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Sliding wear behavior of deformation-processed Cu-15vol.%Cr in situ composites
WEAR ELSEVIER
Wear 195 (1996) 214-222
Sliding wear behavior of deformation-processed Cu-lSvol.%Cr composites
in situ
Z. Chen a, P. Liu a, J.D. Verhoeven b, E.D. Gibson b aSchool of Technology, Eastern Illinois University, Charleston, IL 61920, USA b Ames Laboratory of USDOE. Department of Materials Science and Engineering, Iowa State UniversiTy, Ames, IA 50011, USA Received 17 May 1995; accepted 1 February 1996
Abstract A deformation-processed Cu-lSvol.%Cr in situ composite was made by consumable arc melting and casting followed by extensive deformation. A superior combination of mechanical strength and electrical/thermal conductivity was achieved with the composite since Cr filaments existed in the nearly pure copper matrix. The effects of sliding speed and normal pressure on sliding wear behavior and microstructure of the composite were investigated, with a composite pin rubbing against a hardened AISI 52100 steel disk on a pin-on-disk wear tester. In the studied range of normal pressure and sliding speed, the wear rate increased with increased normal pressure, whereas the wear rate decreased with increased sliding speed. Sliding-induced subsurface deformation occurred not only in the sliding direction but also in the lateral directions perpendicular to the sliding direction. This lateral flow produced a twisting of the Cr filaments. The complex deformation mode was revealed clearly by the morphological change of the ribbon-like filaments. Both constituents of the composite were cooperatively deformed. The thickness of deformed subsurface layer increased with increasing normal pressure and sliding speed. Scanning electron micrographs showed plastic deformation flow on the wear surface. Keywords: Sliding wear; Composite; Cu-Cr; Deformation processing
1. Introduction
Copper
refractory
metal
deformation-processed
in situ
provide a superior combination of high mechanical strength and high thermal/electrical conductivity. Cu-Cr composites were developed because Cr offers lower cost, better corrosion resistance, and higher modulus of elasticity. Verhoeven et al. [ 1] demonstrated that Cu-lSvol.%Cr alloy exhibited deformation characteristics similar to those of Cu-Nb and Cu-Ta alloys. Moreover, Cu-Cr was expected to produce higher strength at a given deformation strain than Cu-Nb because the strength of the composites was shown to increase with higher modulus of the refractory constituent [ 21. The remarkable malleability of the Cr phase, in spite of the brittle nature of pure Cr at room temperature, was first demonstrated by Funkenbusch et al. in Cu-17vol.%Cr alloys drawn into long filaments [ 3,4]. Through appropriate control of initial microstructure, the cast alloy could be drawn to a true drawing strain of 6.3. The tensile strength of Cu17vol.%Cr in situ composite increased with increased true drawing strain, reaching 840 MPa at a true drawing strain of composites
such
as Cu-Nb,
Cu-Ta
and Cu-Cr
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6.3, which was well above that predicted by the rule of mixtures. Kim et al. [5] studied the effect of thermomechanical processing techniques on strength and electrical conductivity of Cu-7vol.%Cr in situ composite wire. It was found that the optimum combination of strength and electrical conductivity could be obtained by the process of DSDAD (deformation, solution treatment, drawing, aging and drawing). A combination of a strength of 925 MPa and an electrical conductivity of 7 1.8% IACS (International Annealed Copper Standard) was achieved, which is superior to commercially available copper alloys. Hardwick et al. [ 61 studied the effect of annealing on the microstructure and mechanical properties of a Cu-Cr microcomposite sheet deformed to a true strain of 6.6. After annealing for 10 h at 650 “C, extensive spheroidization of the Cr ribbons occurred. As a result, yield strength was reduced by 55%, and the ductility increased markedly. Failure in the CuCr composite occurred by a mixture of ductile rupture and cleavage. Friction and wear behavior of metal-matrix composites has attracted considerable interest, most of which is on nonmetallic fiber, particle or whisker-reinforced composites [ 71.
2. Chen et al. /Wear
No investigation of friction and wear behavior of Cu-Cr in situ composites is available. Only a few studies have been conducted for friction and wear behavior of Cu-Nb in situ composites and precipitation-hardened Cu-Cr alloys. Liu et al. [ 81 studied the effects of Nb composition and true deformation strain on wear rate and coefficient of friction. A CuNb in situ composite pin was slid against a hardened 02 tool steel disk at sliding speeds from 0.028 to 2.5 m s- ’and under a normal pressure of 0.68 MPa. It was found that Cu20vol.%Nb showed the best wear resistance. Increasing true deformation strain resulted in increased tensile strength and coefficient of friction and decreased wear rate. The tribological behavior of the Cu-2Ovol.%Nb in situ composite was investigated in terms of sliding-induced plastic deformation flow, effects of Nb filament orientation, sliding speed and annealing temperature [ 91. The wear rate of the specimen with Nb filaments perpendicular to the sliding direction was lower than that with Nb filaments parallel to the sliding direction, which indicated that this in situ composite material had an anisotropic wear resistance as modeled by Hornbogen [ lo]. With increased annealing temperature, the hardness of the composite decreased, and both the coefficient of friction and the wear rate increased. It was found that both the Cu matrix and the Nb filaments underwent sliding-induced deformation with no loss of bonding between the Nb filaments and the Cu matrix. Increasing sliding speed resulted in a lower wear rate, which the authors ascribed to oxidation and work hardening. Nayeb-Hashemi et al. [ 111 investigated the friction and wear behavior of a powder metallurgical Cu-Nb microcomposite rubbing against gray cast iron of hardness 92 RB. They found that the wear rate increased drastically with sliding speed and the coefficient of friction seemed to be independent of fiber orientations. Suh [ 121 studied the tribological behavior of precipitationhardened Cu-O58at.%Cr and Cu-O.8 lat.%Cr alloys aged for different periods at 500 “C. The alloy cylinder rotated against a stationary AISI 52100 steel rod. The hardness of these materials initially increased with aging time from 637 MPa for solid solutions to 1372 Mpa for peak-aged alloys, and then decreased. It was found that the wear rate initially decreased with aging time by a factor of 3 for both Cu-Cr alloys and then increased approximately linearly. The minimum wear rate did not correspond to the maximum hardness. However, the coefficient of friction was fairly constant for all treatment times for both alloys. They attributed these results to the combined effects of hard particles on the hardness and the coherency of particles to the matrix. The dry sliding wear behavior of Cu-lSvol.%Cr in situ composite has been studied in this paper. The effects of normal pressure and sliding speed, the mechanism of subsurface deformation as well as the microstructural change due to sliding are discussed. The thickness of the subsurface deformation layer and microstructural change have been quantitatively analyzed in relation to normal pressure and sliding speed.
195 (19%) 214-222
215
2. Experimental details
Fig. 1 shows the flow chart of the material preparation process and the corresponding microstructure for the CulSvol.%Cr in situ composite. A Cu-lSvol.%Cr alloy ingot was first prepared by the vacuum induction melting method, followed by chill casting. An electrode formed from this casting was then consumable arc melted into a copper mold in order to obtain fine dendrites of Cr in the Cu-Cr alloy ingot. The deformation process started with hot extrusion of an ingot of 102 mm diameter into a rod of 5 1 mm diameter. Then the rod was rolled into a plate of 12.7 mm thickness and finally into a sheet of 1.6 mm thickness. The total true deformation strain was q= 5.4 or a reduction in area of 99.5%, which resulted in a hardness of 74 RB. Dry sliding tests were performed on a pin-on-disk wear tester, with the Cu-lSvol.%Cr in situ composite pin rubbing against a hardened AISI 52100 steel disk. The steel disk was austenitized at 843 “C for 30 min, oil quenched, and then tempered at 150 “C for 1 h. The resultant hardness of the disk was between 58.5 and 62 RC. The pin was cut from the CulSvol.%Cr in situ composite sheet and had a rectangular section of 1.57 X 10 mm. During the wear test, the 10 mm length of the specimen was in the sliding direction and the Cr filaments were perpendicular to the sliding plane. The disk rotated at 600,800,1000, and 1200 rev min- * with an average wear track diameter of 88.5 mm, i.e. the sliding speed ranged from 2.78 to 5.56 m s-i. The normal load varied from 0.98 to 6.86 N, which corresponded to a normal pressure of 0.06-0.44 MPa over the apparent contact area of 15.7 mm*. Before wear testing, the specimen was ground on emery paper to 600 grit in running water, and then was cleaned in methyl alcohol for 5 min using an ultrasonic cleaner. The disk was cleaned with methyl alcohol. The pin was weighed every 5 min with an accuracy of lo-’ g. Processing
Microstructure
I
Consumable arc
Dendritic Cr in Cu/Cu-Cr alloy
melting into water cooled copper mold
m&ix
Deforming: Hot extrusion
I Deforming: Rolling
’
Fig. 1. Flow chart of composite StruCtUWS.
Cr tilamcnts in cldcu-Cr alloy matrix
processing
and corresponding
micro-
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216
195 (1996) 214-222
slidingDkction 4 I
/
LongitudinalSe&on
TransvcxscSection
Fig. 2. Schematic diagram of nomenclature rolling and sliding directions.
definition for specimen sections,
3. Results and discussions
the rolling and sliding directions are illustrated in Fig. 2. Sliding took place on the transverse section. The microstructure of the Cu-lSvol.%Cr in situ composite before sliding is shown in Fig. 3. Fig. 3(a) shows the microstructure taken from the face section. The Cr filaments, with variable ribbonshaped cross-sections, align in the rolling direction. Fig. 3 (b) illustrates that some Cr filaments with coarse knots are present in the longitudinal sections, which indicates that Cr dendrites were not uniformly deformed into thin ribbons. Some Cr phase can also be seen as spherical precipitates in the copper matrix. 3.2. Effect of normal pressure on wear
3.1. Microstructure
of Cu-lSvol.%Cr
in situ composite
The ingot was deformed with a true deformation strain of 5.4, which caused the original Cr dendrites to turn into ribbonlike filaments aligned in the rolling direction. The definitions of the composite specimen sections and relationship between
Fig. 3. Microstructure
of as-processed
Cu-lSvoI.%Cr
in situ composite:
The variation of volume loss of Cu-lSvol.%Cr in situ composites with sliding distance under different normal pressures and at a sliding speed of 3.70 m SK’ is shown in Fig. 4. The volume loss increased with increasing sliding distance, with a rapid increase during the running-in stage and an
(a) face section; (b) longitudinal
section; (c) traverse section
Z. Chen et al. /Wear 195 (1996) 214-222
0
2
4
6 g 10 sliding Distance (km)
12
ot,““““‘,““‘,“‘,“,i 0 0.1
14
Fig. 4. Variation of composite volume loss with sliding distance at a sliding speed of 3.70 m s-’ and under four different normal pressures.
0.2 0.3 Nomal Rcssurc (MPa)
10
20
30 Sliding Tii
40 bin)
Fig. 5. Variation of wear rate of the composite sliding speed of 3.70 m s-l.
50
60
pin with sliding time at a
almost linear increase in the following steady-state stage. Under a normal pressure of 0.44 MPa, the total volume loss was much greater than that under lower normal pressures. A total wear rate can be defined as total volume loss per sliding distance [ 131. The variation of the total wear rate with sliding time is illustrated in Fig. 5. The total wear rate decreased with sliding time, approaching a steady state. It may be seen in Fig. 4 that increasing the pressure from 0.31 to 0.44 MPa produced a sharp increase in initial wear. This transition is illustrated nicely in Fig. 6 which plots total wear rate after a 60 min sliding time vs. the normal pressure. It is seen that a sharp transition from mild wear to severe wear has occurred when the pressure was raised from 0.31 to 0.44 MPa. These data show that the material should be maintained at a normal pressure below 0.44 MPa to avoid severe wear while sliding at a speed of 3.7 m s- ‘. 3.3. Effect of sliding speed on wear The change of the total wear rate with sliding time under a normal pressure of 0.31 MPa is shown in Fig. 7. Fig. 8 shows the variation of the total wear rate with sliding speed under a normal pressure of 0.31 MPa. The total wear rate decreased with increased sliding speed, but it tended to level off at a sliding speed of 4.63 m s-‘. 3.4. Microstructural
change due to sliding
By studying the change in the alignment of the Cr filaments near the wear surface, it was possible to evaluate the nature of the deformation layer produced at the surface. Fig. 9 illus-
0.5
Fig.
6. Variation of steady wear rate with normal pressure for Cu-lSvol.%Cr composite slid at a sliding speed of 3.70 m SK’. 10
0
0.4
ll..,ll,l,.llllllll(,...,,,.,,,~
0
10
20
30 40 Sliding Time (ah)
so
60
Fig. 7. Variation of composite pin wear rate with sliding time under a normal pressure of 0.31 MPa.
02
2.5
3
3.5 4 4.5 Sliding Speed (m i’)
5
5.5
6
Fig. 8. Variation of steady wear rate with sliding speed under a normal pressure of 0.31 MPa.
trates an ideal deformation pattern caused by sliding, which is depicted by changes in the Cr filaments. The deformation layer had a maximum thickness at the faces of the pins, and the layer thickness decreased symmetrically toward the centerline (C.L.) of the pins, which could be expected from simple contact mechanics since a flat punch had maximum stresses at the edges. Fig. 10 shows the face section microstructure of the composite below the wear surface after sliding at a speed of 3.70 m s-’ and under normal pressures of 0.06 MPa and 0.44 MPa. No debonding occurred between the Cu matrix and Cr filaments. The sliding-induced plastic deformation layer is revealed by changes in alignment of the Cr filaments. Under a lower normal pressure the majority of Cr filaments were bent in the sliding direction, as shown in Fig. 10(a). Under a higher normal pressure, however, Cr filament refinement was seen as a major characteristic (Fig. 10(b)). In addition to producing a flow in the sliding direction within the deformation layer, the sliding force also produced
Z. Chen et al. / Wear 195 (I 996) 214-222
218
Slldbg DhCth
Loqltodbd wctbn
CbCkWke Twkt off3
nbment
Fig. 9. Schematic of subsurface plastic deformation pattern on CulSvol.%Cr composite due to sliding. a flow in the lateral directions defined on Fig. 9. This flow was revealed by a twisting of the filaments demonstrated in Fig. 11, which shows two matched sectional micrographs of Cu-lSvol.%Cr in situ composite sliding at a speed of 3.7 m s-’ and under a normal pressure of 0.31 MPa for 60 min. Fig. 11 (a) was taken on the face section and Fig. 11 (b) on the longitudinal section. The sliding direction in Fig. 11 (a) was from the left to the right, as shown by the arrow, whereas the sliding direction in Fig. 11 (b) was from the page toward the reader. The pair of micrographs in Fig. 11 are matched along the edge located at the upper left corner of Fig. 9. The magnification is sufficiently high that only the wear affected layer appears in the micrographs, so that Fig. 11 (a) corresponds to the lower portion of Fig. IO(b). Fig. 11 (b) illustrates the twisting action produced by flow in the lateral direction. As shown in Fig. 9 the lateral flow produced a twisting action in opposite directions on the two sides of the rectangularly shaped pins, a clockwise rotation near the left face and a counter clockwise rotation near the right face. The twisting of the filaments was studied with optical and scanning electron microscopy and Fig. 12 presents a model of the shape change involved in the twisting.
3.5. Thickness of deformed subsurjace layer The thickness of the deformed subsurface layer was defined as the distance between the zone of unchanged microstructure to the wear surface on the face sections, as shown in Fig. lO( b) . Fig. 13 shows the variation of the thickness of deformed subsurface layer with normal pressure at a sliding speed of 3.7 m s - ‘. The thickness data were an average over 18 equal-distance measurements along the entire contact length of the pin in the sliding direction. The thickness increased monotonically from 52 to 246 pm, when normal
Fig. 10. Microstructural change on face sections of Cu-lSvol.%Cr in situ composite pin after sliding against a hardened AISI 52100 disk at a sliding speed of 3.70 m s-’ for 1 h, (a) under a normal pressure of 0.06 MPa; (b) under a normal pressure of 0.44 MPa. The arrow indicates the sliding direction. increased from 0.06 to 0.44 MPa. The rate of thickness increase with normal pressure decreased as normal pressure rose. The variation of the thickness of deformed subsurface layer with sliding speed under a normal pressure of 0.31 MPa is shown in Fig. 14. The thickness of deformed subsurface layer increased linearly with sliding speed. At a sliding speed of 2.78 m s-‘, the thickness of subsurface layer was 159 pm,
pressure
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195 (1996) 214-222
Fig. 11. Micrograph
of Cu-lSvol.%Cr in situ composite pin, after sliding at a speed of 3.7 m s-l and under a normal pressure of 0.31 MPa for 60 min: (a) face section; (b) longitudinal section. The arrow indicates the sliding direction.
0.1
0
0.2
0.3
0.4
0.5
NormalI’msun (MPa)
Fig. 13. Variation of thickness of deformed pressure at a sliding speed of 3.7 m s-r.
subsurface
layer with normal
Sliding Direction
*zeal
Direction
Fig. 12. Schematic model of spatial shape change in ribbon-like during sliding wear process.
Cr filament
and at a sliding speed of 5.56 m s-‘, the thickness was 309 km. From Fig. 13 and Fig. 14, it can be seen that both increased pressure and sliding speed increased the thickness of the deformed subsurface layer. The deformed subsurface layer was caused by the friction when the composite pin slid against the hardened disk. The measured coefficient of friction decreased slightly with increased normal pressure and sliding speed. Generally, increasing normal pressure would cause
2.5
3
3.5
4 4.5 5 Sliding Sped (m 6’)
Fig. 14. Variation of thickness of deformed speed under a normal pressure of 0.31 MPa.
5.5
subsurface
6
layer with sliding
more asperities of the composite to contact with those of the disk and hence increase the friction force even though the average coefficient of friction decreased slightly, which would increase the opportunity for plastic deformation of the composite. The real plastic deformation of Cr filaments in the composite was caused by the local stress. Increasing the applied normal pressure certainly increased the asperity stress
220
Z. Chen et al. /Wear
that caused a larger extent of plastic deformation. Since increasing sliding speed increased surface temperature which was resulted from friction heat and deformation energy, the pin material became softer at higher sliding speed than at lower sliding speed. The temperature increase caused by increased sliding speed would facilitate plastic deformation. 3.6. Study of wear mechanism Although crystalline Cr has brittle characteristics at room temperature [ 31, there was a deformation flow of the composite both in the sliding direction and in lateral directions, as shown in Fig. 10 and Fig. 11. The Cr filaments demonstrated good strain compatibility with the Cu matrix during fabrication of the composites. It is not surprising, therefore, that the composite displayed a good strain compatibility between its constituents during sliding and that no debonding phenomenon was observed between the Cr filaments and the Cu matrix. This sliding-induced deformation found in the Cu-Cr system was similar to that found in the Cu-Nb system [141. The slip mechanism occurring during the lateral twist can be explained with the favored crystal slip systems depicted in Fig. 15. Since ribbon-like Cr filaments were embedded in a softer Cu matrix, plastic deformation of the Cu-lSvol.%Cr in situ composites was controlled by the deformation mode of the harder Cr filaments. According to Hardwick et al. [ 61, a texture was developed in the Cr filaments at the true strain of 6.6, though the texture was not as strong as those found for Nb in Cu-lSvol.%Nb and for Ta in Cu-lSvol.%Ta. The main rolling texture for body-centered cubic Cr could be identified as (001) [ 1lo]. The ( 110) crystal plane of the Cr filaments lay in the sliding surface, and the crystal directions [ ill ] and [ ii1 ] were 45” away from the sliding direction. Therefore, both ( 110) [ i 111 and ( 110) [ ii 1] slip systems could be activated during sliding. Slip motion 45” away from the sliding direction in the Cr filaments could account for both elongation in the sliding direction and twist in the lateral directions. As illustrated by comparing Fig. 10(a) with Fig. lO( b), the deformation layer was less extensive under lower normal pressure than under higher normal pressure. It appears that Fig. lO( a) shows the beginning of plastic deformation of Cr filaments, while in Fig. lO( b) the plastic deformation of Cr filaments is well developed in both the sliding and lateral
Fig. 15. Schematic composite.
crystal
deformation
model for Cu-lSvol.%Cr
in situ
I95 (1996) 214-222
directions. The figures confirm that increasing normal pressure causes more plastic deformation in the contact subsurface region. The scanning electron micrograph of the wear surface shown in Fig. 16 also reveals the plastic deformation flow on the wear surface. This surface deformation flow was affected by wear debris and transferred material. Fig. 17 shows a micrograph of the composite microstructure located just below the wear surface. The specimen was sectioned slightly below but parallel to the wear surface. The black regions indicate the presence of wear debris, back-transferred material or oxide, on the surface. The orientation of the Cr filaments clearly shows that the surface deformation flow was affected by the transferred material. The complex flow pattern in the deformation layer revealed by the change in orientation of the Cr filaments is quite interesting. According to contact mechanics, apparently flow in the direction of the applied force on the wear surface is a maximum at the faces of the pin and decreases to a minimum at the center of the pin. Because the broad faces of the Cr
Fig. 16. Scanning electron micrograph revealed plastic deformation flow, after sliding at a speed of 3.7 m s- ’and under a normal pressure of 0.3 1 MPa for 30 min. The arrow indicates the sliding direction.
Fig. 17. Embedded wear debris inhibited the plastic deformation flow. The specimen was slid at a speed of 3.7 m s-’ and under a normal pressure of 0.19 MPa for 60 min. The arrow indicates the sliding direction.
Z. Chen et al. /Wear 195 (19%) 214-222
221
Fig. 18. Transfer films in the steel disk: (a) the original disk surface; (b) sliding direction normal to the ground groves and (c) sliding direction parallel to the ground groves of the steel disk, sliding at 3.70 m s-’ under a normal pressure of 0.06 MPa. The arrow indicates the sliding direction.
filaments are aligned parallel to the pin faces the bending moment inertia of the filaments in the applied force direction is much higher than it would be perpendicular to this direction. It seems likely, therefore, that the twisting of the filaments occurs because bending in the lateral directions requires a very low stress. Hence, one would predict that the wear resistance of the composites would be better for wear directions oriented parallel to the sheet faces of the composites, as studied here, compared with wear in the orthogonal lateral direction. During sliding, material transfer was an important factor which affected the dry sliding behavior of the studied tribosystem consisting of aCu-lSvol.%Cr in situ composite and a hardened AISI 52100 steel disk. Fig. 18 shows the original disk surface morphology and the composite transfer films on the disk after sliding at a sliding speed of 3.7 m s- ’and under a normal pressure of 0.06 MPa for 1 h. Even under such a low normal pressure as 0.06 MPa, composite transfer to the steel disk was significant (as in Fig. 18(b)). In the orientation where the sliding direction was normal to the ground grooves of the disk, the grooves were covered by the composite transfer film. In the orientation where the sliding direction was parallel to the ground grooves, the transferred film was a little thinner. The ground groves on the steel disk were filled with transferred material and smoothed by the sliding action. Material transfer from the hardened AISI 52100 steel disk to the composite pin also occurred during sliding. Energy
dispersive spectroscopy analysis on the wear surface in the scanning electron microscope revealed significant Fe peaks. This evidence confirmed that steel disk debris was back transferred on to the composite surface.
4. Conclusions The following conclusions can be made from this study of the sliding wear behavior and microstructural changes of deformation-processed Cu-lSvol.%Cr in situ composite. In the studied range of normal pressure (0.06-O&l MPa) and sliding speed (2.78-5.56m s-l), the wear rate increased with increasing normal pressure, whereas the wear rate decreased with increasing sliding speed. The sliding-induced subsurface deformation occurred not only in the sliding direction but also in the lateral directions perpendicular to the sliding direction. The experiments showed that the frictional forces in the pin-on-disk wear test produced a lateral flow that was largest at the surfaces of the sample and fell off toward the centerline between the two surfaces. The complex deformation mode was revealed clearly by the morphological change of the ribbon-like filaments. The thickness of the deformed subsurface layer increased with increasing normal pressure and sliding speed.
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4. Scanning electron microscopy analysis showed plastic deformation flow on the wear surface and illustrated that the surface deformation flow was affected by wear particles back transferred onto the surface.
Acknowledgements This material was based upon work supported by the National Science Foundation under Grant No. MSS9307797. The scanning electron microscopy was performed at the Center for Advanced Cement-Based Materials of the University of Illinois at Urbana-Champaign, The work of J.D. Verhoeven and E.D. Gibson was supported by the Ames Laboratory which is operated for the U.S. Department of Energy by Iowa State University under contract No. W-7405 ENG-82.
References [ l] I.D. Verhoeven, W.A. Spitzig,L.L. Jones, H.L. Downing,C.L.
Trybus, E.D. Gibson, L.S. Chumbley, L.G. Fritzemeierand G.D. Schnittgrund, Development of deformation processed copper-refractory metal composite alloys, J. ofMater. Erg., 12 (1990) 127-139. [2] W.A. Spitzig. A.R. Pelton and F.C. Laabs, Characterization of the strength and microstructure of heavily cold worked Cu-Nb composites, Acta Metall., 35 (1987) 2427-2442.
195 (1996) 214-222
[ 31 P.D. Funkenbusch, T.H. Courtney and D.G. Kubisch, Fabricability of and microstructural development in cold worked metal matrix composites, Ser. Metall., 18 (1984) 1099-l 104. [4] P.D. Funkenbusch and T.H. Courtney, On the strength of heavily cold worked in situ composites, Acta MetalL, 33 (1985) 913-922. [5] S.T. Kim, P.M. Berge and J.D. Verhoeven, Deformation processed Cu-Cr alloys-Optimizing strength and conductivity, J. Mater. Eng. Performance, 4 ( 1995). [6] D.A. Hardwick, C.G. Rhodes and L.G. Fritzemeier, The effect of annealing on the microstructure and mechanical properties of Cu-X microcomposites, Metall. Trans. A, 24A (1993) 27-34. [7] P.K. Rohatgi. Y. Liu and S. Ray, Friction and wear of metal-matrix composites, in ASM Handbook-Friction, Lubrication, and Wear Technology, Vol. 18, ASM International, Ohio, 1990, pp. 801-811. [8] P. Liu, S. Bahadur and J.D. Verhoeven, The mechanical and tribological behavior of Cu-Nb in situ composites, Wear, 166 (1993) 133-139. [9] P. Liu, S. Bahadur and J.D. Verhoeven, Further investigation on the tribological behavior of Cu-20%Nb in situ composite, Wear, 262-164 (1993) 211-219. [lo] E. Hombogen, Description of wear of materials with isotropic and anisotropic microstructures. in K.C. Ludema (ed.), WearofMaterials. The American Society of Mechanical Engineers, New York, 1985, pp. 477-484. [ 111 H. Nayeb-Hashemi and K. Lee, Friction and wear behavior of Cu-Nb microcomposite produced by powder metallurgy, Proc. 1993 ASME Winter Annu. Meeting, New Orleans. LA, 1993, pp. 263-267. [ 121 N.P. Suh, Tribophysics, Prentice-Hall, Englewood Cliffs, NJ, 1986, pp. 209-22 1. [ 131 D.A. Rigney, Shdingwearof metals,Annu. Rev. Mater. Sci., 18 ( 1988) 141-163. [ 141 Z. Chen, P. Liu, J.D. Verhoeven and E.D. Gibson, Sliding wear behavior and microstructure change of deformation-processed Cu20vol.%Nb in situ composite, Wear, 181-183 (1995) 263-270.
Wear 195 (1996) 214-222
Sliding wear behavior of deformation-processed Cu-lSvol.%Cr composites
in situ
Z. Chen a, P. Liu a, J.D. Verhoeven b, E.D. Gibson b aSchool of Technology, Eastern Illinois University, Charleston, IL 61920, USA b Ames Laboratory of USDOE. Department of Materials Science and Engineering, Iowa State UniversiTy, Ames, IA 50011, USA Received 17 May 1995; accepted 1 February 1996
Abstract A deformation-processed Cu-lSvol.%Cr in situ composite was made by consumable arc melting and casting followed by extensive deformation. A superior combination of mechanical strength and electrical/thermal conductivity was achieved with the composite since Cr filaments existed in the nearly pure copper matrix. The effects of sliding speed and normal pressure on sliding wear behavior and microstructure of the composite were investigated, with a composite pin rubbing against a hardened AISI 52100 steel disk on a pin-on-disk wear tester. In the studied range of normal pressure and sliding speed, the wear rate increased with increased normal pressure, whereas the wear rate decreased with increased sliding speed. Sliding-induced subsurface deformation occurred not only in the sliding direction but also in the lateral directions perpendicular to the sliding direction. This lateral flow produced a twisting of the Cr filaments. The complex deformation mode was revealed clearly by the morphological change of the ribbon-like filaments. Both constituents of the composite were cooperatively deformed. The thickness of deformed subsurface layer increased with increasing normal pressure and sliding speed. Scanning electron micrographs showed plastic deformation flow on the wear surface. Keywords: Sliding wear; Composite; Cu-Cr; Deformation processing
1. Introduction
Copper
refractory
metal
deformation-processed
in situ
provide a superior combination of high mechanical strength and high thermal/electrical conductivity. Cu-Cr composites were developed because Cr offers lower cost, better corrosion resistance, and higher modulus of elasticity. Verhoeven et al. [ 1] demonstrated that Cu-lSvol.%Cr alloy exhibited deformation characteristics similar to those of Cu-Nb and Cu-Ta alloys. Moreover, Cu-Cr was expected to produce higher strength at a given deformation strain than Cu-Nb because the strength of the composites was shown to increase with higher modulus of the refractory constituent [ 21. The remarkable malleability of the Cr phase, in spite of the brittle nature of pure Cr at room temperature, was first demonstrated by Funkenbusch et al. in Cu-17vol.%Cr alloys drawn into long filaments [ 3,4]. Through appropriate control of initial microstructure, the cast alloy could be drawn to a true drawing strain of 6.3. The tensile strength of Cu17vol.%Cr in situ composite increased with increased true drawing strain, reaching 840 MPa at a true drawing strain of composites
such
as Cu-Nb,
Cu-Ta
and Cu-Cr
0043-1648/96/$15.00 0 1996 Elsevier Science S.A. All rights resewed PII SOO43-1648(96)06949-9
6.3, which was well above that predicted by the rule of mixtures. Kim et al. [5] studied the effect of thermomechanical processing techniques on strength and electrical conductivity of Cu-7vol.%Cr in situ composite wire. It was found that the optimum combination of strength and electrical conductivity could be obtained by the process of DSDAD (deformation, solution treatment, drawing, aging and drawing). A combination of a strength of 925 MPa and an electrical conductivity of 7 1.8% IACS (International Annealed Copper Standard) was achieved, which is superior to commercially available copper alloys. Hardwick et al. [ 61 studied the effect of annealing on the microstructure and mechanical properties of a Cu-Cr microcomposite sheet deformed to a true strain of 6.6. After annealing for 10 h at 650 “C, extensive spheroidization of the Cr ribbons occurred. As a result, yield strength was reduced by 55%, and the ductility increased markedly. Failure in the CuCr composite occurred by a mixture of ductile rupture and cleavage. Friction and wear behavior of metal-matrix composites has attracted considerable interest, most of which is on nonmetallic fiber, particle or whisker-reinforced composites [ 71.
2. Chen et al. /Wear
No investigation of friction and wear behavior of Cu-Cr in situ composites is available. Only a few studies have been conducted for friction and wear behavior of Cu-Nb in situ composites and precipitation-hardened Cu-Cr alloys. Liu et al. [ 81 studied the effects of Nb composition and true deformation strain on wear rate and coefficient of friction. A CuNb in situ composite pin was slid against a hardened 02 tool steel disk at sliding speeds from 0.028 to 2.5 m s- ’and under a normal pressure of 0.68 MPa. It was found that Cu20vol.%Nb showed the best wear resistance. Increasing true deformation strain resulted in increased tensile strength and coefficient of friction and decreased wear rate. The tribological behavior of the Cu-2Ovol.%Nb in situ composite was investigated in terms of sliding-induced plastic deformation flow, effects of Nb filament orientation, sliding speed and annealing temperature [ 91. The wear rate of the specimen with Nb filaments perpendicular to the sliding direction was lower than that with Nb filaments parallel to the sliding direction, which indicated that this in situ composite material had an anisotropic wear resistance as modeled by Hornbogen [ lo]. With increased annealing temperature, the hardness of the composite decreased, and both the coefficient of friction and the wear rate increased. It was found that both the Cu matrix and the Nb filaments underwent sliding-induced deformation with no loss of bonding between the Nb filaments and the Cu matrix. Increasing sliding speed resulted in a lower wear rate, which the authors ascribed to oxidation and work hardening. Nayeb-Hashemi et al. [ 111 investigated the friction and wear behavior of a powder metallurgical Cu-Nb microcomposite rubbing against gray cast iron of hardness 92 RB. They found that the wear rate increased drastically with sliding speed and the coefficient of friction seemed to be independent of fiber orientations. Suh [ 121 studied the tribological behavior of precipitationhardened Cu-O58at.%Cr and Cu-O.8 lat.%Cr alloys aged for different periods at 500 “C. The alloy cylinder rotated against a stationary AISI 52100 steel rod. The hardness of these materials initially increased with aging time from 637 MPa for solid solutions to 1372 Mpa for peak-aged alloys, and then decreased. It was found that the wear rate initially decreased with aging time by a factor of 3 for both Cu-Cr alloys and then increased approximately linearly. The minimum wear rate did not correspond to the maximum hardness. However, the coefficient of friction was fairly constant for all treatment times for both alloys. They attributed these results to the combined effects of hard particles on the hardness and the coherency of particles to the matrix. The dry sliding wear behavior of Cu-lSvol.%Cr in situ composite has been studied in this paper. The effects of normal pressure and sliding speed, the mechanism of subsurface deformation as well as the microstructural change due to sliding are discussed. The thickness of the subsurface deformation layer and microstructural change have been quantitatively analyzed in relation to normal pressure and sliding speed.
195 (19%) 214-222
215
2. Experimental details
Fig. 1 shows the flow chart of the material preparation process and the corresponding microstructure for the CulSvol.%Cr in situ composite. A Cu-lSvol.%Cr alloy ingot was first prepared by the vacuum induction melting method, followed by chill casting. An electrode formed from this casting was then consumable arc melted into a copper mold in order to obtain fine dendrites of Cr in the Cu-Cr alloy ingot. The deformation process started with hot extrusion of an ingot of 102 mm diameter into a rod of 5 1 mm diameter. Then the rod was rolled into a plate of 12.7 mm thickness and finally into a sheet of 1.6 mm thickness. The total true deformation strain was q= 5.4 or a reduction in area of 99.5%, which resulted in a hardness of 74 RB. Dry sliding tests were performed on a pin-on-disk wear tester, with the Cu-lSvol.%Cr in situ composite pin rubbing against a hardened AISI 52100 steel disk. The steel disk was austenitized at 843 “C for 30 min, oil quenched, and then tempered at 150 “C for 1 h. The resultant hardness of the disk was between 58.5 and 62 RC. The pin was cut from the CulSvol.%Cr in situ composite sheet and had a rectangular section of 1.57 X 10 mm. During the wear test, the 10 mm length of the specimen was in the sliding direction and the Cr filaments were perpendicular to the sliding plane. The disk rotated at 600,800,1000, and 1200 rev min- * with an average wear track diameter of 88.5 mm, i.e. the sliding speed ranged from 2.78 to 5.56 m s-i. The normal load varied from 0.98 to 6.86 N, which corresponded to a normal pressure of 0.06-0.44 MPa over the apparent contact area of 15.7 mm*. Before wear testing, the specimen was ground on emery paper to 600 grit in running water, and then was cleaned in methyl alcohol for 5 min using an ultrasonic cleaner. The disk was cleaned with methyl alcohol. The pin was weighed every 5 min with an accuracy of lo-’ g. Processing
Microstructure
I
Consumable arc
Dendritic Cr in Cu/Cu-Cr alloy
melting into water cooled copper mold
m&ix
Deforming: Hot extrusion
I Deforming: Rolling
’
Fig. 1. Flow chart of composite StruCtUWS.
Cr tilamcnts in cldcu-Cr alloy matrix
processing
and corresponding
micro-
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195 (1996) 214-222
slidingDkction 4 I
/
LongitudinalSe&on
TransvcxscSection
Fig. 2. Schematic diagram of nomenclature rolling and sliding directions.
definition for specimen sections,
3. Results and discussions
the rolling and sliding directions are illustrated in Fig. 2. Sliding took place on the transverse section. The microstructure of the Cu-lSvol.%Cr in situ composite before sliding is shown in Fig. 3. Fig. 3(a) shows the microstructure taken from the face section. The Cr filaments, with variable ribbonshaped cross-sections, align in the rolling direction. Fig. 3 (b) illustrates that some Cr filaments with coarse knots are present in the longitudinal sections, which indicates that Cr dendrites were not uniformly deformed into thin ribbons. Some Cr phase can also be seen as spherical precipitates in the copper matrix. 3.2. Effect of normal pressure on wear
3.1. Microstructure
of Cu-lSvol.%Cr
in situ composite
The ingot was deformed with a true deformation strain of 5.4, which caused the original Cr dendrites to turn into ribbonlike filaments aligned in the rolling direction. The definitions of the composite specimen sections and relationship between
Fig. 3. Microstructure
of as-processed
Cu-lSvoI.%Cr
in situ composite:
The variation of volume loss of Cu-lSvol.%Cr in situ composites with sliding distance under different normal pressures and at a sliding speed of 3.70 m SK’ is shown in Fig. 4. The volume loss increased with increasing sliding distance, with a rapid increase during the running-in stage and an
(a) face section; (b) longitudinal
section; (c) traverse section
Z. Chen et al. /Wear 195 (1996) 214-222
0
2
4
6 g 10 sliding Distance (km)
12
ot,““““‘,““‘,“‘,“,i 0 0.1
14
Fig. 4. Variation of composite volume loss with sliding distance at a sliding speed of 3.70 m s-’ and under four different normal pressures.
0.2 0.3 Nomal Rcssurc (MPa)
10
20
30 Sliding Tii
40 bin)
Fig. 5. Variation of wear rate of the composite sliding speed of 3.70 m s-l.
50
60
pin with sliding time at a
almost linear increase in the following steady-state stage. Under a normal pressure of 0.44 MPa, the total volume loss was much greater than that under lower normal pressures. A total wear rate can be defined as total volume loss per sliding distance [ 131. The variation of the total wear rate with sliding time is illustrated in Fig. 5. The total wear rate decreased with sliding time, approaching a steady state. It may be seen in Fig. 4 that increasing the pressure from 0.31 to 0.44 MPa produced a sharp increase in initial wear. This transition is illustrated nicely in Fig. 6 which plots total wear rate after a 60 min sliding time vs. the normal pressure. It is seen that a sharp transition from mild wear to severe wear has occurred when the pressure was raised from 0.31 to 0.44 MPa. These data show that the material should be maintained at a normal pressure below 0.44 MPa to avoid severe wear while sliding at a speed of 3.7 m s- ‘. 3.3. Effect of sliding speed on wear The change of the total wear rate with sliding time under a normal pressure of 0.31 MPa is shown in Fig. 7. Fig. 8 shows the variation of the total wear rate with sliding speed under a normal pressure of 0.31 MPa. The total wear rate decreased with increased sliding speed, but it tended to level off at a sliding speed of 4.63 m s-‘. 3.4. Microstructural
change due to sliding
By studying the change in the alignment of the Cr filaments near the wear surface, it was possible to evaluate the nature of the deformation layer produced at the surface. Fig. 9 illus-
0.5
Fig.
6. Variation of steady wear rate with normal pressure for Cu-lSvol.%Cr composite slid at a sliding speed of 3.70 m SK’. 10
0
0.4
ll..,ll,l,.llllllll(,...,,,.,,,~
0
10
20
30 40 Sliding Time (ah)
so
60
Fig. 7. Variation of composite pin wear rate with sliding time under a normal pressure of 0.31 MPa.
02
2.5
3
3.5 4 4.5 Sliding Speed (m i’)
5
5.5
6
Fig. 8. Variation of steady wear rate with sliding speed under a normal pressure of 0.31 MPa.
trates an ideal deformation pattern caused by sliding, which is depicted by changes in the Cr filaments. The deformation layer had a maximum thickness at the faces of the pins, and the layer thickness decreased symmetrically toward the centerline (C.L.) of the pins, which could be expected from simple contact mechanics since a flat punch had maximum stresses at the edges. Fig. 10 shows the face section microstructure of the composite below the wear surface after sliding at a speed of 3.70 m s-’ and under normal pressures of 0.06 MPa and 0.44 MPa. No debonding occurred between the Cu matrix and Cr filaments. The sliding-induced plastic deformation layer is revealed by changes in alignment of the Cr filaments. Under a lower normal pressure the majority of Cr filaments were bent in the sliding direction, as shown in Fig. 10(a). Under a higher normal pressure, however, Cr filament refinement was seen as a major characteristic (Fig. 10(b)). In addition to producing a flow in the sliding direction within the deformation layer, the sliding force also produced
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Slldbg DhCth
Loqltodbd wctbn
CbCkWke Twkt off3
nbment
Fig. 9. Schematic of subsurface plastic deformation pattern on CulSvol.%Cr composite due to sliding. a flow in the lateral directions defined on Fig. 9. This flow was revealed by a twisting of the filaments demonstrated in Fig. 11, which shows two matched sectional micrographs of Cu-lSvol.%Cr in situ composite sliding at a speed of 3.7 m s-’ and under a normal pressure of 0.31 MPa for 60 min. Fig. 11 (a) was taken on the face section and Fig. 11 (b) on the longitudinal section. The sliding direction in Fig. 11 (a) was from the left to the right, as shown by the arrow, whereas the sliding direction in Fig. 11 (b) was from the page toward the reader. The pair of micrographs in Fig. 11 are matched along the edge located at the upper left corner of Fig. 9. The magnification is sufficiently high that only the wear affected layer appears in the micrographs, so that Fig. 11 (a) corresponds to the lower portion of Fig. IO(b). Fig. 11 (b) illustrates the twisting action produced by flow in the lateral direction. As shown in Fig. 9 the lateral flow produced a twisting action in opposite directions on the two sides of the rectangularly shaped pins, a clockwise rotation near the left face and a counter clockwise rotation near the right face. The twisting of the filaments was studied with optical and scanning electron microscopy and Fig. 12 presents a model of the shape change involved in the twisting.
3.5. Thickness of deformed subsurjace layer The thickness of the deformed subsurface layer was defined as the distance between the zone of unchanged microstructure to the wear surface on the face sections, as shown in Fig. lO( b) . Fig. 13 shows the variation of the thickness of deformed subsurface layer with normal pressure at a sliding speed of 3.7 m s - ‘. The thickness data were an average over 18 equal-distance measurements along the entire contact length of the pin in the sliding direction. The thickness increased monotonically from 52 to 246 pm, when normal
Fig. 10. Microstructural change on face sections of Cu-lSvol.%Cr in situ composite pin after sliding against a hardened AISI 52100 disk at a sliding speed of 3.70 m s-’ for 1 h, (a) under a normal pressure of 0.06 MPa; (b) under a normal pressure of 0.44 MPa. The arrow indicates the sliding direction. increased from 0.06 to 0.44 MPa. The rate of thickness increase with normal pressure decreased as normal pressure rose. The variation of the thickness of deformed subsurface layer with sliding speed under a normal pressure of 0.31 MPa is shown in Fig. 14. The thickness of deformed subsurface layer increased linearly with sliding speed. At a sliding speed of 2.78 m s-‘, the thickness of subsurface layer was 159 pm,
pressure
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Fig. 11. Micrograph
of Cu-lSvol.%Cr in situ composite pin, after sliding at a speed of 3.7 m s-l and under a normal pressure of 0.31 MPa for 60 min: (a) face section; (b) longitudinal section. The arrow indicates the sliding direction.
0.1
0
0.2
0.3
0.4
0.5
NormalI’msun (MPa)
Fig. 13. Variation of thickness of deformed pressure at a sliding speed of 3.7 m s-r.
subsurface
layer with normal
Sliding Direction
*zeal
Direction
Fig. 12. Schematic model of spatial shape change in ribbon-like during sliding wear process.
Cr filament
and at a sliding speed of 5.56 m s-‘, the thickness was 309 km. From Fig. 13 and Fig. 14, it can be seen that both increased pressure and sliding speed increased the thickness of the deformed subsurface layer. The deformed subsurface layer was caused by the friction when the composite pin slid against the hardened disk. The measured coefficient of friction decreased slightly with increased normal pressure and sliding speed. Generally, increasing normal pressure would cause
2.5
3
3.5
4 4.5 5 Sliding Sped (m 6’)
Fig. 14. Variation of thickness of deformed speed under a normal pressure of 0.31 MPa.
5.5
subsurface
6
layer with sliding
more asperities of the composite to contact with those of the disk and hence increase the friction force even though the average coefficient of friction decreased slightly, which would increase the opportunity for plastic deformation of the composite. The real plastic deformation of Cr filaments in the composite was caused by the local stress. Increasing the applied normal pressure certainly increased the asperity stress
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that caused a larger extent of plastic deformation. Since increasing sliding speed increased surface temperature which was resulted from friction heat and deformation energy, the pin material became softer at higher sliding speed than at lower sliding speed. The temperature increase caused by increased sliding speed would facilitate plastic deformation. 3.6. Study of wear mechanism Although crystalline Cr has brittle characteristics at room temperature [ 31, there was a deformation flow of the composite both in the sliding direction and in lateral directions, as shown in Fig. 10 and Fig. 11. The Cr filaments demonstrated good strain compatibility with the Cu matrix during fabrication of the composites. It is not surprising, therefore, that the composite displayed a good strain compatibility between its constituents during sliding and that no debonding phenomenon was observed between the Cr filaments and the Cu matrix. This sliding-induced deformation found in the Cu-Cr system was similar to that found in the Cu-Nb system [141. The slip mechanism occurring during the lateral twist can be explained with the favored crystal slip systems depicted in Fig. 15. Since ribbon-like Cr filaments were embedded in a softer Cu matrix, plastic deformation of the Cu-lSvol.%Cr in situ composites was controlled by the deformation mode of the harder Cr filaments. According to Hardwick et al. [ 61, a texture was developed in the Cr filaments at the true strain of 6.6, though the texture was not as strong as those found for Nb in Cu-lSvol.%Nb and for Ta in Cu-lSvol.%Ta. The main rolling texture for body-centered cubic Cr could be identified as (001) [ 1lo]. The ( 110) crystal plane of the Cr filaments lay in the sliding surface, and the crystal directions [ ill ] and [ ii1 ] were 45” away from the sliding direction. Therefore, both ( 110) [ i 111 and ( 110) [ ii 1] slip systems could be activated during sliding. Slip motion 45” away from the sliding direction in the Cr filaments could account for both elongation in the sliding direction and twist in the lateral directions. As illustrated by comparing Fig. 10(a) with Fig. lO( b), the deformation layer was less extensive under lower normal pressure than under higher normal pressure. It appears that Fig. lO( a) shows the beginning of plastic deformation of Cr filaments, while in Fig. lO( b) the plastic deformation of Cr filaments is well developed in both the sliding and lateral
Fig. 15. Schematic composite.
crystal
deformation
model for Cu-lSvol.%Cr
in situ
I95 (1996) 214-222
directions. The figures confirm that increasing normal pressure causes more plastic deformation in the contact subsurface region. The scanning electron micrograph of the wear surface shown in Fig. 16 also reveals the plastic deformation flow on the wear surface. This surface deformation flow was affected by wear debris and transferred material. Fig. 17 shows a micrograph of the composite microstructure located just below the wear surface. The specimen was sectioned slightly below but parallel to the wear surface. The black regions indicate the presence of wear debris, back-transferred material or oxide, on the surface. The orientation of the Cr filaments clearly shows that the surface deformation flow was affected by the transferred material. The complex flow pattern in the deformation layer revealed by the change in orientation of the Cr filaments is quite interesting. According to contact mechanics, apparently flow in the direction of the applied force on the wear surface is a maximum at the faces of the pin and decreases to a minimum at the center of the pin. Because the broad faces of the Cr
Fig. 16. Scanning electron micrograph revealed plastic deformation flow, after sliding at a speed of 3.7 m s- ’and under a normal pressure of 0.3 1 MPa for 30 min. The arrow indicates the sliding direction.
Fig. 17. Embedded wear debris inhibited the plastic deformation flow. The specimen was slid at a speed of 3.7 m s-’ and under a normal pressure of 0.19 MPa for 60 min. The arrow indicates the sliding direction.
Z. Chen et al. /Wear 195 (19%) 214-222
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Fig. 18. Transfer films in the steel disk: (a) the original disk surface; (b) sliding direction normal to the ground groves and (c) sliding direction parallel to the ground groves of the steel disk, sliding at 3.70 m s-’ under a normal pressure of 0.06 MPa. The arrow indicates the sliding direction.
filaments are aligned parallel to the pin faces the bending moment inertia of the filaments in the applied force direction is much higher than it would be perpendicular to this direction. It seems likely, therefore, that the twisting of the filaments occurs because bending in the lateral directions requires a very low stress. Hence, one would predict that the wear resistance of the composites would be better for wear directions oriented parallel to the sheet faces of the composites, as studied here, compared with wear in the orthogonal lateral direction. During sliding, material transfer was an important factor which affected the dry sliding behavior of the studied tribosystem consisting of aCu-lSvol.%Cr in situ composite and a hardened AISI 52100 steel disk. Fig. 18 shows the original disk surface morphology and the composite transfer films on the disk after sliding at a sliding speed of 3.7 m s- ’and under a normal pressure of 0.06 MPa for 1 h. Even under such a low normal pressure as 0.06 MPa, composite transfer to the steel disk was significant (as in Fig. 18(b)). In the orientation where the sliding direction was normal to the ground grooves of the disk, the grooves were covered by the composite transfer film. In the orientation where the sliding direction was parallel to the ground grooves, the transferred film was a little thinner. The ground groves on the steel disk were filled with transferred material and smoothed by the sliding action. Material transfer from the hardened AISI 52100 steel disk to the composite pin also occurred during sliding. Energy
dispersive spectroscopy analysis on the wear surface in the scanning electron microscope revealed significant Fe peaks. This evidence confirmed that steel disk debris was back transferred on to the composite surface.
4. Conclusions The following conclusions can be made from this study of the sliding wear behavior and microstructural changes of deformation-processed Cu-lSvol.%Cr in situ composite. In the studied range of normal pressure (0.06-O&l MPa) and sliding speed (2.78-5.56m s-l), the wear rate increased with increasing normal pressure, whereas the wear rate decreased with increasing sliding speed. The sliding-induced subsurface deformation occurred not only in the sliding direction but also in the lateral directions perpendicular to the sliding direction. The experiments showed that the frictional forces in the pin-on-disk wear test produced a lateral flow that was largest at the surfaces of the sample and fell off toward the centerline between the two surfaces. The complex deformation mode was revealed clearly by the morphological change of the ribbon-like filaments. The thickness of the deformed subsurface layer increased with increasing normal pressure and sliding speed.
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4. Scanning electron microscopy analysis showed plastic deformation flow on the wear surface and illustrated that the surface deformation flow was affected by wear particles back transferred onto the surface.
Acknowledgements This material was based upon work supported by the National Science Foundation under Grant No. MSS9307797. The scanning electron microscopy was performed at the Center for Advanced Cement-Based Materials of the University of Illinois at Urbana-Champaign, The work of J.D. Verhoeven and E.D. Gibson was supported by the Ames Laboratory which is operated for the U.S. Department of Energy by Iowa State University under contract No. W-7405 ENG-82.
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Trybus, E.D. Gibson, L.S. Chumbley, L.G. Fritzemeierand G.D. Schnittgrund, Development of deformation processed copper-refractory metal composite alloys, J. ofMater. Erg., 12 (1990) 127-139. [2] W.A. Spitzig. A.R. Pelton and F.C. Laabs, Characterization of the strength and microstructure of heavily cold worked Cu-Nb composites, Acta Metall., 35 (1987) 2427-2442.
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[ 31 P.D. Funkenbusch, T.H. Courtney and D.G. Kubisch, Fabricability of and microstructural development in cold worked metal matrix composites, Ser. Metall., 18 (1984) 1099-l 104. [4] P.D. Funkenbusch and T.H. Courtney, On the strength of heavily cold worked in situ composites, Acta MetalL, 33 (1985) 913-922. [5] S.T. Kim, P.M. Berge and J.D. Verhoeven, Deformation processed Cu-Cr alloys-Optimizing strength and conductivity, J. Mater. Eng. Performance, 4 ( 1995). [6] D.A. Hardwick, C.G. Rhodes and L.G. Fritzemeier, The effect of annealing on the microstructure and mechanical properties of Cu-X microcomposites, Metall. Trans. A, 24A (1993) 27-34. [7] P.K. Rohatgi. Y. Liu and S. Ray, Friction and wear of metal-matrix composites, in ASM Handbook-Friction, Lubrication, and Wear Technology, Vol. 18, ASM International, Ohio, 1990, pp. 801-811. [8] P. Liu, S. Bahadur and J.D. Verhoeven, The mechanical and tribological behavior of Cu-Nb in situ composites, Wear, 166 (1993) 133-139. [9] P. Liu, S. Bahadur and J.D. Verhoeven, Further investigation on the tribological behavior of Cu-20%Nb in situ composite, Wear, 262-164 (1993) 211-219. [lo] E. Hombogen, Description of wear of materials with isotropic and anisotropic microstructures. in K.C. Ludema (ed.), WearofMaterials. The American Society of Mechanical Engineers, New York, 1985, pp. 477-484. [ 111 H. Nayeb-Hashemi and K. Lee, Friction and wear behavior of Cu-Nb microcomposite produced by powder metallurgy, Proc. 1993 ASME Winter Annu. Meeting, New Orleans. LA, 1993, pp. 263-267. [ 121 N.P. Suh, Tribophysics, Prentice-Hall, Englewood Cliffs, NJ, 1986, pp. 209-22 1. [ 131 D.A. Rigney, Shdingwearof metals,Annu. Rev. Mater. Sci., 18 ( 1988) 141-163. [ 141 Z. Chen, P. Liu, J.D. Verhoeven and E.D. Gibson, Sliding wear behavior and microstructure change of deformation-processed Cu20vol.%Nb in situ composite, Wear, 181-183 (1995) 263-270.