Production and performance of metal foil brushes

Wear, 73 (1981) 261 - 282 0 Elsevier Sequoia S.A., Lausanne

PRODUCTION

P. B. HANEY*, Department (Received

261 -

Printed

AND PERFORMANCE

D. KUHLMANN-WILSDORF

ofMaterialsScience,

University

in The Netherlands

OF METAL

FOIL BRUSHES

and H. G. F. WILSDORF of Virginia,

Charlottesville,

VA 22901

(U.S.A.)

June 19,1981)

Summary Electrical brushes which were made of silver, copper and aluminum foils of 12.5 and 25 m thickness and were composed of 15 - 195 individual foils, were tested in purified argon on a polished copper rotor at a speed of 13 m s-l. Brush pressures varied between 3.1 X lo3 and 2.8 X lo4 N me2 and current densities were up to about 700 A cmm2 (about 4500 A in2). The observed dependence of the voltage drop across the brushes as a function of the current densities agreed closely with Helm’s contact theory as applied to foil brushes. The film resistivities were found to be near ur = lo-l2 52 m2 for copper and silver and to be about 3 X lo-l2 fi m2 for aluminum. The projected performance of foil brushes based on these results is very favorable and the future commercial use of foil brushes appears to be possible. The total loss, electrical and mechanical, through the brushes is independent of current density if the brush pressure is chosen to minimize the total loss. If so, the loss depends only on the brush speed, the hardness of the softer of the two materials involved (i.e. of foil and rotor or slip ring), the coefficient of friction and the film resistivity. Microscopic surface examinations of rotor and brushes show that the brush surface is smoothed through running the brush, whereas the rotor remains almost unaf,fected or is mildly roughened, as long as no arcing takes place. Arcing causes considerable surface roughening on both the brush and the rotor surface and debris is thus deposited on the rotor; this can score the brushes. Further experiments are required to determine the rate of brush wear.

1. Introduction The electrical resistance of electric brushes is composed of three parts: R,, the ohmic resistance of the brush body; Rc, the so-called constriction *Present

address:

U.S. Navy Surface

Weapons

Laboratory,

Dahlgren,

VA, U.S.A.

262

resistance; RF, the film resistance, The constriction resistance is present because contact on an atomistic scale takes place over only a fraction of the overall geometrical area of the contact; this fraction consists of islands of atomistic contact which are called ‘%-spots”. The resistance for current conduction through a circular a-spot of radius a may be written [l] R, =

Pb

OF + __

(1)

=a2

4a

where pb and pr are the electrical resistivities of the materials of which the brush and the opposing contact (e.g. the “rotor”) respectively are made and uE’is the resistance of unit area of the film, if any, separating the brush from the opposing contact surface. Evidently, the first term in eqn. (1) is the constriction resistance and the second term the film resistance of the a-spot. For n similar a-spots in the interface, the total brush resistance is therefore R, = R,/n + R,, i.e. Pb +

R,=R,+RF+Ro=

___ 4an

Pr

‘7F

+ __ ma2

+Ro

When a current I is conducted through a brush moving at a velocity u relative to the opposing (rotor) surface, against which it is pressed with the brush force P, the total power loss W is the sum of electrical and mechanical loss, i.e. W = R,12

+ pvP

(3)

where p is the coefficient of friction. From eqn. (3) the total power loss per unit current, measured in watts per ampere, i.e. in volts, is found to be W -=LL,=LE+LM= I

Pb i 4an

+ - (JF +Ro I+?;’ ima2 !

(4)

As shown by Holm [l] , for non-circular a-spots the constriction resistance could be as much as 30% smaller than that given in eqn. (4). Conversely, when the two sides of an a-spot are separated by a film of high resistivity, the constriction resistance may be increased by as much as the factor 4/n. These corrections oppose each other and may certainly be neglected compared with the other uncertainties in the equation, e.g. the uncertainties in aF,nanda. If we assume that the contacts are essentially “clean” and that the surface film has similar properties everywhere, the total conducting contact area A, = n=a2 is related to the hardness H of the softer of the two contacting materials by Ab = ma2 = P/j = P/(H

(5a)

when plastic deformation occurs at a-spots, where p is the average pressure at the contact spot, i.e. 5 = ,$H. By definition, therefore, t < 1.

263

Usually, for stationary contacts at least, 0.1 2 .$ 5 0.3 is assumed (ref. 1, p. 30). Indeed, theoretically l can vary between zero and unity, as the pressure below an asperity gradually rises from the first infinitesimally light touch of a material with negligible yield strength in the initial state to the hardness H, which is reached asymptotically when the depth of indentation approaches and then exceeds 3% of the radius of curvature of the asperity. From that point onwards, @ = H where H is equal to about thrice the saturation flow stress for uniaxial tension or compression in the fully workhardened state. We believe that after “running-in” a monolithic or foil brush, or indeed any contacting objects for which the number of a-spots is not very large, t will generally be close to unity. The reason is that successive contact events during the running-in period generate the fully work-hardened state at the surface (according to the local conditions of temperature, speed of deformation and other boundary conditions) no matter what the initial hardness may have been and, further, that the wear process will always leave either or both of the two contacting surfaces in a roughened state. Therefore, in the great majority of cases, in the fully run-in condition and with a small to medium number of a-spots such that the average a-spot is plastically deformed, t: = 1 and a = (P/nnH)“2 For elastic deformation a = l.l(Pr,

/nE)1’3

(5b) the average a-spot radius is [ 21 (6)

where r, is the average radius of surface curvature and the r, value of the rougher of the two surfaces is taken in case of doubt. Altogether, therefore, for a uniform film the total power loss for plastically deformed a-spots may be written as ,,LT = RJ

+ 7’

(7) and for elastic a-spots as

It is well known, and can be seen clearly from eqns. (7) and (8), that it is highly advantageous if the number n of a-spots is large. Alas, in monolithic brushes, i.e. brushes made out of one solid piece of material, n may be as small as unity [ 3,4] , and it is rarely as large as 20 [ 51. Also, it would be desirable to attain loads low enough to enable us to work in the elastic mode and thus to benefit from lower losses according to eqn. (8) as compared with eqn. (7). For given values of P and H this typically requires very large r~

264

values. Further, due to statistical changes in the size and number of a-spots the brush resistance fluctuates, the more so the lower n is, and the corresponding electrical noise is emitted. Yet, with monolithic brushes rz cannot be controlled. Therefore, under ordinary circumstances n < 20, the constriction resistance contributes at least a si~~ic~t part of the total brush resistance, a-spots are plastically deformed, losses are relatively high and brush noise is severe. In addition to the discussed desirability of large values of IZ in order to reduce brush noise as well as LT, which is given by eqn. (4) in conjunction with eqns. (7) and (8), attempts are made to increase r, , i.e. to make the contacting surfaces as smooth as possible. With respect to the brush force P, it may be adjusted to minimize LT for any given value of the current; this minimum generally occurs when L, and LE are of similar magnitude. However, in order to reduce wear it is usually desirable to keep P below that level, especially at high currents when the value of P determined to minimize L, may be uncomfortably high. Conversely, at low currents and high speeds the load P which minimizes LT for monolithic brushes is often too low to maintain continuous contact between the brush and the rotor (or slip ring or commutator), in which case the brush “bounces”; this causes arcing, much noise and excess wear. Furthermore, in practice it is rarely if ever possible to adjust P in accordance with the moment~y current. The ubiquitous carbon brushes, which often incorporate metal powders to enhance conductivity, have clearly served as an excellent compromise among the above conflicting motivations or requirements. This is so although they are monolithic and therefore have n 5 20, primarily because carbon brushes have a low coefficient of friction, wear well and are mechanically rather soft so as to run without bouncing at relatively low brush loads. Also, in the past, brush noise has been tolerated without much question. Recently, especially in Europe, brush noise has been recognized as objectionable. At least as importantly, plans for rail launchers [6] and for efficient homopolar motors, e.g. for use in ship propulsion, high power pulsing generators or power storage devices in connection with the future use of fusion energy [ 71, call for brushes with reduced losses, at high current densities and speeds, which are at levels well below those attainable with the best available monolithic brushes. In this connection metal fiber brushes have been made and tested, with very promising results [S, 91. The important advantages of metal fiber brushes are as follows. (i) The number of a-spots can be made so large for these brushes that the constriction resistance Rc becomes entirely negligible. (ii) Metal fiber brushes can be made to be very compliant mechanically; this permits the use of very low values of the brush force without bouncing, even at very high speeds, and has the effect that the mechanical loss is greatly reduced. (iii) With low loads used and a very large number of a-spots, the average a-spot is elastically stressed and the correspondingly lowered electrical loss is obtained.

266

f iv) Under favorable conditions the total area of annular zones about the a-spots with gap widths small enough to permit tunneling becomes significant and even dominant. This causes a further reduction in the electric loss below that given in eqn. (8). Altogether, therefore, metal fiber brushes are intrinsically capable of yielding very low losses even under high speeds and current densities.

2. Theoretical

considerations

for metal foil brushes

A further expansion of brush technology beyond monolithic and fiber brushes is possible by employing stacks of parallel metal foils, similar to the leaves in a book, whose edges make the contact with the rotor. Such brushes will be named “foil brushes”. Two of the four benefits of fiber brushes enumerated above, namely a large number of a-spots, which practically eliminates the constriction resistance, and high compliance so as to permit low pressures and thus low mechanical losses, can also be attained readily with foil brushes. Since the prospect of attaining the other two benefits, i.e. significant current tunneling and elastic behavior of a-spots at low loads, are not likely to be achieved with foil brushes, we can ask why foil brushes should be considered at all in competition with metal fiber brushes. The reasons include the following. (1) Foil brushes can be made more cheaply and in a wider range of materials choices than metal fiber brushes. (2) Foil brushes are expected to have increased wear resistance because a higher fraction of the geometrical contact area can be made of the brush metal: the optimum value of this number (the “packing fraction” f) appears to lie at about f = 15% and almost certainly at no more than 20% for fiber brushes, while for foil brushes the useful range of packing fractions might extend beyond f = 60%. (3) Foil brushes exhibit greater ruggedness in use, e.g. relative insensitivity against accidental mechanical and electrical overload, than metal fiber brushes. (4) It is possible that foil brushes can be made indefinitely long (e.g. by feeding stacked foils through rollers) and that any desired life times can thus be achieved, whereas fiber brushes wear out after their fiber length has been reduced by some fixed fraction of their original length; this reduction is typically a few millimeters. The theoretical considerations presented in Section 1 apply to foil brushes, as they do to fiber brushes, and may in that case be simplified as follows: for metal foil brushes R. is typically negligible, and a-spots will be deformed plastically so that eqn. (7) applies but with the film resistance dominant and the number of a-spots of no further interest provided that n is sufficiently large, e.g. larger than 30. For typical foil brushes, therefore, the loss is given by

266

Furthermore, because through variations in the foil thickness the brushes can be given almost any desired value of mechanical compliance, only the wear resistance and the values of (rF and, on occasion, p are of interest in choosing the material for the construction of foil brushes. From previous measurement on metal fiber brushes f9] , c+ is known to have values from a minimum of about 5 X 10-~13 .Q m2 upwards; this lower limit is realized with gold in a protective atmosphere. For very high performance, protective atmospheres are, so far at least, necessary for all metal choices tried except perhaps gold. Further, contrary to superficial impressions it is preferable to choose foil materials which have a large hardness value so as to retain a low wear rate, since H in eqns. (5), (7) and (9) refers to the softer of the two materials and the rotor will become the controlling factor in this respect for hard brush materials. For control of P it is necessary to rely on a suitable choice of foil thickness, foil length and packing fraction. However, wear resistance is not readily predicted and it is not always found that harder materials wear better than softer ones although the general trend is definitely in that direction.

3. Examples

of foil brushes made and exp~r~en~l

results

Foil brushes were made of aluminum, copper and silver with the dimensions listed in Table 1. Figure 1 shows the macroscopic appearance of the brushes. They may be made by three different methods and variants thereof. (i) In the first method, the desired number of similar rectangular foils are stacked and bound between two similarly rectangular rigid pieces of metal. These serve as a clamp, similar to the hard covers of a book. They are tightened together by means of two screws piercing the foils (Fig. 1). Next, the edges of the clamp and foils are trimmed and shaped to the desired form; special care is taken to shape and to smooth the intended face of the brush, i.e. the foil edges which will later contact the rotor. Thereafter the hard covers are trimmed to expose the desired length of foil. Brushes made in this manner have a nominal packing fraction off = 100%. (ii) In order to make brushes with variable packing fractions, we can proceed as in method (i) but with the difference that foils, of the desired thickness and number to yield the intended packing fraction, of a different metal are interspersed. After clamping, shaping and removing part of the clamp to expose the desired length of the foil stack as in method (i), the filler foils are etched away. For example, when interspersing copper and aluminum foils, the aluminum may be removed by etching in NaOH and the . in both cases the unetched foil is not damaged. copper by etching in HNOs _

26’7

TABLE 1 Experimental and calculated data for foil brushes tested on a polished copper rotor in an argon atmosphere Brush and test symbol Foil material Foil thickness t (Gm) Number N of foils Number of a-spots (n = 3N assumed) Nominal packing fraction f Attack angle # Brush width w (in (cm)) Brush area An = wlvtlf 03s 4 (cm2) Foil length L (cm)

A

B

C

D

E

Ag 25 15

Ag 12.5 135

&

cu

Al

25 50

45

405

25 195 585

25 15 46

100% 0” 1 (2.54) 0.095

33.3% 0”

33.3% 0”

0.5 (1.24) 0.64

100% 15” trailing 1 (2.54) 1.28

1.0 3.45 x 10-s

3.8 3.45 x 10-a

1.3 3.6

1 x 10-u

1 x 10-12

1 x 10-12

3 x 10-u

5 x 10s 13

5 x 108

5.5 x 108

3 x 10s

0.27 (27) 4.3 x 103

13 1.10 (112) 8.6 x 103

13 0.14 (14) 3.1 x 103

13 0.21(21) 2.2 x 104

1.73 x 10-a

7.05 x 10-4

3.85 x 1O-3

12

4

34

2

5.20 x 1O-3 (average value) 10

1.7

0.57

1.2

0.55

1.8

1.85 x10-3

1.85 x 10-S

4.5 x 10-4

3.9 x 1O-3

4.3 x 10-S

9.8 x 10-S

3.3 x

10-5

1.3 x

8.2 x 1O-5

1.2 x 10-4

1.95

1.9 x

10-3

4.6 x 1O-4

4.0 x 10-a

4.4 x 10-4

33.4

1.52 156

0.97 16.1

1.18 31

52

122

36.8

323

0.118

0.22

0.12

0.32

0.7 Combined resistivities 3.45 x 10-a Ph + Pr of materials (a m) Assumed film resistivity 1 x 10-12 OF (a

0.2) 0.45

x

1O-8

(a)

Measured mechanical loss 2LMI W) Coefficient of friction /i = &It& Calculated film resistance SF

100% 0” 1 (2.54) 0.095 0.7 4.7 x 10-s

m2)

Average hardness N (N rnm2) 5 x lo8 Test velocity u (m s-l) 13 Brush load P (N (gf)) 0.27 (27) Brush pressure p = P/AB 2.8 x 104 (N mF2) Measured brush resistance 1.51 x 10-S &B

160

= a~HIf’(fi;2)

Calculated constriction resistance &o = (n112/4) X @b

10-5

X

+ P,)(Hwt/3PfAd”2

63

Calculated brush resistance

x

10e3

cRB=oRF+&(Q)

0.77 Inferred current I,,,u, = 62.9 (J.IvP/,R~)~~~ at minimum loss (A) Current density Jmih = 662 In&An at minimum loss (A cmm2) Minimum loss (inferred) 0.190 I&B/&B

fLThnin

= 218~

0.92

dmin

w

(iii) In order to be able to vary not only the packing fraction but also the foil thickness to below that of the initial foils, and in general to make mechanically superior brushes, we proceed as for method (ii), with the difference that the assembly is inserted into rectangular tubing of the same material as the filler foils (or of any other suitable material provided that it is differ-

Fig. 2. Foil brush running on the rotor of a bag-testing [lo] but including a modified brush holder [ 111.

apparatus as described previously

entially etchable) and drawn or rolled down, in a Turk’s head for example, to attain the desirable dimensions. The desired length of foil is exposed by etching away the tubing and filler foils. Figure 2 shows a foil brush running on a polished copper rotor: measurements of LM, LE and L, obtained in purified argon from such a system, using a brush testing apparatus of the type described previously [lo] but employing an advanced modification of the brush holder [ 111, are shown in Figs. 3 - 7. Tests were conducted at u = 13 m s-l within a wide range of brush loads and current densities. Also shown in Figs, 3 - 7 are comparable measurements made with two gold fiber brushes; this allows the foil brush performance to be compared with that of very good exper~ental metal fiber brushes that have yielded excellent results [S] , i.e. brushes made of gold fibers 20 and 22 pm thick at packing fractions of 10.5% and 15.5% run under a brush pressure of 5 ozf in2 (2160 N mm2). The data in Figs. 3 - 7 are presented in terms of brush pressure rather than brush load and of current density rather than current because these are technologic~ly the more me~~gful parameters. However, the dependence of the brush performance on the brush area AB (i.e. the overall geometrical area of contact between brush and rotor) is different for foil and fiber brushes. Specifically, for a given value of P the number of a-spots as well as the electrical brush resistance is (almost) independent of AB for monolithic brushes, i.e. for n 5 20, whereas for fiber brushes with a given fiber diameter d and packing fraction f the number of a-spots is proportional to AB, i.e. [8,9] n = 4fAB/nd2. Finally, for foil brushes, fully plastic a-spots can be assumed so that, for a given foil thickness t and packing fraction f, we have for foil brushes of width w (see Fig. 2) n ;+ ff~A~/~~ with cr assumed to be about 3 whereas a: * 1 for fiber brushes which are running [ 8,9] . Thus, without neglecting constriction resistance, and following eqns. (2), (5b) and (7), the resistance of foil brushes is given by

269

0.6

-

r-l

0

( 200

100

0

I 300

1000

I 400

2000

I

I 500

I 700

600

3000

4OCKJ

I

4500

I

A/cm’ A/in’

Fig. 3. Graphical representation of the results of brush test A. (A silver foil brush in purified argon at 13 m s-l on a polished copper rotor was used. For comparative purposes the total loss L, previously obtained with gold fiber brushes at the same speed (ref. 8, part 2, Fig. 3) is included; this curve was found for both a gold brush with fiber diameter d = 22 pm and packing fraction f = 15.5% and a gold fiber brush with d = 20 pm and f = 10.5% at a pressure p = 5 ozf in -2 (2160 N me2) in both cases. Since the mechanical loss for these brushes (with a brush area of about 0.7 cm2) was too small to be measured, the coefficient of friction was assumed to be /1= 2. This is a definite overestimate, and it is quite probable that the mechanical loss was only one-quarter as large, with a correspondingly reduced value of LT compared with the curve presented. V, is the melting voltage, i.e. the electrical voltage drop which causes local melting at the a-spots on the foils.)

”

VI

I

I

I

07-

0.6

-___---

0 0

50 250

100 500

I

A”

150

750

I

IOCG

I

FIBER

-

BRUSH

200 1250

I

1500

I

1750

I

2ooO

,

A/in’

0

I

250

1

500

I

750

1

loo0

I

1250

I

15CO

I

1750

I

2ooO

Fig. 4. Graphical representation

of the results of brush test B. (Details as for Fig. 3.)

Fig. 5. Graphical representation

of the results of brush test C. (Details as for Fig. 3.)

I

A/in’

270

0 0 /I

20 100

40 200

300 s

60 400 i

60 500

loo 600 z

110 7?9* i

Fig. 6. Graphical representation

of the results of brush test D. (Details as for Fig. 3.)

Fig. 7. Graphical representation

of the results of brush test E. (Details as for Fig. 3.)

rdcm’

giving a loss

It follows that in terms of brush pressure p = P/AB and current density larger monolithic brushes compare less favorably with foil brushes of the same size, and these with fiber brushes, than do smaller monolithic brushes. However, for the areas concerned, i.e. AB of the order of 0.5 cm2, the effect is not dramatic and the relative merits of the two types of brush are adequately illustrated in Figs. 3 - 7. To permit a more detailed evaluation, all experimental data are compiled in Table 1. For clarity, some measured data have been indicated by the subscript m to distinguish them from computed values which are indicated by the subscript c. J = I/A,,

4. Evaluation

of data

4.1. Film resistiuity and first-order test of the theory Previous studies indicate that the film resistances for gold, silver and copper fiber brushes which are run in purified argon atmospheres are comparable and are given by GF = IO-l2 S2 m2 within a factor of perhaps 2 or 3

271

[S, 9,121, Correspondingly, in the evaluation of the foil brush performance data in Table 1, i&?= lo-i2 a m2 is assumed for copper and silver foils. However, the film resistivity for aluminum is doubtless much higher. In Table 1, or = 3 X lo-l2 G m2 has therefore been assumed for aluminum. Also given in Table 1 are average values for H according to Holm [ 131. With these assumed input parameters, the me~urements may be analyzed theoretic~ly. The number of a-spots per foil may be assumed as CY= 3. If so, the foil brushes had 45 or more a-spots and the constriction resistance is indeed negligible for all brushes, at least in a first approximation; this can be seen from the calculated constriction & and brush resistances cR, listed in Table 1. The values of mRB/cRB, i.e. the ratio of the measured to the calculated values using R, according to eqn. (10) with the assumed uF values, are surprisingly close to unity; the average of the four mRB/cRB values which refer to silver and copper foil brushes in Table 1 is 1.05 without discoverable dependence on brush pressure. It is remarkable that the scatter of the mR,/, RB values in Table 1 suggests a rather more uniform value of uF than typically observed with fiber brushes [ 93. We suggest that this indicates the presence of a cleaning effect through the locally strong plastic deformation at a-spots in foil brushes which is absent in fiber brushes where a-spots are deformed elastically. Therefore, since the measurements for fiber and foil brushes of silver and copper were made on the same kind of rotor and in the same testing atmosphere, it may be confidently assumed that the lower limits of the fiF values for the two types of brushes were the same, The uF vaiues employed in Table 1 were the lower limits of eF deduced from a theoretical inte~re~tion of the resistance of fiber brushes assuming elastic behavior of the a-spots [ 8,9] , i.e. using eqn. (6) which does not involve the indentation hardness H nor any possible ambiguity with respect to the parameter g. By ~on~ast, in the computation of RB in Table 1 fully plastic a spots were assumed, i.e. eqn. (5a) with $ = 1 was used. The good agreement between the computed and the measured values of the foil brush resistances may be taken as strong support for both of the theories involved, ie. the theory previously derived for fiber brushes and the present theory of foil brushes. Specifically, the choice of $ = 1 is vindicated because of the absence of any indication that the brush resistance drops with decreasing average pressure on the a-spots. Theoretically, c must decrease with decreasing values of P/N. Therefore, if t is not close to unity for all the measurements, which means that the assumption of fully plastic a spots is unjustified, the values of mRB/,RB should increase with increasing P/N. Correspondingly, mRs/,RB for tests A - D should increase in the following order: B < D < C < A. In fact the order is A < B < D < C. Moreover, the extreme deviations are only by a factor of 1.52:0.77 = 2; this is certainly incompatible with the frequently used value [ = 0.3 We should not be too confident of the agreement between the measurements and the theory embodied in eqn. (9) if they concerned only one or

272

two foil brushes. As it is, the good agreement between theory and experiment for two different foil thicknesses and a wide range of brush areas, number of foils, loads and brush pressures imparts much confidence in the theory, especially when the background of the already published data gained with fiber brushes is considered. It is shown below that further theoretical analysis provides additional support for the basic theory. The only surprise in the data concerns the coefficient of friction. This is much more dependent on testing conditions and, apparently, especially on brush pressure, than had been expected on the basis of earlier experience with fiber brushes. There are indications of two regimes, one at low P/A,H values, withy = 0.5, and one at high pressures with 1-1= 1.5. The lower regime is that with which all previous work with fiber brushes is concerned [ 8, 91 while the higher regime presumably represents adhesive friction. This same type of transition is widely known from the literature on friction and wear (e.g. refs. 14 - 16). (See note added in proof.) High pressures and clean testing conditions favor the adhesive wear regime with its undesired increase in the coefficient of friction and typical increase in wear rates. For practical purposes it is advisable to avoid this regime. At the same time, if we consider the smaller scatter in film resistivities found in foil brushes as compared with fiber brushes, the following connection becomes persuasive: high pressures and clean conditions tend to reduce the thickness of the surface film, thereby reducing the film resistance but also tending to increase wear and the coefficient of friction. Unlubricated metal brushes of any type should therefore be operated in an intermediate regime in which the films are very thin but coherent so that the film resistance as well as the coefficient of friction are low. However, that regime is inevitably somewhat unstable. Consequently, even minor changes in brush pressure and/or surface coverage of the rotor could displace the running conditions either into the regime of too high film resistivities (which is occasionally apparently encountered with fiber brushes) or adhesive behavior with a too high coefficient of friction and wear. However, the connection between the coefficient of friction and the film resistivity has been found not to be close [ 121 ; this is in agreement with theoretical considerations [ 161. In this much it is also understandable that the data of ,,,RBIcRB in Table 1 are not clearly correlated with the respective 1-1values.

4.2. Microscopic studies of brush surface and wear Much more work will be needed in order to evaluate the wear behavior of metal foil brushes. While the testing equipment permits the brush length to be monitored on a continuing basis, it is impossible to distinguish wear from plastic deformation, which is fairly rapid in the early stages of testing. Thus much longer testing periods are needed than have so far been employed. In addition, it appears, in agreement with theoretical expectation, that wear rates depend on the regime of pressure (as indicated above) and also on current density.

273

It is clear that some wear takes place even in the low friction regime at steady current conduction (Figs. 8,9, 11 - 14) and that much more wear ‘takes place with arcing (Figs. 10 and 15). This can be seen from the wear tracks on the copper rotor which exhibit grooving and discoloration and is shown in Figs. 8 - 10. Similarly, micrographs taken of brush surfaces before and after running show signs of wear (Figs. 11 - 14).

Fig. 8. Surface of a copper

rotor after a 270 mm run under a current I = 50 A (brush

Fig. 9. Micrographs of replicas obtained from a copper rotor: (a) as polished with 600 grit emery paper at 13 m s-l ; (b) after a 270 min run under a current I = 200 A (brush test B). (Magnifications, 64x.)

Fig. 10. As Fig. 9(b) but after arcing under the cathodic brush at a current density of 388 A cme2 (2500 A ine2). (Magnification, 64x.)

Shaping of the brush surfaces by the use of 600 grit emery paper gives them a microscopically strongly abraded surface as their initial condition (Figs. 11 - 14). Interestingly, the effect of running the foil brushes under a current load on a copper rotor is to smooth microscopically the brush surfaces. This is evident on all four pairs of micrographs shown in Figs. 11 - 14. Perhaps this is not surprising in view of the fact that increased wear is observed when current is flowing as compared with the case of zero current and that the local current density must generally be very much increased at

(a)

(b)

Fig. 11. Micrograph of the running surface of a silver foil brush (brush test A) after shaping and polishing with 600 grit emery paper (a) and after running on a poiished copper rotor at maximum current density for the test (b). (The pieces of foil missing from the foil edges in a vertical track to the left of the center should be noted,)

275

i4

(b)

Fig. 12. As Fig. 11 but for brush test B: (a) using 600 grit emery paper; (b) using brush test B.

(a)

(b)

Fig. 13. As Fig. 11 but fir brush test D: (a) using 600 grit emery paper; (b) using brush test D.

276

(a)

(b)

Fig. 14. As Fig. 11 but for brush test E: (a) using 600 grit emery paper;(b) test E. (The fine pitting of the surface after running should be noted.)

using brush

asperities. Additionally, it appears that pieces can be torn away from foil edges, as shown in Fig. 11. However, this type of damage might have occurred previously owing to abrading and polishing during the initial brush preparation. The surface of the aluminum foil brush (Fig. 14) exhibits the same smoothing after running under current as the copper and silver foil brushes but the smoothing of the aluminum foil brush is accompanied by a superimposed fine pitting. Pitting may be the result of microscopic events of electrical breakdown in the aluminum oxide layer. However, arcing strongly increases surface damage (Figs. 10 and 15). 4.3. Projected performance capabilities of foil brushes Evidently, the electrical resistance of foil brushes is essentially that of the film resistance. Neither constriction resistance nor ohmic resistance make any significant contribution. Combined with the ease of control of their mechanical compliance, foil brushes can therefore be superior to monolithic brushes, especially at low and intermediate current densities, not only with respect to total losses but even more so with respect to brush noise. This twin advantage of foil brushes over monolithic brushes becomes still more pronounced at higher speeds. For example, at current densities of about 50 A cm-* and less, the total loss LT for a monolithic SG-142 brush (Stackpole, 75 wt.% Ag in graphite) run at optimum load increases by a factor of about 1.6 when the speed is raised from 13 to 26 m s-l [9] while, as shown below, for foil brushes LT ideally rises in proportion to ul/*, i.e. by the factor 1.41 from 13 to 26 m s-r.

277

For the lowest possible loss, regardless of wear rate, it is necessary to work at or near the minimum of LT. The resulting minimum loss (LT)min, realized at the current Imin, is found by differentiating eqn. (11): (LT)min = 2RaImjn = 2(pUPRn)1’2

(12)

which gives, with the use of RB = a,H/P, I min

1’2

=

g

i

1

/Au

P

%-

(13)

UFH

Further, in actual applications, brushes are designed to perform best at some specific current density Jo. It is clearly desirable to let Jo coincide with the minimum value of L,. Therefore, disregarding possible problems of wear, ideal working conditions are achieved if

where p. = PO/A, is the brush pressure chosen to place the minimum of L,(I) at the desired current density Jo.

(4

value

(b)

Fig. 15. Surfaces of anodic (a) and cathodic (b) silver foil brushes after brush test B: the current was raised until arcing began under the cathodic brush at a current density of 388 A cmm2 (2500 A inP2). Comparison of the appearance of the two brush faces reveals clearly the strong roughening of the cathodic brush. This is in marked contrast with the smoothing effect observed after running brushes without arcing (e.g. compare the anodic brush and Figs. 11 - 14). Corresponding roughening was caused on the rotor (Fig. 10). Some debris on the rotor caused by the arcing evidently caused the prominent wear groove seen running across both brushes in the middle of the figure.

278

Practically speaking, in the construction of a foil brush the desired value of Jo will generally be known and the brush load will be adapted accordingly; this will yield p. from eqn. (14a): at;H

lf2

( i __

PO=

JO

I.iv /

( 14b)

The equivalent loss per ampere conducted at po, i.e. (LT)min, is then found from eqns. (12) and (14):

PO

(15)

From the second part of eqn. (15), once the material parameters or, I and H are known the lowest value of the loss per ampere depends only on the velocity, not on the current density Jo. However, the wattage developed

at the brush-rotor interface, first considered in eqn. (3), does of course increase with Jo. The minimum power density at the interface is W -

=

(LT)minJO = ~(o,~uH)~“JO

(16)

AB

Generally there will be some upper limit of WEAN beyond which the system will not perform. It is liable to be that limit which at the same time is the upper limit for foil brush performance for short term use, Long term performance, however, is limited by brush wear; this typically will not allow p. to be raised beyond some specific range at a given current density and will thereby limit achievable current densities. In these considerations the possible failure of the brush material as such, e.g. through oxidation, embrittlement, softening or melting, must be taken into account. Holm 1173 has listed the voltage drops at which various metals soften and melt. There exists an almost exact one-to-one correspondence between the voltage drop across a contact and the temperature at the interface. However, this voltage drop is not our LT but, basically, LE = RBI, albeit somewhat modified by the value of LM [ 121. The appropriate values of V,, the voltage drop at melting for stationary contacts as given by Holm [ 171, are indicated in Figs. 3 - 7. It is conceivable that the current densities of foil brushes could be raised almost to the point LE = V,, after which the brushes would be destroyed by local melting and the fusing of foils. It is interesting that the behavior of the brushes was found to be so very closely ohmic, with the exception of the aluminum brush, and that no change in brush resistance can be detected at the softening voltages Vs. These voltages are given by Holm [ 171 as Vs = 0.09 V, V, = 0.12 V and Vs = 0.1 V for silver, copper and aluminum respectively. It is thought that this occurs because the time intervals during which the a-spots, and thus the locally high temperature, persist at any one point of the surfaces are too short to allow signific~t annealing. For J < Jmin in Figs. 3 - 7 the value of Lr (and of LM by itself) can substantially exceed V, without any damage to the brushes. This is because

279

ZM is far less effective in raising the a-spot temperatures than Lx. The reason why energy input via ZE, i.e. electrical contact resistance, is far more effective in raising the inter-facial temperature than energy input via ZM, i.e. friction, is probably that the heat evolution due to contact resistance is concentrated very closely at the interface while the evolution of frictional heat is more widely spread out through the subsurface region [ 16, 181. At the current density of Jmin, therefore, for which the electrical loss is equal to the mechanical losses, and for higher current densities, the local a-spot temperature will be close to that which would be deduced from ZE alone. However, since the melting voltage for copper, the most commonly used rotor or slip ring material, is 0.43 V it is doubtful whether any value of (ZT)min above 0.43 V can be used safely, and certainly ZT = 0.86 V would be an upper limit for the use of any foil brushes on a copper substrate for J > Jmin, irrespective of the foil material. For silver and aluminum the melting voltages are given [ 171 as 0.37 V and 0.3 V respectively, i.e. they are lower than that for copper. For current densities above Imin, the total loss Zr for silver and aluminum foil brushes can thus certainly not exceed 0.74 V and 0.6 V respectively and for more safety they should be less than 0.37Vand 0.3 Vrespectively. If we keep these limits in mind, the data of Table 2 may be considered. These are correlated values of Jo, po, (ZT)min and W/AB for velocities between TABLE 2 v (m s-l)

JO (A cmP2)

po (gf cmP2)

PO (N me2)

(LT)min

5 5 5 13 13 13 25 25 25 50 50 50 100 100 100

20 150 1000 20 150 1000 20 150 1000 20 150 1000 20 150 1000

21.3 160 1070 13.3 99.6 663 9.6 71.8 479 6.8 50.8 338 4.8 35.9 240

2.10 1.57 1.05 1.30 9.76 6.50 9.38 7.04 4.69 6.63 4.98 3.32 4.69 3.52 2.35

0.105 0.105 0.105 0,169 0.169 0.169 0.234 0.234 0.234 0.332 0.332 0.332 0.469 0.469 0.469

x x x x x x x x x x x x x x x

lo3 lo4 lo5 lo3 lo3 lo4 lo2 lo3 lo4 lo2 lo3 lo4 lo2 lo3 lo4

WI

WIAB

P

cme2)

2.1 15.7 100.5 3.4 26.4 169 4.7 35.1 234 6.6 49.8 332 9.2 70.4 469

Theoretically forecast performance of foil brushes in case it is desired to operate the brush at the current density Jo such that this current density coincides with the loss minimum (LT)min at the desired operating velocity v. In order to let JO be at the loss minimum, a brush pressure PO = (u~H/~u)~~~J~ must be applied. In this table /J = 1, H = 5.5 x 10s N rn-’ (as for copper) and (JF = 1 X lo- l2 a m have been assumed. Under those operating conditions the minimum loss (L,&n and the total loss density W/AB are as given in the table. It should be noted that, contrary to intuitive expectation, the loss at any one velocity is independent of the current density. However, the loss density rises linearly with the current density, as does pg.

280

5 and 100 m s-’ and current densities Jo up to 1000 A cm-’ computed from eqns. (12) - (161, i.e. assuming that both Rs and the coefficient of friction are independent of the velocity. The values of OF, N and I-(used in Table 2 are OF = 10-l’ Q m, Ii = 5.5 X 10’ N rnVV2and p = 1, i.e. they are compatible with those in Table 1. For comparative purposes, and to check the data in Table 1, the values of are included in Table 1 as calculated from the first Irnin, Jmin and (-?&fmin part of eqn. (13) using the measured data, Le. not including any assumed values. The agreement between these calculated values of Jmin and (L/r)min in Table 1 with the curves in Figs. 3 - 7 is partly automatic. Namely, by measurement Ra as well as 1-1were found to be independent of current density, within the experimental error, except with the aluminum foil in which RB mildly decreases with current density, as shown in Fig. 7 by the cu~ature of the LE line. More significant is the pairwise agreement of the (LT)min values between brushes A and C on the one hand and brushes B and D on the other hand. This is in agreement with theory. Specifically, according to eqn. (15), for the same values of or, p , N and u the minimum loss per ampere, i.e. (LT)min, should be the same for all foil brushes (to the extent that R, can be neglected) independent of P, p, AB and Jmin. Now, it must be admitted that in tests A and C the coefficients of friction are not quite alike, and some erratic differences [9] in + may be expected. However, (LT)min is similar in the two tests, as expected from theory. This is even more remarkable because the angle Q of attack differed from brush A to C, and the areas AB of brushes A and C were substantially different, as were brush load and brush pressure; these differences yielded the correspondingly different values of Imin and J,i,. The same phenomenon may be noticed by comparing brushes B and D; these have decidedly similar values of (Lr)min, which is expected theoretically in view of their similar values of OF, fi and H despite considerable differences in AB and p. Inciden~lly, it is not believed that the coefficient of friction was directly influenced by f in such a way that the lower f values in tests B and D caused lower coefficients of friction, which might superficially be thought from Table 1. Instead of this, a lower packing fraction means greater pliability of the brushes; this prompts the application of a lower brush pressure p which in turn causes the lower JJ values (as explained above). It is very probable that it is advantageous to use packing fractions below loo%, even if J should be large. The data in Table 1 appear to indicate that it is better to use, for example, f 5 0.5, together with a shorter free foil length, than f = 100% with longer foils in order to attain the same brush compliance, because this will permit a larger number of a spots per foil (leading to lower mRdcRB values in terms of Table 1) as well as allowing better cooling. The slight deviation from ohmic behavior in the ~uminum brush is believed to be due to surface film breakdown at higher voltages. This interpretation supports the conclusion, given above, that the fine surface pitting visible in Fig. 14 is due to numerous events of electrical breakdown in the oxide.

281

5. Conclusions

and outlook

The overall excellent agreement between theory and experiment inspires much confidence that the behavior of foil brushes is theoretically understood. In this theory the main basic features are that ohmic and constriction resistances are negligible and that the a-spots are fully plastically deformed. By combining the present results with the prior observation (derived from metal fiber brushes made of gold, silver and copper which were tested in a protective atmosphere) that neither RB nor /1 depend significantly on velocity, it follows that the minimum total loss per ampere, although occurring at a current density which rises with applied brush pressure, is independent of current density, brush area and brush pressure. Given fixed values of the material constants, i.e. hardness, film resistivity and coefficient of friction, this minimum loss is proportional to the square root of the velocity (eqn. (15)). The forecast performance of metal foil brushes is inferior to that of the best fiber brushes but potentially superior in principle to monolithic brushes. In practice, realizing this promise depends on finding foil materials which combine low electrical resistivity and low film resistivity with good wear resistance. The coefficient of friction, which also has a significant influence on brush performance, is probably beyond manipulation [ 161 except in that it is desirable to work in the non-adhesive regime in which the coefficient is about 0.5 rather than in the regime of adhesive wear where it is about three times larger. Preliminary indications are that avoiding very high packing fractions will be helpful. The theoretical superiority of foil brushes compared with monolithic brushes derives from three major sources. Firstly, there is no reason why the film resistance for the same brush load of monolithic brushes should be lower than that of foil brushes incorporating the same metal, but monolithic brushes have, in addition, a considerable ohmic and constriction resistance. Secondly, monolithic brushes can be run at optimal pressure only if this pressure is sufficiently high to prevent bouncing, a requirement which is increasingly troublesome at high speeds. This restriction causes mechanical losses that are too high and means that monolithic brushes are very inferior to foil brushes at low current densities, no matter what speed is used. Thirdly, monolithic brushes emit much electrical noise, whereas foil brushes are very “quiet”. Furthermore, the opportunities that may arise from using essentially endless long foil brushes by advancing stacks of foils between, for example, spring-loaded parallel rollers and thereby effectively eliminating the wear problem could be most attractive. This aspect deserves much further thought and experimentation. It is a considerable drawback that the excellent performance of the foil brushes described in the present paper depends on the maintenance of a protective atmosphere. Indeed contamination of the atmosphere by oxygen, sulphur and/or other species is detrimental. Thus the brush resistance for copper and silver brushes is lower in purified argon than in an ordinary argon

282

atmosphere and the resistance of such brushes increases significantly with time, apparently because an oxide or sulfide layer is gradually formed. This effect is noticeable but not pronounced for silver foil brushes. These presumably react in a similar way as Ag-graphite SG 142 brushes. However, during about 4.5 h of running time copper foil brushes showed an increase in brush resistance of 50% in purified argon and of 60% in ordinary argon. Similar effects have been observed with silver and copper fiber brushes. The described increase in metal brush resistance is thus not peculiar to the form of the contacts, i.e. foils, fibers or, probably, metal particles in a graphite matrix, but to the metallic surfaces used. We can add that gold does not show this effect or, if it does, to only a minor degree.

The experiment& part of this research was presented by P. B. Haney for the award of an M. S. degree in materials science at the University of Virginia. The support of this research by the Office of Naval Research, Metallurgy Branch, Arlington, VA, is gratefully acknowledged.

References 1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16

17 18

R. Holm, Electric Contacts, Springer, New York, 1967, pp. 16 - 18. R. Holm, Electric Contacts, Springer, New York, 1967, p. 368. R. A. Burton, Wear, 59 (1980) 1. J. P. Netzel, Wear, 59 (1980) 135. R. Helm, Electric Contacts, Springer, New York, 1967, p. 34. Phys. Today, (December 1980) 19. J. A. Satkowski and W. R. Seng, Segmented magnet homopolar machines, Seminar on Energy Storage, Compression and Switching, Canberra, Australia, 1977. C. M. Adkins III and D. Kuhlmann-Wilsdorf, Electrical Contacts - 1979, Illinois Institute of Technology, Chicago, 1979, pp. 165, 171. C. M. Adkins III and D. Kuhlmann-Wilsdorf, Electrical Contacts - 1980, Illinois Institute of Technology, Chicago, 1980, p. 67. V. Srikrishnan, S. Dillich and D. Kuhlmann-Wilsdorf, Electrical Contacts - 1978, Illinois Institute of Technology, Chicago, 1978, p. 635. C. M. Adkins III and D. Kuhlmann-Wilsdorf, Wear, 71 (1981) 119. S. Dillich and D. Kuhlmann-Wilsdorf, Electrical Contacts -- 1979, Illinois Institute of Technology, Chicago, 1979, p. 185. R. Holm, Electric Contacts, Springer, New York, 1967, p. 436, F. P. Bowden and D. Tabor, Friction - an Introduction to ~~bology, AnchorDoubleday, New York, 1973, Fig. 38. E. Rabinowicz, Friction and Wear ofMaterials, Wiley, New York, 1965. D. Kuhlmann-Wilsdorf, in D. A. Rigney (ed.), Fundamentals of Friction and Wear of Materials, Proc. 1980 ASM Materials Science Semin., American Society for Metals, Metals Park, OH, Chap. 5, in the press. R. Helm, Electric Contacts, Springer, New York, 1967, p. 438. D. A. Rigney and J. P. Hirth, Wear, 53 (1979) 345.

Note added inproof. Recent theoretical considerationsindicate occurs when, through local cold welding, the a-spots aggregate Wilsdorf, 2. Metallkd., (December 1981), in the press).

that the discussed rise in p into groups (D. Kuhlmann-