Effect of electrochemical potential on tribocorrosion behavior of low temperatureplasma carburized 316L stainless steel in 1 M H2SO4 solution

Y. Sun a,⁎, E. Haruman ba Department of Engineering, Faculty of Technology, De Montfort University, Leicester LE1 9BH, UKb Faculty of Industrial Engineering, Bakrie University, Jakarta 12920, Indonesia

a b s t r a c ta r t i c l e i n f o

Article history:Received 2 February 2011Accepted in revised form 15 March 2011Available online 8 April 2011

Keywords:Stainless steelCarburizingTribocorrosionCorrosive wear

Low temperature plasma carburizing is a surface engineering technique developed for austenitic stainlesssteels to achieve combined improvements in tribological and corrosion resistant properties. However, verylittle work has been conducted to study the response of carburized stainless steel to combined actions bywearand electrochemical corrosion. In the present work, experiments have been conducted to study thetribocorrosion behavior of low temperature plasma carburized AISI 316L stainless steel, under unidirectionalsliding in 1 M H2SO4 solution, using a pin-on-disk tribometer integrated with a potentiostat forelectrochemical control. Sliding wear tests were conducted under potentiodynamic and potentiostaticconditions at a wide range of applied potentials. It is found that the carburized layer exhibits much bettertribocorrosion resistance than the untreated specimen at anodic potentials, but is not effective in improvingwear resistance at cathodic potentials. The results are discussed in terms of the accumulation of the third bodywear particles, material transfer and the contribution of mechanical wear and chemical wear to overallmaterial removal by tribocorrosion.

© 2011 Elsevier B.V. All rights reserved.

1. Introduction

Low temperature carburizing is a technique developed in recentyears to engineer the surfaces of austenitic stainless steels to achieveenhanced mechanical and tribological properties without compromis-ing their good corrosion resistance [1–4]. Theprocess is normally carriedout at temperatures below 500 °C in order to avoid chromium carbideprecipitation in the carburized layer [2,5,6]. During low temperaturecarburizing, carbon atoms diffuse into the surface region of thecomponent and are dissolved in the face-center cubic (fcc) austenitelattices to form a hardened layer up to 40 μm thick supersaturated withup to 3 wt.% carbon [6–10]. Such a precipitation-free carburizedstructure, which is also termed the carbon expanded austenite orcarbon S-phase [1,3], ensures that chromium remains in the solidsolution form, and thusmaintains the corrosion resistant characteristicsof stainless steels.

Tribological tests have shown that the carbon expanded austenitehas excellent wear resistance under unlubricated conditions [11–14].Electrochemical tests have also confirmed the good resistance ofcarbon expanded austenite to pitting corrosion in chloride containingsolutions [2,11,15]. More detailed electrochemical depth profilingmeasurements by Sun [16] further confirmed that the precipitation-

free carburized layer possesses excellent corrosion resistancethroughout its whole thickness in 1 M H2SO4, at least as good asuntreated 316 stainless steel. A further step has been moved forwardin the present work to conduct electrochemical testing of carburized316L stainless steel in the 1 MH2SO4 solution under controlled slidingwear conditions. Such a test scheme is more practical because inmanypractical situations, surface engineered stainless steel components aresubjected to combined corrosion and wear actions, such as inbiomedical and food processing applications, where the componentsare subjected to scratching, abrasion, erosion, and other forms of weardamaging in a corrosive environment. These can lead to the damageor even complete removal of the passive film from the contact surface,resulting in accelerated metal dissolution, which in turn can in manycases lead to accelerated wear [17–27]. So far, very few efforts havebeen made to study the tribocorrosion behavior of low temperaturecarburized stainless steels under combined wear and electrochemicalcorrosion conditions.

In the present work, a systematic investigation has been carriedout to study the tribocorrosion behavior of low temperature plasmacarburized AISI 316L stainless steel under unidirectional sliding in 1 MH2SO4 solution, using a pin-on-disk tribometer integrated with apotentiostat for electrochemical control. A series of experiments havebeen conducted, including potentiodynamic polarization measure-ments during sliding wear and sliding tests under potentiostaticconditions at a wide range of potentials.

Surface & Coatings Technology 205 (2011) 4280–4290

⁎ Corresponding author. Tel.:+44 116 2577072.E-mail address: ysun01@dmu.ac.uk (Y. Sun).

0257-8972/$ – see front matter © 2011 Elsevier B.V. All rights reserved.doi:10.1016/j.surfcoat.2011.03.048

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2. Experimental

2.1. Low temperature plasma carburizing

The material used in the present work is AISI 316L austeniticstainless steel with the following nominal chemical compositions (in

wt.%): 0.02% C, 18.61% Cr, 11.83% Ni, 2.87% Mo and balance Fe. Thematerial was received in the form of a hot-rolled plate and wasmachined into test specimens of 20×20×2 mm3 sizes. Before plasmacarburizing, the specimens were wet ground using a series of SiCgrinding papers down to the 1200 grade. The specimens were thencharged onto the cathodic working table of a commercial plasmanitriding unit adapted for carburizing treatments. The low tempera-ture plasma carburizing process was conducted in the SurfaceEngineering Group of The University of Birmingham and has beendescribed elsewhere [2,10]. The carburizing parameters are the sameas those reported recently [16]. Briefly, the carburizing temperatureand time employed in this work are 470 °C and 15 h, respectively,which are a good combination to achieve a relatively thick carburizedlayer comprising a single expanded austenite phase without chromi-um carbide precipitation [2].

Fig. 1 shows the scanning electronmicrograph of the cross section ofthe carburized specimen (after etching in the 50% HCl+25% HNO3+25% H2O reagent) and the corresponding microhardness profilemeasured across the carburized layer. It can be seen that plasmacarburizing at 470 °C for 15 h produced a carburized layer about 25 μmthick at the surfaceof the specimen. Thehardnessof the carburized layeris about 900 HV at the surface, and decreases gradually towards thesubstrate to reach the value of about 200 HV of the untreated material.X-ray diffraction analysis could not detect any carbide in the carburizedlayer; instead diffraction peaks corresponding to expanded austenitewere recorded [16]. This confirms thewell-established fact that the lowtemperature carburized layer is precipitation-free and comprises asingle expanded austenite phasewith an expanded fcc lattice due to thesupersaturation of carbon [4–6,8,10].

2.2. Tribocorrosion testing

Before tribocorrosion testing, the carburized specimens werepolished using 6 μm and then 1 μm diamond polishing compoundsto achieve a mirror-like surface. For comparison purpose, untreatedspecimens were also tested in this work, which were also polished tothe 1 μm finish. After polishing, the specimens were cleaned in soupwater, running water and then methanol. Before charging into thetribo-electrochemical cell, a test area of 1.3×1.3 cm2 was created bymasking all other areas using adhesive insulating tape. This test areawould be exposed to the solution during tribocorrosion testing.

The tribocorrosion test system is illustrated in Fig. 2, which wasassembled by integrating a pin-on-disk tribometer (manufactured by

Fig. 1. Scanning electron micrograph showing the morphology of the carburized layerproduced in the 316L steel at 470 °C for 15 h (a) and microhardness profile measuredacross the carburized layer (b).

Fig. 2. Schematic diagram showing the tribo-electrochemical cell and experimental setup in this work.

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Teer Coating Ltd.) and an electrochemical potentiostat. During thetest, the specimen held in the test cell was rotating against astationary ball, under controlled load and speed conditions. In orderto electrically isolate the specimen from the tribometer duringtribocorrosion testing, the test cell was made of an insulating andcorrosion resistantmaterial, nylon, and the ball holderwas alsomadeof nylon. The ball slider used in thisworkwas an alumina ball of 8 mmin diameter. Such a setup ensured that only the exposed test area ofthe specimen was subjected to electrochemical actions during thetest. All sliding tests were conducted at a contact load of 20 N androtation speed of 60 rpm (contact frequency of 1). The wear track

diameter for all tests was 6 mm. The maximum contact stress resultingfrom a contact load of 20 N is about 1550 MPa, which is much higherthan the yield strength (about 300 MPa) of the untreated 316L steel.Although the yield strength of the carburized layer has not beenmeasured, a value about 1300 MPa seems reasonable because the upperpart of the carburized layer is four times harder than the substrate (seeFig. 1b). Thus the appliedmaximumcontact stress is alsohigher than theyield strength of the carburized layer.

The electrolyte used for the tests was 1 M H2SO4 solution, whichwas prepared from analytical grade chemicals and distilled water. Thetest cell was filled with about 60 ml of the solution. The test specimenserved as a working electrode, a Saturated Calomel Electrode (SCE)was inserted in the test cell 5 mm above the test surface to serve asthe reference electrode, and a platinumwire was used as the auxiliaryelectrode. Electrochemical measurements were realized using an ACMGILL AC potentiostat equippedwith a computer data login, acquisitionand analysis system controlled by the Sequence and Core Runningsoftware. All the tests were performed at room temperature (22 °C)and open to the air. Most of the tests were duplicated and some weretriplicated to ensure reproducibility of test results.

The first series of experiments conducted in this work is cyclicpotentiodynamic measurements, which involved measuring the polar-ization curves of the specimens during sliding and without sliding at asweep rate of 1 mV s−1, starting from 200 mV(SCE) cathodic to theopen circuit potential (OCP) and polarizing anodically until the currentdensity reached 5 mA cm−2. Then the potential scanwas reversed backto the starting point. Under the sliding condition, the polarization curvewas measured during sliding at a contact load of 20 N and a rotationspeed of 60 rpm.Under the conditionwithout sliding, the specimenwasrotating at a speed of 60 rpm without contact with the slider, whichrepresents the condition of the unworn area outside the wear trackduring sliding tests. Before each polarization measurement, the

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Fig. 3. Potentiodynamic polarization curves measured under sliding andwithout slidingfor the untreated and carburized (CT) specimens. Test condition: potential sweeprate: 1 mV s−1; sliding speed: 60 rpm; contact load: 20 N.

Fig. 4. Recorded coefficient of friction curves during sliding under potentiodynamicpolarization. Test condition: same as Fig. 3.

Table 1Summary of test results for carburized and untreated 316L stainless steel at various potentials in 1 M H2SO4 under sliding for 3800 s at 20 N and 60 rpm.

Potential(mV/SCE)

Average current (mA) Average COF Total wear: VT (mm3) Mechanical wear: Vmech (mm3) Chemical wear: Vchem (mm3)

Untreated CT Untreated CT Untreated CT Untreated CT Untreated CT

−900 −3.50 −7.13 0.443 0.428 0.027 0.024 0.027 0.024 0 0−500 −0.54 −0.69 0.412 0.447 0.027 0.023 0.027 0.023 0 0−291 (OCP, raw) 0.081 (Icor)⁎ – 0.431 – 0.039 – 0.030 – 0.009 –−261 (OCP, CT) – 0.018 (Icor)⁎ – 0.384 – 0.023 – 0.021 – 0.00250 0.709 0.219 0.388 0.353 0.255 0.031 0.173 0.006 0.082 0.025500 0.695 0.258 0.338 0.444 0.464 0.040 0.384 0.01 0.08 0.03900 0.765 0.219 0.339 0.454 0.416 0.031 0.328 0.006 0.088 0.025

⁎ Corrosion current measured by standard Tafel extrapolation from the polarization curves shown in Fig. 3.

Fig. 5. Variation of open circuit potential (OCP) with time as a result of sliding wear in1 M H2SO4 solution. Sliding condition: 20 N load and 60 rpm speed. CT: carburizedspecimen.

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specimen was stabilized for 600 s under the respective conditions. Forexample, the polarization curve during sliding was acquired after aninitial sliding for 600 s.

The second series of experiments is potentiostatic measurements,which involved sliding wear tests at constant applied potentials. The

sliding tests were carried out at a fixed contact load of 20 N and speedof 60 rpm for a total duration of 3800 s. The constant potential wasapplied for a total duration of 5600 s: 900 s before the commence-ment of sliding, 3800 s during sliding and 900 s after the terminationof sliding. During each test, the current and coefficient of friction(COF) were recorded continuously at a sampling frequency of 1 Hz.

After each test, the total wear volume from the wear track wasevaluated by measuring the surface profiles across each wear track atfour locations 90° apart, using a stylus surface profilometer. The averagecross-sectional area of the wear track was then estimated from theprofiles by numerical integration and the total wear volume wasobtained by multiplying the cross-sectional area by the circumferentiallengthof thewear track. The surfacemorphologyof the tested specimenswas examined by optical and scanning electron microscopy (SEM). Themorphology of the wear scar on the alumina ball was also examinedmicroscopically.

3. Results

3.1. Potentiodynamic sliding tests

Fig. 3 shows the potentiodynamic polarization curves measuredunder sliding and without sliding. The forward scan curves measuredwithout sliding for both the untreated and carburized specimens aresimilar to those reported previously which were obtained under staticcondition without specimen rotation [16]. Under the sliding

Fig. 6. Current transients recorded before, during and after sliding tests at variouspotentials indicated. Sliding condition: 20 N load and 60 rpm speed. CT: carburizedspecimen.

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Fig. 7. Surface profilesmeasured across thewear tracks produced on the carburized (CT) and untreated specimens at (a)−500 mV(SCE), (b) OCP; (c) 500 mV(SCE) and (d) 900 mV(SCE).Sliding condition: 20 N load and 60 rpm speed for 3800 s.

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condition, the potentiodynamic behavior of both specimens issignificantly different from that observed without sliding. Firstly,the zero-current potential (corrosion potential) of both specimens isshifted cathodically by more than 100 mV(SCE). Such a cathodic shiftof corrosion potential due to sliding has been observed by manyinvestigators in passive systems and is attributed to the destruction ofthe passive state and the activation of the material inside the weartrack [21,23,25,28,29]. Secondly, the anodic current densities areincreased significantly, by nearly two orders of magnitude for bothspecimens. Such a current increase due to sliding is more significantfor the untreated specimen than for the carburized specimen,suggesting the effectiveness of the carburized layer in reducinganodic dissolution during the sliding process. Interestingly, for bothspecimens, the current density levels off over a wide potential rangeafter the initial increase with increasing anodic potential, indicatingthat a certain degree of passivity or “pseudo-passivity” is maintainedin the wear track at contact intervals during the sliding process. The“pseudo-passivity”-maintaining current density is very high: since theanodic current measured during sliding mainly flows through thewear track area, which is much smaller than the overall specimen areaexposed to the solution, the actual current density in the wear track ismuch higher than that shown in Fig. 3. By estimating the final weartrack area from the track width measured after the test, it is estimatedthat the “pseudo-passivity”-maintaining current density in the weartrack during sliding is roughly 3 mA cm−2 for both specimens. Such ahigh anodic current density, as compared to that measured withoutsliding (about 0.005 mA cm−2) can be explained by the repeatedremoval (depassivation at contacts) and re-growth (repassivation atcontact intervals) of the oxide film during the sliding process. FromFig. 3 it can also be seen that the reverse scan curves measured forboth specimens during sliding nearly overlap with the correspondingforward scan curves, suggesting that the increase in current density inthe transpassive region (above 950 mV(SCE)) is most probably dueto oxygen evolution and not to pit formation, as confirmed bymicroscopic examination.

The recorded coefficient of friction curves during the potentiody-namic sliding tests are shown in Fig. 4. Both specimens exhibit similarfrictional behavior at open circuit (OC), in the cathodic region and inthe low-potential anodic region during both forward and reversescans. In the high-potential anodic and transpassive regions, thecoefficient of friction of the carburized specimen is higher than that ofthe untreated specimen.

3.2. Potentiostatic sliding wear test

Having established thepolarization curves under sliding andwithoutsliding, the next series of experiments involved sliding tests underpotentiostatic conditions at two cathodic potentials (−500 mV(SCE)and −900 mV(SCE)), at open circuit and at three anodic potentials(50 mV(SCE), 500 mV(SCE) and 900mV(SCE)). The results obtained aresummarized in Table 1, which lists the values of average currentmeasured during sliding, average coefficient of friction, and total wearvolume (VT) measured by profilometry.

3.2.1. Effect of potential on electrochemical behaviourDuring sliding at the cathodic potentials of −500 mV(SCE) and

−900 mV(SCE), the measured currents from both specimens arenegative (Table 1), confirming that no electrochemical corrosionoccurs at these potentials. Thus, material loss from the wear trackmust be due to pure mechanical wear. Several investigators employedsuch a test scheme to separate the pure mechanical component fromchemical component during tribocorrosion processes [17,30]. How-ever, there are some concerns regarding the correlation between puremechanical wear at cathodic potentials and that occurring at opencircuit and at anodic potentials because of the different surfaceconditions at different potentials [23]. It is also possible that hydrogen

reduction as a main cathodic reaction can lead to hydrogen chargingof the test specimen and thus hydrogen embrittlement during sliding,which is absent at anodic potentials [23].

Under open circuit condition, the open circuit potential (OCP) can berecorded continuously during the test, see Fig. 5. As expected, slidingresults in a cathodic shift of OCP, which is not stable during the slidingperiod. Rather unexpectedly, the OCP increases gradually with increasingsliding time. It is well known that OCP is related to the electrochemicalstate of a material. The OCP recorded during sliding is a mixed potentialreflecting the active state of the material inside the wear track and thepassive state of the unworn area outside the wear track [27]. Thus, OCPshould decrease with increasing wear track area, i.e. increasing slidingtime, as observed by many investigators [19,21,25]. The unexpectedincrease of OCP with sliding time in the present tests could be due to thecovering of the wear track with corrosion and/or wear products, whichwould reduce the active wear track area.

At anodic potentials, a significant increase in current is observedduring sliding for both specimens, as can be seen from the recordedcurrent transients shown in Fig. 6. Corrosion is therefore accelerated bythe sliding action, a phenomenon observed in passive systems bymanyinvestigators [17,19,25,31]. Under sliding action, the oxide film on realcontact areas is either removed or damaged after a single cycle or acertain number of cycles of sliding, leading to accelerated electrochem-ical corrosion and contributing to wear-accelerated corrosion.

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Fig. 8. Total wear volume (a) and mean coefficient of friction (COF) (b) as a function ofapplied potential. Test condition: 1 M H2SO4, 20 N, 60 rpm, 3800 s. CT-carburizedspecimen.

4284 Y. Sun, E. Haruman / Surface & Coatings Technology 205 (2011) 4280–4290

From Fig. 6, it can be seen that in the case of the untreatedspecimen, sliding leads to a large increase in anodic currents whichfluctuates between 0.4 mA and 1.2 mA at the potentials of 500 mV(SCE) and 900 mV(SCE). However, at the same anodic potentials, thecarburized specimen experiences a much lower current increaseduring sliding, which is more stable with a value around 0.2 mA overthe whole sliding period. This demonstrates the benefit of thecarburizing treatment in reducing wear-induced corrosion, in agree-ment with the potentiodynamic polarization curves shown in Fig. 3.

3.2.2. Effect of potential on total wear volume and coefficient of frictionFig. 7 shows typical surface profiles measured across the wear

tracks produced at cathodic, open circuit and anodic potentials. Thewear volumes estimated from these profiles are summarized inTable 1 and plotted in Fig. 8a as a function of applied potential. For theuntreated specimen, the wear track surface is relatively smooth atcathodic and open circuit potentials (Fig. 7a and b), but becomessignificantly roughened at anodic potentials (Fig. 7c and d). This isaccompanied by a significant increase in wear volume at anodicpotentials by up to more than 10 times (Fig. 8a). The wear volumesmeasured at the two cathodic potentials are similar (Table 1),suggesting that in the test solution, hydrogen embrittlement is not akey issue of concern for the untreated specimen, otherwise the wearvolume at−900 mV(SE) should be larger than that at−500 mV(SCE)because more hydrogen charging is expected at −900 mV(SCE). Inaddition, cathodic protection against tribocorrosion is effective for theuntreated specimen in the test solution, as the wear volumemeasuredat cathodic potentials is the smallest among all potentials tested(Fig. 8a).

The response of the carburized specimen to applied potentialduring the sliding process is quite different. Firstly, the wear tracksurface of the carburized specimen is extremely rough at cathodicand open circuit potentials, and becomes very smooth at anodic

potentials, see Fig. 7. This is in contrast to the observation on theuntreated specimen discussed above, also see Fig. 7. Secondly, thewear volumes produced at the two cathodic potentials are onlymarginally smaller than those resulting from the untreated specimen(Fig. 8a and Table 1). This is rather surprising because the hardcarburized layer does not show much benefit in reducing wear atcathodic potentials in the test solution, a phenomenon that has notbeen reported before. In addition, the wear volume at open circuitpotential is similar to those at cathodic potentials for the carburizedspecimen, suggesting that cathodic protection is not effective for thecarburized specimen in the test solution. As comparedwith the resultobtained for the untreated specimen (Table 1), the carburized layerreduces the wear volume by 40% at open circuit potential. Such adegree of improvement in wear resistance by carburizing is muchless than that observed under dry sliding and rolling-slidingconditions [12,13]. However, the beneficial effect of the carburizedlayer becomes very significant at anodic potentials, as can be seenfrom Fig. 8a and Table 1. In addition, the wear volume of thecarburized specimen is not significantly affected by applied poten-tial: the wear volume is increased only by 70% when the potential isincreased from the cathodic region to the anodic region as high as900 mV(SCE). This behavior is different from that of the untreatedspecimen which sees an increase in wear volume by more than 10times (Fig. 8a). Thus, at anodic potentials, the tribocorrosionresistance of the 316L stainless steel is improved by up to morethan 10 times by the low temperature carburizing treatment.

In Fig. 8b, the measured average coefficient of friction is plottedagainst applied potential for both specimens. The variation ofcoefficient of friction with potential is similar to that recorded duringpotentiodynamic sliding shown in Fig. 4 In the cathodic region, at opencircuit and in the low-potential anodic region, the two specimensexhibit similar coefficients of friction. In the high-potential anodicregion, the carburized specimen exhibits higher coefficient of friction

Fig. 9.Microscopic images showing the surface morphology of the wear tracks produced on the untreated specimens at various applied potentials indicated. Sliding condition: sameas Fig. 8.

4285Y. Sun, E. Haruman / Surface & Coatings Technology 205 (2011) 4280–4290

than the untreated specimen. This may suggest that the corrosionproducts produced from the untreated specimen in the high-potentialanodic region have a lubricating ability to reduce friction; while theincrease in coefficient of friction of the carburized specimen in the high-potential anodic region may be explained by the smoothness of thewear track, which can result in an increased real contact area to givehigher friction due to increased adhesion junctions at real contact areas.Such an explanation is not conclusive because inmany situations roughsurfaces result in higher coefficient of friction due to increasedmechanical interlocking.

3.2.3. Microscopic examination of wear tracks and scarsFig. 9 shows microscopic images of selected wear tracks produced

on the untreated specimen. At the cathodic potential of −500 mV(SCE), the wear track is free from corrosion products and is quitesmooth with many fine and a couple of deep abrasion marks, inagreement with the surface profile measured by profilometry(Fig. 9a). At open circuit, the wear track is also quite smooth butcovered with patches of dark corrosion products (Fig. 9b). At theanodic potentials, the wear tracks become much wider, are coveredwith patches of dark corrosion products and populated with manydeep abrasion marks, in consistency with the surface profilesmeasured by profilometry (Fig. 7c and d). Material pile-up at thewear track edges of the untreated specimen is also evident from Figs. 7and 9, which is due to severe plastic deformation inside and aroundthe wear track. Apart from some pit-like craters formed by materialdelamination, no corrosion pits are observed inside the wear track onthe untreated specimen at the applied anodic potentials. In order tohave a better understanding of the wear mechanisms involved, thewear scars on the alumina ball slider are also examined. As shown inFig. 10a, at the cathodic potential of−500 mV(SCE), the wear scar onthe alumina ball sliding against the untreated specimen is relativelysmooth with some fine wear particles, which are obviously respon-sible for the fine abrasion marks observed on the wear track (Fig. 9a).Similar wear scar morphologies are observed on the alumina ball at−900 mV(SCE) and at open circuit. However, at anodic potentials, thewear scar on the alumina ball sliding against the untreated specimenis covered with many wear debris particles, which form a narrow andelongated band (Fig. 10b). This agrees well with the wide and roughwear tracks produced on the untreated specimen at anodic potentials(Fig. 9d).

The morphology of the wear tracks produced on the carburizedspecimen is shown in Fig. 11. No material pile-up is observed at thewear track edges, obviously due to the high hardness of thecarburized layer to resist severe plastic deformation. In agreementwith surface profile measurements (Fig. 7), the wear tracks producedat cathodic (Fig. 11a) and open circuit (Fig. 11b) potentials areextremely rough with many deep and wide abrasion marks. On theother hand, the wear tracks produced at anodic potentials are verysmooth with fine polishing marks (Fig. 11c and d). Only a few deeperabrasion marks are observed on the wear track produced at 500 mV(SCE) (Fig. 11d). It thus seems that severe abrasive wear occurs atcathodic and open circuit potentials, while micro-polishing is theprincipal mechanism involved at anodic potentials. These wearmechanisms are further confirmed by microscopic examination ofthe wear scars on the alumina ball sliding against the carburizedspecimen (Fig. 12). At the cathodic potential of−500 mV(SCE), a flatarea is produced on the alumina ball and many wear debris particleare deposited onto this worn area (Fig. 12a). Furthermore, materialsdelaminated from the carburized layer are also transferred to theedges of the wear scar, and these transferred materials seem to haveinvolved in the subsequent abrasivewear process as evident from thescratching marks on the surface of the transferred material (arrowedin Fig. 12a). Indeed, detailed examination of the wear track showsevidence of delamination of material from the carburized layer, asshown in Fig. 13. Similar wear scar morphologies on the alumina ball

are observed at −900 mV(SCE) and at open circuit potential. Thus, theineffectiveness of the hard carburized layer in improving wear resistanceat cathodic and open circuit potentials can be explained by theaccumulation of wear debris particles and the transfer of delaminatedmaterials onto the aluminaball slider. These thirdbodies can serve ashardabrasives to wear the remaining carburized layer. On the other hand, atanodic potentials, where a thicker oxide film can be formed inside thewear track at contact intervals, the tendency of interaction between thecarburized surface and the alumina ball can be significantly reduced oreven completely prevented, such that wear takes place in the micro-polishingmode by the smooth alumina slider, resulting in a polished-likewear track surface. Indeed, as shown inFig. 12b,nowearflat canbe seen inthewear scaron the aluminaball at 500 mV(SCE) and the accumulationofwear debris is much less than that observed at cathodic potentials. Theminor accumulation of wear debris on the ball wear scar is obviouslyresponsible for the fewdeeperabrasionmarkson thewear trackproducedat 500 mV(SCE) shown in Fig. 11d.

4. Discussion

The experimental results presented above demonstrate that in the1 M H2SO4 solution, the carburized layer offers significant improve-ment in tribocorrosion resistance at anodic potentials, but is noteffective in reducing wear at cathodic potentials. The improvement intribocorrosion resistance at open circuit by carburizing is not assignificant as at anodic potentials. These results were quite surprising,but when the wear tracks on the test specimens and wear scars on thealumina ball were examined, a clear picture emerges. It is the behaviorof the third bodies, i.e. wear particles and delaminated materials that

Fig. 10. Microscopic images showing the surface morphology of the wear scars on thealumina ball slider after sliding tests on the untreated specimen at (a) −500 mV(SCE)and (b) 500 mV(SCE). Note the accumulation of wear debris on the ball surface at theanodic potential of 500 mV(SCE).

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have played a crucial role in determining the wear behavior of thecarburized specimen at cathodic and open circuit potentials. Thesethird bodies resulting from the carburized layer tend to accumulate onthe alumina ball wear scar at cathodic and open circuit potentials andserve as abrasives to accelerate mechanical wear, which is responsiblefor the significant roughening of the wear track surface. It should bepointed out that the behavior of the third body depends very much onthe corrosion environment and test configuration [23,32]. In a parallelstudy in NaCl solution [33], it is found that the carburized layer iseffective in improving wear resistance of 316L steel at cathodicpotentials.

The results summarized in Table 1 and Fig. 8a clearly show that thereare synergistic effects between corrosion and wear, which result incorrosion-induced wear and wear-induced corrosion [17]. It is impor-tant to identify the contribution of corrosion and wear to materialremoval by tribocorrosion in order to minimize material degradation.Efforts have been made by several investigators [22,27,31,32] toquantify such synergistic effects. In the simple mechanistic approach[23], the total material loss from the wear track (VT) is the result ofmechanical wear (Vmech) and chemical wear (Vchem), i.e.

VT = Vmech + Vchem ð1Þ

The chemical wear component comprises material losses inassociation with the evolution of anodic currents, i.e. material loss dueto metal dissolution and material loss from the oxide film (formed atcontact intervals) by the sliding action; while the mechanical wearproduces metallic wear particles from the underlying substrate.Although Vmech and Vchem are mechanical and chemical mechanismsrespectively, they may influence each other by the interplay ofmechanical, chemical and materials factors. On the one hand, corrosionin a passivemetal is usually accelerated bywear because thewear track

area is depassivated by thewearing action. On the other hand, corrosiondoes not necessarily always accelerate mechanical wear, because theoxidefilmandcorrosionproductsmaymodify friction and contact stressfields responsible for mechanical wear [23].

In Eq. (1), the chemical component (Vchem) can be obtainedthrough the integration of the current measured during sliding underpotentiostatic conditions. In a passive system without pitting, thecurrent measured during sliding flows mainly through the wear track[34], and the volume loss of material due to this wear-inducedcorrosion is given by Faraday's law [24]:

Vchem =ItMnFρ

ð2Þ

where t is sliding time, I is the current registered during the slidingperiod, F is the Faraday's constant (96,458 C mol−1), ρ is the densityof the material and M is the atomic mass of the alloy given by∑XiMi,where Xi is the mole fraction and Mi the atomic mass of the alloyconstituents (Fe, Cr and Ni). The valence of dissolution and oxidation,n, depends on the specific anodic reactions involved. For stainlesssteels in the active region, both Fe and Cr dissolve with a valence of 2and in the passive region they form oxides with a valence of 3, whichalso applies to passive film formation [35].

Since no pitting corrosion is observed inside thewear tracks on boththe carburized and untreated specimens in 1 M H2SO4 solution, it ispossible to calculate the contribution of the chemical component(Vchem) using Eq. (2), and then to derive the mechanical component(Vmech) using Eq. (1), where VT is evaluated from the surface profileacross the wear track. The calculated value of the Vchem componentdepends on the valence value (n) used in the calculation. If n=2 is usedto account for active dissolution, an upper bound will be obtained forVchem; while in the case of n=3, a lower bound will be obtained for

Fig. 11. Microscopic images showing the surface morphology of the wear tracks produced on the carburized specimens at various applied potentials indicated. Sliding condition: sameas Fig. 8.

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Vchem. Since the exact valence value is not known for the testconditions [34], based on the observation that “pseudo-passivity” ismaintained during the sliding process (see Fig. 3), a valence value (n) of3 is used to account for oxide film formation and passive dissolution atcontact intervals. The registered average current at each potential(Table 1) during sliding is assumed to be wholly due to these tworeactions, ignoring the contribution of the third bodies (oxidation anddissolutionof third bodymetallic particles remaining in the contact zonecan also contribute to the measured current [22]). The results aresummarized in Table 1 and plotted in Fig. 14. As discussed previously, at

the cathodic potentials, the totalwear volume is due tomechanicalwearsince no electrochemical corrosion occurs (Vchem=0). At open circuitpotential, themechanical component (Vmech) of theuntreated specimenis not affected but there is a contribution of chemical wear (by 25%) tothe total material loss (Fig. 14a). For the carburized specimen at opencircuit, the mechanical wear component is slightly reduced andchemical wear contributes to about 10% of the total material loss(Fig. 14b). At anodic potentials, the untreated specimen experiences asignificant increase in both mechanical wear and chemical wear(Fig. 14a), with the mechanical wear component being predominant.Thus the untreated specimen suffers from wear-accelerated corrosionand corrosion-acceleratedwear: a typical situation in tribocorrosion. Onthe other hand, for the carburized specimen at anodic potentials, themechanicalwear component is significantly reduced and chemicalwearcontributes to most of the material loss from the wear track (Fig. 14b).Thus, the carburized specimen only suffers from wear-acceleratedcorrosion, but benefits from corrosion-decelerated wear.

According to the above analysis, the tribocorrosion mechanisms ofthe carburized specimen can be proposed as follows. At cathodicpotentials, since no oxide film is formed, the bare carburized surfacemakes direct contact with the alumina ball slider and the resultantwear debris particles stick onto the slider surface easily. As the weardebris particles are generated from the upper part of the carburizedlayer, they are harder than the remaining carburized layer in the weartrack. Thus the wear track is abraded by the transferred particles,

Fig. 12. Microscopic images showing the surface morphology of the wear scars on thealumina ball slider after sliding tests on the carburized specimen at (a) −500 mV(SCE)and (b) 500 mV(SCE). Note the transfer of delaminated material from the test specimento the ball surface (arrowed) at the cathodic potential of −500 mV(SCE).

Fig. 13. Microscopic image showing wear debris formation through delamination ofmaterial from the carburized layer during sliding at the cathodic potential of−500 mV.

0

0.1

0.2

0.3

0.4

0.5

Wea

r Tra

ck V

olum

e (m

m3 )

Wea

r Tra

ck V

olum

e (m

m3 )

-900 -500 OCP(-291) 50 500 900Potential (mV/SCE)

-900 -500 OCP(-291) 50 500 900Potential (mV/SCE)

Untreated 316L1 MH2SO4

20 N, 60 rpm

1 MH2SO420 N, 60 rpm

VmechVchem

VmechVchem

0

0.005

0.01

0.015

0.02

0.025

0.03

0.035

0.04

0.045Carburised 316L

(a)

(b)

Fig. 14. Contribution of mechanical wear and chemical wear components to total weartrack volume of the untreated (a) and carburized (b) specimens tested at variousapplied potentials. Test condition: same as Fig. 8.

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resulting in an extremely rough surface. At open circuit, the oxide filmformed is not thick enough to prevent the alumina slider from digginginto the carburized layer, such that mechanical wear is dominant.Since the slider has to overcome the thin oxide film before reachingthe substrate, either by wearing through the thin oxide film or causingit to flake off, the mechanical wear component is slightly reduced ascompared to the situation without an oxide film, i.e. at cathodicpotentials. The chemical wear comes from the repeated removal of thethin oxide film, whichwould re-grow at each contact interval of 1 s (ata rotation speed of 60 rpm). At anodic potentials, the oxide filmformed on the carburized surface at each contact interval becomesthicker. The repeated removal of the thicker oxide film by subsequentsliding cycles results in larger chemical wear, but the thicker oxidefilm, supported by the hard subsurface, can provide certain protectionto the carburized layer against the mechanical action of the slider,such that mechanical wear of the underlying carburized layer isreduced. Since wear is mainly confined to the oxide film and is limitedin the substrate, a smooth wear track surface results.

On the other hand, due to the low hardness of the untreatedspecimen and the high contact pressure, the alumina slider digsthrough the oxide film (formed at anodic potentials) into the softsubstrate easily to cause significant mechanical wear from thesubstrate. Due to the larger and rougher wear track area, repassivationof the completely depassivated wear track surface requires highercurrents and thus corrosion is much accelerated by wear. The weardebris particles accumulate onto the alumina slider (Fig. 10b), leadingto accelerated mechanical wear of the substrate and significantroughening of the wear track surface (Fig. 9d).

Clearly, at anodic potentials, the untreated specimen suffers fromnotonly severe mechanical wear, but also severe chemical wear. Althoughchemical wear is predominant in the carburized specimen, this chemicalwear component is much smaller than that from the untreatedspecimen. In Fig. 15, the chemicalwear components from the carburizedand untreated specimens at various potentials are compared. It can beseen that chemical wear from the untreated specimen is about threetimes larger than that from the carburized specimen. Such a largedifference in chemical wear between the untreated and carburizedspecimens canbeexplainedby the inherently better corrosion resistanceof the carburized layer (Fig. 3) and the much smaller wear track area onthe carburized specimen. As can be seen from Fig. 7c and d, the weartrack area (exposed to the solution) on the carburized specimen atanodic potentials is about one third of that on the untreated specimen.Thusmaterial loss by corrosion from the carburized specimen should beone third of that from theuntreated specimenat the same corrosion rate,in good agreement with the results in Fig. 15. The much enhanced

tribocorrosion resistance of the carburized layer at anodic potentials cantherefore be derived from its high hardness and good corrosionresistance such that both mechanical wear and chemical wear arereduced under the present testing conditions. It should be pointed outthat the principal tribocorrosion mechanisms of the carburizedspecimen in the H2SO4 solution are expected to depend on contactload and frequency, as in untreated stainless steels reported by otherinvestigators [34,36]. Further tribocorrosion tests are currently beingcarried out at various contact loads and frequencies and the results willbe reported in forthcoming publications.

5. Conclusions

The results of this work demonstrate that in the 1 M H2SO4solution, applied potential has a significant effect on the tribocorro-sion behavior of the carburized and untreated 316L stainless steel.Under the present sliding conditions (20 N contact load and 60 rpmrotation speed), carburizing is not effective in improving the wearresistance of 316L steel at cathodic potentials. The wear particles anddelaminated materials from the carburized layer tend to accumulateonto the alumina slider at cathodic and open circuit potentials,causing abrasive wear of the carburized layer and roughening of thewear track surface. However, at anodic potentials, the carburizingtreatment can significantly improve the tribocorrosion resistance of316L steel by up to 10 times, due to different tribocorrosionmechanisms. At anodic potentials, the untreated specimen suffersfrom wear-accelerated corrosion and corrosion-accelerated wear;while material removal from the carburized specimen is dominatedby chemical wear in the form of repeated removal and re-growth ofthe oxide film during the sliding process. Mechanical wear of theunderlying carburized layer is reduced at anodic potentials under thepresent testing conditions. The principal tribocorrosion mechanism ofthe carburized layer thus changes from pure mechanical wear atcathodic potentials, to mixed mechanical and chemical wear withmechanical dominance at open circuit, and then to mixed chemicaland mechanical wear with chemical dominance at anodic potentials.The much enhanced tribocorrosion resistance of the carburized layerat anodic potentials (and thus in more aggressive environments) isderived from its high hardness and good corrosion resistance suchthat both mechanical wear and chemical wear are reduced. Depend-ing on the anodic potential, low temperature carburizing can reducethe mechanical wear component by 30 to 55 times and the chemicalwear component by 3 times under the present testing conditions.

Acknowledgements

The authors would like to thank Dr. X.Y. Li of The University ofBirmingham for providing the plasma carburizing treatment. TheFaculty of Technology, De Montfort University, is also acknowledgedfor providing the research facilities and the technicians in theEngineering Laboratory for their help in sample preparation andequipment maintenance.

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0

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0.03

0.04

0.05

0.06

0.07

0.08

0.09

0.1

-400 -200 0 200 400 600 800 1000Potential (mV/SCE)

Vch

em (

mm

3 )

Carburised

Untreated

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Effect of electrochemical potential on tribocorrosion behavior of low temperature plasma carburized 316L stainless steel in 1M H2SO4 solutionIntroductionExperimentalLow temperature plasma carburizingTribocorrosion testing

ResultsPotentiodynamic sliding testsPotentiostatic sliding wear testEffect of potential on electrochemical behaviourEffect of potential on total wear volume and coefficient of frictionMicroscopic examination of wear tracks and scars

DiscussionConclusionsAcknowledgementsReferences