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Effect of surface roughness on steel roller scuffing
WEAR Wear 184 (1995) 203-212
Effect of surface roughness on steel roller scuffing Jeng Haur Homg a, Jen Fin Lin bT*,Ke Yang Li b aDepartmenf ofPower Mechanical Engineering, National Yunlin Polytechnic Institute, Y&in 70101, Taiwan ’Department of Mechanical Engineering, National Cheng Kung University, Tainan 70101, Taiwan
Received 18 May 1994; accepted 1 December 1994
Abstract Experiments were carried out utilizing a gear-cam adapter to simulate line-contact lubrication and wear. Roller specimens with various asperity heights and roughness patterns were riding on roller plates and sliding over the two lands of the lower specimen. The scuffing loads for the rollers with various roughness conditions were measured while changing the applied load and spindle’s rotational speed. The effects of the roughness parameters on scuffing load, particularly roughness pattern and asperity height, were thus investigated. The displacement of the bottom specimens during the lubrication and wear process were measured to establish their correlation with scuffing behaviour. The variations of bottom specimen displacement can offer information regarding the film thickness formation on the rollers with various roughness patterns and asperity heights. Keywords: Roughness; Steel; Scuffing; Rollers
1. Introduction
Scuffing is an important form of damage that results from lubricant film failure in sliding and rolling contacts, such as for gears, piston rings and cams. When scuffing occurs, there is a sudden rise in frictional force, accompanied with noise and vibration; the damaged surface shows severe wear and plastic deformation. Usually, scuffing damage is so ruinous that the damaged part must be replaced. However up to now, the complete, detailed mechanism of scuffing has not been stated, nor is there any commonly accepted criterion from which the possibility of scuffing may be predicted in advance. It is believed that scuffing behaviour is a complex function of mechanical and chemical effects; the factors that govern scuffing include lubricant viscosity, oil temperature, material hardness, surface treatment, geometry profile, surface roughness, sliding/rolling speed and chemical activity within the working environment. Many studies investigated the effects of several influential factors on scuffing failure. Blok [11 suggested that scuffing would occur if the surface temperature reached a critical value. The critical surface temperature consists of two components: first, the bulk temperature of metal parts; second, the instantaneous temperature increase of the surfaces as they pass through the contact zone. Although the critical contact * Correspondingauthor. 0043.1648/95/$09.50 0 1995 Elsevier Science S.A. All rights reserved SSDIOO43-1648(94)06574-8
temperature is the most widely used criterion of scuffing, many investigators had found that the total contact temperature just before scuffing is in fact not a constant for a given combination of lubricant and materials, but varies with the experimental conditions [ 21. Matveevsky [ 31 presented a study, of the oil lubrication used in previous work, on point and line contact machines. It showed that frictional power is a viable scuffing criterion under boundary lubrication conditions. Both sliding and rolling speeds cause important effects on the scuffing conditions. Bell [ 41 and Carper [ 51 conducted experiments to examine the influences of both sliding and rolling speed on the scuffing of lubricated steel discs. They concluded after experimenting with a wider range of rolling speeds, sliding speeds and slide/roll ratios, that (a) an increase in the sliding velocity, at a constant rolling velocity, decreases the scuffing load and that, conversely, (b) an increase in the rolling velocity, at a constant sliding velocity, increases the scuffing load. Furthermore, Czichos [ 6,7] had made several detailed investigations and constructed a surface failure mode1 for three-dimensional load-velocity-temperature space which determines the limits of thin film lubrication. To improve the scuffing resistance of machine elements, it is generally considered that increasing the surface hardness is good. However, this is not always correct, since the scuffing resistance is significantly affected by a change in surface conditions, especially the surface roughness during the running-in process [ 81.
Jeng HuurHorngetul’ Wecrr 184 (1995) 203-212
204
According to the FZG machine or rotating rolling bench test, the failure load decreases with the decrease in oil viscosity [ 9,101. Scuffing resistance is increased by using high viscosity lubricants with anti-scuffing additives such as sulphur-, phosphorus-, borate ZnDDP- or PAO-base lubricants [ 11,121. Jackson [ 121 hypothesized that the reduced heat generation by low traction lubricants leads to an increase in scuffing load. Oxygen apparently plays an important role in friction and wear. The experiments by Bjerk [ 131, using a gear-roller test machine with steel rollers lubricated by plain mineral oil, indicate that surface protection against scuffing is provided by the formation of oxides in the contact zone. Other important characteristics on scuffing resistance are both the specimen surface preparation and the whole test procedure [ 141. It shows that running-in helps scuffing performance. Kelly [ 151, by using a twin-disc test machine, concluded that the dominant effect of prior running-in, in enhancing the load and temperature that lubricated surfaces can resist without scuffing, is the reduction of surface roughness. Grew [ 161 showed that chemisorption reactions with surface materials can explain the widely-known phenomenon of additive interference and thus help in understanding scuffing and running-in. The interaction between the surface roughness height and scuffing behaviour has already been well recognized. It is believed that scuffing loads increase as surface roughness decreases [ 17,181. However, the effect of roughness pattern on scuffing performances has not been studied much. Experimental results have proved that the surface roughness pattern does have a considerable effect on tribological performance under various lubrication states [ 191. In this study, experiments were carried out by utilizing a gear-cam adapter to simulate line-contact lubrication and wear. Steel rollers with various asperity heights and roughness patterns were tested in oil lubrication, obtaining scuffing load measurements while differing the applied load and the spindle’s rotational speed; hereby investigating the effects of different surface roughness parameters on the roller’s scuffing. The displacements of the bottom specimen during the loading process were also taken to try to establish their correlations with scuffing loads. The variations of displacement can offer information regarding the film thickness formations of the rollers in relation to various roughness patterns and asperity heights.
2. Application
of friction
theory
to gear contacts
The mechanism of two gear teeth contact at any moment can be simulated by two cylinders of different radii being loaded together and separated by a thin film. For convenience, the model of two long cylinders separated by a thin lubricating film is further simplified as a single long cylinder rolling and sliding on a semi-infinite solid with the same gap as in the lubricated region. The test apparatus, illustrated in Fig. 1 (a),
:.r Spindle
Driver
-
n
Roller plate Roller
Lower
1
7
specimen 1
(a)
Fig. 1. (a) Schematic view of the apparatus and arrangement of the principal working parts and (b) diagrammatic representation of the lower specimen. The dimensions are given in mm.
had two rollers, 180” apart, riding on the inner and outer lands of the lower specimen. The driver contacts the two rollers at a nominal position between the inner and the outer lands of the lower specimen, which thus splits the test in half for each roller, and in half again for each land; therefore, 25% of the test load is applied to each land-roller contact. To simplify the analysis, of the rolling-sliding contacts occurring in the working parts of the system, the roller is assumed to be only rolling at all times at point A, centred
Jeng Haur Homg et al. /Wear
Table 1 r, and r,, drive radii and their correspondingcircumferences
Outer land, r, Inner land, r, Drive land
205
Table 2 Physical propertiesof test oil
Radius (cm)
Circumference(cm)
Property
Test oil
1.905 0.714 1.310
11.969 4.469 8.227
Kinematic viscosity,cSt: constant at 40 “C; constant at 100 “C Density at 15 “C, g ml-’ Pour Point, OC Total Acid No.
14; 3.5 0.862 -15 < 0.03
between the two lands (Fig. 1 (b) ) . At that point, the driver rolls 8.22 cm without any slippage (see Table 1). For the inner land, only 4.47 cm is available for rolling at its midpoint (point F) ; so the difference, 3.75 cm, is slippage. In the same manner, the circumference at the midpoint (point C) of the outer land is 11.97 cm and the roller rolls out 8.22 cm, again leaving a 3.75 cm slippage. Although these two slippages have the same absolute value, the outer end of the roller travels a larger distance than the inner end. Therefore, forward sliding occurs on the outer land and backward sliding occurs on the inner land. The friction coefficient of the rolling-sliding contacts can be determined if the applied load and the friction torque are available. The friction torque, according to the detailed analysis demonstrated in Ref. [ 191, can be written as T=O.SW(f,r,
184 (1995) 203-212
-f2r2)
(1)
where W is the applied load;f, andf* are the friction coefficients at the outer and inner lands respectively; r, and r, are the radii of the outer and the inner lands respectively (see Table 1). Since the coefficient of friction on each individual land is not as easy to obtain as suggested in Eq. ( 1), an “average” coefficient of friction f is introduced to replace the “local” coefficient of friction at two lands. Thus, the average friction coefficient is stated as [ 191
f= 1.679TIW
(2)
3. Experimental set-up and operating conditions The line-contact experiments were conducted on a commercial test machine utilizing a gear-cam contact adapter. The working components consisted of a driver, a roller plate and a lower specimen. The lower specimen is stationary when the driver rotates at a constant speed. The assembly of those components in motion is completely immersed in an oil cup with a volume of 100 ml. A long, levered bar connected to the specimen support was utilized to measure the vertical displacement of the bottom specimen. A wear gage was vertically positioned at the end of the lever bar, with the ratio of specimen displacement to the wear gage readings being 1:6. The wear gage was not intended to measure the actual film thickness between two specimens, but was used to obtain successive readings of the bottom specimen displacement. The experimental conditions were electronically monitored by a digital instrumentation system (e.g. the speed of
rotation, the oil temperature, the applied load, the driving torque and the specimen displacement). Measurements of the roller surface roughness and the wear volume were done on a three-dimensional surfcorder. The wear rates of the roller were thus obtained by measuring the volumetric loss. The accuracy of temperature measurements using the thermocouple is + 0.5 “C, for rotating speed measurement +3 rev min-‘. All the roller specimens were made of AISI 8620 steel, with a radius of 0.25 cm and a length of 2 cm. The heat treatments of the specimens followed the requirements described in the Gear Handbook ( 1984). The roller specimens, from the viewpoint of roughness parameters, fell into four classes: ( 1) smooth surface; (2) longitudinally-oriented pattern; (3) transversely-oriented roughness pattern; (4) oblique roughness pattern forming 45” inclined angle with respect to the roller’s axis. The flat plates of the driver and the two lands of the lower specimen were all polished to have mean roughness height of 0.05 ,um. The mean roughness height of the rollers fell into three categories: 0.2 pm for the relatively smooth surface, 0.6 pm for the small roughness, and 1.2 pm for the large roughness. The composite surface roughnesses of lower specimens and rollers include the various gear finishes for gear pairs [ 201. All the drivers and the lower specimens were made of the same material as the rollers. The lubricant was a mineral base oil with the physical properties given in Table 2. The oil temperature before testing was maintained at 70” by means of an electric heater. Then the machine was brought up to the designated rotational speed with no load, and the oil temperature was increased to 8 1 “C for twenty minutes. The first load was then applied to the bottom specimen by an air pressure loading system for 30 s, and then it was released for 3 min in order to lower the oil temperature to 8 1 f 0.5 “C. The sequent loads were imposed step by step following the same procedure until scuffing occurred. The loading schedule is in Table 3.
4. Results and discussion The experimental results of this study, including the scuffing loads, the friction coefficients and the displacements of bottom specimen, are used to evaluate the effects of both a roller’s roughness pattern and asperity height, on the scuffing that takes place in line-contact oil lubrications.
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Jen~HaurHomgeiul./Weur184(1995/203-212
Table 3 Loading schedule Stage
Load (N)
Mean Hertz pressure (CPa)
1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29
445 667 800 1030 1260 1490 1720 1950 2180 2410 2640 2870 3100 3300 3560 3790 4020 4250 4480 4710 4940 5170 5400 5630 5860 6090 6320 6550 6780
0.67 0.82 0.89 I .02 1.13 1.22 1.31 1.40 1.48 1.55 1.63 1.69 1.77
1.82 1.89 1.95 2.01 2.07 2.12 2.17 2.23 2.28 2.32 2.37 2.43 2.47 2.52 2.56 2.61
In Fig. 2, the lines of the friction coefficient shown were obtained from a line-contact lubrication as a function of the applied load upon the steel rollers with various roughness patterns and asperity heights. The spindle maintained a rotational speed of 1500 rev min-‘. The friction coefficients are the records of the entire contact process which began with the load of 445 N and ended at the scuffing load. All these lines, except for the ones for the longitudinal and oblique patterns with the asperity height of 1.2 pm, show the increases in the friction coefficient with applied load first until the uppermost value is reached; then the friction coefficient declines under higher loads. The rises in the friction coefficient occur with the rollers under mixed lubrication, and are accompanied with an initial increase in the applied load. However, the successive declines in the friction coefficient under higher loads, imply that the rollers were in frictional contact under boundary lubrication. The lines marked 5 and 6 show simple declines in the friction coefficient at the beginning of the loading process. The successive drops in the friction coefficient indicate, that the rollers with a longitudinal or oblique pattern with an asperity height of 1.2 pm, were perhaps acting in a boundary lubrication regime. The experimental results for the rollers with the asperity height of 0.6 pm or 1.2 pm exhibit the magnitudes of friction coefficient to be in the order: fT
a larger asperity height always had a higher friction coefficient before scuffing. From a theoretical analysis of partial EHL, the longitudinally oriented bearing contact areas offer little resistance to the pressure flow, permitting only a small amount of side how, as compared with those of a smooth bearing; the increased flow in the valleys more than offsets the decrease of flow under the asperities and at contacts. As a result, the pressure how in the sliding direction is greater than for a similar smooth surface, and hence decreases the film thickness and causes a large increase in the friction coefficient. On the other hand, the transversely oriented roughnesses impede the pressure flow; the restricting effect on the main flow both increases the film thickness and yields a smaller friction coefficient. According to the section of the partial EHL theory related to the roller’s surface roughness, the dependence of film thickness upon the roller’s surface asperity height with a longitudinal roughness pattern, indicates that the film thickness decreases as the roughness parameter decreases, h = h,l a; and this results in an increase in the friction coefficient. However, for the transverse roughness pattern, the Mm thickness characteristic is exactly the opposite to the longitudinal roughness pattern situation. Theoretically the transversely oriented contacts, with asperity heights of 0.6 pm, offer more resistance to mass flow than when compared with a smooth surface. The roughness pattern’s film thickness is thus magnified when the asperity height is increased; due to this, the friction coefficient should drop, to be even lower than for the smooth surface, under high loads. However in reality, when the asperity height is increased to 1.2 pm, the lubrication mode, with reference to the designated load and rotational speed, approaches or even reaches that of a mixed lubrication regime. Predictions based on the theory of elastohydrodynamic lubrication characteristics, do not completely agree with the experimental results. The scuffing load in line-contact lubrication is determined by many controlling factors: the operating conditions, the roller’s asperity height and the roughness pattern. The scuffing loads shown in Fig. 3 are the experimental results obtained
2
0 09
._
5 ._ k
“0
0.08
z
0.07 -
.‘u 2 %
0.06 :
s
0 05 -
2000
3000
Load ( N ) Fig. 2. The variations of friction coefficient with seven roughness conditions.
with applied load in the rollers
Jeng Haur Homg et al. /Wear
Roughness Pattern Fig, 3. The scuffing loads for the rollers with various asperity heights and roughness patterns. for rollers in oil lubrications with various roughness patterns and asperity heights. For the rollers with a transverse pattern, the surfaces having larger asperity heights need higher loads to cause scuffing. However, for the rollers with the longitudinal or oblique pattern, the scuffing loads corresponding to the larger asperity height (1.2 pm) are conversely lower. The scuffing loads, L, which were compared on the basis of the same asperity height, demonstrate magnitudes in the order: L..r> Lo > LL. The rollers with a transverse pattern can endure larger applied loads just before the scuffing appears, perhaps because of the ‘‘pocket effect’ ’. The pockets or valleys of the mating surfaces might serve as lubricant reservoirs; so the adsorption film or the oxide film has a much greater chance to react and develop on contact areas and to pass frictional heat to lubricant. An extensive amount of research in the past had shown that the oxide film or the adsorption film helped to increase the scuffing load [ 14,211. For a transverse pattern surface, the sliding direction is perpendicular to the valleys, resulting in stronger pocket effect, so it needs a higher load to cause scuffing. For a longitudinal surface pattern, the sliding direction is parallel to the valleys, resulting in very little pocket effect, so it has the lowest scuffing load for the same roughness height. The higher the surface roughness height, the stronger the pocket effect is for transverse surface; so the scuffing load increases together with the surface roughness. The rollers with a smooth surface were polished by machine tools so that they actually possessed a slight longitudinal pattern. When compared to the scuffing loads in the rollers with the same roughness pattern but with different asperity heights (0.6 pm and 1.2 pm), the lowering in asperity height to 0.2 pm can indeed increase the scuffing load. A rougher surface permits direct contact more easily than a smoother surface. The bottom specimen displacement variations were obtained during the loading states, by measuring the motions of the bottom specimen relative to its initial position, corresponding to the designated operating conditions. The displacements of the bottom specimen are the moving distance measurements read from the end of the long levered bar. Since
207
184 (1995) 203-212
the initial position at the beginning of motion is dependent upon the rotational speed and the applied load, although the lines all start marked by the same zero value, the lines shown in Figs. 4(a) and 4(b) simply account for the bottom specimens’ displacements relative to their different initial positions; the lines are actually unable to reflect the real film thicknesses, for lack of film thickness measurements at the beginning of motion. These displacements were obtained at the same rotational speed ( 1500 rev min- ‘) but under different loads (Fig. 4(a) is 1260 N whereas Fig. 4(b) is 2180 N). The bottom specimen displacements are the oil lubrication measurements under a load of 1260 N, as shown in Fig. 4(a). Under the designated conditions, the displacements for all seven roughness conditions were obtained. The bar graph in Fig. 3 shows that for the various roughness patterns and asperity heights, the applied load of 1260 N is lower than the scuffing loads. The results of displacement show that the rollers with the transverse pattern are apt to create positive displacements during the lubrication process; whereas, the rollers with either the longitudinal or oblique pattern show the tendency towards negative displacements. The roughness conditions readily producing positive displacement always result in higher scuffing loads; on the other hand, the negative displacements always cause lower scuffing loads. s
-6
b’
Ohliq.
(Ra=l.Z~m)
-8 i:,
4
8
12
16
20
16
20
24
Time (set) 8
-8 0
(b)
4
8
12
Time
24
(set)
Fig. 4. The displacements of bottom specimen varying with the testing time during the loading process.
208
Jeng Huur Horng et al. /Wear
-_1 1000
IS00
Rotation
?OOO 2500
3000
Speed (rpm)
Fig, 5. The variations of scuffing load with spindle’s rotational speed for the rollers with various roughness patterns and asperity heights.
(al
ib) Fig. 6. The worn surface of the driver at the speed of 750 rev min-’ and under the load of 1720 N: (a) negative slippage for the inner portion; (h) positive slippage for the outer portion (the arrow indicates the sliding direction).
In Fig. 4(b) , when the load was raised to 2180 N, only four roughness condition lines are shown (of the seven), because their scuffing load was higher than 2180 N. As seen in Fig. 3, positive displacements, such as in the lines marked by 0, 1 and 4, always resulted in relatively higher scuffing loads. No precise correlation could be established between the changes in displacement and the magnitudes of scuffing load. That is, the loading process that produced a larger increase in displacement does not necessarily lead to a higher scuffing load. The lines in Fig. 5 portray the scuffing loads as a function of spindle’s rotational speed. One can notice that the two lines
184 (1995) 203-212
marked 4 and 0 do not have recorded scuffing loads at the speed of 750 rev min- ‘. Realistically, there nevertheless exists a scuffing load for that specific rotational speed; however, the scuffing loads corresponding to lines 4 and 0 at the rotational speed of 750 rev min- ’ are greater than the load capacity (6800 N) of the wear test machine, and thus they were unable to be measured. For all the rollers with various roughness patterns and asperity heights, scuffing loads show a notable decline as the rotational speed is increased. When the system is under high rotational speeds, even under a small load scuffing still might occur. Evidently, for the rollers with the same asperity height and at the same rotational speed, the scuffing loads L, varying with roughness pattern, show magnitudes in the order: &>Lo>L,. However, the effect of asperity height on the scuffing load is dependent upon the surface roughness pattern. As for the rollers with the transverse pattern, increasing the surface asperity height makes a favourable condition for elevating the scuffing load; conversely, doing the same thing with to the rollers with either the longitudinal or oblique pattern reduces the scuffing load. The differences in the magnitude of the scuffing load due to changes in the surface asperity height are significant, especially when the rollers are acting at low rotational speeds. In fact, the region under a line represents the conditions, of applied load in conjunction with the spindle’s rotational speed, which are available without scuffing occurring. The regions created by the rollers with the transverse pattern are evidently broader than compared with those regions of the rollers with either the longitudinal or oblique pattern. Figs. 6(a) and 6(b) show the worn appearance of a driver after contact with the rollers. In the inner portion, the rolling and sliding motions during the contacts generate a larger slide to roll ratio, and the ratio is negative; plastic deformation along with fractures is on the worn surface. However, in the outer portion of the roller where the slippage was positive and smaller, surface polishing was the dominant wear mechanism. IO
,
0
1000
2000
1000
4000
5000
Load ( N ) Fig. 7. The wear rates of rollers varying with applied load for the rollers with various roughness patterns and asperity heights.
Jeng Huur Homg etul. /Wear
Roughness
pattern
Fig. 8. The coefficients of friction corresponding to the time just before the initiation of scuffing for the rollers with various roughness patterns and asperity heights.
The curves in Fig. 7 portray the wear rates of the rollers varying with applied load, for the rollers with various roughness patterns and asperity heights. All the lines are extended individually to the load level that initiates scuffing. The wear rates associated with the seven roughness conditions show different rises as the applied load increases; the change in the wear rate becomes milder under higher loads. In general, the rollers with a smaller asperity height are apt to produce a relatively lower wear rate under the applied loads. The effect of surface roughness pattern on wear behaviour, regardless of the roller surface asperity height, shows the magnitudes of wear rate corresponding to the higher loads to have the order: W,> W,> W,. This sequence is possibly altered under the lower loads due to a change in the state of lubrication; the boundary lubrication present in the contacts under higher loads is perhaps in mixed lubrication under lower loads. The friction coefficients that were measured prior to the initiation of scuffing for the seven roughness conditions are shown in Fig. 8. Based on the same surface roughness pattern, the increase of the rollers’ asperity height always elevated the friction coefficients for all three roughness patterns. It was found that the differences in friction coefficient were quite limited, even when the surface asperity height was increased from 0.6-1.2 pm. However, the surface roughness pattern seems to be more important for determining the contact friction coefficient than asperity height. The changes are notable in the friction coefficient by differing the roughness pattern, especially for the rollers with the larger asperity height. For the rollers having the same asperity height, the friction coefficients show magnitudes in the order: fT
184 (1995) 203-212
209
and roughness pattern. Concerning the rollers having the same asperity height, for the rollers with the transverse pattern, the change in the average roughness was always the largest of three patterns. The influence of roughness pattern on the average roughness of the worn surface becomes significant for the rollers with large asperity heights. Notable differences in the average roughness were found for the rollers with different roughness patterns but the same asperity height ( 1.2 pm). The photographs of the roller’s worn surfaces are provided in Figs. lO( a)--1 0( g) to illustrate their relationship with the displacement changes shown in Figs. 4(a) and 4(b). Figs. 10(a)-IO(g) portray the wear mechanisms of rollers in oil lubrication under the load of 1260 N, for the surfaces with various asperity heights and roughness patterns. According to the plots shown in Fig. 3, this applied load is still lower than the scuffing loads for any of the rollers with the seven roughness conditions; no scuffing had taken place in these photographs. Of the three roughness patterns, the surface damage shown on rollers with either a longitudinal or oblique pattern were comparatively more severe; the damages are highlighted by the white areas in the photographs. The changes of the surface topography for the various surface roughnesses, are strikingly consistent with the displacement variations shown in Fig. 4(a). Increasing the displacement present in a roller with the transverse pattern can increase the film thickness during the wear process and thus alleviate surface damage significantly. Conversely, decreasing the displacement of the bottom specimen for a roller with either the longitudinal or oblique pattern created negative displacement and thus enhanced surface damage. When the applied load was elevated to 2180 N, only four of the seven roughness conditions are shown in the photographs of Figs. 11 (a)-1 1 (d) . (The other three had already reached the scuffing load threshold.) The surface damage present in the photograph of Fig. 11 (d) for the roller with the oblique pattern is apparently the most severe. As the lines show in Fig. 4(b), the sole negative displacement was also produced in the rollers with the oblique pattern.
Roughness
Pattern
Fig. 9. The comparisons between roller’s initial and final roughnesses the rollers with various asperity heights and roughness patterns.
for
JengHaurHomgetal./Weur184(l995)203-212
210
(b)
(d)
(e 1
Fig. 10. The worn surface appearances of the roller acting at the spindle’s rotational speed of 1500 rev min- ’and under the load of 1260 N (the arrow indicates the sliding direction): (a) smooth surface; (b) transverse pattern (R,=0.6pm); (c) longitudinal pattern (R,=0.6/-an); (d) oblique pattern (R,=0.6pm); (e) transverse pattern (R,= 1.2 pm); (f) longitudinal pattern (R,= 1.2 pm); (g) oblique pattern (R,=1.2pm).
5. Conclusion The friction and wear behaviour of roller surfaces, involving changes in the roughness parameters, was presented for
a wide range of applied loads between partial elastohydrodynamic and boundary lubrications. The controlling factors (the applied load, the asperity height, and the roughness pattern) were selectively varied in the tests to establish the
Jeng Ham Homg et al. / Wear I84 (1995) 203-212
( 1)
(2)
(3)
(4)
ib)
211
The rollers with a smaller asperity height are apt to produce a relatively lower wear rate. The wear rates corresponding to the time just before scuffing starts show magnitudes in the order: W, > W, > W,. Regardless of the roller’s asperity height, the magnitudes of the friction coefficient always show the order: fT Lo > L,. Increasing the surface asperity height forms a favourablecondition for raising the scuffing load; conversely, the same manner toward the rollers with either a longitudinal or oblique pattern shows the tendency of lowering the scuffing load.
References [ 11 H. Blok, Theoretical study of temperature
Fig. 11.The worn surface appearances of the roller acting at the spindle’s rotational speed of 1500 rev min - ’and under the load of 2180 N (the arrow indicates the sliding direction) : (a) smooth surface; (b) transverse pattern (R, = 0.6 ym) ; (c) transverse pattern (R, = 1.2pm) ; (d) oblique pattern (R,=0.6pm)
correlation between operation conditions. follows:
the scuffing behaviour and the The conclusions can be drawn as
rise at surfaces of actual contact under lubricating conditions, general discussion on lubrication and lubricants, Inst. Mech. Eng., 2 ( 1937) 222-235. [2] A. Dyson, Scuffing-a review, Part 1, Tribal. Inf., 9 ( 1) ( 1975) 7787. [ 31 R.M. Matveevsky, Friction power as a criterion of seizure with sliding lubricated contact, Wear, 155 ( 1992) l-5. [4] J.C. Bell, A. Dyson and J.W. Hardley, The effects of rolling and sliding speeds on the scuffing of lubricated steel disc, ASLE Trans., 18 ( 1) ( 1972) 62-73. [5] H.J. Carper and P.M. Ku, Thermal and scuffing bebaviour of disks in sliding-rolling contact, ASLE Trans., 18, ( 1) ( 1974) 39-47. [6] H. Czichos, Failure criteria in thin film lubrication: the concept of a failure surface, Tribal. Int., 7 (1974) 14-20. 171 A. Begelinger and A.W.J. de Gree, Thin film lubrication of sliding point contacts of AISI 52100 steel, Wear, 28 ( 1974) 103-I 14. 181 Y. Yamamoto and F. Hirano, Relation between scuffing resistance and increase in surface hardness during tests under conditions of rolling/ sliding, Wear, 63 ( 1980) 165-173. [9] Y.N. Drozdow and V.G. Archegov, The scoring resistance of sliding metals, Wear, 69 (1981) 299-307. 1101 G. Niemann, H. Rettig and G. Lechner, Scuffing tests on gear oils in the FZG apparatus, ASLE Trans., 4 (1961) 71-86. [ 111 G. Bollani, Failure criteria in thin film lubrication with EP additives, Wear, 36 (1976) 19-23. [ 121 A. Jackson, M.N. Webster and J.C. Enthoven, The scuffing of lubricant traction on scuffing, STLE Trans., 37 (2) ( 1994) 387-395. [ 131 R. Bjerk, Oxygen-an extreme-pressure agent, ASLE Trans., 16 (2) (1973) 97-106. [ 141 F.S. Fein, Operating procedure effects on critical temperatures, ASLE Trans., 10 (1967) 373-385. f151 D.A. Kelly, C.G. Barnes, R.W. Freeman and G.W. Critchlow, Running-in and the enhancement of scuffing resistance, Proc. ~nst. Mech. Eng. C, 206 ( 1992) 425429.
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212 [I61 W.J.S.
Grew and A. Cameron, Thermodynamics of boundary lubrication and scuffing, Proc. R. Sot. London, Ser. A, 327 (1972) 47-59.
[17] Y. Yamamoto, The effect of surface hardness of carbon steels on scuffing resistance in rolling-sliding contact, Wear, 89 ( 1983) 22% 234. [ 181 D.B. Durkee and H.S. Cheng, An examination of possible model of scuffing failure in simple sliding, Wear, 59 (1980) 223-230.
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[ 191 J.H. Homg, J.F. Lin and K.Y. Li, The effect of surface irregularities on the tribological behaviour of steel rollers under rolling-sliding contact, ASMEJ. Tribal., 116 (2) (1994) 209-218. [20] D.W. Dudley, Gear Handbook: The Design, Manufacture, and Applications of Gears, McGraw-Hill, New York, 1984. [21] A. Jahanmir and M. Beltzer, An adsorption model for friction in boundary lubrication, ASLE Trans., 29 (3) (1986) 423430.
Effect of surface roughness on steel roller scuffing Jeng Haur Homg a, Jen Fin Lin bT*,Ke Yang Li b aDepartmenf ofPower Mechanical Engineering, National Yunlin Polytechnic Institute, Y&in 70101, Taiwan ’Department of Mechanical Engineering, National Cheng Kung University, Tainan 70101, Taiwan
Received 18 May 1994; accepted 1 December 1994
Abstract Experiments were carried out utilizing a gear-cam adapter to simulate line-contact lubrication and wear. Roller specimens with various asperity heights and roughness patterns were riding on roller plates and sliding over the two lands of the lower specimen. The scuffing loads for the rollers with various roughness conditions were measured while changing the applied load and spindle’s rotational speed. The effects of the roughness parameters on scuffing load, particularly roughness pattern and asperity height, were thus investigated. The displacement of the bottom specimens during the lubrication and wear process were measured to establish their correlation with scuffing behaviour. The variations of bottom specimen displacement can offer information regarding the film thickness formation on the rollers with various roughness patterns and asperity heights. Keywords: Roughness; Steel; Scuffing; Rollers
1. Introduction
Scuffing is an important form of damage that results from lubricant film failure in sliding and rolling contacts, such as for gears, piston rings and cams. When scuffing occurs, there is a sudden rise in frictional force, accompanied with noise and vibration; the damaged surface shows severe wear and plastic deformation. Usually, scuffing damage is so ruinous that the damaged part must be replaced. However up to now, the complete, detailed mechanism of scuffing has not been stated, nor is there any commonly accepted criterion from which the possibility of scuffing may be predicted in advance. It is believed that scuffing behaviour is a complex function of mechanical and chemical effects; the factors that govern scuffing include lubricant viscosity, oil temperature, material hardness, surface treatment, geometry profile, surface roughness, sliding/rolling speed and chemical activity within the working environment. Many studies investigated the effects of several influential factors on scuffing failure. Blok [11 suggested that scuffing would occur if the surface temperature reached a critical value. The critical surface temperature consists of two components: first, the bulk temperature of metal parts; second, the instantaneous temperature increase of the surfaces as they pass through the contact zone. Although the critical contact * Correspondingauthor. 0043.1648/95/$09.50 0 1995 Elsevier Science S.A. All rights reserved SSDIOO43-1648(94)06574-8
temperature is the most widely used criterion of scuffing, many investigators had found that the total contact temperature just before scuffing is in fact not a constant for a given combination of lubricant and materials, but varies with the experimental conditions [ 21. Matveevsky [ 31 presented a study, of the oil lubrication used in previous work, on point and line contact machines. It showed that frictional power is a viable scuffing criterion under boundary lubrication conditions. Both sliding and rolling speeds cause important effects on the scuffing conditions. Bell [ 41 and Carper [ 51 conducted experiments to examine the influences of both sliding and rolling speed on the scuffing of lubricated steel discs. They concluded after experimenting with a wider range of rolling speeds, sliding speeds and slide/roll ratios, that (a) an increase in the sliding velocity, at a constant rolling velocity, decreases the scuffing load and that, conversely, (b) an increase in the rolling velocity, at a constant sliding velocity, increases the scuffing load. Furthermore, Czichos [ 6,7] had made several detailed investigations and constructed a surface failure mode1 for three-dimensional load-velocity-temperature space which determines the limits of thin film lubrication. To improve the scuffing resistance of machine elements, it is generally considered that increasing the surface hardness is good. However, this is not always correct, since the scuffing resistance is significantly affected by a change in surface conditions, especially the surface roughness during the running-in process [ 81.
Jeng HuurHorngetul’ Wecrr 184 (1995) 203-212
204
According to the FZG machine or rotating rolling bench test, the failure load decreases with the decrease in oil viscosity [ 9,101. Scuffing resistance is increased by using high viscosity lubricants with anti-scuffing additives such as sulphur-, phosphorus-, borate ZnDDP- or PAO-base lubricants [ 11,121. Jackson [ 121 hypothesized that the reduced heat generation by low traction lubricants leads to an increase in scuffing load. Oxygen apparently plays an important role in friction and wear. The experiments by Bjerk [ 131, using a gear-roller test machine with steel rollers lubricated by plain mineral oil, indicate that surface protection against scuffing is provided by the formation of oxides in the contact zone. Other important characteristics on scuffing resistance are both the specimen surface preparation and the whole test procedure [ 141. It shows that running-in helps scuffing performance. Kelly [ 151, by using a twin-disc test machine, concluded that the dominant effect of prior running-in, in enhancing the load and temperature that lubricated surfaces can resist without scuffing, is the reduction of surface roughness. Grew [ 161 showed that chemisorption reactions with surface materials can explain the widely-known phenomenon of additive interference and thus help in understanding scuffing and running-in. The interaction between the surface roughness height and scuffing behaviour has already been well recognized. It is believed that scuffing loads increase as surface roughness decreases [ 17,181. However, the effect of roughness pattern on scuffing performances has not been studied much. Experimental results have proved that the surface roughness pattern does have a considerable effect on tribological performance under various lubrication states [ 191. In this study, experiments were carried out by utilizing a gear-cam adapter to simulate line-contact lubrication and wear. Steel rollers with various asperity heights and roughness patterns were tested in oil lubrication, obtaining scuffing load measurements while differing the applied load and the spindle’s rotational speed; hereby investigating the effects of different surface roughness parameters on the roller’s scuffing. The displacements of the bottom specimen during the loading process were also taken to try to establish their correlations with scuffing loads. The variations of displacement can offer information regarding the film thickness formations of the rollers in relation to various roughness patterns and asperity heights.
2. Application
of friction
theory
to gear contacts
The mechanism of two gear teeth contact at any moment can be simulated by two cylinders of different radii being loaded together and separated by a thin film. For convenience, the model of two long cylinders separated by a thin lubricating film is further simplified as a single long cylinder rolling and sliding on a semi-infinite solid with the same gap as in the lubricated region. The test apparatus, illustrated in Fig. 1 (a),
:.r Spindle
Driver
-
n
Roller plate Roller
Lower
1
7
specimen 1
(a)
Fig. 1. (a) Schematic view of the apparatus and arrangement of the principal working parts and (b) diagrammatic representation of the lower specimen. The dimensions are given in mm.
had two rollers, 180” apart, riding on the inner and outer lands of the lower specimen. The driver contacts the two rollers at a nominal position between the inner and the outer lands of the lower specimen, which thus splits the test in half for each roller, and in half again for each land; therefore, 25% of the test load is applied to each land-roller contact. To simplify the analysis, of the rolling-sliding contacts occurring in the working parts of the system, the roller is assumed to be only rolling at all times at point A, centred
Jeng Haur Homg et al. /Wear
Table 1 r, and r,, drive radii and their correspondingcircumferences
Outer land, r, Inner land, r, Drive land
205
Table 2 Physical propertiesof test oil
Radius (cm)
Circumference(cm)
Property
Test oil
1.905 0.714 1.310
11.969 4.469 8.227
Kinematic viscosity,cSt: constant at 40 “C; constant at 100 “C Density at 15 “C, g ml-’ Pour Point, OC Total Acid No.
14; 3.5 0.862 -15 < 0.03
between the two lands (Fig. 1 (b) ) . At that point, the driver rolls 8.22 cm without any slippage (see Table 1). For the inner land, only 4.47 cm is available for rolling at its midpoint (point F) ; so the difference, 3.75 cm, is slippage. In the same manner, the circumference at the midpoint (point C) of the outer land is 11.97 cm and the roller rolls out 8.22 cm, again leaving a 3.75 cm slippage. Although these two slippages have the same absolute value, the outer end of the roller travels a larger distance than the inner end. Therefore, forward sliding occurs on the outer land and backward sliding occurs on the inner land. The friction coefficient of the rolling-sliding contacts can be determined if the applied load and the friction torque are available. The friction torque, according to the detailed analysis demonstrated in Ref. [ 191, can be written as T=O.SW(f,r,
184 (1995) 203-212
-f2r2)
(1)
where W is the applied load;f, andf* are the friction coefficients at the outer and inner lands respectively; r, and r, are the radii of the outer and the inner lands respectively (see Table 1). Since the coefficient of friction on each individual land is not as easy to obtain as suggested in Eq. ( 1), an “average” coefficient of friction f is introduced to replace the “local” coefficient of friction at two lands. Thus, the average friction coefficient is stated as [ 191
f= 1.679TIW
(2)
3. Experimental set-up and operating conditions The line-contact experiments were conducted on a commercial test machine utilizing a gear-cam contact adapter. The working components consisted of a driver, a roller plate and a lower specimen. The lower specimen is stationary when the driver rotates at a constant speed. The assembly of those components in motion is completely immersed in an oil cup with a volume of 100 ml. A long, levered bar connected to the specimen support was utilized to measure the vertical displacement of the bottom specimen. A wear gage was vertically positioned at the end of the lever bar, with the ratio of specimen displacement to the wear gage readings being 1:6. The wear gage was not intended to measure the actual film thickness between two specimens, but was used to obtain successive readings of the bottom specimen displacement. The experimental conditions were electronically monitored by a digital instrumentation system (e.g. the speed of
rotation, the oil temperature, the applied load, the driving torque and the specimen displacement). Measurements of the roller surface roughness and the wear volume were done on a three-dimensional surfcorder. The wear rates of the roller were thus obtained by measuring the volumetric loss. The accuracy of temperature measurements using the thermocouple is + 0.5 “C, for rotating speed measurement +3 rev min-‘. All the roller specimens were made of AISI 8620 steel, with a radius of 0.25 cm and a length of 2 cm. The heat treatments of the specimens followed the requirements described in the Gear Handbook ( 1984). The roller specimens, from the viewpoint of roughness parameters, fell into four classes: ( 1) smooth surface; (2) longitudinally-oriented pattern; (3) transversely-oriented roughness pattern; (4) oblique roughness pattern forming 45” inclined angle with respect to the roller’s axis. The flat plates of the driver and the two lands of the lower specimen were all polished to have mean roughness height of 0.05 ,um. The mean roughness height of the rollers fell into three categories: 0.2 pm for the relatively smooth surface, 0.6 pm for the small roughness, and 1.2 pm for the large roughness. The composite surface roughnesses of lower specimens and rollers include the various gear finishes for gear pairs [ 201. All the drivers and the lower specimens were made of the same material as the rollers. The lubricant was a mineral base oil with the physical properties given in Table 2. The oil temperature before testing was maintained at 70” by means of an electric heater. Then the machine was brought up to the designated rotational speed with no load, and the oil temperature was increased to 8 1 “C for twenty minutes. The first load was then applied to the bottom specimen by an air pressure loading system for 30 s, and then it was released for 3 min in order to lower the oil temperature to 8 1 f 0.5 “C. The sequent loads were imposed step by step following the same procedure until scuffing occurred. The loading schedule is in Table 3.
4. Results and discussion The experimental results of this study, including the scuffing loads, the friction coefficients and the displacements of bottom specimen, are used to evaluate the effects of both a roller’s roughness pattern and asperity height, on the scuffing that takes place in line-contact oil lubrications.
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Jen~HaurHomgeiul./Weur184(1995/203-212
Table 3 Loading schedule Stage
Load (N)
Mean Hertz pressure (CPa)
1 2 3 4 5 6 7 8 9 10 11 12 13 14 15 16 17 18 19 20 21 22 23 24 25 26 27 28 29
445 667 800 1030 1260 1490 1720 1950 2180 2410 2640 2870 3100 3300 3560 3790 4020 4250 4480 4710 4940 5170 5400 5630 5860 6090 6320 6550 6780
0.67 0.82 0.89 I .02 1.13 1.22 1.31 1.40 1.48 1.55 1.63 1.69 1.77
1.82 1.89 1.95 2.01 2.07 2.12 2.17 2.23 2.28 2.32 2.37 2.43 2.47 2.52 2.56 2.61
In Fig. 2, the lines of the friction coefficient shown were obtained from a line-contact lubrication as a function of the applied load upon the steel rollers with various roughness patterns and asperity heights. The spindle maintained a rotational speed of 1500 rev min-‘. The friction coefficients are the records of the entire contact process which began with the load of 445 N and ended at the scuffing load. All these lines, except for the ones for the longitudinal and oblique patterns with the asperity height of 1.2 pm, show the increases in the friction coefficient with applied load first until the uppermost value is reached; then the friction coefficient declines under higher loads. The rises in the friction coefficient occur with the rollers under mixed lubrication, and are accompanied with an initial increase in the applied load. However, the successive declines in the friction coefficient under higher loads, imply that the rollers were in frictional contact under boundary lubrication. The lines marked 5 and 6 show simple declines in the friction coefficient at the beginning of the loading process. The successive drops in the friction coefficient indicate, that the rollers with a longitudinal or oblique pattern with an asperity height of 1.2 pm, were perhaps acting in a boundary lubrication regime. The experimental results for the rollers with the asperity height of 0.6 pm or 1.2 pm exhibit the magnitudes of friction coefficient to be in the order: fT
a larger asperity height always had a higher friction coefficient before scuffing. From a theoretical analysis of partial EHL, the longitudinally oriented bearing contact areas offer little resistance to the pressure flow, permitting only a small amount of side how, as compared with those of a smooth bearing; the increased flow in the valleys more than offsets the decrease of flow under the asperities and at contacts. As a result, the pressure how in the sliding direction is greater than for a similar smooth surface, and hence decreases the film thickness and causes a large increase in the friction coefficient. On the other hand, the transversely oriented roughnesses impede the pressure flow; the restricting effect on the main flow both increases the film thickness and yields a smaller friction coefficient. According to the section of the partial EHL theory related to the roller’s surface roughness, the dependence of film thickness upon the roller’s surface asperity height with a longitudinal roughness pattern, indicates that the film thickness decreases as the roughness parameter decreases, h = h,l a; and this results in an increase in the friction coefficient. However, for the transverse roughness pattern, the Mm thickness characteristic is exactly the opposite to the longitudinal roughness pattern situation. Theoretically the transversely oriented contacts, with asperity heights of 0.6 pm, offer more resistance to mass flow than when compared with a smooth surface. The roughness pattern’s film thickness is thus magnified when the asperity height is increased; due to this, the friction coefficient should drop, to be even lower than for the smooth surface, under high loads. However in reality, when the asperity height is increased to 1.2 pm, the lubrication mode, with reference to the designated load and rotational speed, approaches or even reaches that of a mixed lubrication regime. Predictions based on the theory of elastohydrodynamic lubrication characteristics, do not completely agree with the experimental results. The scuffing load in line-contact lubrication is determined by many controlling factors: the operating conditions, the roller’s asperity height and the roughness pattern. The scuffing loads shown in Fig. 3 are the experimental results obtained
2
0 09
._
5 ._ k
“0
0.08
z
0.07 -
.‘u 2 %
0.06 :
s
0 05 -
2000
3000
Load ( N ) Fig. 2. The variations of friction coefficient with seven roughness conditions.
with applied load in the rollers
Jeng Haur Homg et al. /Wear
Roughness Pattern Fig, 3. The scuffing loads for the rollers with various asperity heights and roughness patterns. for rollers in oil lubrications with various roughness patterns and asperity heights. For the rollers with a transverse pattern, the surfaces having larger asperity heights need higher loads to cause scuffing. However, for the rollers with the longitudinal or oblique pattern, the scuffing loads corresponding to the larger asperity height (1.2 pm) are conversely lower. The scuffing loads, L, which were compared on the basis of the same asperity height, demonstrate magnitudes in the order: L..r> Lo > LL. The rollers with a transverse pattern can endure larger applied loads just before the scuffing appears, perhaps because of the ‘‘pocket effect’ ’. The pockets or valleys of the mating surfaces might serve as lubricant reservoirs; so the adsorption film or the oxide film has a much greater chance to react and develop on contact areas and to pass frictional heat to lubricant. An extensive amount of research in the past had shown that the oxide film or the adsorption film helped to increase the scuffing load [ 14,211. For a transverse pattern surface, the sliding direction is perpendicular to the valleys, resulting in stronger pocket effect, so it needs a higher load to cause scuffing. For a longitudinal surface pattern, the sliding direction is parallel to the valleys, resulting in very little pocket effect, so it has the lowest scuffing load for the same roughness height. The higher the surface roughness height, the stronger the pocket effect is for transverse surface; so the scuffing load increases together with the surface roughness. The rollers with a smooth surface were polished by machine tools so that they actually possessed a slight longitudinal pattern. When compared to the scuffing loads in the rollers with the same roughness pattern but with different asperity heights (0.6 pm and 1.2 pm), the lowering in asperity height to 0.2 pm can indeed increase the scuffing load. A rougher surface permits direct contact more easily than a smoother surface. The bottom specimen displacement variations were obtained during the loading states, by measuring the motions of the bottom specimen relative to its initial position, corresponding to the designated operating conditions. The displacements of the bottom specimen are the moving distance measurements read from the end of the long levered bar. Since
207
184 (1995) 203-212
the initial position at the beginning of motion is dependent upon the rotational speed and the applied load, although the lines all start marked by the same zero value, the lines shown in Figs. 4(a) and 4(b) simply account for the bottom specimens’ displacements relative to their different initial positions; the lines are actually unable to reflect the real film thicknesses, for lack of film thickness measurements at the beginning of motion. These displacements were obtained at the same rotational speed ( 1500 rev min- ‘) but under different loads (Fig. 4(a) is 1260 N whereas Fig. 4(b) is 2180 N). The bottom specimen displacements are the oil lubrication measurements under a load of 1260 N, as shown in Fig. 4(a). Under the designated conditions, the displacements for all seven roughness conditions were obtained. The bar graph in Fig. 3 shows that for the various roughness patterns and asperity heights, the applied load of 1260 N is lower than the scuffing loads. The results of displacement show that the rollers with the transverse pattern are apt to create positive displacements during the lubrication process; whereas, the rollers with either the longitudinal or oblique pattern show the tendency towards negative displacements. The roughness conditions readily producing positive displacement always result in higher scuffing loads; on the other hand, the negative displacements always cause lower scuffing loads. s
-6
b’
Ohliq.
(Ra=l.Z~m)
-8 i:,
4
8
12
16
20
16
20
24
Time (set) 8
-8 0
(b)
4
8
12
Time
24
(set)
Fig. 4. The displacements of bottom specimen varying with the testing time during the loading process.
208
Jeng Huur Horng et al. /Wear
-_1 1000
IS00
Rotation
?OOO 2500
3000
Speed (rpm)
Fig, 5. The variations of scuffing load with spindle’s rotational speed for the rollers with various roughness patterns and asperity heights.
(al
ib) Fig. 6. The worn surface of the driver at the speed of 750 rev min-’ and under the load of 1720 N: (a) negative slippage for the inner portion; (h) positive slippage for the outer portion (the arrow indicates the sliding direction).
In Fig. 4(b) , when the load was raised to 2180 N, only four roughness condition lines are shown (of the seven), because their scuffing load was higher than 2180 N. As seen in Fig. 3, positive displacements, such as in the lines marked by 0, 1 and 4, always resulted in relatively higher scuffing loads. No precise correlation could be established between the changes in displacement and the magnitudes of scuffing load. That is, the loading process that produced a larger increase in displacement does not necessarily lead to a higher scuffing load. The lines in Fig. 5 portray the scuffing loads as a function of spindle’s rotational speed. One can notice that the two lines
184 (1995) 203-212
marked 4 and 0 do not have recorded scuffing loads at the speed of 750 rev min- ‘. Realistically, there nevertheless exists a scuffing load for that specific rotational speed; however, the scuffing loads corresponding to lines 4 and 0 at the rotational speed of 750 rev min- ’ are greater than the load capacity (6800 N) of the wear test machine, and thus they were unable to be measured. For all the rollers with various roughness patterns and asperity heights, scuffing loads show a notable decline as the rotational speed is increased. When the system is under high rotational speeds, even under a small load scuffing still might occur. Evidently, for the rollers with the same asperity height and at the same rotational speed, the scuffing loads L, varying with roughness pattern, show magnitudes in the order: &>Lo>L,. However, the effect of asperity height on the scuffing load is dependent upon the surface roughness pattern. As for the rollers with the transverse pattern, increasing the surface asperity height makes a favourable condition for elevating the scuffing load; conversely, doing the same thing with to the rollers with either the longitudinal or oblique pattern reduces the scuffing load. The differences in the magnitude of the scuffing load due to changes in the surface asperity height are significant, especially when the rollers are acting at low rotational speeds. In fact, the region under a line represents the conditions, of applied load in conjunction with the spindle’s rotational speed, which are available without scuffing occurring. The regions created by the rollers with the transverse pattern are evidently broader than compared with those regions of the rollers with either the longitudinal or oblique pattern. Figs. 6(a) and 6(b) show the worn appearance of a driver after contact with the rollers. In the inner portion, the rolling and sliding motions during the contacts generate a larger slide to roll ratio, and the ratio is negative; plastic deformation along with fractures is on the worn surface. However, in the outer portion of the roller where the slippage was positive and smaller, surface polishing was the dominant wear mechanism. IO
,
0
1000
2000
1000
4000
5000
Load ( N ) Fig. 7. The wear rates of rollers varying with applied load for the rollers with various roughness patterns and asperity heights.
Jeng Huur Homg etul. /Wear
Roughness
pattern
Fig. 8. The coefficients of friction corresponding to the time just before the initiation of scuffing for the rollers with various roughness patterns and asperity heights.
The curves in Fig. 7 portray the wear rates of the rollers varying with applied load, for the rollers with various roughness patterns and asperity heights. All the lines are extended individually to the load level that initiates scuffing. The wear rates associated with the seven roughness conditions show different rises as the applied load increases; the change in the wear rate becomes milder under higher loads. In general, the rollers with a smaller asperity height are apt to produce a relatively lower wear rate under the applied loads. The effect of surface roughness pattern on wear behaviour, regardless of the roller surface asperity height, shows the magnitudes of wear rate corresponding to the higher loads to have the order: W,> W,> W,. This sequence is possibly altered under the lower loads due to a change in the state of lubrication; the boundary lubrication present in the contacts under higher loads is perhaps in mixed lubrication under lower loads. The friction coefficients that were measured prior to the initiation of scuffing for the seven roughness conditions are shown in Fig. 8. Based on the same surface roughness pattern, the increase of the rollers’ asperity height always elevated the friction coefficients for all three roughness patterns. It was found that the differences in friction coefficient were quite limited, even when the surface asperity height was increased from 0.6-1.2 pm. However, the surface roughness pattern seems to be more important for determining the contact friction coefficient than asperity height. The changes are notable in the friction coefficient by differing the roughness pattern, especially for the rollers with the larger asperity height. For the rollers having the same asperity height, the friction coefficients show magnitudes in the order: fT
184 (1995) 203-212
209
and roughness pattern. Concerning the rollers having the same asperity height, for the rollers with the transverse pattern, the change in the average roughness was always the largest of three patterns. The influence of roughness pattern on the average roughness of the worn surface becomes significant for the rollers with large asperity heights. Notable differences in the average roughness were found for the rollers with different roughness patterns but the same asperity height ( 1.2 pm). The photographs of the roller’s worn surfaces are provided in Figs. lO( a)--1 0( g) to illustrate their relationship with the displacement changes shown in Figs. 4(a) and 4(b). Figs. 10(a)-IO(g) portray the wear mechanisms of rollers in oil lubrication under the load of 1260 N, for the surfaces with various asperity heights and roughness patterns. According to the plots shown in Fig. 3, this applied load is still lower than the scuffing loads for any of the rollers with the seven roughness conditions; no scuffing had taken place in these photographs. Of the three roughness patterns, the surface damage shown on rollers with either a longitudinal or oblique pattern were comparatively more severe; the damages are highlighted by the white areas in the photographs. The changes of the surface topography for the various surface roughnesses, are strikingly consistent with the displacement variations shown in Fig. 4(a). Increasing the displacement present in a roller with the transverse pattern can increase the film thickness during the wear process and thus alleviate surface damage significantly. Conversely, decreasing the displacement of the bottom specimen for a roller with either the longitudinal or oblique pattern created negative displacement and thus enhanced surface damage. When the applied load was elevated to 2180 N, only four of the seven roughness conditions are shown in the photographs of Figs. 11 (a)-1 1 (d) . (The other three had already reached the scuffing load threshold.) The surface damage present in the photograph of Fig. 11 (d) for the roller with the oblique pattern is apparently the most severe. As the lines show in Fig. 4(b), the sole negative displacement was also produced in the rollers with the oblique pattern.
Roughness
Pattern
Fig. 9. The comparisons between roller’s initial and final roughnesses the rollers with various asperity heights and roughness patterns.
for
JengHaurHomgetal./Weur184(l995)203-212
210
(b)
(d)
(e 1
Fig. 10. The worn surface appearances of the roller acting at the spindle’s rotational speed of 1500 rev min- ’and under the load of 1260 N (the arrow indicates the sliding direction): (a) smooth surface; (b) transverse pattern (R,=0.6pm); (c) longitudinal pattern (R,=0.6/-an); (d) oblique pattern (R,=0.6pm); (e) transverse pattern (R,= 1.2 pm); (f) longitudinal pattern (R,= 1.2 pm); (g) oblique pattern (R,=1.2pm).
5. Conclusion The friction and wear behaviour of roller surfaces, involving changes in the roughness parameters, was presented for
a wide range of applied loads between partial elastohydrodynamic and boundary lubrications. The controlling factors (the applied load, the asperity height, and the roughness pattern) were selectively varied in the tests to establish the
Jeng Ham Homg et al. / Wear I84 (1995) 203-212
( 1)
(2)
(3)
(4)
ib)
211
The rollers with a smaller asperity height are apt to produce a relatively lower wear rate. The wear rates corresponding to the time just before scuffing starts show magnitudes in the order: W, > W, > W,. Regardless of the roller’s asperity height, the magnitudes of the friction coefficient always show the order: fT
References [ 11 H. Blok, Theoretical study of temperature
Fig. 11.The worn surface appearances of the roller acting at the spindle’s rotational speed of 1500 rev min - ’and under the load of 2180 N (the arrow indicates the sliding direction) : (a) smooth surface; (b) transverse pattern (R, = 0.6 ym) ; (c) transverse pattern (R, = 1.2pm) ; (d) oblique pattern (R,=0.6pm)
correlation between operation conditions. follows:
the scuffing behaviour and the The conclusions can be drawn as
rise at surfaces of actual contact under lubricating conditions, general discussion on lubrication and lubricants, Inst. Mech. Eng., 2 ( 1937) 222-235. [2] A. Dyson, Scuffing-a review, Part 1, Tribal. Inf., 9 ( 1) ( 1975) 7787. [ 31 R.M. Matveevsky, Friction power as a criterion of seizure with sliding lubricated contact, Wear, 155 ( 1992) l-5. [4] J.C. Bell, A. Dyson and J.W. Hardley, The effects of rolling and sliding speeds on the scuffing of lubricated steel disc, ASLE Trans., 18 ( 1) ( 1972) 62-73. [5] H.J. Carper and P.M. Ku, Thermal and scuffing bebaviour of disks in sliding-rolling contact, ASLE Trans., 18, ( 1) ( 1974) 39-47. [6] H. Czichos, Failure criteria in thin film lubrication: the concept of a failure surface, Tribal. Int., 7 (1974) 14-20. 171 A. Begelinger and A.W.J. de Gree, Thin film lubrication of sliding point contacts of AISI 52100 steel, Wear, 28 ( 1974) 103-I 14. 181 Y. Yamamoto and F. Hirano, Relation between scuffing resistance and increase in surface hardness during tests under conditions of rolling/ sliding, Wear, 63 ( 1980) 165-173. [9] Y.N. Drozdow and V.G. Archegov, The scoring resistance of sliding metals, Wear, 69 (1981) 299-307. 1101 G. Niemann, H. Rettig and G. Lechner, Scuffing tests on gear oils in the FZG apparatus, ASLE Trans., 4 (1961) 71-86. [ 111 G. Bollani, Failure criteria in thin film lubrication with EP additives, Wear, 36 (1976) 19-23. [ 121 A. Jackson, M.N. Webster and J.C. Enthoven, The scuffing of lubricant traction on scuffing, STLE Trans., 37 (2) ( 1994) 387-395. [ 131 R. Bjerk, Oxygen-an extreme-pressure agent, ASLE Trans., 16 (2) (1973) 97-106. [ 141 F.S. Fein, Operating procedure effects on critical temperatures, ASLE Trans., 10 (1967) 373-385. f151 D.A. Kelly, C.G. Barnes, R.W. Freeman and G.W. Critchlow, Running-in and the enhancement of scuffing resistance, Proc. ~nst. Mech. Eng. C, 206 ( 1992) 425429.
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212 [I61 W.J.S.
Grew and A. Cameron, Thermodynamics of boundary lubrication and scuffing, Proc. R. Sot. London, Ser. A, 327 (1972) 47-59.
[17] Y. Yamamoto, The effect of surface hardness of carbon steels on scuffing resistance in rolling-sliding contact, Wear, 89 ( 1983) 22% 234. [ 181 D.B. Durkee and H.S. Cheng, An examination of possible model of scuffing failure in simple sliding, Wear, 59 (1980) 223-230.
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[ 191 J.H. Homg, J.F. Lin and K.Y. Li, The effect of surface irregularities on the tribological behaviour of steel rollers under rolling-sliding contact, ASMEJ. Tribal., 116 (2) (1994) 209-218. [20] D.W. Dudley, Gear Handbook: The Design, Manufacture, and Applications of Gears, McGraw-Hill, New York, 1984. [21] A. Jahanmir and M. Beltzer, An adsorption model for friction in boundary lubrication, ASLE Trans., 29 (3) (1986) 423430.