Friction-stir processing of AISI 440C high-carbon martensitic stainless steel for improving hardness and corrosion resistance

Journal of Materials Processing Tech. 277 (2020) 116448

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Friction-stir processing of AISI 440C high-carbon martensitic stainless steel for improving hardness and corrosion resistance

T

⁎

L. Panb, C.T. Kwoka,b, , K.H. Loa,b a b

Department of Electromechanical Engineering, University of Macau, China Institute of Applied Physics and Materials Engineering, University of Macau, China

A R T I C LE I N FO

A B S T R A C T

Associate Editor: R Mishra

Friction-stir processing (FSP) has been attempted to harden the surface of a high-carbon martensitic stainless steel (AISI 440C) and to improve its corrosion resistance. Owing to high input of heat, the penetration depth of the FSPed zone increases in the range from 1.5 to 2 mm when the translational speed decreases. As the translational speed increases, higher content of retained austenite in the advancing side and center of the processed zone are obtained due to higher cooling rate. The maximum hardness of the FSPed 440C is up to 779 HV1 and higher than that of the conventionally hardened sample austenitized at 1200 ℃ (618 HV1). The pitting corrosion resistance of the sample processed at 300 mm·min−1 in 3.5 wt% sodium chloride solution is the lowest among the FSPed samples. It is mainly attributed to the heterogeneous microstructures at the retreating side (martensite and carbides) and advancing side and center (with martensite and retained austenite) leading to selective attack occurred at the more active region, i.e. the retreating side.

Keywords: Friction-stir surface hardening Stainless steel Microstructure Pitting

1. Introduction High-carbon martensitic stainless steel (MSS) AISI 440C is a kind of chromium steel containing no nickel. It possesses high mechanical strength and hardness, high wear resistance but moderate corrosion resistance thanks to its high carbon content (about 1.1 wt%). It is widely used in various engineering applications including valves, turbines, pump impellers, cutlery, industrial blades, molds and dies, surgical and dental instruments, to name a few. Although it contains a relatively high content of chromium (17 wt%), its corrosion resistance is the lowest among the various stainless steels. High carbon content in AISI 440C has a potential detrimental impact on its corrosion resistance owing to formation of chromium carbides, thus depleting chromium in the metallic phases and degrading passivity. Through solid-state transformation of the surface, engineering components made of MSSs can be modified at specific locations for enhancing the corrosion resistance and hardness while the bulk properties can be preserved. Without changing the overall chemical compositions, surface hardening of MSSs can be done by localized heating and subsequent quenching. The common methods used to harden the surface of MSSs include flame hardening and induction hardening. However, Dossett and Totten (2013) have reported that the drawback of flame hardening is the possibility of part distortion, while induction hardening conducted at high frequencies requires close coupling between the part and the coil, ⁎

which must be precisely maintained. On the other hand, Lo et al. (2003) reported that the surface hardness of AISI 440C was increased to 600–800 HV using laser transformation hardening with a CW Nd:YAG laser. Mahmoudi et al. (2010) reported that surface hardness of laserhardened AISI 420 was increased to 490 HV with a pulsed Nd:YAG laser. However, the penetration depth of the laser-hardened zone of these MSSs is limited to only 250–300 μm. According to Dodds et al. (2013), increasing laser power density or interaction time could increase the penetration depth but it would cause oxidation, melting or vaporization of the surface. Mishra and Ma (2005) derived friction-stir processing (FSP) from friction-stir welding as a solid-state surfacing process. Padhy et al. (2018) has reported that FSP has been developed for grain refinement, homogenization of microstructure and compositions, surface hardening, fabricating surface/bulk composites, repairing cracked and defective components and enhancing superplasticity of the engineering alloys. During FSP, a stirring tool is pricked into a sample and move in a programmed route. Ma (2008) reported that the temperature achieved could be as high as 60–90% of the melting point of the material and the strain was up to 40 and strain rate was in the range of 1–1000 s−1. Padhy et al. (2018) reported that FSP could lead to immense thermal exposure, severe plastic deformation, dissociation and dissolution of secondary phase and inter-mixing of the material. FSP variables including translational and rotation speeds of the stirring tool can control

Corresponding author. E-mail address: [email protected] (C.T. Kwok).

https://doi.org/10.1016/j.jmatprotec.2019.116448 Received 4 March 2019; Received in revised form 17 August 2019; Accepted 30 September 2019 Available online 04 October 2019 0924-0136/ © 2019 Elsevier B.V. All rights reserved.

Journal of Materials Processing Tech. 277 (2020) 116448

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heat generation. Ma (2008) reported that higher rotation speed caused higher heat input leading to coarser grain size and lower hardness in the FSPed zone, whereas increase in translational speed decreases the time to heat exposure (i.e. lower heat input) resulting in reduced grain size and higher hardness. Hence, both factors should be compromised in order to attain sufficient heat generation and material plasticization. Kulekci et al. (2015) also reported that porosity and crack in the ferrous alloys can be eliminated by using FSP. Guo et al. (2019) found that friction surfacing could modify the microstructure by fragmentation of coarse MnS inclusions in austenitic stainless steel AISI 316 L to finer inclusions. Pan et al. (2019) showed that hardness and corrosion resistance of AISI 420 MSS could be enhanced by FSP with a hardness change from 184 to 697 HV1 and a noble shift in pitting potential from 4 to 176 mVSCE as compared with the annealed AISI 420. Moreover, Yao et al. (2015) found that the hardness of cast irons was improved by FSP due to the homogeneous distribution of ultra-refined graphite particles. Cui et al. (2016) studied the relationship between microstructure evolution and mechanical properties during FSP of AISI 201 austenitic stainless steel (ASS). The maximum hardness of the FSPed zones (260 HV) was found to be slightly higher than that of the base material (210 HV) owing to the high dislocation density in such zones induced by severe plastic deformation during the FSP. Hajian et al. (2015) reported that the highest yield and tensile strengths of the FSPed 316 L ASS are 1.6 and 1.2 times higher than those of the base material respectively in spite of a 50% decrease in ductility. The increase in hardness of the FSPed zone agreed with the increase in tensile properties and was attributed to grain size reduction. Moreover, Hajian et al. (2014) found that the cavitation erosion resistance of the FSPed 316 L was 3–6 times of the substrate owing to grain size reduction. On the other hand, some researches on FSP of MSS AISI 420 with 13 wt% Cr and 0.45 wt% C have been conducted. Dodds et al. (2013) reported that the grain size of the FSPed AISI 420 was finer than that of the conventionally hardened one, and some retained austenite was present in the martensitic structure. The hardness of the FSPed AISI 420 achieved (682 HV0.3) was slightly higher than that of the conventionally hardened one (662 HV0.3) while the wear resistance of the former was generally higher than that of the latter. Pan et al. (2019) reported that the corrosion resistance of the FSPed AISI 420 was enhanced as reflected by the increase in corrosion and pitting potentials as compared to the annealed and conventionally hardened AISI 420 while the hardness (697 HV1) is close to that of the finding of Dodds et al. (2013). As a matter of fact, the contents of carbon and chromium in AISI 440C are higher than those in AISI 420. The effect of the FSP conditions on evolution of microstructure, mechanical and corrosion behaviors deserves detailed exploration. In the literature, to no study on FSP of AISI 440C has been reported. However, Puli and Janaki Ram (2012b) reported that AISI 440C coatings have been successfully deposited on a steel substrate via friction surfacing (FS) with rotational speeds of 800–1500 rpm and translational speeds of 1–6 mm/s. In contrast to FSP, successful FS has to be carried out by exerting a rotating consumable rod with a constant axial force on the substrate. Puli and Janaki Ram (2012b) found that the 440C coatings formed by FS showed a martensitic structure with less and finer carbides and higher content of retained austenite (40%) as compared with the conventionally hardened 440C. The FSed AISI 440C coatings exhibited adequate hardness (590 HV0.3), superior corrosion resistance but inferior wear resistance due to the presence of a high content of retained austenite as compared to the conventionally hardened AISI 440C. For the engineering components made of the heat-treatable ferrous alloys like AISI 440C, FSP can be applied for hardening the surface directly and no consumable rod is necessary. In this work, the influence of translational speed of the stirring tool on the hardness and corrosion behavior of AISI 440C using FSP was investigated. Also, the hardness and electrochemical behavior of the annealed and conventionally hardened 440C were studied for comparison.

Fig. 1. Schematic diagram of the FSP setup.

2. Experimental details The as-received AISI 440C plate was in annealed condition (designated as AR440C) with dimensions of 100 × 50 × 6 mm3. Its nominal compositions in weight percent were: 17% Cr; 0.75% Mo; 1% Mn; 1.1% C; 1% Si; 0.049% P; 0.03% S and balance Fe. To achieve different microstructures, the conventionally hardened (as-quenched) samples were obtained via austenitizing the AR440C at 1000 ℃ or 1200 ℃ for 1 h followed by air quenching (designated as AQ440C-1000 and AQ440C1200, respectively). Before FSP, the surface of AR440C was ground with 600-grit emery paper and then rinsed with ethanol. FSP was conducted with a frictionstir welding machine (China FSW Center, FSW-TS-M16) and the FSP setup is depicted in Fig. 1. A W-Re stirring tool with a flat shoulder of 15-mm diameter and 10-mm height was used. Argon was used to avoid oxidation of the surface of the samples at a rate of flow of 15 L·min−1. The stirring tool was discarded after traveling for 200 mm for avoiding contamination of the samples. Behnagh et al. (2012) reported that the stirring tool was tilted at 1.5° for facilitating the forging action at the edge of shoulder. The stirring tool was pricked into the substrate at a depth of 0.1 mm. To compromise between heat input and processing efficiency for FSP of AISI 440C, a high rotation speed (2000 rpm) with various translational speeds (150, 200 and 300 mm·min−1) of the stirring tool was selected as the processing conditions. The cross-section and surface of the FSPed samples were mechanically polished and then chemically etched with the etchant prepared with 25 g FeCl3, 25 ml HCl and 100 ml H2O. The samples AR440C, AQ440C-1000 and AQ440C-1200 were prepared with the same method for microstructural analysis. Their microstructures were studied by a scanning-electron microscope (SEM, Hitachi S-3400 N). The phases present at the surface of AR440C, AQ440C-1000, AQ440C-1200 and the center of FSPed samples were identified using an X-ray diffractiometer (XRD, Rigaku MiniFlex 600) with a Cu Kα source at a scanning rate of 1°·min−1. The SEM equipped with Electron Backscatter Diffraction (EBSD) was used to distinguish different phases through different diffraction patterns obtained from a stationary electron beam interacting with a tilted sample. The volume fraction of austenite (γ) was analyzed using an image analyzing software (Tango) with the EBSD phase maps taken from the FSPed specimens. Each result was obtained by taking the mean value of five fields of dimensions equal to 25 × 28 μm. Based on the thermal modeling of FSW presented by Song and Kovacevic (2003), a modified model was developed for analytically

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estimating the temperature field in FSP of 440C by applying COMSOL multi-physics software in the present study. As there was no pin on the tool, the heat was generated only from the friction between the tool shoulder and the sample. The heat generation at the distance r from the center of the tool is given by Eq. (1).

μ (Fn As ) ωr , if T < Tmelt qshoulder (r , T ) = ⎧ if T > Tmelt 0, ⎨ ⎩

(1)

where Fn is the normal force, AS is the surface area of the shoulder, Tmelt is the melting temperature of AISI 440C (1483 °C), μ is the friction coefficient and ω is the angular velocity of the tool. Eq. (2) below describes the heat transfer in the plate. As the coordinate system is attached to, and hence moving with the processing tool a convection term and the conductive term are added in the heat transfer equation.

ρCp μ⋅+∇T + ∇⋅(−k∇T ) = Q

(2)

where k represents thermal conductivity, ρ is the density, Cp is specific heat capacity of the substrate material, and μ is the transverse speed of the tool. Due to surface-to-ambient radiation and natural convection, the heat loses from upper surface to lower surface. The corresponding heat fluxes at upper and lower surfaces are described in Eqs. (3) and (4). 4 qu = hu (T0 − T ) + εσ (Tamb − T4 )

(3)

4 qd = hd (T0 − T ) + εσ (Tamb − T4 )

(4)

where hu and hd are heat transfer coefficients for natural convection, T0 is an associated reference temperature, ε is the surface emissivity, σ is the Stefan-Boltzmann constant, and Tamb is the ambient temperature. Conforming to ASTM-E384 (2011) standard, hardness of various samples was analyzed with an automatic Vickers hardness analyzer (Wilson VH3100) with a diamond indenter subjected to a 1-kg load. For each FSPed sample, 3 tests at the center of the processed zone along the transverse direction (started from 0.1 mm beneath the surface) and 3 tests from the advancing side (AS) to the retreating side (RS) and also in the longitudinal direction (0.2 mm beneath the surface) were carried out. According to ASTM G61-86 (2018) standard, potentiodynamic polarization tests of AR440C, AQ440C-1000, AQ440C-1200 and the FSPed samples were conducted in 3.5 wt% sodium chloride (NaCl) solution (open to air) at 25 ± 1 °C with a potentiostat (PAR VersaStat 3 F). To keep a constant surface roughness, the sample surfaces were ground using the 1000-grit emery paper. The average surface roughness of the samples after grinding with the 1000-grit emery paper was measured to be 0.12 ± 0.02 μm using a profilometer. To avoid crevice corrosion, passivation treatment with 25% nitric acid at 50 °C for an hour was applied to the surface. A saturated calomel electrode (SCE) and a pair of graphite rods were used as the reference electrode and counter electrodes, respectively. Scanning of potential started at 200 mV under the free corrosion potential followed by increasing the potential in the noble direction at a rate of 10 mV·min−1. Corrosion current density (Icorr) and corrosion potential (Ecorr vs. SCE) were determined from the polarization curves using Tafel extrapolation. Tefal Extrapolation was performed by extending the linear portions of the anodic and cathodic parts of the polarization curve back to their intersection. A potential scan of approximately ± 300 mV about Ecorr was required to determine if a linear section of at least one decade of current was present. These two lines eventually met at a point where the corrosion current density (Icorr) is obtained. Tafel extrapolation was performed by the computer software PowerCORR. The pitting potential (Epit vs. SCE) was obtained at the inflection point. The protection potential (Eprot vs. SCE) stands for the potential at which the scan reverses and intersects the polarization curve. More than 2 tests were carried out for each sample. Pit morphology of the samples after corrosion was characterized using optical

Fig. 2. Microstructure of (a) AR440C; (b) AQ440C-1000 and (c) AQ440C-1200.

microscope and SEM. Electrochemical impedance spectroscopy (EIS) was performed at Ecorr ± 10 mVSCE with an AC frequency in the range 10 kHz – 10 mHz in 3.5 wt% NaCl at 25 ± 1 °C. The EIS results were analyzed and fitted using the software Zview.

3. Results and discussion 3.1. Microstructure The SEM micrographs of AR440C, AQ440C-1000 and AQ440C-1200 are shown in Fig. 2. For AR440C, numerous particles of chromium carbide (M23C6) are randomly distributed in the ferritic grains (Fig. 2a). The average grain size of the ferrite (α) is 16 μm and the size of the carbides is in the range of 0.2–3 μm. For AQ440C-1000 and AQ440C3

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Fig. 3. XRD patterns of AR440C, AQ440C-1000, AQ440C-1200 and FSPed samples (at center) processed at different translational speeds.

plunging a stirring tool into the surface of AISI 440C plate with translational motion until a track of FSPed zone was made. Significant heat was generated in the FSPed zone (including the AR, center and RS) owing to severe plastic deformation and high strain rate. Dodds et al. (2013) reported that the AS and RS of the FSPed 420 fabricated at a low rotational speed (300 rpm) and a translational speed of 150 mm·min−1 were symmetric due to the use of the stirring tool with a pin. Without using the pin, a high rotational speed (2000 rpm) and various translational speeds (150–300 mm·min−1) were used in the present study. From Fig. 5, the FSPed zone of FSP440C-150 fabricated at 150 mm·min−1 is more symmetric than the ones fabricated at 200 and 300 mm·min−1 (Fig. 5b–c). For the samples FSP440C-200 and FSP440C-300, the AS is thicker than the RS. As the stirring tool was moving at high translational speeds, a distinctive asymmetry at the AS and RS was created. Yadav and Bauri (2012) investigated the microstructure and mechanical properties of the FSPed aluminum. They reported that the material at the RS around the stirring tool experienced lower degree of thermal exposure and plastic deformation because the tangential component of rotation and translation of the stirring tool were opposite resulting in smaller frictional force and hence less generated heat. The depth of penetration and width of the FSPed zone increased as the translational speed decreased or heat input increased. At 300 mm·min−1, the depth of penetration and width of the FSPed

1200, as the carbon content is high, both samples contain plate martensite (α’) with an average grain size of 1.3 and 2 μm, respectively. A few fine carbide particles with average size of 1 μm is observed in AQ440C-1000 while no carbide particles can be found in the martensitic matrix of AQ440C-1200. Compared to AR440C, the carbide content in the conventionally hardened samples AQ440C-1000 and AQ440C-1200 is much lower. Fig. 3 shows the XRD patterns of AR440C, AQ440C-1000, AQ440C-1200 and the centers of the FSPed samples. The phases detected in the samples using XRD are consistent with the SEM observation. For AR440C, the ferrite (major phase) and M23C6 (minor phase) are detected. For AQ440C-1000 and AQ440C-1200, martensite (α’) is the major phase with minor phase M23C6 in the former but carbide is not detected in the latter. In addition to the martensite, a low content of retained austenite (γ) and carbides are detected in the FSPed samples (Fig. 3). Fig. 4 typically shows the top view of FSP440C-150 fabricated at 150 mm·min−1. The transverse cross-sections of the FSPed samples fabricated at different translational speeds (150, 200 and 300 mm·min−1) showing the basin-shaped FSPed zones are depicted in Fig. 5. Mishra and Ma (2005) reported that the formation of basinshaped FSPed zones was due to severe plastic deformation and frictional heating from the contact between the surface of the samples and the shoulder of the stirring tool. As shown in Fig. 1, FSP involved

4

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Fig. 4. Top view of FSPed 440C processed at 150 mm/min.

zone of FSP440C-300 were 1.5 mm and 13.6 mm, respectively (Table 1). As the translational speed was reduced to 150 mm·min−1, the penetration depth of FSP440C-150 was increased to 2 mm, the width became 14.7 mm and was very close to the diameter of shoulder of the stirring tool (15 mm). According to Saeid et al. (2008), the fundamental principles of interrelations of FSP parameters with peak temperatures (Tp) and physical metallurgy of AISI 440C are shown below:

Tp

ω2 ⎞ = K⎛ ν × 10 4 ⎠ ⎝ ⎜

Tm

Table 1 FSP parameters, width and penetration depth of FSPed 440C. Translational speed (mm·min−1)

Width (mm)

Processed depth (mm)

FSP440C-150 FSP440C-200 FSP440C-300

150 200 300

14.7 14.5 13.6

2.0 1.7 1.5

α

The peak temperature is inversely proportional to the traverse speed. The lower traverse speed, the higher the traverse speed can be achieved. In addition, Hamilton et al. (2015) reported that the viscosity increased with the increase in traverse speed leading to shallower depth of the processed zone. Avila et al. (2018) reported that the peak temperature, cooling rate and the dwell time in AS are higher than that of RS resulting in the different microstructures in the AS and RS. Longer exposure time, higher peak temperature and higher cooling rate induced more carbides to dissolve into the matrix and more austenite retained in the AS. The difference in temperature became more

⎟

(5)

where Tm (°C) is the melting temperature of the alloy, ω and ν are rotational speed and translational speed respectively, K and α are the constants. From Eq. (5) and the experimental results of the present study under constant ω, a simplified relationship between Tp and v is obtained as given below:

Tp ∝

Sample

1 να

Fig. 5. Cross-sectional views of the FSPed 440C processed at various translational speeds: (a) 150 mm·min−1; (b) 200 mm·min−1; and (c) 300 mm·min−1. 5

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Fig. 6. SEM micrographs of (a) FSP440C-150 and (b) FSP440C-300: (i) AS; (ii) center; and (iii) RS.

translational speed (i.e. higher cooling rate) (Table 2). The retained austenite content at the center of FSP440C-300 was the highest (15.1%), while the retained austenite content at the center of FSP440C150 was the lowest (6.4%). In the present study, increase in the retained austenite content at the AS and center of the FSPed samples is attributed to increase in cooling rate, which is essentially governed by the translational speed of the stirring tool. According to the work of Lo et al. (2003), the laser processing conditions used (including fluence and scanning speed of the laser beam) for AISI 440C caused no melting and it was still in the solid state in laser transformation hardening. It was indicated that the retained austenite content in the laser-hardened 440C increased as the scanning speed of the laser beam (cooling rate) increased. In addition, Puli and Janaki Ram (2012b) also reported that the temperatures for the martensitic transformation start (Ms) and finish (Mf) decreased as the cooling rate increased. High cooling rate suppressed martensitic transformation and austenite was retained.

pronounced as the translation speed of the tool increased. This is due to the cooler material entering the processed zone from RS and then AS, and the increased flow rate more effectively lowers temperature in RS. The SEM and EBSD results at the AS, center and RS of the FSPed samples processed at difference translational speeds are shown in Figs. 6, 7. Puli and Janaki Ram (2012a) found that the austenite in the friction-surfaced AISI 316 L was dynamically recrystallized and its grain size was refined. In the present study, the austenite formed during FSP at austenitizing temperature was also dynamically recrystallized and then transformed into martensite during cooling. Ultimately the average grain size of the martensite formed after quenching at the center of the processed zones of the FSPed samples is reduced to about 0.75 μm as depicted in Fig. 6a(ii) and b(ii). The AS and center of the FSPed samples mainly contained martensite and retained austenite with tiny amount of carbides (Figs. 6, 7). The content of retained austenite in the AS and center of the FSPed samples increased with increasing in 6

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Fig. 7. EBSD band contrast superimposed with phase maps of the cross-sectioned FSPed 440C processed at various translational speeds: (a) FSP440C-150 (150 mm·min−1) at center; (b) FSP440C-200 (200 mm·min−1) at center; (c) FSP440C-300 (300 mm·min−1) at center and (d) FSP440C-300 (300 mm·min−1) at RS. The red phase is austenite (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article).

martensite became 0.6 wt% and 17.5 wt%, respectively and were then sufficient for attaining high hardness and high corrosion resistance. Using COMSOL multi-physics analysis, the surface temperature distribution of the sample FSP440-150 during FSP at a rotation speed of 2000 rpm and translational speed of 150 mm/min is shown in Fig. 8. It shows that the substrate transports the heat away as the tool leaves. While at the front of the tool, new cool material enters. The predicted peak temperatures for the samples FSP440-150, FSP440-200 and FSP440-300 were 1310, 1160 and 962 °C, respectively. It should be noticed that this model only focused on the thermal effect and temperature distribution during FSP. The geometry was supposed to be symmetric around the process line. The samples were assumed to be infinitely long, so the effect near the edges of the plate was neglected. In addition, the stirring effect (plastic deformation) was not taken into consideration due to the complex material flow. Also, the difference in the temperature at AS and RS cannot be analyzed and further investigation is required. According to the results of the thermal model and Fe-Cr-C pseudo-binary diagram at 17 wt% Cr by Lippold and Kotecki (2005), the microstructural changes are as follows:

Table 2 Content of retained austenite in FSPed 440C. Sample

AS (%)

Center (%)

RS (%)

FSP440C-150 FSP440C-200 FSP440C-300

10.6 16.2 18.4

6.4 10.6 15.1

0 0 0

Moreover, Luo et al. (2017) reported that increase in cooling rate would decrease Ms and Mf, which narrowed the martensite transformation range for a high-speed steel (AISI M42). As the time was too short for martensite transformation under higher cooling rate, more austenite was retained in the final microstructure. Lippold and Kotecki (2005) calculated the pseudo-binary phase diagram of Fe-Cr-C at 17 wt% Cr. As an annealed AISI 440C with 1.1 wt % C was heated to 1000 °C, the initial phases (ferrite + M23C6) were transformed into (austenite + M23C6). When the temperature of AISI 440C was further raised to 1100–1310 °C, such initial phases were transformed into (austenite + M7C3). Upon rapid quenching, the austenite was transformed into martensite, and carbide phases (M23C6 or M7C3) were dispersed in the martensitic matrix. It is noticed that the austenite formed at 1000 °C contained only about 0.3 wt% C and 11.7 wt% Cr. The martensite formed upon quenching possessed the same compositions as the austenite formed at 1000 °C. Therefore, heattreating of AISI 440C at 1000 °C (i.e. AQ440C-1000) cannot achieve adequate hardness and corrosion resistance (i.e. C and Cr contents of the matrix should be at least 0.6 wt% and 12 wt%, respectively). When the austenitizing temperature for AISI 440C was increased to 1200 °C (i.e. AQ440C-1200), the contents of C and Cr in the austenite /

α + M23C6 → γ + M23C6 / (M7C3) → α′ + γ + M23C6 During FSP, a peak temperature range of 962–1310 °C could be reached. The temperature was above the solvus temperature, so α in the sample was transformed to γ and the carbides were partially dissolved. After rapid quenching, the specimens mainly contained martensite and some small undissolved carbides. The difference in microstructure at AS, center and RS of the FSPed samples depends on the temperature achieved for austenitizing and cooling rate in FSP. Puli and Janaki Ram (2012b) also reported that the 7

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Fig. 8. Surface temperature of the workpiece being friction processed at rotation speed of 2000 rpm and traverse speed of 150 mm/min.

temperatures for martensitic transformation start (Ms) and finish (Mf) decreased as the cooling rate increased. High cooling rate suppressed martensitic transformation and austenite was retained. The higher cooling rate accounts for the existence of fine martensite and retained austenite at the AS and center [Fig. 6a(i),(ii) and b(i),(ii)]. From Fig. 6a (iii), b(iii), some undissolved carbides can be observed in the center and RS, implying that the temperature achieved there was not as high as that at the AS. Luo et al. (2017) explained the relationship between the undissolved carbides and the austenitizing temperature by checking the microstructures of AISI M42 high-speed steel after different austenitizing temperature. They concluded that lower austenitizing temperature led to decrease in carbide dissolution. For all FSPed specimens, retained austenite was absent in the RS since its temperature was relatively low (800–1100 °C). From the pseudo-binary phase diagram of Fe-Cr-C at 17 wt% Cr calculated by Lippold and Kotecki (2005), as the temperature was raised to 800–1100 °C followed by rapid quenching, the annealed 440C underwent the following transformation sequence: α + M23C6 → γ + M23C6 → α′ + γ + M23C6 The initial phases α + M23C6 in the annealed 440C was transformed to γ + M23C6. M23C6 was partially dissociated and the amount of carbon dissolved in the austenite was low at the RS. Baghjari and Akbari (2013) also found that the diffusion coefficient of carbon decreases with decreasing temperature. Consequently, the ‘low-carbon’ austenite was transformed into martensite and no austenite was retained. 3.2. Hardness The hardness distribution along the centerlines of the processed zones of the FSPed samples is illustrated in Fig. 9a. The highest hardness was observed near the surface and the hardness dropped along the depth. Plateau region is not observed in the FSPed zone suggesting a microstructure with decreasing carbon content along the depth due to decrease in temperature and cooling rate. The hardness at 0.1 mm beneath the surface for FSP440C-150, FSP440C-200 and FSP440C-300 are 779, 746 and 728 HV1, respectively (Table 3). All FSPed samples are harder than AR440C (181 HV1), AQ440C-1000 (510 HV1) and AQ440C-

Fig. 9. Hardness profiles of FSPed samples along (a) the centerline; (b) the longitudinal direction.

Table 3 Maximum hardness and corrosion parameters of various samples. Sample

Max. hardness (HV1)

Ecorr (mVSCE)

Icorr (nA/cm2)

Epit (mVSCE)

Eprot (mVSCE)

Epit - Ecorr (mVSCE)

Eprot - Ecorr (mVSCE)

Pit density (%)

AR440C AQ440C-1000 AQ440C-1200 FSP440C-150 FSP440C-200 FSP440C-300

181 510 618 779 746 728

−232 −129 −122 −110 −132 −106

338.4 231.6 82.4 8.5 12.8 31.7

36 137 137 151 139 70

−256 −166 −245 −190 −213 −265

268 266 259 261 271 176

−24 −37 −123 −80 −81 −159

1.13 0.64 0.64 0.62 0.63 0.67

8

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Fig. 10. Plots of hardness and various corrosion parameters against translational speed.

1200 (618 HV1). Puli and Janaki Ram (2012b) reported that the hardness of the friction-surfaced coating possessed lower hardness (590 HV0.3) than that of the FSPed samples fabricated at a higher rotation speed (2000 rpm) in the present study. The lower hardness in the friction-surfaced coating is due to the existence of very high content of retained austenite (40 vol%) and some undissolved carbides resulting from lower heat input. The surface of AISI 440C went through severe plastic deformation during FSP. The surface hardness at the center of the FSPed 440C (728–779 HV1) is higher than that of AQ440C-1200 (618 HV1) due to severe plastic deformation and grain refinement other than the presence of martensite. Hajian et al. (2014) reported that the hardness of the FSed 316 L was the highest near the surface due to the highly severe plastic deformation there. High dislocation density was introduced in the FSPed zone, thus the hardness was improved and gradually decreased along the depth to hardness of the substrate (181 HV1). To a lower extent, hardness enhancement of the FSPed samples is owing to reduction in grain size. The average grain size of the FSPed samples (0.75 μm) is smaller than that of AQ440C-1200 (2 μm). Furthermore, the maximum surface hardness of the FSPed samples decreased (Fig. 9) as translational speed increased was owing to increase in the retained austenite content. Pan et al. (2019) reported that the maximum hardness achieved in FSPed 440C (728–779 HV1) is higher than that of the FSPed 420 (698 HV1) processed at 150 mm·min−1 (designated as FSP420-150, as shown in Table 3) because of the lower carbon content (0.45 wt%) or lower martensite content in AISI 420. Fig. 9b shows the variation in hardness along the longitudinal direction of the FSPed zone including AS, center and RS at 0.2 mm beneath the top surface. The heat generated was different at the AS, center and RS resulting in non-uniform hardness distribution in the longitudinal direction. The hardness in the RS decreased more significantly than that in the AS. In addition, the hardness at the RS was lower than that at the AS for all FSPed samples. It is because the austenitizing temperature at the RS of the FSPed 440C is relatively low leading to incomplete dissolution of carbides and martensitic transformation.

Fig. 11. Potentiodynamic polarization curves of (a) as-received and conventionally hardened 440C; and (b) FSPed 440C in 3.5% NaCl solution at 25 °C.

3.3. Cyclic polarization and pit morphology The polarization curves of AR440C, AQ440C-1000, AQ440C-1200 and the FSPed 440C in 3.5 wt% NaCl solution at 25 °C are depicted in 9

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Fig. 12. Pit morphology of AR440C after polarization test.

carbide, resulting in a lower Cr content in the solid solution. According to Baroux (1995), the regions depleted in Cr became the active sites for pitting attack in chloride-containing environments. From Fig. 10a and Table 3, the corrosion resistances of the two conventionally hardened samples (AQ440C-1000 and AQ440C-1200) are higher than that of AR440C. While the Ecorr and Epit for AQ440C-1000 and AQ440C-1200 are very close but the Icorr of the former is higher than the latter by a factor of 2.8. For AQ440C-1200 (Ecorr = -122 mVSCE and Epit = +137 mVSCE), a larger amount of carbides was dissolved in the matrix during austenitizing, so more Cr was released into the solid solution. Lu et al (2015) reported that an increase in the austenitizing temperature to 1130 ℃ could enhance the pitting resistance of AISI 420 in 3.5 wt% NaCl solution with a noble shift in Epit from -30 to 400 mVSCE. The enhanced corrosion resistance of AISI 420 heat-treated at high austenitizing temperature was due to a more uniform Cr distribution and a reduced chromium carbides content. On the other hand, Candelária and Pinedo (2003) also studied the corrosion behavior of AISI 420 in 0.5 M H2SO4 using immersion test. A decrease in the corrosion resistance was observed when the austenitizing temperature increased up to 1075 ℃. On the other hand, the corrosion resistance increased above 1100 ℃ owing to the beneficial effects of carbide dissolution and also the internal martensite lattice stresses. Their results are consistent with the present findings. From Figs. 10 and 11b, the pitting corrosion resistance increases as the traverse speed decrease. The pitting corrosion resistance of FSP150 is the highest among the FSPed samples fabricated at different translational speeds, while its retained austenite content in the AS and center is the lowest. This seems to be counter-intuitive as the retained

Fig. 11. For all samples, a passive region can be clearly observed and abrupt increase in current density is recorded as the passive film breaks. Ecorr, Icorr, Epit, Eprot are extracted from the polarization curves, and Epit – Ecorr (representing pit initiation resistance) and (Eprot – Ecorr) (representing pit propagation resistance) are calculated as shown in Table 3. The values of (Epit – Ecorr) of the samples are very close except for FSP440C-300. The smaller the value of (Epit – Ecorr) (i.e. FSP440C300), the larger the quantity of the pits (higher pit density) can be observed as compared with the other FSPed samples (Table 3). From Fig. 11, the polarization curves of all samples do not show repassivation above Ecorr. This means that the pits will propagate in all samples at Ecorr once the pits are formed. Pit propagation means the development of pits in size, i.e. growth of pit. (Eprot – Ecorr) (pit propagation resistance) is the resistance to such development. From Table 3, the magnitude of Eprot – Ecorr of FSP440C-300 is the largest (with a negative sign) indicating its lowest pit propagation resistance. The plot of Ecorr, Epit, Eprot, (Epit – Ecorr) and (Eprot – Ecorr) against the translational speed is shown in Fig. 10. It can be observed that Epit and (Eprot – Ecorr) decrease as the translational speed increases (retained austenite content increases) whereas no significant trend can be observed for the plot of Ecorr and (Epit – Ecorr) vs translational speed (Fig. 10). From Fig. 11a and Table 3, AR440C possesses the lowest corrosion resistance as evidenced by the most active Ecorr (-232 mVSCE), lowest Epit (+36 mVSCE) and highest Icorr (338.4 nA/cm2). Mahmoudi et al. (2010) reported that the pitting resistance of the MSSs was mainly dependent on the distribution of Cr in the solid solution. Since AR440C was in the annealed condition, some Cr combined with carbon as chromium 10

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Fig. 13. Pit morphology of AQ440C-1200 after polarization test.

matrix so the more active ferrite was corroded. Compared to AR440C, the randomly distributed pits on the corroded surface of AQ440C-1200 are smaller and shallower (Fig. 13). The pit densities (area of pits per unit sample surface area) of AR440C and AQ440C-1200 were found to be 1.13% and 0.64% respectively, reflecting that AR440C experienced more severe pitting attack. Most of the chromium carbides were dissociated and dissolved in the martensite of AQ440C-1200, which was more corrosion resistant than AR440C. Among all samples, the Ecorr and Epit of FSP440C-150 is the noblest and its Icorr is the lowest. Compared with AQ440C-1200, the Ecorr and Epit of the FSP440C-150 shifted in the noble direction due to the existence of retained austenite. Puli and Janaki Ram (2012b) found that the effect of severe plastic deformation, frictional heating, fragmentation and dissolution of carbides led to reduction in the size and quantity of the carbides as in the friction-surfaced 440C coating. In addition, the existence of a high retained austenite content in the FSPed 440C is due to the high translation speed (i.e. high cooling rate). Kwok et al. (2000) reported that the retained austenite is more corrosion resistant than the martensite in the laser-surfaced melted AISI 420. The improvement in corrosion resistance of the FSPed 440C is owing to the existence of higher retained austenite content, more Cr in the solid solution in the FSPed 440C as compared to the AQ440C-1200. FSP440C-150 and FSP440C-200 possess higher corrosion potential and pitting potential than that of AQ440C-1200. Puli and Janaki Ram (2012b) reported that the noble shift in Ecorr for the friction-surfaced 440C coating was

austenite possesses higher corrosion resistance than the martensite. It is mainly attributed to the heterogeneous microstructures at the RS (martensite and carbides) and AS and center (with martensite and retained austenite). It could be observed that selective attack occurred at the more active region, i.e. the RS of the FSPed samples (Fig. 14) while the AS and center were cathodic. Compared with FSP300, more carbides were dissolved in the RS of FSP150 as the temperature achieved for austenitizing was higher. FSP150 with fewer undissolved carbides in the RS resulted in higher corrosion resistance although it has lower retained austenite content in the AS and center. This reflects the undissolved carbides in the RS is more detrimental (the anode) to the pitting corrosion resistance than the lower austenite content in the AS and center. The pit morphology of the corroded surfaces of AR440C and AQ440C-1200 after polarization test is illustrated in Figs. 12 and 13, respectively. From Fig. 12, the neighboring regions around the undissolved carbides have been severely corroded. This indicates that the pitting initiated around the undissolved carbides, and would develop into stable pits with fallout of the carbides particles because the area close to Cr-rich carbides was depleted in Cr as evidenced by the EDS line scan (Fig. 12c). When the potential reaches the relatively low Epit for AR440 (36 mVSCE), the pits start to be initiated at the Cr-depleted regions and grow with further increase in the potential. According to Huttunen-Saarivirta et al. (2016), the carbides were reported to possess a nobler corrosion potential than the nearby ferritic 11

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Fig. 14. Pit morphologies of (a) FSP440C-150 and (b) FSP440C-300 after polarization test.

consistent with the present findings. From Table 3, the Epit of FSP440C300 was lower than that of AQ440C-1200. According to Pan et al. (2019), the Ecorr and Epit of the FSP420-150 were nobler than those of the FSPed 440C but the Icorr of the FSP420-150 was higher than that of the FSPed 440C (Table 3). It was due to the higher Cr content (17 wt%) in the FSPed 440C. The low corrosion resistance of FSP440C-300 is due to the large difference in microstructure between its AS/center and the RS. At the AS and center, martensite and retained austenite were present. Martensite and undissolved M23C6 existed in the RS but no retained austenite was found. Since the microstructures of the AS, center and RS of the FSPed samples were heterogeneous, micro-galvanic effect will arise. The RS is more active than the AS and center. At the RS, the undissolved carbides caused a reduction in the chromium content in the martensite, which became the active site for initiation of pits. From Fig. 14, selective attack was evidenced by the pit morphology

of the FSPed samples. The heterogeneous microstructures at the AS, center and RS of the FSPed zones led to selective corrosion attack at the RS. From Fig. 14b, selective attack was more severe at the RS (anodic) of FSP440C-300 with more undissolved carbides in the RS due to lower austenitizing temperature. According to Jiang et al. (2017), the undissolved carbides in the high-nitrogen martensitic stainless steel would result in less stable passive film with lower repassivation ability and more initiation sites for pitting. The pit morphology of the FSPed 440C after the potentiodynamic polarization test is quite different from those of AR440C and AQ440C. From Fig. 14a, the pits mainly are distributed at the RS of FSP440C-150, and the pits were seldom found at the AS and center. From Fig. 14a, a hole with lacy metal covers in FSP440C-150 can be observed. Ernst et al. (1997) reported that the existence of lacy pattern of corrosion pits were related to passivation and undercutting near the pit mouth. While for FSP450-300 (Fig. 14b), larger pits were concentrated at 12

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AS and center are the cathodes. In addition, the pits in the RS of FSP440C-300 are larger than those of FSP440C-150 (Fig. 14). Consequently, selective attack in the FSPed sample fabricated at a higher translational speed was more serious. The pit density of FSP150 and FSP300 is 0.62% and 0.67%, respectively. 3.4. EIS study The Nyquist and Bode plots of the tested samples are showed in Fig. 15. The Nyquist plot (Fig. 15a) exhibited a capacitive loop in the high-medium frequency range, which is a typical behavior of stainless steels as Souza et al. (2017) indicated. It shows a compressing trend of capacitive semicircle from FSP440C-150 to AR440C. The diameter of the loop of AR440C was the smallest representing the lowest corrosion resistance. The diameter of the sample increased after FSP and quenching at 1200 °C. The diameter of FSP440C-300 was smaller as compared to those of AQ440C-1200 and FSP440C-150, indicating its lower corrosion resistance. The Bode plots show typical capacitive like behavior with one time constant (Fig. 15b). Xu et al. (2013) indicated that the closer the phase angle to 90°, the more perfect is the capacitive response of the electrode surface. Also, Liu et al. (2016) reported that the deviation from ideal capacitance is attributable to dispersion effect caused by local inhomogeneities in the surface. FSP440C-300 exhibited the lowest phase angle representing the lowest homogeneities in FSPed region. An equivalent circuit shown in Fig. 16 is based on the one proposed by Lu et al. (2015) for studying the electrochemical properties of a 13%Cr-type martensitic stainless steel in 3.5 wt% NaCl. The impedance CPE represents the double-layer capacitance of the steel/electrolyte interface and is described in Eq. (6).

Z (CPE ) = Q0−1 (jω)−α

Fig. 15. EIS results of AR440C, AQ440C-1200, FSP440C-150 and FSP440C-300: (a) Nyquist plots and (b) Bode plots.

Z = Rs + [Rct−1 + Q0 (jω)−α ]−1

Table 4 Fitted parameter values of experimental specimens obtained from EIS spectra.

AR440C AQ440C-1200 FSP440C-150 FSP440C-300

Rs /Ω cm2

7.63 7.95 6.62 6.29

CPE α (0-1)

Q0/ (Ω−1 cm-2 sα)

0.89 0.90 0.90 0.87

6.23 × 10−5 5.29 × 10−5 4.22 × 10−5 5.35 × 10−5

Rct /Ω cm2

χ2

7.87 × 104 1.86 × 105 3.00 × 105 1.45 × 105

0.0043436 0.00088172 0.0013939 0.0017965

(7)

where Rs is the solution resistance, Rct is the charge transfer resistance. The fitted results from the EIS spectra are listed in Table 4. As can be seen from Table 4, Rct varies from 7.87 × 104 to 3.00 × 105 Ω cm2. The largest Rct of FSP440C-150 (3.00 × 105 Ω cm2) represents the highest corrosion resistance due to film thickening as reported by Liu et al. (2016). On the contrary, the Rct of FSP440C-300 is the lowest, which agrees well with the results of potentiodynamic polarization. When α = 1, the CPE becomes a pure capacitor. However, the double-layer capacitance usually deviates from pure capacitance because of the dispersion effect. As a result, the values of α are in the range of 0.87 to 0.9 for the experimental samples. El-Egamy and Badaway (2004) studied the passivity of 304 austenitic stainless steel in alkaline sodium sulphate solutions. They found that the behavior of CPE is closely related to the heterogeneity of the electrode interface. According to the work of Hefny et al. (1985), when α approached 1, the reciprocal of Q0 is proportional to the thickness of passive film. In the present study, the values of Q0 increased from 4.22 × 10−5 Ω-1 cm-2 sα (FSP440C-150) to 6.23 × 10-5 Ω-1 cm-2 sα (AR440C). The improvement in corrosion resistance of FSP440C-150 is due to thickening of the passive film.

Fig. 16. Equivalent circuit used to fit EIS data.

specimen

(6)

where Q0 is the CPE element, j = −1 , ω = 2πf is the angular frequency and α is the dispersion coefficient related to the surface roughness and homogeneity, where 0 ≤ α ≤1. The total impedance of the system is given by Eq. (7):

the RS. Under high magnification the pits of FSP440-300 have an openmouth morphology (Fig. 14b). The result of line scan by EDS shows that the bottom of a pit is rich in Cr, indicating the existence of the uncorroded carbides in the matrix. Moreover, several micro-pits at the bottom of a primary pit can be observed as indicated by the arrows (Fig. 14b). Selective attack occurs at RS, which is the anode whereas the

4. Conclusions

• The depth of penetration of the FSPed 440C decreased and the retained austenite content at the center/advancing side increased with the increasing translational speed because of decrease in heat input.

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•

• •

Different heat inputs at the advancing side, center and retreating side resulted in variation in microstructure, hardness and corrosion resistance. The surface hardness at centers of the FSPed zones is in the range of 728–779 HV1 and higher than the conventionally hardened 440C (618 HV1) because of the presence of martensite, grain refinement and severe plastic deformation. The heat generated was different at the advancing side, center and retreating side, thus resulting in nonuniform hardness distribution. Compared to the conventionally hardened 440C, pitting resistance of the FSPed 440C processed at 150 and 200 mm·min−1 was improved as reflected by the noble shift in the pitting potential, being attributable to fragmentation and dissolution of chromium carbides. Difference in microstructure at the advancing side and center (with martensite and retained austenite) and retreating side (with martensite and carbides) of the FSPed 440C processed at a high translational speed (300 mm/min) is due to insufficient heat input and low austenitizing temperature (800–1100 °C) at the retreating side. From polarization and EIS results, the pitting resistance of FSP300 was lower than that of AQ440C-1200 and selective attack occurred at the retreating side (more anodic).

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