
1.15
T r T r
mm
3 - 3
2 1 1
10.85 1.15
12
r r T
mm
3 - 4
The IEC 60502.2-1998 and IEC 60092-354 standards give the specific information of
thickness for the copper tape and the inner sheath. Besides these relevant standards,
datasheets of several manufacturers such as Prysmian, UC and TMC use the typical
values of thickness of copper tape and inner sheath for medium voltage power cable.
Therefore, the copper tape with thickness of 0.1mm and the inner sheath with
thickness of 1.4mm are chosen from the general values of manufacturers. The r3 and
r4 can be calculated,
3 2 2
16.5
r r T
mm
3 - 5
21
4 6 4 5
18.5 0.1 1.4
17
r r T T
mm
3 - 6
The thickness of armor is chosen based on IEC 60092-354 standard. The IEC
60092-354 standard requires that if the fictitious diameter under armor is larger than
30mm, the nominal diameter of the braid wire should be 0.4mm as a minimum.
Besides the IEC standard, one of the major marine and offshore cable manufacturers
TMC uses 0.4mm as diameter of braid wire. Therefore, 0.4mm is chosen as the
approximate diameter of braid wire due to the information of standard and cable
manufacturer. At each cross of braid wire, the two individual wires are superposed
with each other. Consequently, the thickness of armor can be approached by doubling
the diameter of braid wire,
6 0.8T mm
3 - 7
The r5, r6 and r7 can be calculated by equations,
5 4 4
17 0.1
17.1
r r T
mm
3 - 8
6 5 5
17.1 1.4
18.5
r r T
mm
3 - 9
7 6 6
18.5 0.8
19.3
r r T
mm
3 - 10
The above geometric information is used in the cable model of ATP-EMTP. The
material properties of cable will be explained in the next section.
22
3.1.3 Material properties
3.1.3.1 Core of Cable
In ATP-EMTP, only solid and hollow conductor can be directly modeled [3]. However,
the design of CCI cable core is stranded. Therefore, the resistivity of core material
should be adjusted to perform the effect of small space between each strand. The
resistivity of common conductor material is shown in Table 3-2,
Table 3 - 2 Resistivity of different metal materials
Material Copper Aluminum Lead Steel
ρ (Ω·m) 1.72×10-8 2.83×10
-8 2.2×10-7 1.8×10
-7
The equation of references [9, 10] provide one possible method to adjust the
resistivity of material of stranded core,
2
' 1
2
38
6
8
10.85( )
101.72 10300
10
2.12 10
c c
c
r
A
m
3 - 11
Where,
Ac: the efficient (nominal) cross sectional area of the core;
r1: the outer radius of core;
ρc’: the resistivity of material of conductors. In this study, the copper of conductor is
chosen based on the datasheet of CCI cable;
3.1.3.2 Semiconducting Layer and Surrounding Insulation of Cable
The different semiconducting layers of cable cannot be automatically included in the
cable model of ATP-EMTP. However, the effect of different semiconducting layers
can be considered through changing the electrical permittivity of insulation and
keeping the capacitance between conductor and metal screen constant [16]. Hard
ethylene propylene rubber is the insulation of cable and the relative permittivity is
around 3.2. The conversion procedure is shown in the equation, [16, 17]
23
4
1
0
10
12
ln( )
2
17ln( )
10.854.4 102 8.85 10
3.55
r
r
rC
3 - 12
Where,
C: the capacitance of cable which is obtained from datasheet of CCI cable;
ε0: 8.85×10-12
F/m;
In the above conversion procedure, the semiconducting layers are represented as part
of insulating materials of main insulation, consequently the main insulation of cable
will extend to the conductor and the metal screen. As a result of this procedure, it
makes the electrical permittivity increase to leave the capacitance unchanged [17].
Besides the relative permittivity of insulation, the relative permittivity of surrounding
materials such as the inner sheath and the outer sheath is also required for the cable
model in ATP-EMTP. As mentioned above, the material of inner sheath and outer
sheath is SHF1. The SHF1 is a kind of halogen free thermoplastic with the ethylene
vinyl acetate (EVA) as the essential polymer to make such thermoplastic [18].
Therefore, the relative permittivity of EVA is used to approach the material of SHF1
and its typical value is around 2.8 [19]. The relative permeability of conductor,
insulation and surrounding material is equal to 1 [8].
3.1.4 Sensitivity Analysis of Semiconducting Layer
The completed model of 1×300mm2 single core cable system is shown in Figure 3-2
and the length of each cable is 180m.
Figure 3 - 2 Cable model in ATP-EMTP
24
This length refers to the distance between the switchgear and KM-T7111 main gas
compressor A motor. Using the open circuit test [20], a step voltage excitation is
applied on the sending end of cable system and the response of voltage excitation is
simulated on the receiving end of cable system, as shown in Figure 3-3. The metal
screen and the armor of each cable are grounded during the open circuit test.
Figure 3 - 3 Open circuit test of cable system
Figure 3 - 4 Response of open circuit test of cable system
The response of Figure 3-4 shows that the semiconducting layers make the
propagation speed decrease, because the HF transients of cable propagate as
decoupled coaxial wave between core and metal screen. Consequently, the modeling
of core, insulation, semiconducting layers and metal screen determine the transient
behavior of cable. If the thickness of insulation is fixed, adding semiconducting layers
will increase the inductance of conductor and metal screen loop without changing the
capacitance. This will lead to decrease propagation speed and increase surge
impedance [17].
25
3.2 Modeling of Generator and Busbar
The selected FPSO vessel has four synchronous generators and each generator is
driven by a gas turbine. During normal production, three of them work together to
feed loads and one of them is used as cold redundancy. The generator is modeled as
an ideal voltage source in series with its subtransient impedance [7], as shown in
Figure 3-5. The effect of saturation, excitation, governor system and mechanical part
are not included.
Figure 3 - 5 Simplified model of generator
The subtransient impedance is modeled by the symmetric resistance and inductance
coupled line: Line_SY3. When the Line_SY3 model is used to represent the
subtransient impedance of generator, the zero sequence impedance and the positive
sequence impedance of generator are required. Cs is the bushing capacitance which is
between the generator windings and the ground. The values of different parameters
are shown in Table 3-3, and all of these values are from the vendor test report.
Table 3 - 3 Electrical data of generator
Parameter Unit Value
Rated Voltage kV 11
Rated Apparent Power kVA 29070
Power Frequency Hz 60
Base Impedance Ω 4.16
R0 Ω 0.1249
X0 Ω 0.3746
R1 Ω 0.0558
X1 Ω 0.8408
Cs μF 0.1
26
3.3 Modeling of Busbar
Rectangular busbars are placed in the high voltage switchgear of the selected FPSO
vessel. The length of each separate busbar is around 85cm to 100cm. The transient
response of busbar is similar to the short length of overhead transmission line because
they have similar environment of installation. Therefore, in the switching transient
analysis, the busbar can be modeled by the transmission line with specific
characteristic impedance [8].
In ATP-EMTP, the busbar can be accurately modeled by the JMarti frequency
dependent model and equivalent PI model. Both of them require the radius and
positions of busbars conductors in X-Y coordinate, as shown in Figure 3-6.
Figure 3 - 6 Model of busbar in ATP-EMTP
The cylindrical conductors are used to approach the performance of rectangular
conductors. Because the impedance of rectangular structure is complicated to estimate
and two models of ATP-EMTP are unable to modify the conductors with rectangular
structure. The cross sectional area of each cylindrical conductor is determined by the
rated current. The rated current of each busbar conductor is 4000A and the cross
sectional area of each conductor is around 25.87cm2. The position of each conductor
is measured from the design report of switchgear. In this study, the PI model is used
because if the busbar is very short such as 100cm or below, the JMarti model has the
internal numerical problem. In fact, this numerical problem can be solved through
decreasing the time step of simulation, but it requires more than one gigabytes
memory space of computer hard disk and causes overflow of ATP-EMTP.
3.4 Conclusions
This chapter has described the modeling of cable, generator and busbar. The cable is
modeled by JMarti frequency dependent model in ATP-EMTP, which includes the
effect of stranded conductors and semiconducting layers. The resistivity of stranded
conductor is slightly increased due to the small space between each strand. The
response of open circuit test shows that the propagation speed is decreased by the
semiconducting layers.
27
The model of generator is represented by an ideal voltage source behind its
subtransient impedance without considering the mechanical part, the excitation and
the governor systems. The busbar is treated the same to the short length overhead line.
The cylindrical conductors that have the same rated current are used to approach the
effect of rectangular busbar by PI model.
28
Chapter 4 Surge Protection Device
4.1 Introduction
The overvoltage of power system can be limited by using surge protection devices. In
general, the protection devices can be divided into two groups. All kinds of surge
arresters belong to the first group. The second group of surge protection device
consists of surge capacitors and C-R suppressors. The method of categorization is
based on the protective way [1]. The surge arrester can be used to limit the amplitude
of different kinds of overvoltages such as the switching overvoltage and the lightning
overvoltage, however it cannot modify the rate of rise of voltage surge. The function
of surge capacitor is to reduce the rate of rise of voltage surge and decrease the rate of
rise of TRV to decrease reignitions. Therefore, both groups of the protection devices
could provide suitable surge protection to electrical equipments such as motors and
transformers. In this study, the surge arrester is modeled and used in later chapters,
because the experience shows that when the overvoltage is generated by switching
operation, the limitation of amplitude of overvoltage could provide better protection
to equipments [8].
4.2 Impulse function for Current Source
The electrical response of surge arrester model should be equal or similar to the data
that is suggested by the manufacturer. In order to check the response of surge arrester
model, the different current impulse waves should be described first. In general, the
impulse wave is determined by its front time Tf and its time to half peak value Th, as
shown in Figure 4-1.
29
Figure 4 - 1 Current impulse wave 1/5μs
The front time Tf is given by the following equation, [8]
0 . 9 0 . 3
0 . 6f
T TT
4 - 1
Where,
T0.9: the time when the wave reaches 90% of the peak value;
T0.3: the time when the wave reaches 30% of the peak value;
The time to half peak value Th is determined by, [8]
0.9 0.30.3
2h c
T TT T T
4 - 2
Where,
Tc: the time when the wave reaches to the half peak value in the wave tail;
The time Tr can be calculated by following equation, [8]
0.9 0.30.3
2r
T TT T
4 - 3
Originally, the above method is applicable for determination of voltage impulse wave.
The time Tr of current impulse wave is determined by the trace that passes through 90%
30
and 10% of peak value rather than 90% and 30% of peak value. However in order to
simplify the work, the current impulse wave is tackled the same to the voltage
impulse wave [8]. In ATP-EMTP, the different current impulse waves can be
accurately generated by Heidler current source, which is shown in the following
equation, [3]
_
_
( )
( ) exp( )
1 ( )
n
f
Peakn
f
t
T tf t I
t tau
T
4 - 4
Where,
Ipeak: the peak value of current impulse;
n: the parameter determines the rate of rise of function. It is ranging from 1 to 10;
T_f: the constant determined by front time of wave;
tau: the constant determined by tail time of wave;
The reference [6] proposed the calculated parameters for Heidler function with
different current impulse waves. The optimized parameters are used to generate three
typical current impulse waves: the fast front wave 1/5μs, the lightning impulse wave
8/20μs and the switching impulse wave 30/60μs, as shown in Table 4-1. The peak
value of the current impulse is chosen from datasheet of ABB MWD surge arrester.
Table 4 - 1 Parameters of Heidler function for different current impulse waves
Impulse
(μs/μs)
Peak Value of
Current
impulse (A)
n T_f (μs) tau (μs)
1/5 1000 5 2.07 4.62
1/5 5000 5 2.07 4.62
1/5 10000 5 2.07 4.62
8/20 1000 5 16.55 9.62
8/20 5000 5 16.55 9.62
8/20 10000 5 16.55 9.62
8/20 20000 5 16.55 9.62
30/60 125 5 56.7 17.5
30/60 250 5 56.7 17.5
30/60 500 5 56.7 17.5
30/60 1000 5 56.7 17.5
31
4.3 Modeling of Surge Arrester
The amplitude of overvoltage in power system can be limited by surge arrester. The
surge arrester provides a temporal path to earth which the superfluous charge is
removed from the system. When the power system is working under normal voltage, a
very small or no leakage current is conducting in the surge arrester. However, if the
overvoltage appears due to the different operations of power system, the surge arrester
can conduct the current without causing fault [21]. The following typical surge
arresters are used,
Open Spark Gap.
The open spark gap is an early overvoltage protection device. It is simple and cheap,
however when it works, a short circuit appears in the power system due to an arc is
conducting between the line and the earth.
SiC Surge Arrester
Silicon carbide (SiC) surge arrester is an improvement of surge arrester. The spark gap
is in series with nonlinear resistance which is made of the stacks of silicon carbide
discs. The spark gap gives high impedance during normal voltage. If the overvoltage
occurs and exceeds the flashover voltage of the spark gap, a current will conduct in
nonlinear resistance, where the surge energy is dissipated and overvoltage is limited
[21].
Metal Oxide Surge Arrester
Metal oxide surge arrester is a very important improvement of SiC surge arrester,
which does not have a spark gap. The major material of metal oxide surge arrester is
the mixture of zinc oxide and additional metal oxide such as MnO [8]. Due to the
large nonlinear volt-ampere characteristic, a very small leakage current flows through
the surge arrester during normal voltage, whereas the overvoltage impulses can be
diverted with current surges of many kA [21].
As mentioned above, the metal oxide surge arrester has a good performance for surge
protection. Therefore, the 11.3kV ABB metal oxide surge arrester is modeled and used
to protect the electrical equipments in power system of the FPSO vessel. The
measured volt-ampere characteristic of 11.3kV surge arrester is shown in Fig 4-2.
32
Figure 4 - 2 V-I characteristic of 11.3kV surge arrester
With the information of volt-ampere characteristic of surge arrester, the reference [9,
22] have proposed two different models of surge arrester which can be applied in
ATP-EMTP. But all of these models require additional information such as
temperature dependent resistance and resistance-inductance filter. The reference [14]
has proposed a simplified Schmidt surge arrester model which is also used in this
study, as shown in Fig 4-3.
Figure 4 - 3 Simplified Schmidt surge arrester model
Where,
R(i): the current dependent resistance which is used to represent the non linear
voltage-ampere characteristic of metal oxide surge arrester;
Ca: the capacitance of surge arrester block.
Ra and La: the physical characteristics of ZnO grains;
The parameters of surge arrester are estimated to be: Ra = 0.06Ω, La = 0.5μH and Ca =
0.14nF based on the reference [14]. The volt-ampere characteristic of R(i) is chosen
from Figure 4-2. The residual voltages of simulations are compared with the data
sheet of ABB 11.3kV surge arrester, as shown in Table 4-2.
33
Table 4 - 2 Comparison of residual voltage between simulation and data sheet
Impulse (μs/μs)
Peak Value of
Current impulse
(A)
Peak Value of
Residual Voltage
(kV) (Simulation)
Peak Value of
Residual Voltage
(kV) (Data Sheet)
1/5 1000 23.97 -
1/5 5000 29.07 -
1/5 10000 33.86 -
8/20 1000 23.87 23.4
8/20 5000 27.54 26.1
8/20 10000 29.81 27.6
8/20 20000 33.49 30.6
30/60 125 21.11 20.3
30/60 250 21.98 21.4
30/60 500 22.89 22.1
30/60 1000 23.86 23.0
The residual voltages of simulations under different current impulse waves are shown
from Figure 4-4 to Figure 4-6.
34
Figure 4 - 4 Upper trace: 1/5μs current waves with different amplitudes; Lower trace: response of
surge arrester after excitation by 1/5μs current waves
35
Figure 4 - 5 Upper trace: 8/20μs current waves with different amplitudes; Lower trace: response of
surge arrester after excitation by 8/20μs current waves
36
Figure 4 - 6 Upper trace: 30/60μs current waves with different amplitudes; Lower trace: response
of surge arrester after excitation by 30/60μs current waves
As shown in above Figures, the residual voltage of surge arrester has relationship with
the shape of current impulse and the peak value of current impulse. If the front
steepness of current impulse or the peak value of current impulse is increased, surge
arrester experiences higher residual voltage. Due to that the model of surge arrester
includes the inductance of surge arrester, the peak value of current impulse is lagging
to the residual voltage. Table 4-2 has shown that the model of surge arrester has a
good tolerance for the 30/60μs switching current impulse, however it has slightly
higher deviation for the current impulse with shorter front time such as the 8/20
lightning current impulse. This because that the model of surge arrester does not
consider the hysteresis of volt-ampere characteristic [8]. In this study, the major
37
purpose is the analysis of switching operation of vacuum circuit breaker. Therefore, it
is sufficient to use this model of surge arrester, because the front time of the reignition
current impulse is not as short as that of the lightning current impulse [8].
4.4 Conclusions
In this chapter, the concepts, advantages and disadvantages of different surge arresters
are compared first. Then the steep front, the lightning and the switching current
impulse is calculated by the Heidler function in ATP-EMTP. Based on several
references, a simplified Schmidt surge arrester model is used to model the ABB
MWD 11.3kV surge arrester. The simulated response of surge arrester model is
compared with the datasheet of ABB and a good tolerance is achieved. This model
of surge arrester will be used in later chapters.
38
Chapter 5 Modeling of Induction Motor
5.1 Introduction
Induction motor is one of the most important electrical loads to allow continued
processing in a FPSO vessel. Based on the driven machines such as compressor and
pump, the rated power of high voltage induction motors are ranging from several
hundreds of kW to several MW. In this study, the high voltage induction motor is
modeled according to three different switching operations. The first one is the
switching operation of motor under starting condition. This situation is usually from
the inappropriate settings of protection relays. Next situation is the switching
operation of motor under full load condition. It may occur due to the process trip. Last
situation is the switching operation of motor under light load condition. It is the most
common situation which disconnects the induction motor from the power system in
the FPSO vessel.
Therefore, the accurate model of induction motor is necessary for taking transient
analysis. The distribution of overvoltage stress inside the motor winding is not
analyzed in this study, because the major purpose of this study is the determination of
switching transient overvoltage in high voltage side of the selected FPSO power
system and the influence of surge arrester.
5.2 Modeling of Induction Motor under Starting Condition
The disconnection of motor under starting condition can cause more severe switching
transient overvoltage than the motor under full and light load working conditions.
When the motor is starting up, the rotor does not acquire high speed to generate
enough back electromotive force (EMF) which is opposite to the source voltage. The
low back EMF cannot keep the TRV at a low level after opening the contacts of VCB.
Therefore, the modeling of induction motor under starting condition can be regarded
as doing switching operation under the situation where the rotor is locked. As shown
in Figure 5-1, the paper [5] proposed the T-equivalent circuit of induction motor to
represent the electrical characteristics of motor under starting condition.
39
Figure 5 - 1 T-equivalent circuit or induction motor
Where,
R1 and X1: the impedance of stator;
R2’ and X2’: the impedance of rotor;
Xm: the magnetizing reactance;
s : the slip of motor;
V: the phase voltage;
In the equivalent network, the slip of motor is equal to 100%. Although the equivalent
network is simple and useful, it does not consider the bushing capacitance for each
phase and the natural oscillation of motor windings. In order to fix this problem, the
modified equivalent network is proposed and shown in the Figure 5-2,
Figure 5 - 2 Modified T-equivalent circuit of induction motor under starting condition
Where,
R1, X1, R2’ and X2’: the impedance of motor for each phase;
Cg: the bushing capacitance of each phase;
Rd: the damping resistance of motor winding for each phase;
In the modified circuit, the stator magnetizing reactance Xm is ignored, because Xm is
larger than the combination of impedance of rotor during locked rotor condition or
starting condition. The values of R1, X1, R2’ and X2’are calculated from the no load and
the locked rotor test reports of vendor [23]. Cg is the bushing capacitance that comes
from the electrical study of project and the information of vendor. Rd is the damping
40
resistance that is given by the equation, [24]
1
2( )
dm
g
m
RR
CL
5 - 1
Where,
Rm : the total resistance of each phase;
Lm : the total inductance of each phase;
τ: the attenuation time constant of natural oscillation of load itself and is related to the
amplitude factor Ψ of recovery voltage by the equation, [24]
2
( ) 11
ln( )1
m mL C
5 - 2
Where,
The typical value of amplitude factor Ψ of motor is 1.5;
5.3 Evaluation Circuit and Simulation Results of Motor
under Starting Condition
The Figure 5-3 shows the evaluation circuit that is used to analyze the switching
transient overvoltage for a typical load in the selected FPSO vessel: main gas
compressor A motor KM-T7111. Three main generators are connected to the busbar
through 4×(3×1×300mm2) cable system, which means that each phase has four
8.7/15kV single core cables. The KM-T7111 motor is connected to the VCB via
2×(3×1×300mm2) cable system. The modeling of generator, the cable and the busbar
is explained in previous chapters.
41
Figure 5 - 3 Evaluation circuit of induction motor
In this case, the VCB is used to disconnect the motor from locked rotor condition with
rated voltage 11kV to approach the starting condition. It means that after the motor is
connected to the power system for an extremely short time period, the rotor can be
regarded as stalled due to the fact that the mechanical load, and then, the terminal
voltage of motor is still around the rated voltage. Under such circumstance, the VCB
performs the switching operations in different arcing time intervals to get comparable
switching overvoltages. Before making the explanation of three typical phenomena of
switching overvoltages, the electrical characteristics of starting motor is shown in
Figure 5-4, three key characteristics are compared with the test report of vendor in
Table 5-1. The phase voltage, the phase current and the power factor show an
acceptable tolerance based on the IEC 60092-101 and the IEC 60034.01-2010
standards. The deviation of these characteristics comes from the parameters of
equivalent network.
Figure 5 - 4 Simulated voltage and current of Phase A of KM-T7111 motor under starting
condition
42
Table 5 - 1 Comparison of characteristics of motor under starting condition between simulation
and data sheet
Item Results of
Simulation
Results of Test Report
and Datasheet
Tolerance of
IEC 60092-101
IEC 60034.01-2010
Peak Value of
Phase Voltage (V) 9016.1 8981.5 +6% to -10%
Peak Value of
Phase Current (A) 2918.5(3.29p.u) 2905.1(3.28p.u) +20%
Power Factor
under Starting
Condition
0.14 0.12 Maximum absolute
value: 0.07
The first switching operation is performed by setting the arcing time around 850μs. In
this case, the phase A of KM-T7111 motor is the first pole to clear. Figure 5-5 shows
that the maximum switching overvoltage of terminal of motor is almost 12kV (1.3pu)
in phase A. Another important factor is that the whole switching operation does not
have any reignitions, this can be explained with cooperation of Figure 5-6, which
shows the voltage across contacts of phase A of VCB. When the contacts of VCB are
opened, around 3ms, the increase of dielectric strength is always larger than the
increase of TRV. Therefore, the VCB could withstand the TRV and make successful
interruption without reignitions.
Figure 5 - 5 Simulated three phase voltage in terminal of KM-T7111 motor; (Red trace: phase A;
Green trace: phase B; Blue trace: phase C)
43
Figure 5 - 6 Simulated voltage across the contacts of phase A of VCB
Besides the successful interruption without reignitions, if the arcing time is decreased
from 850μs to approximately 350μs and while the other parameters of circuit are kept
constant, the switching overvoltage in terminal of motor are increased drastically. It is
also the worst situation that could occur during a switching operation based on Figure
5-7. The peak value of overvoltage could reach to around 56kV (6.2p.u). Such a high
overvoltage could give a large stress to the insulation system of motor and decrease
the life time of motor. Besides the stress of overvoltage, the other special scenario is
shown on the Figure 5-8. Based on the explanation of previous chapters, the current of
phase A reaches to the chopping level before the other two phases, therefore, the
current of phase A should be interrupted first. However, the simulation shows the
totally different results where the current of phase C is interrupted before phase A and
phase B. Both of the above situations are the results of virtual current chopping.
44
Figure 5 - 7 Upper left trace: simulated three phase voltage in terminal of KM-T7111 motor;
Upper right trace: simulated voltage of phase A in terminal of KM-T7111 motor; Lower left trace:
simulated voltage of phase B in terminal of KM-T7111 motor; Lower right trace: simulated
voltage of phase C in terminal of KM-T7111 motor
Figure 5 - 8 Upper left trace: simulated three phase current in VCB; Upper right trace: simulated
current of phase A in VCB; Lower left trace: simulated current of phase B in VCB; Lower right
trace: simulated current of phase C in VCB
45
The virtual current chopping may occur if the first pole to clear has multiple
reignitions before the other two phases are interrupted [25]. The reignition of one
phase could cause the HF current flow into the other two phases through the electrical
coupling of load. The evaluation circuit of Figure 5-3 is simplified to represent the
path of HF current and is shown in Figure 5-9,
Figure 5 - 9 Simplified evaluation circuit for the explanation of virtual current chopping
After the VCB is opened, the current of phase B and C are far from the chopping level
and the power frequency current is still conducting through the arc. Phase A has
multiple reignitions during switching operation, as shown in Figure 5-10. The HF
current It in phase A flows to ground through the bushing capacitance Cg, and then It
could be divided into two parts while each part goes back to phase B and C via the
bushing capacitance Cg, as shown in Figure 5-9.
Figure 5 - 10 Simulated voltage across the contacts of phase A of VCB
46
Figure 5 - 11 Simulated three phase current in VCB; (Red trace: phase A; Green trace: phase B;
Blue trace: phase C)
The HF currents of phase B and C are superimposed on the power frequency current
and both the HF currents of phase B and C have the same polarity and magnitude, but
the polarity is opposite to the HF current in phase A. If the peak value of HF current
in phase B or C is larger than the peak value of power frequency current, the HF
current could force the power frequency current to zero, as shown in Figure 5-11. This
unnatural current zero phenomenon is named virtual current chopping. Compared
with the normal level of current chopping 6A for a modern VCB, the virtual chopping
level could reach to several hundred amperes and lead to larger overvoltage than the
normal case.
After the definition of virtual current chopping, Figure 5-7 and Figure 5-8 can be
reviewed. After around 20 reignitions in phase A, the first virtual current chopping is
observed in phase C due to the HF current flows through the capacitive couplings in
the load side, as shown in lower right trace of Figure 5-11. Then, the HF current of
phase C is temporarily interrupted and TRV with the very steep du/dt appears between
the contacts of phase C in VCB, as shown in lower right trace of Figure 5-12.
47
Figure 5 - 12 Upper trace: simulated voltage across contacts of phase A of VCB; Lower left trace:
simulated voltage across contacts of phase B of VCB; Lower right trace: simulated voltage across
contacts of phase C of VCB
At this moment, the contacts of phase C in VCB do not move to enough distance to
generate a strong dielectric strength which could withstand such high TRV. Therefore,
the first reignition occurs in phase C. If the virtual current chopping occurs several
times, the multiple reignitions appear in phase C until dielectric strength becomes
strong enough, as shown in lower right trace of Figure 5-12 around 6.5ms. Besides
phase C, the virtual current chopping also skips to phase B and causes the similar
multiple reignitions. After several repetitions of virtual current chopping, a
considerable overvoltage is built up in the motor side. In phase A, the HF current is
unsuccessfully interrupted and the power frequency current is conducting through the
arc until the next zero current point. The reason of unsuccessful interruption is that the
multiple reignitions make the current difficult to interrupt and some points which the
power frequency current is larger than the HF current could be reached. Consequently,
the HF current is unsuccessfully interrupted by the VCB due to the numbers of zero
crossing of HF current becomes less and less, as shown in Figure 5-13.
48
Figure 5 - 13 Simulated current of phase A in VCB
If the arcing time is further decreased to around 100μs, the peak value of switching
overvoltage could be very moderate such as 14kV (1.6pu) in Figure 5-14. In fact, the
reason of this special case can be explained from the unsuccessful interruption of
VCB. After the contacts of VCB are opened, phase A is chopped first, as shown in
Figure 5-15. Due to the short arcing time, the distance between two contacts is too
small to withstand the TRV. After the reignition occurs several times, the VCB fails to
interrupt the HF current in phase A due to the same reason as mentioned above and
the power frequency current takes over until the next current zero point. The
conducting period of HF current is shorter than the previous case and does not cause
the virtual current chopping. Consequently, the switching overvoltage is not as serious
as the previous case.
Figure 5 - 14 Simulated three phase voltage in terminal of KM-T7111 motor; (Red trace: phase A;
Green trace: phase B; Blue trace: phase C)
49
Figure 5 - 15 Simulated three phase current in VCB; (Red trace: phase A; Green trace: phase B;
Blue trace: phase C)
5.4 Modeling of Induction Motor under Full Load Condition
In ATP-EMTP, the Universal Machine Type 3(UM3) is used to model the induction
motor under full load condition, as shown in Figure 5-16 [3, 6]. The back EMF could
be represented in Type 3 model because this model is based on the d-q-0 rotor
reference coordination. Unlike the other models, the Type 3 model has not built in a
mechanical part. Therefore, the user should build the mechanical part of motor
separately by different electrical elements, for example the capacitance representing
the moment of inertia and the current source representing the load torque [3].
Figure 5 - 16 Universal machine type 3 model of ATP-EMTP
5.4.1 Optimized Parameters of Motor under Full Load Condition
The UM3 model requires accurate parameters of T-equivalent circuit of motor which
should represent the steady state of running. As has been mentioned in the previous
section, the parameters of T-equivalent circuit are calculated from test reports of
vendor such as the no load test and the locked rotor test based on the reference [23]. It
is a convenient and easy method to calculate parameters that show a good approach
for the electrical characteristics of motor under starting condition, however, it has
limitations for calculating parameters of motor under full load condition.
50
When the induction motor is working under normal running condition, the current of
rotor bar is uniformly distributed and the leakage flux lines are formed as shown in
Figure 5-17 [26]. If the working condition is changed from running to starting or
locked rotor, the mechanical speed of motor is quite low and even equal to zero. Such
a low mechanical speed results in very high slip and slip frequency in the rotor. Some
squirrel cage induction motors have a deep and narrow rotor bar, then the slot leakage
flux is concentrated on the lower part of rotor bar. During the starting or locked rotor
situation, the reactance of lower part of rotor bar is larger than the upper part of rotor
bar because the flux is not homogeneous distributed, while the current of rotor bar is
mainly flowing through the upper part. Therefore, the current change makes the
reactance of rotor conductor decrease and the resistance of rotor conductor increase.
This impedance change is named as skin effect. The large squirrel cage induction
motors are designed to perform a high starting torque and a low starting current by the
application of skin effect during start up.
Figure 5 - 17 Flux and current distribution in rotor conductor
The reference [23] has proposed a method to perform the locked rotor test with 25%
rated frequency to avoid the serious skin effect. In this case, the leakage reactance of
rated frequency could be regarded as proportional to the frequency. However, the
locked rotor test of vendor is finished with 100% rated frequency. If the five
parameters of T-equivalent circuit are calculated by the proposed method [23] without
any changes, a large deviation is inevitable. In order to minimize the deviations
caused by the skin effect, the author proposed a modified method to calculate
parameters of T-equivalent circuit of motor under full load condition with minimum
influence of skin effect. Figure 5-1 shows a typical single phase T-equivalent circuit.
If the induction motor is under no load or locked rotor test, the T-equivalent circuit
could be simplified, as shown in Fig 5-18,
51
Figure 5 - 18 Left trace: locked rotor equivalent circuit based on the Figure 5-1; Right trace: no
load equivalent circuit based on the Figure 5-1
Under the locked rotor condition, the slip is 100%. The magnetizing reactance Xm is
parallel by the rotor impedance. Due to that the Xm is very large, the magnetizing
circuit section could be regarded as open circuit. Therefore, the impedance of motor
under locked rotor condition is,
1
1
BL
VZ
I
5 - 3
2
13
BLBL
PR
I
5 - 4
'
1 2BLR R R 5 - 5
2 2
BL BL BLX Z R 5 - 6
'
1 2BLX X X 5 - 7
Where,
ZBL: the locked rotor impedance;
RBL: the locked rotor resistance;
XBL: the locked rotor reactance;
PBL: the absorbed power of motor;
Under no load condition, the circuit section of rotor could be regarded as open circuit
due to that the slip is very small and the rotor resistance is extremely large. Therefore,
the impedance of motor under no load condition is,
52
1
1
NL
VZ
I
5 - 8
0
2
13NL
PR
I
5 - 9
2 2
NL NL NLX Z R 5 - 10
1NL mX X X 5 - 11
Where,
ZNL: the no load impedance;
RNL: the no load resistance;
XNL: the no load reactance;
P0: the no load power of motor;
Based on the equation 5-7 and 5-11, the T-equivalent circuit can be modified to
eliminate the magnetizing reactance Xm and the rotor reactance X2’, as shown in
Figure 5-18.
Figure 5 - 19 Modified T-equivalent circuit of induction motor
The equivalent impedance of modified T-equivalent circuit is,
'
21 1
1 1 '
21 1
( ) ( )
( ) ( )
NL BL
th
NL BL
Rj X X j X X
sZ R jX
Rj X X j X X
s
5 - 12
53
The real and imaginary parts of equivalent impedance are,
'
221
1 '2 22
1
( )
( ) ( 2 )
NL
th
NL BL
RX X
sR RR
X X Xs
5 - 13
2'
21 1 1 1
1 '2 22
1
( )( )( 2 ) ( )
( ) ( 2 )
NL BL NL BL NL
th
NL BL
RX X X X X X X X X
sX X
RX X X
s
5 - 14
The no load and the locked rotor reactance could be calculated by the equation 5-7
and 5-11. S is the slip of the motor and the stator resistance R1 is measured by test
report. Besides these known parameters, either real or imaginary part of equivalent
impedance has two variables: the stator leakage reactance X1 and the rotor resistance
R2’. The value range of X1 is from 0 to XBL, and from 0 to XNL. Due to the reason of
minimum cover (XBL<XNL), the value of X1 is chosen between 0 and XBL. The
determination of rotor resistance R2’ is cooperated with the theory of skin effect.
Based on the equation 5-5, R2’ is calculated by the locked rotor test. As a result, the
skin effect plays a dominant role and makes the rotor resistance R2’ larger than the one
in full load condition. In the other word, the value of R2’ should be between 0 and
(RBL-R1) in full load condition. Therefore, the basic idea of modified calculation
method is to find out new parameters of X1, X2’, Xm and R2’ which could perform the
accurate characteristics of full load condition through adjusting the values of X1 and
R2’ within their intervals. The flowchart of algorithm is shown in Figure 5-20.
54
Figure 5 - 20 Flowchart of modified method
Two variables X1 and R2’ are set to zero in the initialization. The slip of motor is
chosen from the rated slip, for example the rated slip of KM-T7111 motor is 0.76%.
XNL and XBL are set to values which are calculated based on the equation 5-7 and 5-10.
R1 is equal to the measured value of test report. An identified impedance Zth_i is also
arranged in the initial stage before doing calculations. This identified impedance is the
No
No
No
Yes
X1 ≤ XBL
R2’ ≤ RBL- R1
Calculation of Rth, Xth,
ΔRth, ΔXth
If ΔRth and ΔXth
≤ ε (error)
X1(j) = X1
R2’(j) = R2’
ΔRth(j) = ΔRth
ΔXth(j) = ΔXth
j = j + 1
R2’ = R2’ + Step size
End
X1 = X1 + Step size
Initial Values (s, R1, X1, R2’, XBL, XNL, Rth_i, Xth_i, j)
Yes
Yes
55
equivalent impedance of motor under full load condition and can be estimated by the
equations,
_rated
th i
rated
VZ
I
5 - 15
_ 23
ratedth i
rated
PR
I
5 - 16
2 2
_ _ _th i th i th iX Z R
5 - 17
Where,
Vrated: the rated phase voltage;
Irated: the rated phase current;
Prated: the t rated power of motor;
Rth_i: the real part of identified impedance;
Xth_i: the imaginary part of identified impedance;
After the initialization stage, the algorithm starts to execute the main function. If the
stator leakage reactance X1 is smaller than the locked rotor reactance XBL, the
algorithm steps into next judgment which is about the rotor resistance R2’. If the R2’ is
smaller than the difference between the locked rotor resistance RBL and the measured
stator resistance R1, the first task will be calculated. Besides the Rth and the Xth, the
output of the first task also includes the differences between the equivalent impedance
and the identified impedance,
_th th th iR R R 5 - 18
_th th th iX X X 5 - 19
If both differences are smaller than the defined error ε, the stator leakage reactance X1,
the rotor resistance R2’ and two differences ΔRth and ΔXth are recorded into the array.
Then, the calculation returns back to the judgment of rotor resistance R2’ with one step
size increased. The calculation is repeated until the rotor resistance R2’ is larger than
the difference between the locked rotor resistance RBL and the measured stator
resistance R1. Next, the rotor resistance is given to zero and the stator leakage
reactance is increased by one step size to repeat calculations. After the calculation of
two loops is finished, the relationship of ΔRth and ΔXth are recorded into two arrays.
Finally, If the equivalent impedance Rth and Xth calculated based on the parameters of
56
X1 and R2’ have minimum ΔRth and ΔXth, the parameters of X1 and R2’ could be used to
represent the motor under full load running condition. In the other word, the deviation
of parameters due to skin effect could be minimized.
In order to verify the above modified method, the parameters of an Alstom 14MW
induction motor which is used in Capixaba Golfinho FPSO B project are calculated
and compared with the parameters of datasheet. In total, the algorithm finds out
389956 points of ΔRth and ΔXth which fulfill the requirement of defined error ε, for
example the ε is set to 3Ω for this motor. As shown in Figure 5-21, the star trace is the
ΔRth and the dot trace is the ΔXth. If the ΔRth reaches to the lowest ε, the ΔXth has the
largest ε. When the ΔXth has the lowest ε, the ε of ΔRth is located in the middle part.
Consequently, the ΔXth with the lowest deviation ε and the ΔRth with middle level of
deviation ε are chosen to calculate the parameters and the comparison is shown in
Table 5-2.
Figure 5 - 21 Values of ΔRth and ΔXth
Table 5 - 2 Comparison of parameters from different methods
Item R1 X1 Xm R2’ X2’ Trace on Figure
5-22
Method[18] 0.045 1.39 84.51 0.24 1.39 Blue
Modified Method 0.045 2.07 83.83 0.069 0.708 Red
Information of Data
Sheet 0.045 2.32 92.5 0.074 0.745 Black
57
Figure 5 - 22 Comparison of torque, current and speed characteristics with different parameters
shown in Table 5-2
The result of modified method has very closed approach to the information of vendor,
as shown in red trace of Figure 5-22. However, the blue trace which represents the
parameters from the method [23] gives large deviation for the full load condition,
because the rated torque point is moved to the left. The starting part of torque and
current characteristics are not fully represented, nevertheless the specific starting
model is used to fix this problem.
5.5 Evaluation Circuit and Simulation Results of Motor
under Full Load Condition
When the motor is working under full load condition, the voltage, current, power
factor and mechanical speed should be equal to or around the rated value. The
response of modified parameters and the d-q-0 model is evaluated by the same circuit
in Figure 5-3 to check whether the simulation results such as the phase voltage and
the phase current are fulfilling or not. The load is KM-T7111 Main Gas Compressor A
Motor and the parameters are calculated by the modified method. The electrical and
the speed characteristics under full load condition are shown in Figure 5-23 and Table
5-3.
58
Figure 5 - 23 Voltage and current of phase A of KM-T7111 motor under full load condition
Results of simulation have shown that the modified parameters have an acceptable
tolerance to the response of motor test report. The next procedure is similar to the
motor under starting condition. Typical switching operations are performed to check
the response of overvoltage. Figure 5-24 shows two cases with different arcing time
and results. The first case is the successful interruption without reignition in the VCB
and the arcing time is around 450μs. Second case is the successful interruption with
multiple reignitions and the arcing time is around 100μs.
Table 5 - 3 Comparison of electrical characteristics of KM-T7111 motor under full load condition
between simulation and test report
Item Results of
Simulation
Results of
Test Report
Tolerance of
IEC 60092-101
IEC 60034.01-2010
Peak Value of Phase
Voltage (V) 9015.1 9010 +6% to -10%
Peak Value of Phase
Current (A) 878.1 876.4 -
Power Factor under Full
Load Condition 0.9 0.89
Maximum absolute
value: 0.07
Mechanical Speed (rad/s) 187.1 187.07 -20% to +20%
Slip (%) 0.74 0.76 -20% to +20%
59
Figure 5 - 24 Upper left trace: simulated three phase voltage at terminal of load, 450μs arcing time;
Upper right trace: simulated three phase voltage at terminal of load, 100μs arcing time; Lower left
trace: simulated three phase current in VCB, 450μs arcing time; Lower right trace: simulated three
phase current in VCB, 100μs arcing time; (Red trace: phase A; Green trace: phase B; Blue trace:
phase C)
Unlike the switching operation of motor under starting condition, the amplitude and
the rate of rise of TRV is much lower for the motor under full load condition, as
shown in Figure 5-25. The overvoltage due to multiple reignition does not reach to a
serious level and the virtual current chopping is not observed at this moment. This is
because when the motor is running, the back EMF appears on the winding of motor.
Once the motor is disconnected, the rotor cannot stop rotating instantaneously. It
continues generating the back EMF so that the TRV builds up slowly and reduces the
switching overvoltage into insignificant level.
Figure 5 - 25 Left trace: simulated voltage across the contacts of phase B of VCB, 450μs; Right
trace: simulated voltage across the contacts of phase B of VCB, 100μs
60
When the arcing time is decreased to a very short time around 50μs, the virtual
current chopping is observed in VCB and the considerable overvoltage around 33kV
(3.7p.u) occurs at the terminal of motor, as shown in Figure 5-26. This is because that
the VCB does not have enough time to build up high dielectric strength due to short
arcing time. As a result, the multiple reignitions occur in VCB and HF current is
injected into network to cause virtual current chopping.
Figure 5 - 26 Left trace: simulated three phase voltage at terminal of load, 50μs arcing time; Right
trace: simulated three phase current in VCB, 50μs arcing time; (Red trace: phase A; Green trace:
phase B; Blue trace: phase C)
5.6 Modeling of Induction Motor under Light Load
Condition
The modeling of induction motor under light load condition is performed by adjusting
the slip and the load torque in the UM3 model which is proposed in Section 5.4.
Based on the requirements of different driven machines, the phase current of light
load motor is set to 20% rated phase current for a compressor and 15% rated phase
current for a pump.
5.7 Evaluation Circuit and Simulation Results of Motor
under Light Load Condition
The switching operation of motor under light load condition is the most common
situation in a FPSO vessel. Before doing switching operations due to task such as
maintenance, the load of motor should be decreased first. The evaluation circuit which
is proposed in Figure 5-3 is also used in this part to verify the electrical characteristics
and the switching overvoltage of KM-T7111 motor under light load condition.
In this case, the output power of KM-T7111 motor is decreased to 17.65% rated
power to approach the motor that is working under light load condition. The
61
parameters of d-q-0 model are the same with the motor under full load condition
besides the load torque and the slip.
Figure 5 - 27 Left trace: simulated voltage of phase A of KM-T7111 motor under light load
condition; Right trace: simulated current of phase A of KM-T7111 motor under light load
condition
Table 5 - 4 Comparison of electrical characteristics of KM-T7111 motor under light load condition
between simulation and test report
Item Results of
Simulation
Results of
Test Report
Tolerance of
IEC 60092-101
IEC 60034.01-2010
Peak Value of Phase
Voltage (V) 9060.7 8981.5 +6% to -10%
Peak Value of Phase
Current (A) 175.68 177.39 -
Power Factor under Light
Load Condition 0.71 0.65
Maximum absolute
value: 0.07
Mechanical Speed (rad/s) 188.3 188.3 -20% to +20%
Slip (%) 0.11 0.11 -20% to +20%
Output (%) 17.6 15 -
As shown in Figure 5-27 and Table 5-4, the electrical characteristics of model of
KM-T7111 motor has shown sufficient acceptance to the test report. Hence, this
model is used to represent the KM-T7111 motor under light load condition. The
typical switching operation with around 100μs arcing time is shown in Figure 5-28.
After the current is chopped, the TRV appears between two contacts of phase B which
is the first pole to clear. The level of TRV is much lower than the starting and the full
load conditions. The switching overvoltage appears on the phase B with the peak
value around 9.1kV (1.01pu). The multiple reignitions and the virtual current
chopping are not observed in the motor under light load condition.
62
Figure 5 - 28 Upper trace: simulated three phase voltage at terminal of load, 100μs arcing time;
Lower left trace: simulated three phase current in VCB, 100μs arcing time; (Red trace: phase A;
Green trace: phase B; Blue trace: phase C) Lower right trace: simulated voltage across the
contacts of phase B of VCB;
5.8 Location and Protective Effect of Surge Arrester
As has been discussed in the previous sections, the arcing time of VCB plays an
important role in the switching transient overvoltage, especially for the motor under
starting condition. When the motor is under full and the light load condition, the
switching transient overvoltage shows clear low level, except the situation where the
arcing time of VCB in switching of a motor under full load condition is very short. As
mentioned in Chapter 4, a surge arrester is a suitable device to protect the electrical
equipments from the switching transient overvoltage. If the surge arrester is used in
the network, the major issue is the location of surge arrester. Generally speaking, each
surge arrester has a limited protective zone of only a few to up to several tens meters.
According to a general rule of thumb, the distance between the surge arrester and the
equipment to be protected should be as short as possible [27].
In order to verify the electrical response and the protective effect of surge arrester, the
basic insulation level of 11kV motor should be introduced first. The basic insulation
level (BIL) is the level of full wave lightning impulse that a network should withstand
without having any damage or flashover [21]. As the requirement of insulation
coordination, the IEC 60034-15 standard suggests the BIL of 11kV motor to be 49kV
peak value. This voltage is applied between the coil terminals of motor and the earth.
In the simulation, it is equivalent to the phase voltage in the terminals of motor. As
shown in Figure 5-7, the switching overvoltage of KM-T7111 motor under starting
condition with 350μs arcing time is the worst case at this moment, and the peak value
63
of phase voltage reaches to around 56kV (6.2p.u). Such high overvoltage could
seriously damage the main insulation of motor and decrease the lifetime. Therefore,
the surge arrester is connected to the terminal of KM-T7111 motor or the load side of
VCB, and then the limitation effect of overvoltage in above the worst case is checked.
Figure 5-29 shows that the switching overvoltage is clearly mitigated by using surge
arrester in the terminals of load. The peak value of phase voltage is limited to around
26kV (2.9p.u) and it is below the BIL of motor. Moreover, another important factor is
that the surge arrester cannot decrease multiple reignitions in the VCB.
Figure 5 - 29 Simulated three phase voltage in terminal of KM-T7111 motor, 350μs arcing time,
surge arrester is connected at terminal of KM-T7111 motor; (Red trace: phase A; Green trace:
phase B; Blue trace: phase C)
Figure 5-30 shows the situation that the surge arrester is added to the downstream of
VCB. The distance between the surge arrester and the KM-T7111 motor is around
180m. In this case, the switching transient overvoltage is mitigated to around 36kV
(4p.u). Although the peak value of overvoltage is below the BIL of motor, the
performance is not as good as the case where the surge arrester is connected at the
terminals of motor. It is because that the traveling waves reflect in the cable and make
the phase voltage in terminals of motor is higher than the protective level of surge
arrester [27].
Figure 5 - 30 Simulated three phase voltage in terminal of KM-T7111 motor, 350μs arcing time,
surge arrester is connected at load side of VCB; (Red trace: phase A; Green trace: phase B; Blue
trace: phase C)
64
5.9 Conclusions
This chapter has introduced three different models of motor under starting, full load
and light load conditions. The author proposed modified method to estimate the
accurate parameters of motor under full load condition and the calculated parameters
also show the low deviation for motor under light load condition. The main gas
compressor A motor KM-T7111 is chosen as an example to determine the switching
transient overvoltage under starting, full load and light load condition. Consequently,
when the KM-T7111 motor is disconnected during start up, the most serious
overvoltage could reach to a very high level which is above the BIL of motor.
Moreover, the overvoltage above the BIL of motor is not observed for the
disconnection of KM-T7111 motor under full load and light load conditions. Once the
surge arrester is used, the amplitude of overvoltage is mitigated effectively and the
most appropriate location of surge arrester is the terminal of motor.
65
Chapter 6 Analysis of Switching Transient
Overvoltage in the Power System of A FPSO Vessel
6.1 Introduction
The purpose of this study is to check the switching transient overvoltage of different
motors and confirm the protective effect of surge arrester in the 11kV power system
of the selected FPSO vessel. Due to the large numbers of equipments, a simplified
layout of 11kV power system is used, as shown in Figure 6-1. Based on the client
requirements, the configuration of main generators is 3+1. It means that three main
generators work under normal production and one main generator is used as a cold
redundancy. The simplified layout ignores the cold redundancy of main generator.
Some induction motors and transformers which have the rated power around 1MW to
4MW are also ignored from the simplified layout. Originally, the 3×50mm2 three core
cable is used to connect between the VCB and motors such as PM-V2201B,
KM-T7941, PM-T6711 and PM-V2201A. Due to lack of dimensions inside the three
core cable, the three 1×50mm2 single core cables are applied to replace the three core
cable.
Four typical induction motors under starting, full load and light load conditions are
analyzed in this chapter: the 10.2MW main gas compressor A motor KM-T7111
connected by 180m cable to a VCB, the 10.2MW main gas compressor B motor
KM-T7131 connected by 300m cable to a VCB, the 5.5MW water injection pump
motor PM-T2611 connected by 160m cable to a VCB, and the 1.25MW refrigerant
compressor motor KM-T7941 connected by 240m cable to a VCB. Therefore, the
relationship between motors with different rated power and switching transient
overvoltage is recorded. Besides the effect of rated power of motor, the influence of
length of cable is also considered by doing comparable analysis between the main gas
compressor A motor and the main gas compressor B motor. Finally, the surge arrester
is applied in the terminals of four motors under starting condition. The switching
transient overvoltage of them is compared with the situation where motors do not
have surge arresters. The modeling of different equipments is the same to the previous
chapters. The above information of switching transient overvoltage is calculated by a
statistical way.
66
Figure 6 - 1 Simplified layout of 11kV power system of FPSO vessel
67
6.2 Results of Switching Operation of Motor under
Starting Condition
The interruption of motor under starting condition is usually from the inappropriate
settings of protection relays and a long duration of starting current during
commissioning of a FPSO vessel. When it occurs, as explained in previous chapters,
the multiple reignitions could lead to a serious overvoltage in the windings of motor.
As shown from Figure 6-2 to 6-5, four typical loads are disconnected by the VCB and
the relationship between cumulative probability and maximum phase voltage is
recorded.
Figure 6 - 2 Cumulative probability of maximum phase voltage at the terminal of KM-T7111
motor under starting condition without a surge arrester connected
Figure 6 - 3 Cumulative probability of maximum phase voltage at the terminal of KM-T7131
motor under starting condition without a surge arrester connected
0
0.2
0.4
0.6
0.8
1
1.2
0 2 4 6 8 10
Cu
mu
lati
ve P
rob
abili
ty
Maximum Phase Voltage (p.u)
0
0.2
0.4
0.6
0.8
1
1.2
0 1 2 3 4 5 6 7 8 9
Cu
mu
lati
ve P
rob
abili
ty
Maximum Phase Voltage (p.u)
68
Figure 6 - 4 Cumulative probability of maximum phase voltage at the terminal of PM-T2611
motor under starting condition without a surge arrester connected
Figure 6 - 5 Cumulative probability of maximum phase voltage at the terminal of KM-T7941
motor under starting condition without a surge arrester connected
Figure 6-2 and Figure 6-3 show the influence of length of cable, the maximum phase
voltage of KM-T7131 motor is slightly lower than the KM-T7111 motor, because the
KM-T7131 motor is connected to the VCB by a longer cable than KM-T 7111 motor.
If the length of cable is increased, the lumped capacitance of cable is also increased. It
means that the surge impedance and the rate of rise of transient voltage are reduced.
As a result, the KM-T7131 motor experiences lower overvoltage. Besides the effect of
length of cable, the above figures have shown another interesting phenomenon that
the small machine produces a higher switching overvoltage than the large one.
Generally speaking, it is a common behavior for switching operation of induction
motor [28]. Two basic reasons can be used to explain this phenomenon. First of all, if
the VCB does not experience reignitions, the switching overvoltage is caused by the
current chopping, for example the overvoltage of KM-T7941 is around 1.4p.u and the
overvoltage of KM-T7111 is around 1.1p.u respectively. The equation 2-3 shows that
this overvoltage is proportional to the level of current chopping and surge impedance
of load. Since the trend of surge impedance is inverse proportional to the rated
0
0.2
0.4
0.6
0.8
1
1.2
0 2 4 6 8 10
Cu
mu
lati
ve P
rob
abili
ty
Maximum Phase Voltage (p.u)
0
0.2
0.4
0.6
0.8
1
1.2
0 2 4 6 8 10 12
Cu
mu
lati
ve P
rob
abili
ty
Switching Overvoltage (p.u)
69
apparent power of motor, as shown in Figure 6-6 [29], the small motor produces the
higher switching overvoltage which comes from the current chopping.
Figure 6 - 6 Surge impedance per phase of synchronous machines; (The impedance of induction
motors can be assumed to be less than 80% of the value in this Figure)
Second, if the VCB has multiple reignitions, the overvoltage may reach to 10.4p.u for
the KM-T7941 motor, whereas the other three motors keep the overvoltage below
9p.u. As has been mentioned in Chapter 2, the level of current chopping of modern
VCB is around 6A. When the motor has lower rated power, the phase current of motor
is also reduced, such as the RMS value of phase current of KM-T7111 is 627A and
the RMS value of phase current of KM-T2611 is 346A. As a result, if the VCB
performs switching operation for motors, the phase current of small machine reaches
to the level of current chopping before the large machine, as shown in Figure 6-7. A
time interval T appears between the chopping time of small machine and the chopping
time of large machine. Therefore, when the small machine has multiple reignitions,
the time interval T could make reignition more serious and get more possibilities to
have the virtual current chopping. This explanation is approved by the simulation as
shown in Table 6-1.
Figure 6 - 7 Comparison of current chopping for small and large motors; (Green trace: phase
current of small motor; Red trace: phase current of large motor)
70
Table 6 - 1 Comparison of reignitions characteristics for KM-T7111, PM-T2611 and KM-T7941
motors
Machine
Type
Times of Switching
Operations with
Reignition
Times of Virtual
Cuurent Chopping
Total Times of
Simulation in One
Electrical Period
KM-T7111 12 10 50
PM-T2611 18 18 50
KM-T7941 17 16 50
With a surge arrester connected at the terminal of motor, the switching overvoltage is
obviously decreased, as shown in Figure 6-8 to 6-11. As has been mentioned in
Chapter 5, the surge arrester successfully mitigates the maximum overvoltage to
relatively low level. However the surge arrester does not limit the cumulative
probability for overvoltage due to multiple reignitions
Figure 6 - 8 Cumulative probability of maximum phase voltage at the terminal of KM-T7111
motor under starting condition with a surge arrester connected
Figure 6 - 9 Cumulative probability of maximum phase voltage at the terminal of KM-T7131
motor under starting condition with a surge arrester connected
0
0.2
0.4
0.6
0.8
1
1.2
0 0.5 1 1.5 2 2.5 3 3.5
Cu
mu
lati
ve P
rob
abili
ty
Switching Overvoltage (p.u)
0
0.2
0.4
0.6
0.8
1
1.2
0 0.5 1 1.5 2 2.5 3 3.5
Cu
mu
lati
ve P
rob
abili
ty
Maximum Phase Voltage (p.u)
71
Figure 6 - 10 Cumulative probability of maximum phase voltage at the terminal of PM-T2611
motor under starting condition with a surge arrester connected
Figure 6 - 11 Cumulative probability of maximum phase voltage at the terminal of KM-T7941
motor under starting condition with a surge arrester connected
6.3 Results of Switching Operation of Motor under Full
Load Condition
The switching operation of motor under full load condition can be resulted from the
process trip such as the differences of temperature and pressure in the processing
equipment. As shown from Figure 6-12 to 6-15, four different motors have remote
cumulative probabilities to acquire the phase overvoltage above 3.5p.u, for example,
the phase overvoltages of both the KM-T7111 and the PM-T2611 motors have 4%
probability to get the values between 4 to 4.5p.u, and the KM-T7131 motor has 2%
probability to have the phase overvoltage around 3.9p.u. On the other hand, the small
power motor KM-T7941 has slightly higher cumulative probability, where it is 10%
to have the phase overvoltage between 3.5 and 4.5p.u. This slightly increase can be
explained by the theory that is shown in Figure 6-7 as well. Although the phase
0
0.2
0.4
0.6
0.8
1
1.2
0 0.5 1 1.5 2 2.5 3 3.5
Cu
mu
lati
ve P
rob
abili
ty
Maximum Phase Voltage (p.u)
0
0.2
0.4
0.6
0.8
1
1.2
0 0.5 1 1.5 2 2.5 3 3.5
Cu
mu
lati
ve
Maximum Phase Voltage (p.u)
72
overvoltage due to the multiple reignitions and the virtual current chopping may reach
to the largest value around 4.2p.u (38kV) for the KM-T7111 motor and the
KM-T7941 motor, such overvoltage is still below the BIL of 11kV motor. Therefore,
the surge arrester is not really necessary in this case.
Figure 6 - 12 Cumulative probability of maximum phase voltage at the terminal of KM-T7111
motor under full load condition without a surge arrester connected
Figure 6 - 13 Cumulative probability of maximum phase voltage at terminal of KM-T7131 motor
under full load condition without surge arrester connected
0
0.2
0.4
0.6
0.8
1
1.2
0 0.5 1 1.5 2 2.5 3 3.5 4 4.5
Cu
mu
lati
ve P
rob
abili
ty
Maximum Phase Voltage (p.u)
0
0.2
0.4
0.6
0.8
1
1.2
0 0.5 1 1.5 2 2.5 3 3.5 4 4.5
Cu
mu
lati
ve P
rob
abili
ty
Maximum Phase Voltage (p.u)
73
Figure 6 - 14 Cumulative probability of maximum phase voltage at the terminal of PM-T2611
motor under full load condition without a surge arrester connected
Figure 6 - 15 Cumulative probability of maximum phase voltage at the terminal of KM-T7941
motor under full load condition without a surge arrester connected
6.4 Results of Switching Operation of Motor under Light
Load Condition
As has been mentioned in the beginning of this chapter, the switching operation of
motor under light load condition is the common situation which disconnects the load
from the network in a FPSO vessel. The relationship of cumulative probability and the
maximum phase voltage is shown from Figure 6-16 to 6-19. These simulations of four
motors illustrate that the serious multiple reignitions, the voltage escalations and the
virtual current chopping are not observed in switching operation of motor under light
load condition. The reason is that the rate of rise and the amplitude of TRV in the
VCB are really low, as shown in Figure 5-28. The main switching overvoltage is
caused by the current chopping. The below figures show that the motor with lower
0
0.2
0.4
0.6
0.8
1
1.2
0 0.5 1 1.5 2 2.5 3 3.5 4 4.5
Cu
mu
lati
ve P
rob
abili
ty
Maximum Phase Voltage (p.u)
0
0.2
0.4
0.6
0.8
1
1.2
0 0.5 1 1.5 2 2.5 3 3.5 4 4.5
Cu
mu
lati
ve P
rob
abili
ty
Maximum Phase Voltage (p.u)
74
rated power, the higher phase overvoltage. That is because the motor with low rated
power reaches to the high surge impedance, as shown in Figure 6-6. Consequently, the
small motor such as KM-T7941 experiences slightly higher overvoltage than the other
motors. When compared with the previous two cases of motor working conditions, the
motor under light load condition has relatively low overvoltage and such overvoltage
is far below the BIL of 11kV motor. As a result, the surge arrester has little effect on
this overvoltage and it is unnecessary to add the surge arrester into the network to
protect the motor under light load condition from overvoltage.
Figure 6 - 16 Cumulative probability of maximum phase voltage at the terminal of KM-T7111
motor under light load condition without a surge arrester connected
Figure 6 - 17 Cumulative probability of maximum phase voltage at the terminal of KM-T7131
motor under light load condition without a surge arrester connected
0
0.2
0.4
0.6
0.8
1
1.2
1.005 1.01 1.015 1.02 1.025 1.03 1.035 1.04 1.045
Cu
mu
lati
ve P
rob
abili
ty
Maximum Phase Voltage (p.u)
0
0.2
0.4
0.6
0.8
1
1.2
1 1.05 1.1 1.15 1.2
Cu
mu
lati
ve P
rob
abili
ty
Maximum Phase Voltage (p.u)
75
Figure 6 - 18 Cumulative probability of maximum phase voltage at the terminal of PM-T2611
motor under light load condition without a surge arrester connected
Figure 6 - 19 Cumulative probability of maximum phase voltage at the terminal of KM-T7941
motor under light load condition without a surge arrester connected
0
0.2
0.4
0.6
0.8
1
1.2
1.08 1.1 1.12 1.14 1.16 1.18 1.2
Cu
mu
lati
ve P
rob
abili
ty
Maximum Phase Voltage (p.u)
0
0.05
0.1
0.15
0.2
0.25
0.3
0.35
0.4
0.45
1.15 1.2 1.25 1.3 1.35 1.4 1.45
Cu
mu
lati
ve P
rob
abili
ty
Maximum Phase Voltage (p.u)
76
Chapter 7 Conclusions and Future Work
7.1 Conclusions
In this study, the switching transient overvoltage in the 11kV power system of
Floating Production, Storage and Offloading (FPSO) Vessel is analyzed. The whole
work is starting from the modeling of different equipments in the power system of the
selected FPSO vessel. After the modeling of different equipments is finished, a
simplified layout of 11kV power system in the selected FPSO vessel is described and
used to check the switching transient overvoltage for four typical induction motors.
The first important phenomenon is that the motor under starting condition could have
the most serious overvoltage, which comes from the multiple reignitions and the
virtual current chopping of VCB. In this case, the peak value of overvoltage could be
above the basic insulation level (BIL) of 11kV motor. To be compared with the motor
under starting condition, the motor under full load condition has very remote
probability to have large overvoltage due to the multiple reignitions and the virtual
current chopping. However, the level of overvoltage does not exceed the BIL of 11kV
motor. If the motor is under light load condition, the maximum overvoltage is very
moderate and far below the BIL of 11kV motor.
The second important phenomenon is that the motor with lower rated power produces
the higher overvoltage for the starting, the full load and the light load working
conditions. As has been discussed in the Chapter 6, this phenomenon is clearly
observed in the PM-T2611 and the KM-T7941 motors.
The third important phenomenon is that the VCB has enough capacity to interrupt the
current of four different motors in the starting, the full load and the light load working
conditions. If the current of first pole to clear of VCB is not successfully interrupted
and the power frequency current is conducting through the arc, the successful
interruption will be achieved in the next zero point of power frequency current.
Finally, the surge arrester is connected to the terminals of four motors to protect the
motors from the switching transient overvoltage, especially the motor under starting
condition. The amplitude of overvoltage is limited to the protective level of surge
arrester. The reduced level of overvoltage is below the BIL of 11kV motor.
The recommendations of whole work could be concluded: First of all, the surge
arrester is recommended for all induction motors in the 11kV power system of the
selected FPSO vessel. The most suitable location of surge arrest is the terminal of
motor. Second, due to the most serious overvoltage occurs in the starting condition of
motor, the commissioning of the selected FPSO vessel should be followed the strict
77
principles, especially the impropriate settings and the wire mistakes of protection
relays should be avoided.
7.2 Future Work
In the future, the work should focus on the accumulation of more accurate data of
equipments such as the parameters of VCB, the geometric information of cable and
the parameters of motor. Moreover, the parameters of motor are based on the power
frequency in this study. The frequency of switching transient analysis is ranging from
power frequency to several MHz. Therefore, the high frequency characteristics of
motor should be measured by an impedance analyzer and combined with the power
frequency model to represent the motor with a wide range of frequency. In this study,
the different short circuit faults cannot be truly represented due to the limitation of
generator model. In the future work, if the generator is modeled as accurate as
possible, the switching transient overvoltage during different short circuit faults can
be analyzed.
78
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80
Appendix
Arcing Time
(μs) A (kV/ms) D (kA/μs)
Peak Voltage across the Contacts of
VCB (kV)
0-150 20 0.5 91.59
0-150 20 0.55 91.46
0-150 20 0.6 92.98
0-150 20 0.65 91.55
0-150 20 0.7 91.5
0-150 25 0.5 91.27
0-150 25 0.55 91.34
0-150 25 0.6 92.28
0-150 25 0.65 91.15
0-150 25 0.7 91
0-150 30 0.5 91.13
0-150 30 0.55 91.19
0-150 30 0.6 91.76
0-150 30 0.65 90.98
0-150 30 0.7 90.98
150-300 20 0.5 91.38
150-300 20 0.55 91.43
150-300 20 0.6 92.75
150-300 20 0.65 91.39
150-300 20 0.7 91.29
150-300 25 0.5 91.27
150-300 25 0.55 91.18
150-300 25 0.6 91.33
150-300 25 0.65 90.85
150-300 25 0.7 91.13
150-300 30 0.5 87.67
150-300 30 0.55 90.65
150-300 30 0.6 90.65
150-300 30 0.65 89.16
150-300 30 0.7 89.23
300-600 20 0.5 88.66
300-600 20 0.55 90.89
300-600 20 0.6 91.43
300-600 20 0.65 90.63
300-600 20 0.7 84.75
300-600 25 0.5 83.51
81
300-600 25 0.55 83.11
300-600 25 0.6 87.51
300-600 25 0.65 81.13
300-600 25 0.7 87.26
300-600 30 0.5 70.28
300-600 30 0.55 54.07
300-600 30 0.6 70.1
300-600 30 0.65 61.17
300-600 30 0.7 53.46
