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Scuffing and rolling contact fatigue resistance of discrete laser spot hardened austempered ductile iron
Wear 422–423 (2019) 100–107
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Scuffing and rolling contact fatigue resistance of discrete laser spot hardened austempered ductile iron
T
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Ann Zammita, , Stephen Abelaa, John Charles Bettsa, Remigiusz Michalczewskib, Marek Kalbarczykb, Maurice Grecha a b
Department of Metallurgy and Materials Engineering, Faculty of Engineering, University of Malta, Msida MSD 2080, Malta Tribology Department, Institute for Sustainable Technologies - National Research Institute (ITeE-PIB), Radom, Poland
A R T I C LE I N FO
A B S T R A C T
Keywords: Austempered ductile iron Discrete laser spot hardening Back-tempering Scuffing wear Rolling contact fatigue
This study determined the influence of discrete laser spot hardening on the scuffing wear and rolling contact fatigue resistance of Cu-Ni austempered ductile iron (ADI) specimens. The frequency of the single laser pulses were such so as to produce patterns of separated spots, adjacent spots or overlapping spots. Lubricated sliding wear tests were carried out to determine the behaviour of ADI under starved lubrication conditions, usually occurring during the start-up of gears. ADI specimens hardened with adjacent and separated laser spots displayed higher resistance to scuffing than the surfaces treated with overlapping laser spots and the control samples (those tested in the as-austempered condition). Results obtained also show that laser hardening increased the average rolling contact fatigue life over the as-austempered ductile iron specimens. This improved performance was attributed to the higher volume of martensite formed and hence higher surface hardness, and also the residual compressive stresses induced in the laser-treated surfaces. This benefit outweighed the negative effect of the slight roughness induced by the laser treatment.
1. Introduction Following austempering, nodular cast iron results in the formation of austempered ductile iron. This iron has a microstructure consisting mainly of ferrite, austenite and nodular graphite. ADI exhibits an excellent combination of mechanical properties such as strength and ductility [1,2]. High-energy sources, such as laser beams, electron beams, high frequency pulsed generators used for induction heating, flames and tungsten inert gas arcs, are sometimes used to attain a high level of hardness in the near-surface region of engineering components [3,4]. When hardening ferrous alloys, laser beams exhibit advantages and yield higher qualities compared to other high-energy sources [5,6]. The laser surface hardening (LSH) of steels and cast irons results in the transformation of the microstructure to austenite, followed by the formation of martensite resulting from self-quenching and rapidly cooling to room temperature. It is also possible to have some retained austenite in the final microstructure [7–9]. A number of studies have been carried out on the laser hardening of austempered ductile iron (ADI) [8–14]. These show an increase in hardness to around 620 – 1030 HV to depths ranging from 0.1 to
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1.8 mm. The high hardness after LSH is expected to improve the tribological characteristics of austempered ductile iron (ADI) components. In fact, dry sliding wear tests by Roy and Manna [12] showed that LSH significantly enhanced the wear resistance of unalloyed ADI specimens. This was attributed to the high surface hardness and also the compressive residual stresses induced in the surface. Tan et al. [13] also showed that when laser-hardened, Cu-Mo ADI exhibited a lower weight loss by dry abrasive wear. Lubricated sliding tests conducted by Xu and Lu [9] showed the contact fatigue limit of laser-hardened Cu-Mo ADI to be 14% higher than that of the as-austempered counterparts. The improvement in wear resistance following laser hardening of ADI is not always evident. The high loads induced during testing can result in a strain-induced austenite to martensite transformation at the surface of the specimens. This yields an as-austempered microstructure similar to that of laser-hardened ADI and consequently similar wear characteristics. In fact, dry contact fatigue tests carried out by Lu and Zhang [7] showed that the wear rate of the as-austempered DI was slightly higher than the wear rate of the laser-hardened counterparts. Laser hardening of large surfaces frequently require several adjacent tracks [15], which will lead to overlapping, back-tempering and a loss in hardness in the area between the tracks (Fig. 1) [16].
Corresponding author. E-mail address: [email protected] (A. Zammit).
https://doi.org/10.1016/j.wear.2019.01.061 Received 12 October 2018; Received in revised form 10 January 2019; Accepted 11 January 2019 Available online 14 January 2019 0043-1648/ © 2019 Elsevier B.V. All rights reserved.
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Table 2 Laser processing parameters. Laser processing parameter
Value
Beam power (W) Beam diameter (mm) Beam pulse duration (ms) Nitrogen gas flow rate (l/min)
600 1.75 300 7.5
Fig. 1. Schematic effect of back-tempering on surface hardness [16].
Such a disadvantage can be overcome by laser hardening the surface using discrete spots. Similar processing has been carried out on ferrous materials such as grey cast iron, ductile iron, steel and stainless steel [17–21]. One of the first studies was conducted by Papaphilippou and Jeandin [20] who reported a 300% increase in the wear resistance of ductile iron following laser spot hardening. The laser-hardened specimens were treated with a CO2 laser in pulsed mode and the spots were evenly distributed using six different patterns. No significant difference in the wear rates was noted for the different patterns with different coverage percentages (10%, 20%, 50% and 100%). Another study by Xue et al. [18] showed that discrete laser spot hardening doubled the unlubricated sliding wear resistance of AISI O1, D2 and 4340 steels. In addition, the samples with only 20–40% irradiated coverage showed similar wear rates as those with 100% coverage. These and other studies [18–20] show that complete surface coverage is not necessary for optimum wear resistance and thus shorter production cycles can be utilised. In addition, in case of discrete surface hardening, the areas in between the hardened zones can act as lubricant reservoirs, which helps to decrease the friction coefficient between mating surfaces, thus improving the wear resistance. Several studies have been carried out to investigate phase transformations and hardness distribution of ADI after laser surface engineering using continuous waves. The authors from ref [14] have previously published work aimed at optimising the laser processing parameters for hardening spots on Cu-Ni ADI. Tribo-studies have not been conducted on ADI hardened using discrete laser spots, and hence this study will focus on the scuffing and rolling contact fatigue characteristics of laser-hardened Cu-Ni ADI using hardened spots with different distances between spots.
Fig. 2. Different laser spot arrays: (a) separate spots (LS), adjacent spots (LA) and (c) overlapping spots (LO).
2.2. Scuffing tests Scuffing tests were carried out using a pin-on-disk tribometer. In this configuration, 5 mm diameter cylindrical pins were machined from as-austempered or laser-hardened ADI (LS, LA and LO) while counterpart disks were made from hardened AISI D2 steel disks. The disks were hardened up to 60 HRC in a vacuum furnace using a quench and tempering heat treatment. Tests were carried out under boundary lubrication conditions using AGIP Rotra LSX 75W-90 oil. Boundary lubrication was created by first spreading the oil on the disk through a nozzle, then rotating the disk at 600 rpm for 5 s to remove the excess lubrication. During the test, a constant pressure of 10 MPa was applied and a rotational speed of 1450 rpm was maintained. In order to determine a suitable applied load, a progressive load test was first carried out. 10 MPa was chosen since it was the minimum load with which scuffing occurred within reasonable test duration and without creating plastic deformation on the specimens. On the other hand, the speed was selected based on scuffing tests carried out on gears using an FZG gear test machine [22]. The test was stopped when scuffing occurred. This was characterised by a sharp increase in temperature and friction, a rise in material loss from the pin together with a sharp increase in sound emitted from the pin-on-disk assembly. Tests were conducted at an ambient temperature of 18.3 ± 0.9 °C and relative humidity of 42.4 ± 1.6%. The results show the number of cycles sustained before scuffing occurred and are an average of the five repeats per test condition.
2. Materials and methods 2.1. Heat treatment and laser surface processing The ductile iron used in this study had the composition shown in Table 1. Austempering was carried out by first austenitising at 900 ± 2 °C for two hours, followed by austempering in a salt bath kept at a temperature of 360 ± 5 °C for one hour and subsequently cooling in air to room temperature. Laser hardening was performed using a CO2 laser with a wavelength of 10.6 µm and a focal length of 190.5 mm coupled with a 4-axis CNC system. The discrete spots were produced using stationary laser pulses and processing parameters shown in Table 2. These parameters were established in a previous study [14]. ADI surfaces having different patterns of hardened spots were created. The difference between patterns was in the distance between the spots. Three different laser spot arrays were considered as shown in Fig. 2, namely: separated spots (the distance between the spots being equal to one spot diameter) (LS), adjacent (non-overlapping) spots (LA) and overlapping spots with approximately 50% overlap (LO).
2.3. Rolling contact fatigue tests Rolling contact fatigue tests were carried out using a modified T-03 four-ball pitting tester [23]. In this test, a load of 300 N which translates into a contact stress of 2.6 GPa using Hertz contact theory, was applied to an upper ADI cone. The load was selected by observing the wear track in preliminary tests carried out using various loads. 300 N was the minimum load which ensured the specimens were elastically and not plastically deformed during testing. The cone acted upon three lower 100Cr6 chrome alloy bearing steel balls with a diameter of 12.7 mm
Table 1 Chemical composition of ductile iron in wt%. Element
C
Si
Cu
Ni
Mn
P
Mg
Al
S
Fe
Composition
3.26
2.36
1.63
1.58
0.24
0.011
0.057
0.024
0.006
Bal.
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3.2. Spot matrix hardening
and a roughness value Ra of 0.032 µm. These balls were placed in a race filled with lubricating oil (Shell Transaxle gear oil 75W-90). The cone rotational speed was set at 1450 rpm, so as to have a similar rotational speed as the scuffing tests. At least fifteen runs were carried out for two types of processing namely: polished as-austempered DI and laser-hardened ADI. Samples used to perform the latter test were laser-hardened using parameters which yielded the best results in the lubricated sliding wear tests. The tests were performed at an ambient temperature of 22 ± 3 °C and a relative humidity of 42 ± 4%. Tests were terminated when pitting occurred. This was taken to occur when the level of vibrations exceeded a predefined level, namely 150% the average vibration level that was recorded during the first three minutes of the test. The time to pitting was recorded for each run and then converted into number of cycles. The contact fatigue life of each surface type was determined using the Weibull analysis method.
Discrete laser spots were generated on the entire surface of the ADI specimen. Three different spot patterns were applied, namely separated spots (LS), adjacent spots (LA) and overlapping spots (LO) as shown in Fig. 4. The white dashed circles shown in the images pertaining to the LS (Fig. 4a) and LA (Fig. 4b) specimens represent the laser spots. The overlapping spots for the LO (Fig. 4c) specimens are not clearly identifiable, and hence are not marked in the image. Fig. 5 shows the surface roughness readings for the three different spot patterns. Interestingly, surfaces with overlapping spots exhibited the highest roughness value (Ra: 0.70 ± 0.35 µm; RZ: 6.56 ± 1.81 µm). In comparison, surfaces with separate and adjacent laser spots displayed lower Ra roughness values of 0.43 ± 0.14 µm (RZ: 3.08 ± 0.21 µm) and 0.41 ± 0.08 µm (RZ: 2.69 ± 0.60 µm) respectively. It is noted that these roughness values are much lower than that of laser melted surfaces which typically exhibited Ra values of 1.34 ± 0.39 µm [14]. Fig. 6 shows the hardness profile across one single spot, which represents the hardness for specimens having separated laser spots. Fig. 7 presents the hardness profile taken through the centre of three adjacent laser-hardened spots. It can be noted that the hardness increases to values of around 700–800 HV in the laser irradiated area. The lower value of 370 HV denotes the hardness of the parent as-austempered ausferritic matrix, outside the laser-hardened spots. On the other hand, if measurements were taken along a line horizontally or vertically displaced from the centre line of the spot, the hardness values were lower. This means surfaces with adjacent spots contained regions of hard martensite (the irradiated spots), and softer ausferrite between the spots (the parent substrate material). Fig. 8 shows the hardness profile taken along the centre of six overlapping spots. The hardness varies between 450 and 650 HV. The softer regions can be attributed to back-tempering, a phenomenon commonly reported with overlapping multi-track hardening [15,16,24,25]. The heat input associated with a subsequent spot tempers the structure generated by the preceding spot, thus lowering the hardness. Consequently, the laser-hardened surface consists of regions of martensite and other regions of tempered martensite. Contrary to what was reported [15], no cracks were observed on the surface with overlapping laser spots. This implies better quality surfaces. The work shows that both an adjacent and overlapping spot hardening would result in only a fraction of the overall surface area being hardened. This results in regions of varying hardness. Zhang et al. [19] state that this is beneficial as the softer areas serve as lubricant reservoirs. These help in decreasing the coefficient of friction, hence improving the wear resistance. Jiang et al. [26] also suggest that one could use another processing strategy, one in which a second hardening pattern is applied to areas that were not hardened by the first hardening process. However, the latter could result in tempering the initial matrix of hardened spots. The reduced coverage and reduced energy input associated with surfaces hardened using adjacent spots rather than overlapping spots will lower the overall processing costs resulting from shorter processing time and lower use of consumables. The lower heat input would also lead to less heat distortion [18,19,27,28]. In fact, the surface roughness (Ra) generated by overlapping spots was higher by 0.70 µm, from that of adjacent spots which had a value of 0.31 µm (Fig. 5).
2.4. Material characterisation Surface analysis was carried out using a Nikon Optiphot-100 optical microscope and a Carl Zeiss Merlin scanning electron microscope. Micro-hardness tests were carried out using a Mitutoyo MVK-H12 testing machine, a load of 50 gf and a dwell time of 10 s. Mean surface roughness Ra were taken using a Mitutoyo Surftest501 profilometer. The roughness results are an average of at least three values. The roughness of the surfaces used for the rolling contact fatigue tests was measured before and after the wear tests using a Taylor Hobson TALYSURF CCI non-contact optical interferometer. This equipment was also used to obtain three-dimensional maps of the wear tracks formed and using multiple scans. The surface area, volume and profile of pits were obtained using Taylor Hobson Form Talysurf contact profilometer. All scans were integrated and analysed using MountainsMap® software. 3. Results and discussion 3.1. Single laser spot Single laser-hardened spots had a martensitic structure, that on occasions was supplemented by a ledeburitic structure surrounding the graphite nodules [14]. The nodules were unaltered as shown in Fig. 3. The surface hardness of the laser spots was approximately 770 HV and the hardened depth was circa 150 µm. The resulting martensitic structure was shown to induce compressive stresses in the hardened zone. The surface roughness, Ra of polished as-austempered ductile iron specimens increased from 0.15 ± 0.08 µm to 0.43 ± 0.14 µm as a result of laser hardening using one spot.
3.3. Scuffing resistance Lubricated sliding wear tests were carried out to determine the behaviour of discrete laser spot hardened ADI tested under starved lubrication conditions. As-austempered DI (A) and LS, LA, and LO laserhardened ADI specimens were investigated. Fig. 9 shows that during starved lubricating wear tests, surfaces hardened with separate spots (LS) and with adjacent spots (LA) survived a longer number of cycles
Fig. 3. SEM micrograph of the laser hardened surface (shown for comparative purposes) [14]. 102
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Fig. 4. Spot matrix laser hardening of ADI specimens showing (a) separated, (b) adjacent and (c) overlapping spots; unetched.
Fig. 5. Surface roughness values for different patterns of spot laser hardening.
Fig. 6. Microhardness profile across the surface of one single spot [14].
than the laser-hardened surfaces with overlapping spots (LO) and also the control as-austempered (A) specimens. Surfaces hardened with adjacent spots (LA) survived the longest number of cycles namely 1.5 × 105 cycles, whilst those having separated spots (LS) specimens survived 2.7 × 104 cycles. This is much higher than the 2.3 × 103 cycles survived by the as-austempered DI (A) specimens. The surface with adjacent spots contains the largest volume of the martensitic structure and only a small fraction of unhardened regions. This provides a higher wear resistance, and hence results in a longer number of cycles before failure. In addition, a large volume fraction of martensite implies that compressive stresses are present over a larger area, both at the surface and at sub-surface locations. Compressive stresses formed after LSH delay crack initiation and propagation [6]. Moreover, the coefficient of friction for LA specimens is the lowest
and about 0.08. This value remains constant throughout the sliding wear test (Fig. 10). In contrast, the friction coefficient for the LS specimens increases in steps as the test progresses. This can be attributed to the scuffing failure of the soft ausferritic structure located between the laser spots. As the test continues, the austenite in the ausferrite can transform to martensite, stopping any further increase in the coefficient of friction. In fact, the surface hardness measured after the wear test increased to around 560 HV, suggesting the change in phase from austenite to martensite. This phenomenon may occur a number of times depending on lubricant film failure and every time martensite is formed a low coefficient of friction results. The cold weld junctions formed during testing, between the martensitic pin surface and the hard disk break easily, resulting in lower adhesion between the two surfaces [29].
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Fig. 9. Sliding cycles before scuffing for as-austempered (A) and laser-hardened specimens (LS, LA, LO).
Fig. 7. Microhardness profile across the surface of three adjacent LSH spots.
Fig. 10. Friction coefficient for tested specimens.
regions between the spots, were ausferrite, which can strain harden and improve further the resistance to scuffing failure. Notwithstanding this, the rougher surfaces of the LO specimens were expected to create lubricant pockets, leading to a higher scuffing resistance than the asaustempered specimens. It appears however that the austenite to martensite transformation and absence of lubrication pockets in the asaustempered specimens is counterbalanced by the lower surface hardness of the tempered martensitic regions and the rougher surfaces of the LO specimens. Previously, it has been shown that laser hardening improves the abrasion wear resistance [13] and dry sliding wear resistance [7,12] of the upper ausferritic ADI by approximately 40% and 70% respectively. In the present study, it was shown that discrete spot laser hardening will improve the sliding wear resistance of ADI under starved lubrication conditions. Papaphilippou and Jeandin [20] reported a 17% increase in the wear resistance when spot laser hardening ductile iron. However, these researchers reported no significant difference in the wear rates of surfaces with different laser treatment patterns and different surface coverage percentages. It must be pointed out that tests were carried out under dry conditions and hence the beneficial effect of pockets of lubricants could not be taken advantage of [20]. Similar observations were reported by Xue et al. [18] who studied the influence of discrete laser spot hardening of different steels under unlubricated conditions. The samples that were irradiated with 100% coverage exhibited similar wear rates as those with 20–40% coverage.
Fig. 8. Typical microhardness profile across the surface of six overlapping LSH spots.
The martensite present in the laser-hardened regions provides better heat dissipation due to its good thermal conductivity leading to a reduction in the CoF [30,31]. The initial Ra value of LA specimens was circa 0.31 µm, which is higher than that of the LS specimens which has a roughness of approximately 0.15 µm. The dimples present in specimens with a higher roughness can serve as lubricant reservoirs, which help to decrease the friction coefficient and provide a higher scuffing resistance [19]. Extrapolating, then surfaces with overlapping laser spots which exhibited an even higher surface roughness at circa 0.70 µm, should have resulted in an even lower coefficient of friction and a high resistance to scuffing failure. However, as can be seen in Fig. 9, these specimens exhibited a high coefficient of friction and scuffing resistance comparable to that of as-austempered samples. It was shown in Fig. 8 that LO specimens had very few hard martensitic regions and a high volume fraction of tempered martensite with a hardness of approximately 450 HV. In comparison, in LS and the LA specimens, the softer 104
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Fig. 11. SEM micrographs showing cracking and pitting on typical worn surfaces.
3.3.1. Wear mechanism The worn surfaces showed signs of severe wear and adhesion between the two sliding bodies (Fig. 11). Scuffing is the result of microwelding of the surfaces due to the lack of lubrication, since it is not possible for the oil to separate the two sliding surfaces. Consequently, the asperities weld together due to the applied pressure and excessive heat generated by friction. As the two surfaces continue with sliding, the surfaces alternately weld and tear metal from each other's surfaces. This results in the formation of cracks and pits on the scuffed surfaces. Fig. 12 also shows peeling of the surface as well as surface delamination, depicted by the detachment of thin platelets from the surface. Additionally, evidence of shearing and plastic flow was noted in the sub-surface of the test specimens (Fig. 13). A distorted ausferritic microstructure underneath the martensitic microstructure is visible. The strain hardening of the microstructure discussed in Section 3.3 is caused by plastic deformation, resulting in a stronger material which causes surface flow and the microstructure to distort. Propagation of the cracks was influenced by graphite nodules, as shown in Fig. 14. The weak interface and low strength between the graphite and the metal matrix has a major influence on the fracture path and lives of the specimens.
Fig. 13. Cross-section of typical worn surface showing cracks and a distorted ausferritic microstructure.
counterparts. The L10 and L50 fatigue lives, which were obtained from the Weibull probability plot in Fig. 17, are respectively 4.3 × 105 and 9.2 × 105 cycles. The average fatigue life of the laser-hardened ADI specimens is 6.9 × 105 cycles greater than the fatigue life of the asaustempered DI specimens. Xu and Lu [9] conducted contact fatigue testing of ADI specimens austempered at 370 °C and loaded at a stress of 2.3 GPa. They showed that the improvement brought about as a result of laser hardening is only 0.5 × 105 cycles when carrying out. The discrepancy in the two studies may be attributed to a number of factors. For example, Xu and Lu [9] do not report the surface finish of either the as-austempered or the laser-hardened specimens tested. Also, they did not observe the hard bull's eye ledeburitic structure around the graphite nodules that was noted in the present study. The high hardness of the ledeburitic structure may be one of the reasons for the better results reported in the present study. The polished as-austempered specimens had a surface roughness (Ra) of 0.05 µm, which increased to 0.12 µm after laser hardening. With these roughness values, the specific film thickness λ was calculated to be 1.12. A value of λ above 1 denotes that rolling is taking place in the full lubricated regime. However, considering this value is very close to 1, it is then likely that at some point/s during the test this value changes and reaches a value that is lower than 1. In reality, this implies that the condition of point contact assumed when calculating λ was true. Under conditions of general surface contact or occasional contact of asperities, the value of λ is lower than 1 and mixed lubrication occurs. This would imply that the laser-hardened specimens were being tested under full/ mixed lubrication conditions. The specific film thickness for the smoother as-austempered specimens was calculated to have a value of 2.81, meaning that these specimens were tested under full lubrication, which should give the
3.4. Rolling contact fatigue resistance Rolling contact fatigue (RCF) tests were carried out on as-austempered ADI specimens and on laser-hardened ADI specimens that were irradiated using adjacent laser spots and which exhibited the highest resistance to scuffing. The number of cycles to failure was recorded when pitting occurred on the specimens. Fig. 15 shows a typical pit formed after testing a laser hardened specimen, while Fig. 16 shows numerous cracks present on a typical wear track. Fig. 17 shows that the laser-hardened ADI specimens exhibited the better rolling contact fatigue performance than their as-asutempered
Fig. 12. SEM micrograph showing delamination of typical worn surfaces. 105
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Fig. 14. Influence of graphite nodules on the crack propagation in typical worn surfaces.
specimens a longer life. However, in this work the laser-treated specimens performed better, which can be attributed to the higher surface hardness (~ 770 HV) compared to that of the as-austempered specimens (~ 370 HV). The compressive stresses present in the martensitic regions of the laser-hardened specimens [14] serve to reduce the maximum shear stress when the specimens are in loaded contact. This delays crack nucleation and propagation and as a result increases the fatigue life. The RCF performance of the laser-hardened ADI is superior to that reported in some other studies for carburised steel. For example, Oila and Bull [32], Widmark and Melander [33] and Cavallaro et al. [34] obtained RCF lives respectively of 0.5 × 105, 3 × 105 and 2 × 105 cycles when testing at an applied stress of approximately 2.3 GPa. These values are lower than the average RCF life of 6.9 × 105 cycles obtained for the laser-hardened ADI in the present study. In another study on carburised steel, Ramanathan and Radhakrishnan [35] report an RCF life of 20 × 105 cycles when testing at a stress of 2.5 GPa. These values are comparable to those in this study. Fatigue values reported by other researchers [36,37] have fatigue lives of approximately 500 × 105 cycles. This is much higher than that reported in the present study even if the applied stress is only 1.7 GPa. Same applies to results by Hongbin et al. [38] who reported an RCF value 50 × 105 cycles when testing at an applied stress of 3.05 GPa. The above results show that the RCF life of carburised steel varies widely and depends on the parameters of the carburising treatment and the surface integrity of the specimens following carburising. It has been shown that laser hardening may be a suitable treatment for improving the rolling contact fatigue life of ADI. Laser-treated ADI surfaces have RCF values comparable to that reported by a number of researchers studying carburised steel, a material known to be suitable for engineering components, such as gears.
Fig. 16. Cracks present on a wear track after rolling contact fatigue testing a laser- hardened specimen.
4. Conclusions This study was carried out in order to determine the influence of discrete laser spot hardening on the scuffing wear resistance and rolling contact fatigue resistance of a Cu-Ni austempered ductile iron (ADI). The main conclusions arising from this study are the following: 1. Discrete laser spot hardening is a suitable technique for improving the scuffing wear and the rolling contact fatigue resistance of upper ausferritic austempered ductile iron. 2. Lubricated sliding wear tests under starved lubrication conditions revealed that the laser-hardened ADI specimens treated with adjacent laser spots performed better than the as-austempered and laser-hardened samples treated with other spot patterns. The higher
Fig. 15. SEM and 3D profilometer images of the wear tracks following contact fatigue testing of a laser- hardened specimen. 106
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[6] J.C. Ion, Laser Processing of Engineering Materials, Elsevier ButterworthHeinemann, 2005. [7] G.-X. Lu, H. Zhang, Sliding wear characteristics of austempered ductile iron with and without laser hardening, Wear 138 (1990) 1–12. [8] C. Soriano, J. Leunda, J. Lambarri, V.G. Navas, C. Sanz, Effect of laser surface hardening on the microstructure, hardness and residual stresses of austempered ductile iron grades, Appl. Surf. Sci. 257 (2011) 7101–7106. [9] D.-X. Xu, G.-X. Lu, Contact fatigue behaviour of an austenitic-bainitic ductile iron, Wear 169 (1993) 153–159. [10] S.K. Putatunda, L. Bartosiewicz, R.J. Hull, M. Lander, Laser hardening of austempered ductile cast iron (ADI), Mater. Manuf. Process. 12 (1997) 137–151. [11] A. Roy, I. Manna, Mathematical modeling of localized melting around graphite nodules during laser surface hardening of austempered ductile iron, Opt. Lasers Eng. 34 (2000) 369–383. [12] A. Roy, I. Manna, Laser surface engineering to improve wear resistance of austempered ductile iron, Mater. Sci. Eng. vol. A297, (2001) 85–93. [13] Y.-H. Tan, S.-I. Yu, J.-L. Doong, J.-R. Hwang, Abrasive wear property of bainitic nodular cast iron in laser processing, J. Mater. Sci. 25 (1990) 4133–4139. [14] A. Zammit, S. Abela, J.C. Betts, M. Grech, Discrete laser spot hardening of austempered ductile iron, Surf. Coat. Technol. 331 (2017) 143–152. [15] ASM International, Friction, Lubrication and Wear Technology, vol. 18, (1992). [16] L. Giorleo, B. Previtali, Q. Semeraro, Modelling of back tempering in laser hardening, Int. J. Adv. Manuf. Technol. 54 (2011) 969–977. [17] D.I. Pantelis, G. Pantazopoulos, S.S. Antoniou, Wear behavior of anti-galling surface textured gray cast iron using pulsed-CO2 laser treatment, Wear 205 (1997) 178–185. [18] L. Xue, M.U. Islam, G. McGregor, Dot matrix hardening of steels using a fiber optic coupled pulsed Nd: YAG laser, Mater. Manuf. Process. 14 (1999) 53–65 (1999). [19] X.M. Zhang, H.C. Man, H.D. Li, Wear and friction properties of laser surface hardened EN31 steel, J. Mater. Process. Technol. 69 (1997) 162–166. [20] C. Papaphilippou, M. Jeandin, Spot laser hardening, J. Mater. Sci. Lett. 15 (1996) 1064–1066. [21] L. Tricarico, A. Ancona, G. Palumbo, D. Sorgente, R. Spina, P.M. Lugara, Discrete spot laser hardening and remelting with a high-brilliance source for surface structuring of a hypereutectoid steel, Mater. Des. 115 (2017) 194–202. [22] Standard Test Method for Evaluating the Scuffing Load Capacity of Oils (FZG Visual Method), ASTM D 5182 - 97, Reapproved, 2002. [23] R. Michalczewski, W. Piekoszewski, M. Szczerek, J. Wulczyński, A method for the assessment of the rolling contact fatigue of modern engineering materials in lubricated contact, Trans. Famena 36 (2012) 39–48. [24] R.S. Lakhkar, Y.C. Shin, M.J.M. Krane, Predictive modeling of multi-track laser hardening of AISI 4140 steel, Mater. Sci. Eng.: A 480 (2008) 209–217. [25] G. Tani, L. Orazi, A. Fortunato, Prediction of hypo eutectoid steel softening due to tempering phenomena in laser surface hardening, CIRP Ann. - Manuf. Technol. 57 (2008) 209–212. [26] J. Jiang, L. Xue, S. Wang, Discrete laser spot transformation hardening of AISI O1 tool steel using pulsed Nd:YAG laser, Surf. Coat. Technol. 205 (2011) 5156–5164. [27] P.R. Woodard, J. Dryden, Thermal analysis of a laser pulse for discrete spot surface transformation hardening, J. Appl. Phys. 85 (1999) 2488–2496. [28] K.A. Khashan, O.A. Hamadi, Features of spot-matrix surface hardening of lowcarbon steel using pulsed laser, J. Eng. Technol. 25 (2007) 168–175. [29] M.O.A. Mokhtar, M. Zaki, G.S.A. Shawki, Effect of mechanical properties on frictional behaviour of metals, Tribol. Int. (1979) 165–267. [30] A. Gural, Influence of martensite particle size on dry sliding wear behaviour of low carbon dual phase powder metallurgy steel, Met. Mater. 48 (2010) 25–31. [31] S. Sendooran, P. Raja, Metallurgical investigation on cryogenic treated HSS tool, Int. J. Eng. Sci. Technol. 3 (2011) 3992–3996. [32] A. Oila, S.J. Bull, Assessment of the factors influencing micropitting in rolling/ sliding contacts, Wear 258 (2005) 1510–1524. [33] M. Widmark, A. Melander, Effect of material, heat treatment, grinding and shot peening on contact fatigue life of carburised steels, Int. J. Fatigue 21 (1999) 309–327. [34] G.P. Cavallaro, T.P. Wilks, C. Subramanian, K.N. Strafford, P. French, J.E. Allison, Bending fatigue and contact fatigue characteristics of carburized gears, Surf. Coat. Technol. 71 (1995) 182–192. [35] S. Ramanathan, V.M. Radhakrishnan, Investigation of rolling contact fatigue damage of a case-carburized low alloy steel, Wear 45 (1977) 323–333. [36] D.P. Townsend, E.V. Zaretsky, Effect of shot peening on surface fatigue life of carburized and hardened AISI 9310 spur gears, NASA Tech. Pap. 2047 (1982) 5–12. [37] J. Champaigne, Shot peening overview, Metal Improvement Company, 2001. [38] X. Hongbin, C. Qing, S. Eryu, W. Dengzhen, C. Zhaohong, W. Zhengle, The effect of shot peening on rolling contact fatigue behaviour and its crack initiation and propagation in carburized steel, Wear 151 (1991).
Fig. 17. Weibull probability plot for data obtained from rolling contact fatigue tests.
life provided by the former laser-hardened samples is associated with a load carrying capacity, as a result of the higher volume fraction of martensite on the surface, and also the residual compressive stresses induced in the laser-hardened surface. 3. The process of laser hardening with overlapping spot matrices caused back-tempering leading to regions of martensite and of tempered martensite. 4. ADI specimens treated with laser-hardened spots separated from each other survived 2.7 × 104 cycles before scuffing, which is also an improvement over the performance exhibited by their counterpart as-austempered specimens. In comparison, the surfaces irradiated with overlapping laser-hardened spots displayed the poorest performance, with their endurance limit roughly equal to that of asaustempered specimens. This was attributed to the presence of the softer tempered martensitic regions. 5. It was also shown that laser hardening consisting of adjacent spots increases the rolling contact fatigue life by 6.9 × 105 cycles over that of as-austempered counterparts. This improvement was obtained as a result of the harder laser-treated surfaces and the residual compressive stresses, which outweighed the negative effect of the rougher surface. Acknowledgements The authors would like to acknowledge the ERDF funding for the purchase of the testing equipment through the project: "Developing an Interdisciplinary Material Testing and Rapid Prototyping R&D Facility (Ref. no. 012)". References [1] R.A. Harding, The production, properties and automotive applications of austempered ductile iron, Kov. Mater. 45 (2007) 1–16. [2] AGMA 939-A07, Austempered Ductile Iron for Gears, ed: AGMA, 2007, pp. 1–10. [3] H. Zhang, Y. Shi, C.Y. Xu, M. Kutsuna, Surface hardening of gears by laser beam processing, Surf. Eng. 19 (2003) 134–136. [4] W.M. Steen, J. Powell, Laser surface treatment, Mater. Eng. 2 (1981) 157–162. [5] W.M. Steen, Laser Material Processing, Springer, London, 2003.
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Scuffing and rolling contact fatigue resistance of discrete laser spot hardened austempered ductile iron
T
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Ann Zammita, , Stephen Abelaa, John Charles Bettsa, Remigiusz Michalczewskib, Marek Kalbarczykb, Maurice Grecha a b
Department of Metallurgy and Materials Engineering, Faculty of Engineering, University of Malta, Msida MSD 2080, Malta Tribology Department, Institute for Sustainable Technologies - National Research Institute (ITeE-PIB), Radom, Poland
A R T I C LE I N FO
A B S T R A C T
Keywords: Austempered ductile iron Discrete laser spot hardening Back-tempering Scuffing wear Rolling contact fatigue
This study determined the influence of discrete laser spot hardening on the scuffing wear and rolling contact fatigue resistance of Cu-Ni austempered ductile iron (ADI) specimens. The frequency of the single laser pulses were such so as to produce patterns of separated spots, adjacent spots or overlapping spots. Lubricated sliding wear tests were carried out to determine the behaviour of ADI under starved lubrication conditions, usually occurring during the start-up of gears. ADI specimens hardened with adjacent and separated laser spots displayed higher resistance to scuffing than the surfaces treated with overlapping laser spots and the control samples (those tested in the as-austempered condition). Results obtained also show that laser hardening increased the average rolling contact fatigue life over the as-austempered ductile iron specimens. This improved performance was attributed to the higher volume of martensite formed and hence higher surface hardness, and also the residual compressive stresses induced in the laser-treated surfaces. This benefit outweighed the negative effect of the slight roughness induced by the laser treatment.
1. Introduction Following austempering, nodular cast iron results in the formation of austempered ductile iron. This iron has a microstructure consisting mainly of ferrite, austenite and nodular graphite. ADI exhibits an excellent combination of mechanical properties such as strength and ductility [1,2]. High-energy sources, such as laser beams, electron beams, high frequency pulsed generators used for induction heating, flames and tungsten inert gas arcs, are sometimes used to attain a high level of hardness in the near-surface region of engineering components [3,4]. When hardening ferrous alloys, laser beams exhibit advantages and yield higher qualities compared to other high-energy sources [5,6]. The laser surface hardening (LSH) of steels and cast irons results in the transformation of the microstructure to austenite, followed by the formation of martensite resulting from self-quenching and rapidly cooling to room temperature. It is also possible to have some retained austenite in the final microstructure [7–9]. A number of studies have been carried out on the laser hardening of austempered ductile iron (ADI) [8–14]. These show an increase in hardness to around 620 – 1030 HV to depths ranging from 0.1 to
⁎
1.8 mm. The high hardness after LSH is expected to improve the tribological characteristics of austempered ductile iron (ADI) components. In fact, dry sliding wear tests by Roy and Manna [12] showed that LSH significantly enhanced the wear resistance of unalloyed ADI specimens. This was attributed to the high surface hardness and also the compressive residual stresses induced in the surface. Tan et al. [13] also showed that when laser-hardened, Cu-Mo ADI exhibited a lower weight loss by dry abrasive wear. Lubricated sliding tests conducted by Xu and Lu [9] showed the contact fatigue limit of laser-hardened Cu-Mo ADI to be 14% higher than that of the as-austempered counterparts. The improvement in wear resistance following laser hardening of ADI is not always evident. The high loads induced during testing can result in a strain-induced austenite to martensite transformation at the surface of the specimens. This yields an as-austempered microstructure similar to that of laser-hardened ADI and consequently similar wear characteristics. In fact, dry contact fatigue tests carried out by Lu and Zhang [7] showed that the wear rate of the as-austempered DI was slightly higher than the wear rate of the laser-hardened counterparts. Laser hardening of large surfaces frequently require several adjacent tracks [15], which will lead to overlapping, back-tempering and a loss in hardness in the area between the tracks (Fig. 1) [16].
Corresponding author. E-mail address: [email protected] (A. Zammit).
https://doi.org/10.1016/j.wear.2019.01.061 Received 12 October 2018; Received in revised form 10 January 2019; Accepted 11 January 2019 Available online 14 January 2019 0043-1648/ © 2019 Elsevier B.V. All rights reserved.
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Table 2 Laser processing parameters. Laser processing parameter
Value
Beam power (W) Beam diameter (mm) Beam pulse duration (ms) Nitrogen gas flow rate (l/min)
600 1.75 300 7.5
Fig. 1. Schematic effect of back-tempering on surface hardness [16].
Such a disadvantage can be overcome by laser hardening the surface using discrete spots. Similar processing has been carried out on ferrous materials such as grey cast iron, ductile iron, steel and stainless steel [17–21]. One of the first studies was conducted by Papaphilippou and Jeandin [20] who reported a 300% increase in the wear resistance of ductile iron following laser spot hardening. The laser-hardened specimens were treated with a CO2 laser in pulsed mode and the spots were evenly distributed using six different patterns. No significant difference in the wear rates was noted for the different patterns with different coverage percentages (10%, 20%, 50% and 100%). Another study by Xue et al. [18] showed that discrete laser spot hardening doubled the unlubricated sliding wear resistance of AISI O1, D2 and 4340 steels. In addition, the samples with only 20–40% irradiated coverage showed similar wear rates as those with 100% coverage. These and other studies [18–20] show that complete surface coverage is not necessary for optimum wear resistance and thus shorter production cycles can be utilised. In addition, in case of discrete surface hardening, the areas in between the hardened zones can act as lubricant reservoirs, which helps to decrease the friction coefficient between mating surfaces, thus improving the wear resistance. Several studies have been carried out to investigate phase transformations and hardness distribution of ADI after laser surface engineering using continuous waves. The authors from ref [14] have previously published work aimed at optimising the laser processing parameters for hardening spots on Cu-Ni ADI. Tribo-studies have not been conducted on ADI hardened using discrete laser spots, and hence this study will focus on the scuffing and rolling contact fatigue characteristics of laser-hardened Cu-Ni ADI using hardened spots with different distances between spots.
Fig. 2. Different laser spot arrays: (a) separate spots (LS), adjacent spots (LA) and (c) overlapping spots (LO).
2.2. Scuffing tests Scuffing tests were carried out using a pin-on-disk tribometer. In this configuration, 5 mm diameter cylindrical pins were machined from as-austempered or laser-hardened ADI (LS, LA and LO) while counterpart disks were made from hardened AISI D2 steel disks. The disks were hardened up to 60 HRC in a vacuum furnace using a quench and tempering heat treatment. Tests were carried out under boundary lubrication conditions using AGIP Rotra LSX 75W-90 oil. Boundary lubrication was created by first spreading the oil on the disk through a nozzle, then rotating the disk at 600 rpm for 5 s to remove the excess lubrication. During the test, a constant pressure of 10 MPa was applied and a rotational speed of 1450 rpm was maintained. In order to determine a suitable applied load, a progressive load test was first carried out. 10 MPa was chosen since it was the minimum load with which scuffing occurred within reasonable test duration and without creating plastic deformation on the specimens. On the other hand, the speed was selected based on scuffing tests carried out on gears using an FZG gear test machine [22]. The test was stopped when scuffing occurred. This was characterised by a sharp increase in temperature and friction, a rise in material loss from the pin together with a sharp increase in sound emitted from the pin-on-disk assembly. Tests were conducted at an ambient temperature of 18.3 ± 0.9 °C and relative humidity of 42.4 ± 1.6%. The results show the number of cycles sustained before scuffing occurred and are an average of the five repeats per test condition.
2. Materials and methods 2.1. Heat treatment and laser surface processing The ductile iron used in this study had the composition shown in Table 1. Austempering was carried out by first austenitising at 900 ± 2 °C for two hours, followed by austempering in a salt bath kept at a temperature of 360 ± 5 °C for one hour and subsequently cooling in air to room temperature. Laser hardening was performed using a CO2 laser with a wavelength of 10.6 µm and a focal length of 190.5 mm coupled with a 4-axis CNC system. The discrete spots were produced using stationary laser pulses and processing parameters shown in Table 2. These parameters were established in a previous study [14]. ADI surfaces having different patterns of hardened spots were created. The difference between patterns was in the distance between the spots. Three different laser spot arrays were considered as shown in Fig. 2, namely: separated spots (the distance between the spots being equal to one spot diameter) (LS), adjacent (non-overlapping) spots (LA) and overlapping spots with approximately 50% overlap (LO).
2.3. Rolling contact fatigue tests Rolling contact fatigue tests were carried out using a modified T-03 four-ball pitting tester [23]. In this test, a load of 300 N which translates into a contact stress of 2.6 GPa using Hertz contact theory, was applied to an upper ADI cone. The load was selected by observing the wear track in preliminary tests carried out using various loads. 300 N was the minimum load which ensured the specimens were elastically and not plastically deformed during testing. The cone acted upon three lower 100Cr6 chrome alloy bearing steel balls with a diameter of 12.7 mm
Table 1 Chemical composition of ductile iron in wt%. Element
C
Si
Cu
Ni
Mn
P
Mg
Al
S
Fe
Composition
3.26
2.36
1.63
1.58
0.24
0.011
0.057
0.024
0.006
Bal.
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3.2. Spot matrix hardening
and a roughness value Ra of 0.032 µm. These balls were placed in a race filled with lubricating oil (Shell Transaxle gear oil 75W-90). The cone rotational speed was set at 1450 rpm, so as to have a similar rotational speed as the scuffing tests. At least fifteen runs were carried out for two types of processing namely: polished as-austempered DI and laser-hardened ADI. Samples used to perform the latter test were laser-hardened using parameters which yielded the best results in the lubricated sliding wear tests. The tests were performed at an ambient temperature of 22 ± 3 °C and a relative humidity of 42 ± 4%. Tests were terminated when pitting occurred. This was taken to occur when the level of vibrations exceeded a predefined level, namely 150% the average vibration level that was recorded during the first three minutes of the test. The time to pitting was recorded for each run and then converted into number of cycles. The contact fatigue life of each surface type was determined using the Weibull analysis method.
Discrete laser spots were generated on the entire surface of the ADI specimen. Three different spot patterns were applied, namely separated spots (LS), adjacent spots (LA) and overlapping spots (LO) as shown in Fig. 4. The white dashed circles shown in the images pertaining to the LS (Fig. 4a) and LA (Fig. 4b) specimens represent the laser spots. The overlapping spots for the LO (Fig. 4c) specimens are not clearly identifiable, and hence are not marked in the image. Fig. 5 shows the surface roughness readings for the three different spot patterns. Interestingly, surfaces with overlapping spots exhibited the highest roughness value (Ra: 0.70 ± 0.35 µm; RZ: 6.56 ± 1.81 µm). In comparison, surfaces with separate and adjacent laser spots displayed lower Ra roughness values of 0.43 ± 0.14 µm (RZ: 3.08 ± 0.21 µm) and 0.41 ± 0.08 µm (RZ: 2.69 ± 0.60 µm) respectively. It is noted that these roughness values are much lower than that of laser melted surfaces which typically exhibited Ra values of 1.34 ± 0.39 µm [14]. Fig. 6 shows the hardness profile across one single spot, which represents the hardness for specimens having separated laser spots. Fig. 7 presents the hardness profile taken through the centre of three adjacent laser-hardened spots. It can be noted that the hardness increases to values of around 700–800 HV in the laser irradiated area. The lower value of 370 HV denotes the hardness of the parent as-austempered ausferritic matrix, outside the laser-hardened spots. On the other hand, if measurements were taken along a line horizontally or vertically displaced from the centre line of the spot, the hardness values were lower. This means surfaces with adjacent spots contained regions of hard martensite (the irradiated spots), and softer ausferrite between the spots (the parent substrate material). Fig. 8 shows the hardness profile taken along the centre of six overlapping spots. The hardness varies between 450 and 650 HV. The softer regions can be attributed to back-tempering, a phenomenon commonly reported with overlapping multi-track hardening [15,16,24,25]. The heat input associated with a subsequent spot tempers the structure generated by the preceding spot, thus lowering the hardness. Consequently, the laser-hardened surface consists of regions of martensite and other regions of tempered martensite. Contrary to what was reported [15], no cracks were observed on the surface with overlapping laser spots. This implies better quality surfaces. The work shows that both an adjacent and overlapping spot hardening would result in only a fraction of the overall surface area being hardened. This results in regions of varying hardness. Zhang et al. [19] state that this is beneficial as the softer areas serve as lubricant reservoirs. These help in decreasing the coefficient of friction, hence improving the wear resistance. Jiang et al. [26] also suggest that one could use another processing strategy, one in which a second hardening pattern is applied to areas that were not hardened by the first hardening process. However, the latter could result in tempering the initial matrix of hardened spots. The reduced coverage and reduced energy input associated with surfaces hardened using adjacent spots rather than overlapping spots will lower the overall processing costs resulting from shorter processing time and lower use of consumables. The lower heat input would also lead to less heat distortion [18,19,27,28]. In fact, the surface roughness (Ra) generated by overlapping spots was higher by 0.70 µm, from that of adjacent spots which had a value of 0.31 µm (Fig. 5).
2.4. Material characterisation Surface analysis was carried out using a Nikon Optiphot-100 optical microscope and a Carl Zeiss Merlin scanning electron microscope. Micro-hardness tests were carried out using a Mitutoyo MVK-H12 testing machine, a load of 50 gf and a dwell time of 10 s. Mean surface roughness Ra were taken using a Mitutoyo Surftest501 profilometer. The roughness results are an average of at least three values. The roughness of the surfaces used for the rolling contact fatigue tests was measured before and after the wear tests using a Taylor Hobson TALYSURF CCI non-contact optical interferometer. This equipment was also used to obtain three-dimensional maps of the wear tracks formed and using multiple scans. The surface area, volume and profile of pits were obtained using Taylor Hobson Form Talysurf contact profilometer. All scans were integrated and analysed using MountainsMap® software. 3. Results and discussion 3.1. Single laser spot Single laser-hardened spots had a martensitic structure, that on occasions was supplemented by a ledeburitic structure surrounding the graphite nodules [14]. The nodules were unaltered as shown in Fig. 3. The surface hardness of the laser spots was approximately 770 HV and the hardened depth was circa 150 µm. The resulting martensitic structure was shown to induce compressive stresses in the hardened zone. The surface roughness, Ra of polished as-austempered ductile iron specimens increased from 0.15 ± 0.08 µm to 0.43 ± 0.14 µm as a result of laser hardening using one spot.
3.3. Scuffing resistance Lubricated sliding wear tests were carried out to determine the behaviour of discrete laser spot hardened ADI tested under starved lubrication conditions. As-austempered DI (A) and LS, LA, and LO laserhardened ADI specimens were investigated. Fig. 9 shows that during starved lubricating wear tests, surfaces hardened with separate spots (LS) and with adjacent spots (LA) survived a longer number of cycles
Fig. 3. SEM micrograph of the laser hardened surface (shown for comparative purposes) [14]. 102
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Fig. 4. Spot matrix laser hardening of ADI specimens showing (a) separated, (b) adjacent and (c) overlapping spots; unetched.
Fig. 5. Surface roughness values for different patterns of spot laser hardening.
Fig. 6. Microhardness profile across the surface of one single spot [14].
than the laser-hardened surfaces with overlapping spots (LO) and also the control as-austempered (A) specimens. Surfaces hardened with adjacent spots (LA) survived the longest number of cycles namely 1.5 × 105 cycles, whilst those having separated spots (LS) specimens survived 2.7 × 104 cycles. This is much higher than the 2.3 × 103 cycles survived by the as-austempered DI (A) specimens. The surface with adjacent spots contains the largest volume of the martensitic structure and only a small fraction of unhardened regions. This provides a higher wear resistance, and hence results in a longer number of cycles before failure. In addition, a large volume fraction of martensite implies that compressive stresses are present over a larger area, both at the surface and at sub-surface locations. Compressive stresses formed after LSH delay crack initiation and propagation [6]. Moreover, the coefficient of friction for LA specimens is the lowest
and about 0.08. This value remains constant throughout the sliding wear test (Fig. 10). In contrast, the friction coefficient for the LS specimens increases in steps as the test progresses. This can be attributed to the scuffing failure of the soft ausferritic structure located between the laser spots. As the test continues, the austenite in the ausferrite can transform to martensite, stopping any further increase in the coefficient of friction. In fact, the surface hardness measured after the wear test increased to around 560 HV, suggesting the change in phase from austenite to martensite. This phenomenon may occur a number of times depending on lubricant film failure and every time martensite is formed a low coefficient of friction results. The cold weld junctions formed during testing, between the martensitic pin surface and the hard disk break easily, resulting in lower adhesion between the two surfaces [29].
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Fig. 9. Sliding cycles before scuffing for as-austempered (A) and laser-hardened specimens (LS, LA, LO).
Fig. 7. Microhardness profile across the surface of three adjacent LSH spots.
Fig. 10. Friction coefficient for tested specimens.
regions between the spots, were ausferrite, which can strain harden and improve further the resistance to scuffing failure. Notwithstanding this, the rougher surfaces of the LO specimens were expected to create lubricant pockets, leading to a higher scuffing resistance than the asaustempered specimens. It appears however that the austenite to martensite transformation and absence of lubrication pockets in the asaustempered specimens is counterbalanced by the lower surface hardness of the tempered martensitic regions and the rougher surfaces of the LO specimens. Previously, it has been shown that laser hardening improves the abrasion wear resistance [13] and dry sliding wear resistance [7,12] of the upper ausferritic ADI by approximately 40% and 70% respectively. In the present study, it was shown that discrete spot laser hardening will improve the sliding wear resistance of ADI under starved lubrication conditions. Papaphilippou and Jeandin [20] reported a 17% increase in the wear resistance when spot laser hardening ductile iron. However, these researchers reported no significant difference in the wear rates of surfaces with different laser treatment patterns and different surface coverage percentages. It must be pointed out that tests were carried out under dry conditions and hence the beneficial effect of pockets of lubricants could not be taken advantage of [20]. Similar observations were reported by Xue et al. [18] who studied the influence of discrete laser spot hardening of different steels under unlubricated conditions. The samples that were irradiated with 100% coverage exhibited similar wear rates as those with 20–40% coverage.
Fig. 8. Typical microhardness profile across the surface of six overlapping LSH spots.
The martensite present in the laser-hardened regions provides better heat dissipation due to its good thermal conductivity leading to a reduction in the CoF [30,31]. The initial Ra value of LA specimens was circa 0.31 µm, which is higher than that of the LS specimens which has a roughness of approximately 0.15 µm. The dimples present in specimens with a higher roughness can serve as lubricant reservoirs, which help to decrease the friction coefficient and provide a higher scuffing resistance [19]. Extrapolating, then surfaces with overlapping laser spots which exhibited an even higher surface roughness at circa 0.70 µm, should have resulted in an even lower coefficient of friction and a high resistance to scuffing failure. However, as can be seen in Fig. 9, these specimens exhibited a high coefficient of friction and scuffing resistance comparable to that of as-austempered samples. It was shown in Fig. 8 that LO specimens had very few hard martensitic regions and a high volume fraction of tempered martensite with a hardness of approximately 450 HV. In comparison, in LS and the LA specimens, the softer 104
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Fig. 11. SEM micrographs showing cracking and pitting on typical worn surfaces.
3.3.1. Wear mechanism The worn surfaces showed signs of severe wear and adhesion between the two sliding bodies (Fig. 11). Scuffing is the result of microwelding of the surfaces due to the lack of lubrication, since it is not possible for the oil to separate the two sliding surfaces. Consequently, the asperities weld together due to the applied pressure and excessive heat generated by friction. As the two surfaces continue with sliding, the surfaces alternately weld and tear metal from each other's surfaces. This results in the formation of cracks and pits on the scuffed surfaces. Fig. 12 also shows peeling of the surface as well as surface delamination, depicted by the detachment of thin platelets from the surface. Additionally, evidence of shearing and plastic flow was noted in the sub-surface of the test specimens (Fig. 13). A distorted ausferritic microstructure underneath the martensitic microstructure is visible. The strain hardening of the microstructure discussed in Section 3.3 is caused by plastic deformation, resulting in a stronger material which causes surface flow and the microstructure to distort. Propagation of the cracks was influenced by graphite nodules, as shown in Fig. 14. The weak interface and low strength between the graphite and the metal matrix has a major influence on the fracture path and lives of the specimens.
Fig. 13. Cross-section of typical worn surface showing cracks and a distorted ausferritic microstructure.
counterparts. The L10 and L50 fatigue lives, which were obtained from the Weibull probability plot in Fig. 17, are respectively 4.3 × 105 and 9.2 × 105 cycles. The average fatigue life of the laser-hardened ADI specimens is 6.9 × 105 cycles greater than the fatigue life of the asaustempered DI specimens. Xu and Lu [9] conducted contact fatigue testing of ADI specimens austempered at 370 °C and loaded at a stress of 2.3 GPa. They showed that the improvement brought about as a result of laser hardening is only 0.5 × 105 cycles when carrying out. The discrepancy in the two studies may be attributed to a number of factors. For example, Xu and Lu [9] do not report the surface finish of either the as-austempered or the laser-hardened specimens tested. Also, they did not observe the hard bull's eye ledeburitic structure around the graphite nodules that was noted in the present study. The high hardness of the ledeburitic structure may be one of the reasons for the better results reported in the present study. The polished as-austempered specimens had a surface roughness (Ra) of 0.05 µm, which increased to 0.12 µm after laser hardening. With these roughness values, the specific film thickness λ was calculated to be 1.12. A value of λ above 1 denotes that rolling is taking place in the full lubricated regime. However, considering this value is very close to 1, it is then likely that at some point/s during the test this value changes and reaches a value that is lower than 1. In reality, this implies that the condition of point contact assumed when calculating λ was true. Under conditions of general surface contact or occasional contact of asperities, the value of λ is lower than 1 and mixed lubrication occurs. This would imply that the laser-hardened specimens were being tested under full/ mixed lubrication conditions. The specific film thickness for the smoother as-austempered specimens was calculated to have a value of 2.81, meaning that these specimens were tested under full lubrication, which should give the
3.4. Rolling contact fatigue resistance Rolling contact fatigue (RCF) tests were carried out on as-austempered ADI specimens and on laser-hardened ADI specimens that were irradiated using adjacent laser spots and which exhibited the highest resistance to scuffing. The number of cycles to failure was recorded when pitting occurred on the specimens. Fig. 15 shows a typical pit formed after testing a laser hardened specimen, while Fig. 16 shows numerous cracks present on a typical wear track. Fig. 17 shows that the laser-hardened ADI specimens exhibited the better rolling contact fatigue performance than their as-asutempered
Fig. 12. SEM micrograph showing delamination of typical worn surfaces. 105
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Fig. 14. Influence of graphite nodules on the crack propagation in typical worn surfaces.
specimens a longer life. However, in this work the laser-treated specimens performed better, which can be attributed to the higher surface hardness (~ 770 HV) compared to that of the as-austempered specimens (~ 370 HV). The compressive stresses present in the martensitic regions of the laser-hardened specimens [14] serve to reduce the maximum shear stress when the specimens are in loaded contact. This delays crack nucleation and propagation and as a result increases the fatigue life. The RCF performance of the laser-hardened ADI is superior to that reported in some other studies for carburised steel. For example, Oila and Bull [32], Widmark and Melander [33] and Cavallaro et al. [34] obtained RCF lives respectively of 0.5 × 105, 3 × 105 and 2 × 105 cycles when testing at an applied stress of approximately 2.3 GPa. These values are lower than the average RCF life of 6.9 × 105 cycles obtained for the laser-hardened ADI in the present study. In another study on carburised steel, Ramanathan and Radhakrishnan [35] report an RCF life of 20 × 105 cycles when testing at a stress of 2.5 GPa. These values are comparable to those in this study. Fatigue values reported by other researchers [36,37] have fatigue lives of approximately 500 × 105 cycles. This is much higher than that reported in the present study even if the applied stress is only 1.7 GPa. Same applies to results by Hongbin et al. [38] who reported an RCF value 50 × 105 cycles when testing at an applied stress of 3.05 GPa. The above results show that the RCF life of carburised steel varies widely and depends on the parameters of the carburising treatment and the surface integrity of the specimens following carburising. It has been shown that laser hardening may be a suitable treatment for improving the rolling contact fatigue life of ADI. Laser-treated ADI surfaces have RCF values comparable to that reported by a number of researchers studying carburised steel, a material known to be suitable for engineering components, such as gears.
Fig. 16. Cracks present on a wear track after rolling contact fatigue testing a laser- hardened specimen.
4. Conclusions This study was carried out in order to determine the influence of discrete laser spot hardening on the scuffing wear resistance and rolling contact fatigue resistance of a Cu-Ni austempered ductile iron (ADI). The main conclusions arising from this study are the following: 1. Discrete laser spot hardening is a suitable technique for improving the scuffing wear and the rolling contact fatigue resistance of upper ausferritic austempered ductile iron. 2. Lubricated sliding wear tests under starved lubrication conditions revealed that the laser-hardened ADI specimens treated with adjacent laser spots performed better than the as-austempered and laser-hardened samples treated with other spot patterns. The higher
Fig. 15. SEM and 3D profilometer images of the wear tracks following contact fatigue testing of a laser- hardened specimen. 106
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[6] J.C. Ion, Laser Processing of Engineering Materials, Elsevier ButterworthHeinemann, 2005. [7] G.-X. Lu, H. Zhang, Sliding wear characteristics of austempered ductile iron with and without laser hardening, Wear 138 (1990) 1–12. [8] C. Soriano, J. Leunda, J. Lambarri, V.G. Navas, C. Sanz, Effect of laser surface hardening on the microstructure, hardness and residual stresses of austempered ductile iron grades, Appl. Surf. Sci. 257 (2011) 7101–7106. [9] D.-X. Xu, G.-X. Lu, Contact fatigue behaviour of an austenitic-bainitic ductile iron, Wear 169 (1993) 153–159. [10] S.K. Putatunda, L. Bartosiewicz, R.J. Hull, M. Lander, Laser hardening of austempered ductile cast iron (ADI), Mater. Manuf. Process. 12 (1997) 137–151. [11] A. Roy, I. Manna, Mathematical modeling of localized melting around graphite nodules during laser surface hardening of austempered ductile iron, Opt. Lasers Eng. 34 (2000) 369–383. [12] A. Roy, I. Manna, Laser surface engineering to improve wear resistance of austempered ductile iron, Mater. Sci. Eng. vol. A297, (2001) 85–93. [13] Y.-H. Tan, S.-I. Yu, J.-L. Doong, J.-R. Hwang, Abrasive wear property of bainitic nodular cast iron in laser processing, J. Mater. Sci. 25 (1990) 4133–4139. [14] A. Zammit, S. Abela, J.C. Betts, M. Grech, Discrete laser spot hardening of austempered ductile iron, Surf. Coat. Technol. 331 (2017) 143–152. [15] ASM International, Friction, Lubrication and Wear Technology, vol. 18, (1992). [16] L. Giorleo, B. Previtali, Q. Semeraro, Modelling of back tempering in laser hardening, Int. J. Adv. Manuf. Technol. 54 (2011) 969–977. [17] D.I. Pantelis, G. Pantazopoulos, S.S. Antoniou, Wear behavior of anti-galling surface textured gray cast iron using pulsed-CO2 laser treatment, Wear 205 (1997) 178–185. [18] L. Xue, M.U. Islam, G. McGregor, Dot matrix hardening of steels using a fiber optic coupled pulsed Nd: YAG laser, Mater. Manuf. Process. 14 (1999) 53–65 (1999). [19] X.M. Zhang, H.C. Man, H.D. Li, Wear and friction properties of laser surface hardened EN31 steel, J. Mater. Process. Technol. 69 (1997) 162–166. [20] C. Papaphilippou, M. Jeandin, Spot laser hardening, J. Mater. Sci. Lett. 15 (1996) 1064–1066. [21] L. Tricarico, A. Ancona, G. Palumbo, D. Sorgente, R. Spina, P.M. Lugara, Discrete spot laser hardening and remelting with a high-brilliance source for surface structuring of a hypereutectoid steel, Mater. Des. 115 (2017) 194–202. [22] Standard Test Method for Evaluating the Scuffing Load Capacity of Oils (FZG Visual Method), ASTM D 5182 - 97, Reapproved, 2002. [23] R. Michalczewski, W. Piekoszewski, M. Szczerek, J. Wulczyński, A method for the assessment of the rolling contact fatigue of modern engineering materials in lubricated contact, Trans. Famena 36 (2012) 39–48. [24] R.S. Lakhkar, Y.C. Shin, M.J.M. Krane, Predictive modeling of multi-track laser hardening of AISI 4140 steel, Mater. Sci. Eng.: A 480 (2008) 209–217. [25] G. Tani, L. Orazi, A. Fortunato, Prediction of hypo eutectoid steel softening due to tempering phenomena in laser surface hardening, CIRP Ann. - Manuf. Technol. 57 (2008) 209–212. [26] J. Jiang, L. Xue, S. Wang, Discrete laser spot transformation hardening of AISI O1 tool steel using pulsed Nd:YAG laser, Surf. Coat. Technol. 205 (2011) 5156–5164. [27] P.R. Woodard, J. Dryden, Thermal analysis of a laser pulse for discrete spot surface transformation hardening, J. Appl. Phys. 85 (1999) 2488–2496. [28] K.A. Khashan, O.A. Hamadi, Features of spot-matrix surface hardening of lowcarbon steel using pulsed laser, J. Eng. Technol. 25 (2007) 168–175. [29] M.O.A. Mokhtar, M. Zaki, G.S.A. Shawki, Effect of mechanical properties on frictional behaviour of metals, Tribol. Int. (1979) 165–267. [30] A. Gural, Influence of martensite particle size on dry sliding wear behaviour of low carbon dual phase powder metallurgy steel, Met. Mater. 48 (2010) 25–31. [31] S. Sendooran, P. Raja, Metallurgical investigation on cryogenic treated HSS tool, Int. J. Eng. Sci. Technol. 3 (2011) 3992–3996. [32] A. Oila, S.J. Bull, Assessment of the factors influencing micropitting in rolling/ sliding contacts, Wear 258 (2005) 1510–1524. [33] M. Widmark, A. Melander, Effect of material, heat treatment, grinding and shot peening on contact fatigue life of carburised steels, Int. J. Fatigue 21 (1999) 309–327. [34] G.P. Cavallaro, T.P. Wilks, C. Subramanian, K.N. Strafford, P. French, J.E. Allison, Bending fatigue and contact fatigue characteristics of carburized gears, Surf. Coat. Technol. 71 (1995) 182–192. [35] S. Ramanathan, V.M. Radhakrishnan, Investigation of rolling contact fatigue damage of a case-carburized low alloy steel, Wear 45 (1977) 323–333. [36] D.P. Townsend, E.V. Zaretsky, Effect of shot peening on surface fatigue life of carburized and hardened AISI 9310 spur gears, NASA Tech. Pap. 2047 (1982) 5–12. [37] J. Champaigne, Shot peening overview, Metal Improvement Company, 2001. [38] X. Hongbin, C. Qing, S. Eryu, W. Dengzhen, C. Zhaohong, W. Zhengle, The effect of shot peening on rolling contact fatigue behaviour and its crack initiation and propagation in carburized steel, Wear 151 (1991).
Fig. 17. Weibull probability plot for data obtained from rolling contact fatigue tests.
life provided by the former laser-hardened samples is associated with a load carrying capacity, as a result of the higher volume fraction of martensite on the surface, and also the residual compressive stresses induced in the laser-hardened surface. 3. The process of laser hardening with overlapping spot matrices caused back-tempering leading to regions of martensite and of tempered martensite. 4. ADI specimens treated with laser-hardened spots separated from each other survived 2.7 × 104 cycles before scuffing, which is also an improvement over the performance exhibited by their counterpart as-austempered specimens. In comparison, the surfaces irradiated with overlapping laser-hardened spots displayed the poorest performance, with their endurance limit roughly equal to that of asaustempered specimens. This was attributed to the presence of the softer tempered martensitic regions. 5. It was also shown that laser hardening consisting of adjacent spots increases the rolling contact fatigue life by 6.9 × 105 cycles over that of as-austempered counterparts. This improvement was obtained as a result of the harder laser-treated surfaces and the residual compressive stresses, which outweighed the negative effect of the rougher surface. Acknowledgements The authors would like to acknowledge the ERDF funding for the purchase of the testing equipment through the project: "Developing an Interdisciplinary Material Testing and Rapid Prototyping R&D Facility (Ref. no. 012)". References [1] R.A. Harding, The production, properties and automotive applications of austempered ductile iron, Kov. Mater. 45 (2007) 1–16. [2] AGMA 939-A07, Austempered Ductile Iron for Gears, ed: AGMA, 2007, pp. 1–10. [3] H. Zhang, Y. Shi, C.Y. Xu, M. Kutsuna, Surface hardening of gears by laser beam processing, Surf. Eng. 19 (2003) 134–136. [4] W.M. Steen, J. Powell, Laser surface treatment, Mater. Eng. 2 (1981) 157–162. [5] W.M. Steen, Laser Material Processing, Springer, London, 2003.
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