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steel friction pairs
Wear 267 (2009) 1241–1251
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Experimental study of the effect of microtexturing on oil lubricated ceramic/steel friction pairs K.-H. Zum Gahr ∗ , R. Wahl, K. Wauthier Universität Karlsruhe, Institute of Materials Science and Engineering II and Forschungszentrum Karlsruhe, Institute for Materials Research I, P.O. Box 3640, 76185 Karlsruhe, Germany
a r t i c l e
i n f o
Article history: Received 1 September 2008 Received in revised form 15 December 2008 Accepted 15 December 2008 Keywords: Microtexturing Friction Oil lubrication Film thickness Advanced ceramic In situ observation
a b s t r a c t The objective of the research work was to study the potential of commercial ceramics such as Al2 O3 , Al2 O3 –ZrO2 and SiC in pairing with steel 100Cr6 and laser-assisted deterministic microtexturing of the flat functional surfaces as design element for fast running friction systems, such as multiple disc clutches, under unidirectional oil lubricated sliding. A more fundamental part of the study was oriented towards understanding of the effect of microtexturing on mechanisms at mixed or boundary lubrication and development of lubricating oil film in the contact area. Experiments were carried out with pellet/disc geometry at relatively low velocities ≤0.3 m/s and normal loads ≤10 N using a model test with an attached microscope for in situ observation of the contact area. In the second part, friction behaviour of the different ceramic materials with and without microtexturing was characterised using a friction test rig at sliding velocities up to 10 m/s and normal loads ≤60 N under more practical-oriented conditions. Results showed that frictional behaviour was significantly influenced by the ceramic material and features of microtexturing, e.g. dimples, channels, width, depth and area coverage fraction. Improved understanding of operating mechanisms and design rules of microtexturing for controlling friction coefficient at practical sliding systems was obtained from a descriptive model. © 2009 Elsevier B.V. All rights reserved.
1. Introduction There is a great demand for new materials and improved design of functional surfaces at many sliding and friction systems owing to requirements such as handling of liquids of low viscosity at hermetically sealed pumps or increased power transmission in the power train of automotive engines. Advanced multiple disc clutches are an example of an oil lubricated friction system whereas currently used friction materials meet limits of mechanical and thermal load capacity [1,2]. One reason are the so-called hot spots which are formed locally on the frictional surfaces due to high pressures and high temperatures. Even an excess of the critical sliding velocity for hot spotting during a single short engagement can lead to failure or degradation of the friction materials and as a consequence of the clutch. New friction materials have to fulfil high demands on a sufficiently high and stable friction coefficient for power transmission, running comfort under varying operating conditions with sliding velocities up to about 30 m/s as well as high wear resistance and thermal stability over a long service time. Advanced ceramic materials such as alumina or silicon carbide offer a high potential under such severe loading conditions.
∗ Corresponding author. Tel.: +49 7247 82 3897; fax: +49 7247 82 4567. E-mail address: [email protected] (K.-H. Zum Gahr). 0043-1648/$ – see front matter © 2009 Elsevier B.V. All rights reserved. doi:10.1016/j.wear.2008.12.108
Greater friction values and hence greater power transmission can be achieved under oil lubrication with ceramic/ceramic or steel/ceramic than for example with steel/steel sliding pairs [3]. High scuffing resistance of ceramics is beneficial under lubrication with liquids of low viscosity or at mixed and boundary lubrication at more or less fractions of solid/solid contact [4,5]. Friction and wear are strongly influenced under poor lubrication by topography of the functional surfaces [5–7]. With decreasing film thickness of the liquid a transition to boundary friction can occur depending on the operating conditions and thus tribochemical reactions may be caused on materials such as Al2 O3 or SiC or even phase transformations in the surface of partially stabilised ZrO2 ceramics. Tribochemical films produced at the loaded surface dominate friction and wear characteristic under these conditions. Surface microtexturing of the functional areas can take effect in improving tribological performance of lubricated systems or in friction control under varying operating conditions. Several studies showed that texture pattern of dimples or grooves may cause effects such as removal of wear debris from the contact, improvement of running-in behaviour, improvement of wetting behaviour, action as lubricant reservoirs, increase of load bearing capacity or promotion of hydrodynamic lubrication [8–17]. All these effects can lead to a substantial reduction of friction in sliding systems. It was also reported that shape [15], size [8,9,11,15], area coverage fraction [11,15] or orientation of the texture elements [15] as well as operat-
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Fig. 1. Schematic description of experimental set-up, test parameters, materials and lubrication of (a) model test and (b) friction test rig.
ing conditions or lubricant properties [14,17] play an important role for film thickness of the lubricating liquid and friction behaviour of the system. Also transitions between liquid and mixed or boundary friction may be affected by microtexturing [14]. However, the effectiveness differs substantially and in some cases texturing showed no significant or even a detrimental effect on friction coefficient and load bearing capacity depending on the mated materials and experimental conditions used. The reason is a complex nature of contact phenomena occurring between the surfaces and the liquid in contact. Trustworthy results can be expected at parallel flat contacts but point or line contacts may cause beneficial or detrimental effects on friction, since ratio of scale of contact and texture pattern may become crucial. It follows, that much more knowledge is needed to design an optimised microtexturing for a special application. Aim of the present study was to achieve a deeper knowledge of the mechanisms and most important factors influencing the frictional contact of microtextured surfaces of oil lubricated ceramic/steel pairs during unidirectional sliding. Contact conditions were analysed by using high resolution experimental methods and results were used for description of the frictional contact as a function of influencing factors such as materials mated, parameters of texture pattern, load, sliding velocity or film thickness of lubricating oil. 2. Materials and experimental methods
16 mm for these tests were fabricated from the commercial ceramics SSiC (EKasicF, ESK Ceramics), Al2 O3 (HTC 99.9, Hightech Ceram) and Al2 O3 –15 vol.% ZrO2 (SN80, CeramTec). A flat circular contact area of Ø 6 mm was lapped at the convex (radius of 100 mm) end face of the ceramic pellets using directly the tribological test rig for reducing misalignment. Properties and micrographs of the different materials are given in Table 1 and Fig. 2, respectively. 2.2. Microtexturing of pellets Laser-assisted microtexturing was carried out on the functional area of the pellets by using an Ytterbium-Fibrelaser (ACSYS Lasertechnik) at a wave length of 1.064 m and a power of about 20 W depending on the material. Deterministic texture patterns (Fig. 2) of circular microdimples (CD) or microchannels parallel or crossed to each other (CH) were produced on the lapped or polished areas of the pellets by laser ablation process. Texture parameters such as width, depth, offset and area coverage fraction were varied in different test series. Width of channels was varied from 50 to 300 m depending on the test. Larger widths of the channels were generated by putting several tracks side by side with a 10 m offset. Depth was varied between 5 and 20 m. Values of area coverage fractions accounted for 35, 50 and 75%. Offset values resulted from the width and area coverage fraction selected (Fig. 2c). After laser ablation the texture patterns showed debris, which was removed using gentle grinding and polishing.
2.1. Materials and specimens 2.3. Tribological test Tribological behaviour was studied at unidirectional sliding using a pellet-on-disc geometry. As delivered pellets of the steel 100Cr6 were used through hardened and tempered with a hardness of 790 ± 10 HV. A flat circular contact area of Ø 7.2 mm was machined at the convex (radius of 8 mm) end face. The contact area of the pellet was finally polished to Ra = 0.01 m and mated with a polished sapphire disc (GWI Sapphire) during tests using a model tribometer (Fig. 1a). On the other hand, ceramic pellets were mated with normalized discs of the steel 100Cr6 (300 ± 10 HV) in tests using a larger friction test rig (Fig. 1b). The cylindrical pellets of Ø
The model test and the friction test rig differed in the manner that hardened and tempered 100Cr6 steel pellets were mated with sapphire discs at the model test while pellets fabricated from different ceramics (HTC 99.9, SN80, EKasic F) were mated with normalized 100Cr6 steel discs at the friction test rig (Fig. 2). Using the commercial friction test rig (Fig. 1b; UMT3, CETR), oil lubricated tests were carried out at unidirectional sliding up to relative velocities of 10 m/s and applied pressures of 2.12 MPa calculated for untextured circular contact area of Ø 6 mm and nor-
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Table 1 Marking and properties of materials used for the model test and friction test rig. Marking
Material
Specimen
Grain size (m)
Density (103 kg/m3 )
Hardness (HV500 )
Young’s modulus (GPa)
Sapphire 100Cr6 HTC 99.9 SN80 EKasic F 100Cr6
Al2 O3 -single crystal Hardened 99.9 wt.%–Al2 O3 Al2 O3 –15 vol.% ZrO2 SSiC Normalized
Disc Pellet Pellet Pellet Pellet Disc
– – 1.0–3.0 1.5 1.9 –
3.98 7.84 3.96 4.10 3.17 7.84
2200 790 2084 1610 2540 300
350 212 380 380 410 212
Fig. 2. (a–c) Schematic description of geometry and size of texture pattern as well as SEM micrographs of (d) microtextured surfaces of steel 100Cr6 or Ekasic F and (e) polished microstructure of the ceramic materials used.
mal load of FN = 60 N. Textured or untextured ceramic pellets were mated with the directly in the tribometer lapped contact area with fine ground discs of the pearlitic 100Cr6 steel (Ra = 0.10 m, 300 ± 10 HV). Mineral oil (ISO-VG 100, FVA No. 3, 40◦ C = 85.2 mPa s) without additives was fed into the contact area by drip lubrication. The stationary slip test (Fig. 1b) was characterized in that way that
after loading the friction pair by the normal load, sliding velocity was continuously increased during 5 s up to the value wanted and it was hold constant for 120 s before the velocity was shut down to zero. Each test at a given constant velocity (varied from 0.10 to 10 m/s), was followed by a cooling period down to room temperature before the run for the next greater velocity was started. At
Fig. 3. Polished or microtextured steel pellets mated with polished sapphire discs at the model test (a) friction coefficient and (b) oil film thickness h* versus sliding velocity (CD: circular dimples, CH: crossed microchannels and pol.: untextured, polished (FN = 5 N, atext. = 35%, B = 100 m, T = 10 m)).
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Fig. 4. (a) Friction coefficient and (b) oil film thickness of 100Cr6 pellets with three different oriented (A, B = parallel and C = crossed) microchannels as well as untextured surface (P) mated with sapphire discs (FN = 5 N, v = 0.30 m/s, atext. = 50%, B = 100 m, T = 10 m).
the deceleration test friction behaviour was measured during the slow down period lasting 5 s from a sliding velocity of 10 m/s down to zero. All tests were run in laboratory air at room temperature of 22 ◦ C and 50% relative humidity. Reported values were taken from representative runs. Studies using the model test (Fig. 1a) were run at low velocities and low normal loads up to 0.3 m/s and 10 N, respectively. Martensitic 100Cr6 steel pellets (790 ± 10 HV) with polished or microtextured contact areas of Ø 7.2 mm were mated at oil lubricated unidirectional sliding with polished sapphire discs. The same mineral oil as used in the friction test rig was fed at 2.0 mm3 /s by drip lubrication in front of the contact area. Besides friction force, separation distance h* between the contact area of steel pellet and sapphire disc was continuously measured during each test by using a high accuracy distance transmitter with a resolution of ±0.5 m. Recorded values of separation distance “h* ” represent a measure of
the oil film thickness “h” but it has to be expected that these values can be greater than the real oil film thickness “h” because waviness of the sapphire disc (<1 m after manufacturer data) or alignment variations will contribute to the measured data. In situ observation of the contact area during sliding was possible using a special microscope with magnification up to 200 fold. It was preferentially applied for studying the development of the lubricant film and the oil flow at microtextured contact areas.
3. Results Oriented towards the aim to improve the understanding of effects of microtexturing, model tests were carried out and the transferability of their results to more application-oriented loading conditions was proved by tests using the friction test rig.
Fig. 5. Effect of normal load and sliding velocity on friction coefficient and oil film thickness at 100Cr6/sapphire pairs with (a and b) untextured and (c and d) crossed microchannels textured steel pellets (atext. = 75%, B = 100 m, T = 10 m).
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Fig. 6. (a) Friction coefficient and (b) oil film thickness of 100Cr6/sapphire pairs with untextured (pol) pellets or pellets textured with crossed microchannels of 35, 50 or 75% versus sliding velocity (CH35–CH75: atext. = 35–75%, B = 100 m, T = 10 m, FN = 5 N).
3.1. Model test Friction coefficient and oil film thickness of hardened 100Cr6 steel pellets with and without microtextured functional area mated with sapphire discs were measured in the model test as function of loading and microtexturing parameters. 3.1.1. Texture pattern Fig. 3 shows friction coefficient and thickness of oil film of 100Cr6/sapphire pairs with polished or microtextured steel pellets as function of sliding velocity. At the normal load of 5 N, depth of texture pattern of 10 m and area coverage fraction atext. = 35%, lower values of film thickness and greater values of friction coefficient were measured at pellets textured with crossed microchannels (CH) than with circular microdimples (CD) or untextured only polished surface. Both film thickness and friction coefficient of pairs with CD-textured or untextured pellets increased continuously with increasing sliding velocity. The pellet textured with microdimples showed a greater slope of friction with
sliding velocity than the untextured pellet (Fig. 3b) that pointed on a hydrodynamic effect. However, texturing with microchannels resulted in a minimum of friction coefficient at about 0.10 m/s under the experimental conditions used (Fig. 3a), i.e. a behaviour similar to that described by the “Stribeck curve”. Friction coefficient increased with decreasing sliding velocity below this value in agreement with the simultaneously decreasing film thickness to very low values down to the scale of surface roughness of the mated specimens (Fig. 3b). This means that a transition from full film lubrication to partial (or mixed) lubrication and then to boundary lubrication occurred with decreasing sliding velocity. Since higher friction values are advantageous for applications in oil lubricated friction systems, following studies were focused on effects of microchannels. 3.1.2. Orientation of microchannels Fig. 4 shows friction coefficient and film thickness of 100Cr6/sapphire pairs as function of the orientations of parallel microchannels (A, B) compared with crossed microchannels (C) and
Fig. 7. (a) Friction coefficient and (b) oil film thickness of 100Cr6/sapphire pairs with untextured (pol) pellets or pellets textured with crossed microchannels of (a and b) depth T = 5, 10 or 20 m at the width of 100 m and (c and d) width B = 60, 100 or 300 m at the depth of 10 m versus sliding velocity (atext. = 50%, FN = 5 N).
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Fig. 8. Friction coefficient of different ceramic pellets (a and c) untextured and (b and d) textured with crossed microchannels mated with normalized 100Cr6 steel discs (a and b) versus sliding time, i.e. 5 s acceleration and 120 s at 10 m/s (stationary slip test) and (c and d) versus sliding velocity (deceleration test) at normal load of 60 N (CH: atext. = 75%, B = 100 m, T = 10 m).
without texturing (P) of the steel pellets. Orientation (A) led to eased oil drain and orientation (B) to constrained oil drain with respect to the sliding direction of the rotating sapphire disc. It is expected that a situation of constrained oil drain out of the sliding contact can benefit hydrodynamic lubrication. This is supported by the greater
film thickness and lower friction coefficient measured at 0.30 m/s and 5 N normal load for the orientation (B) than for (A). Values of film thickness measured at the pellet with crossed microchannels (C) ranged between (A) and (B) (Fig. 4b) but friction coefficient was greater than (A) and (B) (Fig. 4a). In comparison, the untextured
Fig. 9. Friction coefficient at 125 s of sliding versus sliding velocity of (a and b) 100Cr6 discs mated with HTC 99.9, SN80 or Ekasic F pellets, (a) untextured and (b) textured, at FN = 60 N as well as SN80/100Cr6 pairs with untextured or textured pellets at (c) normal load of 15 N and (d) 60 N, respectively (CH: atext. = 75%, B = 100 m, T = 10 m, stationary slip test at different sliding velocities).
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Fig. 10. Friction coefficient of SN80 pellets textured with crossed microchannels of area fractions of 35, 50 and 75% or untextured as reference mated with 100Cr6 discs versus (a) sliding time at the stationary slip test, i.e. 5 s acceleration and 120 s at 10 m/s, as well as (b) sliding velocity from 10 m/s to zero at the deceleration test (B = 100 m, T = 10 m, FN = 60 N).
polished pellet resulted in relatively high film thickness and low friction. 3.1.3. Normal load and sliding velocity Effects of normal load and sliding velocity on both friction coefficient and oil film thickness are shown in Fig. 5 at 100Cr6/sapphire pairs with untextured pellets and pellets textured with crossed microchannels of atext. = 75%. In case of untextured pellets, friction coefficient and oil film thickness increased continuously with sliding velocity but decreased with increasing normal load (Fig. 5a and b). The relation between film thickness and loading parameters (FN , v) was not changed by texturing with crossed microchannels, however, the values were substantially smaller (Fig. 5d). In contrast to this, friction behaviour was substantially altered by the texture in the way that friction coefficient increased with textured pellets with decreasing sliding velocity (Fig. 5c). This was caused by transition from liquid to mixed or boundary friction. The increasing friction with decreasing sliding velocity correlates very well with the measured reduction in film thickness down to very low values (Fig. 5d). 3.1.4. Area coverage fraction Fig. 6 shows the influence of area coverage fraction of crossed microchannels on friction coefficient and oil film thickness of 100Cr6/sapphire pairs. Independently of the area fraction of microchannels and sliding velocity, oil film thickness was smaller at textured than untextured pellets. Increasing atext. from 35 to 75% resulted in decreasing film thickness (Fig. 6b). Film thickness values falling below a limit of about 0.5 m led to increasing friction coefficient with further decreasing sliding velocities (Fig. 6a) due to transition to partial lubrication because of breakthrough of the oil film and partial solid/solid contact between the mated surfaces. The value of sliding velocity at this transition was shifted to greater velocities with increasing area fraction of the microchannels (Fig. 6a). 3.1.5. Depth and width of texture pattern Depth and width of the crossed microchannels are texture variables which can be selected system-specific for optimizing friction behaviour. Fig. 7a and b shows the effect of the depth of crossed microchannels on friction coefficient and oil film thickness of the 100Cr6/sapphire pairs at an area coverage fraction of atext. = 50% and a channel width of 100 m. Pellets textured with shallow microchannels of the depth of 5 m showed qualitatively similar friction behaviour as function of sliding velocity as untextured pellets (Fig. 7a). Below a sliding velocity of about 0.15 m/s, values of film thickness and friction coefficient approached those of the untextured pellets, under the applied experimental conditions. Both friction and film thickness curve were substantially altered
at sliding velocities below about 0.15 m/s when the channel depth was enhanced from 5 to 10 m. At 10 m depth, transition to partial lubrication occurred accompanied by increasing friction values. Pellets with microchannels of 20 m depth run under partial or boundary lubrication and high values of friction were measured at all sliding velocities (Fig. 7a). Friction coefficient and oil film thickness of 100Cr6/sapphire pairs with pellets textured with crossed microchannels of different widths of 60, 100 and 300 m but constant depth of 10 m are given in Fig. 7c and d. Pellets with narrow channels of 60 m showed similar values of friction coefficient and oil film thickness as untextured pellets in the velocity range tested. Larger widths of 100 and 300 m decreased substantially film thickness and led to transition from liquid to mixed friction (Fig. 7c). Both the curve of friction coefficient and film thickness of pellets with the very broad channels of 300 m deviated from those of 100 m at sliding velocities below about 0.10 m/s. Results of friction coefficient and film thickness of the pairs with pellets textured by microchannels of the largest depth of 20 m or largest width of 300 m point at possibilities to increase friction coefficient and to reduce dependence of friction coefficient on sliding velocity. 3.2. Friction test rig Tribological tests were run using two different test procedures namely a stationary slip test and a deceleration test for characterisation of frictional behaviour of commercial ceramics as pellets mated with normalized 100Cr6 discs. The tests were focused on application of advanced ceramics with and without microtexturing as materials in oil lubricated friction systems, such as multiple disc clutches. 3.2.1. Materials and operating conditions Fig. 8 shows friction coefficient of pellets fabricated from oxide and non-oxide ceramics (HTC 99.9: Al2 O3 , SN80: Al2 O3 –ZrO2 and Ekasic F: SSiC) mated with normalized 100Cr6 steel discs as function of sliding time or sliding velocity. Using the stationary slip test (Fig. 1b), friction coefficient was measured during 120 s sliding at 10 m/s (Fig. 8a and b). After the acceleration period to 10 m/s during the first 5 s, values of friction coefficient decreased more or less strongly with prolonged time of sliding. Friction coefficient of the pairs with untextured pellets was between 0.08 and 0.10 at the end of the tests, whereas the greatest value was measured at EKasic F and the lowest at HTC 99.9 (Fig. 8a). Microtexturing of the pellets with crossed microchannels of atext. = 75% enhanced friction values substantially, whereby a value of 0.15 was achieved with SN80 pellets at the end of the test (Fig. 8b). Friction behaviour of the different pairs was also characterised using the deceleration test, i.e. sliding velocity was slowed down
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Fig. 11. Friction coefficient of SN80 pellets textured with crossed microchannels of widths varied from 50 to 200 m mated with 100Cr6 discs versus (a) sliding time, i.e. 5 s acceleration and 120 s at 10 m/s (stationary slip test) and (b) versus sliding velocity from 10 m/s to zero (deceleration test) at the normal load of 60 N (atext. = 75%, T = 10 m).
from 10 m/s to zero during 5 s (Fig. 1b). In the case of untextured pellets, friction coefficient decreased continuously with decreasing sliding velocity (Fig. 8c). However, friction coefficient of pairs with microtextured pellets stayed almost constant with decreasing velocity at values between about 0.11 and 0.13 depending on the ceramic used (Fig. 8d). Below sliding velocities of about 2 m/s, friction values increased slightly up to 0.15 when microtextured pellets of Ekasic F or SN80 were used. Hence, a remarkable reduction of friction dependence on sliding velocity was achieved under these operating conditions by texturing of the pellets with microchannels. Results of stationary slip tests on different materials at different sliding velocities are presented in Fig. 9. Friction coefficient of pairs with untextured and with crossed microchannels textured ceramic pellets was determined after 125 s, i.e. acceleration period of 5 s and 120 s sliding at a constant velocity ranging from 0.10 to 10 m/s. Friction values of untextured pellets increased continuously with sliding velocities up to about 4 m/s (Fig. 9a). Above this limit, friction behaviour was strongly influenced by the ceramic pellet used. With further increasing velocity, friction values decreased more or less owing to cumulative input of friction energy (Fig. 9a). Highest friction coefficient at 125 s of sliding was achieved by the SN80/100Cr6 pair. While friction behaviour of pairs with untextured pellets exhibited a strong dependence on sliding velocity up to about 4 m/s, the pairs with textured pellets showed fluctuations of friction values only in a relative narrow band between about 0.11 and 0.14 over the total range of sliding velocities and at all ceramics tested (Fig. 9b). Friction values were substantially raised by microtexturing at low to medium sliding velocities compared with untextured pellets. Effects of microtexturing and applied normal load are presented at SN80/100Cr6 pairs in Figs. 9c and d. At the relatively low load of 15 N, high friction values were measured with the textured pellets, particularly (Fig. 9c), which reached about twice of the values of the untextured pellet. Increase of normal load to 60 N resulted in substantially lower friction values and no significant differences were obtained between behaviour of untextured and textured pellets in the range from 4 to 10 m/s (Fig. 9d). Owing to strong increase of friction coefficient below about 4 m/s by microtexturing, the pellet textured with crossed microchannels showed friction values which were almost independently of sliding velocity. 3.2.2. Effect of texture parameters Area coverage fractions and widths of microchannels were varied at SN80/100Cr6 pairs between 35 and 75% and 50 and 200 m, respectively. Friction behaviour was studied using both the stationary slip test and the deceleration test, according to Fig. 1b. Fig. 10a shows friction coefficient as function of sliding time at the constant velocity of 10 m/s. Friction values of pairs with untextured SN80 pellets decreased almost continuously from the end
of acceleration period to 125 s of sliding. Pellets with microtexturing led to greater friction whereas the values increased with area coverage fractions substantially. At an area fraction of 75%, friction coefficient was about 0.13 at 125 s of sliding compared with 0.08 with the untextured pellet (Fig. 10a). Effect of the area coverage fraction of microtexture on friction behaviour at deceleration from 10 m/s to zero during 5 s is presented in Fig. 10b. Friction of the untextured pellet decreased to very low values with decreasing sliding velocity (read diagram from right to left). Increasing area fractions of microchannels on textured pellets from 35 over 50–75% resulted in enhanced friction. At 75%, friction values were almost independently of sliding velocity, while at 35% a transition from full film lubrication to boundary lubrication occurred similar as described by the Stribeck curve (Fig. 10b). Fig. 11a shows the effect of the width of crossed microchannels on friction characteristic of the SN80/100Cr6 pairs measured at the stationary slip test. Widths of the channels were varied between 50 and 200 m at an area coverage fraction of 75%. Lower friction values at 125 s of sliding at 10 m/s were measured in the case of greater channel widths at the constant depth of 10 m. With broader channels, e.g. 200 m, a substantially improved constancy of friction coefficient with sliding time was achieved compared with the smaller channel widths. Results from deceleration tests are presented in Fig. 11b. Friction coefficient of pellets with 200 or 150 m broad microchannels increased more or less while values of the untextured pellets decreased with decreasing sliding velocities. Smaller channel widths of 50 or 100 m resulted in greater friction values than those of the broader widths of 150 or 200 m at sliding velocities above 1 m/s and its friction values decreased slightly below sliding velocities of about 2 m/s. 4. Discussion Friction behaviour of different ceramics in pairing with steel 100Cr6 was characterised with respect to applications as friction materials in oil lubricated systems such as fast running multiple disc clutches as one example. Besides the practical-oriented experiments using a pellet/disc configuration at high sliding velocity and medium applied normal load, experiments were carried out using a model test, which gave an insight in development of the lubricating oil film during sliding contact up to 0.30 m/s. An important aspect was to investigate the effect of microtexturing on friction behaviour whereas the results should be useful for designing functional areas both at sliding and friction systems, i.e. for achieving low or high friction values, respectively. However, the present study was primarily focused on an increase of friction coefficient and reduction of dependence of friction values on sliding velocity and sliding time by using microtexturing of the functional area of the pellets. The ceramic used as pellet affected friction behavior in two ways namely the friction level and the constancy of friction coefficient
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Fig. 12. Schematic representation of course of (a) friction coefficient and (b) oil film thickness versus sliding velocity of pairs with untextured (polished or lapped) or textured pellets with crossed microchannels as well as (c) equation for describing effective oil film thickness and (d and e) cross-sections through contact area between pellet and disc at two different film thicknesses.
with prolonged time of sliding at high velocity of 10 m/s at the stationary slip test (Fig. 8a and b). With untextured pellets, Ekasic F showed the highest friction values, which decreased more or less at all three ceramics with sliding time. This decline of the value of friction coefficient was observed only at sliding velocities ≥3 m/s and increased with increasing sliding velocity and sliding time. In other stationary slip tests, not presented here, at equal conditions, the temperature measured by thermocouples 0.15 mm below the contact area of the pellets showed a temperature rise after 125 s of sliding at 10 m/s of T = 35–40 K depending on the ceramic. Hence, it was concluded, that increasing energy input (∼FN vt) owing to increasing time and/or sliding velocity resulted in a rise of oil temperature and therefore owing to lower oil viscosity in decreasing friction coefficient. Ranking of the friction level of the ceramics was changed by microtexturing as it is shown in Fig. 8b. The higher energy input due to higher friction coefficient at textured than untextured pellets affected friction constancy versus time of the pair SN80/100Cr6 less than EKasicF/100Cr6. SN80 pellets textured with crossed microchannels offered the highest values and constancy of friction over the time of running. This was attributed to the greater heat capacity of Al2 O3 –ZrO2 (SN80: 780 J/(kg K)) compared with SSiC (EKasic F: 600 J/(kg K)). Generally, the mineral oil (ISO-VG 100) offers a high heat capacity of 1789 J/(kg K) but a low thermal conductivity of 0.134 W/(m K) and hence friction heat should be removed from the contact via oil flow. It has to be expected that thermal conductivity (SSiC, EKasic F: = 125 W/(m K); Al2 O3 –ZrO2 , SN80: 28 W/(m K)) of the ceramics will influence friction behaviour at longer sliding times. Recently published experimental or theoretical work [18–20] was directed to microtexturing for reduction of friction and for increase in load bearing capacity. Hence, tribological behavior of texture pattern such as isolated circular microcavities was analysed. Hydrodynamic behavior of two parallel flat surfaces in sliding contact is explained [19] by suction of oil into the contact because the pressure inside the microcavities is less than the surrounding atmospheric pressure. Cavitation at the inlet side of the microcav-
ity plays an important role for the load support and inlet suction does not occur when cavitation pressure increases to atmospheric pressure [19]. This mechanism can explain the thicker oil film and lower friction of pairs consisting of parallel surfaces with microcavities. In agreement with this mechanism, Fig. 3 of the present study shows that the circular microdimples on the pellet surface resulted in greater oil film thickness and lower friction values with increasing sliding velocities than the only polished, untextured pellets. Shallow microcavities can lead to a substantial increase in film thickness, while deep microcavities can result in a local decrease or even collapse of the lubricant film [18–20]. This was ascribed to the more restricted lubricant flow from deep than from shallow microcavities [19] or to the different pressure distribution inside the microcavity, because the pressure inside a deep cavity is lower [18]. In contrast to numerous research work on isolated microcavities, the present study was concentrated on the effect of interconnected microchannels where the lubricant flow is less constricted. According to Fig. 3, circular microdimples increased but crossed microchannels decreased oil film thickness compared with untextured pellets. Fig. 4 shows that the orientation of parallel microchannels influenced film thickness and friction coefficient owing to eased or constrained oil drain out of the contact area. Microtexturing results in reduction of load bearing area A by the factor (1 − atext. ), since inside the microchannels exists only a low pressure, as mentioned above for circular cavities and in addition due to the low constraint of oil drain along the channels, and hence load bearing capacity is low at the area fraction of channels. The oil drain can be impeded by side wall friction inside the channel, reduced cross-section of the channels and greater oil viscosity. Increase of normal load from 15 N to 60 N (Fig. 9c and d) resulted in an about fourfold increase of friction force on the untextured but only in a twofold increase on the textured pellet at higher sliding velocities. From analysis of the experimental data at 15 N and under the assumption of hydrodynamic lubrication at v ≥about 5 m/s and Newtonian behavior, it was concluded that microtexturing by
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medium sized channels led to an increase in the effective values of film thickness and viscosity compared with the untextured pair. This apparent increase in viscosity and the enhanced friction force was caused by non-laminar oil flow and resulting hydrodynamic losses. Change of the normal load to 60 N led to reduction of oil film thickness at the textured pellet and to transition to common hydrodynamic behaviour with friction values comparable to the untextured pellet at high sliding velocities (Fig. 9d). At low sliding velocities, texturing by microchannels increased friction coefficient because reduction of oil film thickness led to transition from liquid to mixed or boundary friction. As result, friction coefficient was almost independently of sliding velocity at the load of 60 N (Fig. 9d). Oil film thickness became smaller with increasing area coverage fractions of microchannels and hence this transition was shifted to higher sliding velocities (Fig. 6). Results from the application-oriented friction test rig showed that not only friction values were raised with greater area fractions but also friction course versus sliding velocity was substantially altered (Fig. 10b). At a fraction of 75% of microchannels, friction values were almost independently of sliding velocity and the decline of friction coefficient with time of sliding at 10 m/s was reduced (Fig. 10a). Depth and width of microchannels were other parameters of microtexturing which influenced strongly film thickness and friction coefficient (Fig. 7). Increasing depths and widths of the channels reduced oil film thickness. Hence, friction values were raised and the transition to mixed or boundary lubrication occurred already at higher sliding velocities. This transition in lubrication conditions at larger channel widths (B = 150 or 200 m) was also confirmed by results from the friction test rig using the deceleration test (Fig. 11b). At higher sliding velocities however, smaller channel widths of 50 or 100 m resulted in higher friction values, which were almost independently of velocity (Fig. 11). Smaller channel widths, e.g. 60 m in the model test (Fig. 7a and b) and 50 or 100 m in the deceleration test (Fig. 11b) led to greater thicknesses of the oil film and lower friction values at low sliding velocities ≤0.5 m/s. Low depths and small widths of the microchannels promoted hydrodynamic lubrication and hence similar friction and lubrication conditions as the pairs with untextured pellets (Fig. 7). Fig. 12 shows a descriptive model developed from the experimental results and the support by in situ observation of the contact area in the model test using a high resolution microscope. The effective oil film thickness h* consists to a first approximation (Fig. 12c) of different parts: h0 = separation distance between pellet and disc without lubricant and normal load; which is a function (f) of (alignment deviation, surface roughness of contact area etc.), hstat = change of separation distance (film thickness) at lubricated static contact; f (oil charge and viscosity, normal load, capillary effects, wetting behavior etc.), htext = change of film thickness due to microtexturing at lubricated sliding contact; f (texture pattern, depth T, width B, area coverage fraction atext. , etc.) and hhydro = change of film thickness under hydrodynamic lubrication; f (oil viscosity, sliding velocity, characteristic length in sliding direction, normal load etc.). According to hydrodynamic theory, oil film thickness of pairs with untextured, only polished pellets increases continuously with sliding velocity and decreases with increasing normal load (Figs. 5b and 12b). Development of a load bearing oil film can be supported by the inlet suction of oil [19] between the parallel contact surfaces of pellet and disc. In the hydrodynamic region, microtexturing showed relatively little effect on friction behavior (Fig. 9a, b and d), but lowering the normal load from 60 to 15 N results in strong decrease of friction force by a factor of about four at untextured but only by the factor of two at textured pellets (Fig. 9c). With decreasing sliding velocity or increasing normal load the dominance of the hydrodynamic part hhydro on the effective film thickness h* decreases and the effect
of microtexturing htext gains influence on the slope of both film thickness and friction curve versus sliding velocity (Fig. 12a and b). Decreasing film thickness with decreasing sliding velocity, increasing normal load and/or increasing area fraction of microtexture lead to transition from completely oil filled to only partially oil filled microchannels and friction due to oil flow inside the channels when film thickness h* becomes smaller than the channel depth (Fig. 12d and e). As mentioned above completely oil filled microchannels offer only a limited load carrying capacity that can be enhanced by increased constraint of oil drain through the channels. Load carrying capacity of only partially oil filled channels goes to zero, especially at easy oil drain through broad and deep channels. This leads to a steep decrease in film thickness compared with the untextured pellets (Fig. 12b). Further decrease of film thickness with decreasing sliding velocity or increasing normal load as well as area coverage fraction, depth and width of microchannels can result in transition to mixed or even boundary lubrication, i.e. breakthrough of oil film and asperities interactions on the mated solids, when the value of h* approaches the sum of roughness values Rq of pellet and disc surface. This means the specific film thickness becomes smaller than 1 and value of friction coefficient increases (Fig. 12a and b). Hence, curves of oil film thickness and friction coefficient of pairs with textured pellets can run with decreasing sliding velocity over three different regions characterised by different slopes from completely oil filled channels (h > T) to only partially filled channels (h > T) at small oil film thickness and finally to boundary lubrication (h ≈ Rq ). Friction course versus sliding velocity results from superposition of solid friction, friction due to oil flow inside the channels and around the pellet as well as hydrodynamically induced full film friction. According to the model, depth of the texture pattern is a very important factor with respect to friction and load carrying capacity. 5. Conclusions Experiments showed that microtexturing of functional areas can be very useful to adapt tribological behavior to requirements of a friction system. Results from a model test and those from a friction test rig oriented to applications such as fast running oil lubricated friction systems, e.g. multiple disc clutches, led to a deeper understanding of the effect of texturing by microchannels. Comparing sliding pairs containing pellets textured by crossed microchannels (CH) with untextured pellets, it can be concluded that - oil film thickness is reduced by texture parameter such as increasing area coverage fraction, depth and width of microchannels (CH) as well as decreasing sliding velocity and increasing normal load. - crossed microchannels increase friction coefficient at medium to low sliding velocities due to oil flow inside the channels and transition to mixed lubrication. - curves of oil film and friction coefficient versus sliding velocity run at textured pellets (CH) over three different regions characterised by different slopes with decreasing film thickness, namely from full film lubrication with completely oil filled channels to partially oil filled channels with oil flow friction and to mixed or boundary lubrication when film thickness approaches roughness of the mated surfaces, according to the descriptive model presented. - crossed microchannels result in greater friction coefficient but lower oil film thickness than circular microdimples (CD). Acknowledgements The authors would like to thank the Deutsche Forschungsgemeinschaft (DFG) for financial support within the frame of the
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Collaborative Research Centre 483 “High performance sliding und friction systems based on advanced ceramics”. References [1] E.P. Becker, Trends in tribological materials and engine technology, Tribology International 37 (2004) 569–575. [2] P. Zagrodzki, S.A. Truncone, Generation of hot spots in a wet multidisk clutch during short-term engagement, Wear 254 (2003) 474–491. [3] D.H. Hwang, K.-H. Zum Gahr, Tribological behaviour of steel-alumina sliding pairs under boundary lubrication conditions, Tribotest Journal 9–3 (2003) 209–217. [4] J. Qu, J.J. Truhan, P.J. Blau, Investigation of the scuffing characteristics of candidate materials for heavy duty diesel fuel injectors, Tribology International 38 (2005) 381–390. [5] E. Peter, K.-H. Zum Gahr, Influence of surface finish and sliding speed on oscillating sliding wear of laser modified oxide ceramics in aqueous media, Mat.-Wiss. u. Werkstofftech. 36 (2005) 129–135. [6] M. Scholle, Hydrodynamical modelling of lubricant friction between rough surfaces, Tribology International 40 (2007) 1004–1011. [7] P. Nyman, et al., Influence of surface topography on friction characteristics in wet clutch applications, Wear 261 (2006) 46–52. [8] Y. Tejima, A. Ishiyama, A. Ura, The influence of rounded pores diameter and depth at sliding faces on the coefficient of friction under oil lubrication, Japanese Journal of Tribology 44 (5) (1999) 531–545. [9] A.V. Ronen, I. Etsion, Y. Kligerman, Friction reducing surface-texturing in reciprocating automotive components, Tribology Transactions 44 (3) (2001) 359–366.
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[10] X. Wang, K. Kato, K. Adachi, The lubrication effect of micro-pits on parallel sliding faces of SiC in water, Tribology Transactions 45 (3) (2002) 294– 301. [11] M. Wakuda, Y. Yamauchi, S. Kanzaki, Y. Yasuda, Effect of surface texturing on friction reduction between ceramic and steel materials under lubricated sliding contact, Wear 254 (2003) 356–363. [12] I. Etsion, State of the art in laser surface texturing, Transactions of the ASME 127 (2005) 248–253. [13] S. Schreck, K.-H. Zum Gahr, Laser-assisted structuring of ceramic and steel surfaces for improving tribological properties, Applied Surface Science 247 (2005) 616–622. [14] A. Kovalchenko, et al., The effect of laser surface texturing on transitions in lubrication regimes during unidirectional sliding contact, Tribology International 38 (2005) 219–225. [15] H.L. Costa, I.M. Hutchings, Hydrodynamic lubrication of textured steel surfaces under reciprocating sliding conditions, Tribology International 40 (2007) 1227–1238. [16] K.H. Zum Gahr, M. Mathieu, B. Brylka, Friction control by surface engineering of ceramic sliding pairs in water, Wear 263 (2007) 920–929. [17] P. Andersson, et al., Microlubrication effect by laser-textured steel surfaces, Wear 262 (2007) 369–379. [18] L. Mourier, et al., Transient increase of film thickness in micro-textured EHL contacts, Tribology International 39 (2006) 1745–1756. [19] M. Fowell, et al., Entrainment and inlet suction: two mechanisms of hydrodynamic lubrication in textured bearings, Transactions of the ASME 129 (2007) 336–347. [20] I. Krupka, M. Hartl, Experimental study of microtextured surfaces operating under thin film EHD lubrication conditions, Transactions of the ASME 129 (2007) 502–508.
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Experimental study of the effect of microtexturing on oil lubricated ceramic/steel friction pairs K.-H. Zum Gahr ∗ , R. Wahl, K. Wauthier Universität Karlsruhe, Institute of Materials Science and Engineering II and Forschungszentrum Karlsruhe, Institute for Materials Research I, P.O. Box 3640, 76185 Karlsruhe, Germany
a r t i c l e
i n f o
Article history: Received 1 September 2008 Received in revised form 15 December 2008 Accepted 15 December 2008 Keywords: Microtexturing Friction Oil lubrication Film thickness Advanced ceramic In situ observation
a b s t r a c t The objective of the research work was to study the potential of commercial ceramics such as Al2 O3 , Al2 O3 –ZrO2 and SiC in pairing with steel 100Cr6 and laser-assisted deterministic microtexturing of the flat functional surfaces as design element for fast running friction systems, such as multiple disc clutches, under unidirectional oil lubricated sliding. A more fundamental part of the study was oriented towards understanding of the effect of microtexturing on mechanisms at mixed or boundary lubrication and development of lubricating oil film in the contact area. Experiments were carried out with pellet/disc geometry at relatively low velocities ≤0.3 m/s and normal loads ≤10 N using a model test with an attached microscope for in situ observation of the contact area. In the second part, friction behaviour of the different ceramic materials with and without microtexturing was characterised using a friction test rig at sliding velocities up to 10 m/s and normal loads ≤60 N under more practical-oriented conditions. Results showed that frictional behaviour was significantly influenced by the ceramic material and features of microtexturing, e.g. dimples, channels, width, depth and area coverage fraction. Improved understanding of operating mechanisms and design rules of microtexturing for controlling friction coefficient at practical sliding systems was obtained from a descriptive model. © 2009 Elsevier B.V. All rights reserved.
1. Introduction There is a great demand for new materials and improved design of functional surfaces at many sliding and friction systems owing to requirements such as handling of liquids of low viscosity at hermetically sealed pumps or increased power transmission in the power train of automotive engines. Advanced multiple disc clutches are an example of an oil lubricated friction system whereas currently used friction materials meet limits of mechanical and thermal load capacity [1,2]. One reason are the so-called hot spots which are formed locally on the frictional surfaces due to high pressures and high temperatures. Even an excess of the critical sliding velocity for hot spotting during a single short engagement can lead to failure or degradation of the friction materials and as a consequence of the clutch. New friction materials have to fulfil high demands on a sufficiently high and stable friction coefficient for power transmission, running comfort under varying operating conditions with sliding velocities up to about 30 m/s as well as high wear resistance and thermal stability over a long service time. Advanced ceramic materials such as alumina or silicon carbide offer a high potential under such severe loading conditions.
∗ Corresponding author. Tel.: +49 7247 82 3897; fax: +49 7247 82 4567. E-mail address: [email protected] (K.-H. Zum Gahr). 0043-1648/$ – see front matter © 2009 Elsevier B.V. All rights reserved. doi:10.1016/j.wear.2008.12.108
Greater friction values and hence greater power transmission can be achieved under oil lubrication with ceramic/ceramic or steel/ceramic than for example with steel/steel sliding pairs [3]. High scuffing resistance of ceramics is beneficial under lubrication with liquids of low viscosity or at mixed and boundary lubrication at more or less fractions of solid/solid contact [4,5]. Friction and wear are strongly influenced under poor lubrication by topography of the functional surfaces [5–7]. With decreasing film thickness of the liquid a transition to boundary friction can occur depending on the operating conditions and thus tribochemical reactions may be caused on materials such as Al2 O3 or SiC or even phase transformations in the surface of partially stabilised ZrO2 ceramics. Tribochemical films produced at the loaded surface dominate friction and wear characteristic under these conditions. Surface microtexturing of the functional areas can take effect in improving tribological performance of lubricated systems or in friction control under varying operating conditions. Several studies showed that texture pattern of dimples or grooves may cause effects such as removal of wear debris from the contact, improvement of running-in behaviour, improvement of wetting behaviour, action as lubricant reservoirs, increase of load bearing capacity or promotion of hydrodynamic lubrication [8–17]. All these effects can lead to a substantial reduction of friction in sliding systems. It was also reported that shape [15], size [8,9,11,15], area coverage fraction [11,15] or orientation of the texture elements [15] as well as operat-
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Fig. 1. Schematic description of experimental set-up, test parameters, materials and lubrication of (a) model test and (b) friction test rig.
ing conditions or lubricant properties [14,17] play an important role for film thickness of the lubricating liquid and friction behaviour of the system. Also transitions between liquid and mixed or boundary friction may be affected by microtexturing [14]. However, the effectiveness differs substantially and in some cases texturing showed no significant or even a detrimental effect on friction coefficient and load bearing capacity depending on the mated materials and experimental conditions used. The reason is a complex nature of contact phenomena occurring between the surfaces and the liquid in contact. Trustworthy results can be expected at parallel flat contacts but point or line contacts may cause beneficial or detrimental effects on friction, since ratio of scale of contact and texture pattern may become crucial. It follows, that much more knowledge is needed to design an optimised microtexturing for a special application. Aim of the present study was to achieve a deeper knowledge of the mechanisms and most important factors influencing the frictional contact of microtextured surfaces of oil lubricated ceramic/steel pairs during unidirectional sliding. Contact conditions were analysed by using high resolution experimental methods and results were used for description of the frictional contact as a function of influencing factors such as materials mated, parameters of texture pattern, load, sliding velocity or film thickness of lubricating oil. 2. Materials and experimental methods
16 mm for these tests were fabricated from the commercial ceramics SSiC (EKasicF, ESK Ceramics), Al2 O3 (HTC 99.9, Hightech Ceram) and Al2 O3 –15 vol.% ZrO2 (SN80, CeramTec). A flat circular contact area of Ø 6 mm was lapped at the convex (radius of 100 mm) end face of the ceramic pellets using directly the tribological test rig for reducing misalignment. Properties and micrographs of the different materials are given in Table 1 and Fig. 2, respectively. 2.2. Microtexturing of pellets Laser-assisted microtexturing was carried out on the functional area of the pellets by using an Ytterbium-Fibrelaser (ACSYS Lasertechnik) at a wave length of 1.064 m and a power of about 20 W depending on the material. Deterministic texture patterns (Fig. 2) of circular microdimples (CD) or microchannels parallel or crossed to each other (CH) were produced on the lapped or polished areas of the pellets by laser ablation process. Texture parameters such as width, depth, offset and area coverage fraction were varied in different test series. Width of channels was varied from 50 to 300 m depending on the test. Larger widths of the channels were generated by putting several tracks side by side with a 10 m offset. Depth was varied between 5 and 20 m. Values of area coverage fractions accounted for 35, 50 and 75%. Offset values resulted from the width and area coverage fraction selected (Fig. 2c). After laser ablation the texture patterns showed debris, which was removed using gentle grinding and polishing.
2.1. Materials and specimens 2.3. Tribological test Tribological behaviour was studied at unidirectional sliding using a pellet-on-disc geometry. As delivered pellets of the steel 100Cr6 were used through hardened and tempered with a hardness of 790 ± 10 HV. A flat circular contact area of Ø 7.2 mm was machined at the convex (radius of 8 mm) end face. The contact area of the pellet was finally polished to Ra = 0.01 m and mated with a polished sapphire disc (GWI Sapphire) during tests using a model tribometer (Fig. 1a). On the other hand, ceramic pellets were mated with normalized discs of the steel 100Cr6 (300 ± 10 HV) in tests using a larger friction test rig (Fig. 1b). The cylindrical pellets of Ø
The model test and the friction test rig differed in the manner that hardened and tempered 100Cr6 steel pellets were mated with sapphire discs at the model test while pellets fabricated from different ceramics (HTC 99.9, SN80, EKasic F) were mated with normalized 100Cr6 steel discs at the friction test rig (Fig. 2). Using the commercial friction test rig (Fig. 1b; UMT3, CETR), oil lubricated tests were carried out at unidirectional sliding up to relative velocities of 10 m/s and applied pressures of 2.12 MPa calculated for untextured circular contact area of Ø 6 mm and nor-
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Table 1 Marking and properties of materials used for the model test and friction test rig. Marking
Material
Specimen
Grain size (m)
Density (103 kg/m3 )
Hardness (HV500 )
Young’s modulus (GPa)
Sapphire 100Cr6 HTC 99.9 SN80 EKasic F 100Cr6
Al2 O3 -single crystal Hardened 99.9 wt.%–Al2 O3 Al2 O3 –15 vol.% ZrO2 SSiC Normalized
Disc Pellet Pellet Pellet Pellet Disc
– – 1.0–3.0 1.5 1.9 –
3.98 7.84 3.96 4.10 3.17 7.84
2200 790 2084 1610 2540 300
350 212 380 380 410 212
Fig. 2. (a–c) Schematic description of geometry and size of texture pattern as well as SEM micrographs of (d) microtextured surfaces of steel 100Cr6 or Ekasic F and (e) polished microstructure of the ceramic materials used.
mal load of FN = 60 N. Textured or untextured ceramic pellets were mated with the directly in the tribometer lapped contact area with fine ground discs of the pearlitic 100Cr6 steel (Ra = 0.10 m, 300 ± 10 HV). Mineral oil (ISO-VG 100, FVA No. 3, 40◦ C = 85.2 mPa s) without additives was fed into the contact area by drip lubrication. The stationary slip test (Fig. 1b) was characterized in that way that
after loading the friction pair by the normal load, sliding velocity was continuously increased during 5 s up to the value wanted and it was hold constant for 120 s before the velocity was shut down to zero. Each test at a given constant velocity (varied from 0.10 to 10 m/s), was followed by a cooling period down to room temperature before the run for the next greater velocity was started. At
Fig. 3. Polished or microtextured steel pellets mated with polished sapphire discs at the model test (a) friction coefficient and (b) oil film thickness h* versus sliding velocity (CD: circular dimples, CH: crossed microchannels and pol.: untextured, polished (FN = 5 N, atext. = 35%, B = 100 m, T = 10 m)).
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Fig. 4. (a) Friction coefficient and (b) oil film thickness of 100Cr6 pellets with three different oriented (A, B = parallel and C = crossed) microchannels as well as untextured surface (P) mated with sapphire discs (FN = 5 N, v = 0.30 m/s, atext. = 50%, B = 100 m, T = 10 m).
the deceleration test friction behaviour was measured during the slow down period lasting 5 s from a sliding velocity of 10 m/s down to zero. All tests were run in laboratory air at room temperature of 22 ◦ C and 50% relative humidity. Reported values were taken from representative runs. Studies using the model test (Fig. 1a) were run at low velocities and low normal loads up to 0.3 m/s and 10 N, respectively. Martensitic 100Cr6 steel pellets (790 ± 10 HV) with polished or microtextured contact areas of Ø 7.2 mm were mated at oil lubricated unidirectional sliding with polished sapphire discs. The same mineral oil as used in the friction test rig was fed at 2.0 mm3 /s by drip lubrication in front of the contact area. Besides friction force, separation distance h* between the contact area of steel pellet and sapphire disc was continuously measured during each test by using a high accuracy distance transmitter with a resolution of ±0.5 m. Recorded values of separation distance “h* ” represent a measure of
the oil film thickness “h” but it has to be expected that these values can be greater than the real oil film thickness “h” because waviness of the sapphire disc (<1 m after manufacturer data) or alignment variations will contribute to the measured data. In situ observation of the contact area during sliding was possible using a special microscope with magnification up to 200 fold. It was preferentially applied for studying the development of the lubricant film and the oil flow at microtextured contact areas.
3. Results Oriented towards the aim to improve the understanding of effects of microtexturing, model tests were carried out and the transferability of their results to more application-oriented loading conditions was proved by tests using the friction test rig.
Fig. 5. Effect of normal load and sliding velocity on friction coefficient and oil film thickness at 100Cr6/sapphire pairs with (a and b) untextured and (c and d) crossed microchannels textured steel pellets (atext. = 75%, B = 100 m, T = 10 m).
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Fig. 6. (a) Friction coefficient and (b) oil film thickness of 100Cr6/sapphire pairs with untextured (pol) pellets or pellets textured with crossed microchannels of 35, 50 or 75% versus sliding velocity (CH35–CH75: atext. = 35–75%, B = 100 m, T = 10 m, FN = 5 N).
3.1. Model test Friction coefficient and oil film thickness of hardened 100Cr6 steel pellets with and without microtextured functional area mated with sapphire discs were measured in the model test as function of loading and microtexturing parameters. 3.1.1. Texture pattern Fig. 3 shows friction coefficient and thickness of oil film of 100Cr6/sapphire pairs with polished or microtextured steel pellets as function of sliding velocity. At the normal load of 5 N, depth of texture pattern of 10 m and area coverage fraction atext. = 35%, lower values of film thickness and greater values of friction coefficient were measured at pellets textured with crossed microchannels (CH) than with circular microdimples (CD) or untextured only polished surface. Both film thickness and friction coefficient of pairs with CD-textured or untextured pellets increased continuously with increasing sliding velocity. The pellet textured with microdimples showed a greater slope of friction with
sliding velocity than the untextured pellet (Fig. 3b) that pointed on a hydrodynamic effect. However, texturing with microchannels resulted in a minimum of friction coefficient at about 0.10 m/s under the experimental conditions used (Fig. 3a), i.e. a behaviour similar to that described by the “Stribeck curve”. Friction coefficient increased with decreasing sliding velocity below this value in agreement with the simultaneously decreasing film thickness to very low values down to the scale of surface roughness of the mated specimens (Fig. 3b). This means that a transition from full film lubrication to partial (or mixed) lubrication and then to boundary lubrication occurred with decreasing sliding velocity. Since higher friction values are advantageous for applications in oil lubricated friction systems, following studies were focused on effects of microchannels. 3.1.2. Orientation of microchannels Fig. 4 shows friction coefficient and film thickness of 100Cr6/sapphire pairs as function of the orientations of parallel microchannels (A, B) compared with crossed microchannels (C) and
Fig. 7. (a) Friction coefficient and (b) oil film thickness of 100Cr6/sapphire pairs with untextured (pol) pellets or pellets textured with crossed microchannels of (a and b) depth T = 5, 10 or 20 m at the width of 100 m and (c and d) width B = 60, 100 or 300 m at the depth of 10 m versus sliding velocity (atext. = 50%, FN = 5 N).
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Fig. 8. Friction coefficient of different ceramic pellets (a and c) untextured and (b and d) textured with crossed microchannels mated with normalized 100Cr6 steel discs (a and b) versus sliding time, i.e. 5 s acceleration and 120 s at 10 m/s (stationary slip test) and (c and d) versus sliding velocity (deceleration test) at normal load of 60 N (CH: atext. = 75%, B = 100 m, T = 10 m).
without texturing (P) of the steel pellets. Orientation (A) led to eased oil drain and orientation (B) to constrained oil drain with respect to the sliding direction of the rotating sapphire disc. It is expected that a situation of constrained oil drain out of the sliding contact can benefit hydrodynamic lubrication. This is supported by the greater
film thickness and lower friction coefficient measured at 0.30 m/s and 5 N normal load for the orientation (B) than for (A). Values of film thickness measured at the pellet with crossed microchannels (C) ranged between (A) and (B) (Fig. 4b) but friction coefficient was greater than (A) and (B) (Fig. 4a). In comparison, the untextured
Fig. 9. Friction coefficient at 125 s of sliding versus sliding velocity of (a and b) 100Cr6 discs mated with HTC 99.9, SN80 or Ekasic F pellets, (a) untextured and (b) textured, at FN = 60 N as well as SN80/100Cr6 pairs with untextured or textured pellets at (c) normal load of 15 N and (d) 60 N, respectively (CH: atext. = 75%, B = 100 m, T = 10 m, stationary slip test at different sliding velocities).
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Fig. 10. Friction coefficient of SN80 pellets textured with crossed microchannels of area fractions of 35, 50 and 75% or untextured as reference mated with 100Cr6 discs versus (a) sliding time at the stationary slip test, i.e. 5 s acceleration and 120 s at 10 m/s, as well as (b) sliding velocity from 10 m/s to zero at the deceleration test (B = 100 m, T = 10 m, FN = 60 N).
polished pellet resulted in relatively high film thickness and low friction. 3.1.3. Normal load and sliding velocity Effects of normal load and sliding velocity on both friction coefficient and oil film thickness are shown in Fig. 5 at 100Cr6/sapphire pairs with untextured pellets and pellets textured with crossed microchannels of atext. = 75%. In case of untextured pellets, friction coefficient and oil film thickness increased continuously with sliding velocity but decreased with increasing normal load (Fig. 5a and b). The relation between film thickness and loading parameters (FN , v) was not changed by texturing with crossed microchannels, however, the values were substantially smaller (Fig. 5d). In contrast to this, friction behaviour was substantially altered by the texture in the way that friction coefficient increased with textured pellets with decreasing sliding velocity (Fig. 5c). This was caused by transition from liquid to mixed or boundary friction. The increasing friction with decreasing sliding velocity correlates very well with the measured reduction in film thickness down to very low values (Fig. 5d). 3.1.4. Area coverage fraction Fig. 6 shows the influence of area coverage fraction of crossed microchannels on friction coefficient and oil film thickness of 100Cr6/sapphire pairs. Independently of the area fraction of microchannels and sliding velocity, oil film thickness was smaller at textured than untextured pellets. Increasing atext. from 35 to 75% resulted in decreasing film thickness (Fig. 6b). Film thickness values falling below a limit of about 0.5 m led to increasing friction coefficient with further decreasing sliding velocities (Fig. 6a) due to transition to partial lubrication because of breakthrough of the oil film and partial solid/solid contact between the mated surfaces. The value of sliding velocity at this transition was shifted to greater velocities with increasing area fraction of the microchannels (Fig. 6a). 3.1.5. Depth and width of texture pattern Depth and width of the crossed microchannels are texture variables which can be selected system-specific for optimizing friction behaviour. Fig. 7a and b shows the effect of the depth of crossed microchannels on friction coefficient and oil film thickness of the 100Cr6/sapphire pairs at an area coverage fraction of atext. = 50% and a channel width of 100 m. Pellets textured with shallow microchannels of the depth of 5 m showed qualitatively similar friction behaviour as function of sliding velocity as untextured pellets (Fig. 7a). Below a sliding velocity of about 0.15 m/s, values of film thickness and friction coefficient approached those of the untextured pellets, under the applied experimental conditions. Both friction and film thickness curve were substantially altered
at sliding velocities below about 0.15 m/s when the channel depth was enhanced from 5 to 10 m. At 10 m depth, transition to partial lubrication occurred accompanied by increasing friction values. Pellets with microchannels of 20 m depth run under partial or boundary lubrication and high values of friction were measured at all sliding velocities (Fig. 7a). Friction coefficient and oil film thickness of 100Cr6/sapphire pairs with pellets textured with crossed microchannels of different widths of 60, 100 and 300 m but constant depth of 10 m are given in Fig. 7c and d. Pellets with narrow channels of 60 m showed similar values of friction coefficient and oil film thickness as untextured pellets in the velocity range tested. Larger widths of 100 and 300 m decreased substantially film thickness and led to transition from liquid to mixed friction (Fig. 7c). Both the curve of friction coefficient and film thickness of pellets with the very broad channels of 300 m deviated from those of 100 m at sliding velocities below about 0.10 m/s. Results of friction coefficient and film thickness of the pairs with pellets textured by microchannels of the largest depth of 20 m or largest width of 300 m point at possibilities to increase friction coefficient and to reduce dependence of friction coefficient on sliding velocity. 3.2. Friction test rig Tribological tests were run using two different test procedures namely a stationary slip test and a deceleration test for characterisation of frictional behaviour of commercial ceramics as pellets mated with normalized 100Cr6 discs. The tests were focused on application of advanced ceramics with and without microtexturing as materials in oil lubricated friction systems, such as multiple disc clutches. 3.2.1. Materials and operating conditions Fig. 8 shows friction coefficient of pellets fabricated from oxide and non-oxide ceramics (HTC 99.9: Al2 O3 , SN80: Al2 O3 –ZrO2 and Ekasic F: SSiC) mated with normalized 100Cr6 steel discs as function of sliding time or sliding velocity. Using the stationary slip test (Fig. 1b), friction coefficient was measured during 120 s sliding at 10 m/s (Fig. 8a and b). After the acceleration period to 10 m/s during the first 5 s, values of friction coefficient decreased more or less strongly with prolonged time of sliding. Friction coefficient of the pairs with untextured pellets was between 0.08 and 0.10 at the end of the tests, whereas the greatest value was measured at EKasic F and the lowest at HTC 99.9 (Fig. 8a). Microtexturing of the pellets with crossed microchannels of atext. = 75% enhanced friction values substantially, whereby a value of 0.15 was achieved with SN80 pellets at the end of the test (Fig. 8b). Friction behaviour of the different pairs was also characterised using the deceleration test, i.e. sliding velocity was slowed down
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Fig. 11. Friction coefficient of SN80 pellets textured with crossed microchannels of widths varied from 50 to 200 m mated with 100Cr6 discs versus (a) sliding time, i.e. 5 s acceleration and 120 s at 10 m/s (stationary slip test) and (b) versus sliding velocity from 10 m/s to zero (deceleration test) at the normal load of 60 N (atext. = 75%, T = 10 m).
from 10 m/s to zero during 5 s (Fig. 1b). In the case of untextured pellets, friction coefficient decreased continuously with decreasing sliding velocity (Fig. 8c). However, friction coefficient of pairs with microtextured pellets stayed almost constant with decreasing velocity at values between about 0.11 and 0.13 depending on the ceramic used (Fig. 8d). Below sliding velocities of about 2 m/s, friction values increased slightly up to 0.15 when microtextured pellets of Ekasic F or SN80 were used. Hence, a remarkable reduction of friction dependence on sliding velocity was achieved under these operating conditions by texturing of the pellets with microchannels. Results of stationary slip tests on different materials at different sliding velocities are presented in Fig. 9. Friction coefficient of pairs with untextured and with crossed microchannels textured ceramic pellets was determined after 125 s, i.e. acceleration period of 5 s and 120 s sliding at a constant velocity ranging from 0.10 to 10 m/s. Friction values of untextured pellets increased continuously with sliding velocities up to about 4 m/s (Fig. 9a). Above this limit, friction behaviour was strongly influenced by the ceramic pellet used. With further increasing velocity, friction values decreased more or less owing to cumulative input of friction energy (Fig. 9a). Highest friction coefficient at 125 s of sliding was achieved by the SN80/100Cr6 pair. While friction behaviour of pairs with untextured pellets exhibited a strong dependence on sliding velocity up to about 4 m/s, the pairs with textured pellets showed fluctuations of friction values only in a relative narrow band between about 0.11 and 0.14 over the total range of sliding velocities and at all ceramics tested (Fig. 9b). Friction values were substantially raised by microtexturing at low to medium sliding velocities compared with untextured pellets. Effects of microtexturing and applied normal load are presented at SN80/100Cr6 pairs in Figs. 9c and d. At the relatively low load of 15 N, high friction values were measured with the textured pellets, particularly (Fig. 9c), which reached about twice of the values of the untextured pellet. Increase of normal load to 60 N resulted in substantially lower friction values and no significant differences were obtained between behaviour of untextured and textured pellets in the range from 4 to 10 m/s (Fig. 9d). Owing to strong increase of friction coefficient below about 4 m/s by microtexturing, the pellet textured with crossed microchannels showed friction values which were almost independently of sliding velocity. 3.2.2. Effect of texture parameters Area coverage fractions and widths of microchannels were varied at SN80/100Cr6 pairs between 35 and 75% and 50 and 200 m, respectively. Friction behaviour was studied using both the stationary slip test and the deceleration test, according to Fig. 1b. Fig. 10a shows friction coefficient as function of sliding time at the constant velocity of 10 m/s. Friction values of pairs with untextured SN80 pellets decreased almost continuously from the end
of acceleration period to 125 s of sliding. Pellets with microtexturing led to greater friction whereas the values increased with area coverage fractions substantially. At an area fraction of 75%, friction coefficient was about 0.13 at 125 s of sliding compared with 0.08 with the untextured pellet (Fig. 10a). Effect of the area coverage fraction of microtexture on friction behaviour at deceleration from 10 m/s to zero during 5 s is presented in Fig. 10b. Friction of the untextured pellet decreased to very low values with decreasing sliding velocity (read diagram from right to left). Increasing area fractions of microchannels on textured pellets from 35 over 50–75% resulted in enhanced friction. At 75%, friction values were almost independently of sliding velocity, while at 35% a transition from full film lubrication to boundary lubrication occurred similar as described by the Stribeck curve (Fig. 10b). Fig. 11a shows the effect of the width of crossed microchannels on friction characteristic of the SN80/100Cr6 pairs measured at the stationary slip test. Widths of the channels were varied between 50 and 200 m at an area coverage fraction of 75%. Lower friction values at 125 s of sliding at 10 m/s were measured in the case of greater channel widths at the constant depth of 10 m. With broader channels, e.g. 200 m, a substantially improved constancy of friction coefficient with sliding time was achieved compared with the smaller channel widths. Results from deceleration tests are presented in Fig. 11b. Friction coefficient of pellets with 200 or 150 m broad microchannels increased more or less while values of the untextured pellets decreased with decreasing sliding velocities. Smaller channel widths of 50 or 100 m resulted in greater friction values than those of the broader widths of 150 or 200 m at sliding velocities above 1 m/s and its friction values decreased slightly below sliding velocities of about 2 m/s. 4. Discussion Friction behaviour of different ceramics in pairing with steel 100Cr6 was characterised with respect to applications as friction materials in oil lubricated systems such as fast running multiple disc clutches as one example. Besides the practical-oriented experiments using a pellet/disc configuration at high sliding velocity and medium applied normal load, experiments were carried out using a model test, which gave an insight in development of the lubricating oil film during sliding contact up to 0.30 m/s. An important aspect was to investigate the effect of microtexturing on friction behaviour whereas the results should be useful for designing functional areas both at sliding and friction systems, i.e. for achieving low or high friction values, respectively. However, the present study was primarily focused on an increase of friction coefficient and reduction of dependence of friction values on sliding velocity and sliding time by using microtexturing of the functional area of the pellets. The ceramic used as pellet affected friction behavior in two ways namely the friction level and the constancy of friction coefficient
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Fig. 12. Schematic representation of course of (a) friction coefficient and (b) oil film thickness versus sliding velocity of pairs with untextured (polished or lapped) or textured pellets with crossed microchannels as well as (c) equation for describing effective oil film thickness and (d and e) cross-sections through contact area between pellet and disc at two different film thicknesses.
with prolonged time of sliding at high velocity of 10 m/s at the stationary slip test (Fig. 8a and b). With untextured pellets, Ekasic F showed the highest friction values, which decreased more or less at all three ceramics with sliding time. This decline of the value of friction coefficient was observed only at sliding velocities ≥3 m/s and increased with increasing sliding velocity and sliding time. In other stationary slip tests, not presented here, at equal conditions, the temperature measured by thermocouples 0.15 mm below the contact area of the pellets showed a temperature rise after 125 s of sliding at 10 m/s of T = 35–40 K depending on the ceramic. Hence, it was concluded, that increasing energy input (∼FN vt) owing to increasing time and/or sliding velocity resulted in a rise of oil temperature and therefore owing to lower oil viscosity in decreasing friction coefficient. Ranking of the friction level of the ceramics was changed by microtexturing as it is shown in Fig. 8b. The higher energy input due to higher friction coefficient at textured than untextured pellets affected friction constancy versus time of the pair SN80/100Cr6 less than EKasicF/100Cr6. SN80 pellets textured with crossed microchannels offered the highest values and constancy of friction over the time of running. This was attributed to the greater heat capacity of Al2 O3 –ZrO2 (SN80: 780 J/(kg K)) compared with SSiC (EKasic F: 600 J/(kg K)). Generally, the mineral oil (ISO-VG 100) offers a high heat capacity of 1789 J/(kg K) but a low thermal conductivity of 0.134 W/(m K) and hence friction heat should be removed from the contact via oil flow. It has to be expected that thermal conductivity (SSiC, EKasic F: = 125 W/(m K); Al2 O3 –ZrO2 , SN80: 28 W/(m K)) of the ceramics will influence friction behaviour at longer sliding times. Recently published experimental or theoretical work [18–20] was directed to microtexturing for reduction of friction and for increase in load bearing capacity. Hence, tribological behavior of texture pattern such as isolated circular microcavities was analysed. Hydrodynamic behavior of two parallel flat surfaces in sliding contact is explained [19] by suction of oil into the contact because the pressure inside the microcavities is less than the surrounding atmospheric pressure. Cavitation at the inlet side of the microcav-
ity plays an important role for the load support and inlet suction does not occur when cavitation pressure increases to atmospheric pressure [19]. This mechanism can explain the thicker oil film and lower friction of pairs consisting of parallel surfaces with microcavities. In agreement with this mechanism, Fig. 3 of the present study shows that the circular microdimples on the pellet surface resulted in greater oil film thickness and lower friction values with increasing sliding velocities than the only polished, untextured pellets. Shallow microcavities can lead to a substantial increase in film thickness, while deep microcavities can result in a local decrease or even collapse of the lubricant film [18–20]. This was ascribed to the more restricted lubricant flow from deep than from shallow microcavities [19] or to the different pressure distribution inside the microcavity, because the pressure inside a deep cavity is lower [18]. In contrast to numerous research work on isolated microcavities, the present study was concentrated on the effect of interconnected microchannels where the lubricant flow is less constricted. According to Fig. 3, circular microdimples increased but crossed microchannels decreased oil film thickness compared with untextured pellets. Fig. 4 shows that the orientation of parallel microchannels influenced film thickness and friction coefficient owing to eased or constrained oil drain out of the contact area. Microtexturing results in reduction of load bearing area A by the factor (1 − atext. ), since inside the microchannels exists only a low pressure, as mentioned above for circular cavities and in addition due to the low constraint of oil drain along the channels, and hence load bearing capacity is low at the area fraction of channels. The oil drain can be impeded by side wall friction inside the channel, reduced cross-section of the channels and greater oil viscosity. Increase of normal load from 15 N to 60 N (Fig. 9c and d) resulted in an about fourfold increase of friction force on the untextured but only in a twofold increase on the textured pellet at higher sliding velocities. From analysis of the experimental data at 15 N and under the assumption of hydrodynamic lubrication at v ≥about 5 m/s and Newtonian behavior, it was concluded that microtexturing by
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medium sized channels led to an increase in the effective values of film thickness and viscosity compared with the untextured pair. This apparent increase in viscosity and the enhanced friction force was caused by non-laminar oil flow and resulting hydrodynamic losses. Change of the normal load to 60 N led to reduction of oil film thickness at the textured pellet and to transition to common hydrodynamic behaviour with friction values comparable to the untextured pellet at high sliding velocities (Fig. 9d). At low sliding velocities, texturing by microchannels increased friction coefficient because reduction of oil film thickness led to transition from liquid to mixed or boundary friction. As result, friction coefficient was almost independently of sliding velocity at the load of 60 N (Fig. 9d). Oil film thickness became smaller with increasing area coverage fractions of microchannels and hence this transition was shifted to higher sliding velocities (Fig. 6). Results from the application-oriented friction test rig showed that not only friction values were raised with greater area fractions but also friction course versus sliding velocity was substantially altered (Fig. 10b). At a fraction of 75% of microchannels, friction values were almost independently of sliding velocity and the decline of friction coefficient with time of sliding at 10 m/s was reduced (Fig. 10a). Depth and width of microchannels were other parameters of microtexturing which influenced strongly film thickness and friction coefficient (Fig. 7). Increasing depths and widths of the channels reduced oil film thickness. Hence, friction values were raised and the transition to mixed or boundary lubrication occurred already at higher sliding velocities. This transition in lubrication conditions at larger channel widths (B = 150 or 200 m) was also confirmed by results from the friction test rig using the deceleration test (Fig. 11b). At higher sliding velocities however, smaller channel widths of 50 or 100 m resulted in higher friction values, which were almost independently of velocity (Fig. 11). Smaller channel widths, e.g. 60 m in the model test (Fig. 7a and b) and 50 or 100 m in the deceleration test (Fig. 11b) led to greater thicknesses of the oil film and lower friction values at low sliding velocities ≤0.5 m/s. Low depths and small widths of the microchannels promoted hydrodynamic lubrication and hence similar friction and lubrication conditions as the pairs with untextured pellets (Fig. 7). Fig. 12 shows a descriptive model developed from the experimental results and the support by in situ observation of the contact area in the model test using a high resolution microscope. The effective oil film thickness h* consists to a first approximation (Fig. 12c) of different parts: h0 = separation distance between pellet and disc without lubricant and normal load; which is a function (f) of (alignment deviation, surface roughness of contact area etc.), hstat = change of separation distance (film thickness) at lubricated static contact; f (oil charge and viscosity, normal load, capillary effects, wetting behavior etc.), htext = change of film thickness due to microtexturing at lubricated sliding contact; f (texture pattern, depth T, width B, area coverage fraction atext. , etc.) and hhydro = change of film thickness under hydrodynamic lubrication; f (oil viscosity, sliding velocity, characteristic length in sliding direction, normal load etc.). According to hydrodynamic theory, oil film thickness of pairs with untextured, only polished pellets increases continuously with sliding velocity and decreases with increasing normal load (Figs. 5b and 12b). Development of a load bearing oil film can be supported by the inlet suction of oil [19] between the parallel contact surfaces of pellet and disc. In the hydrodynamic region, microtexturing showed relatively little effect on friction behavior (Fig. 9a, b and d), but lowering the normal load from 60 to 15 N results in strong decrease of friction force by a factor of about four at untextured but only by the factor of two at textured pellets (Fig. 9c). With decreasing sliding velocity or increasing normal load the dominance of the hydrodynamic part hhydro on the effective film thickness h* decreases and the effect
of microtexturing htext gains influence on the slope of both film thickness and friction curve versus sliding velocity (Fig. 12a and b). Decreasing film thickness with decreasing sliding velocity, increasing normal load and/or increasing area fraction of microtexture lead to transition from completely oil filled to only partially oil filled microchannels and friction due to oil flow inside the channels when film thickness h* becomes smaller than the channel depth (Fig. 12d and e). As mentioned above completely oil filled microchannels offer only a limited load carrying capacity that can be enhanced by increased constraint of oil drain through the channels. Load carrying capacity of only partially oil filled channels goes to zero, especially at easy oil drain through broad and deep channels. This leads to a steep decrease in film thickness compared with the untextured pellets (Fig. 12b). Further decrease of film thickness with decreasing sliding velocity or increasing normal load as well as area coverage fraction, depth and width of microchannels can result in transition to mixed or even boundary lubrication, i.e. breakthrough of oil film and asperities interactions on the mated solids, when the value of h* approaches the sum of roughness values Rq of pellet and disc surface. This means the specific film thickness becomes smaller than 1 and value of friction coefficient increases (Fig. 12a and b). Hence, curves of oil film thickness and friction coefficient of pairs with textured pellets can run with decreasing sliding velocity over three different regions characterised by different slopes from completely oil filled channels (h > T) to only partially filled channels (h > T) at small oil film thickness and finally to boundary lubrication (h ≈ Rq ). Friction course versus sliding velocity results from superposition of solid friction, friction due to oil flow inside the channels and around the pellet as well as hydrodynamically induced full film friction. According to the model, depth of the texture pattern is a very important factor with respect to friction and load carrying capacity. 5. Conclusions Experiments showed that microtexturing of functional areas can be very useful to adapt tribological behavior to requirements of a friction system. Results from a model test and those from a friction test rig oriented to applications such as fast running oil lubricated friction systems, e.g. multiple disc clutches, led to a deeper understanding of the effect of texturing by microchannels. Comparing sliding pairs containing pellets textured by crossed microchannels (CH) with untextured pellets, it can be concluded that - oil film thickness is reduced by texture parameter such as increasing area coverage fraction, depth and width of microchannels (CH) as well as decreasing sliding velocity and increasing normal load. - crossed microchannels increase friction coefficient at medium to low sliding velocities due to oil flow inside the channels and transition to mixed lubrication. - curves of oil film and friction coefficient versus sliding velocity run at textured pellets (CH) over three different regions characterised by different slopes with decreasing film thickness, namely from full film lubrication with completely oil filled channels to partially oil filled channels with oil flow friction and to mixed or boundary lubrication when film thickness approaches roughness of the mated surfaces, according to the descriptive model presented. - crossed microchannels result in greater friction coefficient but lower oil film thickness than circular microdimples (CD). Acknowledgements The authors would like to thank the Deutsche Forschungsgemeinschaft (DFG) for financial support within the frame of the
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