Materials & Design 83 (2015) 820–828
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Investigation on the solution treated behavior of economical 19Cr duplex stainless steels by Mn addition Yinhui Yang ⇑, Jianchun Cao, Yang Gu School of Materials Science and Engineering, Kunming University of Science and Technology, Kunming 650093, PR China
a r t i c l e
i n f o
Article history: Received 11 November 2014 Revised 18 May 2015 Accepted 23 May 2015 Available online 20 June 2015 Keywords: Duplex stainless steels Mn addition Solution treatment Toughness Pitting corrosion
a b s t r a c t Effect of Mn on microstructure, mechanical property and pitting corrosion of 19% Cr economical duplex stainless steels with solution temperatures ranging from 1040 to1220 °C has been investigated. The austenite content increases with more Mn addition, but decreases by increasing solution temperature, which can be inferred by trend of partition coefﬁcient KMn with solution temperature. Meanwhile, a balanced austenite-ferrite duplex structure of solution-treated specimens was obtained with Mn addition. The impact energy at 20 °C increased with decreasing solution temperature from 1220 °C to 1040 and 1120 °C, and improved by more Mn addition due to more ductile austenite phase formation. These toughness variations were consistent with fracture morphology characteristic changing. The effect of more Mn addition and solution treatment of 1120°Con decreasing of tensile strength and 0.2% offset yield strength were slight. However, the elongation to fracture (%) fell greatly with Mn addition up to 8.1 wt.% for as-rolled and solution treated specimens due to larger deformation strains of austenite than that of ferrite. The decreasing trend of pitting corrosion potential became slower with Mn addition from 3.6 to 8.1 wt.%. The pitting corrosion resistance was lowered by increasing solution temperature due to more weakened repassivation ferrite phase formation. Ó 2015 Elsevier Ltd. All rights reserved.
1. Introduction Duplex stainless steels (DSS) exhibit an attractive combination of mechanical properties and corrosion resistance and are thus widely used in various applications such as power plants, desalination facilities, the marine industry, and chemical plants [1–3]. It is well known that such good properties of duplex steels rely on a two-phase microstructure comprised of approximately equal amounts of austenite (c) and ferrite (d) . Solution treatment is important heat treatment process for DSS, which can promote dissolution of carbide and other precipitation phases after rolling process, thus maintain high mechanical property and corrosion resistance of DSS [5–7]. However, different solution treatment temperatures can result in volumetric equilibrium fraction of d and c phase changes in DSS, and Cr, Mn and Mo content in the two phases will be altered accordingly. These changes have a great inﬂuence on ferrite phase transformation behavior and mechanical, corrosion and formability . Li et al.  reported that both the ferrite content and the decomposition dynamics in ferrite caused by solution treatment
⇑ Corresponding author. E-mail address: [email protected]
(Y. Yang). http://dx.doi.org/10.1016/j.matdes.2015.05.052 0264-1275/Ó 2015 Elsevier Ltd. All rights reserved.
affect the followed thermal aging behavior. Solution treatment is an important process in the production of DSS. It is generally used to re-dissolve the harmful precipitates and eliminate the macro-segregation. In addition, the content, shape and distribution of the ferrite phases, as well as the alloy compositions in ferrite and austenite, can be adjusted by altering the solution temperature and time. The decomposition kinetics in ferrite is closely related to the chemical compositions, especially the Ni content, which could be changed in the solution treatment process. It is known that nitrogen can improve the pitting corrosion resistance as well as enhance the strength of stainless steels . Meanwhile, the manganese has been considered for austenite former and added to increase solubility of nitrogen in stainless steels [11–13], and its detrimental effect on the pitting corrosion resistance is usually associated with MnS inclusion formation. Besides that, Toor reported that the resistance to pitting and metastable pitting corrosion of a new high Mn–Ni free DSS decreased with increasing Mn content since the number of (Mn, Cr) oxides acted as preferential sites of pitting . In addition to commonly used austenite stabilizing element Ni, the Mn, C, and N elements also can be used as substitutions for Ni additions to stabilize the austenitic structure effectively [15,16], but the good welding performance require low carbon content for DSS application. Nitrogen, as an austenite stabilizer, also
Y. Yang et al. / Materials & Design 83 (2015) 820–828
improves resistance and increases strengthening effect of steel [17,18], while its limited solubility in stainless steel increase production difﬁculty of DSS. Manganese is approximately 5–8 times cheaper than Ni, and is an effective austenitic stabilizer with increasing solubility of nitrogen in stainless steels. Thus, the Mn– N duplex stainless steels are actively being developed by replacing expensive Ni with low cost Mn with a certain amount of nitrogen addition. Jang et al. believed that tensile properties of CD4MCU cast duplex stainless steel was determined by the volume fraction of hard ferritic phase and the shape of austenitic phase, and the resistance to pitting corrosion and stress corrosion cracking in 3.5% NaCl + 5% H2SO4 aqueous solution was recovered with the addition of 0.8% Mn to 2% Mn . However, the effect of Mn on the microstructural evolution, as well as the mechanical and corrosion properties, of duplex stainless steels has not been well established. On the other hand, solution treatment is an effective way to maintain the high resistance to localised corrosion, by which the alloying elements can be in solid solution and homogeneously distributed in metal in order to attain passivation effect. Thermal aging behavior of DSS is closely related to the ferrite content and the chemical compositions in ferrite. For a particular grade of steel, these factors are controlled by the fabrication process, including the solidiﬁcation and the following heat treatment, especially by the solution treatment process. However, little information exists regarding the inﬂuence of Mn on the mechanical property and pitting corrosion of DSS under different solution treatment temperatures. Therefore, the present work attempts to provide a further understanding of the effect of Mn additions on microstructure, mechanical property and pitting corrosion resistance of 19% Cr economical duplex stainless steel under different solution temperatures.
the point fell on the boundary was counted as a half. Electrochemical potentiodynamic polarization was performed in a deaerated 3.5 wt.% NaCl solution at temperatures varying from 25 °C (±1 °C). A platinum sheet and a saturated calomel electrode (SCE) were used as the counter and reference electrodes, respectively. The specimens, embedded in epoxy resin with an exposure area of 100 mm2, acted as a working electrode. Prior to each experiment, the specimens were ground mechanically up to 3000 grit, rinsed with distilled water and dried in hot air. Anodic potentiodynamic polarization was tested through linear sweep technique at a sweep rate of 0.1 mV/s, from the free corrosion potential to 1200 mV (SCE). All potentials are given against the SCE. After that, the corrosion morphology was observed by SEM. Corrosion potential (Ecorr), passive ﬁlm breakdown potential (Eb), passivation current density (Ip) were obtained from the polarization curves. Corrosion current density (icorr) is commonly obtained by the extrapolation of the cathodic and anodic slopes between 50 and 100 mV away from Ecorr. Eb is the potential the passive ﬁlm breaks down then leading to rapid increase of current density and it was determined as the current density reached a value of 100 lA/cm2.
3. Results and discussion 3.1. Microstructure and composition analysis Fig. 1 shows optical microstructures obtained from the specimens with different Mn contents after solution treatment at 1040 °C, 1120 °C and 1220 °C. Under such conditions, they show an obvious duplex structure of austenite and ferrite. The bright etched austenite (c) islands were embedded in a gray etched ferrite (d) matrix, and no obvious secondary precipitates were found in the ferrite, austenite matrix and c/d phase boundaries. The microstructures of different Mn content specimens varied signiﬁcantly with different solution treatment temperatures from 1040 to 1220 °C. The volume fraction as a function of solution treatment temperature is plotted in Fig. 2 for the specimens solution treated at different temperatures. With an increase in Mn content from 3.6 to 8.1 wt.%, the amount of austenite phases increased, which was reduced greatly by increasing solution treatment temperature from 1040 to 1220 °C. Under these three different solution treatment temperatures, the volume fractions of austenite phases ranging from 40.1% to 60.9% were obtained, indicating good balance of duplex phases with different Mn addition. When treated with solution at 1040 °C, the austenite volume of sample alloy 1 was 54.1%, while the austenite volume of samples alloy 1 and alloy 2 was 52.5% and 54.6% respectively with solution treatment at 1120 °C, which has a microstructure consisting of ferrite and austenite phases with approximately a 1:1 ratio. This suggested that low Mn content addition in 19% Cr economical DSS can promote austenite phases formation with comparatively wide solution treatment temperature, which covered a commonly used solution treatment temperature of 1050 °C in DSS production. However, increasing solution treatment temperature to 1220 °C greatly reduced austenite volume fraction, indicating that effect of Mn on austenite formation was reduced by higher solution temperature. The Cr, Ni equivalent formulas of Schaefﬂer graph are as follows :
2. Experimental procedures The raw materials were melted in a 25 kg vacuum induction furnace then cast as a single square ingot. The casting ingots were hot forged into 30-mm plates, then the as-forged samples were hot rolled into 12 mm plates at temperature ranging from 1050 to 1200 °C. The chemical compositions of the rolling plates, designated as Alloy 1, Alloy 2 and Alloy 3, are shown in Table 1. Samples with dimension of 12 12 50 mm were cut from these rolling plates then solution treated at 1040 °C, 1120 °C, 1220 °C for 30 min respectively, followed by water quenching. The microstructures of specimens were electrochemically etched by 40 wt.% KOH solution for optical microscopy observation. Tensile tests were performed at room temperature with specimens having a gage length of 25 mm and diameter of 5 mm according to the National Standard of the P.R.C., GB/T228-2002, the specimen gage section was oriented parallel to the rolling direction. Charpy impact tests were performed at 20 °C, and the direction of the V-notch was oriented perpendicular to the rolling direction. The volume fractions of austenite phase were measured using the method of manual point count according to ASTM E 562 as follows : the magniﬁcation of the micrograph was 500 and the grid size (number of points) was 20. Any point that fell on the phase studied was counted as one, otherwise zero. In addition,
Table 1 The chemical composition of hot rolled 19% Cr duplex stainless steel (wt.%). Elements
Alloy 1 Alloy 2 Alloy 3
0.01 0.01 0.01
0.13 0.11 0.12
3.59 5.54 8.10
0.006 0.005 0.005
0.007 0.007 0.006
19.34 19.43 19.55
1.56 1.59 1.55
0.98 0.92 0.87
0.21 0.22 0.20
Y. Yang et al. / Materials & Design 83 (2015) 820–828
Fig. 1. Optical metallographs showing the microstructures of specimens after solution treatment at 1040 °C, 1120 °C and 1220 °C: (a) 3.6 wt.% Mn, 1040 °C; (b) 8.1 wt.% Mn, 1040 °C; (c) 3.6 wt.% Mn, 1120 °C; (d) 8.1 wt.% Mn, 1120 °C; (e) 3.6 wt.% Mn, 1220 °C; and (f) 8.1 wt.% Mn, 1220 °C.
The best dual-phase structure can be obtained when Creq and Nieq meet the condition as follows :
½Creq ½Nieq þ 11:59 < 7
Fig. 2. Austenite volume fraction variation for specimens with different Mn after solution treatment at 1040 °C, 1120 °C and1220 °C.
Creq ¼ Cr þ 1:8Mo þ 2Nb þ 2:5Si
Nieq ¼ Ni þ 30ðC þ NÞ þ 0:5Mn
The Creq value of alloy 1, alloy 2 and alloy 3 is 21.43, 21.36 and 21.42, while the Nieq value of alloy 1, alloy 2 and alloy 3 is 9.96, 11.26 and 11.91 respectively. These values of Creq and Nieq highly meet Eq. (3), having a balance of dual-phase structure, which agree well with above microstructure analysis result. The X-ray diffraction results of the specimens with Mn content variation at solution treatment temperature of 1040 °C and 1220 °C are presented in Fig. 3. It is obviously seen that only austenite and d-ferrite phases were identiﬁed and no other precipitation phase peak was found in these spectra, indicating good solution effect of austenite formation elements (Ni, Mn, N) and ferrite formation elements (Cr, Mo) even with high Mn content addition (8.1 wt.%) in this type of DSS. When the Mn addition increased from 3.6% to 5.5 wt.% as shown in Fig. 3a and b, approximate major diffraction peaks intensity between d-ferrite and austenite can be observed under solution treatment at 1040 and 1120 °C, indicating good balance of two phases, which agree with microstructure analysis results. With Mn addition further increased to 8.1 wt.% (Fig. 3c), the intensity of austenite phase peaks appeared stronger
Y. Yang et al. / Materials & Design 83 (2015) 820–828
Fig. 3. X-ray diffraction results of the specimens with Mn content variation at different solution treatment temperature: (a) the specimens of alloy 1 solution treated at 1040 °C, 1120 °C; (b) the specimens of alloy 2 solution treated at 1040 °C, 1120 °C; and (c) the specimens of alloy 3 solution treated at 1040 °C, 1120 °C.
compared with d-ferrite phases, due to strengthened austenite effect of c-forming element Mn. Table 2 presents the concentrations and partition coefﬁcients (Ki) of major alloying elements in the specimens from different solution temperatures and Mn contents, where Ki = Xdi /Xci , and Xi d and Xci represent the average concentration of element i in the d and c phases respectively. As shown in Fig. 4, the element-partitioning of Cr, Ni between d ferrite and austenite phase became more uniform as the solution temperature increased from 1040 to 1220 °C, while the element-partitioning of Mn was consistent with this trend . Moreover, the values of partition coefﬁcient KMn were less than 0.90 with three different Mn contents addition at solution treatment temperatures of 1040 and 1120 °C, indicating better austenite effect than solution treatment at 1220 °C due to more Mn partitioning in austenite phase. It also
inferred that the alloy 2 with middle Mn addition content had lower value of partition coefﬁcient KMn than the alloy 1 and alloy 3 with different solution treatment temperatures, which can be attributed to great different chemical activity of the elements in two phases with different Mn contents. This indicates that better austenite effect was obtained with more Mn element solution in austenite phase.
3.2. Mechanical property analysis The error bar variety of impact data as a function of different Mn contents with as-rolled and different solution treatment temperatures are given in Fig. 5. Of all of the error-bars, the differences between the upper and lower limitations were relatively smaller for as-rolled and 1220 °C solution treated specimens, and the
Table 2 Concentrations and partition coefﬁcients (K) of major alloying elements in the specimens from different solution temperatures. Solution temperature (°C)
Cr concentration (wt.%) XdCr
Ni concentration (wt.%) XdNi
Mn concentration (wt.%) XdMn
Alloy 1 Alloy 2 Alloy 3
20.33 ± 0.31 21.28 ± 0.29 21.31 ± 0.28
18.72 ± 0.18 18.38 ± 0.11 18.40 ± 0.14
1.09 ± 0.03 1.16 ± 0.02 1.16 ± 0.02
1.41 ± 0.03 1.04 ± 0.02 1.27 ± 0.04
1.81 ± 0.04 1.92 ± 0.04 1.74 ± 0.02
0.78 ± 0.03 0.54 ± 0.02 0.73 ± 0.03
3.21 ± 0.04 4.63 ± 0.05 7.11 ± 0.07
3.75 ± 0.04 5.93 ± 0.05 8.44 ± 0.07
0.86 ± 0.03 0.78 ± 0.02 0.84 ± 0.02
Alloy 1 Alloy 2 Alloy 3
20.22 ± 0.23 21.39 ± 0.25 21.44 ± 0.27
18.85 ± 0.12 18.51 ± 0.14 18.56 ± 0.17
1.07 ± 0.02 1.16 ± 0.02 1.15 ± 0.03
1.45 ± 0.04 1.11 ± 0.02 1.33 ± 0.02
1.77 ± 0.03 1.87 ± 0.04 1.71 ± 0.03
0.82 ± 0.04 0.59 ± 0.02 0.78 ± 0.02
3.26 ± 0.03 4.7 ± 0.04 7.17 ± 0.06
3.70 ± 0.04 5.88 ± 0.06 8.39 ± 0.07
0.88 ± 0.02 0.80 ± 0.01 0.85 ± 0.02
Alloy 1 Alloy 2 Alloy 3
19.83 ± 0.18 20.87 ± 0.21 20.82 ± 0.22
19.11 ± 0.21 18.98 ± 0.15 18.92 ± 0.16
1.04 ± 0.02 1.10 ± 0.03 1.10 ± 0.02
1.53 ± 0.04 1.19 ± 0.02 1.41 ± 0.03
1.68 ± 0.02 1.79 ± 0.03 1.61 ± 0.02
0.91 ± 0.04 0.66 ± 0.02 0.88 ± 0.03
3.35 ± 0.03 4.79 ± 0.05 7.27 ± 0.06
3.62 ± 0.03 5.75 ± 0.05 8.24 ± 0.07
0.93 ± 0.02 0.83 ± 0.02 0.88 ± 0.02
Y. Yang et al. / Materials & Design 83 (2015) 820–828
Fig. 4. The partition coefﬁcient Ki of specimens with different Mn content after solution treated at 1040, 1120 and 1220 °C.
values of standard deviation became higher with increasing impact energy. In compared with as-rolled specimen, high level of impact toughness with more than 180 J impact energy was obtained for specimens solution treated at 1040 and 1120 °C, which can be attributed to strong thermally activated solid solution strengthening effect . Meanwhile, the impact energy was lowered by increasing solution treatment temperature to 1220 °C, which was related to less ductile austenite phase formation as shown by quantitative metallographic analysis in Fig. 2. For as-rolled and solution treatment of 1120 and 1220 °C specimens, as an increase of Mn addition from 3.6 wt.% to 5.5 wt.%, the toughness increased with higher impact energy values. As the manganese content further increased to 8.1 wt.%, the impact energy for as-rolled and 1120 °C solution treated specimens decreases slightly, but which for specimens solution treated at 1040 and 1220 °C obviously increased, indicating enhancement of toughness. This is because increasing manganese content causes an increase in the volume fraction of ductile austenite phase , thus improving impact toughness of alloys. The fracture surfaces of the studied steel after impact testing are shown in Fig. 6. At solution treatment of 1220 °C (Fig. 6a), the fracture mode of alloy 1 was mixed, with large areas consisting of ductile microvoid coalescence intermingled with
some areas of dispersed areas of shallow microvoids, with energy absorption of only 152 J. With solution treatment temperature decreased to 1120 and 1040 °C, the fracture surface of this low Mn content alloy was ductile and mainly consisted of small dimples, which agree with the Charpy impact toughness curve trend. Just as for alloy 2 with 5.5 wt.% Mn under solution treatment of 1120 °C, many small and deep dimples appeared on fracture surface, which is corresponding to high toughness. For alloy 3 with high Mn contents of 8.1 wt.%, the ductile dimpled fracture surface was observed for three different solution treatment temperature, indicating that relative high impact toughness was obtained. Consequently, these changing of fracture morphologies conﬁrmed the variations of toughness with different Mn addition and solution treatment temperatures. In order to investigate the effect of Mn addition on tensile behavior, the true stress–true strain curves at room temperature as a function of Mn content for the as-rolled and solution treatment of 1120 °C specimens are shown in Fig. 7. For the true strain above 0.15, the increment slope of the true stress decreased for solution treated specimen compared with as-rolled specimen. Meanwhile, No obvious yield behavior are observed in the shapes of curves with Mn addition higher than 3.6 wt.%, indicating strain hardening occurring with increasing Mn content , this is related to dislocation strengthening caused by volume fraction difference of two phases with more Mn addition . The region of uniform elongation for as-rolled and solution treated specimens became much shorter with Mn addition increased to 8.1 wt.%, showing comparatively low plasticity with more austenite formation, which can be attributed to larger deformation strains of austenite than that of ferrite . Moreover, the true strain increased with decreasing Mn content for this solution treated and as-rolled specimens. The tensile strength, yield strength (0.2% offset), and elongation to fracture (%) from theses tensile curves were summarized in Table 3. The effect of Mn on tensile property of as-rolled specimen was similar to the solution treated specimen. The tensile strengths and 0.2% offset yield strengths decrease slowly with increasing Mn content, and which decrease slightly by solution treatment of 1120 °C than as-rolled. The elongation to fracture (%) decreases slightly with Mn content increase from 3.6 to 5.5 wt.%, but which fall greatly with Mn addition increasing to 8.1 wt.%. This loss of ductility can be attributed to enhanced intrinsic hardening effect of high Mn addition on DSS . Moreover, The elongation to fracture of 47% attained with solution treatment of 1120 °C and 3.6 wt.% Mn addition, indicating a certain softening effect by solution treatment. 3.3. Corrosion resistance property
Fig. 5. The Charpy impact toughness of experimental alloys with Mn content variation under different solution treatment temperature.
The potentiodynamic polarization curves for specimens in tested solutions are shown in Fig. 8. Pitting potential Eb was observed for as-rolled and different Mn contents specimens under different solution treatment temperatures, indicating pit nucleation and repassivation . A great difference was found between the electrochemical behavior of 3.6 wt.% Mn alloy 1 and 5.5 wt.% Mn alloy 2 in 3.5 wt.% NaCl solution (Fig. 8a and b). It showed more positive breakdown potential and a wider passive region with decreasing solution temperature. But the 8.1 wt.% Mn alloy 3 with higher Mn addition exhibited less wide passive domain at these solution treatment temperature (Fig. 8c). Some report suggested that the pitting corrosion resistance equivalent number (PREN) decreased with more Mn addition, thus lowering the pitting corrosion resistance of stainless steel [30,31]. But this detrimental effect of Mn addition became different with different solution treatment temperature as shown in Table 4. It exhibited that Eb decreased with increasing Mn content from 3.6 wt.% to 5.5 wt.%, while the decrement of which became slower with higher Mn content of
Y. Yang et al. / Materials & Design 83 (2015) 820–828
Fig. 6. The fracture surfaces of the studied steel after impact testing at 20 °C for specimens solution treated at different solution temperatures: (a) alloy 1 solution treated at 1220 °C, (b) alloy 1 solution treated at 1120 °C, (c) alloy 1 solution treated at 1040 °C, (d) alloy 2 solution treated at 1120 °C, (e) alloy 3 solution treated at 1220 °C, (f) alloy 3 solution treated at 1120 °C, and (g) alloy 3 solution treated at 1040 °C.
8.1%, indicating high Mn content addition lowered pitting resistance to some degree. This pitting susceptibility with high Mn concentrations addition was associated with the formation of MnS inclusions due to strong afﬁnity of Mn with S . Under solution treatment at 1220 °C, the values of Eb approach as-rolled specimen with different Mn contents, which increase with decreasing solution treatment temperature to 1120 and 1040 °C, revealing that higher pitting corrosion resistance was achieved with lowering solution treatment temperature. This is involved two phase ratio and alloying elements partitioning variation in austenite and ferrite. The partition coefﬁcients (Ki) of major alloying elements analysis results in Table 2 showed that the KCr decreased with increasing solution treatment temperatures, indicating weakened repassivation ability in ferrite phase compared with in austenite phase, thus more ferrite formation at higher solution treatment
deteriorate pitting corrosion resistance. The passivation current densities, ip, is very similar to different Mn contents with solution treatment from 1040 to 1220 °C, which revealed that more Mn contents additions had little effect on overall passivation ability of two phases. Fig. 9 showed the morphology of corrosion surface after potentiodynamic polarization testing with different Mn contents under solution treatment temperature of 1220 °C and 1120 °C. It was found that the characteristic mechanism of pitting corrosion for specimens solution treated at 1220 °C was selective corrosion (Fig. 9a, c and d). The pitting corrosion was initiated at the interface between the c-phase and the d-phase, and was then propagated into d-phase resulting in deep and elongated stable pits formation. This can be attributed to lower PREN values in ferrite than that in austenite at high annealing temperature [33,34]. However, as
Y. Yang et al. / Materials & Design 83 (2015) 820–828 Table 3 Tensile test results of the alloys for the as-rolled and solution treatment of 1120 °C specimens.
Fig. 7. Tensile curves of the alloys with different Mn addition for the as-rolled and solution treatment of 1120 °C specimens.
solution treatment temperature was decreased to 1120 °C, the pits size in ferrite get smaller, and some pits appeared in austenite phase, which was related to element partitioning in two phases. As shown in previous Fig. 4 analysis, the chromium partitioning coefﬁcients KCr increased with decreasing solution treatment
Elongation to fracture (%)
Heat treatment condition
Tensile strength (MPa)
Yield strength (MPa)
Alloy 1, as-rolled, 3.6 wt.% Mn Alloy 2, as-rolled, 5.5 wt.% Mn Alloy 3, as-rolled, 8.1 wt.% Mn Alloy 1, solution treated at 1120 °C, 3.6 wt.% Mn Alloy 2, solution treated at 1120 °C, 5.5 wt.% Mn Alloy 3, solution treated at 1120 °C, 8.1 wt.% Mn
950 ± 31
715 ± 17
38 ± 1.9
830 ± 28
612 ± 16
42.5 ± 2.1
790 ± 19
580 ± 18
29 ± 1.6
765 ± 18
614 ± 21
47 ± 1.7
693 ± 17
607 ± 17
37.8 ± 1.5
732 ± 27
571 ± 15
27 ± 1.2
temperature, indicating a decrease of Cr content in austenite deteriorate the repassivation ability. In addition, the decreasing of solution treatment temperature also increased the volume fraction of austenite phases, increasing chance of pitting attack in austenite. It is clearly shown that the number of pits increased with increasing Mn content from 3.6 wt.% to 5.5 wt.% and 8.1 wt.% Mn, which was consistent with the decreasing of pitting potential Eb, indicating that higher Mn addition increased pitting susceptibility. This was mainly related to negative effect of Mn on PREN values, more
Fig. 8. Potentiodynamic polarization curves of specimens for as-rolled and different Mn contents specimens under different solution treatment temperatures: (a) alloy 1 with Mn content of 3.6 wt.%; (b) alloy 2 with Mn content of 5.5 wt.%; and (c) alloy 3 with Mn content of 8.1 wt.%.
Y. Yang et al. / Materials & Design 83 (2015) 820–828 Table 4 Pitting and passivation values of the materials tested in 3.5% NaCl solution at different Mn contents. Heat treatment condition
Eb (mV vs. SCE)
ip (lA cm2)
icorr (lA cm2)
Ecorr (mV vs. SCE)
Alloy Alloy Alloy Alloy Alloy Alloy Alloy Alloy Alloy Alloy Alloy Alloy
311 ± 8.2 244 ± 9.1 284 ± 7.9 341 ± 9.2 235 ± 6.9 185 ± 5.3 346 ± 10.1 311 ± 7.8 315 ± 6.9 461 ± 9.7 347 ± 8.6 309 ± 8.2
2.5 ± 0.14 1.95 ± 0.12 4.30 ± 0.19 2.45 ± 0.16 2.09 ± 0.17 2.39 ± 0.12 2.58 ± 0.18 2.50 ± 0.16 3.84 ± 0.21 1.69 ± 0.19 2.39 ± 0.15 2.69 ± 0.22
1.99 ± 0.09 1.25 ± 0.07 2.23 ± 0.11 2.12 ± 0.08 1.28 ± 0.07 1.82 ± 0.08 1.53 ± 0.06 1.99 ± 0.09 2.19 ± 0.11 1.51 ± 0.07 2.08 ± 0.08 2.24 ± 0.06
452.3 ± 9.6 516 ± 13.2 549 ± 12.5 380.6 ± 9.9 456 ± 8.8 443 ± 12.1 476 ± 11.6 452.3 ± 9.7 605 ± 17.3 462 ± 10.8 436 ± 9.3 509 ± 14.2
1, 2, 3, 1, 2, 3, 1, 2, 3, 1, 2, 3,
as-rolled, 3.6 wt.% Mn as-rolled, 5.5 wt.% Mn as-rolled, 8.1 wt.% Mn solution treated at 1220 °C, 3.6 wt.% Mn solution treated at 1220 °C, 5.5 wt.% Mn solution treated at 1220 °C,8.1 wt.% Mn solution treated at 1120 °C, 3.6 wt.% Mn solution treated at 1120 °C, 5.5 wt.% Mn solution treated at 1120 °C, 8.1 wt.% Mn solution treated at 1040 °C, 3.6 wt.% Mn solution treated at 1040 °C, 5.5 wt.% Mn solution treated at 1040 °C, 8.1 wt.% Mn
Fig. 9. The morphology of corrosion surface after potentiodynamic polarization testing with different Mn contents under solution treatment temperature of 1220 °C and 1120 °C: (a) alloy 1 with 3.6 wt.% Mn and solution treated at 1220 °C, (b) alloy 1 with 3.6 wt.% Mn and solution treated at 1120 °C, (c) alloy 2 with 5.5 wt.% Mn and solution treated at 1220 °C, (d) alloy 2 with 5.5 wt.% Mn and solution treated at 1120 °C, (e) alloy 3 with 8.1 wt.% Mn and solution treated at 1220 °C, and (f) alloy 3 with 8.1 wt.% Mn and solution treated at 1120 °C.
Mn addition can lead to the decrease of PREN values of two phases. The lower the value of PREN is, the worse the pitting corrosion resistance can be.
4. Conclusions With solution temperatures ranging from 1040 to1220 °C, the inﬂuence manganese alloying elements on microstructure, mechanical property and pitting corrosion of 19% Cr economical
duplex stainless steel were examined. For the solution treated specimens where the balanced two phases were kept by Mn addition. In addition, the amount of austenite phases increased with increasing Mn content and decreased by increasing solution treatment temperature. The impact energy at 20 °C increased with decreasing solution temperature from 1220 °C to 1040 and 1120 °C, and improved by more Mn addition due to more ductile austenite phases formation. This toughness variations were consistent with morphology characteristic of the fracture surface changing. The effect of more Mn addition and solution treatment of
Y. Yang et al. / Materials & Design 83 (2015) 820–828
1120°C on decreasing of tensile strength and 0.2% offset yield strength were slight, whereas the elongation to fracture (%) fell greatly with Mn addition increasing to 8.1 wt.% for as-rolled and solution treated specimens due to larger deformation strains of austenite than that of ferrite. The decreasing trend of pitting corrosion potential became slower with Mn content from 3.6 to 8.1 wt.%. The pitting corrosion mainly occurs in the ferrite phases in the case of high solution temperature (1220 °C), but it happens in two phases in the case of low solution temperature (1120 °C). The pitting corrosion resistance decreased by increasing solution treatment temperature due to more weakened repassivation ferrite phases formation. Acknowledgement This work is supported by the National Science Foundation of China (Grant No. 51261010). References  K.N. Adhe, V. Kain, K. Madangopal, H.S. Gadiyar, Inﬂuence of sigma-phase formation on the localized corrosion behavior of a duplex stainless steel, J. Mater. Eng. Perform. 5 (1996) 500–506.  A. Kashiwar, N. Phani Vennela, S.L. Kamath, R.K. Khatirkar, Effect of solution annealing temperature on precipitation in 2205 duplex stainless steel, Mater. Charact. 74 (2012) 55–63.  H. Sarlak, M. Atapour, M. Esmailzadeh, Corrosion behavior of friction stir welded lean duplex stainless steel, Mater. Des. 66 (2015) 209–216.  M. Sadeghian, M. Shamanian, A. Shafyei, Effect of heat input on microstructure and mechanical properties of dissimilar joints between super duplex stainless steel and high strength low alloy steel, Mater. Des. 60 (2014) 678–684.  J.K.L. Lai, K.W. Wong, D.J. Li, Effect of solution treatment on the transformation behaviour of cold-rolled duplex stainless steels, Mater. Sci. Eng., A 203 (1) (1995) 356–364.  V.S. Moura, L.D. Lima, J.M. Pardal, A.Y. Kina, R.R.A. Corte, S.S.M. Tavares, Inﬂuence of microstructure on the corrosion resistance of the duplex stainless steel UNS S31803, Mater. Charact. 59 (8) (2008) 1127–1132.  Tuba Karahan, Hayriye Ertek Emre, Mustafa Tümer, Ramazan Kacar, Strengthening of AISI 2205 duplex stainless steel by strain ageing, Mater. Des. 55 (2014) 250–256.  J.K.L. Lai, K.W. Wong, D.J. Li, Effect of solution treatment on the transformation behavior of cold-rolled duplex stainless steels, Mater. Sci. Eng., A A203 (1995) 356–364.  S.l. Li, Y.l. Wang, H.l. Zhang, S.X. Li, G.Q. Wang, X.T. Wang, Effects of prior solution treatment on thermal aging behavior of duplex stainless steels, J. Nucl. Mater. 441 (2013) 337–342.  A. Belfrouh, C. Masson, D. Vouagner, A.M. Debecdelievre, N.S. Prakash, J.P. Audouard, The cumulative effect of alloying elements N, W, Mo and Cu on the corrosion behaviour of 17Cr–13Ni stainless steel in 2NH2SO4, Corros. Sci. 38 (1996) 1639–1643.  G. Lothongkum, P. Wongpanya, S. Morito, Effect of nitrogen on corrosion behavior of 28Cr–7Ni duplex and microduplex stainless steels in air-saturated 3. 5 wt% NaCI solution, Corros Sci. 48 (1) (2006) 137–153.
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