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A wear map of bearing steel lubricated by silver films
Wear 255 (2003) 883–892
A wear map of bearing steel lubricated by silver films Seung Ho Yang a , Hosung Kong a,∗ , Eui-Sung Yoon a , Dae Eun Kim b a
Tribology Research Center, Korea Institute of Science and Technology, 39-1 Hwawolkok-dong Sungbuk-ku, Seoul, South Korea b Department of Mechanical Engineering, Yonsei University, Seoul, South Korea
Abstract Wear map of bearing steel lubricated by silver film has been constructed to delineate the wear transition behavior with the change in operating conditions. Experiments were performed in dry sliding conditions using two types of ball-on-disk type test rigs under the contact pressure of 100–1000 MPa and the sliding speed of 20–1000 mm/s in ambient air. For the silver coating, an ultra thin IBAD silver bond layer was firstly deposited on AISI 52100 steel surfaces and then a relatively soft silver film was deposited by a thermal evaporation method onto the first layer. This functionally gradient film showed a great improvement in the life, mainly owing to the better bond strength. In order to build up a general framework on the tribological behavior of the functionally gradient silver films, all test data were plotted on a map whose axes are contact pressure and sliding speed. As a result, three main regimes were clearly identified: (i) elastic/plastic deformation of silver coating without failure, (ii) mild wear regime after initial failure of silver coating and (iii) severe wear regime. In the mild wear regime, the contact surfaces were covered with transfer layers of agglomerated wear particles. The transfer layer acted as a protective layer and resulted in low friction after initial failure of the coating. The formation of transfer layer was suppressed by several destructive actions, when the sliding speed was high. And, above a critical sliding speed, no transfer layer was able to form. In the discussion, an empirical model that could explain the existence of critical sliding speed was proposed. © 2003 Elsevier Science B.V. All rights reserved. Keywords: Wear map; Transfer layer; Functionally gradient film; Critical velocity
1. Introduction Soft metallic coating has been focused as a good solid lubricating film for the space bearing element [1]. The solid lubricating characteristic of thin soft metallic films has been well explained by the Tabor’s adhesion theory of friction that described how the effective lubricating characteristics could be obtained between two materials in sliding contact [2]. Soft metals also have excellent heat conductivity since they diffuse out the frictional heats generated at the contacting interfaces [3]. However, soft metallic coatings usually have very little mutual solubility with iron-based steel alloys, so they generally show poor adhesion to the steel surfaces. This has been a critical restriction for applying the soft metallic coatings to the practical machine elements. As a result, it has been normally recognized that initial failure of coated layer is the definite end of service life. Recently, the formation of transfer layer has been one of the great concerns to many tribologists. Because it could alter the friction and wear characteristics very seriously even ∗ Corresponding author. Tel.: +82-2-958-5655; fax: +82-2-958-5659. E-mail address: [email protected] (H. Kong).
after the initial failure of coating occurred. Lots of work have focused on the generation mechanism of transfer layer [4–9] and the role of the transfer layer on friction and wear [10–13]. Yet all of those works dealt in a specific point of view under certain operating conditions. Therefore, the generation mechanism and the role of transfer layer need more elaborated works. In this work the tribological characteristics of silver transfer layers were focused. Functionally gradient silver film was tested to study the effect of the film bond strength on the tribological characteristics of the film and also the characteristics of the transfer layer were evaluated. Tests were performed in various ranges of load and speed and the results were summarized in a wear map constructed with the axes of contact pressure and sliding speed.
2. Experimental details Pure (99.99%) silver was coated on the hardened AISI 52100 steel disks with a functionally gradient silver film deposition method. Friction and wear characteristics of the coatings were evaluated with two ball-on-disk type tribo-testers. In order to cover a wide range of load
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conditions, both a macro-tribo-tester and a miniature tribo-tester were used to cover wide range of loads. All the tribo-tests were performed in unidirectional dry sliding under various ranges of load and speed in an ambient condition. After the tests, the size of wear scars of tested balls were measured with an optical microscope and the wear coefficient (K) was calculated as follows: HV K= (1) LD where H is the hardness of hardened AISI 52100 steel, V the worn volume of the ball, L the sliding distance and D the applied normal load. 2.1. Tribo-testers When the maximum Hertzian contact pressure is greater than 400 MPa (the applied normal load is 1.1 N), a macro-tribo-tester was used. Test loads were applied by using dead weight. Friction force was measured with a load cell (maximum load: 3 kgf) installed at a right angle to the direction of sliding. The apparatus was described in detail elsewhere [14]. When the applied load was less than 1.1 N, which is the minimum load of the macro-tribo-tester, tests were performed with a miniature tribo-tester shown in Fig. 1. Maximum load of the miniature tribo-tester was 100 gf. 2.2. Specimens Hardened AISI 52100 steel disks (hardness: HR C 57–61) of 90 mm in diameter and 6 mm in thickness were used for the tests performed with a macro-tribo-tester. For a miniature tribo-tester, hardened AISI 52100 steel disks (hardness: HR C 57–61) of 40 mm in diameter and 6 mm in thickness were used. As for the specimen balls, hardened AISI 52100 steel
Fig. 1. A close-up view of the (a) macro-tribo-tester and (b) miniature tribo-tester.
Fig. 2. Schematic figure showing the ion beam assisted coating system.
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balls (hardness: HR C 55–60) of 12.7 mm in diameter were used. 2.3. Coating methods In order to increase the bond strength of a thermally evaporated silver coating, functionally gradient coating was performed using accelerated argon ion bombardment. The coating system was illustrated schematically in Fig. 2. A hollow cathode type ion gun with the mouth diameter of 30 mm was installed in the thermal evaporation chamber. Specimen disks were ultrasonically cleaned with n-hexane for 2 min. Final cleaning was performed by the argon ion bombardment (1.5 keV) in the coating chamber under vacuum (1 × 10−2 Pa) before coating.
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Table 1 Coating conditions Thin seed layer formation Ion bombardment Vacuum (Pa) Ion beam acceleration (eV) Argon gas flow rate (sccm) Thermal evaporation Evaporant Heating current (A) Thermal evaporation Vacuum (Pa) Evaporant Heating current (A)
1 × 10−2 1500 2.0 Silver (0.05 g) 500 1 × 10−3 Silver (2.0 g) 500
Fig. 3. Wear coefficients and friction coefficients of functionally gradient silver coatings as a function of maximum Hertzian contact pressure.
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Functionally gradient silver coating was performed by using a thermal evaporation of silver assisted by accelerated argon ion bombardment. For functionally gradient coating, a very thin atomically mixed silver layer of 8 nm in thickness was firstly deposited onto the steel disks by (using) the ion beam assisted thermal evaporation method (of silver assisted by the bombardment of accelerated (1.5 keV) argon ions). It was intended that this ultra-thin layer be utilized as an atomically mixed seed layer for the subsequent thermally evaporated silver coating. For the second step, silver film was deposited by a thermal evaporation method onto the thin seed layer. Silver coating of 2 m in thickness was finally deposited on the steel disk specimen. Detailed coating conditions are listed in Table 1.
Table 2 Test conditions Test conditions
Roughness of substrate, Ra (nm) Roughness after coating, Ra (nm) Coating thickness (m) Speed (mm/s) Normal force (N) [maximum Hertzian contact pressure (MPa)] Lubrication Environments Temperature
Specimen Ball
Disk
10 10 (no coating) – 20, 100, 400, 1000 0.02 [100], 0.14 [200], 0.47 [300], 1.10 [400], 2.19 [500], 5.88 [700], 17.64 [1000] Dry Ambient air Room temperature
100 20 2.0
2.4. Test conditions Tests were performed in dry sliding conditions at the load range 0.02–17.64 N and the speed range 20–1000 mm/s for a fixed sliding distance of 1000 m. Detailed test conditions are listed in Table 2. Each test was repeated more than three times.
3. Results and discussion 3.1. Effect of contact pressure and sliding velocity Tests were performed under various normal loads at the constant sliding speed of 20 mm/s with the functionally gradient silver films. The wear coefficients and friction coef-
ficients were shown in Fig. 3. When the contact pressure was less than 400 MPa, no failure was found. As the coating film supported the applied load without the film failure, mild wear (K = 1.0 × 10−5 to 1.0 × 10−6 ) was observed. When the contact pressure was higher than 400 MPa, however, the silver coating failed and the wear coefficient rose abruptly. The frictional behaviors were different from the results of wear coefficients. The coating failure did not always accompany a high friction coefficient. In order to find the changes at the contact surfaces, EPMA line profiling analyses for the disks were performed after tests, as shown in Fig. 4. When the contact pressure was low (Pmax = 100 MPa), silver was not detected at the track surface. On the other hand, under the load of 400 and 700 MPa, transfer of silver was observed at the track surface after the initial coating failure. It was
Fig. 4. EPMA line profiling analysis results of the disk surfaces with functionally gradient silver coatings under the various conditions of contact pressure.
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Fig. 5. Wear coefficients and friction coefficients of functionally gradient silver-coated surfaces as a function of sliding speed.
found that such a transfer layer played an important role on the friction behavior afterward. Fig. 5 shows the wear and friction behaviors of functionally gradient silver films at the various sliding speeds of 20, 100, 400 and 1000 mm/s under the constant contact pressure of 400 MPa. When the sliding speed was less than 100 mm/s, relatively low wear coefficients (1.0 × 10−3 to 1.0 × 10−4 ) were obtained. However, at the higher sliding speeds (400 and 1000 mm/s), the wear coefficient increased by one order of magnitude higher (1.0 × 10−2 to 1.0 × 10−3 ). On the other hand, the friction coefficient increased with the sliding speed. Results of the EPMA line profile analyses (Fig. 6) showed that when the sliding speed was relatively low (less than 100 mm/s), silver transfer layer was noticeable at the track surfaces. The formation of such a transfer layer explained why the low wear coefficient was obtained even after the
initial coating failure. The transfer layer acted as a solid lubricant. When the sliding speed was 1000 mm/s, however, the transfer layer could not be found at the track surfaces but only scattered wear debris was found at the track surfaces, and the resultant wear coefficient and friction coefficient were high. The above results suggested that both the contact pressure and the sliding speed affected seriously the formation of transfer layer and the tribological behavior of silver coating films. 3.2. Construction of wear map The tests were performed in various ranges of load and speed and the results were summarized in a wear map constructed with the axes of contact pressure and sliding speed as shown in Fig. 7. Maximum Hertzian contact pressure and
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Fig. 6. EPMA line profiling analysis results of the disk surfaces with functionally gradient silver-coated surfaces under the various conditions of sliding speeds.
sliding speed were selected as the axes of the map because the P × V (pressure × speed) value corresponds to the mechanical input energy to the coating films [15,16]. All of the wear coefficients and the coefficients of friction were plotted numerically at the different combinations of contact pressure and sliding speed.
Fig. 7. The characteristic wear regimes of the functionally gradient silver-coated surfaces (italicized wear coefficients and coefficients of friction were given numerically).
Based on the relative magnitude of the given wear coefficients, three different characteristic regimes (Regime I, Regime II and Regime III) could be identified as shown in Fig. 7. 3.2.1. Regime I In this regime, silver film sustains the applied load without any failure of the coating. The coated silver film prevented the ball and track surfaces from direct contact, which resulted in mild wear during the sliding (Fig. 8). The onset of the film failure can occur by increasing the contact pressure further. It was found that the criteria for coating failure depend on whether the given mechanical input energy (P × V value) exceeds some critical value or not. 3.2.2. Regime II Regime II demarcated the region where relatively low wear coefficients (order of 10−4 to 10−5 ) were still sustained after the silver coating apparently failed. The low wear coefficients were resulted mainly from the transfer layers of agglomerated wear debris on the contact surfaces (Fig. 9). After the apparent failure of coated silver films, wear particles were re-grinded or plastically deformed to generate much more fine particles. These fine particles are supposed to have very high energy due to its large surface area, so they tend to easily stick to the adjacent surface in order to reduce the total free energy [17]. The transfer layers acted as a protective layer, they decrease the wear and friction to a low and stable value. Therefore, it was found that the formation of transfer layer
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Fig. 8. EPMA line profiling analysis results of the ball and disk surfaces tested in Regime I.
Fig. 9. EPMA line profiling analysis results of the ball and disk surfaces tested in Regime II.
provided beneficial effects to the contact surfaces even after the coating film failed. 3.2.3. Regime III Regime III was demarcated the region of relatively high wear coefficients (K > 10−3 ). As seen in Fig. 10, transfer layer could not be observed in this regime. Friction and wear were higher than those of Regimes I and II. In this regime, the film failure apparently meant the end of service life since there is no protective layer after the failure. 3.3. Wear map of functionally gradient silver coating Based on the above characteristic regimes (I, II and III), a wear map of the functionally gradient silver coating was
constructed as shown in Fig. 11. It was found that the border between Regimes II and III in Fig. 11 is mostly affected by the sliding speed. It also suggested that a critical sliding speed (vcr ) exists in the formation of transfer layer, above which no material transfer is able to form. 3.4. Critical sliding velocity Transfer layer formation means the stable attachment of wear particles on the sliding surfaces. Wear particles generally have very high surface area to volume ratios and high surface energy. So, the wear particles have the tendency to stick to adjacent surface. But not every wear particle can stick to the surface stably because disturbing forces exists. Therefore, the formation of transfer layer could be
Fig. 10. EPMA line profiling analysis results of the ball and disk surfaces tested in Regime III.
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The work of adhesion (γ) between a transfer island and the surface of bearing steel is approximately 70 erg/cm2 when the surface energy of silver (γ A ) is 920 erg/cm2 and the surface energy (γ B ) of bearing steel is 1500 erg/cm2 [18]. Let us consider a cylindrical silver transfer island having the radius (R) of 1 m and height of 1 m. The density of pure silver is 10 g/cm2 . The radius of wear track (r) was 15 mm and the measured water wetting angle of silver (θ A ) and bearing steel (θ B ) was 80◦ and 60◦ , respectively. When the centrifugal force is the dominant detaching force, the critical velocity (vcr ) is calculated about 1000 mm/s. According to Fig. 5, the critical velocity was approximately 100–1000 mm/s. Therefore, it could be concluded that above arguments are reasonable.
4. Conclusions
Fig. 11. A wear map of the functionally gradient silver coating (italicized wear coefficients and coefficients of friction were given numerically).
competitive depending on the two forces, such as detaching forces (centrifugal force and asperity impact force) and attaching forces (adhesion force and gravitational force of wear particles). The centrifugal force and the asperity impact force increase with approximately the square of the sliding velocity. So the tendency of transfer layer formation decreases with the sliding velocity. Therefore, it could be noted that there exists a critical velocity when the transfer layer formation is not able to form. When the centrifugal force is the dominant detaching force, transfer island is stable when the following Eq. (2) is satisfied. Details of the simple analyses used in these calculations are given in Appendix A. 1/2 µr 3 mT g+ πγR+2πRγL (cos θA +cos θB ) ≥vcr mT 2 (2) Or, when the asperity impact force is the dominant detaching force, transfer island is stable when Eq. (3) is satisfied. 1/2 µ(x) 3 mT g+ πγR+2πRγL (cos θA +cos θB ) ≥vcr mT 2 (3) where µ is the static friction coefficient of transfer island, mT the self-mass of transfer island, g the gravitational acceleration, γ the work of adhesion between transfer island and disk surface, R the radius of transfer island, γ L the surface tension of water, θ A the water wetting angle of the transfer island, θ B the water wetting angle of the disk surface, x the deformation by asperity impact, vcr is the critical velocity. In order to evaluate the above reasoning, the critical velocity was calculated roughly using Eqs. (2) and (3) as follows.
The tribological role of transfer layer of silver-coated surfaces was summarized as follows: (1) Silver transfer layers could protect the contact surface beneficially even after the initial coating failure occurs. (2) Wear maps of silver coating films were constructed using the axes of contact pressure and sliding speed, where three different characteristic wear regimes were clearly identified such as Regime I (no failure, mild wear), Regime II (dominant region for transfer layer, moderate wear) and Regime III (no transfer layer, severe wear). (3) The formation of transfer layer was strongly affected by the operating conditions, especially by the sliding speed. It was suppressed when the sliding speed was high and, hence, above a critical sliding speed, no transfer layer was able to form. (4) Centrifugal force and asperity impact force affected seriously the critical velocity.
Acknowledgements The authors would like to thank the Ministry of Science and Technology and Critical Technology 21 program (Machinery Design Technology Enhancement) and the National Research Laboratory Program for their support and interest in this work.
Appendix A In order to explain how the formation of the transfer layer depends on the sliding speed, various forces that may influence the transfer layer formation were discussed. A free body diagram of one transferred mass on a flat disk surface was constructed in Fig. 12. The wear particles are deformed by concentrated contact pressure, so they generally reveal plate-like shapes. In this work, one cylindrical
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So they could be dealt separately. The transfer island is stable when Eqs. (A.4) and (A.5) are satisfied
Fig. 12. A free body diagram of a transfer island formation.
transfer island was simply considered where major forces were shown in the figure. The formation of transfer layer could depend on two competitive forces; detaching forces such as centrifugal force (Fc ), asperity impact force (Fi ) and vertical component of the centrifugal force (Fcz ), and attaching forces such as adhesion force (Fad ) and gravitational force of wear particles (Fg ). In this work, the run-out in the vertical direction was less than 1 m so the vertical component of the centrifugal force (Fcz ) was neglected. When the transfer island is stably stuck to the disk surface, the force equilibrium condition should satisfy Eqs. (A.1) and (A.2) µ(Fg + Fad ) ≥ Fc
(A.1)
µ(Fg + Fad ) ≥ Fi
(A.2)
where µ is a static friction coefficient. If adhesion energy is proportional to the contact area, JKR theory can be applied to predict the adhesion force. Also, additional contribution from a capillary force, which generally has the highest contribution to the adhesion force among the intermolecular forces such as van der Waals force and electrostatic force, was also considered in Eq. (A.3) Fad = 23 πγR + 2πRγL (cos θA + cos θB )
(A.3)
where R is the radius of curvature, θ A the water wetting angle of the transfer island, θ B the water wetting angle of the disk surface and γ the work of adhesion between the transfer island and disk surface. As previously described, two detaching forces such as centrifugal force and the asperity impact force were considered in this work. As seen in Fig. 13 the centrifugal force and the asperity impact force act perpendicular each other.
µ[mT g + 23 πγR + 2πRγL (cos θA + cos θB )] ≥ Fc
(A.4)
µ[mT g + 23 πγR + 2πRγL (cos θA + cos θB )] ≥ Fi
(A.5)
where mT is the mass of transfer island and g the gravitational acceleration. The centrifugal force can be estimated with the formula shown in Eq. (A.6) Fc = mT
(A.6)
where mT is the self-mass of transfer island, vA the sliding velocity of disk and r the radius of the wear track. The impact between the asperity of counter surface and the transfer island could be accommodated by the deformation of the asperity or the deformation of transfer island. As a result, a simplified model for the asperity impact force accommodation could be summarized in Fig. 13. By using the theory of momentum conservation, the asperity impact force could be expressed with Eq. (A.7) Fi t = (mA vA + mT vT ) (A.7) where t is the impact time, mA the mass of the impacting asperity, mT the self-mass of the transfer island, ∆vA the accommodated velocity of impacting asperity and ∆vT the accommodated velocity of transfer island. In this work, the impact deformation mainly occurs at the relatively soft transfer island, so ∆vA could be neglected. And the ∆vT could be the average deformation value (x/2) divided by impact time (t) as expressed in Eq. (A.8). Fi t = mT vT = mT
x 2t
(A.8)
In this case, ∆v could be replaced by the impacting velocity of the asperity (vA ) and Eq. (A.8) could be converted to Eq. (A.9). Fi = mT
v2A x mT vA 2 = (x) = m T 2 x 2x 2(t)2
(A.9)
According to Eqs. (A.6) and (A.9), the centrifugal force and the asperity impact force linearly increase with the square of the sliding velocity. Therefore, the tendency of transfer layer formation decreases with the sliding velocity. And it implied that there exists a critical velocity above which no transfer layer is able to form. When the centrifugal force is the dominant detaching force, transfer island is stable when Eq. (A.10) is satisfied.
Fig. 13. A graphical illustration showing the accommodation of asperity impact force.
(vA )2 r
1/2 µr 3 mT g+ πγR+2πRγL (cos θA +cos θB ) ≥vcr mT 2 (A.10)
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Or, when the asperity impact force is the dominant detaching force, transfer island is stable when Eq. (A.11) is satisfied. 1/2 3 µ(x) mT g+ πγR+2πRγL (cos θA +cos θB ) ≥vcr mT 2 (A.11) In order to evaluate the above reasoning, the critical velocity is calculated roughly using Eqs. (A.10) and (A.11). According to Israelachvili [19], when the surfaces do not have polarity, the work of adhesion (γ) could be calculated with Eq. (A.12) √ γ = γA + γB − 2 γA γB (A.12) References [1] M.R. Hilton, P.D. Fleischauer, Applications of solid lubricant films in space, Surf. Coat. Technol. 54–55 (1992) 435–441. [2] F.P. Bowden, D. Tabor, The Friction and Lubrication of Solids, Part II, Clarendon Press, Oxford (1964) 52–86. [3] A. Erdemir, D.E. Busch, R.A. Erck, G.R. Fenske, R. Lee, Ion-beam-assisted deposition of silver films on zirconia ceramics for improved tribological behavior, STLE 47 (10) (1990) 863–872. [4] D.A. Rigney, Transfer, mixing and associated chemical and mechanical processes during the sliding of ductile materials, Wear 245 (2000) 1–5. [5] S. Wilson, A.T. Alpas, Tribo-layer formation during sliding wear of TiN coating, Wear 245 (2000) 223–229. [6] M.S. Bednar, D. Kuhlmann-Wilsdorf, Amorphous and alloy film formation in sliding of silver on copper, Wear 181–183 (1995) 922– 932.
[7] X.Y. Li, K.N. Tandon, Microstructural characterization of mechanically mixed layer and wear debris in sliding wear of an Al alloy and an Al based composite, Wear 245 (2000) 148– 161. [8] F.H. Stott, The role of oxidation in the wear of alloys, Tribol. Int. 31 (1–3) (1998) 61–71. [9] P. Diss, M. Brendle, A general approach to discontinuous transfer films: influence of sliding speed and stick-slip phenomena, Wear 203–204 (1997) 564–572. [10] A. Kapoor, F.J. Franklin, Tribological layers and the wear of ductile materials, Wear 245 (2000) 204–215. [11] 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 (9) (1997) 663–676. [12] J. Zhang, A.T. Alpas, Transition between mild and severe wear in aluminium alloys, Acta Mater. 45 (2) (1997) 513–528. [13] M. Brendle, P. Turgis, S. Lamouri, A general approach to discontinuous transfer films: the respective role of mechanical and physico-chemical interactions, Tribol. Trans. 39 (1) (1996) 157– 165. [14] E.-S. Yoon, H. Kong, O.K. Kwon, J.-E. Oh, Evaluation of frictional characteristics for a pin-on-disk apparatus with different dynamic parameters, Wear 203–204 (1997) 341–349. [15] S.C. Lim, Recent developments in wear mechanism maps, Tribol. Int. 31 (1–3) (1998) 87–97. [16] H. Kong, M.F. Ashby, Wear mechanisms in brittle solids, Acta Metall. Mater. 40 (1992) 2907–2920. [17] S.H. Yang, H. Kong, E.-S. Yoon, The effect of silver particle addition on the tribological characteristics of rubbing AISI 52100 steel pairs, in: Proceedings of the Second World Tribology Congress, September 2001. [18] E. Rabinowicz, Friction and Wear of Materials, 2nd ed., Wiley, New York, 1995, pp. 285–287. [19] J.N. Israelachvili, Intermolecular and Surface Forces, 2nd ed., Academic Press, New York, 1992, pp. 139–151.
A wear map of bearing steel lubricated by silver films Seung Ho Yang a , Hosung Kong a,∗ , Eui-Sung Yoon a , Dae Eun Kim b a
Tribology Research Center, Korea Institute of Science and Technology, 39-1 Hwawolkok-dong Sungbuk-ku, Seoul, South Korea b Department of Mechanical Engineering, Yonsei University, Seoul, South Korea
Abstract Wear map of bearing steel lubricated by silver film has been constructed to delineate the wear transition behavior with the change in operating conditions. Experiments were performed in dry sliding conditions using two types of ball-on-disk type test rigs under the contact pressure of 100–1000 MPa and the sliding speed of 20–1000 mm/s in ambient air. For the silver coating, an ultra thin IBAD silver bond layer was firstly deposited on AISI 52100 steel surfaces and then a relatively soft silver film was deposited by a thermal evaporation method onto the first layer. This functionally gradient film showed a great improvement in the life, mainly owing to the better bond strength. In order to build up a general framework on the tribological behavior of the functionally gradient silver films, all test data were plotted on a map whose axes are contact pressure and sliding speed. As a result, three main regimes were clearly identified: (i) elastic/plastic deformation of silver coating without failure, (ii) mild wear regime after initial failure of silver coating and (iii) severe wear regime. In the mild wear regime, the contact surfaces were covered with transfer layers of agglomerated wear particles. The transfer layer acted as a protective layer and resulted in low friction after initial failure of the coating. The formation of transfer layer was suppressed by several destructive actions, when the sliding speed was high. And, above a critical sliding speed, no transfer layer was able to form. In the discussion, an empirical model that could explain the existence of critical sliding speed was proposed. © 2003 Elsevier Science B.V. All rights reserved. Keywords: Wear map; Transfer layer; Functionally gradient film; Critical velocity
1. Introduction Soft metallic coating has been focused as a good solid lubricating film for the space bearing element [1]. The solid lubricating characteristic of thin soft metallic films has been well explained by the Tabor’s adhesion theory of friction that described how the effective lubricating characteristics could be obtained between two materials in sliding contact [2]. Soft metals also have excellent heat conductivity since they diffuse out the frictional heats generated at the contacting interfaces [3]. However, soft metallic coatings usually have very little mutual solubility with iron-based steel alloys, so they generally show poor adhesion to the steel surfaces. This has been a critical restriction for applying the soft metallic coatings to the practical machine elements. As a result, it has been normally recognized that initial failure of coated layer is the definite end of service life. Recently, the formation of transfer layer has been one of the great concerns to many tribologists. Because it could alter the friction and wear characteristics very seriously even ∗ Corresponding author. Tel.: +82-2-958-5655; fax: +82-2-958-5659. E-mail address: [email protected] (H. Kong).
after the initial failure of coating occurred. Lots of work have focused on the generation mechanism of transfer layer [4–9] and the role of the transfer layer on friction and wear [10–13]. Yet all of those works dealt in a specific point of view under certain operating conditions. Therefore, the generation mechanism and the role of transfer layer need more elaborated works. In this work the tribological characteristics of silver transfer layers were focused. Functionally gradient silver film was tested to study the effect of the film bond strength on the tribological characteristics of the film and also the characteristics of the transfer layer were evaluated. Tests were performed in various ranges of load and speed and the results were summarized in a wear map constructed with the axes of contact pressure and sliding speed.
2. Experimental details Pure (99.99%) silver was coated on the hardened AISI 52100 steel disks with a functionally gradient silver film deposition method. Friction and wear characteristics of the coatings were evaluated with two ball-on-disk type tribo-testers. In order to cover a wide range of load
0043-1648/03/$ – see front matter © 2003 Elsevier Science B.V. All rights reserved. doi:10.1016/S0043-1648(03)00148-0
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conditions, both a macro-tribo-tester and a miniature tribo-tester were used to cover wide range of loads. All the tribo-tests were performed in unidirectional dry sliding under various ranges of load and speed in an ambient condition. After the tests, the size of wear scars of tested balls were measured with an optical microscope and the wear coefficient (K) was calculated as follows: HV K= (1) LD where H is the hardness of hardened AISI 52100 steel, V the worn volume of the ball, L the sliding distance and D the applied normal load. 2.1. Tribo-testers When the maximum Hertzian contact pressure is greater than 400 MPa (the applied normal load is 1.1 N), a macro-tribo-tester was used. Test loads were applied by using dead weight. Friction force was measured with a load cell (maximum load: 3 kgf) installed at a right angle to the direction of sliding. The apparatus was described in detail elsewhere [14]. When the applied load was less than 1.1 N, which is the minimum load of the macro-tribo-tester, tests were performed with a miniature tribo-tester shown in Fig. 1. Maximum load of the miniature tribo-tester was 100 gf. 2.2. Specimens Hardened AISI 52100 steel disks (hardness: HR C 57–61) of 90 mm in diameter and 6 mm in thickness were used for the tests performed with a macro-tribo-tester. For a miniature tribo-tester, hardened AISI 52100 steel disks (hardness: HR C 57–61) of 40 mm in diameter and 6 mm in thickness were used. As for the specimen balls, hardened AISI 52100 steel
Fig. 1. A close-up view of the (a) macro-tribo-tester and (b) miniature tribo-tester.
Fig. 2. Schematic figure showing the ion beam assisted coating system.
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balls (hardness: HR C 55–60) of 12.7 mm in diameter were used. 2.3. Coating methods In order to increase the bond strength of a thermally evaporated silver coating, functionally gradient coating was performed using accelerated argon ion bombardment. The coating system was illustrated schematically in Fig. 2. A hollow cathode type ion gun with the mouth diameter of 30 mm was installed in the thermal evaporation chamber. Specimen disks were ultrasonically cleaned with n-hexane for 2 min. Final cleaning was performed by the argon ion bombardment (1.5 keV) in the coating chamber under vacuum (1 × 10−2 Pa) before coating.
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Table 1 Coating conditions Thin seed layer formation Ion bombardment Vacuum (Pa) Ion beam acceleration (eV) Argon gas flow rate (sccm) Thermal evaporation Evaporant Heating current (A) Thermal evaporation Vacuum (Pa) Evaporant Heating current (A)
1 × 10−2 1500 2.0 Silver (0.05 g) 500 1 × 10−3 Silver (2.0 g) 500
Fig. 3. Wear coefficients and friction coefficients of functionally gradient silver coatings as a function of maximum Hertzian contact pressure.
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Functionally gradient silver coating was performed by using a thermal evaporation of silver assisted by accelerated argon ion bombardment. For functionally gradient coating, a very thin atomically mixed silver layer of 8 nm in thickness was firstly deposited onto the steel disks by (using) the ion beam assisted thermal evaporation method (of silver assisted by the bombardment of accelerated (1.5 keV) argon ions). It was intended that this ultra-thin layer be utilized as an atomically mixed seed layer for the subsequent thermally evaporated silver coating. For the second step, silver film was deposited by a thermal evaporation method onto the thin seed layer. Silver coating of 2 m in thickness was finally deposited on the steel disk specimen. Detailed coating conditions are listed in Table 1.
Table 2 Test conditions Test conditions
Roughness of substrate, Ra (nm) Roughness after coating, Ra (nm) Coating thickness (m) Speed (mm/s) Normal force (N) [maximum Hertzian contact pressure (MPa)] Lubrication Environments Temperature
Specimen Ball
Disk
10 10 (no coating) – 20, 100, 400, 1000 0.02 [100], 0.14 [200], 0.47 [300], 1.10 [400], 2.19 [500], 5.88 [700], 17.64 [1000] Dry Ambient air Room temperature
100 20 2.0
2.4. Test conditions Tests were performed in dry sliding conditions at the load range 0.02–17.64 N and the speed range 20–1000 mm/s for a fixed sliding distance of 1000 m. Detailed test conditions are listed in Table 2. Each test was repeated more than three times.
3. Results and discussion 3.1. Effect of contact pressure and sliding velocity Tests were performed under various normal loads at the constant sliding speed of 20 mm/s with the functionally gradient silver films. The wear coefficients and friction coef-
ficients were shown in Fig. 3. When the contact pressure was less than 400 MPa, no failure was found. As the coating film supported the applied load without the film failure, mild wear (K = 1.0 × 10−5 to 1.0 × 10−6 ) was observed. When the contact pressure was higher than 400 MPa, however, the silver coating failed and the wear coefficient rose abruptly. The frictional behaviors were different from the results of wear coefficients. The coating failure did not always accompany a high friction coefficient. In order to find the changes at the contact surfaces, EPMA line profiling analyses for the disks were performed after tests, as shown in Fig. 4. When the contact pressure was low (Pmax = 100 MPa), silver was not detected at the track surface. On the other hand, under the load of 400 and 700 MPa, transfer of silver was observed at the track surface after the initial coating failure. It was
Fig. 4. EPMA line profiling analysis results of the disk surfaces with functionally gradient silver coatings under the various conditions of contact pressure.
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Fig. 5. Wear coefficients and friction coefficients of functionally gradient silver-coated surfaces as a function of sliding speed.
found that such a transfer layer played an important role on the friction behavior afterward. Fig. 5 shows the wear and friction behaviors of functionally gradient silver films at the various sliding speeds of 20, 100, 400 and 1000 mm/s under the constant contact pressure of 400 MPa. When the sliding speed was less than 100 mm/s, relatively low wear coefficients (1.0 × 10−3 to 1.0 × 10−4 ) were obtained. However, at the higher sliding speeds (400 and 1000 mm/s), the wear coefficient increased by one order of magnitude higher (1.0 × 10−2 to 1.0 × 10−3 ). On the other hand, the friction coefficient increased with the sliding speed. Results of the EPMA line profile analyses (Fig. 6) showed that when the sliding speed was relatively low (less than 100 mm/s), silver transfer layer was noticeable at the track surfaces. The formation of such a transfer layer explained why the low wear coefficient was obtained even after the
initial coating failure. The transfer layer acted as a solid lubricant. When the sliding speed was 1000 mm/s, however, the transfer layer could not be found at the track surfaces but only scattered wear debris was found at the track surfaces, and the resultant wear coefficient and friction coefficient were high. The above results suggested that both the contact pressure and the sliding speed affected seriously the formation of transfer layer and the tribological behavior of silver coating films. 3.2. Construction of wear map The tests were performed in various ranges of load and speed and the results were summarized in a wear map constructed with the axes of contact pressure and sliding speed as shown in Fig. 7. Maximum Hertzian contact pressure and
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Fig. 6. EPMA line profiling analysis results of the disk surfaces with functionally gradient silver-coated surfaces under the various conditions of sliding speeds.
sliding speed were selected as the axes of the map because the P × V (pressure × speed) value corresponds to the mechanical input energy to the coating films [15,16]. All of the wear coefficients and the coefficients of friction were plotted numerically at the different combinations of contact pressure and sliding speed.
Fig. 7. The characteristic wear regimes of the functionally gradient silver-coated surfaces (italicized wear coefficients and coefficients of friction were given numerically).
Based on the relative magnitude of the given wear coefficients, three different characteristic regimes (Regime I, Regime II and Regime III) could be identified as shown in Fig. 7. 3.2.1. Regime I In this regime, silver film sustains the applied load without any failure of the coating. The coated silver film prevented the ball and track surfaces from direct contact, which resulted in mild wear during the sliding (Fig. 8). The onset of the film failure can occur by increasing the contact pressure further. It was found that the criteria for coating failure depend on whether the given mechanical input energy (P × V value) exceeds some critical value or not. 3.2.2. Regime II Regime II demarcated the region where relatively low wear coefficients (order of 10−4 to 10−5 ) were still sustained after the silver coating apparently failed. The low wear coefficients were resulted mainly from the transfer layers of agglomerated wear debris on the contact surfaces (Fig. 9). After the apparent failure of coated silver films, wear particles were re-grinded or plastically deformed to generate much more fine particles. These fine particles are supposed to have very high energy due to its large surface area, so they tend to easily stick to the adjacent surface in order to reduce the total free energy [17]. The transfer layers acted as a protective layer, they decrease the wear and friction to a low and stable value. Therefore, it was found that the formation of transfer layer
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Fig. 8. EPMA line profiling analysis results of the ball and disk surfaces tested in Regime I.
Fig. 9. EPMA line profiling analysis results of the ball and disk surfaces tested in Regime II.
provided beneficial effects to the contact surfaces even after the coating film failed. 3.2.3. Regime III Regime III was demarcated the region of relatively high wear coefficients (K > 10−3 ). As seen in Fig. 10, transfer layer could not be observed in this regime. Friction and wear were higher than those of Regimes I and II. In this regime, the film failure apparently meant the end of service life since there is no protective layer after the failure. 3.3. Wear map of functionally gradient silver coating Based on the above characteristic regimes (I, II and III), a wear map of the functionally gradient silver coating was
constructed as shown in Fig. 11. It was found that the border between Regimes II and III in Fig. 11 is mostly affected by the sliding speed. It also suggested that a critical sliding speed (vcr ) exists in the formation of transfer layer, above which no material transfer is able to form. 3.4. Critical sliding velocity Transfer layer formation means the stable attachment of wear particles on the sliding surfaces. Wear particles generally have very high surface area to volume ratios and high surface energy. So, the wear particles have the tendency to stick to adjacent surface. But not every wear particle can stick to the surface stably because disturbing forces exists. Therefore, the formation of transfer layer could be
Fig. 10. EPMA line profiling analysis results of the ball and disk surfaces tested in Regime III.
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The work of adhesion (γ) between a transfer island and the surface of bearing steel is approximately 70 erg/cm2 when the surface energy of silver (γ A ) is 920 erg/cm2 and the surface energy (γ B ) of bearing steel is 1500 erg/cm2 [18]. Let us consider a cylindrical silver transfer island having the radius (R) of 1 m and height of 1 m. The density of pure silver is 10 g/cm2 . The radius of wear track (r) was 15 mm and the measured water wetting angle of silver (θ A ) and bearing steel (θ B ) was 80◦ and 60◦ , respectively. When the centrifugal force is the dominant detaching force, the critical velocity (vcr ) is calculated about 1000 mm/s. According to Fig. 5, the critical velocity was approximately 100–1000 mm/s. Therefore, it could be concluded that above arguments are reasonable.
4. Conclusions
Fig. 11. A wear map of the functionally gradient silver coating (italicized wear coefficients and coefficients of friction were given numerically).
competitive depending on the two forces, such as detaching forces (centrifugal force and asperity impact force) and attaching forces (adhesion force and gravitational force of wear particles). The centrifugal force and the asperity impact force increase with approximately the square of the sliding velocity. So the tendency of transfer layer formation decreases with the sliding velocity. Therefore, it could be noted that there exists a critical velocity when the transfer layer formation is not able to form. When the centrifugal force is the dominant detaching force, transfer island is stable when the following Eq. (2) is satisfied. Details of the simple analyses used in these calculations are given in Appendix A. 1/2 µr 3 mT g+ πγR+2πRγL (cos θA +cos θB ) ≥vcr mT 2 (2) Or, when the asperity impact force is the dominant detaching force, transfer island is stable when Eq. (3) is satisfied. 1/2 µ(x) 3 mT g+ πγR+2πRγL (cos θA +cos θB ) ≥vcr mT 2 (3) where µ is the static friction coefficient of transfer island, mT the self-mass of transfer island, g the gravitational acceleration, γ the work of adhesion between transfer island and disk surface, R the radius of transfer island, γ L the surface tension of water, θ A the water wetting angle of the transfer island, θ B the water wetting angle of the disk surface, x the deformation by asperity impact, vcr is the critical velocity. In order to evaluate the above reasoning, the critical velocity was calculated roughly using Eqs. (2) and (3) as follows.
The tribological role of transfer layer of silver-coated surfaces was summarized as follows: (1) Silver transfer layers could protect the contact surface beneficially even after the initial coating failure occurs. (2) Wear maps of silver coating films were constructed using the axes of contact pressure and sliding speed, where three different characteristic wear regimes were clearly identified such as Regime I (no failure, mild wear), Regime II (dominant region for transfer layer, moderate wear) and Regime III (no transfer layer, severe wear). (3) The formation of transfer layer was strongly affected by the operating conditions, especially by the sliding speed. It was suppressed when the sliding speed was high and, hence, above a critical sliding speed, no transfer layer was able to form. (4) Centrifugal force and asperity impact force affected seriously the critical velocity.
Acknowledgements The authors would like to thank the Ministry of Science and Technology and Critical Technology 21 program (Machinery Design Technology Enhancement) and the National Research Laboratory Program for their support and interest in this work.
Appendix A In order to explain how the formation of the transfer layer depends on the sliding speed, various forces that may influence the transfer layer formation were discussed. A free body diagram of one transferred mass on a flat disk surface was constructed in Fig. 12. The wear particles are deformed by concentrated contact pressure, so they generally reveal plate-like shapes. In this work, one cylindrical
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So they could be dealt separately. The transfer island is stable when Eqs. (A.4) and (A.5) are satisfied
Fig. 12. A free body diagram of a transfer island formation.
transfer island was simply considered where major forces were shown in the figure. The formation of transfer layer could depend on two competitive forces; detaching forces such as centrifugal force (Fc ), asperity impact force (Fi ) and vertical component of the centrifugal force (Fcz ), and attaching forces such as adhesion force (Fad ) and gravitational force of wear particles (Fg ). In this work, the run-out in the vertical direction was less than 1 m so the vertical component of the centrifugal force (Fcz ) was neglected. When the transfer island is stably stuck to the disk surface, the force equilibrium condition should satisfy Eqs. (A.1) and (A.2) µ(Fg + Fad ) ≥ Fc
(A.1)
µ(Fg + Fad ) ≥ Fi
(A.2)
where µ is a static friction coefficient. If adhesion energy is proportional to the contact area, JKR theory can be applied to predict the adhesion force. Also, additional contribution from a capillary force, which generally has the highest contribution to the adhesion force among the intermolecular forces such as van der Waals force and electrostatic force, was also considered in Eq. (A.3) Fad = 23 πγR + 2πRγL (cos θA + cos θB )
(A.3)
where R is the radius of curvature, θ A the water wetting angle of the transfer island, θ B the water wetting angle of the disk surface and γ the work of adhesion between the transfer island and disk surface. As previously described, two detaching forces such as centrifugal force and the asperity impact force were considered in this work. As seen in Fig. 13 the centrifugal force and the asperity impact force act perpendicular each other.
µ[mT g + 23 πγR + 2πRγL (cos θA + cos θB )] ≥ Fc
(A.4)
µ[mT g + 23 πγR + 2πRγL (cos θA + cos θB )] ≥ Fi
(A.5)
where mT is the mass of transfer island and g the gravitational acceleration. The centrifugal force can be estimated with the formula shown in Eq. (A.6) Fc = mT
(A.6)
where mT is the self-mass of transfer island, vA the sliding velocity of disk and r the radius of the wear track. The impact between the asperity of counter surface and the transfer island could be accommodated by the deformation of the asperity or the deformation of transfer island. As a result, a simplified model for the asperity impact force accommodation could be summarized in Fig. 13. By using the theory of momentum conservation, the asperity impact force could be expressed with Eq. (A.7) Fi t = (mA vA + mT vT ) (A.7) where t is the impact time, mA the mass of the impacting asperity, mT the self-mass of the transfer island, ∆vA the accommodated velocity of impacting asperity and ∆vT the accommodated velocity of transfer island. In this work, the impact deformation mainly occurs at the relatively soft transfer island, so ∆vA could be neglected. And the ∆vT could be the average deformation value (x/2) divided by impact time (t) as expressed in Eq. (A.8). Fi t = mT vT = mT
x 2t
(A.8)
In this case, ∆v could be replaced by the impacting velocity of the asperity (vA ) and Eq. (A.8) could be converted to Eq. (A.9). Fi = mT
v2A x mT vA 2 = (x) = m T 2 x 2x 2(t)2
(A.9)
According to Eqs. (A.6) and (A.9), the centrifugal force and the asperity impact force linearly increase with the square of the sliding velocity. Therefore, the tendency of transfer layer formation decreases with the sliding velocity. And it implied that there exists a critical velocity above which no transfer layer is able to form. When the centrifugal force is the dominant detaching force, transfer island is stable when Eq. (A.10) is satisfied.
Fig. 13. A graphical illustration showing the accommodation of asperity impact force.
(vA )2 r
1/2 µr 3 mT g+ πγR+2πRγL (cos θA +cos θB ) ≥vcr mT 2 (A.10)
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Or, when the asperity impact force is the dominant detaching force, transfer island is stable when Eq. (A.11) is satisfied. 1/2 3 µ(x) mT g+ πγR+2πRγL (cos θA +cos θB ) ≥vcr mT 2 (A.11) In order to evaluate the above reasoning, the critical velocity is calculated roughly using Eqs. (A.10) and (A.11). According to Israelachvili [19], when the surfaces do not have polarity, the work of adhesion (γ) could be calculated with Eq. (A.12) √ γ = γA + γB − 2 γA γB (A.12) References [1] M.R. Hilton, P.D. Fleischauer, Applications of solid lubricant films in space, Surf. Coat. Technol. 54–55 (1992) 435–441. [2] F.P. Bowden, D. Tabor, The Friction and Lubrication of Solids, Part II, Clarendon Press, Oxford (1964) 52–86. [3] A. Erdemir, D.E. Busch, R.A. Erck, G.R. Fenske, R. Lee, Ion-beam-assisted deposition of silver films on zirconia ceramics for improved tribological behavior, STLE 47 (10) (1990) 863–872. [4] D.A. Rigney, Transfer, mixing and associated chemical and mechanical processes during the sliding of ductile materials, Wear 245 (2000) 1–5. [5] S. Wilson, A.T. Alpas, Tribo-layer formation during sliding wear of TiN coating, Wear 245 (2000) 223–229. [6] M.S. Bednar, D. Kuhlmann-Wilsdorf, Amorphous and alloy film formation in sliding of silver on copper, Wear 181–183 (1995) 922– 932.
[7] X.Y. Li, K.N. Tandon, Microstructural characterization of mechanically mixed layer and wear debris in sliding wear of an Al alloy and an Al based composite, Wear 245 (2000) 148– 161. [8] F.H. Stott, The role of oxidation in the wear of alloys, Tribol. Int. 31 (1–3) (1998) 61–71. [9] P. Diss, M. Brendle, A general approach to discontinuous transfer films: influence of sliding speed and stick-slip phenomena, Wear 203–204 (1997) 564–572. [10] A. Kapoor, F.J. Franklin, Tribological layers and the wear of ductile materials, Wear 245 (2000) 204–215. [11] 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 (9) (1997) 663–676. [12] J. Zhang, A.T. Alpas, Transition between mild and severe wear in aluminium alloys, Acta Mater. 45 (2) (1997) 513–528. [13] M. Brendle, P. Turgis, S. Lamouri, A general approach to discontinuous transfer films: the respective role of mechanical and physico-chemical interactions, Tribol. Trans. 39 (1) (1996) 157– 165. [14] E.-S. Yoon, H. Kong, O.K. Kwon, J.-E. Oh, Evaluation of frictional characteristics for a pin-on-disk apparatus with different dynamic parameters, Wear 203–204 (1997) 341–349. [15] S.C. Lim, Recent developments in wear mechanism maps, Tribol. Int. 31 (1–3) (1998) 87–97. [16] H. Kong, M.F. Ashby, Wear mechanisms in brittle solids, Acta Metall. Mater. 40 (1992) 2907–2920. [17] S.H. Yang, H. Kong, E.-S. Yoon, The effect of silver particle addition on the tribological characteristics of rubbing AISI 52100 steel pairs, in: Proceedings of the Second World Tribology Congress, September 2001. [18] E. Rabinowicz, Friction and Wear of Materials, 2nd ed., Wiley, New York, 1995, pp. 285–287. [19] J.N. Israelachvili, Intermolecular and Surface Forces, 2nd ed., Academic Press, New York, 1992, pp. 139–151.