50 Years of TMF Contacts Design · PDF file50 Years of TMF Contacts Design Considerations. ......
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Technical collection
H. Schellekens
2008 - Conferences publications
50 Years of TMF Contacts
Design Considerations
ISBN ________________ XXIIIrd Int. Symp. on Discharges and Electrical Insulation in Vacuum-Bucharest-2008
50 Years of TMF Contacts Design Considerations
H. Schellekens
Schneider Electric, Medium Voltage Development, Usine 38V, ZAC Champ Saint-Ange, Varces, France
Abstract- The historical evolution of TMF or RMF contacts
is discussed. Today the very basic guideline as prescribed
by Harold. N. Schneider, 50 years ago, still prevails. This
allows for simple, compact and robust contact designs.
Better understanding of the thermodynamics of the arc has
helped to design contacts coping with short circuit currents
of up to 100 kA. Contact material improvement has helped
to boost up the short circuit interruption performance,
which today is equivalent to the AMF contact structures.
Progress in the theoretical understanding of the arcing
process, and the increasing base of experimental data
permit to give guidelines to design a correct functioning
TMF contact.
I. INTRODUCTION
A. History of TMF contacts
Commercialization of sealed vacuum interrupters
started in the 1950’s. The first VI’s were mainly used as
load break switches for capacitor bank switching.
General Electric Company started development work on
VI’s in 1952. Within GEC, Harold N. Schneider was in
1958 the first to propose a compact contact design, the
“spiral” contact, fig. 1a, to move the constricted arc [1].
At about the same time in 1962 Anthony A. Lake and
Michael P. Reece [2] proposed a “cup” shaped contact
which creates a sufficiently strong magnetic field to
move the arc on the ridge, fig.1b. Here, the many slots
that prevail in the cup extend into the ridge which forms
the contact surface. As a variant the cup contact can
be topped with a contact disc. The constricted arc,
although rotating, projects vapour and droplets onto the
surrounding walls. This is attributed by Richard L.
Hundstad of Westinghouse Electric Company in 1972 to
the radial component of the force on the arc [3]. He,
therefore, proposed a “folded petal” contact system that
reduces the radial force on the arc by bending the
contact arms in the 3rd dimension, fig. 1c.
B. Relation between TMF, RMF and AMF
Often TMF contacts are also called “radial magnetic
field” or RMF contacts, as the magnetic field
component that makes the arc rotate points in the radial
direction. A TMF contact system is composed of a set of
2 non-identical contacts.
AMF contacts are designed to create a mainly axial
magnetic field between the contacts. This field tends to
keep the arc in a diffuse and stable arcing mode. An
AMF contact system is composed of a set of 2 identical
contacts, which resemble the contacts of fig. 1b and 1c.
C. TMF and Contact Materials
The development of vacuum interrupters in general is
strongly related to the development of materials; the
TMF contacts are no exception to this rule. Three
different time periods can be distinguished. In the
beginning the main contact material was pure copper or
copper-bismuth alloy, an optimized alloy to minimize
the force necessary to break contact welds. These
materials suffer from arc erosion, which create large
droplets that influence negatively the dielectric
characteristics of the VI. The development of sintered
copper-chromium (CuCr) alloy (in the 1970’s) was a
great improvement [4]. With this material the generation
of large droplets was reduced as this material has a grain
size in the order of 100 µm; so, only micro droplets are
emitted by the arc. The grainy structure of the material,
though, makes it more susceptible to gas adsorption,
which interferes with current interruption in other ways.
A CuCr contact fabricated by arc melting in a low
pressure environment [5] created a virtually gas free
material, which has been the standard for interruption
performance since the mid 1980’s. Due to patent
protection, this material was not available for the
Fig. 1a. Spiral contact by Harold
Schneider
Fig. 1b. Cup contact by Anthony A. Lake
and Michael P. Reece
Fig. 1c. Folded petal contact by
Richard Hundstad
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large community. Only recently, a vacuum casting
technique [6] has reduced the cost of the fabrication of
gas-free CuCr and mass production has popularized its
application.
D. Renewed interest for TMF
At the end of the 1970’s AMF contacts became
popular. Higher current ratings and higher voltage
ratings and longer electrical switching live could be
attained with the AMF technology. The success of AMF
attracted more research, which led to a reasonably good
understanding of the arc behavior [7,8,9]. As a
consequence, the application of TMF contacts regressed
relatively. Yet development on TMF contacts continued.
For low voltage applications the interruption current
was raised to 100kA [10]. By applying the development
methodology, which was so successful for the AMF
contacts [11], a breakthrough was obtained in the design
of TMF contacts. As a consequence, near to identical
interruption performance was obtained for TMF and
AMF contacts. Recently also the first generator breaker
based on TMF contacts with rated current of 75 kA at
15 kV has been commercialized. [12]
II. ARC MODELS
Models of arc motion are used to relate contact design
to interruption performance. To describe the arc motion
across the contact surface, an obvious approach would
be to use electro-mechanics: the Lorentz-force on a
“solid” conductor to explain the arc motion. Yet, the arc
can not be treated as a solid conductor, as the plasma is
composed of electrons, ions and gas. Movement of these
species within the plasma is far more complex and the
displacement of the arc becomes less evident. Below
different arc model categories are presented that deal
with arc motion.
Fig. 2. Solid arc (blue) model of TMF contact geometry with 4 curved petals.
Fig. 3. Direction of Lorentz force for a “solid”arc body on 4 curved petal contact for any possible position.
A. Arc models based on force balance equations
3D magnetic field modeling tools, make it possible
to calculate the force on a “solid” body for complex
electrode geometries. Fig. 2 shows a typical 3D finite
element model of the arc between curved contacts. This
approach allows adapting the contact geometry to, for
instance, maximize the force on the arc in the rotating
direction, fig. 3 [11].
B. Models based on energy balance equations
A sophisticated approach mixes a force balance model
of the arc with an energy balance model for the contacts.
The arc is mainly characterized by the physical
evaporation process at the anode and cathode foot point.
The ionized species are confined to the arc by the strong
self-magnetic field. The neutral particles are lost from
the arc due to diffusion. Evaporation has to compensate
for this loss (energy balance). The neutral mass loss
represents a loss of momentum (force balance).
This leads to a direct relation between arc speed and
contact design [13]. Scaling laws have been derived that
relate these parameters to interruption capability [14].
C. 2D and 3D arc models
A 2D plasma dynamic model combines full plasma
modeling with an exact description of the plasma
contact interaction including thermal energy balance in
the contacts. This model gives an impressively realistic
image of how the arc could move fig. 4. [15]
In the phenomenological arc model Nike [9,16], the
arc is governed by a minimum energy criterion and
adapts its size to the prevailing magnetic field. In a
transverse magnetic field, the arc reduces its
cross-section and moves in the amperian direction. Fig.
5 shows the arc for three sequential time steps in a
folded-petal contact system. The time necessary to
move between the positions depends on arc voltage and
inductivity of the contact system [17].
Fig. 4. Arc motion predicted by plasma dynamic model
Fig. 5. Arc motion predicted by 3D model Nike.
D. Comparison between the models
In table I a comparison is given of the above presented
models in terms of their predictive powers. All models
are capable to predict the arc speed. Only models that
include a description of the arc plasma behaviour are
capable to predict the arc diameter. Energy balance
equations relate the arc speed to the material properties
of the contact material. Only extended, 2D/3D physical
models can explain the arc motion from one contact arm
to the other. With respect to scaling laws, force
calculations contain not sufficient elements. The 2D/3D
physical models mask the essential relationships in their
numerical output but multiple simulations allow
revealing scaling laws. Albeit 2D/3D physical models
predict arc movement, in [15] arc motion is conditioned
by the ability of the arc to generate a hot electron
emissive contact surface up front, whereas in [16] the
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main current displaces, due to electrodynamics and arc
energy balance, which results in an apparent arc motion. TABLE I : Comparison of predictive powers of the models. Type of Model
Force Calculation
Force and Energy Balance
2D and 3D
models
Arc Diameter No Yes Yes
Arc Speed Yes Yes Yes
Time to Cross a contact slot
No No Yes
Scaling laws No Yes Yes
III. DESIGN CONSIDERATIONS
The design considerations given below are based on
experimental results, theoretical models and numerical
simulations.
A. Contact shape
The contact system has to be assessed [11] on the
effectiveness of its current interruption capability. This
is promoted by its ability to move the arc and to force it
to rotate. It is evident that a contact fails if the arc does
not run at all, or suddenly stops and stays immobile. To
achieve this the main design rule, as formulated by
Harold Schneider [1], has to be followed: “The disc
(contact) should be slotted from the outer peripher y inward,
and the slot configuration should be such that the current
path between the conductor and an arc terminal loca ted at
substantially any angular point on the outer periph eral
region has a net component extending generally
tangentially with respect to the periphery in the v icinity of
the arc”. All three TMF contact types of fig. 1 obey to
this rule.
Fig. 6. Anode arc root diameter variation with arc current for a Siemens type cup contact.
B. Width of the contact arm
The width of the arc on a contrate cup contact. is
given in fig. 6 for the anode side of the arc. The width of
the arc at the cathode side is square root of 2 larger. This
sets a lower limit for the width of the contact arms [13].
C. Number of contact slots
The time to make one revolution is composed of the
time for the arc to cross the total circumferential
distance with the proper arc speed and of the time to
cross each of the contact slots. The proper arc speed
depends on parameters like momentary arc current, arc
diameter, local magnetic field and contact distance. The
local magnetic field strength depends on the form of the
contacts. In some designs, like the folded petal, this
field is maximized. The proper arc speed is between 100
to 1000 m/s for a vacuum arc. Fig. 7 shows the
evolution of time depending on the angular position of
the arc; moments of rapid displacement across the
contact arms are interrupted by the larger time it takes to
cross the contact slot [17]. The number of slots has a
direct influence on the time to make a complete
revolution.
Fig. 7. Evolution of time depending on the angular arc position. The proper arc speed is 1030m/s and the apparent speed is 290 m/s.
D. Contact diameter
Interruption performance is intimately related to
contact diameter, D, and contact distance, d. The
theoretical relation of eq. 1 is derived by [14]. The
constant CRMF depends on contact geometry and
material.
7.0max d
DCI RMF= (1)
E. Width of slots
The slots in the contact are an essential feature to
create their form. Yet, the width of the slots has no
influence on the behaviour of the TMF arc or on its
ability to cross the slot. The latter solely depends on the
length of the contact arm. Due to contact erosion the
slot fills up progressively with melt. Therefore an
appropriate choice of slot width should be made
depending on the expected lifetime of the contacts.
F. Contact thickness
Contact thickness has a direct influence on the
magnetic field in the arc. For a stationary or stalled arc
the field reduces with increased contact thickness; so, to
set an arc in motion thin contacts should be favored. For
the running arc current flow is close to the surface and
the contact thickness has nearly no influence on its
speed. Fig. 8 shows the current distribution on a rail
contact an arc speed of 50 m/s.
G. Contact distance
The contact distance has a direct influence on the arc
speed. On the contact arm the constricted arc will
always run. Yet, a minimum contact gap is necessary to
make the arc move across the slot. With increasing
contact distance the proper speed of the arc as well as
the number of revolutions will increase. The rotation
frequency depends on the contact distance and on the
number of contact arms.
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Fig. 8. Current distribution in a rail contact for arc speed of 50 m/s. Current flow close to contact surface.
H. Contact heating
As the constricted arc is running on the edge of the
contact, the contact heating is very time and place
dependent. The notion of the relative heating time
defined as the ratio of heating time and cooling time
becomes useful in order to compare different contact
designs. A heating period is followed by a cooling
period. The heating time is the ratio of arc diameter and
proper arc speed; typically for a current of 50 kA the arc
has 8 mm diameter and the proper arc speed of 200 m/s
this time is 40µs. As the arc stalls at the end of the
contact arm, the heating time at this position is of the
order of 100µs, fig.7. The cooling time for both
positions is determined by the time of one complete
revolution 550µs. For this example the relative heating
time at the end of the contact arm is 100/550= 18%.
Reduction of the relative heating time is an efficient
way to increase the interruption performance.
I. Shield heating
As the arc heats the contact surface, the contact
surface temperature will increase and can surpass easily
the melting temperature. Due to the high magnetic
pressure in the arc the liquid will be ejected from the arc
root. This process of erosion cooling is one of the
advantages for TMF contacts with respect to AMF
contacts. The disadvantage is the deposition of the
erosion products on the shields. During the arcing
process the shields are under a constant spray of liquid
material. This material will cool upon impact with the
shield. Depending on the thermal accommodation
between shield and contact material, the shield will heat
up beyond the melting temperature and start to erode.
For a correct performance of the VI the shield thickness
and composition are important parameters.
J. Electrical endurance
Contact erosion increases exponentially with current
on TMF contacts. Therefore, the nominal short circuit
current rating depends on the final application. Fig. 9
compares the electrical endurance for AMF and TMF
contacts for a 20kA rated vacuum interrupter.
K. Opening speed
A minimum contact distance, ~ 4 mm, is necessary
before the arc crosses a slot. So, the opening speed of
the circuit breaker should be designed such that this
minimum opening is attained in a short time to facilitate
the interruption process.
10
100
1000
10000
0.1 1 10 100
Current [kA]
Ope
ratio
ns
AMF TMF
Fig. 9. Electrical endurance of AMF and TMF contacts.
IV. CONCLUSIONS
Improved contact materials and a better understanding
of the multi-physics governing the TMF help in the
conception of compact and efficient vacuum interrupters.
In the continuing process to reduce the cost of vacuum
interrupters the TMF technology is a challenging
alternative to the AMF technology for applications
where electrical endurance is of second concern.
ACKNOWLEDGMENT
The author thanks the students Julien Fontchastagner, David Gonin and Guillaume Gomez for their valuable contributions.
REFERENCES 1. H. N. Schneider, US. Patent 2,949,520, filed 1958 2. A. A. Lake and M. P. Reece, UK Patent 997,384, filed 1963. 3. R. L. Hundstad, US. Patent 3,845,262 , filed 1972 4. A. A. Robinson, British Patent 1,194,674, 1970 5. R Muller, “Arc-Melted CuCr Alloys as Contact Materials for
Vacuum Interrupters, Siemens Forsch. Entwicklungsber. 17-33, 105-111, 1988
6. B. Miao, Z. Yan, Z. Yumin, L. Guoxun, D. Shenlin, H. Yu , “Two new Cu-Cr alloy contact materials”, XIXth ISDEIV, p. 729-732, Xi’an, China
7. S. Yanabu, S. Souma, T. Tamagawa, S. Yamashita, T. Tsutsumi, “Vacuum arcs under axial magnetic field and its interrupting ability”, Proc. IEE, 126, p.313-320, 1979
8. H. Fink, M. Heimbach, W. Shang, “A new contact design based on a quadrupolar axial magnetic field and its characteristics”, European Trans. on Electrical Power Engineering, 2000
9. H. Schellekens, "Arc Behaviour in Axial Magnetic Field Vacuum Interrupters Equipped With an External Coil", Proc. XVIIIth ISDEIV 2, 514-517, Eindhoven, 1998
10. H. Fink and R. Renz, “Future trends in vacuum technology applications”, Proc. XXth ISDEIV, 25-29, Tours, 2002
11. E. Dullni, E. Schade, W. Shang, “Vacuum arcs driven by cross-magnetic fields (RMF)”, ISDEIV pp. 60-66, Tours, 2002
12. Eaton / Cutler-Hammer Product Bulletin BR01301001E “Medium Voltage Generator Vacuum Circuit Breakers”, 2006
13. W. Haas and W. Hartmann, “Investigation of arc roots of constricted high current vacuum arcs”, IEEE transactions on plasma science, Vol. PS-27, No. 4, pp 954-60, August 1999
14. R. Renz, “ Thermodynamic Models for TMF and AMF vacuum arcs”, XXIIth ISDEIV, pp. 443-7, Matsue, 2006
15. T. Delachaux, O. Fritz, D. Gentsch, E. Schade and D.L. Shmelev, „Numerical simulation of a moving high-current vacuum arc driven by a transverse magnetic field”, XXIIth ISDEIV, pp. 273-7, Matsue, 2006
16. H. Schellekens, “Phenomenological arc model NIKE”, private communication Current zero club meeting 2003, Grenoble.
17. J. Fontchastagner, O. Chadebec, H. Schellekens, G. Meunier, and V. Mazauric “Coupling of an Electrical Arc Model With FEM for Vacuum Interrupter Designs”, IEEE TRANSACTIONS ON MAGNETICS, VOL. 41, NO. 5, MAY 2005
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