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

E-mail of author:

hans.schellekens@schneider-electric.com

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