Friction and wear of metal-composite electrical contacts

Friction and wear of metal-composite electrical contacts

119 Wear, 158 (1992) 119-140 Friction and wear of metal-composite N. IL Myshkin electrical contacts and V. V. Konchits Metal-Polymer Research In...

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119

Wear, 158 (1992) 119-140

Friction and wear of metal-composite N. IL Myshkin

electrical

contacts

and V. V. Konchits

Metal-Polymer Research Institute, Belamssian Academy of Sciences, Gomel (Belams)

Abstract

The effect of electric current on friction and wear in the sliding contacts of composite material against metal is considered. It has been shown that this effect is similar to the lubrication of contact but there are certain differences in tribological behaviour of materials with non-carbon polymer binder and metal-filled composites. 1. Introduction

Electrical sliding contacts are known to be moving assemblies for which conventional lubricants cannot be used and self-lubricating composite materials are considered as the most efficient way to solve the problem of current collection and wear resistance. These composites traditionally are fabricated by the sintering of carbon-based compounds or hot pressing [l]. There are three basic types of material used in brushes: metal-graphite, carbon-graphite and electrographite. In most cases the development of new brush materials for modern current pickups is accomplished by empirical selection of the composition of composites that are most resistant to wear under the action of the current load and have the required electrical characteristics. The investigations performed in this field concern primarily the study of the dependence of the friction and electrical characteristics of the material on the quality and the number of components, the load-velocity parameters of the current pick-up assembly, the ambient medium conditions and other factors. Rational contact material selection is made difficult by the inadequate level of knowledge about the physics of friction current passage to the sliding contact (SC). This applies particularly to the study of the patterns of electrical current passage through the contact with the presence of transition layers in the friction zone. For the friction assemblies of metal-composite type the formation on the metal surface of an intermediate transition layer (a “third body”) is a general principle for the assurance of high antifriction qualities. However, the electroconductive properties of the transition layers, the nature of their influence on the current passage through the contact, and the behaviour of the electrical characteristics have received very little study. The electrical current passing through the friction zone has significant influence on transition layer formation, on the friction processes, and in the final analysis on the friction behaviour of the pair. However, the mechanism of the electric current influence on friction interaction in the SC has also received little study. The features of current passage through a contact covered by transition layers have been considered elsewhere [4-81. This paper has been inspired by J. Lancaster’s ideas relating to the tribology of brush materials. The authors have tried to evaluate their own data as well as data presented elsewhere, emphasizing the interrelation between the electrical phenomena and tribological behaviour of the composite brush materials.

Elsevier Sequoia

120

2. Effect of the electric current

on friction

in metal-compo+te

sliding

contact

In most cases the passage of an electric current through the metal U.S. the selflubricating composite material SC is accompanied by a significant change in the friction characteristics. For the traditional electrobrush materials an increase in the current load generally leads to a reduction in the brush noise (“chattering”) and a reduction in the brush friction against the collector [9, lo]. In electrical machine operational practice these changes are so significant that in many cases we speak of the “lubricating” action of electric current, or current “lubrication”. Analysis of the available data shows that this nature of the electric current influence on friction is most typical of the carbon electrobrush materials. However, for many metal-containing brush types the decreasing dependence of the friction coefficient f on the current density j is less marked (sometimes it is not observed at all) [ll]. The described dependence is also not typical for some new types of electrobrush material nor for metallic contacts which prevents consideration of the universal “lubricating” action of current as an additional factor influencing friction. In fact our experiments [12] have shown that, in the region of the current loads (up to 50 A cm-“) actually used in the SC of electrical machines for different contact materials, three basic forms of the dependence of f on j are possible (Fig. 1). A smooth reduction in the friction with an increase in the current load (the abovementioned electric current “lubrication” effect) is characteristic of carbon electrobrush materials of the graphite, electrographite and carbon-graphite classes, and also of the metal-graphite brushes under conditions where they contain limited amounts of the metallic components and binder (Fig. 1, curves 1 and 2). With a high non-carbonized binder content (20-30 wt.%) in the brushes obtained using the hot-pressing method the friction is reduced to a lesser degree, but there is a marked reduction in f upon achieving the critical current density (curve 3). For the composite materials with a high metallic component content (over 90 wt.%) the “lubrication” effect is not generally observed; conversely, at high current loads there is a tendency towards an increase in the friction (curve 4).

u

0.30

4

0.05L

I

5

I

10

1

15

I

I

20

25

I 30

j,Aicm2

Fig. 1. Dependence of the influence that the nature of brush material has on the friction electrographite brush; curve 2, MGS-7 coefficient on the current density: curve 1, EG-2A copper-graphite brush; curve 3, brush with a high non-carbonized polymer binder content (25 wt.%); curve 4, brush with a very high copper content (94 wt.%); -, curves for an anodepolarized brush; ---, curves for a cathode-polarized brush.

121

In accordance with the concept of ref. 13, at the electrical SC the current load can be examined as an additional friction system input factor (together with the load, velocity, temperature, intermediate medium etc.), leading to a change in its internal factors (roughness, heat release in the friction zone, structure and mechanical properties of the near-surface layers etc.), which in turn influences the friction characteristics of the SC. From this viewpoint the establishment of the causes of the dependence of the friction at the SC on the current magnitude (and in many cases on its direction) observed in practice should be based on study of the electric current influence on the internal factors of the system. Analysis of the data in the literature shows that in most cases a change in the friction characteristics of the brush collector (slip ring) type of SC when loaded with current is associated with additional heating of the friction zone because of the Joule heat and the change in the collector surface state. However, the different nature of the influence of the current load on the frictional behaviour of the different types of friction material and the unsuccessful attempts to explain several experimental facts by the action of only these two factors indicate that the electric current also influences other internal factors, which under certain circumstances may play a decisive role. 2.1. Mechanism

of the “lubricating” action of the electric current The cause of friction reduction on increase in the current load in the brush collector contact have been studied by several researchers; however, they have not arrived at a common point of view. According to Holm [14], who studied the operation of graphite brushes, the surface temperature increase upon current passage leads to a reduction in the energy of the bond between the oriented graphite flakes in the collector film and on the brush surface, thereby facilitating sliding. Lancaster and Stanley [15] say that there is no direct influence of the electric current on the intermolecular interaction and suggest that the influence of the current reduces to an irreversible change (oxidation) in the material in the brush surface layer, leading in the final analysis to a change (reduction) in the real contact area (RCA). Hypotheses have also been proposed [16, 171 in accordance with which current “lubrication” may be due to softening of the collector material on the contact spots and facilitation of the “tearing out” of the weld bridges which arise, to an increase in the amount of wear products in the contact zone, to the transformation of sliding friction to rolling friction, and to the heating, softening and fracture of the wear products between the rubbing surfaces. The brittleness and the relatively low (in comparison with the metals) elastic modulus and elastic limit of the brush materials for which the effect of “lubrication” by the electric current is observed suggest that the work of the frictional forces is expended on overcoming intermolecular interaction on the segments of elastic contacting and on destruction of the surface of the material by microcutting [18, 191. Our model experiments [4, 121 showed that with a realization of a predominantly elastic contact and an unchanged state of the surface layer on the metallic counterbody (sliding of electrographite specimens over a “fresh”, i.e. not having friction transfer layers, polished copper surface with low specific loads was realized) the friction coefficient remains almost unchanged even with current densities exceeding 100 A cm-‘, while for multiple brush ring contact under actual conditions a reduction in f is noted even for jl5 A cm-’ (Fig. 2); nor is there any influence of polarity on the magnitude of the friction coefficient. A marked reduction in f is observed only with very high current densities tj> 103 A cm-‘), which are not normally used in current pick-up assemblies with contact materials of the given type. However, we observe the reversible influence of

122

Fig. 2. Friction co&icient f as a function of current density j for anode-polarized electrographite EG-2A sliding over a copper ring (curve 3) (diameter, 40 mm; p==loO hPa; w=20 T rad s-l) and over a flat copper specimen with a “worked” wear product film (curve 2) and without a film (curve 1) @= 100 hPa; v = 17 urn s-l).

Fig. 3. Conditionai ultimate stress c (curves 1-3) and friction track width d (curves 1-3) vs. contact current density j for brushes VT-I (curves 1 and 1’1, MGS (curves 2 and 2’) and EG2A (curves 3 and 3’).

electric current on the resistance of the surface layers of composite materials to destruction with realization of the microcutting regime (Fig. J). As the current load is increased, the ultimate stress o (the ratio of the tangential force I; necessary for

brittle fracture scratch formation to the scratch cross-sectional area: a= 1Wr0/d3) decreases rapidly with simultaneous increase in the friction track width and the temperature in the contact zone. In the region of current densities of 104-105 A C111-2,

typical of individual significant. The calculated

contact spots temperature

[Xl, 211, the aforementioned rise in the contact zone

changes

are very

may exceed 500 K

for carbon materials. Experiments (Fig. 4) with current-loaded and currentbess specimens of the test materials sliding along the same friction track on a slip ring showed that the counterbody surface changes taking pIace on passage of an electric current through the contact

123 II 0.20 f,

f,

(4 f,

0.16 0.12

LW~

0.20 0.16 0. I2

m

II I

----_--

i

i

Ej-----+ ----

0.20

0.20 fo 0.16 1

I------

---

0.16 0.12 ~~

F 1 4 2

-

:

I t .---------=_ 1

CP F

1 es

1

IC r

2

._ i-

d

Fig.

4. Friction coefficient of a EG-2A test brush (shown in black on the diagram) sliding over the same friction track with an identical secondary brush for secondary brush loads of (a) 300 hPa and (b) 1500 hPa: curves 1, test brush current loaded, secondary brush currentless; curves 2, test brush currentless, secondary brush current loaded; I, III, j-0; II, j= +20 A cm-‘.

a decrease or an increase of the friction coefficient, depending on the operating regime (specimen polarity, mechanical load, vibration etc.). The results obtained show that the primary factor leading to friction reduction on increase in the current load in the brush contact is the thermal action of the electric current on the brush surface layer, and this action is reversible. The reversibility of the current’s thermal action is indicated, for example, by the rapid change (1-2 s) in the magnitude of the test brush friction coefficient upon activation or deactivation of the current (Figs. 4(a) and [email protected]), curves 1). It is unlikely that, in such a short time, there could take place on the brush surface irreversible changes leading, for example, to a change in the RCA, a suggested in ref. 15. Considering the high local temperatures developed on the contact spots with the indicated current densities, and also the wellknown fact of the reduction in the strength of solid bodies on increase in temperature, we can assume that the current passage leads to weakening of the bond between the structural elements of the material in the microvolumes adjacent to the RCAs. The resuit is facilitation of the brush surface damage process on the part of the contact spots where micro~uttin~ is realized. At the same time the elastic modulus of the brush material examined, dete~ining the dimension of the contact areas in the case of elastic deformation, depends comparatively little on the temperature. Therefore there is no significant influence of the electric current on the frictional interaction of the pair with realization of a predominantly elastic contact. Thus the primary factor leading to manifestation of the “lubrication” effect for the group of materials examined is the thermai action of the electric current on the surface of the material, leading to reversible weakening of its structure in the microvolumes adjacent to the contact regions and to reduction in the work of the frictional forces on the part of the contact spots where the microcutting regime is realized, At the same time the magnitude of the limiting stress leading to realization of microcutting decreases. As a result, together with the reduction in the friction in the brush-coilector SC, another well-known effect appears - an increase in the intensity of brush wear with increase in the current density f3, 10, ZO]. The i~uence of the operating regime on the state of the transferred layers leads to a situation in which there is no unique may lead to either

124

relationship between coefficient [ 121.

the thermal energy power released by the current

and the friction

2.2. Characteristics of frictional behaviour of brush materials with low thermal and electrical conductivities. Microtribometric studies modelling the influence of electric current on the elementary frictional interaction events [4, 121 have shown that for each brush material type there is a limiting value of the current density on the contact spots (Fig. 3). Upon reaching the limiting current density the friction track width increases markedly, the clarity of the track outline disappears, friction becomes unstable, and the quantity 0 decreases by one to two orders of magnitude. It is clear that, in this case, the Joule heat release leads to destruction of the connective skeleton of the brush material in the region adjacent to the contact spot, as a consequence of which the monolithicity of the material is lost. We shall examine the factors determining the magnitude of the limiting (critical) current density jcr on the contact spot. It was shown in ref. 21 that in the case of a single circular contact travelling at a low speed the surface temperature increase due to current passage can be estimated from the formula

‘= njh,

UJ + &)a

Considering

j= 2q4+

that I=fn#j,

A*)

Ucd

or Jcr =

a=dl2,

lJ,=IR,,

Wcr - T,)(h + A,)

we have

lR

R,d3

Consequently, for contact spot dimensions, jcr will be lower for lower thermal conductivities and lower thermal stabilities of the brush material and for higher contact resistances. As was shown above, with the current loads actually used in electrical machines for the basic brush material types, the surface changes due to current heating are reversible, i.e. the contact spot temperature does not reach the critical value. However, on reduction in the thermal stability or the thermal and electrical conductivity of the brush material, the critical value of the temperature on the contact spots will be reached at lower current densities in accordance with eqn. (2). An experimental verification on a microtribometer confirmed that for the “hot-pressed” brush materials with a binder of the same type (FFS) the value of j_ is lower for larger amounts of binder, i.e. for lower thermal and electrical conductivities. With high non-carbonized binder content (above 20 wt.%) in the brush the values of jcr on the contact spots decrease so significantly that they may be reached even with the current loads actually used in the SCs of electrical machines. It is precisely this circumstance that leads to anomalies in the behaviour of the frictional and electrical characteristics of the contact and reduction in the allowable current density for brush materials with a high noncarbonized thermoreactive binder content. According to ref. 22 such materials are serviceable only at relatively low current densities (2-6 A cm-* for the nominal contact area). On increase in the current load we observe instability of operation of the current pick-up assembly, deterioration of commutation and accelerated brush wear. Our experiments (41 relating to the multiple brush-collector contact confirmed that jn depends on the specific electrical resistance of the material and the contact voltage drop (resistance). At the same time it was shown in experiments current-conducting and currentless brushes sliding over a common friction track that the change in the

125

counterbody surface condition (intensification of the abrasivity) for j>jcr is not the primary cause of the change in the wear intensity. Summarizing the obtained results, we can conclude that, in the case of the group of materials examined, with current densities above the critical value (j>j,) the temperature on most of the contact spots reachesvalues sufficient for intensive destruction of the binder. Destruction of the connective skeleton in the surface layer microvolumes adjacent to the contact areas leads to weakening of the bond between the filler particles and to their easy removal, which explains the intensification of the wear. The reduction in the magnitude of the shear strength of the material in the surface is the cause of the marked reduction in the friction coefficient. Thus for composite materials with a low thermal stability and a low electrical conductivity the Lancaster and Stanley concepts [lo] are applicable, in accordance with which the friction behaviour of the SC is determined by the intensity of destruction of the binder in the contact material surface layer under the action of Joule heat.

2.3. Influence of the electric current on the frictional interaction of highly filled metalcontaining brush materials

It was noted above that the frictional behaviour of brushes with a high metallic component content (over 90 wt.%) differs from the behaviour of the carbon materials; the decreasing dependence off on j does not show up in practice and at high loads the friction increases in many cases. Typically, the surface layers formed by such brushes on the counterbody surface have only a slight amount of transferred brush material. Consequently, the probability of formation of contact spots of metal-metal type during sliding is high. It is obvious that in this case most of the electric current is transmitted through the metallic contact spots, and the nature of the frictional interaction in the limits of these spots determines the frictional behaviour of the multiple contact as a whole. Microtribometric studies [21] have shown that for metallic contact spot current densities not exceeding 103-104 A cm-’ there is almost no change in either the overall friction coefficient or its molecular and deformational components. However, at higher current loads, there is a tendency towards an increase in the friction coefficient and the track width and the “jumpy” nature of the friction intensities (Table 1). Calculation of the contact spot temperature increase owing to current passage using eqn. (1) showed that because of the high thermal conductivity of the metals and the low contact voltage drop the friction zone temperature is relatively low (no higher than 320-330 K) and cannot lead to the marked change in the physical and mechanical propertics of the metal. With use of the well-known fact that the free electrons drifting through the metal lattice influence the dislocation movement and consequently the metal is capable of plastic deformation it was suggested in ref. 23 that there is direct action of the electric current passing through the friction zone on the nature of the plastic deformation of the contacting surfaces. It was shown [21] that for a metallic SC there is a region of contact spot current loads (104-106 A cm-“) in which the influence of the electric current on friction facilitates plastic deformation of the surfaces, leading in the final analysis to increases in the RCA and the friction coefficient. Upon further increase in the current load (to 106 A cm-’ or above) the thermal action of the electric current intensifies and becomes dominant, leading to softening of the surface layers, an increase in the RCA, and scuffing. With contact spot current densities of less than 103 A cm-‘, both the plasticities and the thermal actions of the current are not significant. Under these conditions a more significant factor may be the influence

0.08 0.20 0.26 0.32 0.38

0.1 0.5 1.0 1.5 2.0

0

(A)

Contact voltage drop U, (V)

Contact area (X 10e6 cm2)

3.4 3.4 3.6 4.0 4.5 5.2

21 21 21.4 22.5 23.9 25.8

on friction characteristics

Friction track width d (pm)

of electric current

Current I

Influence

TABLE 1

0.062 0.062 0.067 0.080 0.085 0.090

f

Friction coefficient

WS. copper

0.009 0.009 0.009 0.009 0.010 0.011

Deformation component of friction coefficient fd

of steel indenter

0.053 0.053 0.058 0.071 0.075 0.079

fm 0.1 0.6 7 17 30 44

(K)

Surface temperature rise 0

pm)

0.3 1.4 2.5 3.3 3.8

Current density j (X10’ A cm-‘)

N; u =2.25 pm s-r; r,=500

Molecular component of friction coefficient

system (P=O.15

z o\

127

of the external electric field in the clearance on the kinetics of the oxidative processes, as has been noted repeatedly in studies of weak-current SCs. Thus the peculiarities of the frictional behaviour of the group of materials examined cause the friction zone to be dominated by the metallic contact spots and are explained by the specific nature (noted above) of the influence of electric current on them.

3. Effect of the electric current

on wear in metal-composite

sliding

contact

Numerous studies of the operation of pure carbon electrobrush materials (brushes of the “black” grades (graphite and carbon) and electrographite) have shown [18, 191 that the low elastic limit and the brittleness of these materials lead to two forms of contact interaction in the de-energized state: predominantly microcutting in the runningin process and elastic deformation in the steady state regime after the wear products fill the microdepressions of the metal counterbody surface. In the latter case, wear is the result of contact fatigue. Direct confirmation of the fatigue wear mechanism for the carbon electrobrush materials of the basic types was obtained in refs. 24-26. Scobert [3] and Meyer [17], while not denying the existence of fatigue wear, consider that the primary contribution to wear of the de-energized carbon materials in the steady state operating regime is that of microcutting and grinding of their surface by the abrasive wear products and by the counterbody microasperities with disruption of the external friction conditions. This viewpoint agrees with the fact that on the surface of the brush and the collector along the direction of sliding there is always a considerable number of grooves (striae), located precisely opposite one another. In the general case for the self-lubricating contact materials, wear in the absence of an electric current is apparently the sum of the fatigue and abrasive forms of wear. The passage of an electric current through the contact significantly complicates the processes leading to wear of the contact elements. The general pattern is rapid increases in the intensity of their wear on increase in the current density [3, 10, 201. The additional wear of the contact elements arising upon loading the SC with electrical cured is termed “electrical” wear in the absence of a current [26, 271. This term is introduced on the basis of the final result without accounting for the fact that the actual causes of the additional wear may have both an electrical nature (electroerosion with sparking or arcing) and a mechanical nature (change in the roughness of the contact surfaces and their physical and mechanical properties, and so on) 120, 26, 283. Individual attempts to find the fundamental relationship between the wear intensity and the magnitude of the current load have not been successful. It was impossible to predict the wear intensity for various sliding conditions on the basis of experiments conducted under the specified conditions even for electrobrushes of the same type 1181; this was due to the great variety of factors in~uencing the wear process at the SC.

3.1. Factors

leading

to electrical

wear in the absence

of sparking

In the first investigations devoted to studying the patterns of frictional interaction in the contact of a carbon brush with a copper collector, the primary cause of the unfavourable influence of an electric current on brush wear was considered to be the increase in the roughness and intensification of the abrasive properties of the metal counterbody surface [29, 301. This viewpoint was later confirmed in several studies

128

[20, 31, 321. A high collector temperature, deformation of the crystal lattice of copper during friction (causing stressing of the lattice and increased reactivity), the processes of electrolyte decom~sition of the adsorbed moisture in the friction zone and also the directional displacement of the ions under the action of the electrical field lead to significant oxidation of the wear track part of the collector. Current passage is accompanied by continuous alternation of the processes of fretting of the oxide films on the areas of direct contact with formation of the co-called spots [20] and their reoxidation upon emergence from below the brush. The result is roughening of the collector surface and the appearance of hard oxidized particles in the friction zone, which leads to intensification of the abrasive wear. In many cases this factor may actually play the governing role, particularly for brittle carbon brushes, the wear intensityofwhich is extremely sensitive to the counterbody surface conditions [25, 331. The following experimentally obtained results indicate this. (1) In many cases with current-carrying brushes sliding over the same friction track the wear of the track is almost completely determined by the quantity j and by the nature and polarity of the current-loaded brush [30]. (2) Dependence of the brush wear intensity on the collector material is observed, and for metals that are characterized by a high formation rate and abrasivity of the oxides (e.g. A1203 on aluminium) the wear intensity is considerably higher (sometimes by an order of magnitude) than for copper and steel [9]. (3) In the case of operation against slip rings the wear of the cathode-polarized carbon brushes is higher than that of the anode-polarized brushes, with a more highly oxidized and rougher working surface of the ring beneath the brushes j20, 301. However, there are several facts indicating that abrasive wear often does not play the governing role, particularly for metal-containing brushes. For example, an increase in the wear intensity on increase in the contact current density is observed under conditions when oxidation of the metal counterbody surface is eliminated: in the operation of the brushes against gold or silver collectors and slip rings in a carbon dioxide atmosphere and in vacuum [34, 351. According to the data in ref. 18, wear fragment separation from the brush surface is the result of weakening of the material in the near-surface layer during mechanical inputs and adhesion of the segments on contacting. On the basis of several experiments it is shown that adhesion of the carbon materials to the copper oxides is considerably weaker than to pure copper. Current passage causes breakdown of the oxide film and an increase in the adhesion on the segments of brush contact with the copper, and therefore an increase in the tangential stresses at the interface and a reduction in the number of deformation cycles causing wear. Intensification of the adhesion on the contact spots may also take place because of dissociation (under the action of current) of the thin (1.5-3 nm) film of adsorbed and chemisorbed water molecules or organic substances, separating the contact elements on the areas of elastic contact 136, 371. Here, microcontact welding and brush material transfer to the counterbody may take place 1371. If on an increase in current load such conditions arc reached at most of the contact spots, then we observe a transition from “weak” to “strong” wear or to abrasion with the formation of a large amount of powder-like wear products ]37]. In ref. 27 in explaining the detected effect of the higher (by two orders of magnitude) wear of carbon brushes in an oxygen medium in comparison with that in a hydrogen medium it was suggested that electrical wear is caused by weakening of the brush surface layer as a consequence of oxidation. The possibility of realization of such processes and their influence in increasing brush wear with the action of a

129

current load was later confirmed in ref. 15. Current passage causes at the contact spots and the microvolumes of the brush material adjacent to them a marked increase in the temperature, the values of which may exceed the critical magnitude at which the brush material (or more often one of its phases) oxidizes noticeably. The diffusion of atmospheric oxygen into the porous skeleton of the brush contributes to this process. As a rule the brush binder, having the lowest critical temperature of the ingredients used in brush production, will most probably be subject to oxidation. Preferential oxidation of the binder weakens the bond between the grains in the material of the brush surface layer, thereby contributing to spalling of the grains in the event of mechanical inputs. The role of this factor in increasing wear under the action of cured the composite contact materials with a low thermal stability and electroconductivity is evident. The data presented in Section 2.2 have shown that the wear intensity on thermal oxidative degradation of non-graphitized binder results in the friction behaviour of composite materials with a low heat stability and a low electrical conductivity (materials with a high content of synthetic binder). At current densities higher than the critical value, this effect becomes dominant with respect to brush wear 1381. Figure 5 is a good illustration of this case and presents the wear and electrical performances of copper-graphite brushes based on phenol-formaldehyde binder. The difference between the wear of the anode and cathode brushes when rubbing along the same wear track illustrates the fact that the countersurface of a ring in this case has little effect. The wear rate is clearly related to the value AU controlling the heat behaviour in the contact and hence the thermal oxidative degradation of the brush binder. The anode polarity can also result in oxidation of the brush due to oxygen transport under the effect of the electric field gradient. For traditional brush materials, having a high thermal conductivity and a high electrical conductivity, in the normal regime of operation the thermal action on the material in the near-surface layer basically has a reversible nature [4, 12, 391. Nevertheless, its contribution to electrical wear must be considered. The reason is that in the operation of the real contact, because of mechanical or electrical factors, fluctuations in the electric current are always observed as a result of which the current density through the individual contact spots may exceed the statistical average value by several times. As a result, in the brush material microvolumes

b

0

3/

/

l

0,64

\

ZI .

m

I (4

I

I

I

10

20

30

~,Alcm~

@I

J,A/cm2

Fig. 5. Contact voltage drop AU and wear rate Ih of copper-graphite brushes functions of current density j when rubbing along the same track on the contact Ml; +=40 mm; v=1.5 m s-‘; p=lOO hPa).

(MGS-7) as ring (copper

adjacent tQ these spots,the temperaturemay be very high. The lower the thermal stability of the brush material, and in particular of its binder, the higher is the ~robabili~ that brush material oxidation occurs in the ~divi~ual mi~rovolnmes and the greater is their influence on the wear intensity. Careful experiments [40] have shown that at high temperatures, high current densities or high sliding velocities an important factor influencing the wear of the carbon materials may be direct oxidation, leading to carbon loss in the form of gaseous oxides. McKee ed a& 1401 found a correlation at temperatures of 460-473 K between the wear rate of carbon brushes and their oxidation temperature. It was shown that a sufFiciently high temperature for oxidation of the carbon material can be created by the current passing through the high resistance in the contact zone, The copper oxide particles migrating within the brnsh act as catalysers, reducing the temperature at which intensive oxidation begins. Thus the current determines the oxidation rate and the migration of the copper catalyscrs. The ability of the surface of a compostite to withstand abrasive wear is determined by its strength properties. It was shown in refs. 12 and 3.5 that in a wide range of current densities (103-105 A cm-‘) at the contact spots before the critical values are reached, there is a reversible reduction in the strength ofthe material in the microvolumes adjacent to the contact areas experiencing the current load. As a result the fraction of contact spots where the conditions for the onset of microeutting are realized will increase. Moreover, on reduction in the strength properties which is a function of the surface flash temperature, deformation of the surface to~o~aphy may take place as a result of the normal and tangential furces. There is no direct experimental confirmation of the action of this electrical wear factor; however, its existence can be hypothesized on the basis of the data presented in ref. 1 on the influence of electric current on frictional interaction, and its role wiI1 obviously be determined by the temperature dependence of the strength properties of the contact materials. Some workers consider that the “thermal shocks” arising as a result of Joule heat release on the smafl electrically conductive contact areas 1411 are one of the factors leading to intensification of the wear of the contact elements under the action of an electric current. The considerable heating in the course of a short time interval (microseconds) of the small composite material volumes adjacent to the conductive spots leads to high temperature gradients and mechanical stresses because of thermal expansion. This applies in particular to the materials whose components have significantly different thermal expansion coefficients, e.g. copper-graphite brushes (the coefficients of thermal expansion for copper, graphite and Bakelite are respectively 17 X 10e6 K-“, 7.8 x 10w6 K-i and 22 X 10V6 K-i). The appearance of mechanical stresses may lead to the formation of microcracks at the interfaces of the components. The combination of thermal and mechanical shocks creates conditions for the development of these cracks and failure of the surface layer. In addition to the difference between the coefficients of thermal expansion of the components, another condition for thermal shock onset is the low thermal conductivity of the rhaterial, causing the heat release to be of a local nature. This applies, for example, to the carbon-black electrical brush grades (EG-74, EG-54 etc.) with pitch as the binder. It should be noted that almost all the explanations examined above for the onset of electrical wear in the brnsh-collector SC in the absence of sparking and arcing postuiate that the basic cause is Joule heat released in the friction zone and leading (depending on the operating regime and the nature of the material) to mechanical weakening (disinte~at~on~ of the surface layer of the brush, to intens~~~tion of the adhesive interaction at the contact spots, to ~~tensi~cation of the abrasivity of the

131 collector surface or, finally, to direct loss of brush material as a result of chemical erosion (Fig. 6). It should be mentioned that the nature and consequence of heat release due to current passage and due to friction are not equivalent. 3.2. Characteristics of thermal action of electric current and its influence on wear in the brush contact In the sliding of a brush against a collector (slip ring), conduction may be realized through the segments of brush surface in contact with the clear metal, with metal covered with conductive films of a different nature (adsorption, passivating etc.), and with the friction transfer film. Depending on the structure of the intermediate layer formed on the collector [S], one of the above-mentioned contact spot types is usually statistically dominant, although spots of the other types are present to one degree or another. In the case of clean metallic contact spots (e.g. in the SC of copper-graphite or silver-graphite brushes with a copper collector) the cured density is variable over the contact spot surface, increasing towards the edge of the spot. The maximal temperature increase due to the Joule heat is observed at the interface. If we neglect the temperature dependences of the electroconductivity, then according to Holm [20]

The temperature ,+

k$)

distribution

in the symmetric

constriction

region obeys the law [20]

(4)

According to eqn. (4), a temperature rise that is 100 times smaller than the rise on the contact surface is observed at distances of only about 0.5 of the contact spot radius, i.e. there is almost no temperature rise. The very small mass of the metal heated in the constriction region leads to a very rapid increase in the temperature rise after the initiation of current flow, after about 10-4-10-6 s according to the data in refs. 42-44. However, even in these situations with relatively slow sliding speeds the contact spot lifetime is comparable with these times, as a consequence of which thermal equilibrium cannot be reached. As a result the contact spot temperature rise will be somewhat less than the value determined using eqn. (3). With the presence of thin “tunnelling” conducting films on the metal the distribution of the current lines at the contact spot becomes more uniform. For the limiting case, when the contact spot resistance is almost entirely determined by the film, Joule heat is released within the contact spot plane and the heat source is uniformly distributed over the spot. Therefore we cannot used eqn. (3) to calculate the maximal temperature increase. According to ref. 21 in many cases we can use for this purpose eqn. (1). A qualitatively different picture arises in the case of current passage through the segments of a contact between metal and carbon, which has much lower thermal and electrical conductivities. In this case the magnitudes of R, and of the Joule heat released are determined by the brush material, while the metal acts as an ideal conductor and heat absorber, preventing the maximal temperature from being reached at the contact surface. The most highly heated region lies within the brush at a distance of the order of the mean radius of the conducting spots (Fig. 7). Calculations [ZO] show that the magnitude of the maximal temperature rises for the contact of metals with carbon may exceed the magnitude of the contact surface temperature rise by several times. This situation undoubtedly facilitates realization of the processes examined

-

r

133

Fig. 7. Temperature increase in metal M VS. carbon C constriction region. I is the maximal temperature isotherm. in Section 3.1., leading to electrical wear. For example, oxidation of the binder and chemical erosion or thermal stresses because of thermal shocks under the corresponding conditions will be realized not on the interface but rather at some depth, leading to the formation of cracks and particles weakly bonded with the substrate, and in the final analysis to mechanical weakening of the near-surface layer of the composite and more rapid wearing of the latter. The probability that these processes are realized also increases because for the graphite-based materials the quantities h and p appearing in eqn. (3) and determining the intensity of heat generation and dissipation respectively depend on the temperature. Specifically, after an initial reduction, p remains at the same level or increases with increase in the temperature, while A usually decreases with increase in T [41]. Consequently, on increase in the current load the heat generation increases, but the heat dissipation does not accelerate to the same degree. As was shown in ref. 41, this leads to the existence for any given graphite-metal contact spot dimension of a critical magnitude of the current, namely I,,, above which thermal equilibrium is not possible. For Z&Z,, the increase in the spot surface temperature is limited only by the destruction temperature of the material. The consequences of thermal instability upon reaching Z,, are intensified by the generation of extremely high temperatures only in the graphite at some depth from its surface. The formation of a friction transfer film on the collector surface has a dual influence on brush wear. On the one hand, the friction and intermotecular interaction between the materials of the contact elements decrease, which leads to a reduction in the wear intensity. On the other hand, there is the contact resistance on the segments of brush contact with the transfer film (because of its high specific resistance R, of the contact spots of the brush with the metal). Consequently, the critical magnitude of the current that leads to disruption of thermal equilibrium decreases by a factor of more than 2 for such contact spots [41]. Intensification of the thermal action on rhe surface of the composite inevitably leads to the appearance of factors (Fig. 2 of ref. 1) that lead to an increase in the wear intensity. The dimension of the near-surface layer brush zones that are subject to thermal action is important for the wear processes under the action of an electric current. In the no-current state the mechanical action on the surface layers of the contact elements

134

is due to the process in the real contact zones, and specifically to the appearance of tangential and shearing stresses. The maximal tangential stresses arise at a definite depth h (from a few micrometres to hundreds of micrometres), comparable in magnitude with the dimensions of the real contact spots (Fig. S), and the quantity h defines the thickness of the weakened layer and the dimension of the wear products. The electric current in the SC almost always passes not through the entire RCA experiencing the mechanical load, but rather only through some part of this area 1201. Therefore only some part of the RCA is subjected to the additional thermal action, and the process reflected in Fig. 6 takes place in brush material microvolumes that are much smaller than the microvolumes subject to the mechanical actions. This is shown schematically in Fig. 8. It is evident that the dimension of the wear products, the formation of which was due to the action of the current, will therefore be smaller than the dimension of mechanical wear products, which is always observed in practice [lo, 161. There are many data indicating that electrical wear is primarily the result of increasing severity of the thermal regime precisely on the individual conducting areas and is not the result, for example, of the increase in the average temperature in the interface. Thus it is known [45] that, if the brushes are subdivided into elementary brushes with individual pre-loads (the specific load and the current density remain the same), their overall contact resistance decreases with simultaneous reduction in the wear intensity. In a heavy-current SC with the use of brushes with different nominal contact areas A, the wear depends more on the magnitude of the current than on the current density, defined by the ratio I/A, [46]. We can explain these results if we consider that subdivision of the brush increases the RCA (in addition to the reduction in R,, this is also indicated by an increase in the friction coefficient [47]) and for unchanged magnitudes of the mechanical load and the current change in A, (and correspondingly the quantity j) has little influence on the RCA and (we can assume) the conducting area. An increase in the thermal loading on the conducting segments leads to an increase in the wear intensity. 3.3. Influence of the electric field in the clearance The condition of the contacting surfaces undergoing the wear processes at the SC may be influenced not only directly by the electric current in the form of a directed flow of charge carriers but also by the electric field in the contact clearance. Special experiments [1’7] have shown that, in the case of a high potential gradient, carbon powder is released from the surface of the carbon brushes, influencing the wear and the contact resistance. It was shown in [48] in a study of the electric strength of micron clearances and the conductance in a brush contact that, with clearances of

s

Fig. 8. Schematic diagram of electric current passage through a contact spot experiencing a mechanical load: 1, conduction zones; 2, maximal temperature rise; C, collector; S, contact surface; B, brush.

135

up to 1 pm, deformation of the electrode surface elements and shorting of the gap are possible because of the mechanical stresses created by the forces of the electrostatic field, The presence, on the surfaces, of particles that are weakly bonded to the substrate facilitates this process. As a result, as the contact spots separate, the onset of sparking or arcing is preceded by a “bridging” stage of circuit interruption. The particles forming the bridge are in essence already wear products. Thus under certaitl conditions the electric field may lead directly to the formation of wear particles. However, further studies are necessary in order to evaluate the significance of the role of this factor in electrical wear. The influence of the electric field on surface layer formation in the SC, and primarily on the oxidation and mass transfer processes, is more obvious. The oxidation of several electrotechnical metals (copper and its alloys, steel and nickel) is a diffusion process, in which the metal ions migrate through the oxide film to the surface, where they interact with atmospheric oxygen. The oxidation rate is in general determined by the average surface temperature. However, on the metallic surface segments adjacent to the current-carrying areas, the potential gradient (depending on the field direction) may contribute to (the metal is anode polarized) or prevent (the metal is cathode polarized) movement of the ions through the film, thus influencing the oxidation of copper collectors as has been confirmed experimentally under laboratory conditions [32, 491 and also shows up in practice. For example, the cathode-polarized carbon brushes form on the copper collector a light-brown film, consisting basically of copper oxide, while the anode-polarized brush forms a dark film with a low CuZO content and higher transferred brush material content. This last fact also reflects another general pattern - the influence of the electric field on the transfer of the brush material to the surface of the metal. Thus, in the normal operation of electrical machines, much more material is always transferred from the anode brush to the collector than from the cathode brush. A significant part of the brush material deposited on the collector by the anode-polarized brush is transferred to the surface of the cathode-polarized brush [SO]. Conversely, the transfer of material from the cathode brush to the collector surface is slight and there is almost no transfer from the collector to the anode brush. It is typical that more intensive transfer under the anode-polarized brush is observed regardless of the counterbody material, the ambient medium and the load velocity parameters of the current pick-up assemblies. In spite of the large number of publications describing the external manifestations and consequences of this phenomenon, its mechanism is not yet completely clear. Transfer of the brush material takes place in a direction which can be predicted for the positively charged ions or particles. Attempts have been made to explain this as an electrolytic effect or by physical migration of the charged particles under the action of the potentia1 gradient, but these explanations are not entirely satisfactory 131. A consequence of the influence of the field on the formation of the separation surface at the SC are the differences in the nature of the mechanical contact and, what is still more important, the differences in the mechanism of current passage for brushes of different polarities, particularly in the case of their operation along separate friction tracks [8, 511. Both factors have a very direct influence on the electrical wear. In the case of cathode polarization of the brush, current passage is realized basically on the segments of physical contact of the surfaces of the brush and the slip ring metal while, in the case of anode polarization, current passage takes place primarily through the intermediate layer of transferred brush material 181. For the metalcontaining brushes this !eads to manifestation of the inequality AU, > AU_ and

136

consequently more intensive heat release under the anode-polarized brush. It is evident that this in turn leads to intensification of the electrical wear of the anode-polarized brushes .(in accordance with Fig. 6) and shows up in practice in the existence for most metal-containing brush types of a polar inequality Zh_ >Zh+ in the wear. Electrical wear of carbon brushes is to a large degree determined by the oxide film fritting process. The more intensive oxidation under the cathode-polarized brush and the peculiarities of current passage exacerbate the unfavourable influence of this factor, which is one of the primary causes of the inequality Zh- >I,,+ observed for the “black” brush grades. The presence of the electric field in the clearance also leads to “metallization” of the cathode-polarized brushes which is observed in many cases. This phenomenon, which is particularly characteristic for machines with a closed ventilation cycle involves settling of the metallic particles on the working surface of the brushes, hardening of the particles on this surface and subsequent abrasive action on the collector surface. The primary cause of metallization can obviously be considered to be so-called “bridging” transfer. On part of the contact spots as a consequence of fritting or excessive crowding of the current lines the current constriction region in the metal may soften. As the spots separate molten metallic bridges form between them, as a result of the breaking of which there is transfer of metal from the anode to the cathode. Metal transfer to the brush surface may also be realized by passage of metal ions from the anode during microdischarges injection of the molten metal (particularly from the anode during arcing), vaporization of the metal and its transfer by the thermal wind flux to the brush 1403. Metallization of the brush causes roughening of the counterbody surface, which is one of the causes of the inequa~i~ Zh- >Ih+, typical of carbon brushes. Dissimilar mechanical conditions may cause more intensive sparking under the cathodepolarized brush, which also contributes to manifestation of the inequality between Zhand Zh+ [27]. The polarity of the elements at the SC of electrical machines also determines the mechanism of interelectrode gap breakdown during sparking and arcing [20, 28, 521. This is explained by the stronger erosion of the cathode-polarized brushes during sparking [ZS]. The positive metal ions bombard the brush surface and cause a greater increase in the local surface temperature than is for the anode-polarized brushes. The electroerosion component of electrical wear increases correspondingly. 3.4. Some ways to reduce electrical wear Wear at the SC of electrical machines is due to the action of many factors which are often mutually influencing and difficult to separate. However, if we can determine in each specific case the principal factors and explain the mechanism of wear under the action of current, it becomes possible not only to select the optimal combination of contact elements but also to formulate the requirements on the physicomechanical, electrophysical and other properties of the element materials and to determine ways to create such materials, and also to specify the requirements on the design of both the assemblies and the individual elements. It follows from the data presented in Section 3.1 that the action of several electrical wear factors is facilitated by the presence of oxygen in the ambient medium (the formation of non-conductive films with subsequent fritting, chemical erosion, selective oxidation of the binder etc.). The search for ways to ensure effective operation of the SC and to reduce the harmful influence of oxygen has shown that by using inert gaseous media (carbon dioxide, nitrogen, helium, argon etc.) with a controllable moisture content we can achieve a significant increase in the wear resistance of the carbon and

137

metal-graphite contact materials [27,53,54]. In recent years it has been shown [55-591 that brushes created using the traditional technologies from relatively widely used materials can operate with much higher current densities and sliding speeds than were previously considered possible. In addition to control of the atmospheric conditions, the temperatures (both the ambient temperature and the temperature of the brush), the mechanical conditions of contact, and selection of the optimal materials should also be subjected to careful monitoring. An increase in the current at the SC is limited not only by the resistance of the brush material to thermal actions but also by the dynamic stability of the contact system and its ability to avoid intensive sparking with increases in j and v [16]. The dynamic stability is determined by the velocity regimes, the elastoplastic properties of the contact materials, the dynamic characte~stics of the brushes and the elements of the contact devices, the spectral composition, and the intensity of the vibrations of the current-transferring equipment [60]. An increase in the brush pre-load, the use of many parallel brushes, and good alignment of the machine minimize the problem. In addition, many special ways to increase the contact stability (and as a result the current density) at increased rotational speeds have been proposed and used with success in practice: selection of optimal brush angles and brush holder designs, use of circumferential rocking, dual and subdivided (compliant) brushes, a multitude of small, independently sprung brushes etc. [45, 611. Significant achievement in this direction appears to be the development of fibre brushes: carbon, metallic and metallized carbon [59, 62-641. Solid (monolithic) brushes have limitations with respect to speed because of their mass and stiffness (upon reaching some speed the ability to track slight defects on the surface of the rotating collector is lost). The basic difference in the new brush construction from the traditional design lies in the fact that current transfer takes place through a large number of electrically parallel but mechanically independent contacts of thin (tens of micrometres) fibres. Because of the small intertia and high elasticity of the fibres, this ensures reliable contact with the collector regardless of the condition of its surface and vibration, which makes it possible to avoid sparking as a result of “bouncing”. In the course of improvement of the design and reduction of their cost the fibre brushes will be used more widely in the traditional electrotechnical fields.

4. Conclusions Summarizing the results presented above, we can note that the frictional characteristics of the brush-collector (slip ring) type are basically determined by the action of the electric current on the brush surface, and the nature of this action is different for brush materials which differ in their nature. The reversible reduction in the strength of the near-surface layer of carbon brushes upon heating by the Joule heat (reducing the work of the frictional forces on part of the contact spots with realization of the microcutting regime) leads to the effect of “lubrication” by the electric current. The reversible changes in the surface of the brushes with a high non-carbonized polymer binder content, because of destruction of the latter at some critical magnitude of the current load, lead to an anomalous reduction in the friction and an increase in the intensity of wear of these materials. Faciiitation of plastic deformation of the surfaces, leading to an increase in the contact area, explains the frictional behaviour characteristics in the case of high electrical conductivity brush materials with a high metallic component

138

content. The change in the collector surface (transition layer) state upon current passage influences the frictional characteristics to a lesser degree. As was shown earlier [5], the transition layers on the collector surface determine the current passage mechanism and the electrical characteristics of the contact. In the absence of sparking, the basic cause of brush wear intensification under the action of electrical current is the Joule heat released in the friction zone. Depending on the operating regime and the nature of the contact materials the additional heat release resulting from current flow may lead to mechanical weakening of the nearsurface layer of the brush, roughening of the collector surface, intensification of adhesive interaction on the contact spots, direct loss of brush material as a result of chemical erosion, reversible reduction in the brush material strength in the microvolumes adjacent to the contact areas, and mechanical stresses because of thermal expansion. The action of one or more of these factors leads to the appearance of the so-called electrical wear in the brush contact. The electric field in the contact area can be influential factor of wear rate control because of the effect on oxidation, mass transfer and debris formation.

1 P. S. Livshits, handbook on Electrical Machine Brushes, in ~nergoeatomizdat, Moscow, 1983 (in Russian). 2 I. V. Tyomkin, Production of Electra-Carbon Articles, Vysshaya Skola, Moscow, 1975. 3 E. Scobert, Carbon Brushes, Chemical Publishing, New York, 1965. 4 V. V. Konchits, V. V. Meshkov and N. K. Myshkin, Tribolagy of Electrical Contacts, Nauka i Tekhnika, Minsk, 1986. 5 V. V. Konchits and V. G. Savkin, Trenie Iznos, 4 (6) (1983) 972-982. 6 V. A. Belyi, V. V. Konchits, V. V. Meshkov and V. G. Savkin, EIectrotechnika, I2 (1977) 34-36. 7 V. A. Belyi, V. V. Konchits, V. V. Meshkov, V. G. Savkin and A. 1. Sviridyonik, Electrotechniku, 12 (1977) 43-46. 8 V. A. Belyi, V. V. Konchits and V. G. Savkin, Wear, 78 (1982) 249-258. 9 J. Fisher, K. J. Campbell and T. F. J. Quinn, ASLE Trans., 1.5 (3) (1972) 192-200. 10 P. S. Livshits, Electrical Machine Sliding Contacts, Energiya, Moscow, 1974 (in Russian). 11 J. R. Braiiford, Wear, 25 (1) (1973) 85-97. 12 V. V. Konchits, Problems of Friction and Wear, Vof. 18, Tekhnika, Kiev, 1980, pp. 7683 (in

Russian). 13 I. V. Kragel’skii, M. N. Dobychin and V. S. Konbalob, Fundamentals of Friction and Wear Calculations, Mashinostroenie, Moscow, 1977 (in Russian). 14 E. Helm, .7. Appi. Phys., 33 (1) (1962) 156-163. 15 3. K. Lancaster and S. W. Stanley, Brit. I. Appl. Phys., I5 (1) (1964) 2941. 16 V. I. Nellin, N. Ya. Bogatyrev, L. V. Lozhkin er at, Mechanics of Siiding Contacf Transport, Moscow, 1966 (in Russian). 17 R. Mayeur, Rev. G&z. Electr., 64 (1) (1955) 213-221. 18 J. K. Lancaster, Br. .I. Appl. Phys., 13 (3) (1962) 468477. 19 W. T. Clark, A. Connolly and W. Hirst, Br. J. Appl. Phys., 14 (1) (1963) 20-27. 20 R. Helm, Electricat Contacts, Springer, Berlin, 1967. 21 V. V. Konehits, Trenie Iznos, 2 (I) (1981) 170-176. 22 R. R. Paxton, Electrochem. Technol., 5 (5-6) (1967) 174-182. 23 0. A. Troitskii and A. G. Rozno, Fiz. Tekh. Pufuprovodn., 12 (1) (1970) 203-210. 24 W. T. Clark and L. K. Lancaster, Wear, 6 (6) (1963) 467-482. 25 P. V. K. Porgess and I-I. W. Wilman, Proc. Phys. Sot., London, 76 (3) (1960) 513-520.

139 26 J. K. Lancaster,

Nature, 196 (1962) 368-380. 27 R. M. Baker and G. W. Hewitt, Electr. .Z., 33 (6) (1936) 287-289. 28 J. K. Lancaster, Wear, 6 (2) (1963) 341-352. 29 V. P. Hessler, Electr. Eng., 54 (6) (1935) 1050-1058. 30 V. P. Hessler, Electr. Eng., 56 (1) (1937) 8-12. 31 E. Hahn, J. Appl. Phys., 28 (10) (1957) 1171-1176. 32 P. M. Scherer and W. J. Spry, Proc. 5th Conf on Carbon, Vol. 1, Pergamon, Oxford, 1962, pp. 553-560. 33 J. K. Lancaster, Br. J. Appl. Phys., I4 (8) (1963) 497-505. 34 E. Holm, IEEE Trans., Power Appar. Syst., 84 (5) (1965) 65-74. 35 P. K. Lee and J. L. Johnson, IEEE Trans. Comport., Hybrids Manuf Technol., I (1) (1978) 40-%5. 36 J. K. Lancaster, Wear, 34 (2) (1975) 275-290. 37 J. L. Johnson and J. Scheurs, Wear, 78 (2) (1982) 219-232. 38 V. V. Konchits and V. G. Savkin, Problems of Friction and Wear, Vol. 16, Tekhnika,

39 40 41 42 43 44 45 46

47 48

49 50 51 52

53 54 55 56 57

Kiev, 1979, pp. 75-79 (in Russian). V. V. Konchits, Trenie Zznos, 5 (1) (1984) 59-67. D. W. McKee, R. H. Savage and G. Gunnoe, Wear, 22 (2) (1972) 193-214. J. B. P. Williamson and N. Allen, Wear, 78 (1) (1982) 39-48. G. Hilgarth, Elektrotech, 2, Ausg. A, 6 (1957) 176-179. G. Hilgarth, Elektrotech. 2, Ausg. A, 13 (1958) 464-467. F. P. Bowden and J. B. P. Williamson, Proc. R. Sot., Ser. A, 246 (1) (1958) 1-12. I. I. Tuktaev and M. F. Khlystov, Elektrotekhnika, I (1974) 96-101. J. M. Cosstevens, H. G. Rylander and Z. Eliezer, Wear, 48 (1) (1978) 121-130. I. V. Kragel’skii and I. E. Vinogradova, Friction Coefficients GITIMIL, Moscow, 1962 (in Russian). V. V. Postnov, I. L. Gufel’d and Ya. G. Davidovich, Ways to Improve the Quality and Reliability of Electrical Contacts: Summaries of Reports of All-Union Scientific Res. Conf, Leningrad, 1978, pp. 65-66 (in Russian). 0. Kubashewski and B. E. Hopkins, Oxidation of Metah and Alloys (Russian translation). Metallurgiya, Moscow, 1965. K. Wagner, R. Lenz, F. Brischin and H. Scheerbarth, Elektrie, 8 (1964) 248-252. A. E. Stebbens, Proc. Conf on Commutation in Rotating Machines, Institution of Electrical Engineers, London, 1964, pp. 74-77. B. R. G. Swinnerton, M. J. B. Turner and J. E. Thompson, Proc. Conference on Rotating Machines, London, 1964, pp. 68-73. J. L. Jhonson and L. E. Moberly, Proc. [email protected] on Electrical Contacts, Vol. 2, Illinois Institute of Technology, Chicago, IL, 1967, pp. 109-116. J. L. Johnson and J. L. McKinley, Proc. Con$ on Electrical Contacts, Chicago, Proc. Conf on Electrical Contacts, Vol. 3, Illinois Institute of Technology, Chicago, IL, 1970, pp. 155-162. J. L. Johnson and L. E. Moberly, IEEE Trans. Compon., Hybrids Manuf Technol., I (1) (1978) 3640. P. K. Lee and J. L. Johnson, IEEE Trans. Compon., Hybrids [email protected] Technol., I (1) (1978) 40-45. J. R. McNab and J. L. Johnson, IEEE Trans. Compon., Hybrids [email protected] Technol., 2 (1) (1979)

84-89. 58 I. R. McNab, Wear, 78 (1) (1982) 16. 59 M. F. Khlystov, I. I. Tuktaev and 0. V. Ruzaikina, Zntervuz Scientific-Technical Collection, Study of Special Electrical Machines and Rotary Rectifier Systems, TPI, Tomsk, 1980, pp. 140-145

(in Russian). 60 I. I. Tuktaev, M. F. Khlystov, Yu. S. Levashov and 0. V. Ruzaikina, Elektrotekhnika,

4 (1980)

48-50. 61 I. R. McNab, Z’roc. Helm

Conf

on Electrical

IL, 1977, pp. 105-114. 62 A. Marcus, Wear, 78 (1) (1982) 93-107.

Contacts,

Illinois Institute of Technology,

Chicago,

140 63 L. Boger, J. P. Chabreric and J. Saint-Michel, Wear, 78 (1) (1982) 59-68. 64 B. R. G. Swinnerton, Wear, 78 (1) (1982) 81-92.

Appendix:

Nomenclature

P F h

I I CT

I h+,

Ih

i lcr 1 L

N P r0

RC T

T TM To UC AU+,

AU_

V

contact spot radius nominal contact area width of friction track formed by indenter friction coefficient tangential load depth at which tangential stress occurs in microasperity contact zones current critical current wear intensities of anode-polarized and cathode-polarized brushes respectively current density critical contact spot current density distance of isotherms from the centre of the contact spot 2.4~ lo-’ V2 Kp2, Lorentz coefficient load pressure indent top radius contact resistance temperature contact spot temperature increase due to the passage of current density maximal temperature volume temperature voltage drop in a single contact spot voltage drops at anode-polarized and cathode-polarized brushes respectively sliding speed

Greek letters

8 4

oh.3 A Al,

P

u 0 w

A2

contact surface temperature rise running value of temperature rise maxima1 temperature rise specific thermal conductance heat transfer coefficients of contacting specific resistance conditional ultimate stress ring diameter angular velocity

materials