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Corrosion behavior of cold-rolled and post heat-treated 316L stainless steel in 0.9wt% NaCl solution

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  • Corresponding author:

    K.M. Deen    E-mail: kmdeen.ceet@pu.edu.pk;deen@mail.ubc.ca

  • Received: 14 February 2020Revised: 31 March 2020Accepted: 1 April 2020Available online: 3 April 2020
  • The effect of cold rolling and post-rolling heat treatment on the microstructural and electrochemical properties of the 316L stainless steel was investigated. Two-pass and four-pass cold-rolled stainless steel specimens were heat-treated by annealing at 900°C followed by quenching in water. During the cold rolling, the microstructure of the as-received specimen transformed from austenite to strain-induced α′-martensite due to significant plastic deformation that also resulted in significant grain elongation (i.e., ~33% and 223% increases in the grain elongation after two and four rolling passes, respectively). The hardness of the heat-treated as-received specimen decreased from HV 190 to 146 due to the recovery and recrystallization of the austenite grain structure. The cyclic polarization scans of the as-rolled and heat-treated specimens were obtained in 0.9wt% NaCl solution. The pitting potential of the four-pass rolled specimen was significantly increased from 322.3 to 930.5 mV after post-rolling heat treatment. The beneficial effect of the heat treatment process was evident from ~10-times-lower corrosion current density and two-orders-of-magnitude-lower passive current density of the heat-treated specimens compared with those of the as-rolled specimens. Similarly, appreciably lower corrosion rate (3.302 × 10−4 mm/a) and higher pitting resistance (1115.5 mV) were exhibited by the post-rolled heat-treated specimens compared with the as-rolled 316L stainless steel specimens.
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Corrosion behavior of cold-rolled and post heat-treated 316L stainless steel in 0.9wt% NaCl solution

  • Corresponding author:

    K.M. Deen    E-mail: kmdeen.ceet@pu.edu.pk;deen@mail.ubc.ca

  • 1. Corrosion Control Research Cell, Department of Metallurgy and Materials Engineering, University of the Punjab, Lahore, Pakistan
  • 2. Department of Materials Engineering, The University of British Columbia, Vancouver, V6T 1Z4, BC, Canada

Abstract: The effect of cold rolling and post-rolling heat treatment on the microstructural and electrochemical properties of the 316L stainless steel was investigated. Two-pass and four-pass cold-rolled stainless steel specimens were heat-treated by annealing at 900°C followed by quenching in water. During the cold rolling, the microstructure of the as-received specimen transformed from austenite to strain-induced α′-martensite due to significant plastic deformation that also resulted in significant grain elongation (i.e., ~33% and 223% increases in the grain elongation after two and four rolling passes, respectively). The hardness of the heat-treated as-received specimen decreased from HV 190 to 146 due to the recovery and recrystallization of the austenite grain structure. The cyclic polarization scans of the as-rolled and heat-treated specimens were obtained in 0.9wt% NaCl solution. The pitting potential of the four-pass rolled specimen was significantly increased from 322.3 to 930.5 mV after post-rolling heat treatment. The beneficial effect of the heat treatment process was evident from ~10-times-lower corrosion current density and two-orders-of-magnitude-lower passive current density of the heat-treated specimens compared with those of the as-rolled specimens. Similarly, appreciably lower corrosion rate (3.302 × 10−4 mm/a) and higher pitting resistance (1115.5 mV) were exhibited by the post-rolled heat-treated specimens compared with the as-rolled 316L stainless steel specimens.

    • Stainless steel is considered one of the most important classes of engineering materials nowadays, as it is widely used in surgical industries because of its good corrosion resistance and mechanical properties. Austenitic stainless steels are mostly utilized for biomedical applications such as implants, fixtures, pharmaceuticals, and surgical instruments. The extensive use of 316L stainless steel can be ascribed to its biocompatibility, affordable cost, easy fabrication, good mechanical strength, and good corrosion resistance [12]. High corrosive resistance and toughness remain the key prerequisite for the production of surgical instruments [3]. For instance, austenitic stainless steel is widely used in manufacturing spatulas and scissors because of its nonmagnetic behavior and better corrosion resistance. The hardenability of such stainless steel is achieved not by heat treatment but via cold working. Heat treatment of stainless steel can cause Cr and C depletion from the grain boundaries and may cause intergranular corrosion. However, the grain refinement of stainless steel could play a significant role in improving the corrosion resistance by decreasing the diffusion path for carbon and chromium across the fine grains. This condition eliminates the possibility of chromium depletion at the vicinity of grain boundaries, which would promote the formation of uniform passive oxide film on the surface as a result, thus restricting the dissolution of the 316L stainless steel [2,4]. Corrosion resistance is ascribed to the stability of the film generated on the surface. The outer part of the generated passive film is formed by iron oxides, while the inner part is formed by chromium oxides because of its low mobility, which is due to the decreased diffusion property of chromium compared with that of iron [57]. The growth of a protective film is facilitated by the decrease in the diffusion length for chromium atoms due to a reduction in grain size. The literature [812] reported that a fine austenite grain structure ensures the formation of a uniform chromium-enriched surface film. A fine microstructure due to rapid interface boundary diffusion can cause more uniform and compact film generation [7,13]. Stainless steel and Vitallium (CoCrMo alloy) produce similar results in the manufacturing of surgical equipment [14]. One of the advanced usages of stainless steel is found in biopharmaceutical companies. Austenitic stainless steels are not heat-treatable; therefore, mechanical properties can be increased only by work hardening [15]. Several methods of refining the grains of stainless steel are reported in the literature; these methods include nano-texturing, mechanical vibrations, electromagnetic forces, arc oscillation, and cold rolling [7,1617]. Cold working can help in the formation of strain-induced martensite from the austenite phase, where the level of formation of martensite is connected to the range of cold working [15,1819]. The mechanical martensitic modification takes place whenever an external field-induced strain is applied instead of thermal shock-induced stress. This martensite develops in two different states: the ε-martensite phase with a hexagonal close-packed (hcp) structure and the ferromagnetic α′-martensite phase with a body-centred cubic (bcc) structure. The α′-martensite phase is thermodynamically more substantial than the ε-martensite phase [15]. Moreover, with increasing the strength and hardness, cold working considerably influences the corrosion behavior of austenitic stainless steels [15,20]. Previous studies [18,21] found that, unlike stainless steel in an annealed state, cold-worked steel might be more liable or even more corrosion-resistant when exposed to a corrosive environment. 316L stainless steel is shielded by a thin oxide film, which is resistant to corrosion in aggressive environments. The breakdown of this passive film can cause localized corrosion, i.e., crevice/pitting corrosion. The capability of stainless steel to repassivate after the breakdown of the film determines its corrosion resistance [2224]. Some researchers highlighted that pitting potential is reduced by cold working, whereas others revealed that pitting potential is enhanced as a result of cold working [2526].

      This research aims to investigate the effects of plastic deformation induced during the cold rolling process and the influence of post-heat treatment (i.e., annealing and quenching) on the microstructure, hardness, and corrosion behavior of 316L stainless steel in the 0.9wt% NaCl solution. This information may be used in the selection of suitable parameters to process and/or to manufacture medical grade instruments from 316L stainless steel. In addition, the objective of this study is to highlight the corrosion-related issues of as-rolled 316L stainless steel surgical instruments that interact with the human body during surgical procedures or function as orthopedic implant material within the body. The 0.9wt% NaCl solution, which contains 0.009 kg of total salts dissolved in 0.991 kg of water, was used to simulate the total salinity of the human blood. The same physiological solution is most often used as a saline infusion for humans.

    2.   Experimental
    • A rectangular stainless steel 316L sheet with a thickness of 3.35 mm and dimensions of 0.61 m × 0.15 m was purchased from the local market. Before undergoing cold rolling, the sheets were annealed at a temperature of 900°C. The furnace cooled sheets were cleaned in a 5vol% sulfuric acid solution to remove any oxide scale before the cold rolling process. During the rolling process, the thickness of the stainless steel sheets was reduced to 1.9 mm (43% reduction in thickness) and 1.6 mm (52% reduction in thickness) after two and four passes of rolling steps, respectively, as illustrated schematically in Fig. 1.

      Figure 1.  Schematic of the cold rolling process.

      After rolled, the sheets were allowed to air cool and then subjected to heat treatment at 900°C for an hour in a muffle furnace followed by quenching in tap water, as elaborated in Fig. 2.

      Figure 2.  Schematic of the post-rolling heat treatment process showing the heating and cooling cycle.

      The specimens were designated based on the number of rolling passes and post-rolling heat treatment process as given in Table 1.

      For microstructure and hardness analysis, small specimens of 1 cm2 surface area were cut from the rolled sheet samples. These specimens were cold mounted in polyester resin, and grinding was performed sequentially by using silicon carbide paper with grit sizes ranging from coarse P60 to fine P2000. All the specimens were chemically etched in a freshly prepared aqua regia (5vol% HNO3, 95vol% HCl) solution. Optical microscopy was used to examine the microstructural details (Leica Model DM 15000M, Germany). The grain size (length of the grains) was measured from the micrographs (at 100× magnification) obtained from the optical microscope by using the Image J software package by following the ASTM E112 standard practice. The grain size is the general term that may represent either the grain area or length of the elongated grains. In this article, the grain size values correspond to the length of the elongated grains produced during the rolling process. X-ray diffraction (XRD) of these specimens was performed to observe any phase transformation during the cold rolling and quenching processes. Hardness was measured by using a Vickers hardness tester (Shimadzu, Japan) by applying a 2.94 N load for 10 s and by using a diamond pyramid indenter.

      TreatmentSpecimen designation
      As-receivedAR
      Two rolling passes2R
      Four rolling passes4R
      Quenched as-receivedRQ
      Quenched two rolling passes2RQ
      Quenched four rolling passes4RQ

      Table 1.  Designations of the specimens used in the study

      For electrochemical testing, disk-shaped (16 mm in diameter) as-received, cold-rolled, and rolled heat-treated 316L stainless steel specimens were carefully cut by using a diamond wire without generating a significant amount of heat that could otherwise deteriorate the microstructure. To prepare the working electrodes, these disk specimens were individually connected with the copper wire by using silver epoxy and mounted in the polyester resin by leaving the top surface (2.0096 cm2) exposed. A three-electrode electrochemical cell that has a working electrode, Ag/AgCl reference electrode, and graphite rod as a counter electrode was used as shown in Fig. 3. Commercially prepared 0.9wt% NaCl solution (UNISA pharmaceuticals Industries Ltd., Pakistan) was used as an electrolyte for electrochemical analysis. Open circuit potential (OCP, Eoc) was measured for 1 h to stabilize the surface potential and to maintain the equilibrium between the working electrode and electrolyte. Cyclic polarization tests were performed by sweeping the potential from −0.5 V vs. Eoc to an apex potential of +1.5 V vs. Eoc with a forward scan rate of 2.5 mV·s−1. The potential was reversed from the apex potential to the final potential (−0.5 V vs. Eoc) with the same scan rate.

      Figure 3.  Electrochemical test system showing the electrical connection of a three-electrode cell with a potentiostat (SBF—Simulated body fluid).

    3.   Results and discussion
    • Fig. 4 illustrates the optical micrographs of AR, 2R, and 4R 316L stainless steel specimens. The AR specimen (Fig. 4(a)) consisted of equiaxed austenite (γ) grains that had an average grain size of (21.31 ± 0.5) µm. The 2R and 4R specimens, shown in Figs. 4(b) and 4(c), respectively, presented elongated grains in the rolling direction.

      Figure 4.  Microstructures of 316L stainless steel specimens: (a) AR; (b) 2R; (c) 4R. The white arrows in (b) and (c) represent the rolling direction.

      As shown in Fig. 5, the average grain size of the 2R specimen was (28.28 ± 1) µm due to two-pass rolling, which resulted in a 43% decrease in thickness and a 33% increase in grain length compared with the AR specimen. For the 4R specimen, the larger deformation during the rolling process led to a 52% reduction in the sheet thickness and produced a preferentially elongated grain structure with an average grain size of (68.87 ± 4.5) μm. The excessive deformation of the 4R specimen resulted in a 144% increase in the grain length compared with the 2R specimen and a 223% increase in grain elongation compared with the AR specimen. A preferentially oriented lamellar grain structure was also formed in the 4R specimen due to the severe plastic deformation of the austenite grain structure as illustrated in Fig. 4(c).

      Figure 5.  Average grain sizes of AR, 2R, 4R, RQ, 2RQ, and 4RQ 316L stainless steel specimens.

      During the rolling process, the applied stress level could range from 350 to 500 MPa (estimated based on the reduction percentage in thickness and yield strength (210 MPa) of the annealed 316L stainless steel). A stress level greater than the yield strength of 316L stainless steel but lower than the ultimate strength (550 MPa) is always required to induce plastic deformation during the cold rolling process. However, the strain hardening of the 316L stainless steel specimens during the rolling process would increase as a function of rolling passes, and the required stress level in the successive rolling passes would increase. Given the application of multiple rolling passes, the exact stress level after and before each rolling pass was not calculated due to the complexity in the phase transformation and induced strain hardening. However, the detailed microstructural analysis revealed the content of strain-induced martensite increased with an increase in the number of rolling cycles.

      Fig. 6(a) presents the presence of a well-formed polygonal shape grain structure that was formed due to sequential annealing and quenching of as-received 316L stainless steel [27]. The austenite grain showed the formation of annealing twins as indicated by the arrow signs in the microstructure. In the RQ specimen during quenching, the fast cooling rate restricts the formation of martensite due to the larger solubility of carbon in the austenite phase and its slow diffusion from the austenite solid solution during the quenching process. However, the carbon depletion at the austenite grain boundaries and the formation of precipitated chromium carbide cannot be neglected [28], as highlighted in Fig. 6(a). The average grain size in the RQ specimen was (43.93 ± 2.0) µm, which was higher than that of the AR specimen due to the annealing heat treatment (at 900°C for 1 h) and recrystallization of the grains before quenching. During the annealing process, the recrystallization and growth of new austenite grains would occur depending on the extent of the preexisting stresses in the grains associated with the prior thermomechanical treatment history of the specimen. In the 2RQ specimen, the formation of relatively coarse austenite grains was observed during the annealing process and the average grain size was (54.72 ± 0.5) µm. This size was approximately 25% larger than that of the austenite grains present in the RQ specimen. In the case of the 4RQ specimen, the relatively coarse austenite grain structure and recrystallization of the pre-deformed grains (grains that experienced excessive deformation during the rolling process) would occur along with the annealing twins, as evident in Fig. 6(c). The relatively dark grains in Figs. 6(b) and 6(c) indicated the restoration of the excessively deformed grains during the recovery and recrystallization process. These grains were the retained austenite grains in the microstructure that had experienced severe mechanical deformation during the rolling process. These grains would release induced stresses during thermal treatment and coexist with the recrystallized austenite grains. In other words, due to excessive mechanical deformation of 316L stainless steel during the rolling process, the appreciable recovery of the austenitic grains and recrystallization would take place along the twin lines, as evaluated from the microstructural analysis. This condition also resulted in the formation of relatively coarse grains (average grain size of (55.10 ± 1.5) µm), as shown in Fig. 5, and presented a slight elongation (0.7%) along the rolling direction in the 4RQ specimen compared with the 2RQ specimen. A twin structure with the austenite grains would form due to severe mechanical deformation or due to the mismatch in the grain orientation during the recrystallization process [29], as shown in Figs. 6(a)6(c).

      Figure 6.  Microstructures of 316L stainless steel specimens: (a) RQ; (b) 2RQ; (c) 4RQ (T—Twinning in the austenite grain structure; RC—Recrystallized grain).

    • To illustrate the grain orientation and any phase transformation during the cold rolling and quenching process, the XRD patterns were obtained as shown in Figs. 7(a) and 7(b). The diffraction pattern of the AR specimen confirmed the existence of the austenite matrix phase. However, a small collar peak at 44.5° could be affiliated with the formation of α′-martensite in the microstructure, as indexed by the (110) crystallographic plane. The diffraction peaks observed at 43.6°, 50.7°, 74.6°, and 90.5° correspond to the austenite (γ-phase) with preferential grain orientation along the (111), (200), (220), and (222) crystal planes, respectively.

      Figure 7.  XRD patterns of 316L stainless steel specimens: (a) AR, 2R, and 4R; (b) RQ, 2RQ, and 4RQ.

      Austenite to α′-martensite conversion could happen due to the arrangement of Fe and C atoms and the transformation of face-centred cubic (fcc) to bcc crystal structure during thermomechanical treatment [29] of 316L stainless steel. As evident in the diffraction pattern, the characteristic diffraction peaks corresponding to the hcp ε-martensite phase were not observed, whereas the direct γ → α′ transformation could result due to severe plastic deformation of the grain structure as reported in the literature [3031]. In the case of 2R and 4R specimens, the relatively small diffraction peak at 2θ = 44.5° highlighted the formation of α′-martensite phase within the γ-austenite matrix phase. Unlike in the AR specimen, in addition to the diffraction peaks at 43.6°, 50.6°, and 90.3°, an additional peak at 95.8° appeared, which corresponds to the (311) crystallographic plane of the γ-phase. However, the diffraction peak at 74.6° associated with the γ-phase peak disappeared after the rolling process, as shown in Fig. 7(a). As a result of mechanical deformation, the increase in peak intensities at 43.7°, 50.8°, and 90.5° and the emergence of a peak at 95.8° in the diffraction patterns of the 2R and 4R specimens compared with AR highlighted the preferred grain orientation in the (111), (200), (222), and (311) directions, respectively. The austenite contents were calculated by using Eq. (1) [32]:

      where ${V}_{\rm{A}},{I}_{{{\text{α}}}^{'}},\;\;{\rm{and}}\;{I}_{\text{γ}}$ are the austenite volume fraction in the matrix and integrated intensities of the ${{{\text{α}}}}^{'}\left(110\right)\;{\rm{and}}\;{{\text{γ}}}\left(111\right)$ diffraction peaks, respectively. The relatively larger amount of austenite contents (~82vol%) was calculated from the diffraction patterns of the 2R and 4R specimens compared with the austenite content (68vol%) in the AR specimen.

      Compared with the diffraction pattern of the AR specimen, an appreciable increase in the peak intensity at 2θ = 74.6° in the RQ specimen indicated the preferred orientation of austenite grain structure in the (220) direction. Similarly, the diffraction peak (at 96.1°) associated with the (311) plane was also present in the XRD patterns of 2RQ and 4RQ, as shown in Fig. 7(b). In contrast to the 2R and 4R specimens, the origin of the peak associated with the (220) plane was evident in the diffraction patterns of the 2RQ and 4RQ specimens. However, the relative increase in the peak intensities associated with the γ-phase at 43.6°, 50.8°, 74.7°, 90.5°, and 95.8° corresponded to the recrystallization of the γ-phase. The diffraction peak at 44.5° related to the formation of α′-martensite was evident in both 2RQ and 4RQ specimens. Based on Eq. (1), no significant change in the austenite contents was observed in the RQ (68vol%), 2RQ (80vol%), and 4RQ (83vol%) specimens compared with the AR, 2R, and 4R specimens, which highlighted that post-rolling heat treatment had the least effect on the austenite contents. Austenite-to-martensitic transformation is possible when the applied stress level during the mechanical deformation is higher than the yield strength of the austenite phase; this phenomenon is known as strain-induced martensitic transformation. The austenite-to-martensite transformation that occurs at a lower applied stress level than its yield strength is called stress-induced martensite transformation [3334].

    • The hardness of AR, 2R, 4R, RQ, 2RQ, and 4RQ 316L stainless steel specimens are illustrated in Fig. 8. The severe plastic deformation could induce excessive stresses in the grains, which could affect the hardness of the 316L stainless steel [3536]. The hardness of the 4R specimen was HV 285, which was greater than those of the 2R (HV 267) and AR (HV 190) specimens, as shown in Fig. 8. As a result of the excessive deformation of the grain structure caused by the repetitive rolling passes, the considerable strain hardening of the grain structure increased the hardness of the 2R and 4R 316L stainless steel specimens [3738].

      Figure 8.  Hardness profiles of the 316L stainless steel specimens.

      The hardness of the RQ, 2RQ, and 4RQ specimens was significantly decreased to HV 146, 151, and 147, respectively, as shown in Fig. 8. The relatively lower hardness of RQ, 2RQ, and 4RQ specimens than those of the AR, 2R, and 4R specimens was due to the formation of retained austenite during the annealing and quenching process. The α′-martensite formation was favored by a diffusionless shear mechanism that may be produced during the cold rolling process. However, the presence of strain-induced martensite would revert to austenite during the annealing process. The large solubility of carbon in the austenite grain structure and its restricted diffusion during the quenching process would stabilize the austenite phase. Microstructural changes occur during the annealing recovery process, which could appreciably reduce the strain hardening within the grain structure [3940].

    • The OCP values or free potentials of the AR, 2R, 4R, RQ, 2RQ, and 4RQ specimens were measured in 0.9wt% NaCl solution, and the values are shown in Fig. 9. The 316L stainless steel AR specimen showed a more negative OCP of −150.7 mV vs. Ag/AgCl as compared with the 2R (−102.7 mV vs. Ag/AgCl) and 4R (−108.1 mV vs. Ag/AgCl) specimens, which represents its higher corrosion tendency in 0.9wt% NaCl solution. In other words, the relatively negative OCP of the AR specimen and a gradual shift in potential toward a positive value indicated its delayed electrochemical process. In the case of the 2R and 4R 316L stainless steel specimens, the OCP slightly shifted toward a positive potential possibly due to the rapid reaction of these specimens with the electrolyte and the formation of chromium oxide passive film, which could hinder their further dissolution. The negative OCP values of AR, 2R, and 4R specimens in 0.9wt% NaCl solution corresponded to the active state of the surface that showed a high tendency to react with the electrolyte. In simple words, the highly deformed grains would preferably promote the dissolution of ionic species from the surface. Under applied conditions, on the surface of 316L stainless steel, the following reactions are expected to occur (reactions (2)−(4)) [41]:

      Figure 9.  OCP values of the 316L stainless steel specimens.

      After the AR and rolled specimens underwent annealing and quenching, the recovery and recrystallization process would decrease the surface reactivity of 316L stainless steel. Significantly positive OCP values of RQ (−70.6 mV vs. Ag/AgCl) and 2RQ (4.70 mV vs. Ag/AgCl) highlighted their lower corrosion tendency compared with the AR and 2R specimens under the same conditions. However, the OCP behavior indicated that the severely deformed grain structure of the 4RQ specimen and the possible formation of strain-induced martensite within the austenite matrix was not fully retrieved during the recovery and recrystallization process. In other words, the restoration of the pre-deformed austenite grain structure was incomplete during the annealing process. Grain size also influences the corrosion tendency of 316L stainless steel, as described in the literature [2,7]. The relatively finer grain of 2RQ ((54.72 ± 0.5) μm) compared with that of the RQ and 4RQ specimens may also be the possible reason for its low corrosion tendency.

    • The cyclic polarization trends of AR, 2R, and 4R specimens in 0.9wt% NaCl solution are shown in Fig. 10. The quantitative information obtained from these polarization curves is given in Table 2.

      To measure the effect of mechanical deformation and post-heat treatment, the intrinsic corrosion tendency of the specimen was estimated from the polarization curves. In this regard, the scans were initiated from the more negative potential (−500 mV lower than the OCP values of the respective specimens) to estimate the influence of preexisting oxide species on the surface [42]. These species may abruptly form on the surface of 316L stainless steel upon exposure to air or electrolytes. No significant difference in the cathodic polarization curves was observed, as shown in Fig. 10, thereby indicating the possible reduction of oxide species (reverse of reactions (2)−(4)) and water during cathodic polarization. These reduction reactions depend on the thermodynamic stability of water and kinetic activity of the surface species toward water reduction according to reaction (5).

      A considerable difference in the anodic polarization curves was evident in Fig. 10, which corresponds to the surface oxidation and/or active dissolution of 316L stainless steel in 0.9wt% NaCl solution as a function of mechanical deformation induced during the cold rolling process. The current generated by both AR and 2R specimens and the anodic polarization trends were almost similar.

      Figure 10.  Comparison of cyclic polarization curves of the AR, 2R, 4R, RQ, 2RQ, and 4RQ 316L stainless steel specimens.

      Specimenicorr /
      (μA·cm−2)
      Corrosion rate /
      (mm·a−1)
      Ecorr /
      mV
      Eprot /
      mV
      Epit* /
      mV
      Pitting
      resistance / mV
      Protection
      tendency / mV
      AR23.4010.258−200.0−187.512.5
      2R4.4400.049−185.0−145.040.0
      4R1.5500.017−112.5322.3434.8
      RQ0.7828.636 × 10−3−185.5975.31160.8
      2RQ0.1321.448 × 10−3−175.0762.3937.3
      4RQ0.0303.302 × 10−4−185.0930.51115.5
      Note: * shows measured at 10 μA/cm2.

      Table 2.  Electrochemical parameters calculated from the cyclic polarization curves of 316L stainless steel specimens

      The corrosion current density (icorr) and the corrosion rate were calculated by fitting the cyclic polarization curves within the Tafel region with the use of Echem Analyst software. The quantitative information of the kinetic parameters is given in Table 2. The results show that the icorr value for AR was appreciably higher (23.401 μA·cm−2) than that of the as-rolled and heat-treated specimens. Interestingly, compared with AR, the sixfold decrease in icorr was presented by the 2R specimen (4.440 μA·cm−2). After four-pass rolling of 316L stainless steel (4R specimen), the icorr value was further decreased to 1.550 μA·cm−2, which highlighted the positive effects of mechanical deformation on its corrosion resistance. The lower corrosion rate (0.017 mm/a) value registered by the 4R specimen also indicated its appreciably low dissolution rate in the 0.9wt% NaCl solution. When the AR 316L specimens were heat-treated and quenched, the formation of a twin structure was nucleated from the strain-induced martensite phase as evident in RQ, which could further decrease the icorr value, i.e., 0.782 μA·cm−2, relative to the AR, 2R, and 4R specimens. Similarly, the icorr values of the annealed and water quenched 2RQ and 4RQ specimens were further decreased to 0.132 and 0.030 μA·cm−2, respectively. This result indicated the positive effects of the heat treatment procedure, which significantly improved the corrosion resistance of as-rolled 316L stainless steel.

      During the anodic polarization of these specimens, a gradual increase in current with potential was observed, which may be attributed to the formation and growth of the passive oxide film on their surface. However, an abrupt increase was observed in the current by the 4R specimen at a relatively low potential of 322.3 mV vs. Ag/AgCl (designated as breakdown or pitting potential, Epit). This increase in current represented the instability of the passive oxide film that may break above this potential, leading to the accelerated localized dissolution of the surface [43]. This result highlights the deleterious effects of mechanical deformation on the passive film stability formed on the 4R specimen. Both Fe and Cr oxide/hydroxides may coexist in the passive film that may form on the surface of 316L stainless steel at a potential of >−0.2 V vs. Ag/AgCl. However, the Fe(OH)3 species could interact with the Cl ions present in the solution under applied conditions (at the potential of >−0.2 V vs. Ag/AgCl) and may form Fe(OH)Cl+ complex via reaction (6) [44].

      As a result of the severe mechanical deformation, as evident from the microstructure of the 4R specimen, the continuity of the surface oxide film may be disturbed, which is indicated by the origin of a large positive current loop and the very negative protection potential (Eprot) during reverse polarization scan. The inability of the 4R specimen surface to re-passivate in the 0.9wt% NaCl solution is indicated from Eprot < corrosion potential (Ecorr), and the development of a large positive current loop corresponds to its accelerated pitting tendency. The rapid increase in current during anodic polarization is attributed to the formation of soluble FeCl2+ species (${\rm{FeC}}{{\rm{l}}_{{\rm{(aq)}}}^{2 + }}$ in reaction (7)) at the local defect sites (where local pH is significantly dropped) that is depleted with Cr resulting in the pit nucleation and growth [44].

      Also, at potentials larger than Epit, the oxidation of Cr(III) to Cr(VI) species (reaction (8)) could further promote the accelerated dissolution of the 316L stainless steel [41].

      The AR and 2R specimens presented almost similar anodic polarization trends and the current density increased gradually with an increase in potential, approaching 0.1 mA·cm−2 at approximately 1200 mV vs. Ag/AgCl. The continuous increase in anodic current density in the range of 1–100 μA/cm2 highlighted the greater dissolution tendency of AR and 2R specimens. The abrupt increase in current at a potential beyond 1200 mV vs. Ag/AgCl could be associated either with the water oxidation or the localized dissolution of the specimens due to the breakdown of a passive oxide film. This increase in current and formation of a positive current loop during the reverse scan validated the hindrance in the repassivation of the surface. This behavior also indicated the occurrence of vigorous localized reactions that could accelerate the pitting corrosion of AR and 2R specimens in 0.9wt% NaCl solution.

      The appreciably large current response (a current greater than the forward anodic current) during the reverse scan dictates that once the passive film is broken, the pitting corrosion may proliferate in an uncontrolled manner. These polarization trends indicate that the continuous increase in current during anodic polarization and the abrupt increase in current at a high potential (>1000 mV vs. Ag/AgCl) suggested the formation of an unstable passive film that decreased the protection tendency of the AR and 2R specimens. In other words, a reverse scan of both AR and 2R specimens showed that the reverse anodic curves formed a hysteresis loop that intersected the forward anodic curve at Eprot (−187.5 and −145.0 mV vs. Ag/AgCl, respectively) slightly higher than the Ecorr corresponded to the protection tendency of 12.5 and 40 mV, respectively, as given in Table 2. The protection tendency indicates that once the potential of AR and 2R specimens is increased by 12.5 and 40 mV, respectively, from their respective Ecorr, the uncontrolled pitting corrosion on the surface has a high chance of occurring under applied conditions. On the other hand, the significant mechanical deformation of the grain structure in 4R specimens deleteriously affected the passive film stability, as confirmed by the very low Epit and the rapid increase in current density beyond 322.3 mV (measured at 10 μA·cm−2). This reverse anodic scan of the 4R specimen intersected the forward scan at a more negative potential than its Ecorr value (−112.5 mV), which indicates its poor passivation tendency and accelerated localized dissolution.

      Fig. 10 shows the cyclic polarization curves of the RQ, 2RQ, and 4RQ specimens. Significant improvement in the passive film stability and appreciable decrease in current response during anodic potential sweep was evident because of the decreased dissolution tendency of post-rolled heat-treated 316L stainless steel. The almost similar polarization trends of the RQ, 2RQ, and 4RQ specimens indicate the beneficial effects of the recovery and recrystallization process, which may homogenize the microstructure, thereby increasing the passive film’s stability. For instance, at 500 mV vs. Ag/AgCl, the current density of the RQ specimen was almost three times lower than that of the AR specimen. At the same applied potential, the anodic current density of the 4R specimen was approximately two orders of magnitude higher than the 4RQ specimen. Similarly, the approximately 10 times lower icorr of these post-rolled heat-treated specimens highlighted the decrease in the dissolution rate. Most interestingly, the Epit of 4RQ increased considerably to approximately 930.5 from 322.3 mV vs. Ag/AgCl (as registered by the 4R specimen), as evaluated from the polarization trends. Based on these results, the pitting resistance of the 4RQ specimen was calculated to be 1115.5 mV, which was almost three times higher than the 4R specimen. However, slightly low pitting resistance (937.3 mV) was observed in the case of 2RQ after annealing + quenching treatment. During reverse polarization scans, the origin of positive hysteresis loop and very low intersection potential (at a potential more negative than Ecorr) presented by all the RQ, 2RQ, and 4RQ specimens ensured that once the pit is initiated on the surface, it would grow rapidly in an uncontrolled manner [4546]. Also, these trends suggest once the passive film that formed on the rolled and post-heat-treated specimens is broken on the surface, the repassivation/self-healing tendency of these specimens would become very difficult in the 0.9wt% NaCl solution.

    4.   Conclusions
    • (1) The austenite grain structure of 316L stainless steel was transformed to strain-induced martensite during cold rolling, and the hardness considerably increased from HV 190 to 285. However, the post-annealing + quenching of the as-rolled specimens relieved the induced stresses and converted the strain-induced martensite back to austenite due to the recovery and recrystallization process.

      (2) As evident from the cyclic polarization scans, the positive shift in Ecorr and the appreciable decrease in icorr and anodic current density were registered by the post-rolled heat-treated (annealing + quenching) specimens. In other words, the corrosion rate and pitting tendency of the post-rolled heat-treated specimens (2RQ and 4RQ) were significantly decreased. This result also highlighted the beneficial effects of post-heat treatment on improving the passive film stability of the rolled 316L stainless steel.

      (3) The formation of a large hysteresis loop and the intersection of reverse polarization curve at the potential < Ecorr indicated that once the passive film is broken, the localized dissolution of the as-rolled and post-rolled heat-treated specimens would proceed in an uncontrolled manner.

      (4) The beneficial effects of post-rolling heat treatment were evident from the considerably large pitting resistance of the 4RQ specimen (1115.5 mV), which is approximately three times higher than that of the severely deformed 4R specimen.

Reference (46)

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