Wear performance of oil lubricated silicon nitride sliding against various bearing steels

Wear 255 (2003) 657–668

Wear performance of oil lubricated silicon nitride sliding against various bearing steels L. Wang a,∗ , R.J.K. Wood a , T.J. Harvey a , S. Morris a , H.E.G. Powrie b , I. Care c a

Surface Engineering and Tribology Group, School of Engineering Sciences, University of Southampton, Southampton SO17 1BJ, UK b Smiths Aerospace Electronic Systems, School Lane, Chandlers Ford, Southampton, Hampshire SO53 4YG, UK c Rolls-Royce plc, ML-77, PO Box 31, Derby DE24 8BJ, UK

Abstract The selection of bearing steel surfaces for use with silicon nitride rolling elements within hybrid bearings is critical to the performance and life of such components, which have potential applications in advanced high speed aircraft. The wear and friction performance of these combinations is a major factor currently being considered for the next generation hybrid bearings. This paper reports on hybrid bearing contacts that have lubricated Si3 N4 elements, which have been loaded against various bearing steels under pure sliding contact conditions on a fully instrumented pin-on-disc wear test rig. The wear and friction performance of Si3 N4 has been compared to a baseline case of bearing steel M50 ball sliding against a M50 disc. Both hybrid and steel on steel contacts were lubricated by an aircraft engine oil Mobil Jet II. Wear mechanisms were determined by post-test analysis of the pin wear scars, disc wear surface and wear debris using optical microscopy, surface profilometry and FEG-SEM (scanning electron microscopy). The wear rates of Si3 N4 sliding against different bearing steels are ranked by performance and related to their wear mechanisms, hardness and microstructure. Typical sliding contact wear mechanisms were found for the steel on steel combination while Si3 N4 sliding against steel showed that transgranular and sub-micron-cracking mechanisms predominate. Evidence of material transfer (steel onto the silicon nitride) was found. Friction values for the various combinations are also reported and found to be substantially lower (µ = 0.04) than bearing steel on bearing steel combinations (µ = 0.17). The disc and pin wear was monitored on-line by an electrostatic wear sensor, LVDT and laser displacement probe, a friction strain gauge, and an infrared thermometer. Correlations between wear rate and charge generation/level, friction, contact temperature, and disc hardness are presented. © 2003 Elsevier Science B.V. All rights reserved. Keywords: Silicon nitride; Wear; Electrostatic sensing; Bearing steel

1. Introduction Silicon nitride ceramics were developed for use in rolling bearings and hybrid ceramic ball bearings, with Si3 N4 balls and steel rings, and have been used in machine tool spindles under high speed operation for approximately 10 years [1]. Silicon nitride has also been proposed for very high temperature rolling element bearings on aircraft engines [2]. Yui et al. [3] have tested hybrid ball bearings under high temperature (300 ◦ C) and high speed (3.5 million DN—bearing inner race rib diameter (D) × number of revolutions (N)) conditions. The results of the tests confirmed that the bearings function normally with tripentaerythritol ester oil at high temperatures that exceed the limiting temperature of conventional ester-based gas turbine engine oils, and that the sintered powder carburising steel is well suited to bearing applications in which present levels of temperature and ∗ Corresponding author. E-mail address: [email protected] (L. Wang).

speed are exceeded. The properties of the selected steels for bearing races are also very critical to the performance of hybrid bearings. To improve the rolling contact fatigue (rcf) resistance, wear and scuffing resistance, and tolerance to poor lubrication conditions of the well-established bearing steels M50 and M50NiL, a plasma nitriding process was developed to enhance the surface hardness and microstructure. There is some evidence that the resistance is proportional to hardness according to Archard wear law [4]. Plasma nitriding results in the formation of a diffusion zone below the surface and, depending on the nitrogen concentration during nitriding, in the formation of a so called compound layer (white layer) on the surface. The compound layer, which is predominately iron nitride and only several microns thick, is very hard and brittle, and thus is prone to break down under the influence of the high cyclic stresses experienced by a bearing during service. Therefore, this white layer must be removed by grinding from the race of a bearing. The surface hardness of M50 DH and M50NiL DH achieved by plasma nitriding is about 1000 Hv, the normal

0043-1648/03/$ – see front matter © 2003 Elsevier Science B.V. All rights reserved. doi:10.1016/S0043-1648(03)00045-0

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surface hardness of M50 and M50NiL is only about 750 Hv [5]. The wear mechanisms of ceramics have been widely discussed and several mechanisms have been suggested, which include micro-cracking either transgranular or intergranular [6–8], fracture [9], tribochemical reaction [10] and sub-micron chipping and polishing wear [11]. The wear states for ceramics are recognised as mild wear and severe wear [12]. For mild wear, tribochemical reactions take place and form fine wear debris. The debris is normally smaller than the grain size by one-tenth or less. The wear rate remains below 10−6 mm3 /N m. Under severe wear, the wear debris is formed by mechanical cracking of grains or delamination of tribo-films in the scale of the grain size. The wear rate is larger than 10−6 mm3 /N m. By simulating loading conditions in the Si3 N4 hybrid ball bearing system, a wear process was proposed [7]. Wear in the Si3 N4 bearing ball was initiated by contact fatigue stresses. Generated debris is subsequently entrained and suspended in the lubricant and creates a second wear mechanism of three-body abrasive wear, i.e. scratching or cutting in both Si3 N4 and M50 steel race surfaces. Large pieces of debris in the lubricant act as sharp indenters and create larger local inelastic deformation regions in the ball or race surface due to their indentation. The lateral cracks from this process could be initiated sub-surface due to local inelastic deformation. Therefore, inducing spallation, a third wear mechanism. The level of the coefficient of friction is an important factor to be considered in wear tests. A large amount of research has been devoted to measuring the coefficient of friction for the hybrid contacts. Compared to steel on steel, hybrid contact generally has lower level of friction and it has been found that the coefficient of friction for hybrid contact is normally under 0.1 for oil lubricated conditions, and between 0.1 and 1.0 for dry conditions [12–14]. Electrostatic sensing technology has been developed to detect charged debris present in the gas path of turbine engines. The electrostatic charge level in the exhaust gas changes when the amount of debris changes, for example, due to debris produced from a gas path component fault. This change can be detected by installing electrostatic sensors in the gas path and thus is a non-intrusive detecting system for monitoring gas path component deterioration and hence condition [15]. This paper presents novel work looking at applying this technology to monitor the distress levels within fully lubricated hybrid contacts. The electrostatic sensing technology is based on the assumption that a change of the static charge level in the system reflects changes in the conditions of the dynamic system. The theory and details about this technology have been given in Refs. [16–20]. Fig. 1 shows a schematic illustration of the electrostatic charge sensing system to explain its operation [21]. The electrostatic monitoring system comprises a passive sensor connected to signal conditioning (a charge amplifier) from which voltage signals may be recorded and processed. If electric (E) field lines due to charge Q pass-

Fig. 1. Schematic illustration of the electrostatic charge sensing system.

ing the sensor terminate on the sensor face, the electrons in the sensor redistribute to balance the additional charge in the vicinity of the sensor and hence the presence of the charge is measurable. The charge Q is shown as positive in Fig. 1. The signal conditioning converts the detected charge into a proportional voltage signal, which is collected and processed. When a charged surface or charged particle (e.g. debris) moves across a sensor face, an electrostatic charge is induced on the sensor face. The magnitude of induced charge is proportional to the flux of field lines terminating on sensor face. The advantages of using this technique compared with conventional condition monitoring technologies are: (a) It uses a direct measurement of debris produced by contact degradation rather than secondary effects such as excessive vibration and/or temperature exceedances. (b) It is sensitive to most material types including ferrous, metallic and non-metallic such as ceramics [22]. Tasbaz et al. [16] have investigated the use of electrostatic sensors for early detection of scuffing in oil lubricated metal/metal contacts. Electrostatic events were detected prior to the onset of severe scuffing. That result further demonstrated the potential of using electrostatic detection technology as an on-line machinery condition monitoring system. In this current study, the main objective was to seek correlation between the specific wear rates (SWRs) of the contact materials and the electrostatic charge levels. The other objective was to compare the wear performance of a range of advanced bearing steels against silicon nitride under lubricated sliding tests on the laboratory pin-on-disc (PoD) tribometer. The wear mechanisms of silicon nitride and the bearing steels are also discussed. The tests were conducted

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on a fully instrumented test rig capable of on-line disc wear track depth sensing and profilometry, as well as disc temperature monitoring.

2. Experimental setup and test conditions Lubricated Si3 N4 balls were loaded against various bearing steels under pure sliding conditions on a PoD wear test rig. The performance of the various Si3 N4 balls on steel contacts was compared to the baseline case of lubricated M50 bearing steel sliding against itself.

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Table 1 A list of the bearing steels tested on the PoD tribometer Bearing steel

Type/hardening details

Hardness (Hv0.1 kg)

M50 M50 NiL M50 DH

VIM-VAR VIM-VAR, case hardened VIM-VAR, duplex-hardened (through hardened and plasmanitrided) VIM-VAR, duplex-hardened (case hardened and plasmanitrided) Corrosion resistant carburised steel Nitrogen stainless steel Powder metallurgy corrosion resistant steel

784 793 1040

M50 NiL DH

XT1 XT2 XT3

911

748 777 766

2.1. PoD tribometer The test pins (both M50 and Si3 N4 ) were 6.35 mm diameter balls and the bearing steel discs were all 100 mm in diameter and 14 mm thick. The balls were tested in the ‘as-received’ state while the discs were ground and lapped (diamond), to Ra = 0.1–0.2 ␮m, before testing. To enable an extensive investigation of the wear as well as the charge generation, additional instrumentation was installed on the PoD including: (a) a tachometer to monitor the disc rotation speed; (b) an electrostatic sensor to monitor the charge level on the wear track; (c) a strain gauge to monitor the contact friction; (d) an LVDT (linear-variable-differential-transformer) to monitor the vertical displacement of the pin towards the disc from which the linear pin and disk wear is calculated; (e) a laser probe to monitor the depth and profile of the wear track on the disc (accuracy ± 0.2 ␮m, laser spot size of 65 ␮m diameter); (f) an infrared thermometer to monitor the disc surface temperature (about 10 mm from the contact point). The positions of these sensors on the rig are illustrated in Fig. 2.

Fig. 2. Schematic diagram of the laboratory PoD tribometer.

2.2. Test materials The materials for the tests are as follows: (a) Ball materials: M50 and Si3 N4 (HIPSN). The Si3 N4 test balls are commercially available bearing-grade silicon nitride balls. The toughness is 6–8 MPa m1/2 and the grain size is normally 5–6 ␮m but less than 1 ␮m in the ball centre. (b) Disc materials: M50 plus six other bearing materials are described in Table 1. These materials represent a range of surface technologies and experimental surface treatments for use as bearing surfaces. Micro-Vickers hardness (0.1 kg) was tested for all the disc surfaces and is also listed in Table 1. (c) The lubricant tested was Mobil Jet II—a first generation advanced aircraft engine oil with viscosity of 25.3 cSt at 40 ◦ C and 5.0 cSt at 100 ◦ C. 2.3. Test conditions and test programme All the tests were carried out at ambient temperature (18–28 ◦ C) and at a relative humidity of 40–70%. The tests started with rotating the disc without pin contact or lubrication to obtain background levels on the instrumentation. After a few minutes running, the lubricant was sprayed onto the disc before the pin was brought into contact with the disc. The rig was run under this condition for a few minutes before an initial load of 10 N was applied. The load was then gradually increased to the maximum load then held constant for the remainder of the test. Further details about the test conditions are listed in Table 2. The disc rotation speed was varied to maintain a sliding velocity at 7 m/s for all tests. This speed was chosen to replicate the sliding conditions experienced within the outer parts of the ball/race contact which for aircraft bearings vary between 7 and 20 m/s. Although the contact within the hybrid bearing normally experiences sliding and rolling, under low bearing loads pure sliding can occur. However, pure sliding conditions were chosen here

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Table 2 Details of the wear test conditions (100% sliding) Condition

Values

Sliding velocity (m/s) Maximum load (N) Duration under maximum load (m) Sliding distance (m) Lubricant Temperature (◦ C) Humidity (%) Maximum Hertzian contact stress (GPa) Initial lambda (λ) Surface roughness (␮m)

7 60 30 12600 Mobil Jet II Ambient 40–70 2.50 for M50/steel, 2.83 for Si3 N4 /steel 2.4–3.3 0.06 for M50 and Si3 N4 balls, 0.1–0.2 for all discs

120 90 37800

3.56 for Si3 N4 /steel 2.3–3.5

primarily to accelerate the tests and generate wear debris to investigate the sensitivity of the electrostatic sensor technology to these wear rates and mechanisms. A load of 60 N was applied to the M50 on steel as well as the silicon nitride on steel contacts. This particular load was chosen to achieve a contact pressure of 2.5 GPa for the steel on steel and over 2.5 GPa for the Si3 N4 on steel. A higher load of 120 N was applied to the Si3 N4 /steel combinations to differentiate them due to the low wear rates generated at 60 N.

be calculated according to Eq. (3). Thus, the SWRs of the contact pair were obtained, see Eq. (4), and the performance of the couples were ranked accordingly:

2.4. FEG-SEM test and EDX analysis

where D is the track diameter, A the average track cross-section area, R the radius of the pin, and h the height of the worn pin. VL is the volume loss of the disc or pin, F the load, SD the sliding distance and SWR the specific wear rate in mm3 /N m.

The new and worn ball surfaces were observed by a JSM 6500 FEG-SEM to investigate the wear mechanisms. Elemental mapping of the various sample surfaces were obtained by energy dispersive X-ray (EDX) analysis.

VLdisc = πDA

(1)

VLpin = 13 (πh2 (3R − h))

(2)

h = R − (R2 − 41 d 2 )1/2

(3)

SWR =

VL (mm3 ) F (N) × SD (m)

(4)

4. Results and discussion 3. Data analysis methods 4.1. Wear performance of the material combinations On-line monitoring signals were processed to examine the instantaneous change in charge, pin/disc linear wear, friction and temperature. Correlations between these signals were investigated. The data acquisition and signal processing layout is shown in Fig. 3. Amplifiers were used for the signals from the electrostatic sensor, the laser probe, the strain gauge and the LVDT probe. An A/D sampling rate of 8 kHz was used for all channels. A four-channel oscilloscope was used to visualise the on-line readings from the tachometer, electrostatic sensor, laser probe and temperature probe. A DAT recorder was used to record the raw data as a backup. The disc wear tracks were measured with a 2D Talysurf profilometer. For each wear track, four radial measurements were carried out around the wear track circumference and the average of the four measurements used to calculate the disc volume loss (see Eq. (1)). The pin wear scar diameters were measured using optical microscopy. The average of two diameters (the largest and the smallest diameter across the wear scar) was used to calculate the volume loss of the pin (see Eq. (2) [23]), while the height of the worn pin, h, can

The wear performance of the test material combinations was evaluated by calculating the SWR. The relative wear resistance (RWR) of the disc materials was calculated by comparing the SWR of the test material with the SWR values of the M50 pin sliding on an M50 disc under 60 N load using Eqs. (5) and (6): 
SWRtest RWRdisc = (5) SWRM50 at 60 N disc
 SWRtest (6) RWRpin = SWRM50 at 60 N pin Fig. 4 shows the RWRs of both the pin and disc during the sliding tests for the M50 pin loaded at 60 N. It shows that the powder metallurgy corrosion resistant steel XT3 has the best wear resistance, about 30 times better than the baseline material M50. For Si3 N4 balls, two figures were plotted to show the two load conditions. Fig. 5 shows the load of 60 N and Fig. 6

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Fig. 3. Block diagram of the data acquisition system for the PoD tests.

Fig. 4. Pin/disc RWRs comparison: M50 pin, 60 N load, 7 m/s sliding velocity, Mobil Jet II oil lubricated.

Fig. 5. Pin/disc RWRs comparison: Si3 N4 pin, 60 N load, 7 m/s sliding velocity, Mobil Jet II oil lubricated.

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Fig. 6. Pin/disc RWRs comparison: Si3 N4 pin, 120 N load, 7 m/s sliding velocity, Mobil Jet II oil.

shows the results under 120 N load. The M50NiL DH is the best wear resistant bearing steel under 60 N load, which agrees with the results of Dodd et al. [5] based on actual bearing tests. However, the wear resistance under 120 N, relative to M50 against M50 under 60 N, of M50NiL DH is 35 while for M50NiL it is 18,000, clearly making M50NiL the best material. 4.2. The coefficient of friction The averages of the coefficients of friction during the tests were calculated for the hybrid and steel on steel contacts, and the results are listed in Table 3. The coefficients of friction for the oil lubricated Si3 N4 on steels are between 0.04 and 0.09, while for the lubricated M50 on steels the values are between 0.10 and 0.17.

out and the average of the four measurements was used to represent the disc’s hardness, shown in Table 1. The SWRs are plotted against the disc hardness in Fig. 7. Fig. 7(a) shows the tests with M50 pin under 60 N load while Fig. 7(b) and (c) shows the hybrid contacts results under 60 and 120 N loads, respectively. There is no obvious relationship between the disc hardness and its SWR, which agrees with previous work by Zum Gahr [24]. Apart from XT1 and XT2, all other bearing steels show better performances with hybrid contacts. Their wear rates when they are loaded against silicon nitride are at least two orders of magnitudes less than when they are loaded against M50. The pin generally wears less than the disc by one to two orders of magnitudes. 4.4. Correlation between electrostatic charge levels and wear processes

4.3. Hardness tests Micro-hardness tests of 0.1 kg load were performed for each disc material. For each disc, four tests were carried Table 3 Coefficient of friction for hybrid and all steel lubricated sliding contacts Pin

Disc

Load (N)

Coefficient of friction

M50 M50 M50 Si3 N4 Si3 N4 Si3 N4 Si3 N4 Si3 N4 Si3 N4 Si3 N4 Si3 N4

M50 M50NiL XT3 M50 M50NiL M50DH M50NiL DH XT3 M50 M50NiL M50DH

60 60 60 60 60 60 60 60 120 120 120

0.16 0.17 0.10 0.09 0.04 0.07 0.06 0.06 0.08 0.05 0.08

The electrostatic on-line monitoring results from the wear test of the Si3 N4 /M50 contact under a load of 120 N with Mobil Jet II lubrication are shown in Fig. 8. From plots (a)–(e), the traces show the change of contact temperature, coefficient of friction, disc wear track depth obtained from the laser probe, electrostatic sensor response and the SWR of the pin (obtained from the LVDT) vs. time, respectively. It should be noted that the time axis starts when the constant load of 120 N (maximum load) was applied after 43 min. The previous loading period is not shown in this figure. It can be seen that, under the constant load of 120 N, the SWR decreased gradually until the 68th minute, see plot (e), this is possibly due to the running-in ‘polishing’ wear mechanism [11] is dominating, generating sub-micron debris which may act as a grinding paste. Then, there are some sudden increases followed by a steady increase in SWR for about 12 min before another period of instantaneous

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Fig. 7. SWRs of both pin and disc vs. the disc surface micro-Vickers’s hardness (Hv0.1): (a) 60 N load, M50 pin; (b) 60 N load, Si3 N4 pin; (c) 120 N load, Si3 N4 pin.

fluctuations. During this period, more and more fine debris are generated and agglomerated, larger particles (as evidenced by post-test debris inspection) formed on both contact surfaces as well as suspended in the lubricant acting as sharp indenters and creating local elasto-plastic deformation regions in the ball or disc surface [7]. Therefore, a spallation wear mechanism is dominant in this period. After

the 90th minute, the SWR started to level off. It was noted that the contact temperature at this time is above 70 ◦ C, the temperature at which the anti-wear additive in the oil is thought to become active [25]. A similar pattern occurred on the electrostatic charge plot (d). There is general agreement between the SWR trends and charge with some fairly high charge levels being generated

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in charge between 100 and 110 min that leads up to a sudden increase in SWR at 110 min. This could represent a growth in surface phase transformation or differences in additive film thicknesses and represents a precursor signal to the high wear event. The contact temperature at this point is above 70 ◦ C which is sufficient to activate additive film formation. Possible film failure within the disc wear tracks could have been responsible for the high wear indication at 110 min. Fig. 8(c) shows the change in depth of the disc wear track, obtained from the laser probe, during the test. The value in the plot becomes more negative for a deepening track. At the 68th minute, there was a large drop in the laser probe response indicating a significant increase in wear rate. This is thought to be the result of spalling of the silicon nitride ball and the contact becoming less conformal which in turn increases the cutting type wear of the disc surface. Subsequently, the wear track was filled with the fine rubbing debris, indicated by a decrease in the wear track depth. After a further 12 min of steady running a similar feature was observed. The contact temperature and the coefficient of friction also show very strong correlations with wear of the contact (see Fig. 8(a) and (b)) with both increasing with increasing wear. The coefficient of friction settled at about 0.07 after 110 min. At about 90 min, the surface temperature went over 70 ◦ C at which the anti-wear additives become active and the SWR stopped increasing [25]. 4.5. Wear mechanisms

Fig. 8. The test results for Si3 N4 /M50/120 N contact with Mobil Jet II lubrication: (a) contact temperature; (b) coefficient of friction; (c) laser probe reading; (d) electrostatic sensor response; (e) pin SWR vs. time.

between 80 and 90 min coinciding with sudden increases in SWR noted earlier. The increased charge levels reflect the increasing amount of charged debris being generated. However, the possibility of surface phase transformation or differences in additive film thickness within the disc wear tracks could cause surface contact potential differences [19] and thus, also could contribute to the charge levels. Tribocharging is likely to be constant unless the film thickness varies dramatically and therefore is unlikely to contribute to change of charge level. A correlation between the charge level and the SWR can been seen by comparing the two lines in Fig. 8(d) and (e), respectively. Before 68th minute, the wear rate went down slowly where the charge level did not change very much either. When high wear happened at about 68th minute, the electrostatic sensor detected this wear by giving a high charge level as well. This situation happened again between 80 and 90 min. It is worth noting an increase

Scanning electron micrographs were taken of the new and worn ball surfaces. The chemical compositions of the worn silicon nitride pin surfaces were also inspected using an EDX analysis. Prior to the SEM examination the silicon nitride balls were gold coated to stop charging. Fig. 9 shows the FEG-SEM images of a new Si3 N4 ball surface at two different magnifications. It can be seen that the silicon nitride ball has fine acicular Si3 N4 grains with a size range of <1.5 ␮m in length and about 0.1 ␮m in diameter. Furthermore, the grain orientation is randomly distributed. Fig. 10 shows the SEM images of a worn Si3 N4 surface after being tested on XT3 under 120 N load. Other test conditions are as in Table 2. Fig. 10(a) shows the wear scar on the Si3 N4 ball. Under higher magnification in (b), holes and particles a few microns in size are seen on the surface. In Fig. 10(c), smaller sized holes (<1 ␮m) are seen. By closely looking at one of the particles in (b) under higher magnification (d), the micron sized particles are composed of sub-micron particles. Further elemental analysis by EDX has shown that these sub-micron particles include both Fe from the disc and Si3 N4 from the ball (see Fig. 11). Fig. 10(e) and (f) shows that another wear mechanism is involved at the initial wear stage, i.e. whole Si3 N4 grains plucked out from the ball surface perpendicular to the sliding direction. Since the pin wear rate is about 10−11 mm3 /N m and much lower than 10−6 mm3 /N m, it is within the mild wear regime, fine

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Fig. 9. FEG-SEM images of Si3 N4 new ball surface under different magnifications showing the fine acicular Si3 N4 grain distribution and size range: (a) lower magnification; (b) higher magnification.

Fig. 10. FEG-SEM images of Si3 N4 worn surface after sliding tested on XT3, 120 N load, 7 m/s sliding velocity, Mobil Jet II oil lubricated: (a) the wear scar, diameter ∼0.6 mm; (b) worn surface under higher magnification; (c) showing ‘holes’ on the worn surface; (d) showing a particle on the surface, which is a cluster of finer/sub-micron particles; (e) showing a fine acicular grain dropped out on the surface; (f) showing an acicular hole on the surface.

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Fig. 11. EDX analysis results Si3 N4 worn surface after sliding tested on XT3, 120 N load, 7 m/s sliding velocity, Mobil Jet II oil lubricated.

wear debris at smaller than the grain size by one-tenth or less is generated [12]. The worn silicon nitride surface is very smooth (see in Fig. 10), it shows that ‘polishing’ wear or tribo-chemical wear is involved at the initial or mild wear stage [26–29]. The fractured grains are in the sub-micron range and will probably act as a grinding paste. This procedure is illustrated in Fig. 12. Si3 N4 grains which are above or partially above the surface, will fracture and be removed by the sliding action or pulled out when asperity–asperity wear begins; grains which are on the top of the surface and across the sliding direction can be plucked out. Therefore, transgranular and intergranular cracking predominate at the initial wear period. However, as the procedure is repeated, more and larger debris are generated which could lead to severe wear. On the other hand, particles on the contact surfaces also cause abrasive wear which can be seen in Fig. 10(a), (b), (e) and (f). Similar wear models for silicon nitride in sliding wear have been suggested by Braza et al. [11] and Chen et al. [6]. The M50 has a totally different microstructure. FEG-SEM images of a M50 new and worn ball surfaces are shown in

Fig. 12. A schematic illustration of silicon nitride sliding on lubricated bearing steel initial wear mechanism: random distributed Si3 N4 grains near the surface are (a) broken; (b) pulled out as a whole/partial grain; (c) rolled out a whole grain.

Figs. 13 and 14, respectively. Typical adhesive and abrasive steel on steel sliding wear mechanisms were found in the form of cracking and delamination type wear leading to platelet wear particles.

Fig. 13. SEM image of M50 new ball.

Fig. 14. Image of M50 worn surfaces: tested on M50, 60 N load, 7 m/s sliding velocity, Mobil Jet II oil lubricated.

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5. Conclusions

References

Correlations between the simultaneous SWR and the wear rate of the disc, charge levels on the disc wear track or within the oil film covering the wear track, coefficient of friction, and the temperature close to the contact were found for silicon nitride on steel lubricated contacts. The lubricated wear rates of silicon nitride sliding against various bearing steels were found to be 10−7 to 10−11 mm3 /N m. The wear rates of various bearing steels are also evaluated and ranked relative to the M50 couple baseline. A correlation between the wear of silicon nitride and bearing steel under oil lubricated contacts and electrostatic charge levels has been found. This result encourages further investigation of electrostatic sensing technology for hybrid bearing condition monitoring. The conclusions from this study are summarised as follows:

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(a) Si3 N4 /M50NiL shows the best wear resistance under 120 N load while Si3 N4 /M50NiL DH was the best combination under 60 N load at low temperature (18–28 ◦ C), both have ∼10−11 and ∼10−9 mm3 /N m SWRs for the pin and disc, respectively. (b) Measurable electrostatic charge was generated during the lubricated pure sliding wear tests for the Si3 N4 and M50 balls on bearing steels. (c) Si3 N4 has at least two orders of magnitude less wear rate than M50. (d) Si3 N4 on steels has lower friction than that of M50 steel on steels. (e) Correlations between the wear of the hybrid contact and the electrostatic charge levels, the profile/depths of the wear track, the coefficient of friction, and the surface temperature have been found. (f) The wear mechanism for Si3 N4 on steel under lubrication was suggested to be started with polishing (tribo-chemical) wear including transgranular and intergranular cracking leading to sub-micron debris ejection. Spallation starts when larger particles dropped off from the silicon nitride ball. Typical sliding contact wear mechanisms were found for the steel on steel contacts.

Acknowledgements The authors would like to thank the European Commission and Rolls-Royce for funding this work within the ATOS project under Framework V. Acknowledgements also go to the project partners in the work package Rolls-Royce Derby, FAG and Hispano-Suiza for supplying the test materials, and FAG for their help with SEM scan information on silicon nitride. Finally the authors would like to thank Smiths Aerospace (ES-S) for supplying the electrostatic sensors and for their help with analysing and interpreting the electrostatic sensor data.

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