The contact characteristics of rough surfaces in line contact during running-in process

Wear 253 (2002) 899–913

The contact characteristics of rough surfaces in line contact during running-in process Jeng-Haur Horng∗ , Mu-Long Len, Jian-Shing Lee Department of Power Mechanical Engineering, National Huwei Institute of Technology, Huwei, Yunlin 632, Taiwan, ROC Received 29 May 2001; received in revised form 4 February 2002; accepted 4 February 2002

Abstract The contact parameters between mating surfaces during running-in process were investigated using a ring-on-block line contact device. The experimental and theoretical results indicate that both contact width and real contact area substantially increased during the running-in process. This increase in contact width might be the main reason that the line contact running-in process increases tribological performance between relative surfaces. During the process, the real contact area was multiplied and its variation can be evaluated by dimensionless correlation distance only for relatively close contact under lubrication conditions. The true friction power intensity (TFPI) is defined as the ratio between the friction power and the real contact area for two relative surfaces. Regardless of loads and roughness values, TFPI index value increases to a maximum with time and then decreases to a low stable value during the running-in process. The initial running-in stage is the key stage determining whether or not the run-in will fail for constant loading condition based on the variation of TFPI index. In general, the running-in for temperature needs a longer time to reach a stable condition than the running-in for wear. The results of skewness and fluid retention index show that the running-in process transforms the original surface into a beneficial surface having many small valleys to accommodate lubrication oil and protect relative surfaces, so that the oil-bearing capability of total valley volume per unit sampling area remains the same. © 2002 Elsevier Science B.V. All rights reserved. Keywords: Running-in process; Friction coefficient; Real contact area; Flash temperature; TFPI index

1. Introduction The performance of machine elements in relative sliding is mainly affected by the properties of the motion interface. If the interface can be maintained under controlled conditions, then the machine can achieve the required performance. The running-in process is defined as the initial operating procedure, which enables machines to improve the surface profile and frictional compatibility [1]. Although the running-in process is the initial process before cylinders, gears and the other new machines are put into operation on a regular basis, the reason for the running-in effect remains unknown. It is unclear, how the microgeometry changes to improve performance and prolong the operating life of machines. Past research reveals that obtaining longer life for engines relies on a suitable run-in process. Moreover, surface roughness is the main factor that affects the run-in if there are no apparent defects or surface waviness from the manufacturing [2]. Duck’s [3] experimental results suggest that the best valley to peak height for diesel engine cylinders should be ∗ Corresponding author. Fax: +886-5-631-2110. E-mail address: [email protected] (J.-H. Horng).

Rt = 4–7 ␮m. He also suggests that gasoline engine cylinder will have the best run-in performance when valley to peak height is Rt = 2.5 ␮m. The sliding line contact rolling experiment by Kang and Ludema [4] indicates that the initial average value should be Ra = 0.1 ␮m in order to improve the formation of an oxide protective film in the run-in process and reduce the low cycle fatigue wear. A two-disk experiment was performed by Kelly and Critchlow [5] to study the relationship between run-in process and scuffing resistance. It indicated that the reason that the running-in can strengthen scuffing resistance is that the roughness decreases and the specific film thickness increases. However, this result contradicted the phenomena that a smoother surface has a poor effect on scuffing resistance. Malburg and Raja [6] suggested using the roughness curve height, the valley distribution and bearing curve as the standard for estimating cylinder wear. Pawlus [7] using a real engine to conduct his experiment, indicated that under the same roughness height, the wear of a cylinder is proportional to the emptiness coefficient Rp /Rt (Rp is defined as the maximum profile peak height). He also indicated that the higher oil storage ability in the higher peaks could promote an engine’s performance. The two-disk fatigue experiment also revealed that the larger

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the negative value on skewness (RSk ), the longer the operating life [8]. Horng’s scuffing experiment [9] explained that the two main factors for scuffing resistance are the real contact area and valley’s oil-bearing capability. The surface parameters can be characterized by the rms roughness Rq of the height distribution and the correlation distance β which describes the spatial structure of a surface [10]. In the experimental works by Hirst and Hollander [11] and by Poon [12], the surface damage was found to be related to the surface parameters Rq and β. When two surfaces come into sliding contact, the friction heating produced at the interface can cause rapid temperature rises. The temperature rises may cause surface oxidation, peak deformation and thermomechanical wear during running process. Earlier temperature analyses were concentrated on the maximum and average temperature rises. Blok [13] was probably the first to model the friction-induced temperature in a contact between two machine element. Jaeger [14] and Carslaw and Jaeger [15] extended Blok’s model to obtain temperature distribution for different ranges of the sliding speed. The later works released some limitations to suit more real conditions [16–18]. Kuhlmann-Wilsdorf [16] developed approximate expressions of average temperature rise for a single heat source in the low, intermediate and high Peclet number regions. Bos and Moes [18] developed a numerical algorithm to solve the steady state heat partitioning and the associated flash temperatures for arbitrary shaped contacts by matching the surface temperatures of the two contacting solids at all points inside the contact area. Friction power [19] and critical contact temperature [13] had been provided as an explanation of the failure of lubricant layers. Horng [9] experimentally and theoretically studied the relations of true friction power intensity (TFPI) with surface roughness, sliding speed and scuffing. In these analyses, the TFPI index, the scuffing and running-in effects are all dependent upon each other. The above temperature and friction energy parameters at interfaces will be examined during the running-in process. In order to reach a balance among the surfaces, it is necessary to stabilize properties of friction and wear during contact. There must be surface contact in running-in process, so that any tiny changes on surface contact will become an important factor. In order to understand how changes on the contact surfaces during the lubricated wear, this study utilizes the line contact experiment to discuss the relationship and changes among the skewness, contact temperature, TFPI index, correlation distance, oil-bearing index, real contact area and contact width during the running-in process.

2. Theory 2.1. Skewness (Rsk ) Skewness, Rsk , is the measure of asymmetry of surface deviation about a mean plane. It is defined as

Rsk

1 = 3 Rq




∞ −∞

z3 p(z) dz

(1)

where z is the profile coordinate and p(z) is probability density function. This parameter can give some indication of the existence of spiky features. A comparatively large negative skewness may come from such surfaces which have a few significant valley (pits, troughs) and good bearing property. Conversely, a comparatively large positive skewness may indicate the presence of comparatively few spikes on the surface which could quickly wear away in contact with another surface. 2.2. Correlation distance (β) The correlation distance, β, is a measure of the decay rate of the profile and is defined as the distance required for autocorrelation function to decay to a value of 1/e. 2.3. Retention index (Svi , Sci and St ) [20] Valley fluid retention index is a parameter used to indicate the fluid retention property in the valley zone. It is defined as the ratio of the void volume per unit sampling area at the valley zone over the rms deviation, i.e. Svi =

Vv (h0.8 ) Rq × sampling area

(2)

where h0.8 is the normalised height at the 80% bearing area in bearing area ratio curves and Vv (h0.8 ) denotes the void volume under the dimensionless height of 0.8. A larger Svi indicates a good fluid retention capability feature. For a Gaussian surface this index is about 0.11. Core fluid retention index is a parameter used to indicate the fluid retention property in the core zone. It is defined as the ratio of the void volume per unit sampling area at the core zone over the rms deviation, i.e. (Vv (h0.05 ) − Vv (h0.8 )) Sci = (3) Rq × sampling area A larger Sci indicates a good fluid retention property in the core zone. For a Gaussian surface this index is about 1.56. The total fluid retention index (St ) is the sum of the values of core fluid retention index and valley fluid retention index. 2.4. Real contact area (At ) When two surfaces are brought together, surface roughness causes contact to occur at discrete contact spots. The real area of contact is generally much less than the apparent area. The real area of contact of microscopically rough surfaces is an important parameter in the understanding of many tribological phenomena, such as wear, adhesion, friction force, frictional heating, thermal and electrical contact resistance and fluid leakage. An advanced elliptic elastic–plastic microcontact model that takes into account the directional

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nature of surface roughness was proposed by Horng [21] for elliptic contact spots between anisotropic rough surfaces. According to the elliptic elastic–plastic asperity microcontact model, the real contact area can be written as

2.5. FP and TFPI index


At (d) = ηAn πRm f1 (e)

FP = fWVs

d+ωc d


+ ηAn πRm f3 (e)

(z − d)φ(z) dz

∞ d+ωc

The friction power (FP) is defined as the friction heat input per unit time into the contact area and is given by (10)

where Vs is the sliding velocity, W the applied load and f is the friction coefficient. The TFPI gives the definition in terms of the real contact area (At ) and then

[2(z − d)

− ωc (2 − f4 (e))]φ(z) dz

(4)


d+ωc 4 0.5 F (d) = ηAn E  Rm f2 (e) (z − d)1.5 φ(z) dz 3 d
∞ + ηAn πRm P0 f3 (e) [2(z − d) d+ωc

− ωc (2 − f4 (e))]φ(z) dz

901

(5)

TFPI =

FP At

(11)

The TFPI index is represented as TFPI× (Vs )−1.2 (σ )−0.6 [9]. The lower the TFPI index value, the higher the scuffing resistance of the lubricated system. The run-in will fail when the maximum TFPI index value exceeds the critical value. It not only acts as a scuffing criterion but also can be considered as a parameter for evaluating the running-in process.

where f1 (e) = f2 (e) =

E(e) K(e)(1 − e2 )0.5 π E(e)0.5 2K(e)1.5 (1 − e2 )0.5

(6)

(7)

f3 (e) =

E(e)e2 2(1 − e2 )0.5 [E(e) − K(e)(1 − e2 )]

(8)

f4 (e) =

f1 (e) f3 (e)

(9)

where the parameters can be obtained according to Horng [21].

3. Test apparatus and operating variables The experiments were performed on a disk-on-block test rig as shown in schematic representation of Fig. 1. The lower block is stationary while the upper disk rotates at a constant speed. All the test specimens were made of JIS SCM 415 steel. The disks, case-hardened to approximately 60 HRc, were 75.0 mm in diameter and 13.4 mm in total width. The blocks, 12.4 mm × 12.4 mm × 12.4 mm in dimension were heat treated like the disks and completely immersed in an oil cup with a volume of 70 ml. The mechanical properties of test specimen were given in Table 1. The lubricant was a plain mineral oil with physical properties as given in Table 2. These data in tables of test specimens were used to calculate the contact parameters and real contact area ratio.

Fig. 1. Experimental apparatus.

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Table 1 Mechanical properties of the test specimen Property

Values

Poisson ratio Elastic modulus (GPa) Thermal conductivity (W/(m ◦ C)) Specific heat (J/(kg ◦ C)) Hardness (HRc)

0.3 230 48 460 60

contact zones. The oil temperature was maintained at 55 ◦ C by means of a heater in this experiment and the accuracy of temperature measurements using the thermocouple was ±1.5 ◦ C.

4. Results and discussion 4.1. Characterization of wear and friction

The disk drive power consisted of a 5 hp motor driven by a continuously variable speed controller. The experimental conditions were electronically monitored by a digital instrumentation system which included the speed of rotation, oil and specimen temperatures, applied load, friction force, specimen displacement and electrical contact resistance. A levered bar connected to the specimen support was utilized to measure the wear depth of the bottom specimen. In order to understand the lubricating condition between rotating disk and lower block, the test specimens were electrically isolated from the rest of the equipment. The electric circuit was connected with the constant electric resistance and a Wheatstone bridge to constitute an instrument. The contact resistance between the rotating disk and fixed block is transferred through amplifier to a recorder. In this experiment, a normal load is applied to the contact surface between the specimens by an air pressure loading system working on the bottom of the block. The applied load be varied from 200 to 1000 N which is equivalent to nominal Hertzian stress from 115 to 257 MPa. The operating sliding speed was 200 rpm (0.785 m/s). Measurements of the disk surface roughness was done with a portable surfcorder. The surfcorder was connected to a personal computer through an A/D converter for data processing. The upper disk used in the experiments had surface roughness values of Ra = 0.04–0.06 ␮m (measured across the lay). The mean roughness height of the block fell into three categories: 0.20 ␮m (0.15–0.21 ␮m) for smooth surfaces, 0.60 ␮m (0.55–0.65 ␮m) for medium surfaces and 1.10 ␮m (1.00–1.25 ␮m) for rough surfaces. The rotational direction was parallel to the surface roughness direction between

Table 2 Physical properties of the test lubricant Property

Values

Density (15 ◦ C, g/ml) Flash point (◦ C)

0.881 240

Kinematic viscosity (cSt) 40 ◦ C 100 ◦ C

62.1 8.31

Viscosity index Thermal conductivity (W/(m ◦ C)) Pressure–viscosity coefficient (GPa−1 )

98 0.145 2.37

According to the mechanism of the ring-on-block test rig, the upper specimen is a rotating disk and the lower block is a fixed square. As a result, the wear on the lower specimen is more severe than the upper specimen. To compare the variations on surface roughness of both specimens before and after the running-in process, Fig. 2 indicates the average roughness value during the running-in process for the initial roughness value of 1.10 ␮m with different loads. As we can see, the surface of the upper specimen is hardly worn. However, under the same condition, surface roughness value of the lower specimen after 5 min has changed dramatically. The results show that the wear effect has been mostly completed at the initial stage of the running-in process. The rotating condition in Fig. 2 is under the most difficult situation on contact area in all tests, but the severity of its wear between upper and lower samples is totally different. As a result, we can see that in this experiment the wear occurs mostly on the lower specimen. Therefore, the characteristic changes in surface roughness can be observed by the roughness changes on the lower specimen during the overall running-in process. Fig. 3(a) shows the variation of friction coefficient, electrical contact resistance and wear depth with a load of 400 N and average roughness value of 0.20 ␮m for the running-in process. No experimental data for friction coefficient is presented at the time of 0 min because the reading is unstable. The friction coefficient declined rapidly at the initial stage, then gradually decreased to a stable value. The initial wear depth is around 2.5 ␮m and then slowly increases to 10 ␮m. This reveals that at the initial stage of running-in process surface peaks touch and wear intensely and then the interaction effects of mating surfaces decrease gradually. In terms of the characteristic changes of the wear and friction in the overall running-in process, the process indeed causes the increase of the wear and frictional compatibility. According to the decrease on wear and friction in Fig. 3(a) and the variation of the roughness value in Fig. 2, the flattness of roughness peaks and then the decrease of contact pressure is considered as the one of the reasons for the increase of contact resistance. These results could be explained by the variation of electrical contact resistance during the running-in process. The electrical contact resistance value increases gradually in Fig. 3(a). The higher the value of electrical contact resistance, the larger the separation between contacting surfaces. The low initial resistance means that the relative surfaces begin to contact each other [22]. After the initial stage,

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Fig. 2. Variation of average roughness value of both specimens at initial roughness value of 1.10 ␮m with different applied loads.

the contact resistance value increases gradually, indicating that an oil film is gradually created between interfaces. Therefore, the flattness of roughness peaks and the creation of the oil film occur simultaneously during the running-in process. The variation of friction coefficient, electrical contact resistance and wear depth at average roughness value of 0.20 ␮m for the loads of 200 and 1000 N during running-in process are shown in Fig. 3(b) and (c). From Fig. 3(a–c), the initial wear depth is in the order of: (8.3 ␮m)1000 N > (2.5 ␮m)400 N > (0.3 ␮m)200 N . The sequence of wear depth is similar to the initial specific film thickness: (1.03 ␮m)1000 N > (1.15 ␮m)400 N > (1.29 ␮m)200 N The total wear depths at the roughness value of 0.2 ␮m for different loads are (14.7 ␮m)1000 N , (8.5 ␮m)400 N and (2.5 ␮m)200 N . These results show that the percentage of initial wear to total wear increases with increases in the applied load. The contact resistance curves, as described in Fig. 3(a–c), show that the lubrication regimes change from boundary lubrication to mixed lubrication for three cases. However, for the load of 200 N, boundary lubrication only occurs at the starting moment, whereas for the load of 1000 N, boundary lubrication can extend to test time of 30 min. The friction coefficient curves, as described in Fig. 3(a–c), show that the higher the load, the faster decreases will reach a stable value during the operating time. It is significant to note that the value of friction coefficient not only depends on the initial operating conditions, but

also depends on the running-in process. Therefore, the final friction coefficients do not have a similar trend to the specific film thickness. Fig. 3(d) shows the variation of friction coefficient and electrical contact resistance for the loads of 400 N when the surface roughness value increases from 0.2 to 1.10 ␮m. A comparison with the results of Fig. 3(a–c) reveals that the initial wear and total wear in Fig. 3(d) are the largest among the four operating conditions. This is because the wearing process is so severe that the friction coefficient decreases progressively during the overall running-in process. Although the time for boundary lubrication in Fig. 3(c) and (d) is almost equal, the running-in effect is clearly different. From the variation of contact resistance in Fig. 3(c) and (d), the separation between opposing surfaces is always small at high load and low roughness values. For the medium load and high roughness value in Fig. 3(d), the contact resistance increases sharply after boundary lubrication. But at the same time the friction coefficient still decreases and wearing occurs progressively. These results indicate that high surface roughness value may be easy to create an oil film, even though wearing occurs. In the past, it was generally believed that the decrease and stability of the wear and friction, as well as the increase of the contact resistance were caused by gradually flattening roughness peaks so that the contact pressure was decreased and an oil film was created between interfaces. Below, we discuss if this is the true in line contact condition.

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Fig. 3. Variation of friction coefficient, contact resistance and wear depth during the running-in process with different operating conditions. (a) Load = 400 N, Ra = 0.20 ␮m; (b) load = 200 N, Ra = 0.20 ␮m; (c) load = 1000 N, Ra = 0.20 ␮m; (d) load = 400 N, Ra = 1.10 ␮m.

4.2. Characterization of contact width and real contact area The effects of the load on contact width during the running-in process are shown in Fig. 4(a). The data of wear contact width in the figure represent the average value of three readings measured at different positions on the contact line. The figure shows that no matter how large the load is, the contact width rapidly increases at the initial stage of the

running-in process and then becomes stable. The higher the loading, the more contact width there is. Meanwhile, when the applied load is 1000 N, the contact width in the late period still slowly increases. That means in overall running-in process, in addition to the flattening of the peak to cause the real area increase, the apparent area also increases, causing the apparent area pressure to decrease gradually. From Fig. 4(a), the inaccuracy rate of the calculated value by Hertz theory and the value of measuring its real surface

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Fig. 3. (Continued ).

are nearly 300% at the initial stage of running-in process and the inaccuracy rate at the end of the process is 300% greater. In general for tribological design and manufacture, we utilize the Hertz width to calculate and design. However, this method should be modified since it is obvious that the Hertz width used to calculate the contact area is too conservative. Fig. 4(b) shows the effect of surface roughness on contact width during the running-in process, indicating that different roughness values make the contact width different. However, the effect of roughness value on contact width is smaller than that of applied load on contact width.

Real contact area plays a significant role in friction, wear and lubrication in machine components. The applied external loads in contacts are actually carried by the highest peaks of the asperities. The variation of dimensionless correlation distance β/Rq and real contact area ratio At /An at the low loading, as well as small roughness during the running-in process are shown in Fig. 5(a). The value of the theoretical At /An is calculated from Eqs. (4)–(9). It increases gradually, although the variation is too small to show up on the figure. The value of the theoretical At /An multiplied by the ratio between wear contact width and Hertz width equals the value

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Fig. 4. (a) Effects of load on the contact width during the running-in process. (b) Effects of surface roughness on the contact width during the running-in process.

of the experimental At /An . The experimental At /An reveals that the real contact area ratio increases three times rapidly during the running-in process and reaches a stable value at the middle stage of the running-in process. The roughness peak flattening causes the increase of theoretical At /An . Both roughness peak flattening and the wear scars cause

the increase of the experimental At /An , which indicates that the decrease of the real contact pressure of the running-in process is due to the wear scar area. This is why the contact resistance increases in the running-in process and it can effectively increase tribological performance. This indicates that during the line contact running-in process, the key

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Fig. 5. Variation of real contact area ratio and dimensionless correlation distance in the running-in process with different operating conditions. (a) Load = 200 N, Ra = 0.20 ␮m; (b) load = 1000 N, Ra = 0.20 ␮m; (c) load = 200 N, Ra = 1.10 ␮m; (d) load = 1000 N, Ra = 1.10 ␮m.

factor affecting the contact properties is contact width and the flattening effect of the roughness peak is a subordinate factor. The variation of At /An ratio in Fig. 5(b), at the load of 1000 N, also shows a similar trend to that in Fig. 5(a). Theoretical study [23] found that the real contact area ratio and the dimensionless correlation distance β/Rq are dependent on each other. These two figures show that the β/Rq value increases gradually during the running-in process. However, the variation of values of β/Rq and experimental At /An are more consistent for high loads than for low loads. Surface contact is much closer together when the roughness value is 1.10 ␮m, as shown in Fig. 5(c) and (d). The variation of At /An in Fig. 5(c) and (d) is similar to that in Fig. 5(a) and (b) under the roughness value of 0.20 ␮m. However, the initial and final values of At /An are both smaller due to high roughness value. The value of β/Rq is larger for the load of 1000 N than for the load of 200 N. At the same time, the tendency of β/Rq is also similar to Fig. 5(a) and (b). Comparing Fig. 5(c) at load = 200 N and Ra = 1.10 ␮m with Fig. 5(a), the value of β/Rq more quickly reaches a stable value during the running-in process because the relative surfaces of 5(c) are much closer together than those of Fig. 5(a). The

variation of β/Rq in Fig. 5(b) and (d) also show the same trend. According the above results, it can be seen that the closer the relative surfaces, the more consistency there is between the trend of At /An and β/Rq and the value of At /An become stable more rapidly than the value of β/Rq . Furthermore, the value β/Rq increases with the increase in applied load in the running-in process, indicating that the stable value of At /An in the running-in process dose not depend on the value of β/Rq . The larger load makes the value of At /An and β/Rq stabilize rapidly. Regardless of the size of the applied load, the smaller the roughness value, the larger will be the initial and final At /An during the running-in process. 4.3. Characterization of friction heating and scuffing The flash temperature and friction power between contacting surfaces with the average roughness value of 0.20 ␮m at the load of 200 N are presented in Fig. 6(a). The contact condition for the load of 200 N is observed from Fig. 3(b) to have relatively mild wear and high contact resistance. Thus, it shows that the flash temperature is not more than 7 ◦ C and the deviation is very small during the whole process.

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Fig. 6. (a) Variation of flash temperature and friction power between contacting surfaces with the average roughness value of 0.20 ␮m at the loads of 200 and 1000 N during the running-in process. (b) Variation of flash temperature and friction power between contacting surfaces with the average roughness value of 1.10 ␮m at the loads of 200 and 1000 N during the running-in process.

This indicates that the temperature effect on the running-in process is extremely small in this condition. When the contacting load is added to 1000 N, the initial flash temperature at interfaces is about 43 ◦ C and then decreases to a stable value, as shown in Fig. 6(a). The contact temperature, which is the bulk temperature plus the flash temperature, is almost 100 ◦ C at the initial running-in stage. This high temperature is enough to affect the lubricant viscosity, film reaction, wear and material property. The decrease of contact temperature leads to the increase of lubricant viscosity and hence hydrodynamic lubrication occurs between mating surfaces.

As described in Fig. 3(c), the contact electrical resistance increases gradually after 30 min due to the decrease of contact temperature. Regardless of applied loads, the variation trend of flash temperature and friction power is similar at the roughness value of 0.20 ␮m during running-in process, as shown in Fig. 6(a). Fig. 6(b) illustrates the effects of applied loads on the flash temperature and friction power at the roughness value of 1.10 ␮m during the running-in process. It also shows that a high load gives rise to an increase in flash temperature and friction power. The variation trend of flash temperature and

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Fig. 7. (a) Variation of TFPI index between contacting surfaces with the roughness value of 0.20 ␮m for different loads during the running-in process. (b) Variation of TFPI index between contacting surfaces with the load of 1000 N for different roughness values during the running-in process.

friction power from observing Fig. 6(a) and (b) are similar in the different load and roughness values. The maximum flash temperature is 45.5 ◦ C at mating surfaces with the roughness value of 1.10 ␮m and the load of 1000 N in Fig. 6(b). These results indicate that in terms of flash temperature and friction power the effect of load is larger than the effect of roughness value in our experiments. Fig. 6(b) shows that

flash temperature and friction power arrive at a stable values after 70 min. However, the surface with low roughness value of 0.20 ␮m arrives at stable values only after 30 min, as shown in Fig. 6(a). This indicates that severe contact condition delays the running-in time. Comparing the running-in time in Fig. 6 with that in Figs. 2–5, the running-in times of individual contact characteristics, including friction, wear,

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Fig. 8. (a) Variation of skewness with different surface roughness values under the load of 200 N during the running-in process. (b) Variation of skewness with different load values under the average roughness value of 0.60 ␮m during the running-in process.

temperature and film thickness are different. In general, wear needs a shorter running-in time to reach a stable condition compared to temperature. One of the goals of the running-in process is to increase the scuffing resistance between mating surfaces. The TFPI index value can be considered as a performance parameter that represents both the scuffing failure resistance and the effectiveness index of the running-in process. The effects of applied loads and roughness values on TFPI index are

presented in Fig. 7(a) and (b). Regardless of loads and roughness values, the TFPI index increases to a maximum with time at the initial stage and then decreases a low stable value, indicating that run-in can effectively raise the scuffing resistance for all operating conditions. The initial stage is the key point whether or not the TFPI index exceeds the critical value. The maximum deviation occurs at the load of 1000 N and the roughness value of 1.10 ␮m at the initial stage, whereas at the same stage, the minimum deviation

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Fig. 9. Variation of valley fluid retention index, core fluid retention index and total fluid retention index during the running-in process with different operating conditions. (a) Load = 200 N, Ra = 0.20 ␮m; (b) load = 1000 N, Ra = 0.20 ␮m; (c) load = 200 N, Ra = 1.10 ␮m; (d) load = 1000 N, Ra = 1.10 ␮m.

occurs at the load of 2000 N and the roughness value of 0.20 ␮m. This result suggests that a relatively severe contact condition is not suitable for the initial running-in process. 4.4. Characterization of fluid retention parameters The valleys of surface roughness may serve as lubricant reservoirs that promote the formation of a protective film on the real contact area. The oxidation film or extreme pressure film both contribute significant benefit for the interface friction, wear and scuffing. Fig. 8(a) shows variation of skewness with different surface roughness under a load of 200 N during running-in process, indicating that three lines decrease at the initial running-in stage and then slightly rise to a stable value. Compare the Fig. 8(a) with (b), a similar trend is seen. All the final values of skewness differ from −0.6 to −1.1 and below the initial value. These results mean that different running-in procedures all make the surface transform into a beneficial surface with many small valleys to accommodate lubrication oil and protect relative surfaces. However, under the condition of large loading and roughness value (Ra = 1.10 ␮m, load = 200 N or Ra = 0.60 ␮m,

load = 1000 N), the value of skewness has the tendency to decline rapidly in the initial running stage. It might be that at the initial running-in stage, except for flattening, peaks are being broken in order to establish many small valleys so that the skewness becomes smaller. However, in later periods of the process, small cracks might gradually flatten and make the value of skewness rise slightly. These cracks would not appear for small loads or smooth surface at the initial running-in stage, so that the initial value of skewness decreases slightly. Other parameters describing oil-bearing properties of the surface are valley fluid retention index (Svi ), core fluid retention index (Sci ) and total fluid retention index (St ). For the load value of 200 N and roughness value of 0.20 ␮m, Fig. 9(a) shows that Svi increases at the initial running-in stage and then declines to a stable value. This slightly different value reveals that the effect of Svi on the running-in process is small. The core fluid retention index (Sci ) and total fluid retention index (St ) both decline in the running-in process and then rise slowly. At the same time, the slight difference between the initial and final values of St indicates that the running-in procedure has negligible effect

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on oil-bearing parameters. When the loading changes into 1000 N or the roughness value changes to 1.10 ␮m, as indicated in Fig. 9(b–d), the variation of Svi , Sci and St during the running-in process is as the same as the tendency of Fig. 9(a). However, when the value of Ra changes from 0.20 to 1.10 ␮m, the time must be decreased to the minimum value of Sci and St is clearly delayed. After comparing the changes of skewness in Fig. 8 with the oil-bearing index in Fig. 9, it can be concluded that the running-in process is to change large valleys to many small valleys. Thus, the total quantity of fluid retention as void volume per unit sampling area did not change much.

initial value, which means that the total oil-bearing quantity in each valley volume during the process did not change much.

Acknowledgements The author gratefully acknowledges the National Science Council of Taiwan, ROC, which supported this research under Grant NSC 89-2212-150-E-034. The personal support of J.L. Chen, J.C. Lai and Z.T. Lin is also appreciated. References

5. Conclusions This study uses the line contact oil bath mechanism to conduct a running-in operating experiment under different loads and roughness values. It also uses a profile curve to calculate contact performance parameters. The wear scar shows a pronounced increase in contact width and the contact width is much larger than the Hertz width. Overall, it is mainly affected by the applied loads, regardless of the roughness value. This may be main one of the reasons that promotes lubricant performance effectively during the line contact running-in process. The dimensionless correlation distance and the real area ratio tended to have the same trend of variation under relatively severe contact condition during the running-in process. The value of At /An gives stable value more quickly than did β/Rq . When the load is larger, the value of β/Rq will increase more quickly, which means that the real contact area cannot be determined solely by the value of β/Rq . The real contact area value that will multiply during the line contact running-in process is determined mainly by the wear contact width. The variation trend of flash temperature and friction power is similar among the different loads and roughness values. In general, the run-in of temperature needs a longer time to reach a stable condition than that of wear. In order to avoid the damage of scuffing, the maximum value of TFPE index should be less than the critical value of TFPE index of lubricated sliding surfaces. Regardless of loads and roughness values, the TFPI index increases to a maximum with time at the initial stage and then decreases a low stable value during the running-in process. Therefore, the initial running-in stage is the key stage determining whether or not the TFPI index exceeds the critical value. No matter how large the roughness or loads are, they will always cause the skewness to decline after the running-in process. This indicates that the running-in surface contains more small valleys to bear lubricant, thus benefiting the formation of the reaction film among contact areas so that film can protect the contact surface. The core fluid retention index and the total fluid retention index during the running-in process first declined and then increased slowly back to the

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Jeng-Haur Horng graduated in naval architecture and marine engineering in 1981, at the National Cheng-Kung University and received his PhD degree in mechanical engineering at the same university. He is head of the department of power mechanical engineering and professor at the National Huwei Institute of Technology. His research interests are mainly in the area of microtribology, especially the effect of surface roughness on tribology, contact temperature, contact mechanics and industrial applications.