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The influence of lubrication on ceramic and steel sliding contacts
Materials Science and Engineering, A 109 (1989) 407-413
407
The Influence of Lubrication on Ceramic and Steel Sliding Contacts* P. ANDERSSON
Technical Re.warch Centre of Finland, Laboratory of Engineering Production Technology, Metallimiehenkuja 6, SF-02150 Espoo (Finland) O. YLOST~ LO
Helsinki University of Technology, Laboratory of Machine Design, Puumiehenkuja 5 A, SF-02150 Espoo (Finland) (Received June 2, 1988)
Abstract Several reports show that lubrication improves the sliding properties of ceramics. This paper compares' dry and lubricated wear test results. Test materials were alumina, silicon carbide and steel, all sliding on steel. Dry tests were performed with a pin-on-disk machine and the lubricated tests with a journal bearing test machine under conditions of boundary lubrication. In dry wear tests, material transfer from one of the surfaces to the other was found to occur while in lubricated tests this was either reduced or inhibited. Wear rates during lubricated tests were about three orders of magnitude smaller than during unlubricated tests. The coefficient/~ of friction at the end of a test was found to fie in the range /~ = 0.37-0.08 for different sliding pairs in the dry tests while in lubricated tests it remained at/z = 0.02-0.03, showing the benefits of lubrication when appficable.
viscosity lubricant combined with a certain geometry. Boundary lubrication is named after the polarized boundary films which are formed particularly on steel surfaces as the lubricant interacts with the metal. Thus the expression of boundary lubrication is derived from a contact taking place in the boundary region instead of at the substrate surface. It has, however, been suggested in several contexts that the formation of boundary films is insufficient on ceramic surfaces, leading to less reliable boundary lubrication than can be counted on for steel surfaces. A wear mechanism typical of ceramics and present in unlubricated sliding contacts is that related to crack formation caused by traction forces. As lubrication reduces the traction forces, the extent of wear resulting from these will also be decreased [2]. The overall influence of a lubricant on a sliding pair including a ceramic surface will finally depend on several interacting factors.
1. Introduction
2. Experimental procedure
For unlubricated surfaces sliding on each other, the physical and chemical interactions may be quite extended and material transfer from one surface to another are significant signs of these reactions [1 ]. By introducing a lubricant between the sliding surfaces their interaction can, in many respects, be restricted. The surfaces remain less affected by each other and by the surrounding air as a result of the insulating effects of the lubricant. The boundary lubrication conditions occur at low sliding speeds, high loads and with a low-
2.1. Apparatus In order to describe the influence of a lubricant on the friction and wear of ceramic-steel pairs, results from unlubricated and lubricated tribological examinations were compared. The unlubricated tests were performed at the Technical Research Centre of Finland using a pin-on-disk machine with a fixed ball against a horizontally rotating disk geometry. The lubricated tests were performed at the Helsinki University of Technology using a journal bearing test machine with a cylindrical shaft rotating in a cylindrical bearing with a continuous oil supply. During the tests the friction force was recorded using load cells, amplifiers and pen plotters.
*Paper presented at the Symposium on Ceramic Materials Research at the E-MRS Spring Meeting, Strasbourg, May 31-June 2, 1988. 0921-5093/89/$3.50
© Elsevier Sequoia/Printed in The Netherlands
408
2.2. Specimens In dry tests, fixed balls with diameters of 10-12.7 mm were used as pins. The nominal size of the journal bearings used in the lubricated tests was 30 mm diam. × 20 mm. Every test was performed using new surfaces with a surface roughness Ra=0.04-0.12 /zm. The materials used in both dry and lubricated tests were sintered alumina, sintered silicon carbide and steel, all sliding on steel. The alumina and silicon carbide specimens were supplied by commercial engineering ceramics producers. The purities of the alumina and silicon carbide were about 99%, the densities 3.8 g c m -3 and 3.2 g cm -3, and the hardnesses about 1600 H V and 2200 H V respectively. The grade of steel was through hardened 100Cr6 (670 HV) in the dry tests and carburized case hardened 21NiCrMo2 (710 HV) in the lubricated tests. 2.3. Experimental details Dry pin-on-disk tests were performed using a normal force of 10 N, a sliding velocity of 0.2 m s-1 and a sliding distance of 250 m in laboratory air at 20°C and 45%-55% relative humidity. In lubricated journal bearing tests a normal force of 6 kN, a sliding velocity of 0.1 m s- 1 and a sliding distance of 36 km were used. Conditions of boundary lubrication were achieved by using a pure mineral oil with a viscosity of 3.8 × 10 - 3 N s m -2 at the test temperature of 60°C. Because of the low sliding velocity no hydrodynamic effect is estimated to be present. The number of dry and lubricated tests for each material combination was at least three and two respectively. Prior to testing all specimens were cleaned by ultrasonic agitation in 1,1,2-trichloro-l,2,2trifluoroethane. After drying, specimens for pinon-disk tests were further rinsed in ethanol and dried in hot air. Subsequent to the tests, wear rates for the specimens were determined. Diameters of wear tracks on balls used in unlubricated tests were measured in an optical microscope. Wear volumes were calculated using the wear track and ball diameters. Wear volumes of disks from unlubricated tests were calculated from Talysurf profilograms taken across the wear track. Wear volumes of shafts and bearings from lubricated tests were calculated from mass differences determined by weighing cleaned specimens before and after testing. After the measurements, certain
important surfaces were inspected by optical microscopy or scanning electron microscopy (SEM) and analysed using X-ray diffraction (XRD), energy-dispersive spectrometry (EDS) or Auger electron spectroscopy (AES) techniques. 3. Results and discussion
Coefficients of wear and friction are presented in Fig. 1. Wear rates were calculated by dividing the wear volume by the normal force and the sliding distance. In the graphs, wear rates of pins in dry tests or shafts in lubricated tests are illustrated with columns upwards from the horizontal line while wear rates for their counterfaces are shown downwards. Negative wear (layer formation) is illustrated with a shaded column on the opposite side of the horizontal line. Graphs presenting the coefficients of friction at the end of a test are situated below the wear-rate graphs.
3.1. Steel sliding on steel To start from a more traditional point of view, in the sense of materials selection, we begin with steel sliding on steel. Under dry conditions, it is obvious from the patterns seen on the photomicrographs in Fig. 2(a) that the contact mechanism is dominated by adhesive wear of the pin (ball). A layer of metal and oxidized wear debris is formed on the disk during material transfer from the pin. The wear track formed on the pin is characterized by a high density of small scratches formed during the removal of material and a debris shoe in front of the surface. When a lubricant is introduced in the steel-steel contact as sliding begins, both material and energy losses are reduced. No material trans-
V
(m / N m )
a
b
c
d
•
f
g
h
Fig. 1. (Top) Wear rates V* of pins in dry tests or shafts in lubricated tests are illustrated with columns upwards while positive wear rates for their counterfaces are shown downwards or negative wear rates (layer formation) with a shaded column upwards. (Bottom) The coefficients p of friction at the end of a test.
409
I.~ I v
v~*~ '"
17
OIL
Fig. 2. Wear surface photomicrographs and contact mechanism illustrations (a) for an unlubricated steel pin (top) against a steel disk (bottom) test and (b) for a lubricated steel shaft (top) against a steel bearing (bottom) test.
fer between the surfaces is observed, and wear mainly occurs on the steel bearing at the area which is most intensively loaded. As seen from Fig. 1, columns a and b, the wear rate of the lubricated steel bearing is about three orders of magnitude smaller than the wear rate of the dry steel pin. Correspondingly the coefficient of friction is reduced from 0.53 to about 0.03 when introducing boundary lubrication, also showing the wellknown benefits of using lubrication when possible. 3.2. A l u m i n a sliding on steel
By replacing either of the two specimens in the unlubricated pin-on-disk tests with an alumina specimen, two different steel-alumina test
geometries can be obtained. Using two inverse geometries it is possible to analyse the geometry sensitivity of a wear mechanism characteristic of a combination of materials. The dominating observation concerning the alumina-steel dry sliding contacts is the ferrous transfer to the ceramic surfaces, as seen in Fig. 3(a) and (b). Layers consisting of distributed metal areas surrounded by oxidized wear debris are formed on the alumina surfaces. Previous investigations have shown that metals and ceramics may form strong bonds [3], a fact which is also used in different joining techniques. Layer formation is most efficient for a steel pin (ball) in dry sliding on an alumina disk. Contemporary to running-in of the surfaces, a successive
J
Fig. 3. Wear surface photomicrographs and contact mechanism illustrations (a) for an unlubricated alumina pin (top) against a steel disk (bottom) test, (b) for an unlubricated steel pin (top) against an alumina disk (bottom) test and (c) for a lubricated steel shaft (top) against an alumina bearing (bottom) test.
OIL-
Fig. 4. Wear surface photomicrographs and contact mechanism illustrations (a) for an unlubricated silicon carbide pin (top) against a steel disk (bottom) test, (b) for an unlubricated steel pin (top) against a silicon carbide disk (bottom) test and (c) for a lubricated steel shaft (top) against a silicon carbide bearing (bottom) test.
OI
412 evolution towards a steel-steel contact occurs. The wear rate of the steel pin, the negative wear rate of the alumina disk, wear surface observations and the coefficient of friction are very similar to those found for steel sliding on steel. The inverse geometry with an alumina pin dry sliding against a steel disk has a higher coefficient of friction. This can be explained by the introduction of a ploughing component to the friction force as the harder pin slides on the softer disk. Photomicrographs show (Fig. 3(a)) that the layer formation on the alumina pin and the grooving of the steel disk are similar to those appearing on surfaces of the same materials used in the inverse geometry. Mechanically the layer formed on the alumina pin is very weak, and some wear of the substrate has taken place under the layer. The rate of ferrous layer formation on the pin is outside the values reported. Boundary lubrication reduces the rate of ferrous layer formation on alumina but it does not inhibit it, as seen on the alumina bearing wear surface photomicrograph in Fig. 3(c). Wear also occurs on the steel shaft, the rate exceeding that of a steel shaft against a steel bearing (Fig. 1, columns b and e). Test results show that boundary lubrication reduces the wear rate of the steel specimens and the negative wear rate of the alumina specimens by three orders of magnitude. The reduction of the coefficient of friction is also considerable (Fig. 1, columns c-e). 3.3. Silicon carbide sfiding on steel Uniubricated pin-on-disk tests with silicon carbide in contact with steel have also been performed using inverse geometries. For silicon carbide in contact with steel, the dominating contact mechanism is the formation of a transfer layer on steel (Fig. 4(a) and (b)). According to Auger analysis of steel pins slid against silicon carbide the transfer layer composition is complex, containing locally variable rates of silicon carbide, oxygen (or oxides) and free carbon. The steel disk transfer layer volume is larger than the silicon carbide pin wear volume, indicating porosity or oxidation products in the layer. The silicon carbide counterfaces are relatively smooth after the dry tests, containing only small distributed scratches. The transfer layer on a steel disk against a silicon carbide pin is thicker than for the inverse geometry. The coefficient of friction (Fig. 1, column f) is also lower for the first geometry, and similar to that found for silicon carbide against
itself. The friction force similarity can be explained as the result of a silicon carbide pin sliding on a silicon-carbide-originated layer on a steel substrate. The transfer layer formed on a steel pin in dry sliding on a silicon carbide disk is not sufficient to reduce the level of the coefficient of friction. However, the wear rate of the steel pin was very small, indicating a protecting effect of the silicon-carbide-originated transfer layer. Surfaces from boundary lubricated tests with steel shafts and silicon carbide bearings were analysed but no transfer layers were detected (Fig. 4(c)). Analysis methods which have been used on these surfaces are microscopies, EDS and XRD. It seems clear that the steel surface is protected by the lubricant from receiving a transfer layer. In dry tests the transfer layers on steel were advantageous for reducing the wear rates of the steel specimens. In boundary lubricated tests the wear-reducing effects of lubrication are much stronger and the reduction of the coefficient of friction is also evident.
4. Conclusions A comparison of results from dry pin-on-disk tests and boundary lubricated bearing tests with alumina, silicon carbide and steel is presented. A primary comment based upon the test results is that ceramics perform well as sliding surfaces. Lubrication improves the performance of a ceramic-steel sliding pair and should be used when applicable. Wear decreases by three orders of magnitude and friction by one order of magnitude when using pure mineral oil lubrication instead of unlubricated surfaces. Boundary lubrication reduces the material transfer in the alumina-steel contact and eliminates it for silicon carbide against steel. It is to be assumed that a change of sliding parameters leading to a more efficient lubrication mechanism would result in a further increase of the tribological performance of the sliding pair.
Acknowledgments The authors are grateful for financial support from the Technology Development Centre of Finland and the Federation of Finnish Metal and Engineering Industries. The authors also want to thank Mr. Risto O. Toivanen of the Helsinki University of Technology, Laboratory of Electron Microscopy for skilful Auger analysis.
413
References I C. Yust and F. Carignan, Observations on the sliding wear of ceramics, ASLE Trans., 28 (2) (1985) 245-252.
2 D. Buckley and K. Miyoshi, Friction and wear of ceramics, Wear, 100(1984) 333-353. 3 K. Miyoshi and D. Buckley, Friction and wear behaviour of single crystal silicon carbide in sliding contact with various metals, ASLE Trans.. 22 (3) (1979) 245-256.
407
The Influence of Lubrication on Ceramic and Steel Sliding Contacts* P. ANDERSSON
Technical Re.warch Centre of Finland, Laboratory of Engineering Production Technology, Metallimiehenkuja 6, SF-02150 Espoo (Finland) O. YLOST~ LO
Helsinki University of Technology, Laboratory of Machine Design, Puumiehenkuja 5 A, SF-02150 Espoo (Finland) (Received June 2, 1988)
Abstract Several reports show that lubrication improves the sliding properties of ceramics. This paper compares' dry and lubricated wear test results. Test materials were alumina, silicon carbide and steel, all sliding on steel. Dry tests were performed with a pin-on-disk machine and the lubricated tests with a journal bearing test machine under conditions of boundary lubrication. In dry wear tests, material transfer from one of the surfaces to the other was found to occur while in lubricated tests this was either reduced or inhibited. Wear rates during lubricated tests were about three orders of magnitude smaller than during unlubricated tests. The coefficient/~ of friction at the end of a test was found to fie in the range /~ = 0.37-0.08 for different sliding pairs in the dry tests while in lubricated tests it remained at/z = 0.02-0.03, showing the benefits of lubrication when appficable.
viscosity lubricant combined with a certain geometry. Boundary lubrication is named after the polarized boundary films which are formed particularly on steel surfaces as the lubricant interacts with the metal. Thus the expression of boundary lubrication is derived from a contact taking place in the boundary region instead of at the substrate surface. It has, however, been suggested in several contexts that the formation of boundary films is insufficient on ceramic surfaces, leading to less reliable boundary lubrication than can be counted on for steel surfaces. A wear mechanism typical of ceramics and present in unlubricated sliding contacts is that related to crack formation caused by traction forces. As lubrication reduces the traction forces, the extent of wear resulting from these will also be decreased [2]. The overall influence of a lubricant on a sliding pair including a ceramic surface will finally depend on several interacting factors.
1. Introduction
2. Experimental procedure
For unlubricated surfaces sliding on each other, the physical and chemical interactions may be quite extended and material transfer from one surface to another are significant signs of these reactions [1 ]. By introducing a lubricant between the sliding surfaces their interaction can, in many respects, be restricted. The surfaces remain less affected by each other and by the surrounding air as a result of the insulating effects of the lubricant. The boundary lubrication conditions occur at low sliding speeds, high loads and with a low-
2.1. Apparatus In order to describe the influence of a lubricant on the friction and wear of ceramic-steel pairs, results from unlubricated and lubricated tribological examinations were compared. The unlubricated tests were performed at the Technical Research Centre of Finland using a pin-on-disk machine with a fixed ball against a horizontally rotating disk geometry. The lubricated tests were performed at the Helsinki University of Technology using a journal bearing test machine with a cylindrical shaft rotating in a cylindrical bearing with a continuous oil supply. During the tests the friction force was recorded using load cells, amplifiers and pen plotters.
*Paper presented at the Symposium on Ceramic Materials Research at the E-MRS Spring Meeting, Strasbourg, May 31-June 2, 1988. 0921-5093/89/$3.50
© Elsevier Sequoia/Printed in The Netherlands
408
2.2. Specimens In dry tests, fixed balls with diameters of 10-12.7 mm were used as pins. The nominal size of the journal bearings used in the lubricated tests was 30 mm diam. × 20 mm. Every test was performed using new surfaces with a surface roughness Ra=0.04-0.12 /zm. The materials used in both dry and lubricated tests were sintered alumina, sintered silicon carbide and steel, all sliding on steel. The alumina and silicon carbide specimens were supplied by commercial engineering ceramics producers. The purities of the alumina and silicon carbide were about 99%, the densities 3.8 g c m -3 and 3.2 g cm -3, and the hardnesses about 1600 H V and 2200 H V respectively. The grade of steel was through hardened 100Cr6 (670 HV) in the dry tests and carburized case hardened 21NiCrMo2 (710 HV) in the lubricated tests. 2.3. Experimental details Dry pin-on-disk tests were performed using a normal force of 10 N, a sliding velocity of 0.2 m s-1 and a sliding distance of 250 m in laboratory air at 20°C and 45%-55% relative humidity. In lubricated journal bearing tests a normal force of 6 kN, a sliding velocity of 0.1 m s- 1 and a sliding distance of 36 km were used. Conditions of boundary lubrication were achieved by using a pure mineral oil with a viscosity of 3.8 × 10 - 3 N s m -2 at the test temperature of 60°C. Because of the low sliding velocity no hydrodynamic effect is estimated to be present. The number of dry and lubricated tests for each material combination was at least three and two respectively. Prior to testing all specimens were cleaned by ultrasonic agitation in 1,1,2-trichloro-l,2,2trifluoroethane. After drying, specimens for pinon-disk tests were further rinsed in ethanol and dried in hot air. Subsequent to the tests, wear rates for the specimens were determined. Diameters of wear tracks on balls used in unlubricated tests were measured in an optical microscope. Wear volumes were calculated using the wear track and ball diameters. Wear volumes of disks from unlubricated tests were calculated from Talysurf profilograms taken across the wear track. Wear volumes of shafts and bearings from lubricated tests were calculated from mass differences determined by weighing cleaned specimens before and after testing. After the measurements, certain
important surfaces were inspected by optical microscopy or scanning electron microscopy (SEM) and analysed using X-ray diffraction (XRD), energy-dispersive spectrometry (EDS) or Auger electron spectroscopy (AES) techniques. 3. Results and discussion
Coefficients of wear and friction are presented in Fig. 1. Wear rates were calculated by dividing the wear volume by the normal force and the sliding distance. In the graphs, wear rates of pins in dry tests or shafts in lubricated tests are illustrated with columns upwards from the horizontal line while wear rates for their counterfaces are shown downwards. Negative wear (layer formation) is illustrated with a shaded column on the opposite side of the horizontal line. Graphs presenting the coefficients of friction at the end of a test are situated below the wear-rate graphs.
3.1. Steel sliding on steel To start from a more traditional point of view, in the sense of materials selection, we begin with steel sliding on steel. Under dry conditions, it is obvious from the patterns seen on the photomicrographs in Fig. 2(a) that the contact mechanism is dominated by adhesive wear of the pin (ball). A layer of metal and oxidized wear debris is formed on the disk during material transfer from the pin. The wear track formed on the pin is characterized by a high density of small scratches formed during the removal of material and a debris shoe in front of the surface. When a lubricant is introduced in the steel-steel contact as sliding begins, both material and energy losses are reduced. No material trans-
V
(m / N m )
a
b
c
d
•
f
g
h
Fig. 1. (Top) Wear rates V* of pins in dry tests or shafts in lubricated tests are illustrated with columns upwards while positive wear rates for their counterfaces are shown downwards or negative wear rates (layer formation) with a shaded column upwards. (Bottom) The coefficients p of friction at the end of a test.
409
I.~ I v
v~*~ '"
17
OIL
Fig. 2. Wear surface photomicrographs and contact mechanism illustrations (a) for an unlubricated steel pin (top) against a steel disk (bottom) test and (b) for a lubricated steel shaft (top) against a steel bearing (bottom) test.
fer between the surfaces is observed, and wear mainly occurs on the steel bearing at the area which is most intensively loaded. As seen from Fig. 1, columns a and b, the wear rate of the lubricated steel bearing is about three orders of magnitude smaller than the wear rate of the dry steel pin. Correspondingly the coefficient of friction is reduced from 0.53 to about 0.03 when introducing boundary lubrication, also showing the wellknown benefits of using lubrication when possible. 3.2. A l u m i n a sliding on steel
By replacing either of the two specimens in the unlubricated pin-on-disk tests with an alumina specimen, two different steel-alumina test
geometries can be obtained. Using two inverse geometries it is possible to analyse the geometry sensitivity of a wear mechanism characteristic of a combination of materials. The dominating observation concerning the alumina-steel dry sliding contacts is the ferrous transfer to the ceramic surfaces, as seen in Fig. 3(a) and (b). Layers consisting of distributed metal areas surrounded by oxidized wear debris are formed on the alumina surfaces. Previous investigations have shown that metals and ceramics may form strong bonds [3], a fact which is also used in different joining techniques. Layer formation is most efficient for a steel pin (ball) in dry sliding on an alumina disk. Contemporary to running-in of the surfaces, a successive
J
Fig. 3. Wear surface photomicrographs and contact mechanism illustrations (a) for an unlubricated alumina pin (top) against a steel disk (bottom) test, (b) for an unlubricated steel pin (top) against an alumina disk (bottom) test and (c) for a lubricated steel shaft (top) against an alumina bearing (bottom) test.
OIL-
Fig. 4. Wear surface photomicrographs and contact mechanism illustrations (a) for an unlubricated silicon carbide pin (top) against a steel disk (bottom) test, (b) for an unlubricated steel pin (top) against a silicon carbide disk (bottom) test and (c) for a lubricated steel shaft (top) against a silicon carbide bearing (bottom) test.
OI
412 evolution towards a steel-steel contact occurs. The wear rate of the steel pin, the negative wear rate of the alumina disk, wear surface observations and the coefficient of friction are very similar to those found for steel sliding on steel. The inverse geometry with an alumina pin dry sliding against a steel disk has a higher coefficient of friction. This can be explained by the introduction of a ploughing component to the friction force as the harder pin slides on the softer disk. Photomicrographs show (Fig. 3(a)) that the layer formation on the alumina pin and the grooving of the steel disk are similar to those appearing on surfaces of the same materials used in the inverse geometry. Mechanically the layer formed on the alumina pin is very weak, and some wear of the substrate has taken place under the layer. The rate of ferrous layer formation on the pin is outside the values reported. Boundary lubrication reduces the rate of ferrous layer formation on alumina but it does not inhibit it, as seen on the alumina bearing wear surface photomicrograph in Fig. 3(c). Wear also occurs on the steel shaft, the rate exceeding that of a steel shaft against a steel bearing (Fig. 1, columns b and e). Test results show that boundary lubrication reduces the wear rate of the steel specimens and the negative wear rate of the alumina specimens by three orders of magnitude. The reduction of the coefficient of friction is also considerable (Fig. 1, columns c-e). 3.3. Silicon carbide sfiding on steel Uniubricated pin-on-disk tests with silicon carbide in contact with steel have also been performed using inverse geometries. For silicon carbide in contact with steel, the dominating contact mechanism is the formation of a transfer layer on steel (Fig. 4(a) and (b)). According to Auger analysis of steel pins slid against silicon carbide the transfer layer composition is complex, containing locally variable rates of silicon carbide, oxygen (or oxides) and free carbon. The steel disk transfer layer volume is larger than the silicon carbide pin wear volume, indicating porosity or oxidation products in the layer. The silicon carbide counterfaces are relatively smooth after the dry tests, containing only small distributed scratches. The transfer layer on a steel disk against a silicon carbide pin is thicker than for the inverse geometry. The coefficient of friction (Fig. 1, column f) is also lower for the first geometry, and similar to that found for silicon carbide against
itself. The friction force similarity can be explained as the result of a silicon carbide pin sliding on a silicon-carbide-originated layer on a steel substrate. The transfer layer formed on a steel pin in dry sliding on a silicon carbide disk is not sufficient to reduce the level of the coefficient of friction. However, the wear rate of the steel pin was very small, indicating a protecting effect of the silicon-carbide-originated transfer layer. Surfaces from boundary lubricated tests with steel shafts and silicon carbide bearings were analysed but no transfer layers were detected (Fig. 4(c)). Analysis methods which have been used on these surfaces are microscopies, EDS and XRD. It seems clear that the steel surface is protected by the lubricant from receiving a transfer layer. In dry tests the transfer layers on steel were advantageous for reducing the wear rates of the steel specimens. In boundary lubricated tests the wear-reducing effects of lubrication are much stronger and the reduction of the coefficient of friction is also evident.
4. Conclusions A comparison of results from dry pin-on-disk tests and boundary lubricated bearing tests with alumina, silicon carbide and steel is presented. A primary comment based upon the test results is that ceramics perform well as sliding surfaces. Lubrication improves the performance of a ceramic-steel sliding pair and should be used when applicable. Wear decreases by three orders of magnitude and friction by one order of magnitude when using pure mineral oil lubrication instead of unlubricated surfaces. Boundary lubrication reduces the material transfer in the alumina-steel contact and eliminates it for silicon carbide against steel. It is to be assumed that a change of sliding parameters leading to a more efficient lubrication mechanism would result in a further increase of the tribological performance of the sliding pair.
Acknowledgments The authors are grateful for financial support from the Technology Development Centre of Finland and the Federation of Finnish Metal and Engineering Industries. The authors also want to thank Mr. Risto O. Toivanen of the Helsinki University of Technology, Laboratory of Electron Microscopy for skilful Auger analysis.
413
References I C. Yust and F. Carignan, Observations on the sliding wear of ceramics, ASLE Trans., 28 (2) (1985) 245-252.
2 D. Buckley and K. Miyoshi, Friction and wear of ceramics, Wear, 100(1984) 333-353. 3 K. Miyoshi and D. Buckley, Friction and wear behaviour of single crystal silicon carbide in sliding contact with various metals, ASLE Trans.. 22 (3) (1979) 245-256.