Mechanical effects on the corrosion resistance of ferritic stainless steels during microabrasion-corrosion

Mechanical effects on the corrosion resistance of ferritic stainless steels during microabrasion-corrosion

Wear 426–427 (2019) 1474–1481 Contents lists available at ScienceDirect Wear journal homepage: www.elsevier.com/locate/wear Mechanical effects on th...

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Wear 426–427 (2019) 1474–1481

Contents lists available at ScienceDirect

Wear journal homepage: www.elsevier.com/locate/wear

Mechanical effects on the corrosion resistance of ferritic stainless steels during microabrasion-corrosion

T

W.S. Labiaparia,b, M.A.N. Ardilaa, C. Binderc, H.L. Costaa,d, J.D.B. de Melloa,c a

Universidade Federal de Uberlândia, Uberlândia, Brazil Aperam South America, Brazil c Universidade Federal de Santa Catarina, Brazil d Universidade Federal do Rio Grande, Brazil b

A R T I C LE I N FO

A B S T R A C T

Keywords: Abrasion-corrosion Ferritic stainless steel Synergy H2SO4

Mechanical events affect the corrosion behaviour of metals in a wide range of industrial applications, but the synergy between corrosion and mechanical events do not allow them to be studied individually. For ferritic stainless steels in H2SO4 corrosive media, a negative synergy between microabrasion and corrosion has been reported, where low friction passive layers reduce wear rates. However, for the abrasion-corrosion of ferritic stainless steels, little attention has been paid to mechanical effects on their corrosion resistance. To investigate this, potentiodynamic curves were obtained for ferritic stainless steels during standard corrosion tests, which were compared to potentiodynamic curves obtained in two other conditions: i) aerated corrosion tests (nonstatic electrolyte); and ii) microabrasion-corrosion tests using silica + H2SO4 solution slurries. The specimens were ferritic stainless steels with different chemical compositions (11 wt%Cr with and without Ti stabilization; 16 wt%Cr with and without Nb stabilization, 17.5 wt%Cr with Nb + Ti stabilization) and, for comparison, one austenitic stainless steel (18 wt%Cr-8 wt%Ni) and one carbon steel (0.2 wt%C). The results showed that the use of aerated conditions increased passivation current, which was further increased for microabrasion-corrosion conditions. Moreover, although the potentiodynamic curves for all the stainless steel specimens presented a clear passive region under standard corrosion test conditions, only the austenitic stainless steel showed a classic passive region when mechanical effects were present. For the low Cr ferritic stainless steel under microabrasion conditions, the potentiodynamic curves showed a region where the current increased little with potential increased, but this was called a pseudopassivation, since current increased steadily, although little, with potential increase. For higher Cr contents, after the pseudopassive region, a real passive region was present at higher potentials. The potentials for achieving real passive regions were higher under microabrasion conditions than under non-static condition, evidencing a contribution of both mechanical effects (fluid flow and mechanical wear) on their corrosion resistance.

1. Introduction Corrosion and wear are degradation phenomena that are mostly studied separately as isolated systems. However, in many industrial applications, these phenomena can occur concomitantly. In this sense, the wear action needs to be broadly understood: The fluid dynamic response induced by the flow of the electrolyte acting on the cathodic reactions through the acceleration of the ionic diffusion and also by the depolarization of the reactions and the mechanical action of the wear that acts on the anodic dissolution reactions of the materials by the depassivation. When they happen together, their synergy is quite complex. In one hand, wear influences corrosion, for example by depassivation of plastically deformed asperities; on the other hand,

corrosion influences wear, for example due to the Roscoe effect [1], where oxide passive films lead to surface embrittlement. Therefore, in certain cases, the synergy is positive, i.e., the resulting material loss wear can be higher than the sum of the losses promoted by each phenomena. In this case, mechanical wear leads to the exposure of the metal, often removing protective passive layers, and therefore accelerating corrosion [2]. In addition, stressed structures tend to develop different electrochemical responses when compared with the nonstressed state [3]. In other cases, the synergy between corrosion and mechanical wear can be negative, so that their interaction results in lower material loss than the sum of their individual effects separately [4,5]. The complex mechanism of corrosion and wear is dependent on the

E-mail address: [email protected] (J.D.B. de Mello). https://doi.org/10.1016/j.wear.2018.12.057 Received 3 September 2018; Received in revised form 14 December 2018; Accepted 20 December 2018 0043-1648/ © 2018 Elsevier B.V. All rights reserved.

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microstructure and chemical composition of the surface of the materials, the solution pH, the electrochemical imposed conditions and, for the case of abrasive wear, the abrasive properties (size, type and concentration) [6]. The investigation of the synergy between mechanical wear and corrosion is particularly interesting for stainless steels, since they show a protective passive layer and great repassivation capacity when the passive layer is eventually removed. Studies can be found in the literature regarding the abrasion-corrosion behaviour of more traditional austenitic stainless steels (AISI 304 and 316) [7–9], and more recently duplex stainless steels (AISI 2205) [3,6,10]. Cheaper ferritic stainless steels have been less in focus, despite their proved potential in sectors such as the bioenergy sector, for example in sugar and ethanol plants [11]. On one hand, corrosion can lead to the formation of passive layers that affect abrasion resistance. When comparing the behaviour of 316L stainless steel in reciprocating sliding tests in 0.5 M H2SO4 solution under cathodic and passive potentials, some authors [8,9] have found higher wear under passive conditions. They suggested that the thin and hard passive film induced more near-surface deformation, promoting strain accumulation and therefore increasing wear when compared with cathodic potentials, where no passive film was present. Our group has investigated extensively the microabrasion corrosion of ferritic stainless steels. For that, Santos et al. [12] developed a new instrument, enabling the measurement of forces acting on the contact. The test rig consists of a fixed-ball microabrasion tester, where the specimen is supported by a three-axis load cell, which is linked to the electrochemical cell via a very flexible membrane. Microabrasion corrosion conditions (tests using 10%wtSiO2 in 1 N H2SO4 solution slurries under potentiodynamic conditions) were compared with purely mechanical damage occurring in conventional microabrasion tests, i.e., tests in distilled water without externally applied potentials [5]. It was found that friction coefficients under micro-abrasion conditions were two times lower than under pure microabrasion conditions. In addition, microabrasion wear coefficients of stainless steels were four times higher than the coefficients measured in the microabrasion-corrosion tests. It seemed that corrosion induced the formation of a low friction, protective passive layer that reduced wear under microabrasion-corrosion conditions. On the other hand, besides the effects of corrosion on mechanical wear, it is crucial to understand how rubbing affects the electrochemical variables during tribocorrosion, i.e., the effects of the mechanical component on the electrochemical component. For example, for microabrasion-corrosion of WC–10Co–4Cr sprayed coatings [13], it was found that when the samples were merely exposed to the corrosive slurry, allowing them to freely corrode without mechanical action, their potentiodynamic curves showed lower values of passive current density when compared with potentiodynamic curves obtained during microabrasion-corrosion tests. This was attributed to the damage of the passive layer. For ferritic stainless steels, current density increased under abrasivecorrosive environments when compared with abrasion tests using slurries with distilled water without externally applied potentials [4]. This behaviour seems attributable to a competition between removal of the passive layer and repassivation, but the phenomena involved need to be better investigated. This paper investigates further how the mechanical action influences the electrochemical variables by comparing potentiodynamic curves of ferritic stainless steels obtained during microabrasion corrosion tests with pure corrosion tests following standard procedures and corrosion tests under aerated conditions, where the electrolyte flow becomes non-static. The effects of Cr content and stabilization (with Ti and Nb) are analyzed.

Fig. 1. Scheme of the methodological approach.

2. Methodology Fig. 1 summarizes the methodological approach used in this paper. It shows that, for the different materials selected, potentiodynamic curves were obtained for three test conditions. The first involved pure corrosion tests following the standard ASTM G5-94 [14]. The second involved modified corrosion tests in a non-static condition and aerated media. The third involved micro-abrasion corrosion tests. Each test condition was repeated at least three times to ensure reproducibility of the results. The materials tested were six ferritic stainless steels with different Cr contents, with and without stabilization. The stabilizing elements were either Nb or Ti. For comparative purposes, one austenitic stainless steel (18Cr8Ni) and one carbon steel (A36) were also tested. The chemical composition of the specimens is presented in Table 1, which was assessed using various experimental techniques, depending on the element under analysis. Carbon (C) was detected by Infrared Absorption using a Leco CS444® equipment. Nitrogen (N) was analyzed by Thermal Conductivity analysis using a Leco equipment model TC436®. All other elements were analyzed by Optical Emission Spectrometry using an equipment model ThermoARL 4460. Their mechanical properties are presented in Table 2.

2.1. Standard corrosion tests First, potentiodynamic curves were obtained for the different materials following the procedures described in the standard ASTM G5-94 [14] using a potentiostat/galvanostat Biologic SP150. Platinum gauze was used as counter electrode and a saturated calomel electrode (SCE) as the reference electrode. The electrolyte was a 1 N H2SO4 solution in distilled water. All specimens were ground down to sandpaper 600 mesh 8 h before the tests and then ultrasonically cleaned in acetone for 15 min. The tests started by immersing the specimens in the electrolyte during 15 min for stabilization of the open circuit potential (OCP). Table 1 Chemical composition of the specimens.

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Element [wt%]

11Cr

11CrTi

16Cr

16CrNb

17CrTiNb

C Mn Si P S Cr Ni Mo Nb Ti N

0.011 0.009 0.049 0.025 0.011 0.61 0.13 0.32 0.21 0.16 0.49 0.52 0.36 0.46 0.36 0.0247 0.0182 0.0395 0.0358 0.0219 0.0002 0.0005 0.0015 0.0013 0.0024 11.23 11.29 16.10 16.19 17.62 0.31 0.12 0.27 0.19 0.17 0.021 0.005 0.020 0.035 1.75 0.006 0.002 0.014 0.416 0.183 0.003 0.144 0.003 0.004 0.177 0.0145 0.0087 0.0528 0.0202 0.0112

18Cr8Ni

A36

0.055 1.15 0.42 0.0251 0.0009 18.28 8.01 0.063 0.005 0.001 0.0421

0.138 1.06 0.01 0.0154 0.0075 0.01 0.01 0.003 0.001 0.001 0.0026

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details in [18], evidencing slightly angular particles with a mean diameter of 3.4 µm, with 85% of the values between 1 and 10 µm. The preparation of the specimens followed the same procedures described in the previous sub-sections. For the tests, initially the specimen was placed in the electrochemical cell, partially submerged in the 1 N H2SO4 electrolyte. After 15 min for stabilization of the OCP, the normal load was applied (1.42 N), the motor was activated to start the motion of the ball (150 rpm), and the slurry was pumped over the contact between the ball and the specimen at a flow of 1.70 ml/min. In addition, the potentiostat started scanning the potentiodynamic curve from 100 mV below the OCP until the transpassive region. When the scanning rate of 0.167 mV/s specified by ASTM G5-94 [14] was used in the standard corrosion tests, the time necessary to scan the whole polarization curve was around 4 h. However, when the test is very long, the damage caused in the zirconia balls is excessive and it is no longer possible to ensure that the ball roughness remains in the range 0.32 µm < Ra < 0.38 µm. Various scanning rates were tested [18] and it was found that for 1 mV/s the total scanning time was approximately 42 min and the potentiodynamic curves obtained were very similar to those obtained with 0.167 mV/s. Previous pure microabrasion tests of stainless steels [5] showed that a steady-state wear regime with regular craters was reached after around 15 min. Therefore, it was decided that a scanning rate of 1 mV/s would be adequate both in terms of the potentiodynamic curves and of forming regular wear craters. A scanning rate of 1 mV/s has already been used in other works for other materials [13,19–21] and even a higher value of 2 mV/s for stainless steels [7]. After the tests, the surfaces were analyzed by scanning electron microscopy (SEM) and electron backscatter diffraction (EBSD). For the EBSD analysis, transversal cross sections were prepared across the tested regions, as depicted in Fig. 3. Then, the cross sections were polished with sandpaper (down to grit size 600 mesh), then polished with diamond paste (diameters of 9, 3 and 1 µm) and finally polished with colloidal silica. The EBSD analyzer was coupled to a field emission gun (FEG) SEM, model Quanta 250 FEG. The voltage was 20 kV and the step was 4 µm, along the specimen thickness. The software OIM Analysis® was used for the analysis of the EBSD measurements.

Table 2 Mechanical properties of the specimens; YS 0.2 = Yield Strength at 0.2% deformation; TS = Tensile Strength; El = Elongation; HV10 = Vickers hardness under 10 kgf normal load. Material

YS 0.2% [MPa]

TS [MPa]

11Cr 11CrTi 16Cr 16CrNb 17CrTiNb 18Cr8Ni A36

323.8 316.3 343.1 336.2 361.3 347.3 305.3

411.6 400.1 488.3 446.8 483.2 706.2 444.8

± ± ± ± ± ± ±

5.7 4.1 3.9 3.3 1.4 9.4 4.2

± ± ± ± ± ± ±

El [%] 2.0 1.1 2.9 1.7 0.8 7.1 3.2

37.7 40.4 31.3 35.9 36.2 62.3 34.1

± ± ± ± ± ± ±

HV10 [MPa] 0.7 1.9 1.8 1.7 0.2 1.2 1.0

1510 1435 1630 1514 1734 1992 1358

± ± ± ± ± ± ±

13 11 29 15 15 9 11

Then, the potential was scanned from 100 mV below the OCP (cathodic region) until the transpassive region was reached. Potential scanning used a scanning rate of 0.167 mV/s as specified in ASTM G5-94 [14]. 2.1.1. Corrosion tests under non-static, aerated conditions The objective of these tests was the separate the effects of the electrolyte flow that occurs in an abrasion-corrosion test from the mechanical effects caused by abrasion. The procedures for preparation of the specimens were the same as used in the previous tests. The same potentiostat/galvanostat was used, but it was necessary to increase the potential scanning rate from 0.167 mV/s (standard test) to 1 mV/s, because the subsequent microabrasion-corrosion tests required the use of higher potential scanning rates, as explained in the next sub-section. Since the only difference aimed between these tests and the microabrasion-corrosion tests was the occurrence of abrasion, the scanning potential rate was therefore increased to 1 mV/s. The corrosion area (area unprotected by the insulating resin) of 2 cm2 was partially submerged in the electrolyte, as shown in Fig. 2. A rubber tube directed the electrolyte at a pumping rate of 1.70 ml/min towards the corrosion area in the specimen that remained above the fluid level. 2.1.2. Microabrasion-corrosion tests The microabrasion-corrosion tests used a specially designed test rig [12]. A 3D load cell attached to the electrochemical cell via a very flexible membrane allowed precisely applying and monitoring the forces during the tests. Zirconia balls with diameter ϕ = 25.4 mm were used and their roughness was controlled (0.32 µm < Ra < 0.38 µm). This was necessary because the surface roughness of the ball influences particle entrainment into the contact and therefore the test results, as reported by some authors [15,16] and quantified by Costa et al. [17]. The slurry used in the tests was composed of SiO2 particles dispersed in 1 N H2SO4 solution. The abrasive silica was fully characterized, see

3. Results and discussion The focus of the present work is to analyze how the mechanical effects involved in microabrasion-corrosion tests affect the electrochemical behaviour of ferritic stainless steels. The effects of corrosion on abrasion have already been investigated in a previous work [5]. The potentiodynamic curves obtained for the three different test conditions are exemplified in Fig. 4. For each condition and material tested, only

Fig. 2. Aerated tests: (a) scheme showing the partially submerged specimen with a suitable tube for directing the electrolyte; and (b) real condition. 1476

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tests. The current density remained constant during the passive plateau, but again agitation of the fluid increased the passivation current and agitation + rubbing increased it further. The values of passivation current are summarized in Fig. 6. For the lower Cr content (11%), the passivation currents in fact referred to a pseudo-passive behaviour. First, Fig. 6 shows that, as expected, an increase in the amount of Cr reduced the passivation current, for all the conditions tested. The increase in chromium content tends to accentuate the chemical adsorption through covalent coordinate bonds between the chrome atoms and sulphur (S) atoms in the group of SO4-2. This fact is in agreement with molecular dynamics simulations by Diawara et al. [27], where the stability of the passive film grows from with Cr content, reaching its maximum at 20% Cr. It is also notorious the increase in passivation current density Ip with the increase of the intensity of the mechanical event on the surface, first by agitation and then by abrasion. Fig. 7 summarizes the values of passive potentials for the ferritic stainless steels tested. Under standard corrosion conditions, the passivation potential decreased with the amount of Cr, evidencing a stronger tendency for passivation and therefore an increased corrosion resistance. Under non-standard conditions, the specimens first presented a pseudo-passive behaviour and the values presented in Fig. 7 refer therefore to the pseudo-passivation potentials. This figure shows that when mechanical effects were added to the tests, no significant change in the passivation potential was observed. Although the passivation current density (Ip) was strongly affected by the mechanical effects in the tests, the potential at which passivation started (Ep) did not change. In addition, under microabrasion-corrosion conditions, no correlation was found between the Cr content and the passivation potential Ep, which was evident for the other conditions. On the other hand, it had been observed from Fig. 4 that the amount of Cr influenced the potential for secondary passivation, which reduced as the amount of Cr increased. Fig. 8 SEM of the worn craters after microabrasion-corrosion tests, images on the left under lower magnification and images on the right under higher magnification: (a) and (b) 11Cr; (c) and (d) 11CrTi; (e) and (f) 16Cr; (g) and (h) 18Cr8Ni. Fig. 8 exemplifies some of the worn craters obtained after the microabrasion-corrosion tests. They are secondary electron (SE) SEM images obtained at the centre of the craters. The images on the left were obtained under lower magnification and the images on the right under higher magnification. They evidence the abrasive action of the abrasive articles, responsible for removing the passive films. However, the images show different wear mechanisms depending on the Cr content in the stainless steels. For the lowest Cr content (11Cr, Fig. 8.a and Fig. 8.b), wear is due mainly to multiple indentations on the surface. According to Trezona et al. [28], this mechanism is the result of abrasive particles indenting the ball very lightly, so that they can roll in the contact, producing multiple indentations in the specimen. This occurs, for example, when the number of abrasive particles is large and therefore the load per abrasive is low, so that they do not get embedded in the ball. The same mechanism was observed for the specimen stabilized with Ti (11CrTi, Fig. 8.c and Fig. 8.d), showing that the presence of stabilizing elements did not influence the wear mechanism. On the other hand, the Cr content had a significant effect on the wear mechanisms. For 16 wt%Cr (Fig. 8.e and Fig. 8.f), grooves start to appear on the specimen surface, showing a mixed mechanism, with grooves, but also multiple indentations between the grooves (Fig. 8.f). This suggests that fewer abrasive particles entrained into the contact, so that the load per abrasive particle increased. As largely observed in the literature [17,28,29], higher load per abrasive particle induces grooving (sliding of particles) in microabrasion tests instead of multiple indentations (rolling of abrasive particles). The potentiodynamic curves obtained during the microabrasion tests showed that for low Cr contents only a pseudo-passive behaviour was observed, showing that the passive film was not easily repassivated

Fig. 3. Scheme for the preparation of the specimens for EBSD analysis.

one typical curve is presented, but all repetitions showed very reproducible potentiodynamic curves. Two exceptions were the specimens with 16% Cr (16Cr and 16CrNb) when tested under microabrasion-corrosion conditions. The different repetitions for those tests resulted in different curves, as shown in Fig. 5. For comparison purposes, the curves of the three repetitions for the specimens with 11% Cr are also presented, where the different repetitions indeed show very reproducible behaviour. All the stainless steel specimens showed regions of cathodic behaviour, anodic behaviour, passive behaviour and transpassive behaviour. As expected, the carbon steel (A36) did not show a clear passive behaviour (Fig. 4.g). Comparing the standard corrosion tests with the corrosion tests using aerated conditions, it is possible to observe a substantial increase in the passivation current density (Ip) for all the materials tested. This suggests that the mechanical effect of the electrolyte flow generated by pumping the solution rendered passivation more difficult. This effect of aerated environments was observed for the corrosion of stainless steels in NaCl solutions by Qiao et al. [22], and by Le Bozec et al. [23] in saturated solution of oxygen. The latter [23] found that under conditions of oxygen saturation, the anodic and cathodic reactions were accentuated. There is an increase in mass transport by oxygen, accelerating the corrosive process [24]. For the microabrasion-corrosion tests, Ip was further increased, attributable to the abrasion component, which removes the passive film. On the other hand, additional interesting features could be observed for the polarization curves obtained under non-standard conditions. For steels with approximately 11% Cr (11Cr and 11CrTi, independently of stabilization), the potentiodynamic curves obtained in the microabrasion-corrosion condition did not show a regular passivation current density plateau, but instead Ip increased slightly but steadily with the potential (Fig. 4.a and Fig. 4.b). Various authors [6,25,26] discussed this increase in passivation current density Ip as a result of the competition between the removal of the passive layer and repassivation. In fact, since it is not an effective passivation, it has been referred to as pseudo-passivation. When the amount of Cr in the stainless steel increased to 16% (16Cr and 16CrNb, Fig. 4.c and Fig. 4.d), the aerated tests presented a pseudo passivation region for lower potentials (between around − 400 mV and + 550 mV). After that, a secondary passivation led the current density to stabilize at lower values. Under microabrasion-corrosion conditions, in the pseudo-passive region the current density again increased slightly with the potential. The secondary passivation still occurred, but at higher potentials. In fact, the different repetitions for the microabrasion-corrosion tests of the specimens with 16 wt%Cr (Fig. 5.c and d) showed a great variation in the potential for secondary passivation, suggesting an unstable behaviour. Increasing further the amount of Cr to 17% (Fig. 4.e), the pseudopassive region and the secondary passivation region were still present for the aerated and microabrasion-corrosion conditions, but the potential for secondary passivation reduced, i.e., the mechanical effects influenced less the electrochemical behaviour of the material. Aerated conditions increased the passivation current, which was further increased under microabrasion-corrosion conditions. The reference austenitic stainless steel (Fig. 4.f) with 18 wt%Cr and 8 wt%Ni did not present a pseudo-passive region for the non-standard 1477

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Fig. 4. Polarization curves obtained for the three test conditions: i) standard tests based on ASTM G5-94 [14], ii) aerated and non-static condition corrosion conditions, and iii) microabrasion-corrosion tests: (a) 11Cr; (b) 11CrTi; (c) 16Cr; (d) 16CrNb; (e) 17CrTiNb; (f) 18Cr8Ni; (g) A36.

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Fig. 5. Different repetitions for the polarization curves obtained in the microabrasion-corrosion tests: (a) 11Cr, (b) 11CrTi, (c) 16Cr, (d) 16CrNb.

Fig. 6. Results of passivation current density for the stainless steels under the three conditions. Fig. 7. Summary of passivation potentials.

after removal. As the Cr content increased, a passive behaviour was observed at higher potentials, showing that the film was more easily repassivated. Therefore, it seems that for higher chromium contents, a more “stable” passive film exists during microabrasion. This film was previously shown to present low friction [5]. It is possible that the presence of more stable low friction passive films at higher Cr contents lower make particle entrainment more difficult. This could increase the load per particle, promoting a transition from multiple indentations (rolling of particles) to a mixed regime (sliding + rolling of particles). The effect of friction on particle entrainment has been recognized since the test was first proposed [30], where the authors proposed a simple model for the conditions for particle entrainment involving the friction

coefficients between ball, abrasive and specimen. This hypothesis is corroborated by the analysis of the worn surfaces of the austenitic stainless steel (Fig. 8.g and Fig. 8.h), which showed predominance of grooving and only very few indentations. Its potentiodynamic curve did not show a pseudo-passive behaviour, but only a passive behaviour, evidencing a more stable passive film, and probably further difficulty for particle entrainment. Interestingly, EBSD maps of the worn surfaces (Fig. 9) show that under pure microabrasion conditions, deformation of the grains is observed near the surface. Under microabrasion-corrosion conditions, the deformation of grains at the surface is negligible, corroborating the 1479

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Fig. 8. SEM of the worn craters after microabrasion-corrosion tests, images on the left under lower magnification and images on the right under higher magnification: (a) and (b) 11Cr; (c) and (d) 11CrTi; (e) and (f) 16Cr; (g) and (h) 18Cr8Ni.

The abrasion effect that removes the passive layer contributes to a further increase in the current density. For higher Cr contents, after the pseudopassive region, a real passive region was present at higher potentials. The potentials for achieving real passive regions were higher under microabrasion conditions than under non-static condition, evidencing a contribution of both mechanical effects (fluid flow and mechanical wear) on their corrosion resistance. Passivation current (Ip) showed to be a suitable parameter to evaluate corrosion resistance during microabrasion-corrosion tests, but not the passivation potential (Ep).

lower friction values under microabrasion-corrosion conditions when compared with pure abrasion conditions published previously [5]. This lower friction made particle entrainment more difficult under microabrasion-corrosion than under pure microabrasion, as Fig. 8 had evidenced. 4. Conclusions This work showed that for stainless steels the agitation of the fluid that occurs during microabrasion tests contributes to increase the current density, suggesting that passivation becomes more difficult. 1480

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Fig. 9. EBSD maps of the specimen 11Cr after the tests; the upper part of the maps correspond to the worn surfaces: (a) pure microabrasion; (b) microabrasion-corrosion.

For ferritic stainless steels, it is recommended to analyze the secondary passivation phenomenon that occurs when mechanical effects (fluid flow and mechanical wear) take place, since it captures well their electrochemical behaviour during microabrasion-corrosion tests. Acknowledgements The authors would like to thank Coordination for the Improvement of Higher Education Personnel (CAPES (Proex)), Brazil, National Council for Scientific and Technological Development (CNPq), Brazil, (Grant 477286/2011-9), Companhia Brasileira de Metalurgia e Mineração (CBMM S.A.), Brazil, (Grant FEMEC01-2011) and Foundation for Research of the State of Minas Gerais (Fapemig), Brazil, for financial support. References [1] K.L. Zum Gahr, Microstructure and Wear of Materials, Elsevier Science, New York, USA, 1987. [2] D. Landolt, S. Mischler, M. Stemp, S. Barril, Third body effects and material fluxes in tribocorrosion systems involving a sliding contact, Wear 256 (2004) 517–524. [3] V. Vignal, N. Mary, P. Ponthiaux, F. Wenger, Influence of friction on the local mechanical and electrochemical behaviour of duplex stainless steels, Wear 261 (2006) 947–953. [4] J.O. Bello, R.J.K. Wood, J.A. Wharton, Synergistic effects of micro-abrasion–corrosion of UNS S30403, S31603 and S32760 stainless steels, Wear 263 (2007) 149–159.

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