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The influence of ball size on tribological behaviour of MoS2 coating tested on a ball-on-disk wear rig
Wear 243 (2000) 1–5
The influence of ball size on tribological behaviour of MoS2 coating tested on a ball-on-disk wear rig Jiaren Jiang a,∗,1 , R.D. Arnell a , Gajendra Dixit b a
Centre for Advanced Materials and Surface Engineering, University of Salford, Salford M5 4WT, UK b Department of Applied Mechanics, Maulana Azad College of Technology, Bhopal 462 007, India Received 8 April 1999; received in revised form 8 February 2000; accepted 9 February 2000
Abstract The effect of ball diameter on wear and friction of a molybdenum disulphide (MoS2 ) coating deposited using the closed-field magnetron sputtering technique has been investigated on a ball-on-disk wear rig sliding against uncoated steel ball-bearing balls. It was observed that the wear rate of the coating decreased significantly with increase in ball diameter. Correspondingly, the average friction coefficient in the steady state of sliding increased with increase in ball diameter. Under the present experimental conditions, the effect of ball size on wear and friction of the coating vanished when the diameter of the ball exceeded 6.35 mm. This phenomenon is thought to be a result of wear mode transition as a function of contact pressure at the apparent area of contact from a low-rate ‘mild’ wear regime to a high-rate ‘severe’ wear regime when the contact pressure exceeds some critical value. Simulation results using a previously presented model based on the above assumption showed a good agreement with the experimental observations. © 2000 Elsevier Science S.A. All rights reserved. Keywords: Molybdenum disulphide (MoS2 ) coating; Magnetron sputtering; Tribological behaviour; Wear; Effect of experimental conditions
1. Introduction The pin-on-disk wear rig is one of the most commonly used configurations for sliding wear tests in the laboratory. To facilitate accurate alignment between the pin and the disk surfaces, a dome-ended pin specimen is often used. While this arrangement reduces errors in wear measurements, some complications are brought into the interpretation of the obtained experimental results by the non-uniform distribution of contact pressures over the apparent area of contact. For example, although the resultant wear and wear rate would not be influenced by the geometry of the pin if the Archard’s wear behaviour is followed, the wear results will be different for different geometry of the pin specimen if wear of the materials being studied is dependent on contact pressures. There have been few studies carried out to elucidate the effect of the use of the dome-ended pin on the observed wear behaviour. Molybdenum disulphide (MoS2 ) is a unique material as a solid lubricant applied in a vacuum and in inert gases and has been widely studied over the past several decades [1–7]. The application of sputter deposition techniques in produc∗
Corresponding author. Current address: IMTI-NRC, 800 Collip Circle, London, ON, Canada N6G 4X8. 1
ing thin MoS2 coatings has significantly improved the tribological properties of the coatings and provided the potential of modifying and tailoring properties of the coatings by incorporating various elements/compounds [7,8], changing the chemical compositions [9,10] or modifying crystal structure and intercrystallite slip [11,12]. In this work, the effect of ball diameter on tribological behaviour of a MoS2 coating deposited on M42 tool steel substrate using magnetron sputtering technique has been investigated on a ball-on-disk wear rig. The results are explained on the basis of a model considering the effect of contact pressure on wear and wear transition. 2. Experimental details The MoS2 coating was deposited onto M42 tool steel disk substrate using the closed field magnetron sputtering technique. Fig. 1 shows the surface morphology of the as-deposited coating observed using an SEM. According to Buck [13], the morphology corresponds to the stoichiometry for MoS2 coatings/films deposited using the conventional r.f. sputtering technique. The morphology in Fig. 1 seems to suggest that this coating had an amorphous structure and probably a non-stoichiometric composition of MoSx with x≥1. However, since this coating has been deposited
0043-1648/00/$ – see front matter © 2000 Elsevier Science S.A. All rights reserved. PII: S 0 0 4 3 - 1 6 4 8 ( 0 0 ) 0 0 3 4 1 - 0
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J. Jiang et al. / Wear 243 (2000) 1–5
Fig. 1. SEM photograph showing the surface morphology of the as-deposited MoS2 coating.
using a different technique with higher bombardment energy, more detailed study may be necessary fully to characterise the coating. The thickness of the MoS2 layer was approximately 2.6 m measured using the ball crater technique. The hardness and reduced elastic modulus of the MoS2 coating, measured using a nanoindenter, were 7.7±0.3 and 150±4 GPa, respectively. The surface roughness of the coating was Ra 0.37 m. Wear tests were carried out on a ball-on-disk wear rig in dry air with a relative humidity of approximately 7%; this humidity was maintained by passing compressed dry air through saturated sodium hydroxide solution before inleting the gas to the chamber enclosing the test specimens. Uncoated steel ball bearing balls were used as the counter specimen. The diameters of balls used were 3.18, 4.75, 6.35, 7.93, 9.52 and 12.69 mm. Before wear tests, the ball and the disk specimens were ultrasonically cleaned in acetone for approximately 10 min and rinsed using acetone to remove grease and contaminants that may be present on the specimen surfaces. A normal load of 35 N was applied by applying a dead weight to the ball holder. Under this load condition, the initial mean Hertzian contact pressures at the rubbing interface for the various ball sizes were 1180, 903, 744, 642, 568, and 469 MPa, respectively. The sliding speed was 0.25 m s-1 and the total sliding distance was between 400 and 1500 m. During sliding, friction forces were continually measured and logged onto a microcomputer. Wear volumes of the disk specimens were measured using a profilometer linked to a computer. At least eight measurements were made along the circumference of each wear track. The fluctuations in cross-section areas obtained on one wear track were normally within 12% of the average value; the apparently large fluctuations were mainly due to the unevenness of the wear track along the circumference. However, the relative precision of the measurement technique itself was within 2.4% as estimated by repeated measurements on the same wear track. Wear of the ball was estimated by measuring the sizes of the major and the minor axes of the elliptical-shaped wear scar under a microscope and assuming that the scar surface
Fig. 2. The effect of ball diameter on specific wear rates of the MoS2 coated disks (䉱) and the balls (+). The solid lines show the simulation results using the elastic foundation wear model presented in Ref. [23].
was ideally flat. Apparently, this method over-estimates wear volumes of the balls because scars formed on the ball normally had some convex curvatures. Wear has been presented in the form of Archard’s specific wear rate in mm3 N-1 m-1 . In the wear curve measurements, sliding was stopped at certain time intervals to measure the wear volume of the specimens and the sliding was restarted again. A profilometer was attached to the wear rig; this allowed the specimens to remain fitted on the specimen holders during the wear volume measurement so that the re-alignment between the ball and the disk contact areas after the measurement was very good.
3. Results The variation of specific wear rate of the coating as a function of ball diameter is shown in Fig. 2. In general, the specific wear rate decreased with increase in ball diameter. Wear rate at the ball diameter of 3.18 mm was significantly higher than that for the larger ball sizes, although the scatter was larger for the small ball. Some typical wear curves at the different ball diameters are shown in Fig. 3. Wear increased almost linearly with
Fig. 3. Some typical wear curves of the MoS2 coating using balls of different diameters.
J. Jiang et al. / Wear 243 (2000) 1–5
Fig. 4. Typical variation of friction coefficient as a function of sliding distance at different ball sizes. The general trend was similar for the various ball sizes.
sliding distance at small and large ball sizes; however, transitions in wear rate from a high initial rate to a lower rate after some time of sliding was apparently observed at the intermediate diameter of balls. The variation in friction coefficient as a function of sliding distance was very similar for the various diameters of balls. Fig. 4 shows two typical examples. After a short distance of sliding, the friction coefficient increased from some low initial value to slightly higher values and then the friction coefficient fluctuated around some constant average value. However, the average values of friction coefficient showed a steady increase with increase in diameter of the ball, as shown in Fig. 5. The values in Fig. 5 were calculated using friction data after a sliding distance of 200 m.
4. Discussion In sliding wear, the Archard’s Wear Law [14,15] is one of the most basic and most widely used equations; it was derived on the basis of ‘adhesive wear’ concept and predicts a linear increase in wear volume as a function of sliding distance. The same form of wear equation has been obtained based on abrasive cutting wear [16], fatigue wear [17–19]
Fig. 5. Variation of average friction coefficient as a function of ball diameter. The values were calculated using data for sliding beyond 200 m.
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and delamination theory of wear [20]. The Archard’s type of wear cannot explain the significant effect of ball diameter on wear observed in the present study because according to this type of prediction, the total wear volume is only dependent on the total load irrespective of the distribution of pressures over the apparent area of contact. In practice, a transition in wear rate from a high rate at the initial stages of sliding to a low rate after certain distance of sliding is frequently observed. For metals, this transition has been classified as a transition from ‘severe’ wear to ‘mild’ wear and has been attributed to the formation of oxide layers on wear surfaces by Archard [14] and Archard and Hirst [15]. Nevertheless, according to the systematic study of Welsh [21–23] on wear transitions of steels under various conditions, it was shown that the establishment of oxide on the rubbing surfaces was not necessary for the severe to mild wear transition to occur when the substrate hardness exceeded a certain value. In addition, such wear transition has also been observed on non-metal materials (e.g. Fig. 3 and Refs. [24,25]). Most recently, Williams et al. [25] analysed the wear rate transition of carbon-graphite materials as a function of sliding distance. In their model, it is assumed that with sliding, wear debris particles are accumulated/agglomerated on the rubbing surfaces; such agglomeration reduces the maximum contact stresses at the real areas of contact. It was further assumed that wear at areas contacting the agglomerated particles with low contact stresses is negligible and the Archard’s wear law is applicable at the points of intense contact load. Thus, the wear rate decreases with sliding due to the decrease in contact stress. However, during the sliding of the MoS2 against the very smooth steel balls in this study, accumulation/agglomeration of wear debris particles on the MoS2 disk specimen was not significant. Although some material transfer from the disk to the ball surface indeed occurred, this transfer film/layer could not have caused any significant decrease in contact pressures at the real areas of contact because the initial surface of the ball was already very smooth. The development of compact wear debris particle layers have also been shown to have significant and beneficial effect on wear transitions in dry sliding wear of metals (e.g. Ref. [26]) where comminution of wear debris particles are necessary for such layers to form on wear surfaces. However, in the present study, it was noticed that transfer layers were developed on the pin/ball surface after a very short sliding distance and were observed on balls of all the sizes used. Little wear was observed on the ball specimens, presumably due to the protection effect of the transfer layer. Thus, it is reasonable to state that the development of compact wear debris particle layers or transfer layers on wear surfaces was not the dominant factor in controlling wear transitions of the MoS2 coating under the current experimental conditions. According to Welsh’s [21–23] observations, the Archard’s specific wear rate of carbon steels is a function of normal load; the dominant wear mode changed from ‘mild’ wear to ‘severe’ wear when load exceeded some critical value,
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J. Jiang et al. / Wear 243 (2000) 1–5
resulting in an increase in specific wear rate by approximately two orders. This critical transition load was probably associated with the deformation mode in the generation of wear debris particles at the rubbing surfaces, extensive plastic deformation being involved at the real contact region in the ‘severe’ wear mode. Similar relationship between wear mode and normal load or contact pressure can be expected in other materials such as MoS2 , although the characteristics of wear surfaces in the two wear modes may be different for different materials. When the maximum stresses in the region of real areas of contact exceeds some critical value, depending on the properties of the material, severe plastic deformation and probably low cycle fatigue become the dominant mechanism in the generation of wear debris particles. Under such conditions, the wear rate is considerably higher than that at pressures when elastic contact prevails. Based on such argument, the authors [24] presented a wear model describing the running-in process of wear on a ball-on-disk wear rig and showed a good agreement with experimental observations. In this model, the actual contact pressures and their distribution across the rubbing interface were described using contact mechanics. Specific wear rate was represented as a function of the actual contact pressure at a given point within the rubbing interface. Based on Welsh’s [21–23] findings, it was assumed that the dominant wear regime would change from a low-rate regime to a high-rate regime when the real contact pressure exceeds some critical value — the critical contact pressure (pc ). At the same time, the specific wear rate would increase from a lower value to a significantly higher value. Values of the critical contact pressure and the ratio between the high-rate wear regime and the low-rate wear regime are dependent on properties of the material(s) being studied. Thus, on a ball-on-disk type of wear machine, the specific wear rate near the central part of the wear track is expected to be much higher than that at regions towards the edges of the track because the contact pressure near the centre is the highest and may exceed the critical contact pressure for the material being studied. With the progress of sliding, both the ball and the disk are worn, leading to increased apparent area of contact and greatly decreased contact pressures within the rubbing interface. As a result, the size of high-rate-wear-regime-dominated region decreases with sliding distance. Depending on load and wear properties of the rubbing materials, the low-rate wear regime may eventually become dominant over the whole apparent area of contact and a transition in wear rate from a high rate to a low rate can thus be observed (e.g. when using a 6.35-mm ball in this study as shown in Fig. 3). Obviously, under certain conditions where the load is high and/or the critical contact pressure of the material is low, the high-rate wear regime will be dominant during most of the sliding time, e.g. at a ball diameter of 3.18 mm in the present study (Fig. 3), the wear rate being high throughout the sliding. On the other hand, the low-rate wear regime may be dominant over the whole apparent area of contact and throughout most of the sliding distance if the contact pressure is low and/or
the critical contact pressure for the material is high (Fig. 3, using a 12.69-mm ball). Based on this model, when a small ball is used, the contact pressures over a large proportion of the apparent area of contact can be expected to be above the critical contact pressure. As a result, a high-rate wear regime dominates much of the sliding wear process (Fig. 2), the overall wear rate at a small ball size being much higher than that when a larger ball was used. Based on the above concept, if the distribution of contact pressures within the rubbing interface and the transition property of the material are known, then the wear process of the sliding system can be simulated. In Ref. [24], the elastic-foundation method developed in contact mechanics [27] was applied to calculate the contact pressures between the ball and the disk specimens. Applying this elastic-foundation wear model presented in Ref. [24], using the mechanical properties for the MoS2 coating and the steel ball, the variation of specific wear rate as a function of ball diameter has been simulated. The results for sliding distances of 750 and 1500 m are plotted in Fig. 2 as solid lines. In the simulation, values for the critical contact pressure, pc , and the ratio of specific wear rates between the high-rate wear regime and the low-rate wear regime, K2 /K1 , were obtained by fitting the model to the experimental results. The fitting results gave pc =0.187HMoS2 and K2 /K1 =12, respectively, where HMoS2 is the hardness of the MoS2 coating. From Fig. 2, the general trend of the simulation results agrees fairly well with the experimental observations. In Ref. [24], where wear of diamond-like carbon (DLC) coatings was studied, a value of 0.206HDLC was found to fit experimental results for the critical contact pressure, pc . This is equivalent to 0.63YDLC , where YDLC is the yield strength of the DLC coating. For the MoS2 coating used in this study, the critical contact pressure is equivalent to 0.57YMoS2 . Such levels of contact pressure are much less than that required for overall yielding in the contact surface (≥1.6Y [27]). However, fast micro-crack initiation and propagation via the micro-fatigue mechanism is very likely. For metals, the fatigue limit is normally half the ultimate tensile strength for smooth specimens and is far below 0.5 times the ultimate strength for non-smooth specimens [28]. Although no similar data was available for DLC and MoS2 coatings, the order of magnitude should be valid. This wear mechanism change when contact pressure exceeds the critical contact pressure will obviously lead to a transition in wear rate from low values to some significantly higher values. For MoS2 coatings, the mechanical properties are anisotropic and it is expected that its fatigue limit is lower than that of isotropic materials like DLC. Thus, micro-fatigue crack initiation is easier and it is not unreasonable to expect that the value for the critical contact pressure is lower for MoS2 coating (0.186H MoS2 ) than that for DLC coatings (0.206HDLC ). For DLC coatings, the same value of 12 was obtained from fitting for the ratio of specific wear rates between the high-rate wear regime and the low-rate wear regime, K2 /K1 . The magnitude of this
J. Jiang et al. / Wear 243 (2000) 1–5
value fits many the normal observations in the literature for metals that often shows a difference of one to two orders between ‘mild’ wear and ‘severe’ wear. According to Fig. 2, the size effect becomes less significant for large balls. However, it can be expected that the critical radius of curvature will depend on properties of specimen materials and other experimental conditions. As a general guide, in wear tests using a pin-on-disk wear rig, it is advisable that pin specimens with large radius of curvature be used where possible. In Fig. 5, it is shown that the average friction coefficient of the MoS2 coating increased with increase in ball diameter. This presumably resulted mainly from the fact that the average contact pressure at the rubbing interface decreases with increase in ball diameter. In the literature, it is a common observation [6,29,30] that the friction coefficient of MoS2 coatings decreases with increase in contact pressure as a result of the decrease in shear strength of MoS2 with increased compressive stresses. From Fig. 5, it can be noticed that the average friction coefficient did not increase with ball diameter linearly but tended to level off when the ball diameter exceeded 6.35 mm. This coincides with the variations in specific wear rate as a function of ball diameter in Fig. 2 which shows that the effect of ball diameter vanishes when the ball diameter exceeds 6.35 mm. This observation indicates that contact pressures can indeed play important roles in wear of the MoS2 coating and probably in other tribological systems.
5. Concluding remarks A significant effect of ball size on specific wear rate and average friction coefficient of a MoS2 coating has been observed, the wear rate being significantly higher at the smaller ball sizes than that when larger balls were used. This phenomenon is thought to be a result of wear mode transition as a function of contact pressure at the apparent area of contact from a low-rate ‘mild’ wear regime to a high-rate ‘severe’ wear regime when the contact pressure exceeds some critical value. Simulation results using a previously presented model showed a good agreement with the experimental observations. As a general guide, when carrying out sliding wear tests using a pin-on-disk type of wear rig, small radius of curvature for the pin specimen should be avoided. References [1] W.O. Winer, Molybdenum disulphide as a lubricant: a review of the fundamental knowledge, Wear 10 (1967) 422–452. [2] M.J. Todd, Solid lubrication of ball bearings for spacecraft mechanisms, Tribol. Int. 15 (1982) 331–337. [3] J.-P. Hirvonen, J. Koskinen, J.R. Jervis, M. Nastasi, Present progress in the development of low friction coatings, Surf. Coat. Technol. 80 (1996) 139–150.
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[4] C. Donnet, Advanced solid lubricant coatings for high vacuum environments, Surf. Coat. Technol. 80 (1996) 151–156. [5] T. Spalvins, A review of recent advances in solid film lubrication, J. Vac. Sci. Technol., A 5 (1987) 212–219. [6] E.W. Roberts, Thin solid films in space 23 (1990) 95–104. [7] D.-Y. Yu, J.-A. Wang, J.-L. Ou Yang, Thin Solid Films 293 (1997) 1–5. [8] J.R. Lince, M.R. Hilton, A.S. Bommannavar, Metal incorporation in sputter-deposited MoS2 films studied by extended X-ray-absorption fine-structure, J. Mater. Res. 10 (1995) 2091–2105. [9] G. Weise, N. Mattern, H. Hermann, A. Teresiak, I. Bacher, W. Bruckner, H.D. Bauer, H. Vinzelberg, G. Reiss, U. Kreissig, M. Mader, P. Markschlager, Preparation, structure and properties of MoSx films, Thin Solid Films 298 (1997) 98–106. [10] J. Moser, F. Levy, F. Bussy, Composition and growth mode of MoSx sputtered films, J. Vac. Sci. Technol., A 12 (1994) 494–500. [11] P.D. Fleischauer, R. Bauer, Chemical and structural effects on the lubrication properties of sputtered MoS2 films, Tribol. Trans. 31 (1988) 239–250. [12] P.D. Fleischauer, Effects of crystallite orientation on environmental stability and lubrication properties of sputtered molybdenum disulphide thin films, ASLE Trans. 27 (1983) 82–88. [13] V. Buck, Preparation and properties of different types of sputtered MoS2 films, Wear 114 (1987) 263–274. [14] J.F. Archard, Contact and rubbing of flat surfaces, J. Appl. Phys. 24 (1953) 981–988. [15] J.F. Archard, W. Hirst, The wear of metals under unlubricated conditions, Proc. R. Soc. London, Ser. A 236 (1956) 397–410. [16] E. Rabinowicz, Friction and Wear of Materials, Wiley, New York, 1965. [17] J. Halling, A contribution to the theory of mechanical wear, Wear 34 (1975) 239–249. [18] D.J. Whitehouse, J.F. Archard, The properties of random surfaces of significance in their contact, Proc. R. Soc. London, Ser. A 316 (1970) 97–121. [19] V.K. Jain, S. Bahadur, Development of wear equation for polymer metal sliding in terms of fatigue and the topography of the sliding surfaces, Wear 60 (1980) 237–288. [20] N.P. Suh, H.-C. Sin, On prediction of wear coefficients in sliding wear, ASLE Trans. 26 (1983) 360–366. [21] N.C. Welsh, Frictional heating and its influence on the wear of steel, J. Appl. Phys. 28 (1957) 960–968. [22] N.C. Welsh, The dry wear of steels: I. The general pattern of behaviour, Philos. Trans. R. Soc. London, Ser. A 257 (1965) 31– 50. [23] N.C. Welsh, The dry wear of steels: II. Interpretation and special features, Philos. Trans. R. Soc. London, Ser. A 257 (1965) 51–70. [24] Jiaren Jiang, R.D Arnell, On the running-in behaviour of DLC coatings under the ball-on-disk contact geometry, Wear 217 (1998) 190–199. [25] J.A. Williams, J.H. Morris, A. Ball, The effect of transfer layers on the surface contact and wear of carbon-graphite materials, Tribol. Int. 30 (1997) 663–676. [26] Jiaren Jiang, F.H. Stott, M.M. Stock, The role of triboparticulates in dry sliding wear, Tribol. Int. 31 (1998) 245–256. [27] K.L. Johnson, Contact Mechanics, Cambridge Univ. Press, Cambridge, 1985, pp. 153–155, 104–106. [28] N.E. Dowling, in: Mechanical Behaviour of Materials, Prentice-Hall, Englewood Cliffs, NJ, 1993, pp. 361–363. [29] J.-P. Hirvonen, J. Koskinen, J.R. Jervis, M. Nastasi, Present progress in the development of low friction coatings, Surf. Coat. Technol. 80 (1996) 139–150. [30] C. Donnet, J.M. Martin, Th. Le Monge, M. Belin, Super-low friction of MoS2 coatings in various environments, Tribol. Int. 29 (1996) 123–128.
The influence of ball size on tribological behaviour of MoS2 coating tested on a ball-on-disk wear rig Jiaren Jiang a,∗,1 , R.D. Arnell a , Gajendra Dixit b a
Centre for Advanced Materials and Surface Engineering, University of Salford, Salford M5 4WT, UK b Department of Applied Mechanics, Maulana Azad College of Technology, Bhopal 462 007, India Received 8 April 1999; received in revised form 8 February 2000; accepted 9 February 2000
Abstract The effect of ball diameter on wear and friction of a molybdenum disulphide (MoS2 ) coating deposited using the closed-field magnetron sputtering technique has been investigated on a ball-on-disk wear rig sliding against uncoated steel ball-bearing balls. It was observed that the wear rate of the coating decreased significantly with increase in ball diameter. Correspondingly, the average friction coefficient in the steady state of sliding increased with increase in ball diameter. Under the present experimental conditions, the effect of ball size on wear and friction of the coating vanished when the diameter of the ball exceeded 6.35 mm. This phenomenon is thought to be a result of wear mode transition as a function of contact pressure at the apparent area of contact from a low-rate ‘mild’ wear regime to a high-rate ‘severe’ wear regime when the contact pressure exceeds some critical value. Simulation results using a previously presented model based on the above assumption showed a good agreement with the experimental observations. © 2000 Elsevier Science S.A. All rights reserved. Keywords: Molybdenum disulphide (MoS2 ) coating; Magnetron sputtering; Tribological behaviour; Wear; Effect of experimental conditions
1. Introduction The pin-on-disk wear rig is one of the most commonly used configurations for sliding wear tests in the laboratory. To facilitate accurate alignment between the pin and the disk surfaces, a dome-ended pin specimen is often used. While this arrangement reduces errors in wear measurements, some complications are brought into the interpretation of the obtained experimental results by the non-uniform distribution of contact pressures over the apparent area of contact. For example, although the resultant wear and wear rate would not be influenced by the geometry of the pin if the Archard’s wear behaviour is followed, the wear results will be different for different geometry of the pin specimen if wear of the materials being studied is dependent on contact pressures. There have been few studies carried out to elucidate the effect of the use of the dome-ended pin on the observed wear behaviour. Molybdenum disulphide (MoS2 ) is a unique material as a solid lubricant applied in a vacuum and in inert gases and has been widely studied over the past several decades [1–7]. The application of sputter deposition techniques in produc∗
Corresponding author. Current address: IMTI-NRC, 800 Collip Circle, London, ON, Canada N6G 4X8. 1
ing thin MoS2 coatings has significantly improved the tribological properties of the coatings and provided the potential of modifying and tailoring properties of the coatings by incorporating various elements/compounds [7,8], changing the chemical compositions [9,10] or modifying crystal structure and intercrystallite slip [11,12]. In this work, the effect of ball diameter on tribological behaviour of a MoS2 coating deposited on M42 tool steel substrate using magnetron sputtering technique has been investigated on a ball-on-disk wear rig. The results are explained on the basis of a model considering the effect of contact pressure on wear and wear transition. 2. Experimental details The MoS2 coating was deposited onto M42 tool steel disk substrate using the closed field magnetron sputtering technique. Fig. 1 shows the surface morphology of the as-deposited coating observed using an SEM. According to Buck [13], the morphology corresponds to the stoichiometry for MoS2 coatings/films deposited using the conventional r.f. sputtering technique. The morphology in Fig. 1 seems to suggest that this coating had an amorphous structure and probably a non-stoichiometric composition of MoSx with x≥1. However, since this coating has been deposited
0043-1648/00/$ – see front matter © 2000 Elsevier Science S.A. All rights reserved. PII: S 0 0 4 3 - 1 6 4 8 ( 0 0 ) 0 0 3 4 1 - 0
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J. Jiang et al. / Wear 243 (2000) 1–5
Fig. 1. SEM photograph showing the surface morphology of the as-deposited MoS2 coating.
using a different technique with higher bombardment energy, more detailed study may be necessary fully to characterise the coating. The thickness of the MoS2 layer was approximately 2.6 m measured using the ball crater technique. The hardness and reduced elastic modulus of the MoS2 coating, measured using a nanoindenter, were 7.7±0.3 and 150±4 GPa, respectively. The surface roughness of the coating was Ra 0.37 m. Wear tests were carried out on a ball-on-disk wear rig in dry air with a relative humidity of approximately 7%; this humidity was maintained by passing compressed dry air through saturated sodium hydroxide solution before inleting the gas to the chamber enclosing the test specimens. Uncoated steel ball bearing balls were used as the counter specimen. The diameters of balls used were 3.18, 4.75, 6.35, 7.93, 9.52 and 12.69 mm. Before wear tests, the ball and the disk specimens were ultrasonically cleaned in acetone for approximately 10 min and rinsed using acetone to remove grease and contaminants that may be present on the specimen surfaces. A normal load of 35 N was applied by applying a dead weight to the ball holder. Under this load condition, the initial mean Hertzian contact pressures at the rubbing interface for the various ball sizes were 1180, 903, 744, 642, 568, and 469 MPa, respectively. The sliding speed was 0.25 m s-1 and the total sliding distance was between 400 and 1500 m. During sliding, friction forces were continually measured and logged onto a microcomputer. Wear volumes of the disk specimens were measured using a profilometer linked to a computer. At least eight measurements were made along the circumference of each wear track. The fluctuations in cross-section areas obtained on one wear track were normally within 12% of the average value; the apparently large fluctuations were mainly due to the unevenness of the wear track along the circumference. However, the relative precision of the measurement technique itself was within 2.4% as estimated by repeated measurements on the same wear track. Wear of the ball was estimated by measuring the sizes of the major and the minor axes of the elliptical-shaped wear scar under a microscope and assuming that the scar surface
Fig. 2. The effect of ball diameter on specific wear rates of the MoS2 coated disks (䉱) and the balls (+). The solid lines show the simulation results using the elastic foundation wear model presented in Ref. [23].
was ideally flat. Apparently, this method over-estimates wear volumes of the balls because scars formed on the ball normally had some convex curvatures. Wear has been presented in the form of Archard’s specific wear rate in mm3 N-1 m-1 . In the wear curve measurements, sliding was stopped at certain time intervals to measure the wear volume of the specimens and the sliding was restarted again. A profilometer was attached to the wear rig; this allowed the specimens to remain fitted on the specimen holders during the wear volume measurement so that the re-alignment between the ball and the disk contact areas after the measurement was very good.
3. Results The variation of specific wear rate of the coating as a function of ball diameter is shown in Fig. 2. In general, the specific wear rate decreased with increase in ball diameter. Wear rate at the ball diameter of 3.18 mm was significantly higher than that for the larger ball sizes, although the scatter was larger for the small ball. Some typical wear curves at the different ball diameters are shown in Fig. 3. Wear increased almost linearly with
Fig. 3. Some typical wear curves of the MoS2 coating using balls of different diameters.
J. Jiang et al. / Wear 243 (2000) 1–5
Fig. 4. Typical variation of friction coefficient as a function of sliding distance at different ball sizes. The general trend was similar for the various ball sizes.
sliding distance at small and large ball sizes; however, transitions in wear rate from a high initial rate to a lower rate after some time of sliding was apparently observed at the intermediate diameter of balls. The variation in friction coefficient as a function of sliding distance was very similar for the various diameters of balls. Fig. 4 shows two typical examples. After a short distance of sliding, the friction coefficient increased from some low initial value to slightly higher values and then the friction coefficient fluctuated around some constant average value. However, the average values of friction coefficient showed a steady increase with increase in diameter of the ball, as shown in Fig. 5. The values in Fig. 5 were calculated using friction data after a sliding distance of 200 m.
4. Discussion In sliding wear, the Archard’s Wear Law [14,15] is one of the most basic and most widely used equations; it was derived on the basis of ‘adhesive wear’ concept and predicts a linear increase in wear volume as a function of sliding distance. The same form of wear equation has been obtained based on abrasive cutting wear [16], fatigue wear [17–19]
Fig. 5. Variation of average friction coefficient as a function of ball diameter. The values were calculated using data for sliding beyond 200 m.
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and delamination theory of wear [20]. The Archard’s type of wear cannot explain the significant effect of ball diameter on wear observed in the present study because according to this type of prediction, the total wear volume is only dependent on the total load irrespective of the distribution of pressures over the apparent area of contact. In practice, a transition in wear rate from a high rate at the initial stages of sliding to a low rate after certain distance of sliding is frequently observed. For metals, this transition has been classified as a transition from ‘severe’ wear to ‘mild’ wear and has been attributed to the formation of oxide layers on wear surfaces by Archard [14] and Archard and Hirst [15]. Nevertheless, according to the systematic study of Welsh [21–23] on wear transitions of steels under various conditions, it was shown that the establishment of oxide on the rubbing surfaces was not necessary for the severe to mild wear transition to occur when the substrate hardness exceeded a certain value. In addition, such wear transition has also been observed on non-metal materials (e.g. Fig. 3 and Refs. [24,25]). Most recently, Williams et al. [25] analysed the wear rate transition of carbon-graphite materials as a function of sliding distance. In their model, it is assumed that with sliding, wear debris particles are accumulated/agglomerated on the rubbing surfaces; such agglomeration reduces the maximum contact stresses at the real areas of contact. It was further assumed that wear at areas contacting the agglomerated particles with low contact stresses is negligible and the Archard’s wear law is applicable at the points of intense contact load. Thus, the wear rate decreases with sliding due to the decrease in contact stress. However, during the sliding of the MoS2 against the very smooth steel balls in this study, accumulation/agglomeration of wear debris particles on the MoS2 disk specimen was not significant. Although some material transfer from the disk to the ball surface indeed occurred, this transfer film/layer could not have caused any significant decrease in contact pressures at the real areas of contact because the initial surface of the ball was already very smooth. The development of compact wear debris particle layers have also been shown to have significant and beneficial effect on wear transitions in dry sliding wear of metals (e.g. Ref. [26]) where comminution of wear debris particles are necessary for such layers to form on wear surfaces. However, in the present study, it was noticed that transfer layers were developed on the pin/ball surface after a very short sliding distance and were observed on balls of all the sizes used. Little wear was observed on the ball specimens, presumably due to the protection effect of the transfer layer. Thus, it is reasonable to state that the development of compact wear debris particle layers or transfer layers on wear surfaces was not the dominant factor in controlling wear transitions of the MoS2 coating under the current experimental conditions. According to Welsh’s [21–23] observations, the Archard’s specific wear rate of carbon steels is a function of normal load; the dominant wear mode changed from ‘mild’ wear to ‘severe’ wear when load exceeded some critical value,
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resulting in an increase in specific wear rate by approximately two orders. This critical transition load was probably associated with the deformation mode in the generation of wear debris particles at the rubbing surfaces, extensive plastic deformation being involved at the real contact region in the ‘severe’ wear mode. Similar relationship between wear mode and normal load or contact pressure can be expected in other materials such as MoS2 , although the characteristics of wear surfaces in the two wear modes may be different for different materials. When the maximum stresses in the region of real areas of contact exceeds some critical value, depending on the properties of the material, severe plastic deformation and probably low cycle fatigue become the dominant mechanism in the generation of wear debris particles. Under such conditions, the wear rate is considerably higher than that at pressures when elastic contact prevails. Based on such argument, the authors [24] presented a wear model describing the running-in process of wear on a ball-on-disk wear rig and showed a good agreement with experimental observations. In this model, the actual contact pressures and their distribution across the rubbing interface were described using contact mechanics. Specific wear rate was represented as a function of the actual contact pressure at a given point within the rubbing interface. Based on Welsh’s [21–23] findings, it was assumed that the dominant wear regime would change from a low-rate regime to a high-rate regime when the real contact pressure exceeds some critical value — the critical contact pressure (pc ). At the same time, the specific wear rate would increase from a lower value to a significantly higher value. Values of the critical contact pressure and the ratio between the high-rate wear regime and the low-rate wear regime are dependent on properties of the material(s) being studied. Thus, on a ball-on-disk type of wear machine, the specific wear rate near the central part of the wear track is expected to be much higher than that at regions towards the edges of the track because the contact pressure near the centre is the highest and may exceed the critical contact pressure for the material being studied. With the progress of sliding, both the ball and the disk are worn, leading to increased apparent area of contact and greatly decreased contact pressures within the rubbing interface. As a result, the size of high-rate-wear-regime-dominated region decreases with sliding distance. Depending on load and wear properties of the rubbing materials, the low-rate wear regime may eventually become dominant over the whole apparent area of contact and a transition in wear rate from a high rate to a low rate can thus be observed (e.g. when using a 6.35-mm ball in this study as shown in Fig. 3). Obviously, under certain conditions where the load is high and/or the critical contact pressure of the material is low, the high-rate wear regime will be dominant during most of the sliding time, e.g. at a ball diameter of 3.18 mm in the present study (Fig. 3), the wear rate being high throughout the sliding. On the other hand, the low-rate wear regime may be dominant over the whole apparent area of contact and throughout most of the sliding distance if the contact pressure is low and/or
the critical contact pressure for the material is high (Fig. 3, using a 12.69-mm ball). Based on this model, when a small ball is used, the contact pressures over a large proportion of the apparent area of contact can be expected to be above the critical contact pressure. As a result, a high-rate wear regime dominates much of the sliding wear process (Fig. 2), the overall wear rate at a small ball size being much higher than that when a larger ball was used. Based on the above concept, if the distribution of contact pressures within the rubbing interface and the transition property of the material are known, then the wear process of the sliding system can be simulated. In Ref. [24], the elastic-foundation method developed in contact mechanics [27] was applied to calculate the contact pressures between the ball and the disk specimens. Applying this elastic-foundation wear model presented in Ref. [24], using the mechanical properties for the MoS2 coating and the steel ball, the variation of specific wear rate as a function of ball diameter has been simulated. The results for sliding distances of 750 and 1500 m are plotted in Fig. 2 as solid lines. In the simulation, values for the critical contact pressure, pc , and the ratio of specific wear rates between the high-rate wear regime and the low-rate wear regime, K2 /K1 , were obtained by fitting the model to the experimental results. The fitting results gave pc =0.187HMoS2 and K2 /K1 =12, respectively, where HMoS2 is the hardness of the MoS2 coating. From Fig. 2, the general trend of the simulation results agrees fairly well with the experimental observations. In Ref. [24], where wear of diamond-like carbon (DLC) coatings was studied, a value of 0.206HDLC was found to fit experimental results for the critical contact pressure, pc . This is equivalent to 0.63YDLC , where YDLC is the yield strength of the DLC coating. For the MoS2 coating used in this study, the critical contact pressure is equivalent to 0.57YMoS2 . Such levels of contact pressure are much less than that required for overall yielding in the contact surface (≥1.6Y [27]). However, fast micro-crack initiation and propagation via the micro-fatigue mechanism is very likely. For metals, the fatigue limit is normally half the ultimate tensile strength for smooth specimens and is far below 0.5 times the ultimate strength for non-smooth specimens [28]. Although no similar data was available for DLC and MoS2 coatings, the order of magnitude should be valid. This wear mechanism change when contact pressure exceeds the critical contact pressure will obviously lead to a transition in wear rate from low values to some significantly higher values. For MoS2 coatings, the mechanical properties are anisotropic and it is expected that its fatigue limit is lower than that of isotropic materials like DLC. Thus, micro-fatigue crack initiation is easier and it is not unreasonable to expect that the value for the critical contact pressure is lower for MoS2 coating (0.186H MoS2 ) than that for DLC coatings (0.206HDLC ). For DLC coatings, the same value of 12 was obtained from fitting for the ratio of specific wear rates between the high-rate wear regime and the low-rate wear regime, K2 /K1 . The magnitude of this
J. Jiang et al. / Wear 243 (2000) 1–5
value fits many the normal observations in the literature for metals that often shows a difference of one to two orders between ‘mild’ wear and ‘severe’ wear. According to Fig. 2, the size effect becomes less significant for large balls. However, it can be expected that the critical radius of curvature will depend on properties of specimen materials and other experimental conditions. As a general guide, in wear tests using a pin-on-disk wear rig, it is advisable that pin specimens with large radius of curvature be used where possible. In Fig. 5, it is shown that the average friction coefficient of the MoS2 coating increased with increase in ball diameter. This presumably resulted mainly from the fact that the average contact pressure at the rubbing interface decreases with increase in ball diameter. In the literature, it is a common observation [6,29,30] that the friction coefficient of MoS2 coatings decreases with increase in contact pressure as a result of the decrease in shear strength of MoS2 with increased compressive stresses. From Fig. 5, it can be noticed that the average friction coefficient did not increase with ball diameter linearly but tended to level off when the ball diameter exceeded 6.35 mm. This coincides with the variations in specific wear rate as a function of ball diameter in Fig. 2 which shows that the effect of ball diameter vanishes when the ball diameter exceeds 6.35 mm. This observation indicates that contact pressures can indeed play important roles in wear of the MoS2 coating and probably in other tribological systems.
5. Concluding remarks A significant effect of ball size on specific wear rate and average friction coefficient of a MoS2 coating has been observed, the wear rate being significantly higher at the smaller ball sizes than that when larger balls were used. This phenomenon is thought to be a result of wear mode transition as a function of contact pressure at the apparent area of contact from a low-rate ‘mild’ wear regime to a high-rate ‘severe’ wear regime when the contact pressure exceeds some critical value. Simulation results using a previously presented model showed a good agreement with the experimental observations. As a general guide, when carrying out sliding wear tests using a pin-on-disk type of wear rig, small radius of curvature for the pin specimen should be avoided. References [1] W.O. Winer, Molybdenum disulphide as a lubricant: a review of the fundamental knowledge, Wear 10 (1967) 422–452. [2] M.J. Todd, Solid lubrication of ball bearings for spacecraft mechanisms, Tribol. Int. 15 (1982) 331–337. [3] J.-P. Hirvonen, J. Koskinen, J.R. Jervis, M. Nastasi, Present progress in the development of low friction coatings, Surf. Coat. Technol. 80 (1996) 139–150.
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