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a-C:H tribological coating on micropitting wear of bearing steel
Wear 350-351 (2016) 107–115
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Influence of a WC/a-C:H tribological coating on micropitting wear of bearing steel B. Mahmoudi a, G.L. Doll a,n, C.H. Hager Jr.b, R.D. Evans b a b
Timken Engineered Surfaces Laboratories, The University of Akron, Akron, OH 44325, USA The Timken Company, Canton, OH 44720, USA
art ic l e i nf o
a b s t r a c t
Article history: Received 23 September 2015 Received in revised form 13 January 2016 Accepted 17 January 2016 Available online 29 January 2016
The focus of this work was to evaluate the use of a WC/a-C:H coating, specifically designed to withstand rolling contact fatigue, to mitigate micropitting damage on through-hardened AISI 52100 bearing steel. This was accomplished by conducting experiments with various combinations of coated discs and rollers using a three-disc-on-roller rolling contact fatigue machine. In addition to testing coated and uncoated contacts, discs and rollers with three hardness levels and different surface roughness values were also evaluated. Upon the completion of this work, it was observed that the onset of micropitting damage was affected by the hardness of the asperities on the mated roller and disc samples. It was also observed that the coated surface can cause abrasive wear damage on an uncoated surface if the coating is applied to a surface that has a roughness Rq Z 0.5 μm. If the coating is applied to a smoother roller surface (Rq r0.26 μm), the coating is able to mitigate micropitting damage on the roller surface without causing high abrasive wear rates on the disc surfaces. & 2016 Elsevier B.V. All rights reserved.
Keywords: Micropitting Diamondlike carbon Abrasive wear Bearing steel
1. Introduction Micropitting is a form of surface damage in rolling and mixed mode contacts consisting of numerous pits and associated cracks on a scale smaller than that of the Hertz elastic contact [1]. Once initiated, micropitting damage can alter the contact stress profile of the mated components, which may eventually cause stress concentrations so large that they can exceed the yield strength of the material. As a specific example, micropitting has been identified as a root cause of raceway spalling of main shaft spherical roller bearings in wind turbines. This spalling often occurs at fractions of the designed bearing life [2]. Several studies have been reported on micropitting of steel contacts, and all have correlated the onset of micropitting wear to the λ ratio (the ratio of the lubricant film thickness to the composite surface roughness) [3–7]. Decreases in λ and increases in slide-to-roll ratio up to 2% can elevate the risk of micropitting and reduce the number of stress cycles required to initiate micropitting [8,9]. Slide-to-roll ratio (SRR) describes the ratio of the sliding velocity to the entrainment velocity. If two surfaces are moving at the same surface speed, the motion is pure rolling (0.0% sliding). The n
Corresponding author. E-mail address: [email protected] (G.L. Doll).
http://dx.doi.org/10.1016/j.wear.2016.01.010 0043-1648/& 2016 Elsevier B.V. All rights reserved.
entrainment velocity is the mean velocity of the contacting surfaces, as described in Eq. (1). Positive and negative sliding occurs when the discs are rotating with faster and slower surface speeds than the roller, respectively. SRR is calculated based from the following equations: ue ¼
ðud þ ur Þ 2
SRR ¼
ðud ur Þ ue
ð1Þ ð2Þ
where ud , ur are the surface speeds of a disc and roller, and ue is the entrainment velocity. Based upon the results of twin disc experiments, Oila and Bull have proposed a mechanism for micropitting [10,11]. Their proposed mechanism suggests that plastic deformation associated with contacting asperities extends beneath the asperities and causes an increase in the dislocation density, which causes work hardening. Simultaneously, frictional heating is sufficient to activate the diffusion of carbon. At the boundary of the plastically deformed region (PDR), recrystallization occurs. Dislocations pile up against the PDR boundary and a slip band forms. With continuing cycles, dislocations accumulate in the slip band and a crack initiates. The crack propagates along the PDR boundary until it reaches the surface, where it generates a micropit.
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Considering the Oila and Bull suggested mechanism for micropitting and its experimental correlation with the λ ratio, it is hypothesized that a tribological coating may be able to mitigate micropitting damage if it can increase the λ ratio and reduce the frictional heating between asperities. Candidate coatings for this task need to be able to withstand a high number of stress cycles without fracturing, be strongly adherent to the steel substrates, and have deposition temperatures that are compatible with the heat treatment practices of the steel alloy. One coating that meets all of these requirements is a WC/a-C:H coating that consists of nanometer-sized β-W2C precipitates in an amorphous hydrocarbon matrix [12]. Based on nanoindentation with a Berkovich tip using the continuous stiffness measurement method, the coating is moderately harder ( 13 GPa) than hardened bearing steel, and functionally graded to ensure strong adhesion to steel [13]. Although the nanoindentation hardness measurement of 52100 from Ref. [13] was around 10.6 GPa, most indentation measurements show that 61 71 HRC equates to about 7 GPa. The objective of this study was to evaluate the ability of a WC/ a-C:H coating, which satisfies all the above requirements, to mitigate micropitting damage on mated through-hardened 52100 bearing steel surfaces. The performance of the WC/a-C:H coating was evaluated in boundary lubrication at various λ ratios, slide-toroll (SRR) ratios and steel hardness values.
2. Experimental details Testing was conducted using a PCS Instruments Micropitting Rig (MPR) with three through-hardened AISI 52100 1 cm wide discs (disc radius Rx ¼27 mm and crown Ry ¼1) and one throughhardened AISI 52100 steel roller (Rx ¼6 mm, Ry ¼12 mm). Images of the MPR test chamber and the roller geometry are shown in Fig. 1a and b, respectively. The description of this test rig can be found in [14,15]. The geometries of the discs and rollers create elliptical contact where the contact width between the roller and disc is between 500 and 700 μm depending upon load, wear mode and number of cycles. Discs were heat-treated to a hardness of 62 HRC, while the roller hardness values were 53, 57 and 62 HRC. It is acknowledged that the 53 and 57 HRC roller hardness values are lower than those typically specified for rolling element bearing steels. However, using this wide range of roller hardness in the MPR tests allowed for a clearer observation of the influence of hardness on this damage mode and confirmation of the efficacy of bearing steel hardness specifications [7]. The influence of surface roughness on wear was studied by using discs and rollers with two root mean square (RMS) surface roughness values, referred to as either smooth (Rq ¼0.26 μm) or rough (Rq ¼0.56 μm). The variation of surface roughness values is 70.08. Surface geometries, topographies, and roughness were characterized using a Zygo NewView 7300 3D white light interferometry
microscope. The complete test matrix is shown in Table 1. The tests were designed to probe how micropitting and wear depend upon various combinations of material pairs. Specifically, Series 1 (Tests 1.1– 1.9) explores how the hardness difference between the rollers and discs affects micropitting and wear, Series 2 (Tests 2.1–2.2) evaluates micropitting and wear of rollers when paired against coated discs, and Series 3 (Tests 3.1–3.7) probes micropitting and wear of rollers coated with WC/a-C:H. The entrainment velocity (ue) and oil temperature were chosen to achieve specific central lubricant film thicknesses calculated according to Eq. (3) [16]. h
min λ ¼ qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi
R2ar þ R2ad
0:49 0:073 hmin ¼ 1:79 R0 U 0:68 Σ GΣ W Σ U η0 P UΣ ¼ GΣ ¼ 2γ E0 ; W Σ ¼ 0 0 ; 0 2 2E R 2E R
ð3Þ
where λ is the film thickness ratio, hmin is film thickness, Rar is average rms roughness for roller, Rad is average rms roughness for disc, U is the sliding velocity, η0 is the dynamic viscosity, E0 is the equivalent modulus of elasticity, R0 is the reduced radius of curvature, γ is the pressure–viscosity constant, and P is the applied force. The test oil for all tests was an ISO viscosity grade 10 polyalphaolefin (PAO ISO 10) containing no extreme pressure (EP) or anti-wear (AW) additives except that used in Test 1.9 where the lubricant was PAO ISO 68 with no EP and AW additives. The viscosity of the PAO ISO 10 was measured to be 8.72 and 2.06 cP at 38 and 100 °C, respectively, which yielded pressure–viscosity coefficients of 13.5 GPa 1 and 10.15 GPa 1 at 38 and 100 °C. The WC/a-C:H coating was deposited onto roller and disc specimens in a closed-field unbalanced magnetron sputtering system. Strondl et al. have provided an overview of this deposition technology [17]. Roller bearings with this particular coating are able to achieve low λ (i.e., λ o1) fatigue lives more than three times greater than uncoated steel bearings, can tolerate a significant amount of debris damage without a reduction in calculated L10 life, and are very resistant to adhesive wear damage [18,19]. The coating is a layered material consisting of a 100 nmthick Cr layer that forms strong chemical bonds to the FeO surface layer on steel [20]. In the next layer, the composition is gradually changed from Cr to WC/a-C:H. The thickness of this gradient layer is also about 100 nm thick. Finally, the top layer of WC/a-C:H is approximately 900 nm thick. This yields a total coating thickness of approximately 1.1 μm. Testing on the MPR was interrupted periodically in order to perform a visual inspection of the wear track. If no evidence of micropitting or serious wear was observed, the test was restarted. This procedure was carried out until wear was observed or the test was suspended at 8 million cycles. In one case, the test was terminated at 30 million cycles with no significant wear or micropitting in order to measure the durability of the coating.
Fig. 1. (a) MPR chamber, where three larger discs (Rx: 27 mm) are running against one smaller roller in the middle. (b) Details of roller's geometry (Rx: 6 mm and Ry: 12 mm), which leads to elliptical Hertzian contact with cylindrical discs.
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Table 1 Micropitting test matrix. pmax represents the maximum calculated contact stress, ue is the entrainment velocity, SRR is the slide-to-roll ratio, and λ is the ratio of the calculated central lubricant film thickness to the composite surface roughness (Rq). Test
Coating
Rq (μm)
HRC
Discs
Roller
Discs
Roller
Discs
Roller
1.1 1.2 1.3 1.4 1.5 1.6 1.7
N N N N N N N
N N N N N N N
62 62 62 62 62 62 62
62 57 53 57 57 57 57
0.56 0.56 0.26 0.56 0.56 0.56 0.56
0.26 0.26 0.26 0.26 0.26 0.26 0.29
1.8 1.9 2.1 2.2 3.1 3.2 3.3 3.4 3.5
N N Y Y N N N N N
N N N N Y Y Y Y Y
62 62 62 62 62 62 62 62 62
62 62 57 57 62 53 57 57 57
0.56 0.56 0.56 0.26 0.26 0.26 0.56 0.56 0.56
0.26 0.56 0.26 0.26 0.56 0.26 0.26 0.26 0.26
3.6
N
Y
62
57
0.56
0.29
3.7
N
Y
58
62
0.35
0.27
P (GPa)
ue (m/s)
SRR (%)
T (°C)
λ
N106
μ
Comments
1.7 1.5 1.5 1.7 2 2 2 2.25 2.25 2.25 1.5 1.5 1.5 1.5 1.7 2 2 2.25 2.55 2 2.25 2.42 1.7 2 2.25
1 1 2 2 2 2 2
2 2 2 2 10 0 10
40 40 75 75 40 40 40
10 30 2 8 0.5 2 2/1
0.07 0.05 0.06 0.06 0.09 0.05 0.13
Mild wear Mild wear Wear Mild wear Wear þ cracking Micropitting þ cracking Micropitting
2 2 1 1 1 2 2 2 2
0 0 2 0 0 2 2 10 0
40 40 40 40 40 75 75 40 40
Micropitting Micropitting Wear Micropitting Wear No wear or micropitting No wear or micropitting No wear or micropitting; partial wear of coating No wear or micropitting
10
40
0.12
Wear of coating
2
2
75
1 1 2 1.5 7 10 8.5 1 15 7 6 2 6 6 15 15 15
0.06 0.03 0.07 0.05 0.08 0.05 0.05 0.09 0.04
2
0.09 0.09 0.12 0.07 0.14 0.14 0.13 0.13 0.13 0.4 0.09 0.15 0.09 0.12 0.07 0.14 0.16 0.16 0.16 0.14 0.13 0.13 0.10 0.10 0.09
0.06
No wear or micropitting
Fig. 2. Optical microscopy images of the wear tracks on uncoated rollers produced during Tests 1.1 to 1.9. (a) Roller surface from Test 1.1 after 10 million cycles. (b) Roller surface from Test 1.2 after 30 million cycles. (c) Roller surface from Test 1.3 after 1 million cycles. (d) Roller surface from Test 1.4 after 8 million cycles. (e) Roller surface from Test 1.5 after 0.5 million cycles. (f) Roller surface from Test 1.6 after 0.5 million cycles. (g) Roller surface from Test 1.7 after 1 million cycles. (h) Roller surface from Test 1.8 after 0.5 million cycles. (i) Roller surface from Test 1.9 after 0.5 million cycles. All photos have same scale. The arrows indicate the direction of rolling (R) and sliding (S). Positive sliding means when rolling and sliding are pointing to a same direction.
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3. Results and discussion 3.1. Series 1 Fig. 2a–i display optical micrographs taken of the surfaces of uncoated rollers that have run against uncoated discs in test Series 1, and results of these tests are displayed in Table 1. Surfaces of the mated discs from this Series are not shown since no measurable changes occurred during testing. In this study, wear refers to a change in the roughness and profile of counterparts and mild wear refers to change only in surface roughness but not in the profile of the mated contacts. The arrows indicate the direction of rolling (R) and sliding (S). Roller surface damage is attributed to wear or mild wear in (a), (b), (d) and (e) corresponding to Tests 1.1, 1.2, 1.4, and 1.5, respectively. Tests 1.1, 1.2, and 1.4 were performed at SRR r2% and λ o0.1. Test 1.5 was performed at SRR ¼ þ10% and λ ¼0.4. Roller surface damage is attributed to micropitting in (f), (g), (h), and (i), corresponding to Tests 1.6, 1.7, 1.8, and 1.9, respectively. These tests were all performed at values of λ 4 0.1. The surface damage on the roller shown in (c) from Test 1.3 appears to be severe wear, plastic deformation, and tearing. In Series 1, wear or mild wear of the roller occurred when λ was less than 0.1, and micropitting occurred when λ 40.1 for SRR ¼2% or less. An exception to this trend is Test 1.3, where the roller was 9 points softer than the discs on the HRC scale. Test 1.5 was performed with SRR ¼ þ10%, and although the initial λ value was 0.4, abrasive wear occurred. The competition between abrasive wear and micropitting (or surface fatigue) was noted by Morales-Espejel et al. [9], who performed MPR experiments similar to those in Series 1. Micropitting is a surface fatigue phenomenon that requires a number of stress cycles to initiate and propagate cracks. If the rate of material removal by wear is large enough, near surface material does not maintain residence long enough in the contact to accrue the requisite number of stress cycles for crack initiation. Therefore under these test conditions, the transition from a wear mode to micropitting appears to occur near λ ¼ 0.1. If the roller hardness is much less than the counter face (as is the case in Test 1.3), the lambda threshold should be much larger. Additionally, with increasing SRR, the lambda threshold between wear and micropitting also increases. For example, Fig. 3a and b show roller profiles before and after testing from Tests 1.5 and 1.6, respectively. The roller from Test 1.5 has experienced a significant loss of material while the roller from Test 1.6 has maintained its original profile, but with micropitting. In MPR tests performed under boundary lubrication, asperities on the surface of the roller plastically deform under load, which according to the mechanism put forth by Oila and Bull can create work hardening and eventually lead to micropitting [10]. Within the framework of this proposed mechanism, surface fatigue and micropitting initiation should depend upon the yield strength of the materials in contact and the contact stress. Comparing the results of Tests 1.2, 1.4, 1.6, and 1.8, the contact stress that produced micropitting was relatively close to the yield strength of the
roller. Test 1.2 and 1.4 rollers were 57 HRC (or HV 6.6 GPa), and did not exhibit micropitting damage even after as many as 30 million stress cycles. According to Tabor [21], the yield strength (σy) of a material is approximately one-third the Vickers hardness (H). Using this approximation, the yield strength of materials with H¼ 6.6 GPa is σy ¼2.2 GPa. Therefore, if the magnitude of the stresses on the asperities is similar to the calculated maximum Hertzian stress in the contact, tests performed with pmax o2.2 GPa may be insufficient to work harden the asperities and create micropitting. When pmax approached the value of 2.2 GPa in Tests 1.5 and 1.6, the roller surfaces exhibited a combination of wear and cracking.
3.2. Series 2 Fig. 4a and b show the surfaces of uncoated rollers after 1.5 million cycles that were tested against WC/a-C:H coated (Rq ¼0.56 μm) discs (Test 2.1) and WC/a-C:H coated (Rq ¼0.26 μm) discs (Test 2.2), respectively. The WC/a-C:H coating on the rough discs has not only removed the grind lines in the wear track on the roller, but has also removed a considerable amount of the profile of the roller, which can be observed in the profile trace shown in Fig. 4c. The original roller profile radius (Ry ¼12 mm) is no longer present, and the disc-roller contact is much wider. As a result, the Hertzian contact stress of the test was significantly reduced from its initial value of 1.5 GPa during the testing. In comparison, the uncoated roller that was tested against the Rq ¼0.26 μm coated discs in Test 2.2 exhibited micropitting after just 1 million stress cycles, as shown in Fig. 4b and d. Again, the transition between wear and micropitting for the two tests performed in this Series appears to be near an initial λ value of 0.1. The reduction in the magnitude of the grinding lines in the wear track of Fig. 4b indicates that the coated disc was also polishing the uncoated roller. Although polishing of the counter face increases the λ value slightly (from 0.09 to 0.11), in this case, the increase was not enough to avoid micropitting. These results indicate that adding the coating on the disc surface will not prevent micropitting damage on the roller under these test conditions. Almost all of the coating was removed from the discs in Test 2.1, where the WC/a-C:H was applied to a rough roller surface (Rq ¼0.56 μm). Subsequent to this test, submicron-sized particles of the coating were found suspended in the tested oil. Conversely, no coating wear was measured after 1.5 million cycles in Test 2.2. In a previous study [22], wear of the WC/a-C:H coating under this test condition was determined to be associated with fracture that initiated at high aspect ratio features at the substrate/coating interface. These results indicate that the Rq ¼ 0.56 μm surface is too rough for the WC/a-C:H coating to function effectively in rolling contact.
Fig. 3. Profiles of rollers showing the higher rate of wear as slide-to-roll ratio increases from 0.0% to 10.0%. (a) Comparison of the profile of roller from Test 1.5 before and after running 0.5 million cycles. (b) Comparison of roller profile before and after running 0.5 million cycles from Test 1.6.
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3.3. Series 3 Although adding the coating to the discs did not mitigate micropitting damage on the rollers, adding the coating to the rollers did. Fig. 5 shows the surfaces of coated rollers that ran against uncoated discs. The arrows indicate the direction of rolling (R) and sliding (S). Fig. 5b shows the undamaged roller surface after Test 3.2. The coated rollers were able to mitigate wear and micropitting damage in every test except Test 3.6 (shown
111
in Fig. 5f). When the slide-to-roll ratio was 10% in boundary lubrication and the contact stress was greater than 2 GPa, the coated surface exhibited surface damage. Table 1 shows that the friction measured during Test 3.6 (SRR ¼ 10%) was about 30% greater than the friction measured during Test 3.4 (SRR ¼ þ10%). The initiation of micropitting and coating delamination on the roller may be due to a synergistic effect of normal stress from Hertzian contact and shear stress due to friction. For a positive SRR (ud 4 ur), the shear force is in the rolling direction. However, in
Fig. 4. (a) Optical microscopy images of the wear tracks of uncoated rollers after 1.5 million cycles tested against WC/a-C:H-coated rough discs (Rq ¼ 0.56 μm) in Test 2.1. (b) Wear track of uncoated roller tested against WC/a-C:H-coated smooth discs (Rq ¼ 0.26 μm) in Test 2.2. (c) Comparison of roller profiles (Ry) from Test 1.2 (Rq ¼ 0.56 μm uncoated disc against Rq ¼ 0.26 μm uncoated roller) and Test 2.1 (Rq ¼ 0.56 μm coated disc against Rq ¼ 0.26 μm uncoated roller). (d) Comparison of roller profiles (Ry) from Test 1.1 (Rq ¼0.56 μm uncoated disc against Rq ¼ 0.26 μm uncoated roller) and Test 2.2 (Rq ¼ 0.26 μm coated disc against Rq ¼0.26 μm uncoated roller). uD/uR-R means uncoated discs/uncoated roller-rough surfaces. cD/uR-S means coated discs/uncoated roller-smooth surfaces.
Fig. 5. Optical microscopy images of the wear tracks on coated rollers after Tests 3.1 to 3.7. (a) Roller surface from Test 3.1 from Table 1 after 7 million cycles. (b) Roller surface from Test 3.2 after 10 million cycles. (c) Roller surface from Test 3.3 after 8.5 million cycles. (d) Roller surface from Test 3.4 after 1 million cycles. (e) Roller surface from Test 3.5 after 15 million cycles. (f) Roller surface from Test 3.6 after 8 million cycles. (g) Roller surface from Test 3.7 after 45 million cycles. All photos have same scale. The arrows indicate the direction of rolling (R) and sliding (S). Positive sliding means when rolling and sliding are pointing to a same direction.
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Fig. 6. (a) Comparison of the Ry profiles of an uncoated disc before and after the testing against rough coated roller (Test 3.1). (b) Wear scar on disc from Test 3.1 (uncoated Rq ¼ 0.26 μm discs versus coated Rq ¼0.56 μm rollers) after 7 million cycles of roller contact (each 13.5 contacts of roller is equal to one contact of a disc). uD/cR-R means uncoated discs/coated roller-rough surfaces.
negative SRR where ud our, the shear stress is opposite to the rolling direction. It is likely that the direction of the shear stress with respect to the rolling direction affects the damage progression of the coating. Errichello has pointed out that the dedendum of gear teeth (which experiences negative sliding) is more prone to micropitting than the mating gear surface that undergoes positive sliding (the tooth addendum) [23]. Since crack initiation is generally associated with tensile stresses, the shear stress due to positive sliding should generate tensile stress in the coated roller at the inlet position of the contact where the lubricant film thickness is equal to or greater than the central lubricant film thickness. On the other hand, negative SRR generates a maximum tensile stress at the outlet position on the coated roller where the lubricant film thickness is at a minimum. In this study, the lubricant film thickness at the outlet position of the contact is estimated to be about 30% thinner than the central film thickness. Fig. 6a and b show wear scars that were produced on the discs mated with the coated roller in Test 3.1 after 7 million roller stress cycles. In this test, the coated roller was much rougher than the mated discs. This led to a greater amount of wear on the disc surfaces than tests conducted with smoother rollers. Fig. 6a compares the Ry profile for the wear scars of the discs before and after the test. As with Test 2.1 (Fig. 4c), the coating on the rough roller surface in Test 3.1 caused a high rate of wear on the uncoated surface of the disc. 3.4. Wear of uncoated counter face When two surfaces are subjected to high Hertzian contact pressure, the asperities of the mating surfaces elasto-plastically deform to support the applied load. The plastic deformation of the WC/a-C:H asperities should be considerably less than that of the uncoated part, since the coating is harder but more elastic than through-hardened 52100 steel. Therefore, if the roughness of the coated component is relatively large, the coating asperities will continue to remove material from the uncoated counter face until either the contact pressure minimizes due to a change in the profile of the uncoated counter face, or the lubricant film thickness completely separates the asperities of the mating surfaces. This effect is displayed graphically in Fig. 7. Here, the contact stress has been calculated from periodic Ry measurements of the roller surfaces during testing involving Rq ¼0.56 μm discs with and without the WC/a-C:H coating (Tests 1.2 and 2.1). The contact stress between the coated discs and roller in Test 2.1 decreased during testing from 1.5 to nearly 1 GPa. The reduction of the contact stress through wear involving the uncoated contacts in Test 1.2 was much less (a much smaller change in the roller profile). Additionally, the Figure shows that most of the wear on the uncoated rollers occurred during the early stages of Test 2.1.
Fig. 7. Calculated maximum Hertzian contact stress plotted versus cycles for Tests 1.2 (uncoated Rq ¼ 0.5 μm discs versus uncoated Rq ¼0.26 μm roller) and 2.1 (coated Rq ¼0.5 μm discs versus uncoated Rq ¼ 0.26 μm roller). pmax is calculated from Ry measurements of the roller surfaces taken at different cycles during the tests. uD/ uR-R means uncoated discs/uncoated roller-rough surfaces.
The wear volume of the uncoated counter face was significantly reduced when the coating was applied to smoother surfaces. Since the wear rate of the coated part was very low, the total wear volume of the system was associated primarily with the uncoated steel counter faces. The change in the separation between the centers of the discs and the rollers (i.e., the displacement) measured during Tests 1.3 (uncoated Rq ¼ 0.26 μm discs vs. uncoated Rq ¼0.26 μm 53 HRC rollers) and 3.2 (uncoated Rq ¼ 0.26 μm discs vs. coated Rq ¼0.26 μm 53 HRC rollers) is plotted versus stress cycles in Fig. 8a after the first 2 million test cycles. Fig. 8b shows similar results for Tests 1.4 and 3.3. The displacement curves approximate the wear depth and indicate that about 33% less material was removed from the uncoated steel specimens when they were in contact with the coating instead of the uncoated steel. Moreover, Test 3.2 (uncoated Rq ¼0.56 μm discs vs. coated Rq ¼0.26 μm 53 HRC rollers) achieved more than five times more cycles without micropitting or wear than Test 1.3 (uncoated Rq ¼0.56 μm discs vs. uncoated Rq ¼0.26 μm 53 HRC rollers), which exhibited severe wear before 2 million cycles. Fig. 9a–f illustrate the surface morphologies of the discs that were run against an uncoated roller (Test 1.4) and a WC/a-C:Hcoated roller (Test 3.3) after about 2 million cycles. The grind lines are still visible on the surfaces of a disc from Test 1.4 (Fig. 9a and b) and a disc from Test 3.3 (Fig. 9c and d), although the coating on the roller in Test 3.3 appears to have visually altered the appearance of the grind lines on the disc. Ra, Rq and Rsk values obtained from the 3D surface measurements in Fig. 9 are presented in Table 2. A significant change in the Rsk values has occurred as a result of the testing. Rsk values became negative on both disc surfaces, and a disc that ran against the coated roller attained an Rsk value approaching 1. Negative Rsk values have been reported to have a
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Fig. 8. The change in the separation between the centers of the discs and the rollers (i.e., the displacement) measured during (a) Tests 1.3 (uncoated Rq ¼ 0.56 μm discs vs. uncoated Rq ¼0.26 μm 53 HRC rollers) and 3.2 (uncoated Rq ¼0.56 μm discs vs. coated Rq ¼0.26 μm 53 HRC rollers) plotted versus stress cycles. (b) The same results for the rollers in Tests 1.4 and 3.3 from Table 1, with hardness of 57 HRC. uD/cR-S means uncoated discs/coated roller-smooth surfaces.
Fig. 9. (a) The wear scar of an uncoated disc after running against an uncoated roller in Test 1.4 (uncoated Rq ¼ 0.56 μm discs vs. uncoated Rq ¼ 0.26 μm 57 HRC roller) after about 0.58 million cycles. 3D surface morphology and line scan of the disc from Test 1.4 are shown in (b) and (c), respectively. (d) Wear scar of an uncoated disc after running against a coated roller in Test 3.3 (uncoated Rq ¼ 0.56 μm discs vs. coated Rq ¼0.26 μm 57 HRC roller) after about 0.58 million cycles. 3D surface morphology and line scan of the disc from Test 3.3 are shown in (e) and (f), respectively.
positive influence on the fatigue life of rolling element bearings [24] in boundary lubrication. It has been postulated that oil residing in the surface valleys can be squeezed out during elastic deflections of the surface, and bring lubricant to an otherwise starved contact region [25]. Coatings such as WC/a-C:H are known to exhibit no adhesive wear against steel at low to moderate temperatures [26]. On the other hand, steel–steel contacts can experience both abrasive and adhesive wear in low λ or dry conditions. Many studies have reported that the dry friction coefficient of diamond-like carbon–steel contacts is less than that of steel–steel contacts [8,27–31]. Results of this study indicate that the measured friction coefficients of the steel-WC/a-C:H contacts in boundary lubrication depend upon the roughness of the coated component.
Table 2 The surface roughness parameters of a disc before a test and after running 2 million cycles against an uncoated smooth roller and a WC/a-C:H coated smooth roller. Discs' and rollers' hardness values were 62 and 57 HRC, respectively. Value
Disc (initial values)
Rq (μm) 0.555 Ra (μm) 0.445 Rsk 0.028
Disc (Test 1.4) Steelsteel
Disc (Test 3.3) Steel-WC/ a-C:H
0.419 0.334 0.299
0.531 0.416 0.877
Interestingly, a comparison of the friction coefficients involving rough surfaces in Tests 1.2 and 2.1 indicates that the presence of the coating slightly increased the friction. This effect is probably
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due to abrasive wear occurring in these tests (especially since the measured friction in Tests 3.2 and 1.3 is similar). 3.5. Coated discs vs. coating roller The surface of a softer uncoated disc (H ¼58 HRC and Rq ¼0. 35 μm) from Test 3.7 that was run against a WC/a-C:H coated smooth roller at stress increments from 1.7 to 2.25 GPa for a total of 45 million cycles (equivalent to 3.3 million cycles on a disc) is shown in Fig. 10. There is no sign of micropitting or fatigue on the surface of disc. The coated roller counterpart is shown in Fig. 5g. Comparing Tests 2.2 and 3.7, it is concluded that the manifestation of micropitting on the uncoated counter face when the coating is applied to the discs could be different from when the coating is applied to the roller. Although λ is smaller in Test 3.7 than in Test 2.2, the contact stress and number of stress cycles on the uncoated counter face are larger. Additionally, micropitting did not occur on the uncoated disc in Test 3.7 (Fig. 10), while early stage micropitting occurred on the uncoated roller in Test 2.2 (Fig. 4b). Test results of the WC/a-C:H coating on smooth surfaces indicate that although the presence of the coating on the discs promotes the early onset of micropitting on the roller (before about 1 million cycles), when the coating is applied to the roller, no micropitting or high rate of wear occurs on either the roller or the discs – even after 45 million or more cycles. Each disc experiences one stress cycle compared to an equivalent 13.5 stress cycles of the roller, so asperities on the roller are subjected to more stress cycles than those on the disc surfaces. In boundary lubrication (0.1 o λ o0.2), the convective heat flow by the oil is negligible compared to the conduction of heat by the oil and the contacting surfaces. Since the thermal conductivity of hydrogen-containing DLC ( 1 W/mK [32]) is considerably smaller than 52100 steel ( 42 W/mK [33]), it is expected that most of the frictional heat generated in the contact will flow through the steel counter face when it is in contact with a DLC coated surface. Although the frictional heat generation in the contact is transient and of very short duration, it could still promote a rapid diffusion of carbon from the martensite matrix and the creation of a heterogeneous surface region containing ferrite and cementite or other FeC phases such as that observed and reported by Oila and Bull [10,11]. The time interval between thermal shocks is at least an order of magnitude shorter for an element of area on the roller than an element on the disc. Local temperatures on the surface or near-surface region of the rollers should be higher than those on the discs in these experiments. In addition to the time interval between the thermal shocks, the respective volumes of the disc and roller have an impact on the amount of heat that can be absorbed and dissipated through the components. During operation, two of the three discs are partially submerged in the oil sump, but the roller that is in the middle of
three discs is more thermally isolated. Therefore, when discs are coated with the DLC, the amount of frictional heat that flows into the uncoated roller should be larger than the amount of heat that is dissipated in uncoated discs when the roller is coated with the DLC. This qualitative thermal analysis is consistent with the observations that DLC coated discs cause an earlier onset of micropitting on an uncoated roller than when uncoated discs are tested against uncoated rollers with the same test parameters. Conversely, when the coating is applied to the roller, the discs do not exhibit micropitting but can experience an amount of abrasive wear that is consistent with the surface roughness of and hardness difference between the coated roller and the uncoated 52100 steel.
4. Conclusions The objective of this study was to evaluate the ability of a WC/ a-C:H coating to mitigate micropitting and wear damage on mated through-hardened 52100 bearing steel surfaces. The performance of the WC/a-C:H coating was evaluated in boundary lubrication at various λ ratios, slide-to-roll (SRR) ratios and steel hardness values. To determine the origin of micropitting in experiments of this type, it is desirable to understand the evolution of the surface texture during the testing. Although it was not the objective of this study to determine the origin of micropitting in bearing steel, it is certainly worthwhile to pursue that objective in future study. The major conclusions drawn from this study are: 1. Abrasive wear was the dominant surface damage mechanism whenever at least one of contacts included a coated rough surface (Rq ¼0.56 μm). 2. The WC/a-C:H coating experienced a very high rate of wear during testing when it was applied to rough (Rq ¼0.56 μm) substrates, but virtually no wear when the coating was applied to smoother (Rq ¼0.26 μm) substrates. 3. Under boundary lubrication, surface damage of rollers was primarily wear for λ o0.1 and micropitting for λ 40.1. Micropitting of rollers was observed when pmax was comparable to or greater than the estimated yield strength of the 52100 steel, 4. Whereas wear of the roller occurred when the SRR ¼ þ10%, micropitting was the surface damage mode on the roller when the slide-to-roll ratio was 10%. The effect was attributed to whether tensile shear stresses resided in the inlet or outlet positions of the lubricated contact, and was found to be consistent with the observation that micropitting of gear teeth occurs in the region of negative rather than positive sliding. 5. The WC/a-C:H coating eliminated the occurrence of roller wear and micropitting when the coating was applied to the roller. However, when the coating was applied to the discs, the onset of micropitting on the uncoated roller occurred sooner than on rollers paired with uncoated discs. A possible explanation for
Fig. 10. Optical microscopy images of the surface of the uncoated disc from Test 3.7 after running 45 million cycles at 1.7/2/2.25 GPa contact stress against a WC/a-C:H coated roller with hardness of 62 HRC and roughness of Rq ¼ 0.29 μm. (a) Magnification 50 . (b) Magnification 200x.
B. Mahmoudi et al. / Wear 350-351 (2016) 107–115
this behavior was provided. The thermal conductivity difference between the coating and the 52100 steel allows more heat to flow into the uncoated counter face than the coated one. When the discs are coated, most of the frictional heat flows into the roller, increasing the diffusion rate of C in the martensite matrix and driving the phase changes that can give rise to micropitting. When the roller is coated, most of the frictional heat flows into the three discs that are not as thermally isolated and have considerably more mass than the roller.
Acknowledgments The authors are grateful to The Timken Company for their generosity in funding this study. We wish to thank Dr. Luc Houpert and Dr. Young Kang of The Timken Company, and Prof. Yalin Dong of the University of Akron, for beneficial discussions. We are indebted to Mr. R. Fowler and Dr. B. Tury of the University of Akron for their valuable assistance in this project.
References [1] ASTM G40-15, Standard Terminology Relating to Wear and Erosion, 2015. [2] M.N. Kotzalas, G.L. Doll, Tribological advancements for reliable wind turbine performance, Philos. Trans. R. Soc. A 368 (2010) 4829–4850. [3] P.H. Dawson, Effect of metallic contact on the pitting of lubricated rolling surfaces, J. Mech. Eng. Sci. 4 (1) (1962) 16–21. [4] P.H. Dawson, Further experiments on the effect of metallic contact on the pitting of lubricated rolling surfaces, Proc. Insti. Mech. Eng. 180 (Pt 3B) (1965&;1966) 95–112. [5] T. Ahlroos, H. Ronkainen, A. Helle, R. Parikka, J. Virta, S. Varjus, Twin disc micropitting tests, Tribol. Int. 42 (10) (2009) 1460–1466. [6] R.S. Zhou, H.S. Cheng, T. Mura, Micropitting in rolling and sliding contact under mixed lubrication, J. Tribol. 111 (1989) 605–613. [7] H.A. Spikes, A.V. Olver, P.B. Macpherson, Wear in rolling contacts, Wear 112 (1986) 121–144. [8] B. Bhushan, Contact mechanics of rough surfaces in tribology: multiple asperity contact, Tribol. Lett. 4 (1998) 1–35. [9] G.E. Morales-Espejel, V. Brizmer, Micropitting modelling in rolling-sliding contacts: application to rolling bearings, Tribol. Trans. 54 (2011) 625–643. [10] A. Oila, S.J. Bull, Phase transformations associated with micropitting in rolling/ sliding contacts, J. Mater. Sci. 40 (2005) 4767–4774. [11] A. Oila, S.J. Bull, Assessment of the factors influencing micropitting in rolling/ sliding contacts, Wear 258 (2005) 1510–1524.
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[12] T.W. Scharf, M.C. Romanes, K.C. Mahdak, J.Y. Hwang, R. Banerjee, R.D. Evans, G. L. Doll, Atomic-scale structure and composition of tungsten carbide reinforced diamondlike carbon films, Appl. Phys. Lett. 93 (15) (2008) 151909–151909-3. [13] R. Evans, C. Hager Jr., Y. Kang, G. Doll, Comparison of black oxide and tungsten carbide – reinforced diamond like carbon (WC/a-C:H) surface treatments for rolling element bearings, Tribol. Trans. 58 (2015) 444–453. [14] E. Laine, A.V. Olver, T.A. Beveridge, Effect of lubricants on micropitting and wear, Tribol. Int. 41 (2008) 1049–1055. [15] C. Benyajati, A.V. Olver, C.J. Hamer, An experimental study of micropitting, using a new miniature test-rig, Tribol. Series 43 (2003) 601–610. [16] J.A. Williams, Engineering Tribology, Oxford University Press, Oxford, 1994. [17] C. Strondl, G.J. van der Kolk, T. Hurkmans, W. Fleischer, T. Trinh, N.M. Carvalho, J.Th.M. de Hosson, Properties and characterization of multilayers of carbides and diamond-like carbon, Surf. Coat. Technol. 142–144 (2001) 707–713. [18] M. Eckels, M.N. Kotzalas, G.L. Doll, Attaining high levels of bearing performance with a nanocomposite diamond-like carbon coating, Tribol. Trans. 56 (3) (2013) 410–416. [19] R.D. Evans, T.A. Barr, L. Houpert, S.V. Boyd, Prevention of smearing damage in cylindrical roller bearings, Tribol. Trans. 56 (2013) 703–716. [20] R.D. Evans, P.J. Shiller, J.Y. Howe, Adhesion of tungsten carbide reinforced amorphous hydrocarbon thin films (WC/aC: H) to steel substrates for tribological applications, J. Appl. Phys. 109 (2) (2011) 023518. [21] D. Tabor, Hardness of Metals, Clarendon Press, Oxford, 1951. [22] B. Mahmoudi, B. Tury, C.H. Hager, G.L. Doll, Effects of black oxide and a WC/aC:H coating on the micropitting of SAE 52100 bearing steel, Tribol. Lett. 58 (2) (2015) 1–9. [23] R. L. Errichello, Morphology of micropitting, AGMA 11FTM17, 2011. [24] Y. Akamatsu, N. Tsushima, T. Goto, K. Hibi, Influence of surface roughness skewness on rolling contact fatigue life, Tribol. Trans. 35 (4) (1992) 745–750. [25] X. Zhai, L. Chang, M.R. Hoeprich, H.P. Nixon, On mechanisms of fatigue life enhancement by surface dents in heavily loaded rolling line contacts, Tribol. Trans. 40 (4) (1997) 708–714. [26] A. Grill, Tribology of diamond-like carbon and related materials: an updated review, Surf. Coat. Technol. 94 (1997) 507–513. [27] K. Holmberg, A. Matthews, Coatings Tribology: Properties, Mechanisms, Techniques and Applications in Surface Engineering (Tribology and Interface Engineering Series, v. 56), Elsevier, Amsterdam, 2009. [28] C. Donnet, A. Erdemir, Tribology of Diamond-Like Carbon Films: Fundamentals and Applications, Springer, Berlin, 2008. [29] B. Podgornik, J. Vižintin, S. Jacobson, S. Hogmark, Tribological behaviour of WCyC coatings operating under different lubrication regimes, Surf. Coat. Technol. 177–178 (2004) 558–565. [30] C.A. Manier, G. Theiler, D. Spaltmann, M. Woydt, H. Ziegele, Benchmark of thin film coatings for lubricated slip-rolling contacts, Wear 268 (10–11) (2010) 1442–1454. [31] P. Kodali, K.C. Walter, M. Nastasi, Investigation of mechanical and tribological properties of amorphous diamond-like carbon coatings, Tribol. Int. 30 (8) (1997) 591–598. [32] K.T. Wojciechowski, R. Zybala, R. Mania, Application of DLC layers in 3-omega thermal conductivity method, J. Achiev. Mater. Manuf. Eng. 37 (2) (2009) 512–517. [33] M.F. Ashby, J. Abulawi, H.S. Kong, Temperature maps for frictional heating in dry sliding, Tribol. Trans. 34 (4) (1991) 577–587.
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Wear journal homepage: www.elsevier.com/locate/wear
Influence of a WC/a-C:H tribological coating on micropitting wear of bearing steel B. Mahmoudi a, G.L. Doll a,n, C.H. Hager Jr.b, R.D. Evans b a b
Timken Engineered Surfaces Laboratories, The University of Akron, Akron, OH 44325, USA The Timken Company, Canton, OH 44720, USA
art ic l e i nf o
a b s t r a c t
Article history: Received 23 September 2015 Received in revised form 13 January 2016 Accepted 17 January 2016 Available online 29 January 2016
The focus of this work was to evaluate the use of a WC/a-C:H coating, specifically designed to withstand rolling contact fatigue, to mitigate micropitting damage on through-hardened AISI 52100 bearing steel. This was accomplished by conducting experiments with various combinations of coated discs and rollers using a three-disc-on-roller rolling contact fatigue machine. In addition to testing coated and uncoated contacts, discs and rollers with three hardness levels and different surface roughness values were also evaluated. Upon the completion of this work, it was observed that the onset of micropitting damage was affected by the hardness of the asperities on the mated roller and disc samples. It was also observed that the coated surface can cause abrasive wear damage on an uncoated surface if the coating is applied to a surface that has a roughness Rq Z 0.5 μm. If the coating is applied to a smoother roller surface (Rq r0.26 μm), the coating is able to mitigate micropitting damage on the roller surface without causing high abrasive wear rates on the disc surfaces. & 2016 Elsevier B.V. All rights reserved.
Keywords: Micropitting Diamondlike carbon Abrasive wear Bearing steel
1. Introduction Micropitting is a form of surface damage in rolling and mixed mode contacts consisting of numerous pits and associated cracks on a scale smaller than that of the Hertz elastic contact [1]. Once initiated, micropitting damage can alter the contact stress profile of the mated components, which may eventually cause stress concentrations so large that they can exceed the yield strength of the material. As a specific example, micropitting has been identified as a root cause of raceway spalling of main shaft spherical roller bearings in wind turbines. This spalling often occurs at fractions of the designed bearing life [2]. Several studies have been reported on micropitting of steel contacts, and all have correlated the onset of micropitting wear to the λ ratio (the ratio of the lubricant film thickness to the composite surface roughness) [3–7]. Decreases in λ and increases in slide-to-roll ratio up to 2% can elevate the risk of micropitting and reduce the number of stress cycles required to initiate micropitting [8,9]. Slide-to-roll ratio (SRR) describes the ratio of the sliding velocity to the entrainment velocity. If two surfaces are moving at the same surface speed, the motion is pure rolling (0.0% sliding). The n
Corresponding author. E-mail address: [email protected] (G.L. Doll).
http://dx.doi.org/10.1016/j.wear.2016.01.010 0043-1648/& 2016 Elsevier B.V. All rights reserved.
entrainment velocity is the mean velocity of the contacting surfaces, as described in Eq. (1). Positive and negative sliding occurs when the discs are rotating with faster and slower surface speeds than the roller, respectively. SRR is calculated based from the following equations: ue ¼
ðud þ ur Þ 2
SRR ¼
ðud ur Þ ue
ð1Þ ð2Þ
where ud , ur are the surface speeds of a disc and roller, and ue is the entrainment velocity. Based upon the results of twin disc experiments, Oila and Bull have proposed a mechanism for micropitting [10,11]. Their proposed mechanism suggests that plastic deformation associated with contacting asperities extends beneath the asperities and causes an increase in the dislocation density, which causes work hardening. Simultaneously, frictional heating is sufficient to activate the diffusion of carbon. At the boundary of the plastically deformed region (PDR), recrystallization occurs. Dislocations pile up against the PDR boundary and a slip band forms. With continuing cycles, dislocations accumulate in the slip band and a crack initiates. The crack propagates along the PDR boundary until it reaches the surface, where it generates a micropit.
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Considering the Oila and Bull suggested mechanism for micropitting and its experimental correlation with the λ ratio, it is hypothesized that a tribological coating may be able to mitigate micropitting damage if it can increase the λ ratio and reduce the frictional heating between asperities. Candidate coatings for this task need to be able to withstand a high number of stress cycles without fracturing, be strongly adherent to the steel substrates, and have deposition temperatures that are compatible with the heat treatment practices of the steel alloy. One coating that meets all of these requirements is a WC/a-C:H coating that consists of nanometer-sized β-W2C precipitates in an amorphous hydrocarbon matrix [12]. Based on nanoindentation with a Berkovich tip using the continuous stiffness measurement method, the coating is moderately harder ( 13 GPa) than hardened bearing steel, and functionally graded to ensure strong adhesion to steel [13]. Although the nanoindentation hardness measurement of 52100 from Ref. [13] was around 10.6 GPa, most indentation measurements show that 61 71 HRC equates to about 7 GPa. The objective of this study was to evaluate the ability of a WC/ a-C:H coating, which satisfies all the above requirements, to mitigate micropitting damage on mated through-hardened 52100 bearing steel surfaces. The performance of the WC/a-C:H coating was evaluated in boundary lubrication at various λ ratios, slide-toroll (SRR) ratios and steel hardness values.
2. Experimental details Testing was conducted using a PCS Instruments Micropitting Rig (MPR) with three through-hardened AISI 52100 1 cm wide discs (disc radius Rx ¼27 mm and crown Ry ¼1) and one throughhardened AISI 52100 steel roller (Rx ¼6 mm, Ry ¼12 mm). Images of the MPR test chamber and the roller geometry are shown in Fig. 1a and b, respectively. The description of this test rig can be found in [14,15]. The geometries of the discs and rollers create elliptical contact where the contact width between the roller and disc is between 500 and 700 μm depending upon load, wear mode and number of cycles. Discs were heat-treated to a hardness of 62 HRC, while the roller hardness values were 53, 57 and 62 HRC. It is acknowledged that the 53 and 57 HRC roller hardness values are lower than those typically specified for rolling element bearing steels. However, using this wide range of roller hardness in the MPR tests allowed for a clearer observation of the influence of hardness on this damage mode and confirmation of the efficacy of bearing steel hardness specifications [7]. The influence of surface roughness on wear was studied by using discs and rollers with two root mean square (RMS) surface roughness values, referred to as either smooth (Rq ¼0.26 μm) or rough (Rq ¼0.56 μm). The variation of surface roughness values is 70.08. Surface geometries, topographies, and roughness were characterized using a Zygo NewView 7300 3D white light interferometry
microscope. The complete test matrix is shown in Table 1. The tests were designed to probe how micropitting and wear depend upon various combinations of material pairs. Specifically, Series 1 (Tests 1.1– 1.9) explores how the hardness difference between the rollers and discs affects micropitting and wear, Series 2 (Tests 2.1–2.2) evaluates micropitting and wear of rollers when paired against coated discs, and Series 3 (Tests 3.1–3.7) probes micropitting and wear of rollers coated with WC/a-C:H. The entrainment velocity (ue) and oil temperature were chosen to achieve specific central lubricant film thicknesses calculated according to Eq. (3) [16]. h
min λ ¼ qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi
R2ar þ R2ad
0:49 0:073 hmin ¼ 1:79 R0 U 0:68 Σ GΣ W Σ U η0 P UΣ ¼ GΣ ¼ 2γ E0 ; W Σ ¼ 0 0 ; 0 2 2E R 2E R
ð3Þ
where λ is the film thickness ratio, hmin is film thickness, Rar is average rms roughness for roller, Rad is average rms roughness for disc, U is the sliding velocity, η0 is the dynamic viscosity, E0 is the equivalent modulus of elasticity, R0 is the reduced radius of curvature, γ is the pressure–viscosity constant, and P is the applied force. The test oil for all tests was an ISO viscosity grade 10 polyalphaolefin (PAO ISO 10) containing no extreme pressure (EP) or anti-wear (AW) additives except that used in Test 1.9 where the lubricant was PAO ISO 68 with no EP and AW additives. The viscosity of the PAO ISO 10 was measured to be 8.72 and 2.06 cP at 38 and 100 °C, respectively, which yielded pressure–viscosity coefficients of 13.5 GPa 1 and 10.15 GPa 1 at 38 and 100 °C. The WC/a-C:H coating was deposited onto roller and disc specimens in a closed-field unbalanced magnetron sputtering system. Strondl et al. have provided an overview of this deposition technology [17]. Roller bearings with this particular coating are able to achieve low λ (i.e., λ o1) fatigue lives more than three times greater than uncoated steel bearings, can tolerate a significant amount of debris damage without a reduction in calculated L10 life, and are very resistant to adhesive wear damage [18,19]. The coating is a layered material consisting of a 100 nmthick Cr layer that forms strong chemical bonds to the FeO surface layer on steel [20]. In the next layer, the composition is gradually changed from Cr to WC/a-C:H. The thickness of this gradient layer is also about 100 nm thick. Finally, the top layer of WC/a-C:H is approximately 900 nm thick. This yields a total coating thickness of approximately 1.1 μm. Testing on the MPR was interrupted periodically in order to perform a visual inspection of the wear track. If no evidence of micropitting or serious wear was observed, the test was restarted. This procedure was carried out until wear was observed or the test was suspended at 8 million cycles. In one case, the test was terminated at 30 million cycles with no significant wear or micropitting in order to measure the durability of the coating.
Fig. 1. (a) MPR chamber, where three larger discs (Rx: 27 mm) are running against one smaller roller in the middle. (b) Details of roller's geometry (Rx: 6 mm and Ry: 12 mm), which leads to elliptical Hertzian contact with cylindrical discs.
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Table 1 Micropitting test matrix. pmax represents the maximum calculated contact stress, ue is the entrainment velocity, SRR is the slide-to-roll ratio, and λ is the ratio of the calculated central lubricant film thickness to the composite surface roughness (Rq). Test
Coating
Rq (μm)
HRC
Discs
Roller
Discs
Roller
Discs
Roller
1.1 1.2 1.3 1.4 1.5 1.6 1.7
N N N N N N N
N N N N N N N
62 62 62 62 62 62 62
62 57 53 57 57 57 57
0.56 0.56 0.26 0.56 0.56 0.56 0.56
0.26 0.26 0.26 0.26 0.26 0.26 0.29
1.8 1.9 2.1 2.2 3.1 3.2 3.3 3.4 3.5
N N Y Y N N N N N
N N N N Y Y Y Y Y
62 62 62 62 62 62 62 62 62
62 62 57 57 62 53 57 57 57
0.56 0.56 0.56 0.26 0.26 0.26 0.56 0.56 0.56
0.26 0.56 0.26 0.26 0.56 0.26 0.26 0.26 0.26
3.6
N
Y
62
57
0.56
0.29
3.7
N
Y
58
62
0.35
0.27
P (GPa)
ue (m/s)
SRR (%)
T (°C)
λ
N106
μ
Comments
1.7 1.5 1.5 1.7 2 2 2 2.25 2.25 2.25 1.5 1.5 1.5 1.5 1.7 2 2 2.25 2.55 2 2.25 2.42 1.7 2 2.25
1 1 2 2 2 2 2
2 2 2 2 10 0 10
40 40 75 75 40 40 40
10 30 2 8 0.5 2 2/1
0.07 0.05 0.06 0.06 0.09 0.05 0.13
Mild wear Mild wear Wear Mild wear Wear þ cracking Micropitting þ cracking Micropitting
2 2 1 1 1 2 2 2 2
0 0 2 0 0 2 2 10 0
40 40 40 40 40 75 75 40 40
Micropitting Micropitting Wear Micropitting Wear No wear or micropitting No wear or micropitting No wear or micropitting; partial wear of coating No wear or micropitting
10
40
0.12
Wear of coating
2
2
75
1 1 2 1.5 7 10 8.5 1 15 7 6 2 6 6 15 15 15
0.06 0.03 0.07 0.05 0.08 0.05 0.05 0.09 0.04
2
0.09 0.09 0.12 0.07 0.14 0.14 0.13 0.13 0.13 0.4 0.09 0.15 0.09 0.12 0.07 0.14 0.16 0.16 0.16 0.14 0.13 0.13 0.10 0.10 0.09
0.06
No wear or micropitting
Fig. 2. Optical microscopy images of the wear tracks on uncoated rollers produced during Tests 1.1 to 1.9. (a) Roller surface from Test 1.1 after 10 million cycles. (b) Roller surface from Test 1.2 after 30 million cycles. (c) Roller surface from Test 1.3 after 1 million cycles. (d) Roller surface from Test 1.4 after 8 million cycles. (e) Roller surface from Test 1.5 after 0.5 million cycles. (f) Roller surface from Test 1.6 after 0.5 million cycles. (g) Roller surface from Test 1.7 after 1 million cycles. (h) Roller surface from Test 1.8 after 0.5 million cycles. (i) Roller surface from Test 1.9 after 0.5 million cycles. All photos have same scale. The arrows indicate the direction of rolling (R) and sliding (S). Positive sliding means when rolling and sliding are pointing to a same direction.
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3. Results and discussion 3.1. Series 1 Fig. 2a–i display optical micrographs taken of the surfaces of uncoated rollers that have run against uncoated discs in test Series 1, and results of these tests are displayed in Table 1. Surfaces of the mated discs from this Series are not shown since no measurable changes occurred during testing. In this study, wear refers to a change in the roughness and profile of counterparts and mild wear refers to change only in surface roughness but not in the profile of the mated contacts. The arrows indicate the direction of rolling (R) and sliding (S). Roller surface damage is attributed to wear or mild wear in (a), (b), (d) and (e) corresponding to Tests 1.1, 1.2, 1.4, and 1.5, respectively. Tests 1.1, 1.2, and 1.4 were performed at SRR r2% and λ o0.1. Test 1.5 was performed at SRR ¼ þ10% and λ ¼0.4. Roller surface damage is attributed to micropitting in (f), (g), (h), and (i), corresponding to Tests 1.6, 1.7, 1.8, and 1.9, respectively. These tests were all performed at values of λ 4 0.1. The surface damage on the roller shown in (c) from Test 1.3 appears to be severe wear, plastic deformation, and tearing. In Series 1, wear or mild wear of the roller occurred when λ was less than 0.1, and micropitting occurred when λ 40.1 for SRR ¼2% or less. An exception to this trend is Test 1.3, where the roller was 9 points softer than the discs on the HRC scale. Test 1.5 was performed with SRR ¼ þ10%, and although the initial λ value was 0.4, abrasive wear occurred. The competition between abrasive wear and micropitting (or surface fatigue) was noted by Morales-Espejel et al. [9], who performed MPR experiments similar to those in Series 1. Micropitting is a surface fatigue phenomenon that requires a number of stress cycles to initiate and propagate cracks. If the rate of material removal by wear is large enough, near surface material does not maintain residence long enough in the contact to accrue the requisite number of stress cycles for crack initiation. Therefore under these test conditions, the transition from a wear mode to micropitting appears to occur near λ ¼ 0.1. If the roller hardness is much less than the counter face (as is the case in Test 1.3), the lambda threshold should be much larger. Additionally, with increasing SRR, the lambda threshold between wear and micropitting also increases. For example, Fig. 3a and b show roller profiles before and after testing from Tests 1.5 and 1.6, respectively. The roller from Test 1.5 has experienced a significant loss of material while the roller from Test 1.6 has maintained its original profile, but with micropitting. In MPR tests performed under boundary lubrication, asperities on the surface of the roller plastically deform under load, which according to the mechanism put forth by Oila and Bull can create work hardening and eventually lead to micropitting [10]. Within the framework of this proposed mechanism, surface fatigue and micropitting initiation should depend upon the yield strength of the materials in contact and the contact stress. Comparing the results of Tests 1.2, 1.4, 1.6, and 1.8, the contact stress that produced micropitting was relatively close to the yield strength of the
roller. Test 1.2 and 1.4 rollers were 57 HRC (or HV 6.6 GPa), and did not exhibit micropitting damage even after as many as 30 million stress cycles. According to Tabor [21], the yield strength (σy) of a material is approximately one-third the Vickers hardness (H). Using this approximation, the yield strength of materials with H¼ 6.6 GPa is σy ¼2.2 GPa. Therefore, if the magnitude of the stresses on the asperities is similar to the calculated maximum Hertzian stress in the contact, tests performed with pmax o2.2 GPa may be insufficient to work harden the asperities and create micropitting. When pmax approached the value of 2.2 GPa in Tests 1.5 and 1.6, the roller surfaces exhibited a combination of wear and cracking.
3.2. Series 2 Fig. 4a and b show the surfaces of uncoated rollers after 1.5 million cycles that were tested against WC/a-C:H coated (Rq ¼0.56 μm) discs (Test 2.1) and WC/a-C:H coated (Rq ¼0.26 μm) discs (Test 2.2), respectively. The WC/a-C:H coating on the rough discs has not only removed the grind lines in the wear track on the roller, but has also removed a considerable amount of the profile of the roller, which can be observed in the profile trace shown in Fig. 4c. The original roller profile radius (Ry ¼12 mm) is no longer present, and the disc-roller contact is much wider. As a result, the Hertzian contact stress of the test was significantly reduced from its initial value of 1.5 GPa during the testing. In comparison, the uncoated roller that was tested against the Rq ¼0.26 μm coated discs in Test 2.2 exhibited micropitting after just 1 million stress cycles, as shown in Fig. 4b and d. Again, the transition between wear and micropitting for the two tests performed in this Series appears to be near an initial λ value of 0.1. The reduction in the magnitude of the grinding lines in the wear track of Fig. 4b indicates that the coated disc was also polishing the uncoated roller. Although polishing of the counter face increases the λ value slightly (from 0.09 to 0.11), in this case, the increase was not enough to avoid micropitting. These results indicate that adding the coating on the disc surface will not prevent micropitting damage on the roller under these test conditions. Almost all of the coating was removed from the discs in Test 2.1, where the WC/a-C:H was applied to a rough roller surface (Rq ¼0.56 μm). Subsequent to this test, submicron-sized particles of the coating were found suspended in the tested oil. Conversely, no coating wear was measured after 1.5 million cycles in Test 2.2. In a previous study [22], wear of the WC/a-C:H coating under this test condition was determined to be associated with fracture that initiated at high aspect ratio features at the substrate/coating interface. These results indicate that the Rq ¼ 0.56 μm surface is too rough for the WC/a-C:H coating to function effectively in rolling contact.
Fig. 3. Profiles of rollers showing the higher rate of wear as slide-to-roll ratio increases from 0.0% to 10.0%. (a) Comparison of the profile of roller from Test 1.5 before and after running 0.5 million cycles. (b) Comparison of roller profile before and after running 0.5 million cycles from Test 1.6.
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3.3. Series 3 Although adding the coating to the discs did not mitigate micropitting damage on the rollers, adding the coating to the rollers did. Fig. 5 shows the surfaces of coated rollers that ran against uncoated discs. The arrows indicate the direction of rolling (R) and sliding (S). Fig. 5b shows the undamaged roller surface after Test 3.2. The coated rollers were able to mitigate wear and micropitting damage in every test except Test 3.6 (shown
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in Fig. 5f). When the slide-to-roll ratio was 10% in boundary lubrication and the contact stress was greater than 2 GPa, the coated surface exhibited surface damage. Table 1 shows that the friction measured during Test 3.6 (SRR ¼ 10%) was about 30% greater than the friction measured during Test 3.4 (SRR ¼ þ10%). The initiation of micropitting and coating delamination on the roller may be due to a synergistic effect of normal stress from Hertzian contact and shear stress due to friction. For a positive SRR (ud 4 ur), the shear force is in the rolling direction. However, in
Fig. 4. (a) Optical microscopy images of the wear tracks of uncoated rollers after 1.5 million cycles tested against WC/a-C:H-coated rough discs (Rq ¼ 0.56 μm) in Test 2.1. (b) Wear track of uncoated roller tested against WC/a-C:H-coated smooth discs (Rq ¼ 0.26 μm) in Test 2.2. (c) Comparison of roller profiles (Ry) from Test 1.2 (Rq ¼ 0.56 μm uncoated disc against Rq ¼ 0.26 μm uncoated roller) and Test 2.1 (Rq ¼ 0.56 μm coated disc against Rq ¼ 0.26 μm uncoated roller). (d) Comparison of roller profiles (Ry) from Test 1.1 (Rq ¼0.56 μm uncoated disc against Rq ¼ 0.26 μm uncoated roller) and Test 2.2 (Rq ¼ 0.26 μm coated disc against Rq ¼0.26 μm uncoated roller). uD/uR-R means uncoated discs/uncoated roller-rough surfaces. cD/uR-S means coated discs/uncoated roller-smooth surfaces.
Fig. 5. Optical microscopy images of the wear tracks on coated rollers after Tests 3.1 to 3.7. (a) Roller surface from Test 3.1 from Table 1 after 7 million cycles. (b) Roller surface from Test 3.2 after 10 million cycles. (c) Roller surface from Test 3.3 after 8.5 million cycles. (d) Roller surface from Test 3.4 after 1 million cycles. (e) Roller surface from Test 3.5 after 15 million cycles. (f) Roller surface from Test 3.6 after 8 million cycles. (g) Roller surface from Test 3.7 after 45 million cycles. All photos have same scale. The arrows indicate the direction of rolling (R) and sliding (S). Positive sliding means when rolling and sliding are pointing to a same direction.
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Fig. 6. (a) Comparison of the Ry profiles of an uncoated disc before and after the testing against rough coated roller (Test 3.1). (b) Wear scar on disc from Test 3.1 (uncoated Rq ¼ 0.26 μm discs versus coated Rq ¼0.56 μm rollers) after 7 million cycles of roller contact (each 13.5 contacts of roller is equal to one contact of a disc). uD/cR-R means uncoated discs/coated roller-rough surfaces.
negative SRR where ud our, the shear stress is opposite to the rolling direction. It is likely that the direction of the shear stress with respect to the rolling direction affects the damage progression of the coating. Errichello has pointed out that the dedendum of gear teeth (which experiences negative sliding) is more prone to micropitting than the mating gear surface that undergoes positive sliding (the tooth addendum) [23]. Since crack initiation is generally associated with tensile stresses, the shear stress due to positive sliding should generate tensile stress in the coated roller at the inlet position of the contact where the lubricant film thickness is equal to or greater than the central lubricant film thickness. On the other hand, negative SRR generates a maximum tensile stress at the outlet position on the coated roller where the lubricant film thickness is at a minimum. In this study, the lubricant film thickness at the outlet position of the contact is estimated to be about 30% thinner than the central film thickness. Fig. 6a and b show wear scars that were produced on the discs mated with the coated roller in Test 3.1 after 7 million roller stress cycles. In this test, the coated roller was much rougher than the mated discs. This led to a greater amount of wear on the disc surfaces than tests conducted with smoother rollers. Fig. 6a compares the Ry profile for the wear scars of the discs before and after the test. As with Test 2.1 (Fig. 4c), the coating on the rough roller surface in Test 3.1 caused a high rate of wear on the uncoated surface of the disc. 3.4. Wear of uncoated counter face When two surfaces are subjected to high Hertzian contact pressure, the asperities of the mating surfaces elasto-plastically deform to support the applied load. The plastic deformation of the WC/a-C:H asperities should be considerably less than that of the uncoated part, since the coating is harder but more elastic than through-hardened 52100 steel. Therefore, if the roughness of the coated component is relatively large, the coating asperities will continue to remove material from the uncoated counter face until either the contact pressure minimizes due to a change in the profile of the uncoated counter face, or the lubricant film thickness completely separates the asperities of the mating surfaces. This effect is displayed graphically in Fig. 7. Here, the contact stress has been calculated from periodic Ry measurements of the roller surfaces during testing involving Rq ¼0.56 μm discs with and without the WC/a-C:H coating (Tests 1.2 and 2.1). The contact stress between the coated discs and roller in Test 2.1 decreased during testing from 1.5 to nearly 1 GPa. The reduction of the contact stress through wear involving the uncoated contacts in Test 1.2 was much less (a much smaller change in the roller profile). Additionally, the Figure shows that most of the wear on the uncoated rollers occurred during the early stages of Test 2.1.
Fig. 7. Calculated maximum Hertzian contact stress plotted versus cycles for Tests 1.2 (uncoated Rq ¼ 0.5 μm discs versus uncoated Rq ¼0.26 μm roller) and 2.1 (coated Rq ¼0.5 μm discs versus uncoated Rq ¼ 0.26 μm roller). pmax is calculated from Ry measurements of the roller surfaces taken at different cycles during the tests. uD/ uR-R means uncoated discs/uncoated roller-rough surfaces.
The wear volume of the uncoated counter face was significantly reduced when the coating was applied to smoother surfaces. Since the wear rate of the coated part was very low, the total wear volume of the system was associated primarily with the uncoated steel counter faces. The change in the separation between the centers of the discs and the rollers (i.e., the displacement) measured during Tests 1.3 (uncoated Rq ¼ 0.26 μm discs vs. uncoated Rq ¼0.26 μm 53 HRC rollers) and 3.2 (uncoated Rq ¼ 0.26 μm discs vs. coated Rq ¼0.26 μm 53 HRC rollers) is plotted versus stress cycles in Fig. 8a after the first 2 million test cycles. Fig. 8b shows similar results for Tests 1.4 and 3.3. The displacement curves approximate the wear depth and indicate that about 33% less material was removed from the uncoated steel specimens when they were in contact with the coating instead of the uncoated steel. Moreover, Test 3.2 (uncoated Rq ¼0.56 μm discs vs. coated Rq ¼0.26 μm 53 HRC rollers) achieved more than five times more cycles without micropitting or wear than Test 1.3 (uncoated Rq ¼0.56 μm discs vs. uncoated Rq ¼0.26 μm 53 HRC rollers), which exhibited severe wear before 2 million cycles. Fig. 9a–f illustrate the surface morphologies of the discs that were run against an uncoated roller (Test 1.4) and a WC/a-C:Hcoated roller (Test 3.3) after about 2 million cycles. The grind lines are still visible on the surfaces of a disc from Test 1.4 (Fig. 9a and b) and a disc from Test 3.3 (Fig. 9c and d), although the coating on the roller in Test 3.3 appears to have visually altered the appearance of the grind lines on the disc. Ra, Rq and Rsk values obtained from the 3D surface measurements in Fig. 9 are presented in Table 2. A significant change in the Rsk values has occurred as a result of the testing. Rsk values became negative on both disc surfaces, and a disc that ran against the coated roller attained an Rsk value approaching 1. Negative Rsk values have been reported to have a
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Fig. 8. The change in the separation between the centers of the discs and the rollers (i.e., the displacement) measured during (a) Tests 1.3 (uncoated Rq ¼ 0.56 μm discs vs. uncoated Rq ¼0.26 μm 53 HRC rollers) and 3.2 (uncoated Rq ¼0.56 μm discs vs. coated Rq ¼0.26 μm 53 HRC rollers) plotted versus stress cycles. (b) The same results for the rollers in Tests 1.4 and 3.3 from Table 1, with hardness of 57 HRC. uD/cR-S means uncoated discs/coated roller-smooth surfaces.
Fig. 9. (a) The wear scar of an uncoated disc after running against an uncoated roller in Test 1.4 (uncoated Rq ¼ 0.56 μm discs vs. uncoated Rq ¼ 0.26 μm 57 HRC roller) after about 0.58 million cycles. 3D surface morphology and line scan of the disc from Test 1.4 are shown in (b) and (c), respectively. (d) Wear scar of an uncoated disc after running against a coated roller in Test 3.3 (uncoated Rq ¼ 0.56 μm discs vs. coated Rq ¼0.26 μm 57 HRC roller) after about 0.58 million cycles. 3D surface morphology and line scan of the disc from Test 3.3 are shown in (e) and (f), respectively.
positive influence on the fatigue life of rolling element bearings [24] in boundary lubrication. It has been postulated that oil residing in the surface valleys can be squeezed out during elastic deflections of the surface, and bring lubricant to an otherwise starved contact region [25]. Coatings such as WC/a-C:H are known to exhibit no adhesive wear against steel at low to moderate temperatures [26]. On the other hand, steel–steel contacts can experience both abrasive and adhesive wear in low λ or dry conditions. Many studies have reported that the dry friction coefficient of diamond-like carbon–steel contacts is less than that of steel–steel contacts [8,27–31]. Results of this study indicate that the measured friction coefficients of the steel-WC/a-C:H contacts in boundary lubrication depend upon the roughness of the coated component.
Table 2 The surface roughness parameters of a disc before a test and after running 2 million cycles against an uncoated smooth roller and a WC/a-C:H coated smooth roller. Discs' and rollers' hardness values were 62 and 57 HRC, respectively. Value
Disc (initial values)
Rq (μm) 0.555 Ra (μm) 0.445 Rsk 0.028
Disc (Test 1.4) Steelsteel
Disc (Test 3.3) Steel-WC/ a-C:H
0.419 0.334 0.299
0.531 0.416 0.877
Interestingly, a comparison of the friction coefficients involving rough surfaces in Tests 1.2 and 2.1 indicates that the presence of the coating slightly increased the friction. This effect is probably
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due to abrasive wear occurring in these tests (especially since the measured friction in Tests 3.2 and 1.3 is similar). 3.5. Coated discs vs. coating roller The surface of a softer uncoated disc (H ¼58 HRC and Rq ¼0. 35 μm) from Test 3.7 that was run against a WC/a-C:H coated smooth roller at stress increments from 1.7 to 2.25 GPa for a total of 45 million cycles (equivalent to 3.3 million cycles on a disc) is shown in Fig. 10. There is no sign of micropitting or fatigue on the surface of disc. The coated roller counterpart is shown in Fig. 5g. Comparing Tests 2.2 and 3.7, it is concluded that the manifestation of micropitting on the uncoated counter face when the coating is applied to the discs could be different from when the coating is applied to the roller. Although λ is smaller in Test 3.7 than in Test 2.2, the contact stress and number of stress cycles on the uncoated counter face are larger. Additionally, micropitting did not occur on the uncoated disc in Test 3.7 (Fig. 10), while early stage micropitting occurred on the uncoated roller in Test 2.2 (Fig. 4b). Test results of the WC/a-C:H coating on smooth surfaces indicate that although the presence of the coating on the discs promotes the early onset of micropitting on the roller (before about 1 million cycles), when the coating is applied to the roller, no micropitting or high rate of wear occurs on either the roller or the discs – even after 45 million or more cycles. Each disc experiences one stress cycle compared to an equivalent 13.5 stress cycles of the roller, so asperities on the roller are subjected to more stress cycles than those on the disc surfaces. In boundary lubrication (0.1 o λ o0.2), the convective heat flow by the oil is negligible compared to the conduction of heat by the oil and the contacting surfaces. Since the thermal conductivity of hydrogen-containing DLC ( 1 W/mK [32]) is considerably smaller than 52100 steel ( 42 W/mK [33]), it is expected that most of the frictional heat generated in the contact will flow through the steel counter face when it is in contact with a DLC coated surface. Although the frictional heat generation in the contact is transient and of very short duration, it could still promote a rapid diffusion of carbon from the martensite matrix and the creation of a heterogeneous surface region containing ferrite and cementite or other FeC phases such as that observed and reported by Oila and Bull [10,11]. The time interval between thermal shocks is at least an order of magnitude shorter for an element of area on the roller than an element on the disc. Local temperatures on the surface or near-surface region of the rollers should be higher than those on the discs in these experiments. In addition to the time interval between the thermal shocks, the respective volumes of the disc and roller have an impact on the amount of heat that can be absorbed and dissipated through the components. During operation, two of the three discs are partially submerged in the oil sump, but the roller that is in the middle of
three discs is more thermally isolated. Therefore, when discs are coated with the DLC, the amount of frictional heat that flows into the uncoated roller should be larger than the amount of heat that is dissipated in uncoated discs when the roller is coated with the DLC. This qualitative thermal analysis is consistent with the observations that DLC coated discs cause an earlier onset of micropitting on an uncoated roller than when uncoated discs are tested against uncoated rollers with the same test parameters. Conversely, when the coating is applied to the roller, the discs do not exhibit micropitting but can experience an amount of abrasive wear that is consistent with the surface roughness of and hardness difference between the coated roller and the uncoated 52100 steel.
4. Conclusions The objective of this study was to evaluate the ability of a WC/ a-C:H coating to mitigate micropitting and wear damage on mated through-hardened 52100 bearing steel surfaces. The performance of the WC/a-C:H coating was evaluated in boundary lubrication at various λ ratios, slide-to-roll (SRR) ratios and steel hardness values. To determine the origin of micropitting in experiments of this type, it is desirable to understand the evolution of the surface texture during the testing. Although it was not the objective of this study to determine the origin of micropitting in bearing steel, it is certainly worthwhile to pursue that objective in future study. The major conclusions drawn from this study are: 1. Abrasive wear was the dominant surface damage mechanism whenever at least one of contacts included a coated rough surface (Rq ¼0.56 μm). 2. The WC/a-C:H coating experienced a very high rate of wear during testing when it was applied to rough (Rq ¼0.56 μm) substrates, but virtually no wear when the coating was applied to smoother (Rq ¼0.26 μm) substrates. 3. Under boundary lubrication, surface damage of rollers was primarily wear for λ o0.1 and micropitting for λ 40.1. Micropitting of rollers was observed when pmax was comparable to or greater than the estimated yield strength of the 52100 steel, 4. Whereas wear of the roller occurred when the SRR ¼ þ10%, micropitting was the surface damage mode on the roller when the slide-to-roll ratio was 10%. The effect was attributed to whether tensile shear stresses resided in the inlet or outlet positions of the lubricated contact, and was found to be consistent with the observation that micropitting of gear teeth occurs in the region of negative rather than positive sliding. 5. The WC/a-C:H coating eliminated the occurrence of roller wear and micropitting when the coating was applied to the roller. However, when the coating was applied to the discs, the onset of micropitting on the uncoated roller occurred sooner than on rollers paired with uncoated discs. A possible explanation for
Fig. 10. Optical microscopy images of the surface of the uncoated disc from Test 3.7 after running 45 million cycles at 1.7/2/2.25 GPa contact stress against a WC/a-C:H coated roller with hardness of 62 HRC and roughness of Rq ¼ 0.29 μm. (a) Magnification 50 . (b) Magnification 200x.
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this behavior was provided. The thermal conductivity difference between the coating and the 52100 steel allows more heat to flow into the uncoated counter face than the coated one. When the discs are coated, most of the frictional heat flows into the roller, increasing the diffusion rate of C in the martensite matrix and driving the phase changes that can give rise to micropitting. When the roller is coated, most of the frictional heat flows into the three discs that are not as thermally isolated and have considerably more mass than the roller.
Acknowledgments The authors are grateful to The Timken Company for their generosity in funding this study. We wish to thank Dr. Luc Houpert and Dr. Young Kang of The Timken Company, and Prof. Yalin Dong of the University of Akron, for beneficial discussions. We are indebted to Mr. R. Fowler and Dr. B. Tury of the University of Akron for their valuable assistance in this project.
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