Hardness, friction and wear of multiplated electrical contacts

Hardness, friction and wear of multiplated electrical contacts

538 Wear, 162-164 (19Y3) 538-553 Hardness, friction and wear of multiplated electrical contacts Peter A. Engel mechanical and ~nd~~~l Engineerin...

2MB Sizes 35 Downloads 87 Views


Wear, 162-164 (19Y3) 538-553


friction and wear of multiplated

electrical contacts

Peter A. Engel mechanical and ~nd~~~l

Engineering ~~a~~

B~g~arnto~ ~n~e~i~,

B~n~hamfo~ NY ~3~2-~


Eugene Y. Hsue IBM Development Laboratory, Endicott, NY 13760 (USA)

Raymond G. Bayer Co~ultant,

Vnraf, NY f385t3 (USA)

Abstract Full inte~r~tation of the composite (e.g. Vickers) mi~ohardne~ is given for typical muIt~ayer platings used for electrical contacts. For relatively thick layers plated on a Cu substrate, an effective highly stressed region (delimited by the “plastic boundary”) is identified, which introduces a significant modification in the prior interpretive theory of thin platings (P. A. Engel, A. R. Chitsaz and E. Y. Hsue, Interpretation of superficial hardness for multilayer platings, Thin Solid Films, 207 (1992) 144-152). An approximate computation for the internal plastic work is devised. The work done on the underlying substrate is shown to correlate well with findings of the subsurface deformation theory of multiplated contacts. (R. G. Bayer, E. Y. Hsue and J. Turner, A motion-induced subsurface defo~ation wear mechanism, Wear, 154 (1992) 193-204, R. G. Bayer, A general model for sliding wear in electrical contact, Proc. Wear of Muferia& Con& San Francisco, CA, 2993). In particular, wear and friction are shown to be, respectively, proportional and inversely proportional, to the plastic work done on the substrate during multiple passes of a slider.

1. Introduction Separable





in the

electronics industry must meet multiple criteria to insure reliabili~ with respect to current conduction, low friction, and wear resistance through the projected design life [I]. Such criteria have been fulfilled by a judicious combination of the following consecutive platings [l-4]: a noble metal (e.g. soft and/or hard Au) or a seminoble alloy (Pd or Pd-Ni) for surface film, plated on top of Ni, the latter plated on Cu or a copper alloy, e.g. Be-Cu. This sandwich has the advantage of offering a ductile non-oxidizing surface layer, compatible with a hard protective sublayer (Ni); the copper substrate, which may itself be the top of a layered structure, represents a cushion against repeated mechanical loads arising in the contact. To verify the quality of precision-surfaces, a microhardness test is often performed. In the case of typical electrical contacts, the underlying substrate has a strong influence on composite hardness, yet Vickers microhardness tests can reveal significant information on the thickness and hardness of the over-plated layers. For thin layers with t/d<0.2, where t is the total layer

thickness and d the diagonal of the Vickers indentation, a diamond surface impression is preserved relatively unchanged down to the substrate (Figs. l(a), (b)); this geometric property suggested a computationallpredictive “thin plating”-meth~ [S, 6] of establishing a relation between known layer properties (HS, Hi, ti) and the composite hardness H,. For thicker platings (t/d> 0.2) the indentation shape is increasingly distorted into the depth (Figs. l(c), (d)). A modification, or rather, generalization, of the thin plating theory, based on the concept of an effective substrate, will be given in the present paper. Investigations on the friction and wear behavior of this type of multiplated sliding contact have uncovered a subsurface deformation mechanism which forces underlying material to the surface, and gives the appearance of upsurgence 17, 81. When close enough to the surface, the substrate Cu is squeezed upward, giving rise to buckling of the intermediate Ni and the surface film, and penetration by the underlying copper. In extreme cases extrusions of the copper substrate itself were seen to appear on the surface, with very little wear of the surface layers. The present paper intends to show the fundamental relationships among the phenomena of hardness, friction Q 1993 - Eisevier Sequoia. All rights resewed

P. A. Engel et al. / Multiplated electrical contacts



(4 Fig. 1. Vickers microhardness indentations of multiplated media: (a) “thin” plating assembly: Cu (52.5 pm), Ni (1.7 pm), Pd70-Ni30 (1.7 wm), Au flash, top view; (b) cross section of (a); (c) “thick” plating assembly: Cu (87.5 pm), Ni (8.5 pm), Au (6.7 pm), top view; (d) cross section of (c). Top views originally magnified at 1000X, and cross sections at 800 X.

and wear in multilayer contacts. As a first step, we investigate the hardness under an indentation probe (such as the Vickers microhardness tester) which helps identity a depth-zone of large stress and significant plastic deformation occurring under a slider; the position of an “effective substrate” is thus found. We next introduce the concept of plastic work W, as applied to the typical multilayer contact subjected to a slider on top. The plastic work done on the substrate is a measure of its mobility. Mobility plays a decisive role in the tribological behavior of a contact [4, 71; both friction and wear are greatly influenced by the movement of substrate elements toward the surface. We therefore conveniently introduce plastic work as a computational tool to determine a measure of mobility of the substrate. The proportionality between plastic

work and the friction coefficient is subsequently shown by plotting the experimental and analytical results. The relationship between the friction, wear and subsurface mobility is demonstrated through experimental findings.

2. Experimental methods 2.1. vickers microhardness testing Multiplated electrical contact specimens [5] were used; their bases were epoxy-glass printed circuit board material, with a 25 pm peel-apart Cu layer attached to it; upon this base, Cu of variable thickness was plated. The total sandwich had subsequent layers plated upon this foundation, called a substrate. A variation


P. A. Engel et al. I Multiplated electrical contacts

of the plating qualities was induced by regulating the current intensity and duration of plating baths. The mechanical data (hardness Hi and thickness ti) of successive plating materials were checked. After sample cross sectioning, Knoop microhardness tests of the sectioned platings yielded their intrinsic hardnesses Hi, and thickness readings were enhanced by etching. Superficial Vickers microhardness tests were performed at load sequences of P= 10, 25, 50, 100, 200 g, automated for the tester. Two opposite Vickers diamond diagonals d, and d, for each sample were measured, and three sets of data averaged to yield a final value d. Since indentation reading difficulties at low loads tend to arise on account of surface roughness [5], only the 50-200 g load range was further used for the theoretical work. 2.2. Friction and wear experimental methods Friction and wear tests referred to in this paper were performed with a reciprocating ball-plane (modified Bowden-Leben) apparatus. In these tests the springs from a circuit-card edge connector [4] replaced the “ball”. The contact surface of the spring was spherical (R = 0.6 mm radius), with a nominal roughness of 6= 0.4 ,um CLA. The metallurgy of the spring was a Be-Cu substrate (H, =350 kg rnrr~-~) with three layers of platings. The innermost was a 1 pm Ni layer (H=400 kg mm-‘); the middle, a 2.5 pm Pd-Ni layer (H=400 kg mm-‘); the outer surface was a 0.25 pm soft Au layer (H= 90 kg nnre2). The plane specimens consisted of sections of Cu coated printed circuit boards, plated TABLE

2(a). Comparison

of measured,


with different combinations of Ni, Pd-Ni and Au layers. Descriptions of the hardness test results for the latter combinations are shown in Tables 1, 2(a) and 3. The roughness of these specimens was in the range of 6= 0.25-0.4 pm CLA. TABLE 1. Intrinsic and superficial materials [5] Material




Coating (Instrinsic) cu Ni 70P630Ni 80Pd-30Ni Soft Au Hard Au


(kgf mm-‘)


1225 2253 3919 3919 834.3 1958

125 230 400 400 85 200

177.6 326.8 568.4 568.4 121.0 284.0









Substrate (Superficial) 10 pm (400 pinch) Cu on PPCu (P=loO g) 961.2 22.5 pm (900 pinch) Cu on PPCu (P= 100 g) 1077 10 pm (400 pinch) Cu on PPCu (P=200 g) 777.8 22.5 pm (900 Finch) Cu on PPCu (P=200 g) 932.9

In the friction and wear apparatus the springs were mounted to a strain gaged beam which enabled both the normal load and friction to be measured. The plane

and iterated hardness (MPa):

load P200 g







t1 (w)


1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22

0.094 0.151 0.442 0.090 0.179 0.114 0.396 0.084 0.235 0.341 0.528 0.110 0.272 0.396 0.560 0.126 0.335 0.163 0.336 0.082 0.321 0.165

1222 1450 1590 1234 1423 1489 1739 1193 1455 1703 1801 1081 1484 1888 2073 1193 1714 1436 1732 1196 1623 1484

1070 1149 1360 984 1151 1103 1332 1028 1196 1318 1421 1086 1224 1356 1425 1093 1266 1144 1277

0.924 0.803 0.167 0.894 0.757 0.860 0.290 0.902 0.654 0.426 - 0.086 0.866 0.580 0.290 - 0.187 0.842 0.441 0.784 0.438 0.905 0.473 0.781

1240 1257 1875 1119 1263 1177 1740 1106 1353 1649 1916 1155 1426 1793 2058 1161 1542 1244 1561 1074 1516 1245

2.0 5.0 8.5 2.0 1.8 3.5 7.0 1.8 1.8 10.6 11.2 2.5 2.5 12.0 11.0 2.9 4.5 5.2 5.8 1.6 4.2 5.6

3.0 3.0 9.0 3.0 9.0 3.0 9.0 3.0 9.0 3.0 9.0 3.0 9.0 3.0 9.0 3.0 9.0 3.0 9.0 3.0 9.0 3.0

1283 1154





0 0 0 0 0 0 0 0.12 0.12 0.12 0.12 0.12 0.12 0.12 0.12 0.75 0.75 0.75 0.75 0.75 0.75 0.75


1110 1236 1690 1110 1292 1172 1581 1099 1292 1501 1977 1130 1328 1580 1943 1142 1430 1243 1503 1090 1417 1232

P. A. Engel et al. / Multiphted electrical contacts TABLE

2(b). Composition

substrate: Cus22.5 No.

1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22

of test specimens for Table 2(a) (note

Soft Au

Hard Au



(elm (Pin))

(km ([email protected])

3.0( 120) 3.0( 120 9.0(360) 3.0( 120) 9.0(360) 3.0(120) 9.0(360) 3.0( 120) 9.0(360) 3.0( 120) 9.0(360) 3.0( 120) 9.0(360) 3.0( 120) 9.0(360) 3.0( 120) 9.0(360) 3.0( 120) 9.0(360) 3.0( 120) 9.0(360) 3.0(120)

2.0(80) 5.0(200) 8.5(340) 2.0(80) 1.8(70) 3.5(140) 7.0(280) 1.8(70) 1.8(70) 10.6(425) 11.2(450) 2.5(100) 2.5( 100) 12.0(480) 11.0(440) 2.9(115) 4.5(180) 5.2(210) 5.8(230) 1.6(65) 4.2(170) 5.0(200)


0 0 0 0 0 0 0 0.12(5) 0.12(5) 0.12(5) 0.12(5) 0.12(5) 0.12(5) 0.12(5) 0.12(5) 0.75(30) 0.75(u)) 0.75(30) 0.75(30) 0.75(30) 0.75(30) 0.75(30)

Ai = 1/2[d *- (d - 2. ~2bi)*]

bi=(Sin /3*COSP)‘ti


The application of eqn. (1) to ideally thin platings permits the approximation A,=P/Hs=d2J2=A

It has been shown [5, 61 that in multilayer thin film assemblies, where the total film thickness t satisfies t/ d < 0.2, good predictability for the composite hardness results from eqn. (l), an extension of Jonsson and Hogmark [9]: (i= 1,2,. . . n, #s)



3. Hardness of a thin multilayer assembly



For simplicity, by hardness H we shall mean load P divided by the projected indentation area A, thus the composite hardness H,=P/A. H, is the substrate composite hardness, A, the (projected) indentation area of the substrate; Hi and ti are the subsequent layers’ (intrinsic) hardnesses and thicknesses, respectively, and Ai the perimetral area of a layer (Fig. 2), which is subjected to the flow pressure of that material. It is calculated from the indenter angle (/3=22”), the indentation diagonal d, and layer thickness ti

specimen was mounted to a steel holder, which oscillated beneath the spring. Normal loads used in these tests ranged from 50 to several hundred grams; however, most data was generated using a load of 175 gf. In tests used to characterize wear behavior, a 0.25 mm stroke length was used at a repetition rate of 40 cycles min-I. While friction was monitored in these tests, additional test runs using a longer stroke length (2.3 mm) and a lower oscillation rate (10 cycles min-‘) were used to determine the coefficient of friction. The friction tests tended to be of shorter duration (e.g. 100 cycles), than the wear tests, (e.g. lo4 cycles). Wear and friction behavior for these two conditions were compared and found to be similar. Both lubricated and unlubricated tests were run; as a lubricant, either a polyoester or a perfluoropolyether oil were used. Additional details regarding these tests may be found in refs. 4, 7 and 8.



pm (900 pin))




which, when inserted in eqn. (l), produces a unique function of H, in terms of H,, Hi and ti; this is true throughout the practical microhardness load range (50 g QP Q200 g) because, for the considered type of platings deposited on a soft base, H, is a unique and monotonically decreasing function of P. Tests for intrinsic hardness properties of Cu, Ni, Pd-Ni, hard Au and soft Au plated specimens used in previous work [3, 5, 61 established the values of Table 1; the table also includes the superficial hardness ranges of two Cu substrates, one a 22.5 pm thick, the other a 10 pm thick Cu plating both deposited on peelapart Cu. Sample test combinations of the plating materials were then subjected to hardness testing, and the measured and analytical values compared [5, 61.

4. General hardness theory for thick platings In ref. 6, a thick plating configuration of tl = 7.5 pm Ni deposited on a Cu substrate and a t2 = 7.5 pm thick hard Au surface film on top of the Ni were considered in light of eqn. (1). Under P=200 g load, the experimentally measured diagonal was d=47.0 pm. Thus t/d=(7.5+7.5)/ 47 = 0.319 > 0.2. The measured H,, m= 181.1 kgf mm-* was then much larger than the analytically achieved value (by eqn. (l)), H,, a= 126.0 kgf mm-*. The discrepancy is readily explained: it was unrealistic to expect the faraway, soft Cu substrate to participate in load carrying to the extent of being imprinted parallel to the Vickers indentation. For similarities, see Fig. l(d). If, however, the substrate was assumed to be the Ni material (thus eliminating the underlying Cu from con-


P. A. Engel et al. / Multiplated electrical contacts

TABLE 3. Plated configuration (note: Pd-Ni is 70Pd-30 Ni) Specimen

(from top down) studied for friction and plastic work; P= 17S9, R = 0.6 mm; substrate:


H, (computed)

D 0’ pm)

Cu 22.5 pm

us, (lO-‘j N mm)








Gold series (a) Hard Au Ni

3.05 2.54

120 100



(b) Hard Au Ni

3.05 7.62

120 300






(c) Hard Au Ni

9.14 2.54

360 100






(d) Hard Au Ni

9.14 7.62

360 300






Pd-Ni series (e) Soft Au Pd-Ni

0.51 5.59

20 220






(f) Soft Au Pd-Ni Ni

0.51 5.59 2.54

20 220 100






(g) Hard Au Pd-Ni

0.51 5.59

20 220






(h) Hard Au Pd-Ni Ni

0.51 5.59 2.54

20 220 100






sideration, and taking the intrinsic hardness of Ni as the substrate superficial hardness), then the measured composite hardness H,, m= 181.1 kgf mm-’ turned out to be substantially lower than the computed value, H,, a= 225.0 kgf mm-‘. The reason may be sought in that this time the substrate taken into account was excessively hard. To obtain the right position of an “effective substrate”, the influence of its depth under the indenter is sought now. It stands to reason that within that depth the likes of eqn. (1) would operate. Finite element methods can also help identify such a highly plastically stressed region; in ref. 5, for example, a multilayer sandwich similar to the Cu/Ni/Au example above was computed. By considering the equilibrium of the assembly, Fig. 2(d), it is upset by a value A,=A; from the figure we get P acting downward, and H, *A, + ZZHi.Ai acting upward, so that ZHi .Ai is an extra force. Let us, however, set A,[email protected]


to achieve the value of A, needed for equilibrium. A method which would establish a function [email protected]) for substrate stress VS.depth sensitivity as follows. Suppose the load is fanned out from the interior quarter-point of the crater according to Fig. 3; then by eqn. (6) the plastically deformed area at depth z satisfies


I,+) =A,(z)/A = 1 - (z/d). (1 + [Uf])


We take only positive regions of $, and so the plastic area is terminated at the depth z = d/2, where eqn. (7) yields $=O. The equilibrium of forces acting on the indentation region is satisfied if P= *Hi

+ ZHi .Ai

and at the plastic boundary z=d/2 P=ZHi.Ai

(8) we get (9)

In Table 2 measured plating configurations were evaluated for $ by noting the respective d values (Fig. 4); substituting these into eqn. (7) we then computed A,= @A, and these values were used in eqn. (1). Note that for negative I+VS obtained from eqns. (7), $= 0 was substituted. The agreement over the “thin plating” results improved dramatically, as demonstrated by Figs. 5(a), 5(b). This generalized “thick plating theory” can also be used in a purely analytical, predictive mode, based on the following iterative scheme. 1. Establish a trial indentation area A =d2/2, by using a trial composite hardness H, equal to the superficial hardness H, of the substrate: d = (2PIHc)1’2

which yields the trial f/d ratio.



P. A. Engel et aL / Multiplated electrical contacts








l-+----d4 @)


Fig. 2. Schematics of Vickers indentation in a multilayered thin plating (t/d < 0.2): (a) cross section through main diagonal; (b) plan view at the 1st layer; (c) free body of the 1st layer; (d) approximate free body diagram of the multilayer assembly above the substrate.


Fig. 3. Illustration of plastic stress 2)s. depth distribution: cross section at depth z= t; (b) the function tit/d).

2. Choose the effective plastic stress boundary at a depth Z = d/2, and locate it in the multi-layer geometric scheme. This position falls either in the (Cu) substrate (step 3), or in one of the parallel films plated onto it (step 4). 3. If Z >t (Fig. 4(a)), the plastic stress starts in the substrate. Now evaluate ~9at the top of the substrate I = t by eqn. (7), yielding A, at that point, by eqn. (6). Since H,, Hi, and A,, Ai are known, H, can now be calculated by eqn. (1). 4. If Z

H, is now chosen as the intrinsic hardness Hj of the jth layer, and H, can be evaluated by eqn. 1. 5. The H, value, obtained from either step 3 or 4, is now used in eqn. (10) (step l), and the cycle of calculation is repeated until satisfactory convergence is reached. 4.1. Example 1 We may. follow the computational aspect of the iterative procedure leading to an analytically established composite hardness for a “thick plating” (i.e. t/d > 0.2). Let us consider the Cu/Ni/IIard Au sandwich (Ctr, 22.5 pm plated on peel-apart Cu; Ni, t1 = 7.5 pm; Au, fz = 7.5 pm), described in ref. 5 and at the beginning of this section, indented by a P= 100 g load.


P. A. Engel et al. I Multiplated electtical contacts 2400.0




: a.!




? ” 7 u

Fig. 4. Two cases for plastic boundary position: (a) Z > t: plastic boundary in the substrate; (b) 2
By the iterative procedure outlined in Section 3, the initial guess for d is obtained by equating H, with H. = 1958 MPa. In the first step we get d={(2x lOO)/ [lo0 x 1958]}lD= 0.03196 mm. Thus, flu’= 0.00151 0.003196=0.469, and the indentation area is A = 0.03196’/2 = 0.0005107 nun’. As the second step, we choose the plastic boundary at 2 =d/2= 0.01598 mm; it thus falls in the Cu since Z>t. In step 3, we calculate tJ(z=t) from eqn. (7), and get the value 0.09108. Thus A,= +A = (0.09108). (0.000051072) = 0.0000465 mm2. A, =A, =0.0002083 mm2 is obtained from eqn. (2) making A, = (46.5 + 2 X 208.3) X 10m6 mm’. Now by eqn. (1) we get the refined value of H, =2002.5 MPa, to restart the iteration cycle. Convergence was achieved after four cycles, yielding H, = 2027 MPa, to compare with the measured composite hardness of 2097 MPa, a 3.4% discrepancy. Table 2(a) shows, for 22 thick plating configurations (Table 2(b)), a comparison of four corresponding composite hardnesses H, at P = 200 g: those values measured (H,); those evaluated by thin plating theory (Hc) applied to eqn. (1); those evaluated by thick plating theory (4 evaluated at the Cu substrate) applied to eqn. (1) and by use of the measured diagonals (H,,); and finally, the H(Z) values obtained analytically, by the above iterative procedure. A more rigorous treatment of the plastic boundary concept is under work.

5. Plastic work The experience attained in Vickers microhardness tests can assist us in describing affairs under an arbitrary-





















r /













measured hardness (best fit) 00000 t/d < 0.2 (thin plating theory) q case

z E

d 2000.0






z z z ii a s cu “7 2 cl?i

fa 2k 80








I’~“‘~I1’/~I’1’~I~IlIII’1’1I~I 0.3 0.4 0.5





Fig. 5. Hardness& calculated by three methods, in 23 multilayered plating configurations (load, 200 g): (a) H, vs. the measured hardness; (b) H, vs. the t/d ratio.

shaped, say spherical, indenter as well. In particular, great interest is attached to a popular type of sliding contacts with a spherical head (the “Hertz-dot”) used in electrical connectors [2-4, 6-81; as we shall see, the tribological (friction and wear) roles of such contacts are strongly related to their indentation behavior. We shall introduce plastic work as a concept and qualitative measure useful in describing indentation conditions in a multilayered contact. By plastic work we shall mean the energy dissipation during the deformation of a plastically flowing contact. The parallel layers that have entered the plastic region are assumed to be in the collapse state, and suffer ever increasing flexural deformations which include rotation of “plastic

P. A. Engel et al. I Mdtiplated

hinges” [lo] at the perimeter of the deforming plug of material, Fig. 6. Let us define the plastic work Ui of layer i as the work done at the plastic hinge by rotation ei of the originally horizontal layer against the full plastic moment Mp. i: Ui=Mp,ie2m*8i


A constant radius a for all the parallel “plastic disks” is taken, because a flexural rather than compressional contact stress effect is targeted. An expression of the plastic plate moment M,,, i for homogeneous, isotropic, linearly elastic/perfectly plastic disks is readily obtained (eqn. (12)). Table 1 gives the intrinsic hardness Hi of materials used as the test specimens; the yield stresses a,,, i were taken as one-third of the respective Hi values, using an approximate general rule between yield stress and flow pressure [ll]: A4p,i=






The small rotation of each parallel disk follows from the indenter’s radius of curvature R: ei = u/(R + ~tj + ti/2)


The disk radius is calculated composite hardness Z-I,:

from the load P and

a = (PIdQ’”


Suppose the boundary Z of the plastically deforming region falls within the substrate where it defines a disk B IP



electrical contacts


of the substrate having a thickness t,=Z_8ti


The plastic work U, of this disk can then be calculated from eqn. (12) as uY,S*t,‘/4.

6. Friction Several platings, in increasingly thicker (and thus, stiffer) order, have been studied [4,7] for their friction behavior. Each sandwich is based on Cu, with subsequent platings of Ni and, on top, either Au or Pd-Ni. Table 3 lists dimensions of these specimens. The table gives the initial friction coefficient CLachieved when slowly (0.8 mm s-l) passing a Pd-Ni slider (with a 0.25 pm Au flash) of spherical radius R = 0.6 mm under P= 175 g load. The value of plastic work of the substrate (U, = U,, per pass), in addition to the composite elastic plate rigidities D (see Appendix B) of the total platings deposited on a Cu substrate are also noted in consecutive columns of Table 3. Photographic illustrations of the effect of repeated sliding passages over the multiplated contacts are shown in Fig. 7. Figure 8 shows a schematic of the process; the Cu substrate tends to surge upward, and if the platings are thin and weak, they eventually buckle and penetrate the surface. The surface protuberances, resulting from the upsurgence, will reduce conformance of the contacting surfaces on the microscopic level, tending to reduce the real area of contact. As a con-



J= -l)t’ Q







1 Z-I


Fig. 6. Indentation of a multilayered medium by spherical indenter: various layers.

04 deformation pattern; (b) plastic hinges hypothesized in


P. A. Engel et al. / Multiplated electrical contacts



Fig. 7. Examples of upsurgence (from ref. 7): (a) localized movement of Ni and Cu with local Ni exposure: Au surface film: (b) similar system with local exposure of Cu; (c) and (d): similar features as (a) and (b), but with Pd-Ni surface film.

sequence, subsurface deformation will result in reduced friction. Behind the empirical evidence connecting a smaller friction coefficient with a weak plating lies the supporting role of the Cu substrate. We shall present two arguments for the interaction of substrate and plating affecting the friction. The first argument is based upon the plastic work concept; the second one treats the assembly as an elastic one, and computes the elastic work performed on the plating and substrate. From the values of I_Land Ucu listed in Table 3, the diagram of Fig. 9 was plotted. A robust relationship for p vs.l/U, is evident for both Au and Pd-Ni surface films. In both cases the depth of plastic influence passed higher and higher towards the surface of the Cu substrate

as the platings grew thicker, as shown in Table 3. In contrast, no significant variation of the total plastic work of the contact was noted with plating geometry variation. Since increasing plastic work of the substrate tends to break up the surface, it destroys a tendency for conformance between the neighboring sliding surfaces, causing a decrease of the friction coefficient. As a crude approximation, we may also model the contact as an elastic plate (the plating) supported on an elastic “Winkler”- foundation (the Cu substrate). In Appendix B this analysis was performed, for a rigid spherical indenter (R) pressing down on an area_4 = razz of the plating in axial symmetry. The flexural strain energy V,, in the plating and the work performed on the foundation US, are found:

P. A. Engel et al. / Multiplated electrical contacts -MOTION

to be the sum of thinning and subsurface deformation. It was also found that wear depth increased with subsurface deformation. In some cases the wear depth resulted almost totally from deformation of the substrate. When the surface layers were sufficiently thick to eliminate subsurface deformation, the wear was solely of the thinning type [7]. In a composite model [S] for the wear of electrical contacts the simultaneous occurrence of surface wear mechanisms, which lead to thinning of the surface layers, and of subsurface deformation is considered. It is shown that the total wear for contact systems of the type considered can be described by a relationship of the following type:




(1- DIDo)=P=.~L= +~~%rt=~ N,.,.2

(1 - D/D,)P”%~,i, + K,,,

Fig. 8. Gradual development of the upsurgence process (from ref. 7).

Up, = ?rDa‘f2R2


Us, = Q.243ddlR2


The ratio Usu/UP, exhibits the relative magnitude of the substrate-foundation work with respect to that of the plating work: UJJ,,

= 0.121 kd/D = (0.01226kp2) - l/DH,



It is clear that both D and H, are inversely proportional to U,,lU,,. Thus, the elastically computed substrate work exhibits the same tendency as the plastic work did in the previous plastic analysis. Experimentally, the friction coefficient varies proportionally to D as shown in Fig. 9(b) for the platings described in Table 3. This characteristic variation is to be compared with the p ~,s. Us, relation.

7. Wear By wear we shall mean the removal of volume from the contact area. The concept of subsurface plastic work implies that there should be two contributors to wear depth. One is thinning of the surface layers by such mechanisms as abrasion, adhesion, surface fatigue or delamination; the second contribution would be substrate deformation or indentation. This is the condition that has been observed with the thin and weak surface layers, discussed earlier. Scar depths were found


where h is the wear depth; the K values are material wear coefficients for the subsurface and surface wear; D, the plate rigidity of the surface layers; P is the load. Do is identified as the minimum plate rigidity that eliminates subsurface deformation, thus making the substrate line up with the plastic boundary, 2. N is the number of cycles. Since the plastic energy dissipated in the substrate is inversely related to the stiffness, this relationship implies that wear is related to the plastic energy dissipated in the substrate, and decreases as the amount of that energy decreases. This is an inverse of the relationship found for friction.

8. Results and discussion The hardness of multilayer platings was interpreted on the basis of simple geometric models, suggested by physical symmetries. A correct interpretation of superficial hardness is needed in numerous industrial applications, such as acceptance procedures for electrical contacts. The feasibility of such an on-line quality assurance process, with the possibility of error estimation, has been explored in ref. 6, and is further enhanced by the present more general theory. The new, general multilayer-hardness model involves the location Z of a plastic boundary which demarcates the effective substrate. For homogeneous substrates, the existence of a circular plastic boundary was calculated and experimentally shown by Shaw and DeSalvo

[la Having dealt with the hardness properties, computation of the plastic work in the sandwich was the next task. The plastic work of the (Cu) substrate was found to be an indicator of the mobility displayed by the materials of the multilayered assembly. The plastic work of parallel flexurally deformed disks is at the heart of

P. A. Engel et al. / Multiplated electrical contacts

548 I





H Au + NI Sews S. AU + Pd - NI + NI Series H Au+Pd-NitNISerles




**. *.

0 F v) 2



d 0

*. .. 0 A) U,, vs. II






















1.0 0 0 -------A 0 _V 0

3 .



H. Au + Ni Series S. Au + Pd - Ni + Ni Series H. Au + Pd - Ni + Ni Series


B) D VS. p 1o-3


I 0.2





Fig. 9. Friction vs. plastic work of substrate,






and vs. plate-rigidity

the phenomenon, since this reflects the influence of the substrate upon the higher plating layers. Computed values of U,, for both a Cu/Ni/Au series and a Cu/Ni/F’d-Ni/Au series (Table 3) exhibited a robust inversely proportional relationship V.K the respective friction coefficients CL.Figure 9(a) illustrates this for both series. Simultaneously, an observed proportional relationship [4, 71, (see Fig. 9(b)) between I_Land the plate rigidity D of the total plating was also heuristically proven in the present work, by a model of a plate on elastic foundation. The resulting D vs.F diagram displays an ascending slope. The key to the above behavior is seen as increased mobility of the substrate leading to decreased con-



of plating: (a) Uc, vs. CL;(b) D vs. p.

formance of the interface on the microscopic level. This decrease in conformance tends to lower friction. In the meantime, the underlying substrate deformation leads to increased wear depth. The mobility, in fact migration, of material in plastically deforming but not necessarily multilayered contacts was noted by Wellinger and Breckel [13] in connection with the impact wear of copper specimens. Contact stress aspects and their relation to wear of such impacted contacts was considered by Engel [14]. The semi-empirical relationship of the wear of layered substrates VS. their hardness profiles was considered for engineering modeling purposes in ref. 15. The coefficient of friction of the typical electrical contact interfaces discussed tends to change with wear

P. A. Engel et al. I [email protected] electkal

[4]* For the meta~urgies studied, several trends were observed. When a lubricant was present, the coefficient of friction tended to decrease with wear. Under unlubricated conditions, a more complex behavior was often observed. Initially, the coefficient would tend to decrease; however, as wear progressed, the coefficient would often be seen to increase again. While the coefficient of friction changed in these ways, the effect of subsurface deformation and upsurgence on friction persisted. This is illustrated by the data shown in Table 4. These results indicate that the change of the coefficient of friction caused by subsurface defo~ation is in addition to any changes caused by other factors, such as wear or lubrication. Because of the effect that subsurface deformation has on friction, it would appear to be desirable to use thinner Ni layers and softer substrates in contact systems to reduce insertion forces. However, this is generally not the case as it is apparent from ref. 16 which proves Ni to be beneficial in fretting situations. In addition, we enlist the following arguments. As had been reported previously [7, S], the presence of subsurface deformation leads to premature exposure of the base metal underlayers. It also results in increased wear scar depth and a stronger load dependency than is observed when this mechanism is not present. All of this results in reduced wear life. In addition, while a lower coefficient of friction is desirable from the stand~int of insertion, it is undesirable in appli~tions where there is exposure to vibration. In the latter case a lower coefficient of friction tends to increase the amount of motion occurring, which results in increased wear. Higher coefficients of friction tend to reduce the amplitude of vibration-induced motions by providing increased damping. They also tend to increase the excitation threshold that is required for such relative motion to occur between contact interfaces. Both effects have been observed in practical applications. As a consequence, it is generally desirable to either use harder substrates or thicker Ni layers when there is concern with wear life and corrosion.

TABLE 4. Comparative fluoropolyether) Configuration

Initiai Final







was per-









0.5 1.0

0.5 0.4

0.4 0.5

0.28 0.20

0.20 0.16

0.18 0.11

“Config’n 1: 7.5 pm Ni, 7.5 pm Pd-Ni, 0.25 pm Soft Au. bConfig’n 2: 2.5 pm Ni, 3 Nrn Pd-Ni, 0.25 pm Soft Au. ‘Config’n 3: 3 pm Pd-Ni, 0.25 &m Soft Au.



9. Conelufions Great interdependence of the hardness, friction and wear properties of multilayered electrical contacts has been demonstrated. Layer configurations of larger thickness (t/d> 0.2) exhibit a plastic boundary which is si~~cant as for demarcating an effective substrate. The plastic work concept was introduced to compute the mobility of a substrate. This plastic work was seen to be inversely proportional to the friction coefficient, and proportional to the wear rate of the multilayer contact.

Acknowledgments Discussions with W. Lafontaine, R. Schumacher and R. Topa of the IBM Endicott Laboratory are acknowledged. K. R. Wu assisted with computations. The hardness work was done under contract with the IBM Endicott Laboratory.

References 1 J. L. Piechota, Electrical contacts and design, in D. P. Seraphim, R. Lasky and C-Y Li, (eds.), Principles of Electronic Packaging, McGraw Hill, New York, 1989, pp. 196-215. for zero 2 W. L. Brodsky, Testing and design parameters insertion force connector, in Proc 37th IEEE Electmraic Components Con& 1987, pp. 32-40. method to 3 P. A. Engel and D. L. Questad, Indentation measure plating ductility, ASME J. Electron. Packaging, 112 (3) (1990) 272-277. 4 E. Y. Hsue and R. G. Bayer, Metallurgical study and tribological properties of edge card connector spring/tab interface, IEEE Trans. PHI, 12 (2) (1989) 206-214. 5 P. A. Engel, A. R. Chitsaz and E. Y. Hsue, Interpretation of superficial hardness for multilayer platings, Thin Solid Films, 207 (1992) 144-152. evaluation of multilayer 6 P. A. Engel, Vickers microhardness platings for electrical contacts, in Froc. 37th IEEE Helm Conf. on Electrical Contacts, Chicago, IL, 1991, pp. 66-72. 7 R. G. Bayer, E. Y. Hsue and J. Turner, A motion-induced sub-surface deformation wear mechanism, Wear, 154 (1992) 193-204. 8 R. G. Bayer, A general model for sliding wear in electrical contact, in Proc. Wear of Materials Con$, San Fmncisco, CA, 1993. 9 3. Jonsson and S. Hogmark, Hardness measurements of thin films, Thin Solid Films, 181 (1984) 257-269. 10 C. R. Calladine, Plastici for Engineem, Wiley, New York, 1985. 11 D. Tabor, The hardness of solids, Rev. Phys. Technol., 1 (1970) 145-179. 12 M. C. Shaw and G. J. DeSaivo, ASME .T. Eng. Ind., (1970) 46w94. 13 K. Wellinger and H. Breckel, Kenngroessen und Verschfeiss beim Stoss Metallischer Werkstoffe, Wear, f3 (1969) 257-281. 14 P. A. Engel, Impact Wear of Matetiah, Elsevier, New York, 1976.

P. A. Engel et al. t Multiplated electrical contacts

550 15 R. G. Bayer and T. C. Ku, Handbook for Wear, Plenum, New York, 1964.

of Analyrical Design

16 M. Antler and M. H. Drozdowicz, Fretting corrosion of goldplated connector, Wear, 74 (1981) 27-50.

Appendix A: Nomenclature

1 4 AT bi d D E h H HC Hi H, k K sub K surf 4

N P P ; t ti u W

W W, z Z

contact radius indentation area plastically deformed area of layer i total plastic area width of ledge i Vickers indentation diagonal elastic plate rigidity modulus of elasticity wear depth hardness composite hardness intrinsic hardness of layer i measured hardness elastic foundation modulus subsurface wear coefficient (pm N-‘.6 cy~-~.~) surface wear coefficient (pm N-O.’ cy~-O.~) plastic moment number of cycles flow pressure indenting (normal) force radial coordinate radius of slider plating thickness thickness of layer i energy plate displacement wear plastic work depth position of plastic boundary

Greek letters Vickers indenter angle (22”) P 6 surface finish 8 rotation friction coefficient P yield stress a7 $, [email protected]) stress ZIS.depth sensitivity Subscripts s subscript denoting substrate

Appendix B: Approximate elastic contact analysis A rough elastic analysis of the plating supported on the substrate is performed as a problem of an axially symmetric plate on an elastic foundation. The approximate solution is needed in view of the great mathematical complexities [Al] of an analytical solution; however, the essential information on relative energy distribution in the contact will be attainable by the simple model. We indent the plate by a rigid, spherical indenter (R) which is assumed to put a uniform pressure p = H, over a circular area A = m2 of the surface, Fig. Al. For example, the composite rigidity of a two-plate composite is [AZ!]: E2t3 D= 12(1- v22) c2 C = 1 + E&3(1 - ~2~) 2 E2t23(1 - 2)


k . .


. . .


. .

Fig. Al. Model of plate (elastic plating) on elastic foundation





(Cu substrate).

P. A. Engel et al. / Multiplated elecm’cal contacts

while the elastic foundation of the Cu substrate is represented by the foundation modulus k (N mrne3). The displacement of the surface ,is wl(r) and of the foundation: w&). The axial compression of the plating is u(r), and we consider a truncated structure of radius a detached from the rest, Fig. Al. Writing the surface displacement w1= w* + (a”/ 2R)(l -~‘/a”), we get w,=w, --u. Here w* is the depression (sinking-in) of the contact area and u the axial compression of the plating. Since the midsurface of the plating has the same meridional curvature w”= l/ R as the indenter, the flexural strain energy of the plating can be evaluated. Meanwhile, we may neglect the compressional strain energy, and get for the plating


energy U,,,= Da 2(w’a)/2 = nDa “12R2


The foundation energy is calculated from the expression (l/2) 2mGl~,~(r) -r *dr, yielding U,, = 0.243&a 6/R2


References for Appendix B Al S. P. Timoshenko and S. Woinowski-Krieger, Theory ofPlates and Shells, McGraw Hill, New York, 2nd edn., 1959. A2 R. J. Roark and W. C. Young, Formulas for Stress and Strain, McGraw Hill, New York, 5th edn., 1975, p. 377.