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Tribological behavior of gray cast iron textured by maskless electrochemical texturing
Wear 376-377 (2017) 1601–1610
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Tribological behavior of gray cast iron textured by maskless electrochemical texturing L.R.R. da Silva a, H.L. Costa a,b,n a b
Universidade Federal de Uberlândia, Laboratório de Tribologia e Materiais, Av. João Naves de Ávila, 2121, Bloco 1M, 38408-100 Uberlândia, MG, Brazil Universidade Federal do Rio Grande, Escola de Engenharia, Av. Itália, km. 8, Rio Grande, RS, Brazil
art ic l e i nf o
a b s t r a c t
Article history: Received 30 August 2016 Received in revised form 4 January 2017 Accepted 9 January 2017
Gray cast iron is widely used in diesel automotive engines due to its combination of low cost and good fusibility, but also to the presence of graphite flakes in their microstructure. Besides their role as mechanical dampers, graphite flakes can act as solid lubricant, reducing friction and wear of moving components. Surface texturing has been intensively investigated in recent years to improve lubrication and reduce friction and wear in tribological applications, but little attention has been given to surface texturing of gray cast irons. On the other hand, surface texturing adds costs, which can supplant the benefits achieved, in particular when high cost surface texturing techniques are employed. Maskless electrochemical texturing (MECT) is a promising technique, which combines simplicity, low cost and high speed. Recently, MECT was adapted to texture gray cast irons. This work investigates the tribological performance of gray cast iron textured by MECT using lubricated block-on-ring tests under two different rotational speeds and a constant normal load. The blocks consisted of textured cast iron and the rings of SAE 4620 steel. For comparative purposes, untextured polished gray cast iron blocks, as well as gray cast iron blocks subjected to electrochemical polishing (employing the same electrolyte used during MECT) were tested. The results showed a large reduction in friction (up to 2.5x) and wear (higher than 5x) when comparing the polished and the textured samples. Comparing the textured samples and the electropolished samples, this improvement in performance reduced, showing that the benefits were partly due to the texture pattern itself (arrays of pockets) and partly to the exposure of the graphite flakes during electrochemical dissolution of the metal matrix. As the wear progressed and the contact between the block and the ring became more conformal, the benefits obtained by MECT (and electro polishing) were reduced. & 2017 Elsevier B.V. All rights reserved.
Keywords: Gray cast iron Surface texturing Maskless electrochemical texturing Graphite
1. Introduction In vehicles that use internal combustion engines, friction of internal components account for about a third of the energy loss [1,2]. Some alternatives to reduce friction between these components are the use of materials with self-lubricating properties such as gray cast iron [3], of low friction coatings and the application of surface texturing, as well as synergistic responses between them [4,5]. Gray cast iron liners are of particular use in diesel engines [6–8]. The effect of surface texturing under lubricated conditions is strongly dependent on the lubrication regime and on the amount of lubricant supply available. Under full film lubrication conditions, the lubricant pockets may act as microbearings, increasing the n Corresponding author at: Universidade Federal do Rio Grande, Escola de Engenharia, Av. Itália, km. 8, Rio Grande, RS, Brazil. E-mail address: [email protected] (H.L. Costa).
http://dx.doi.org/10.1016/j.wear.2017.01.028 0043-1648/& 2017 Elsevier B.V. All rights reserved.
hydrodynamic pressure, and therefore helping to prevent asperity contact [9–11]. However, due to the reduction in the contact area, which increases contact pressure [9], and also to the occurrence of cavitation [12], surface texturing may lead to negligible [13] or even deleterious effects [12,14]. Under mixed lubrication conditions, the reported effects seem more significant [15–18]. The effect of the pockets as microbearings increases film thickness, reducing the speed or load range for which the lubricant does not separate the surfaces completely, i.e., shifts the Stribeck curves, expanding the ranges for which full film lubrication occurs [13]. Besides, surface texturing increases the time before scuffing occurs when the supply of lubricant is limited [19,20], which is particularly relevant under starved lubrication conditions [21]. Under boundary lubrication, hydrostatic reaction forces in the pockets may help to increase film thickness [22,23], besides the additional lubricant supply and the entrapment of wear debris [24], but the high stresses induced around the pockets [25] may lead to detrimental effects [7].
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The literature is abundant in works regarding the effect of the texture pattern geometry on the tribological performance of textured surfaces under lubricated conditions. They include works regarding the effect of the width of the motifs [11,21,26], the ratio between the height and width of the motifs [9,16], the percentage of area coverage [9,10,12,16,26] and the shape of the motifs [9,18,26], particularly circular dimples, grooves and interrupted grooves. Costa and Hutchings [9] showed a superior performance of chevron-like motifs to increase load capacity under hydrodynamic lubrication when compared with arrays of circular pockets and grooves. Later, Morris et al. [17] used numerical simulations to optimize the geometry of chevron arrays aiming to reduce friction coefficient and to increase load capacity in pistonring systems. The use in industrial scale of surface texturing for reduction of frictional losses depends on the existence of techniques capable of texturing surfaces rapidly and inexpensively in order to make the method cost-effective. Although laser texturing is the technique most widely used industrially [27], it may present some drawbacks, such as the high cost of the equipment involved in the process, large processing time when texturing large areas [21] and the presence of heat affected zones and recast areas, which can act as abrasives if not removed by subsequent process [13]. This last effect can be largely reduced with the use of femtosecond lasers [4,28], but this adds extra cost to the equipment. Among alternative surface texturing techniques, Maskless Electrochemical Texturing method (MECT) was developed as a simple, low cost, high speed texturing technique, particularly suited to metals [29,30]. Initially, it was developed to texture plane surfaces, mainly for carbon steel [29,30] and copper [31]. Very recently, MECT was adapted to texture diesel cylinder liners [32], producing arrays of chevrons, which faced two challenges. The first was the cylindrical geometry and the second was texturing a material containing phases with very different chemical characteristics: the metallic matrix and the graphite flakes. Despite the difficulty to texture gray cast iron, texture patterns with good resolution were produced cheaply and quickly. The results showed that anodic dissolution mainly occurred in the metallic matrix, but not in the graphite. Graphite was only removed after it lost the mechanical support provided by the metallic matrix. This resulted in exposure of the graphite flakes, in addition to the formation of the texture pattern. Therefore, there is potential for an extra beneficial effect when using MECT to texture grey cast iron, since the graphite flakes may act as solid lubricant, in addition to the benefits obtained by the presence of the texture pattern. This paper aims to investigate the tribological behavior of gray cast iron textured by MECT, since this material is widely used in lubricated sliding conditions. For that, lubricated sliding tests of plane textured surfaces were carried out using a block-on-ring tribometer. In addition, mechanically polished surfaces, as well as surfaces obtained by conventional electrochemical polishing were also tested, since this method is used to highlight the graphite flakes in gray cast iron [33]. This comparison helped to separate the contribution originated from the graphite flakes from that from the texture pattern itself. The results evidenced a strong reduction in friction and wear when MECT is used to texture gray cast iron.
spacing of 0.33 70.11 mm. Vickers hardness tests using normal load of 612.5 N showed and average value of 1.810 70.025 GPa. The dimensions of the blocks followed recommendations from ASTM D271-94(2014) standard [34], which were 6.35 mm (width), 15.76 mm(length) and 10 mm (height). All blocks were ground and polished, leading to a 3D RMS surface roughness (Sq) of 0.09870.004 μm. The rings were provided by Falexs according to ASTM D271-94 standard [34]. They were manufactured using SAE 4620 carburized steel, with a Vickers hardness ranging from 7.379 to 7.595 GPa. The width was 8.5 mm and the external diameter was 35 mm. The surface roughness of the rings was assessed measuring three different areas of 1 mm2 in each ring, using a UBM Microfocus Expert IVs laser interferometer, and only rings with Sq between 0.19 and 0.21 μm were used in the tests. Considering the hardness values for the block (1.810 GPa) and the rings (7.487 GPa), wear was mainly expected to occur in the blocks and the ring should remain intact [35]. The lubricant used was a SAE 40 mineral oil with a viscosity (ηo) ranging from 159.9 cSt at 40 °C to 15.2 cSt at 100 °C, and a viscosity pressure coefficient (α) of 2.93 GPa 1. 2.2. Surface texturing The gray cast iron blocks were textured by Maskless Surface Texturing method (MECT). The basic apparatus used is described in Fig. 1 [29], where (1) is a constant voltage source, (2) is the electrolyte reservoir, (3) is the sample to be textured, (4) is the signal modulator, (5) is the electrolyte pump, (6) is the representation of the connections between the electrolyte pump and reservoir and (7) is the tool containing the pattern to be transferred to the sample. The tool cover was made with AISI 304 austenitic stainless steel. The pattern to be transferred to the sample was cut into this cover by laser micro drilling. One of the faces of the tool cover received electrostatic painting after drilling, as shown in Fig. 2, which was responsible for the electrical insulation between the tool and the workpiece. The texture pattern consisted of an array of chevrons, based on the optimization results presented in [17]. During MECT, the electrolyte was a 200 g L 1 NaCl solution in distilled water, as suggested in [9,12], the voltage was 7.5 V, the texturing time was 60 s and the gap between tool and specimen was 100 μm [12]. For comparison, mechanically polished gray cast iron blocks were also tested. In addition, a third group was electrochemically polished after the mechanical polishing. Electrochemical polishing
2. Experimental methods and materials 2.1. Samples and lubricant The blocks used in the sliding tests were manufactured using gray cast iron presenting a pearlitic matrix with an interlayer
Fig. 1. MECT Apparatus used for surface texturing the blocks; from [30], with permission.
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with Nital (Fig. 3.b). This figure is presented only to evidence the material microstructure, but the polished samples used in the sliding tests were not etched. The electrochemically polished surface is illustrated in Fig. 3.c, where the pearlitic matrix was revealed by the dissolution during electrochemical polishing, as well as the graphite flakes, which are in relief in relation to the pearlitic matrix. This is a result of the preferential anodic dissolution of the metallic matrix by the electropolishing process, exposing the graphite flakes [33]. A textured sample is shown in Fig. 3.d, which shows the chevron patterns and the material microstructure, evidenced during electrochemical dissolution that also occur in the areas outside the chevrons (although at a much lower rate). Graphite flakes are evidenced both inside and outside the chevrons. Fig. 3.e presents a magnified view of the areas outside (but near) the chevrons, where the pearlitic nature of the matrix is again evidenced and the graphite flakes appear elevated in relation to the matrix. Therefore, MECT may come to possess synergistically the advantages of the surface texturing (pockets) as well as a secondary effect (exposure of graphite flakes) conferred by electrochemical polishing, suggesting that MECT could result in further gains compared with other texturing methods such as laser texturing for gray cast iron. An example of the 3D surface topography generated by MECT of gray cast iron is illustrated in Fig. 4. Fig. 2. Tool used to surface texture the gray cast iron blocks by MECT.
was carried out at a voltage of 7.5 V, during 60 s using a 200 g L 1NaCl solution as the electrolyte. The mechanically and electrochemically polished blocks had an initial surface roughness of 0.098 70.004 and 0.201 70.032 mm, respectively. The textured blocks had a surface roughness of 0.171 70.027 mm in the areas outside the imposed surface patterns. Micrographs for the three surface conditions used in this work are presented in Fig. 3. Fig. 3.a shows the mechanically polished surface, where only the graphite flakes can be observed. The pearlitic matrix of the polished sample is evidenced by etching
2.3. Block on ring tests The tribological testes were carried in a Falexs block on ring tribometer (Fig. 5), where (1) is the load cell, (2) is the lever arm used to apply the normal load at a ratio of 1:10, (3) is the thermocouple, (4) is the test ring, (5) is the test block and (6) is the hermetically closed test chamber. For the three groups of samples (mechanical polishing, electrochemical polishing and MECT) tests were carried out using a normal load of 343 N and ring rotational speeds of 200 and 600 rpm. Each test had a duration of 20 min. The lubricant temperatures during the tests were measured as approximately 90
Fig. 3. Optical micrographs for the surface conditions used in this work: (a) mechanically polished surface; (b) mechanically polished surface etched with NITAL; (c) electrochemically polished surface; (d) chevrons produced by MECT; (e) amplified view of the surface near the chevrons produced by MECT.
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Table 1 λ parameter values for each condition tested.
Fig. 4. 3D surface map of the textured samples; laser interferometry.
and 75 °C for the 200 and 600 rpm tests, respectively. Each test was replicated three times to ensure reproducibility. The initial mean contact pressure (Pm) and the semi contact width (a) can be calculated using hertz equation for a line contact, as shown in Eq. (1) [36]. 1
Pm=
⎛ wR* ⎞ 2 π E*w ⎟ ; a = 2⎜ 4 R*Lπ ⎝ LE* ⎠
(1)
where E* is the contact young modulus, R* is the contact radius of curvature, L is the contact length and w is the normal load. Using this equation a mean contact pressure of 0.74 GPa between the block and the ring and a semi contact width of 280 mm were found to operate at the beginning of the tests. The initial fluid film thickness (H*) was calculated using the Dowson-Hamrock equation [37–39], as shown in Eq. 2.
H *=1. 6G 0.6U 0.7W −0.13
G=
αE *
(3)
U=
Vη0 E*R*
(4)
W=
w E*A
Rotational speed (rpm)
H* (μm) S*q (μm)
Mechanically Polished Electrochemically Polished Textured Mechanically Polished Electrochemically Polished Textured
200 200
0.324 0.324
0.2337 0.004 1.455 0.284 7 0.032 1.143
200 600 600
0.324 0.699 0.699
0.263 7 0.027 1.231 0.2337 0.004 3.138 0.284 7 0.032 2.465
600
0.699
0.263 7 0.027 2.656
where α is the lubricant pressure viscosity coefficient (GPa 1), V is the speed of the ring in relation to the block (m s 1), ηo is the lubricant viscosity (Pa s), w is the normal load and A is the contact area. For 200 and 600 rpm, the initial film thicknesses (H*) were calculated as 0.223 and 0.311 mm, respectively. Using those values and the 3D RMS surface roughness for each condition, it was possible to calculate the lambda parameters (λ) for each condition (Table 1), defined in the Eq. (6), where S*q is the combined roughness of the block (SqB) and ring (SqR), given by the Eq. (7) [40].
H* Sq*
λ=
Sq*=
(6) SqB2 + SqR2
(2)
where G, U and W are respectively, the non-dimensional material, speed and load parameters [37,38], as shown in Eqs. (3)–(5).
λ parameter
Surface condition
(7)
The majority of the conditions tested presented λ values between 1 and 3, placing the initial test conditions in the mixed EHL lubrication regime [41,42]. One exception is the mechanically polished specimen at 600 rpm, which showed a value slightly above 3. However, due to the wear caused by the initial sliding, it is expected a rapid increase in the surface roughness, so that the λ value should decrease quickly to values below 3. On the other hand, it is expected that wear should also change the contact geometry, making the contact more conformal with increased sliding. This will be discussed in the Results section.
(5)
s
Fig. 5. Block on ring Falex tribometer.
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Fig. 6. Friction coefficient versus time: (a) 200 rpm; (b) 600 rpm.
3. Results and discussion Fig. 6 shows the friction coefficient values measured for the tests at 200 rpm (a) and 600 rpm (b). Only one curve is presented for each condition, but the different repetitions carried out showed very reproducible curves. The evolution of friction coefficient with time showed the typical behaviour commonly reported in the literature for block-on-ring tests [35,43], where an initial peak is followed by a sharp reduction in friction. This initial running-in period is observed for all surface conditions and speeds. The running-in period was longer for the polished specimens than for electropolishing and MECT conditions. This is probably due to graphite flakes, which were exposed by the anodic dissolution of the pearlitic matrix around them. The solid lubricant properties of the graphite flakes exposed during electropolishing [33] and MECT [32] probably contribute to reduce friction in the initial stages of the test, where the sliding speed is lower and therefore lubricant film thickness is thinner [42], acting in the boundary regime. The occurrence of boundary lubrication in the initial stages of block-on-ring tests has been elegantly demonstrated in [35]. After the short initial transient, a drastic reduction in friction was observed for the textured blocks when compared with the polished blocks, whereas the electropolished blocks presented an intermediate behavior. Therefore, two effects must be contributing to improve the performance of the textured blocks. The first is related to the chevron-like pockets. Under mixed lubrication conditions, surface texturing can been very successful, where the contribution of the pockets on hydrodynamic pressure distribution
and the additional lubricant supply seem to contribute to reduce friction [15,17,44], including for cast irons [45]. The second are the exposed graphite flakes, which acts as a solid lubricant in the contact. This second effect also occurs for the electropolished blocks, justifying their intermediate behavior. The positive effect of graphite during sliding of cast irons [33], including for block-onring tests [46], has been largely reported in the literature. Another aspect that is common to all the curves presented is that after the running-in period, friction increases with test time and then decreases. This friction increase is particularly accentuated for the polished block tested at 600 rpm. In order to investigate further this behaviour, some tests were interrupted after only 250 s of test, in order to identify possible phenomena within this period where friction increases. The wear scars produced in the interrupted tests were measured by 3D laser interferometry. Despite the longer sliding distances that were experienced under higher rotational speeds, the volume loss after 250 s of sliding was much larger for 200 rpm (0.03 mm3) than for 600 rpm (0.01 mm3). This occurs because at higher speeds the lubricant film thickness increases [42], reducing asperity contact and then wear. Another point is that the measured temperature was lower for tests at 600 rpm (75 °C) than at 200 rpm (90 °C), probably due to a more efficient convection. At lower temperatures, lubricant viscosity increases, increasing film thickness [42]. Since the rings were much harder than the blocks, wear was mostly restricted to the blocks and wear scars could not be identified in the rings. Therefore, wear progressed in terms of changing contact geometry from an initial line contact to a more conformal contact, which occurred faster for the tests at 200 rpm. The
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Fig. 7. 3D surface topography of the worn scars in the mechanically polished blocks after 250 s of rubbing: (a) 200 rpm; (b) 600 rpm.
occurrence of preferential wear of a softer block and therefore the transition from a line non-conformal contact to a more conformal contact was assessed and discussed in [35] using a displacement sensor. As an example, Fig. 7 presents the wear scars produced after those interrupted tests for the polished blocks. The length of the wear scars (around 5 mm) were smaller than the ring width (8.5 mm). Although we cannot precise the reason for this behaviour, other works in the literature [35] have also showed wear tracks that are smaller than the ring width. One possible explanation is a more efficient lubricant entrainment near the scar extremes, making wear negligible in these areas. In [35], low magnification stereo micrographs near the contact edges indeed suggest better lubricant entrainment near the width extremes. With the software MontainsMap s, the average radius of the wear scars were computed using 10 equally-spaced radius measurements (30 measurements for the 3 repetitions). For the test at 200 rpm, the radius of the scar in the block was 18.1 70.22 mm, which is very close to the ring radius (17.7 mm), whereas for 600 rpm it was 22.4 70.97 mm, i.e., the worn block was less conformal to the ring. Contact pressure is substantially reduced in conformal contacts when compared with non-conformal contacts [47]. So, it seems that as the wear of the blocks proceeds, contact pressure reduces, increasing film thickness. Therefore, after a certain level of conformity, film thickness may increase so that asperity contact reduces significantly and therefore friction reduces with further sliding. Further analysis of Fig. 6 shows that for the tests at 200 rpm, after around 500 s of sliding, the friction behavior of the textured and electropolished blocks became undistinguishable, although friction remained lower than for the mechanically polished block. It is plausible that after 500 s, the wear of the blocks was sufficient to completely remove the chevron-like pockets, so that the only beneficial effect that still remained was that of the exposed graphite flakes. Indeed, at the end of tests at 200 rpm, the chevrons were nearly completely removed inside the wear scars, as shown
in the images presented in Fig. 8.c. Fig. 8 consists of scanning electron microscopy (SEM) images of the wear scars using back scattered electrons (BSE). 3D surface topography of the wear scars showed small traces of reminiscent chevrons (Fig. 9), but they were probably too shallow to have any effect on lubricant entrainment or hydrodynamic pressure. After around 1000 s of test, all blocks showed similar low friction values of around 0.09. Since the exposed graphites are no longer visible within the wear scars at the end of the tests for the textured and electropolished blocks (Fig. 8.b and c), their friction behaviour becomes very similar to that presented by the mechanically polished block. For the tests at 600 rpm, friction coefficients of the textured blocks were the lowest and the electropolished blocks again showed an intermediate behaviour. After around 1000 s of test, the three blocks presented the same friction values. As shown in Fig. 10, chevron-like pockets and exposed graphite flakes still remained inside the wear scars at the end of the tests. With the 3D interferometric maps of the wear scars (Fig. 10), the average scar radii were measured at the end of the tests and all values were very similar to the ring curvature. With high conformity between ring and wear scar, contact pressures were quite low at the end of the tests. On the other hand, the lubricant film should be thicker at 600 rpm than at 200 rpm. Therefore, it is postulated that for the tests at 600 rpm, the lubrication regime shifted to nearly full film lubrication at the end of the tests, with negligible asperity contact. This hypothesis is corroborated by the low friction values in this region for the three blocks (around 0.06). Various authors have proposed that under full film lubrication, the effects of surface texturing can be very small. Costa and Hutchings [9] did not find differences in the friction coefficients presented by smooth and textured surfaces, despite a small increase in load carrying capacity for the textured surfaces. In very recent works [14,48], the simultaneous measurement of friction coefficient and film thickness (by ultrathin laser interferometry) in tests under full film lubrication showed that although surface texturing boosts lubricant entrainment, the effects on friction are
Fig. 8. BSE SEM images of the wear tracks on the blocks after the tests at 200 rpm: (a) mechanical polishing; (b) electrochemical polishing; (c) MECT.
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Fig. 9. 3D surface topography of the wear scars on the blocks after the tests at 200 rpm: (a) mechanical polishing; (b) electrochemical polishing; (c) MECT.
Fig. 10. BSE SEM images of the wear tracks on the blocks after the tests at 600 rpm: (a) mechanical polishing; (b) electrochemical polishing; (c) MECT.
Fig. 11. 3D surface topography of the wear scars on the blocks after the tests at 600 rpm: (a) mechanical polishing; (b) electrochemical polishing; (c) MECT.
negligible or even slightly detrimental. One point to be noted is that the total width of each chevronlike motif that composes the texture pattern (around 2.3 mm) is
substantially wider the initial elastic contact width (560 mm). The literature [7,49] mainly suggests that the size of the individual motifs should be smaller than the contact width, in particular
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Fig. 12. Volume wear rates for the blocks.
Fig. 13. BSE SEM images showing detail of the worn surfaces after the tests: (a) mechanical polishing at 200 rpm; (b) electrochemical polishing at 200 rpm; (c) MECT at 200 rpm; (d) mechanical polishing at 600 rpm; (e) electrochemical polishing at 600 rpm; (f) MECT at 600 rpm.
under full film conditions, due to the fact that the lubricant may flow away from the contact. However, in this work the lubrication regime is mostly mixed, so that wear progresses quickly, increasing the contact width substantially. As seen in Figs. 9 and 11, by the end of the tests the wear craters are wider than the individual motifs. Another point is that the angle of the chevrons ensures that even while the contact width is narrower there is substantial lubricant entrapment within the contact. The motifs used in this work were the smallest chevrons that could be produced, but this can be reduced if finer laser drilling is used to produce the tools that are used in MECT. Another feature that clearly emerges from the 3D maps of the wear scars on the blocks is that wear decreases when the rotational speed increases from 200 rpm to 600 rpm. Wear rates (k) were calculated using the Archard equation [50], as shown in the Eq. (8), where d is the sliding distances and V is the volume loss, calculated using the software MontainsMaps.
k=
V wd
(8)
The global wear rates, measured at the end of the tests, are presented in Fig. 12, which shows an approximate reduction of around 82% of the wear rates between the tests at 200 rpm and 600 rpm, which occurs due to thicker hydrodynamic films that are formed under higher sliding speeds. More importantly, this graph shows a reduction of around 81% in the wear rates found for the textured block when compared with the polished block both at 200 rpm and 600 rpm. For comparison, the reduction when electropolishing was used instead of MECT was of 72% at 200 rpm and of 65% at 600 rpm. Therefore, the wear reduction of the textured blocks is not only due to the exposure of the graphite flakes but also due to presence of the chevron-like pockets. Fig. 13 shows BSE SE images of the wear scars produced after the tests at 200 and 600 rpm using higher magnification. For the mechanically polished block (Fig. 13.a and d), large dark
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agglomerates are seen and EDS analysis confirmed that they are very rich in carbon (almost 90%C). For the electrochemically polished (Fig. 13.b and e) and MECT blocks (Fig. 13.c and f), instead of isolated carbon agglomerates, very fine and evenly distributed carbon-rich regions exist forming a fairly continuous carbon-rich tribolayer, which is probably responsible for the low friction coefficients and wear rates found in the experiments.
4. Conclusions This work investigated the tribological performance of gray cast iron surfaces textured by Maskless Electrochemical Texturing using block-on-ring tests. The results showed that: 1. Surface texturing by MECT reduces drastically friction and wear, which was clearly attributed to two effects. The first was the exposure of the graphite flakes by the preferential anodic dissolution of the metallic matrix and the same effect was obtained when the surfaces were electrochemically polished. The second was due to the chevron-like pockets generated by MECT, which probably boosted lubricant entrainment, increased lubricant supply and increased hydrodynamic pressure. 2. As the block-on-ring tests progressed, the preferential wear of the block generated more conformal contacts. In addition, when the rotational speed was smaller, the pockets were worn out. Both effects contributed to no difference in the friction presented by the different surface conditions after prolonged rubbing. 3. For the higher rotational speed, the pockets remained for longer periods. As the test progressed, the contact became more conformal and the lubrication regime probably shifted to full film conditions, where the effects of surface texturing on friction were negligible. 4. The wear scars found in gray cast iron blocks showed a fairly continuous carbon-rich tribolayer, which was not present for electrochemically polished cast iron.
Acknowledgements The authors are grateful to the Brazilian research agencies CNPq (309481/2014-7) Capes PROEX: FAPEMIG (TEC APQ-0174812) for financial support and to Dr. Nick Morris and Prof. Ramin Rahmani from Loughborough University/UK for valuable discussions.
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Reddyhoff, Combined friction and wear reduction in a reciprocating contact through laser surface texturing, Wear 358–359 (2016) 51–61. [15] A. Kovalchenko, O. Ajayi, A. Erdemir, G. Fenske, I. Etsion, The effect of laser texturing of steel surfaces and speed-load parameters on the transition of lubrication regime from boundary to hydrodynamic, Tribol. Trans. 47 (2004) 299–307. [16] A. Ronen, I. Etsion, Y. Kligerman, Friction-reducing surface-texturing in reciprocating automotive components, Tribol. Trans. 44 (2001) 359–366. [17] N. Morris, M. Leighton, M. De la Cruz, R. Rahmani, H. Rahnejat, S HowellSmith, Combined numerical and experimental investigation of the micro-hydrodynamics of chevron-based textured patterns influencing conjunctional friction of sliding contacts, Proc. Inst. Mech. Eng., Part J: J. Eng. Tribol. (2014). [18] R. Rahmani, I. Mirzaee, A. Shirvani, H. 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Tribological behavior of gray cast iron textured by maskless electrochemical texturing L.R.R. da Silva a, H.L. Costa a,b,n a b
Universidade Federal de Uberlândia, Laboratório de Tribologia e Materiais, Av. João Naves de Ávila, 2121, Bloco 1M, 38408-100 Uberlândia, MG, Brazil Universidade Federal do Rio Grande, Escola de Engenharia, Av. Itália, km. 8, Rio Grande, RS, Brazil
art ic l e i nf o
a b s t r a c t
Article history: Received 30 August 2016 Received in revised form 4 January 2017 Accepted 9 January 2017
Gray cast iron is widely used in diesel automotive engines due to its combination of low cost and good fusibility, but also to the presence of graphite flakes in their microstructure. Besides their role as mechanical dampers, graphite flakes can act as solid lubricant, reducing friction and wear of moving components. Surface texturing has been intensively investigated in recent years to improve lubrication and reduce friction and wear in tribological applications, but little attention has been given to surface texturing of gray cast irons. On the other hand, surface texturing adds costs, which can supplant the benefits achieved, in particular when high cost surface texturing techniques are employed. Maskless electrochemical texturing (MECT) is a promising technique, which combines simplicity, low cost and high speed. Recently, MECT was adapted to texture gray cast irons. This work investigates the tribological performance of gray cast iron textured by MECT using lubricated block-on-ring tests under two different rotational speeds and a constant normal load. The blocks consisted of textured cast iron and the rings of SAE 4620 steel. For comparative purposes, untextured polished gray cast iron blocks, as well as gray cast iron blocks subjected to electrochemical polishing (employing the same electrolyte used during MECT) were tested. The results showed a large reduction in friction (up to 2.5x) and wear (higher than 5x) when comparing the polished and the textured samples. Comparing the textured samples and the electropolished samples, this improvement in performance reduced, showing that the benefits were partly due to the texture pattern itself (arrays of pockets) and partly to the exposure of the graphite flakes during electrochemical dissolution of the metal matrix. As the wear progressed and the contact between the block and the ring became more conformal, the benefits obtained by MECT (and electro polishing) were reduced. & 2017 Elsevier B.V. All rights reserved.
Keywords: Gray cast iron Surface texturing Maskless electrochemical texturing Graphite
1. Introduction In vehicles that use internal combustion engines, friction of internal components account for about a third of the energy loss [1,2]. Some alternatives to reduce friction between these components are the use of materials with self-lubricating properties such as gray cast iron [3], of low friction coatings and the application of surface texturing, as well as synergistic responses between them [4,5]. Gray cast iron liners are of particular use in diesel engines [6–8]. The effect of surface texturing under lubricated conditions is strongly dependent on the lubrication regime and on the amount of lubricant supply available. Under full film lubrication conditions, the lubricant pockets may act as microbearings, increasing the n Corresponding author at: Universidade Federal do Rio Grande, Escola de Engenharia, Av. Itália, km. 8, Rio Grande, RS, Brazil. E-mail address: [email protected] (H.L. Costa).
http://dx.doi.org/10.1016/j.wear.2017.01.028 0043-1648/& 2017 Elsevier B.V. All rights reserved.
hydrodynamic pressure, and therefore helping to prevent asperity contact [9–11]. However, due to the reduction in the contact area, which increases contact pressure [9], and also to the occurrence of cavitation [12], surface texturing may lead to negligible [13] or even deleterious effects [12,14]. Under mixed lubrication conditions, the reported effects seem more significant [15–18]. The effect of the pockets as microbearings increases film thickness, reducing the speed or load range for which the lubricant does not separate the surfaces completely, i.e., shifts the Stribeck curves, expanding the ranges for which full film lubrication occurs [13]. Besides, surface texturing increases the time before scuffing occurs when the supply of lubricant is limited [19,20], which is particularly relevant under starved lubrication conditions [21]. Under boundary lubrication, hydrostatic reaction forces in the pockets may help to increase film thickness [22,23], besides the additional lubricant supply and the entrapment of wear debris [24], but the high stresses induced around the pockets [25] may lead to detrimental effects [7].
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The literature is abundant in works regarding the effect of the texture pattern geometry on the tribological performance of textured surfaces under lubricated conditions. They include works regarding the effect of the width of the motifs [11,21,26], the ratio between the height and width of the motifs [9,16], the percentage of area coverage [9,10,12,16,26] and the shape of the motifs [9,18,26], particularly circular dimples, grooves and interrupted grooves. Costa and Hutchings [9] showed a superior performance of chevron-like motifs to increase load capacity under hydrodynamic lubrication when compared with arrays of circular pockets and grooves. Later, Morris et al. [17] used numerical simulations to optimize the geometry of chevron arrays aiming to reduce friction coefficient and to increase load capacity in pistonring systems. The use in industrial scale of surface texturing for reduction of frictional losses depends on the existence of techniques capable of texturing surfaces rapidly and inexpensively in order to make the method cost-effective. Although laser texturing is the technique most widely used industrially [27], it may present some drawbacks, such as the high cost of the equipment involved in the process, large processing time when texturing large areas [21] and the presence of heat affected zones and recast areas, which can act as abrasives if not removed by subsequent process [13]. This last effect can be largely reduced with the use of femtosecond lasers [4,28], but this adds extra cost to the equipment. Among alternative surface texturing techniques, Maskless Electrochemical Texturing method (MECT) was developed as a simple, low cost, high speed texturing technique, particularly suited to metals [29,30]. Initially, it was developed to texture plane surfaces, mainly for carbon steel [29,30] and copper [31]. Very recently, MECT was adapted to texture diesel cylinder liners [32], producing arrays of chevrons, which faced two challenges. The first was the cylindrical geometry and the second was texturing a material containing phases with very different chemical characteristics: the metallic matrix and the graphite flakes. Despite the difficulty to texture gray cast iron, texture patterns with good resolution were produced cheaply and quickly. The results showed that anodic dissolution mainly occurred in the metallic matrix, but not in the graphite. Graphite was only removed after it lost the mechanical support provided by the metallic matrix. This resulted in exposure of the graphite flakes, in addition to the formation of the texture pattern. Therefore, there is potential for an extra beneficial effect when using MECT to texture grey cast iron, since the graphite flakes may act as solid lubricant, in addition to the benefits obtained by the presence of the texture pattern. This paper aims to investigate the tribological behavior of gray cast iron textured by MECT, since this material is widely used in lubricated sliding conditions. For that, lubricated sliding tests of plane textured surfaces were carried out using a block-on-ring tribometer. In addition, mechanically polished surfaces, as well as surfaces obtained by conventional electrochemical polishing were also tested, since this method is used to highlight the graphite flakes in gray cast iron [33]. This comparison helped to separate the contribution originated from the graphite flakes from that from the texture pattern itself. The results evidenced a strong reduction in friction and wear when MECT is used to texture gray cast iron.
spacing of 0.33 70.11 mm. Vickers hardness tests using normal load of 612.5 N showed and average value of 1.810 70.025 GPa. The dimensions of the blocks followed recommendations from ASTM D271-94(2014) standard [34], which were 6.35 mm (width), 15.76 mm(length) and 10 mm (height). All blocks were ground and polished, leading to a 3D RMS surface roughness (Sq) of 0.09870.004 μm. The rings were provided by Falexs according to ASTM D271-94 standard [34]. They were manufactured using SAE 4620 carburized steel, with a Vickers hardness ranging from 7.379 to 7.595 GPa. The width was 8.5 mm and the external diameter was 35 mm. The surface roughness of the rings was assessed measuring three different areas of 1 mm2 in each ring, using a UBM Microfocus Expert IVs laser interferometer, and only rings with Sq between 0.19 and 0.21 μm were used in the tests. Considering the hardness values for the block (1.810 GPa) and the rings (7.487 GPa), wear was mainly expected to occur in the blocks and the ring should remain intact [35]. The lubricant used was a SAE 40 mineral oil with a viscosity (ηo) ranging from 159.9 cSt at 40 °C to 15.2 cSt at 100 °C, and a viscosity pressure coefficient (α) of 2.93 GPa 1. 2.2. Surface texturing The gray cast iron blocks were textured by Maskless Surface Texturing method (MECT). The basic apparatus used is described in Fig. 1 [29], where (1) is a constant voltage source, (2) is the electrolyte reservoir, (3) is the sample to be textured, (4) is the signal modulator, (5) is the electrolyte pump, (6) is the representation of the connections between the electrolyte pump and reservoir and (7) is the tool containing the pattern to be transferred to the sample. The tool cover was made with AISI 304 austenitic stainless steel. The pattern to be transferred to the sample was cut into this cover by laser micro drilling. One of the faces of the tool cover received electrostatic painting after drilling, as shown in Fig. 2, which was responsible for the electrical insulation between the tool and the workpiece. The texture pattern consisted of an array of chevrons, based on the optimization results presented in [17]. During MECT, the electrolyte was a 200 g L 1 NaCl solution in distilled water, as suggested in [9,12], the voltage was 7.5 V, the texturing time was 60 s and the gap between tool and specimen was 100 μm [12]. For comparison, mechanically polished gray cast iron blocks were also tested. In addition, a third group was electrochemically polished after the mechanical polishing. Electrochemical polishing
2. Experimental methods and materials 2.1. Samples and lubricant The blocks used in the sliding tests were manufactured using gray cast iron presenting a pearlitic matrix with an interlayer
Fig. 1. MECT Apparatus used for surface texturing the blocks; from [30], with permission.
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with Nital (Fig. 3.b). This figure is presented only to evidence the material microstructure, but the polished samples used in the sliding tests were not etched. The electrochemically polished surface is illustrated in Fig. 3.c, where the pearlitic matrix was revealed by the dissolution during electrochemical polishing, as well as the graphite flakes, which are in relief in relation to the pearlitic matrix. This is a result of the preferential anodic dissolution of the metallic matrix by the electropolishing process, exposing the graphite flakes [33]. A textured sample is shown in Fig. 3.d, which shows the chevron patterns and the material microstructure, evidenced during electrochemical dissolution that also occur in the areas outside the chevrons (although at a much lower rate). Graphite flakes are evidenced both inside and outside the chevrons. Fig. 3.e presents a magnified view of the areas outside (but near) the chevrons, where the pearlitic nature of the matrix is again evidenced and the graphite flakes appear elevated in relation to the matrix. Therefore, MECT may come to possess synergistically the advantages of the surface texturing (pockets) as well as a secondary effect (exposure of graphite flakes) conferred by electrochemical polishing, suggesting that MECT could result in further gains compared with other texturing methods such as laser texturing for gray cast iron. An example of the 3D surface topography generated by MECT of gray cast iron is illustrated in Fig. 4. Fig. 2. Tool used to surface texture the gray cast iron blocks by MECT.
was carried out at a voltage of 7.5 V, during 60 s using a 200 g L 1NaCl solution as the electrolyte. The mechanically and electrochemically polished blocks had an initial surface roughness of 0.098 70.004 and 0.201 70.032 mm, respectively. The textured blocks had a surface roughness of 0.171 70.027 mm in the areas outside the imposed surface patterns. Micrographs for the three surface conditions used in this work are presented in Fig. 3. Fig. 3.a shows the mechanically polished surface, where only the graphite flakes can be observed. The pearlitic matrix of the polished sample is evidenced by etching
2.3. Block on ring tests The tribological testes were carried in a Falexs block on ring tribometer (Fig. 5), where (1) is the load cell, (2) is the lever arm used to apply the normal load at a ratio of 1:10, (3) is the thermocouple, (4) is the test ring, (5) is the test block and (6) is the hermetically closed test chamber. For the three groups of samples (mechanical polishing, electrochemical polishing and MECT) tests were carried out using a normal load of 343 N and ring rotational speeds of 200 and 600 rpm. Each test had a duration of 20 min. The lubricant temperatures during the tests were measured as approximately 90
Fig. 3. Optical micrographs for the surface conditions used in this work: (a) mechanically polished surface; (b) mechanically polished surface etched with NITAL; (c) electrochemically polished surface; (d) chevrons produced by MECT; (e) amplified view of the surface near the chevrons produced by MECT.
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Table 1 λ parameter values for each condition tested.
Fig. 4. 3D surface map of the textured samples; laser interferometry.
and 75 °C for the 200 and 600 rpm tests, respectively. Each test was replicated three times to ensure reproducibility. The initial mean contact pressure (Pm) and the semi contact width (a) can be calculated using hertz equation for a line contact, as shown in Eq. (1) [36]. 1
Pm=
⎛ wR* ⎞ 2 π E*w ⎟ ; a = 2⎜ 4 R*Lπ ⎝ LE* ⎠
(1)
where E* is the contact young modulus, R* is the contact radius of curvature, L is the contact length and w is the normal load. Using this equation a mean contact pressure of 0.74 GPa between the block and the ring and a semi contact width of 280 mm were found to operate at the beginning of the tests. The initial fluid film thickness (H*) was calculated using the Dowson-Hamrock equation [37–39], as shown in Eq. 2.
H *=1. 6G 0.6U 0.7W −0.13
G=
αE *
(3)
U=
Vη0 E*R*
(4)
W=
w E*A
Rotational speed (rpm)
H* (μm) S*q (μm)
Mechanically Polished Electrochemically Polished Textured Mechanically Polished Electrochemically Polished Textured
200 200
0.324 0.324
0.2337 0.004 1.455 0.284 7 0.032 1.143
200 600 600
0.324 0.699 0.699
0.263 7 0.027 1.231 0.2337 0.004 3.138 0.284 7 0.032 2.465
600
0.699
0.263 7 0.027 2.656
where α is the lubricant pressure viscosity coefficient (GPa 1), V is the speed of the ring in relation to the block (m s 1), ηo is the lubricant viscosity (Pa s), w is the normal load and A is the contact area. For 200 and 600 rpm, the initial film thicknesses (H*) were calculated as 0.223 and 0.311 mm, respectively. Using those values and the 3D RMS surface roughness for each condition, it was possible to calculate the lambda parameters (λ) for each condition (Table 1), defined in the Eq. (6), where S*q is the combined roughness of the block (SqB) and ring (SqR), given by the Eq. (7) [40].
H* Sq*
λ=
Sq*=
(6) SqB2 + SqR2
(2)
where G, U and W are respectively, the non-dimensional material, speed and load parameters [37,38], as shown in Eqs. (3)–(5).
λ parameter
Surface condition
(7)
The majority of the conditions tested presented λ values between 1 and 3, placing the initial test conditions in the mixed EHL lubrication regime [41,42]. One exception is the mechanically polished specimen at 600 rpm, which showed a value slightly above 3. However, due to the wear caused by the initial sliding, it is expected a rapid increase in the surface roughness, so that the λ value should decrease quickly to values below 3. On the other hand, it is expected that wear should also change the contact geometry, making the contact more conformal with increased sliding. This will be discussed in the Results section.
(5)
s
Fig. 5. Block on ring Falex tribometer.
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Fig. 6. Friction coefficient versus time: (a) 200 rpm; (b) 600 rpm.
3. Results and discussion Fig. 6 shows the friction coefficient values measured for the tests at 200 rpm (a) and 600 rpm (b). Only one curve is presented for each condition, but the different repetitions carried out showed very reproducible curves. The evolution of friction coefficient with time showed the typical behaviour commonly reported in the literature for block-on-ring tests [35,43], where an initial peak is followed by a sharp reduction in friction. This initial running-in period is observed for all surface conditions and speeds. The running-in period was longer for the polished specimens than for electropolishing and MECT conditions. This is probably due to graphite flakes, which were exposed by the anodic dissolution of the pearlitic matrix around them. The solid lubricant properties of the graphite flakes exposed during electropolishing [33] and MECT [32] probably contribute to reduce friction in the initial stages of the test, where the sliding speed is lower and therefore lubricant film thickness is thinner [42], acting in the boundary regime. The occurrence of boundary lubrication in the initial stages of block-on-ring tests has been elegantly demonstrated in [35]. After the short initial transient, a drastic reduction in friction was observed for the textured blocks when compared with the polished blocks, whereas the electropolished blocks presented an intermediate behavior. Therefore, two effects must be contributing to improve the performance of the textured blocks. The first is related to the chevron-like pockets. Under mixed lubrication conditions, surface texturing can been very successful, where the contribution of the pockets on hydrodynamic pressure distribution
and the additional lubricant supply seem to contribute to reduce friction [15,17,44], including for cast irons [45]. The second are the exposed graphite flakes, which acts as a solid lubricant in the contact. This second effect also occurs for the electropolished blocks, justifying their intermediate behavior. The positive effect of graphite during sliding of cast irons [33], including for block-onring tests [46], has been largely reported in the literature. Another aspect that is common to all the curves presented is that after the running-in period, friction increases with test time and then decreases. This friction increase is particularly accentuated for the polished block tested at 600 rpm. In order to investigate further this behaviour, some tests were interrupted after only 250 s of test, in order to identify possible phenomena within this period where friction increases. The wear scars produced in the interrupted tests were measured by 3D laser interferometry. Despite the longer sliding distances that were experienced under higher rotational speeds, the volume loss after 250 s of sliding was much larger for 200 rpm (0.03 mm3) than for 600 rpm (0.01 mm3). This occurs because at higher speeds the lubricant film thickness increases [42], reducing asperity contact and then wear. Another point is that the measured temperature was lower for tests at 600 rpm (75 °C) than at 200 rpm (90 °C), probably due to a more efficient convection. At lower temperatures, lubricant viscosity increases, increasing film thickness [42]. Since the rings were much harder than the blocks, wear was mostly restricted to the blocks and wear scars could not be identified in the rings. Therefore, wear progressed in terms of changing contact geometry from an initial line contact to a more conformal contact, which occurred faster for the tests at 200 rpm. The
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Fig. 7. 3D surface topography of the worn scars in the mechanically polished blocks after 250 s of rubbing: (a) 200 rpm; (b) 600 rpm.
occurrence of preferential wear of a softer block and therefore the transition from a line non-conformal contact to a more conformal contact was assessed and discussed in [35] using a displacement sensor. As an example, Fig. 7 presents the wear scars produced after those interrupted tests for the polished blocks. The length of the wear scars (around 5 mm) were smaller than the ring width (8.5 mm). Although we cannot precise the reason for this behaviour, other works in the literature [35] have also showed wear tracks that are smaller than the ring width. One possible explanation is a more efficient lubricant entrainment near the scar extremes, making wear negligible in these areas. In [35], low magnification stereo micrographs near the contact edges indeed suggest better lubricant entrainment near the width extremes. With the software MontainsMap s, the average radius of the wear scars were computed using 10 equally-spaced radius measurements (30 measurements for the 3 repetitions). For the test at 200 rpm, the radius of the scar in the block was 18.1 70.22 mm, which is very close to the ring radius (17.7 mm), whereas for 600 rpm it was 22.4 70.97 mm, i.e., the worn block was less conformal to the ring. Contact pressure is substantially reduced in conformal contacts when compared with non-conformal contacts [47]. So, it seems that as the wear of the blocks proceeds, contact pressure reduces, increasing film thickness. Therefore, after a certain level of conformity, film thickness may increase so that asperity contact reduces significantly and therefore friction reduces with further sliding. Further analysis of Fig. 6 shows that for the tests at 200 rpm, after around 500 s of sliding, the friction behavior of the textured and electropolished blocks became undistinguishable, although friction remained lower than for the mechanically polished block. It is plausible that after 500 s, the wear of the blocks was sufficient to completely remove the chevron-like pockets, so that the only beneficial effect that still remained was that of the exposed graphite flakes. Indeed, at the end of tests at 200 rpm, the chevrons were nearly completely removed inside the wear scars, as shown
in the images presented in Fig. 8.c. Fig. 8 consists of scanning electron microscopy (SEM) images of the wear scars using back scattered electrons (BSE). 3D surface topography of the wear scars showed small traces of reminiscent chevrons (Fig. 9), but they were probably too shallow to have any effect on lubricant entrainment or hydrodynamic pressure. After around 1000 s of test, all blocks showed similar low friction values of around 0.09. Since the exposed graphites are no longer visible within the wear scars at the end of the tests for the textured and electropolished blocks (Fig. 8.b and c), their friction behaviour becomes very similar to that presented by the mechanically polished block. For the tests at 600 rpm, friction coefficients of the textured blocks were the lowest and the electropolished blocks again showed an intermediate behaviour. After around 1000 s of test, the three blocks presented the same friction values. As shown in Fig. 10, chevron-like pockets and exposed graphite flakes still remained inside the wear scars at the end of the tests. With the 3D interferometric maps of the wear scars (Fig. 10), the average scar radii were measured at the end of the tests and all values were very similar to the ring curvature. With high conformity between ring and wear scar, contact pressures were quite low at the end of the tests. On the other hand, the lubricant film should be thicker at 600 rpm than at 200 rpm. Therefore, it is postulated that for the tests at 600 rpm, the lubrication regime shifted to nearly full film lubrication at the end of the tests, with negligible asperity contact. This hypothesis is corroborated by the low friction values in this region for the three blocks (around 0.06). Various authors have proposed that under full film lubrication, the effects of surface texturing can be very small. Costa and Hutchings [9] did not find differences in the friction coefficients presented by smooth and textured surfaces, despite a small increase in load carrying capacity for the textured surfaces. In very recent works [14,48], the simultaneous measurement of friction coefficient and film thickness (by ultrathin laser interferometry) in tests under full film lubrication showed that although surface texturing boosts lubricant entrainment, the effects on friction are
Fig. 8. BSE SEM images of the wear tracks on the blocks after the tests at 200 rpm: (a) mechanical polishing; (b) electrochemical polishing; (c) MECT.
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Fig. 9. 3D surface topography of the wear scars on the blocks after the tests at 200 rpm: (a) mechanical polishing; (b) electrochemical polishing; (c) MECT.
Fig. 10. BSE SEM images of the wear tracks on the blocks after the tests at 600 rpm: (a) mechanical polishing; (b) electrochemical polishing; (c) MECT.
Fig. 11. 3D surface topography of the wear scars on the blocks after the tests at 600 rpm: (a) mechanical polishing; (b) electrochemical polishing; (c) MECT.
negligible or even slightly detrimental. One point to be noted is that the total width of each chevronlike motif that composes the texture pattern (around 2.3 mm) is
substantially wider the initial elastic contact width (560 mm). The literature [7,49] mainly suggests that the size of the individual motifs should be smaller than the contact width, in particular
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Fig. 12. Volume wear rates for the blocks.
Fig. 13. BSE SEM images showing detail of the worn surfaces after the tests: (a) mechanical polishing at 200 rpm; (b) electrochemical polishing at 200 rpm; (c) MECT at 200 rpm; (d) mechanical polishing at 600 rpm; (e) electrochemical polishing at 600 rpm; (f) MECT at 600 rpm.
under full film conditions, due to the fact that the lubricant may flow away from the contact. However, in this work the lubrication regime is mostly mixed, so that wear progresses quickly, increasing the contact width substantially. As seen in Figs. 9 and 11, by the end of the tests the wear craters are wider than the individual motifs. Another point is that the angle of the chevrons ensures that even while the contact width is narrower there is substantial lubricant entrapment within the contact. The motifs used in this work were the smallest chevrons that could be produced, but this can be reduced if finer laser drilling is used to produce the tools that are used in MECT. Another feature that clearly emerges from the 3D maps of the wear scars on the blocks is that wear decreases when the rotational speed increases from 200 rpm to 600 rpm. Wear rates (k) were calculated using the Archard equation [50], as shown in the Eq. (8), where d is the sliding distances and V is the volume loss, calculated using the software MontainsMaps.
k=
V wd
(8)
The global wear rates, measured at the end of the tests, are presented in Fig. 12, which shows an approximate reduction of around 82% of the wear rates between the tests at 200 rpm and 600 rpm, which occurs due to thicker hydrodynamic films that are formed under higher sliding speeds. More importantly, this graph shows a reduction of around 81% in the wear rates found for the textured block when compared with the polished block both at 200 rpm and 600 rpm. For comparison, the reduction when electropolishing was used instead of MECT was of 72% at 200 rpm and of 65% at 600 rpm. Therefore, the wear reduction of the textured blocks is not only due to the exposure of the graphite flakes but also due to presence of the chevron-like pockets. Fig. 13 shows BSE SE images of the wear scars produced after the tests at 200 and 600 rpm using higher magnification. For the mechanically polished block (Fig. 13.a and d), large dark
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agglomerates are seen and EDS analysis confirmed that they are very rich in carbon (almost 90%C). For the electrochemically polished (Fig. 13.b and e) and MECT blocks (Fig. 13.c and f), instead of isolated carbon agglomerates, very fine and evenly distributed carbon-rich regions exist forming a fairly continuous carbon-rich tribolayer, which is probably responsible for the low friction coefficients and wear rates found in the experiments.
4. Conclusions This work investigated the tribological performance of gray cast iron surfaces textured by Maskless Electrochemical Texturing using block-on-ring tests. The results showed that: 1. Surface texturing by MECT reduces drastically friction and wear, which was clearly attributed to two effects. The first was the exposure of the graphite flakes by the preferential anodic dissolution of the metallic matrix and the same effect was obtained when the surfaces were electrochemically polished. The second was due to the chevron-like pockets generated by MECT, which probably boosted lubricant entrainment, increased lubricant supply and increased hydrodynamic pressure. 2. As the block-on-ring tests progressed, the preferential wear of the block generated more conformal contacts. In addition, when the rotational speed was smaller, the pockets were worn out. Both effects contributed to no difference in the friction presented by the different surface conditions after prolonged rubbing. 3. For the higher rotational speed, the pockets remained for longer periods. As the test progressed, the contact became more conformal and the lubrication regime probably shifted to full film conditions, where the effects of surface texturing on friction were negligible. 4. The wear scars found in gray cast iron blocks showed a fairly continuous carbon-rich tribolayer, which was not present for electrochemically polished cast iron.
Acknowledgements The authors are grateful to the Brazilian research agencies CNPq (309481/2014-7) Capes PROEX: FAPEMIG (TEC APQ-0174812) for financial support and to Dr. Nick Morris and Prof. Ramin Rahmani from Loughborough University/UK for valuable discussions.
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