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Strengthening behaviors of V and W modified Cr19 series duplex stainless steels with transformation induced plasticity
Author’s Accepted Manuscript Strengthening behaviors of V and W modified Cr19 series duplex stainless steels with transformation induced plasticity Huawei Zhang, Yulai Xu, Pengfei Hu, Wanjian Xu, Jun Li, Kunfang Li, Xueshan Xiao www.elsevier.com/locate/msea
PII: DOI: Reference:
S0921-5093(17)31084-5 http://dx.doi.org/10.1016/j.msea.2017.08.071 MSA35419
To appear in: Materials Science & Engineering A Received date: 8 May 2017 Revised date: 16 August 2017 Accepted date: 17 August 2017 Cite this article as: Huawei Zhang, Yulai Xu, Pengfei Hu, Wanjian Xu, Jun Li, Kunfang Li and Xueshan Xiao, Strengthening behaviors of V and W modified Cr19 series duplex stainless steels with transformation induced plasticity, Materials Science & Engineering A, http://dx.doi.org/10.1016/j.msea.2017.08.071 This is a PDF file of an unedited manuscript that has been accepted for publication. As a service to our customers we are providing this early version of the manuscript. The manuscript will undergo copyediting, typesetting, and review of the resulting galley proof before it is published in its final citable form. Please note that during the production process errors may be discovered which could affect the content, and all legal disclaimers that apply to the journal pertain.
Strengthening behaviors of V and W modified Cr19 series duplex stainless steels with transformation induced plasticity Huawei Zhang1, Yulai Xu1, 2, *, Pengfei Hu1, Wanjian Xu1, Jun Li1, 2, Kunfang Li1, Xueshan Xiao1, 2 1. Institute of Materials, Key Laboratory for Microstructures, Shanghai University, Shanghai 200072, China 2. Shanghai University Xinghua Institute of Special Stainless Steels, Jiangsu 225721, China
Abstract: The V and W modified Cr19 series duplex stainless steels with transformation induced plasticity (TRIP) have been newly developed and the effects of microstructures on room temperature mechanical properties have been investigated. The addition of ferrite forming elements V and W slightly increases the volume fraction of ferrite phase from about 55 to 57-60%, however, the cold rolling process hardly influence the fraction. When the duplex phases exhibit a coarse banding morphology before cold rolling, the ultimate tensile strength and elongation are improved significantly with the addition of V and W elements mainly due to the occurrence of TRIP effect. The ultimate tensile strength and yield strength of Cr19+V increase very slightly, but the elongation increases to about 61% after cold rolling. Both the ultimate tensile strength and elongation of Cr19+W significantly increase to about 800 MPa and 67%, respectively. The improvement is attributed to TRIP effect and refinement of banded ferrite and austenite phases. After aging heat treatment, the precipitate of VN and Cr23C6 particles contributes to the further increase of ultimate tensile strength to about 1000 and 920 MPa of Cr19+V and Cr19+W DSSs, respectively. Strain induced αʹ -martensite is transformed from austenite directly with an orientation relationship. The strengthening behaviors have been discussed based on the microstructural evolutions. Key words: Duplex stainless steels, Transformation induced plasticity, Mechanical properties, Microstructures, Strengthening behaviors.
*
Corresponding author E-mail address: [email protected] 1
1. Introduction Duplex stainless steels (DSSs) composed of body-centered cubic ferrite (α) and face-centered cubic austenite (γ) exhibit good mechanical properties and corrosion resistance [1-3]. In order to reduce weight and improve safety, the TRIP steels show great potential for the application in automotive industry due to the good combination of strength and ductility [4]. Considerable research efforts have been carried out to improve both mechanical and corrosion properties, particularly by controlling alloying elements. The substitution of the costly alloying elements such as Ni and Mo with N and Mn makes the austenite phase more prone to the occurrence of the TRIP effect [5]. Ran et al. [6] had investigated the influence of heat treatment on TRIP effect of Fe-19.1Cr-1.3Ni-6.2Mn-1.0Mo-0.173N-0.027C-0.37Si and found that the ultimate tensile strength and elongation are about 1100 MPa and 50%, respectively. Herrera et al. [7] had designed a ductile
economical
Mn-based
duplex
stainless
TRIP
steel
Fe-19.9Cr-0.42Ni-0.16N-4.79Mn-0.11C-0.46Cu-0.35Si with ultimate tensile strength of 1000 MPa and an elongation value higher than 60%. The Mn element is often added into the DSSs to increase the solubility of N which gives potent solid solution strengthening and improves the resistance to pitting corrosion [8]. However, Jang et al. [9] had investigated the tensile and corrosion behavior of CD4MCU cast DSSs with different Mn contents, and found that the addition of 0.8 wt. % Mn was detrimental to both pitting corrosion and stress corrosion cracking properties. Yang et al. [10] had studied the effect of Mn on microstructure, mechanical and pitting corrosion resistant properties of economical Cr19 DSS with solution temperatures ranging from 1040 to 1220 °C, which indicated that the impact energy at 20 °C increased to more than 180 J when the solution temperature decreased from 1220 to 1040 and 1120 °C, and that the impact energy could be improved by Mn addition due to increased amount of austenite phase. But it also exhibited that pitting potential decreased with the increased Mn content from 3.6 to 8.1 wt. %. Interstitial elements such as carbon and nitrogen are always employed to form precipitates which pin the movement of dislocations to achieve good mechanical properties. Choi et al. [11] had investigated the effect of nitrogen content on TRIP effect of the new series of nearly Ni-free DSSs and found that the TRIP effect led to the enhancement of ultimate tensile strength over 1000 MPa and 2
ductility of about 60% for 0.35 wt. % N DSSs. Fraction of strain induced martensite increased with the N content resulting in a better strain hardening. Under severe corrosion conditions, sufficient amount of chromium, molybdenum and nitrogen elements are needed to obtain a high pitting resistance equivalent value. Zhang et al. [12] had designed a new series of Mo-free 21.5Cr-3.5Ni-xW-0.2N economical DSSs and shown that the ultimate tensile strength decreased from about 900 to 790 MPa, but the elongation increased from 35 to 61% and the pitting corrosion potential increased from 760 to 1011 mV with the increase of W content, which indicated that the ductility and corrosion resistance property can be improved with the addition of W element. Tensile properties of DSSs also have great relationship with the volume fraction of ferrite phase and the shape of austenite phase. The chemical compositions, heat treatment steps, and pre-treatment processes are key factors in influence the phase ratios. The DSSs show high strength and resistance to corrosion when the ferrite to austenite phase ratio is approximately 1:1 [13-15]. Solution heat treatment promotes dissolution of carbides and other precipitates after rolling process, which is an important process for DSSs to control the suitable phase ratio and hence obtain a good combination of mechanical properties and corrosion resistance. Lai et al. [16] studied the effect of solution heat treatment on the transformation kinetics of γ to α in DSSs at the temperature range of 1050-1250 °C and found that higher temperature increased the amount of ferrite and modified the morphology of ferrite phase in the matrix. The effects of precipitate strengthening which caused by dislocation and interstitial carbon or nitrogen atoms may be reduced during this period. The mechanical properties of DSSs have great relationship with deformation modes of ferrite and austenite phases. Plastic deformation of ferrite is mainly dominated by movement of dislocation due to high stacking fault energy [17]. The deformation of austenite includes transformation induced plasticity, twinning induced plasticity and planar glide of dislocations with the increase in stacking fault energy [18, 19]. The influence of V and W elements on mechanical properties of Cr19 series DSSs has been hardly investigated, and the new Cr19 series DSSs with additions of V and W elements have been designed to obtain a possible good combination of strength × elongation value and corrosion resistance property. The newly developed Cr19 series DSSs with 0.2 wt. % N without the addition of Mn element have been prepared, and this paper aims to study the effects of V and W elements on microstructure 3
evolutions and mechanical properties, and the strengthening behaviors after different deformation and heat treatments processes have also been discussed. 2. Experimental The Cr19 series DSSs modified with V and W elements were melted in a 50 kW induction furnace in an argon atmosphere. The chemical compositions of the castings designated as Cr19, Cr19+V, and Cr19+W are shown in Table 1. The casting ingots were hot forged into Φ 30 mm sticks, and the as-forged samples were solution heat treated at 1050 °C for 5 min followed by water quenching. The plates with the thickness of about 3.2 mm were cut from the sticks and then cold rolled to the sheets with the thickness of about 1.2 mm followed by solution heat treatment at 1050 °C for 5 min and water quenching. After that the samples are aged at 750 °C for 0, 0.5, 2, 4, and 6 h. The Charpy V-notch specimens with 10 mm × 10 mm × 55 mm were cut from the solution heat treated sticks along the forging direction according to P.R.C., GB/T229-2007 [20]. Charpy impact tests were carried out at 25, 0, and -40 ºC by using AHC3000/2-AT impact test machine with maximum capacity of 450 J. Polarization tests of the steels after hot-forging and solution heat treatment at 1050 °C were performed at least three times to ensure the reproducibility
of
the
results
by
using
an
EG&G
Princeton
Applied
Research
Potentiostat/Galvanostat Model 273A. Pitting corrosion samples were progressively wet ground up to 1200 grit SiC papers and then well rinsed in water, and the edges of the exposed samples were mounted with an epoxy resin. Then the samples with an area of 1 cm2 were tested in 1 mol/L NaCl solution at 25 ºC. The saturated calomel electrode (SCE) was used as a reference electrode, a platinum foil was served as the counter electrode, and open circuit potential (VOC) measurements were carried out for 2500 s. The polarization scan started at -0.4VSCE below VOC until 1.0VSCE with a scan rate of 1 mV/s. The specimens were electrochemically etched by 10 wt. % KOH solution and the optical micrographs were observed by KEYENCE VHX-100 microscopy. The volume fraction of ferrite was measured by quantitative metallographic analysis systems and the average value of ten measurements on each sample was taken as the fraction of ferrite. The fracture surface morphologies were analyzed by using HITACHI SU-1510 scanning electron microscope (SEM). 4
The tensile tests were carried out at room temperature according to National Standard of the P.R.C. GB/T228.1-2010 with the specimens having a gauge length of 30 mm, width of 10 mm, and thickness of 1 mm [21]. The microstructures of the samples after tensile tests were analyzed by JEM-2010F transmission electron microscope (TEM). The specimens were sliced from the tensile plates along the stress direction near the fracture surface, polished with abrasive papers from no. 800 to 2000, and thinned using a twin-jet electro-polishing in the solution composed of 10 vol. % HClO4 and 90 vol. % C2H5OH at -30 ºC and 45 V. 3. Results and discussions The stress-strain curves of hot-forged Cr19 series DSSs after solution heat treatment at 1050 °C for 5 min are shown in Fig. 1. The ultimate tensile strength significantly increases from about 720 MPa of Cr19 to 805 MPa of Cr19+V and 750 MPa of Cr19+W, and the corresponding elongation value also increases from about 43 to 52 and 57%. Addition of V element increases the yield strength from about 545 MPa of Cr19 to 590 MPa of Cr19+V. Although the yield strength decreases from about 590 MPa of Cr19+V to 520 MPa of Cr19+W, the corresponding elongation value increased from about 52 to 57%. It is evident that the addition of V and W elements is beneficial to room temperature tensile mechanical properties. Fig. 2 shows the stress-strain curves of cold-rolled Cr19+V and Cr19+W after solution heat treatment at 1050 °C for 5 min and aging at 750 °C for up to 6 h. It is clear that both the ultimate tensile strength and yield strength of Cr19+V slightly increase by about 10 MPa compared with those before cold rolling, but both the ultimate tensile strength and yield strength of Cr19+W increase by about 70 MPa. The ultimate tensile strength increases while the elongation decreases significantly with the increased aging time of Cr19+V and Cr19+W DSSs. The ultimate tensile strength remains at about 1000 MPa with the corresponding elongation value about 23% after aging over 2 h for Cr19+V. However, the elongation keeps the value higher than 38% even after aging for 6 h for Cr19+W with the improved ultimate tensile strength of about 920 MPa. It can be found that mechanical properties of Cr19+V and Cr19+W get better significantly after cold rolling, because the elongation value increases from about 52% to 61% for Cr19+V although the ultimate tensile strength increases very slightly from about 805 to 815 MPa, while both the ultimate tensile 5
strength and elongation respectively increase from about 750 and 57% to 800 MPa and 67% of Cr19+W. So the cold rolling process improves both the tensile strength and ductility at the same time. The strength × elongation (MPa%) values of cold rolled V and W modified Cr19 DSS after solution treatment at 1050 °C for 5 min are respectively about 49715 and 53600 MPa%, which indicates that the modified Cr19 series DSSs obtain a very good combination of strength × elongation value. The
combination
of
strength
Fe-19.1Cr-1.3Ni-6.2Mn-1.0Mo-0.173N-0.027C-0.37Si
×
elongation
value
of and
Fe-19.9Cr-0.42Ni-0.16N-4.79Mn-0.11C-0.46Cu-0.35Si DSSs are respectively about 55000 MPa% and 66300 MPa% according to Ran et al. [6] and Herrera et al. [7], and these values are very high at present. However, the addition of Si and high content of C may be harmful to the impact toughness and pitting corrosion resistant properties. At the same time, the high Mn contents may also do harm to the pitting corrosion resistant property. Fig. 3 shows the impact energy at 25, 0, and -40 ºC of the hot-forged Cr19, Cr19+V, Cr19+W DSSs after solution heat treatment at 1050 °C for 5 min. It can be obtained that addition of V reduces the impact energy of Cr19 DSS especially at -40 °C. Although W reduces the energy at 25 °C, the impact energy hardly varies at 0 °C and the value even increases from about 128 to 145 J at -40 °C. Fig. 4 shows the SEM fracture surface morphologies after impact test at 0 °C and -40 °C. Large amounts of ductile dimples uniformly distribute on the fracture surface of Cr19 DSS, which indicates a good toughness (Fig. 4(a)). The typical SEM morphologies vary slightly of Cr19+V and Cr19+W DSSs (Fig. 4(b) and (c)). The amount of ductile dimples significantly reduces at -40 °C for both the Cr19 and Cr19+V DSSs (Fig. 4(d) and (e)). However, it should be noticed that the fracture surface morphology of Cr19+W shows a typical ductile fracture at -40 °C because both the amount and size of the dimples increases and decreases respectively compared with that of Cr19 DSS at 0 °C. Fig. 5 shows the representative polarization curves of Cr19 series DSSs after hot-forging and solution heat treatment at 1050 °C. The pitting corrosion potential (Epit) is about 700 mV for Cr19 DSS, and Epit respectively increases to about 820 and 930 mV for Cr19+V and Cr19+W DSSs. It has been suggested that pitting corrosion resistance depend basically on Cr, Mo and N contents, and resistance to pitting corrosion can be characterized in terms of a pitting resistance equivalent 6
(PRE) value, which is often defined as [22, 23]: PRE = wt. % Cr + 3.3 (wt. % Mo + 1/2 wt. % W) + 30 wt. % N
(1)
The PRE values of the experimental alloys calculated according to Eq. (1) are respectively about 30.7, 30.8, and 32.1. It can be seen that the PRE values increase slightly due to the chemical composition vibration and addition of W element, which is in good agreement with the results of the pitting corrosion test. So the yield strength and elongation values of the newly developed Cr19 series DSSs are high, the impact toughness, and pitting corrosion resistance is also very good especially for Cr19+W steel. Fig. 6 shows the typical optical micrographs of the hot forged and cold-rolled specimens after solution heat treatment at 1050 °C. The typical duplex phases of ferrite and austenite have been identified. The volume fraction of ferrite phase which has been measured is respectively about 55, 57, and 60% for Cr19, Cr19+V, and Cr19+W DSSs (Fig. 6(a)-(c)). So it confirms that both the V and W are ferrite forming elements. Fig. 6(d)-(f) shows that both the thickness of the banded ferrite and austenite has been refined dramatically along the rolling direction, but the volume fraction of ferrite phase hardly changes compared with that before cold rolling. The TEM micrographs of hot-forged Cr19+V DSS after solution heat treatment at 1050 °C for 5 min and tensile tests are shown in Fig. 7. The elongated austenite phase has been observed in Fig. 7(a) and the αʹ -martensite has also been identified among the austenite phase (Fig. 7(b)). The corresponding selected area diffraction pattern (SADP) in Fig. 7(c) shows that the face centered cubic austenite phase and the near cubic αʹ -martensite phase co-exist following the orientation relationship of <220>γ // <211>αʹ , which confirms that formation of αʹ -martensite is due to the deformation of austenite along tensile direction. Lots of precipitates marked with dotted circles in Fig. 7(d) have been identified in the matrix, and dislocation lines distribute around these particles. The corresponding TEM-SADP (Fig. 7(e)) and TEM-EDS (Fig. 7(f)) indicate that these precipitates belong to VN particles. Saenarjhan et al. [24] also pointed out that formation of the ɛ -martensite wasn’t a pre-requisite for transforming austenite into αʹ -martensite and suggested that austenite directly transformed to αʹ -martensite during deformation rather than via the deformation bands due to a relatively high driving force for αʹ -martensite transformation in Fe-0.17N-0.02C-20.5Cr-2.0Ni-1.8Mn-0.7Cu-0.6Mo-0.5Si. At the tensile strain of 30%, the 7
αʹ -martensite appeared in austenite with <110>γ // <001>αʹ orientation relationship. For specimen deformed up to failure, a large number of dislocations and stacking faults were observed and both of Nishiyama-Wasserman orientation relationship and Kurdjumov-Sachs orientation relationship of <110>γ // <111>αʹ existed between austenite and αʹ -martensite. In addition, Zhang and Hu [25] demonstrated that deformation induced martensite transformation (DIMT) followed γ →
hexagonally
indexed
deformation
bands
→
αʹ
in
0.14N-0.02C-20Cr-2.1Ni-5.1Mn-0.2Cu-0.3Mo-0.4Si DSS which had a higher concentration of Mn. Deformation bands appeared at the early stage of deformation and αʹ -martensite formed at the intersection of the deformation bands later. Calculation showed that the transformation driving force of austenite into ɛ -martensite was always positive and much bigger than that of austenite into αʹ -martensite because higher Mn content significantly reduced driving force of αʹ -martensite transformation. Therefore, the DIMT sequence seemed to be determined by the value of driving force. But no ɛ -martensite has been identified in Cr19+V steel according to TEM observations. Fig. 8 shows the TEM micrographs of hot-forged Cr19+W after solution heat treatment at 1050 °C for 5 min and tensile tests. The typical TEM-EDS of “Region 1” and “Region 2” in Fig. 8(a) is shown in Fig. 8(b), and W element which has been identified confirms that addition of W mainly solutes in the matrix. The SADPs of “Region 1” and “Region 2” in Fig. 8(a) are respectively shown in Fig. 8(c) and 8(d). Only αʹ -martensite exists at “Region 1” while both γ-austenite and αʹ -martensite are found at “Region 2” following the orientation relationship of <110>γ // <110>αʹ , which also proves the formation of strain induced martensite due to the TRIP effect. It is known that the austenite with different stability has two different transformation paths during deformation controlled by elemental partitioning and austenite grain size distribution. One way is a direct transformation from austenite to αʹ -martensite in medium Mn-TRIP steels [26-28] and quenching and partitioning steels [29, 30], and another way through an intermediate ε-martensite and subsequently the formation of αʹ -martensite in Mn-N DSS [31, 32] and high-Mn austenitic steels [33]. Now that ɛ -martensite wasn’t found in the Cr19 series DSSs, αʹ -martensite was transformed from austenite directly. Fig. 9 shows the typical TEM micrographs of Cr19+V and Cr19+W near the fracture surface 8
after aging at 750 °C for 2 h and tensile tests. The approximately spherical precipitate with the size about 70 nm had been observed in the grain interior in Fig. 9(a), both the corresponding TEM-EDS and TEM-SADP in Fig. 9(b) and (c) confirm that this precipitate belongs to VN. The size of VN in the aged specimen is bigger than that found in Fig. 7(d) due to the coarsening of VN particles to minimize the interfacial energy [34]. A large precipitate had been observed in “Region 1” from the bright filed TEM image of aged Cr19+W in Fig. 9(d), the length and width are respectively about 1200 and 630 nm. The corresponding TEM-EDS pattern shown in Fig. 9(e) indicates that the large particle is enrich of Cr element which can be identified as Cr23C6 carbide by TEM-SADP in Fig. 9(f). Dong et al. [35] showed that a considerable amount of M7C3 carbides transformed into M23C6 carbides while M23C6 continued to grow up when the carbon partitioning time was extended to 15 min at 600 °C in Nb-V-Ti microalloyed ultra-high strength steel. In the studies of Kucharova and Shankar et al. [36, 37], W mostly enriched in M6C and played an important role in suppressing the coarsening of Cr23C6. But no peak of W element had been observed from the EDS patterns, which confirms that W element still solutes in the matrix during aging heat treatment. The SADP of “Region 2” is shown in Fig. 9(g), which proves the existence of αʹ -martensite because of TRIP effect during tensile tests. The above results indicate that addition of V and W elements has great influence on room temperature tensile mechanical properties (Fig. 1), impact toughness (Fig. 3), and pitting corrosion resistant property (Fig. 5), because the V and W elements change the microstructures of the TRIP-assisted Cr19 series DSSs. After hot forging and solution heat treatment, the morphologies of the duplex phase vary slightly except that the volume fraction of ferrite phase increases from about 55 to 57-60% (Fig. 6). The TEM observations also indicate that the αʹ -martensite has been identified in austenite in both the V and W modified DSSs and the αʹ -martensite exhibit an orientation relationship with γ, and VN has been observed in the matrix in Cr19+V steel (Figs. 7 and 9). So the significantly increased elongation from about 43% of hot-forged Cr19 to higher than 52% of hot-forged Cr19+V and Cr19+W DSSs is primarily due to the easier occurrence of TRIP effect during deformation, because addition of ferrite forming elements V and W reduces the stability of austenite. Among the three alloys, Cr19+W shows the highest ferrite fraction and corresponding the least stability of austenite, which leads to the best elongation. Increased tensile 9
strength of Cr19+V results from the TRIP effect, precipitate strengthening of VN particles which hinders the dislocation moving, and solid solution strengthening of V element. However, no carbide particles have been observed in the tensile samples of Cr19+W after solution heat treatment, and the increased tensile strength of Cr19+W is attributed to the TRIP effect because of the phase transformation of αʹ -martensite from austenite, and to the solid solution strengthening of W because the W element mainly solute in the matrix according to the TEM-EDS patterns. The Charpy impact test shows that all the DSSs exhibit good toughness because the impact energy is 200 ± 15 J at 25 °C, the impact energy hardly varies with the addition of W element, and the most important one is that the low temperature impact toughness of Cr19+W is the best among the three steels. Wan et al. [38] showed that impact toughness of 15Cr DSSs decreased with Al content at the same temperature mainly resulting from the ferrite-austenite ratio in duplex phase structure. Austenite phase has a better toughness than that of ferrite phase, so ferrite phase is the phase that exhibits ductile-brittle transition (DBT) phenomenon in DSSs. Pecker et al. [39] also suggested that higher impact energy at 0 °C and -40 °C of 22Cr-2.0Ni was because of the increased stacking fault energy. Fig. 3 shows that the impact energy of Cr19+V is slightly lower than that of Cr19, especially at low temperatures. The reduced toughness of Cr19+V steel may be primary due to the slightly increased volume fraction of ferrite phase and precipitate of VN particles. As to the Cr19+W steel, no precipitate has been identified in the matrix which may benefit the improved toughness at low temperatures. After cold rolling and heat treatment, the ultimate tensile strength, yield strength, and elongation further increase to a very high level. The optical micrographs indicate that the volume fraction of ferrite hardly varies compared with that before cold rolling, but the duplex phases exhibit a banded morphology and the thickness dramatically reduces (Fig. 6). Hence, the enhancement of room temperature tensile mechanical properties after cold rolling is primarily due to the refinement of both ferrite and austenite phases compared with those after hot forging. The reduced thickness of the phases significantly contributes to the increased elongation of Cr19+V and the improved tensile strength and elongation of Cr19+W DSS. So the reduced thickness of the duplex phase does not play an important role in improving the tensile strength because the tensile strength is in a relatively high level due to the precipitate strengthening of VN nano-sized particles. 10
However, the reduced thickness great influence the tensile strength of Cr19+W steel because the strengthening mainly comprised of solid solution strengthening and TRIP effect. After aging heat treatment, the tensile strength and ductility of Cr19+V is respectively stable at about 1000 MPa and 23%. The improved tensile strength of Cr19+V and Cr19+W after aging heat treatment is primary due to the precipitate of VN and Cr23C6 carbides, respectively. Conclusion The microstructures, room temperature tensile mechanical properties, impact toughness, and pitting corrosion resistance of TRIP-assisted V and W modified Cr19 series DSSs have been investigated and the corresponding strengthening behaviors have also been discussed. Addition of ferrite forming elements V and W increase the volume fraction of ferrite from 55 to 57-60%, but the fraction hardly varies after cold rolling and following heat treatment. The V and W elements reduce the stability of austenite and result in an easier occurrence of TRIP effect of Cr19 series DSSs because the ultimate tensile strength and elongation can be improved obviously when the duplex phases exhibit a coarse banding morphology: the V element increases the ultimate tensile strength to 805 MPa with elongation of 52%, and W improves the ultimate tensile strength of 750 MPa with elongation about 57% after solution heat treatment at 1050 °C of the hot-forged Cr19 series DSSs. Addition of W increases the impact energy at -40 °C of hot-forged steels while V reduces impact energy of Cr19 DSSs due to the precipitate of VN particles. The improvement of room temperature mechanical properties before cold rolling is mainly due to the phase transformation strengthening. After cold rolling and heat treatment at 1050 °C, the improvement of tensile strength and elongation is mainly due to the TRIP effect and refinement of the banded duplex phases especially for the Cr19+W, because the ultimate tensile strength and yield strength of Cr19+V increase very slightly, while both the ultimate tensile strength and elongation of Cr19 +W increases to about 800 MPa and 67%, respectively. The significantly increased value of strength × elongation is primarily due to the phase refining strengthening and transformation strengthening. Strain induced αʹ -martensite is transformed from austenite directly without the formation of ɛ -martensite with an orientation relationship existing between austenite phase and strain induced martensite. 11
The precipitate strengthening plays an important role after aging heat treatment, the ultimate tensile strength of Cr19+V is promoted notably to about 1000 MPa due to the precipitation of VN particles, but both the yield strength and elongation dramatically decrease to 475-525 MPa and about 23%, respectively. The ultimate tensile strength of Cr19+W increases to as high as 920 MPa, but the yield strength and elongation respectively decrease by about 10-45 MPa and 22-29%, which is primary due to the precipitate of Cr23C6 carbide.
Acknowledgements Sponsored by Shanghai Sailing Program [17YF1405800] of Shanghai Municipality. Prospective Joint Research Project of Department of Science and Technology of Jiangsu Province [BY2015068-01].
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Table and Figure Captions Table 1 Chemical composition of V and W modified Cr19 series DSSs Fig. 1 Stress-strain curves of Cr19, Cr19+V and Cr19+W DSSs after hot forging and solution heat treatment at 1050 °C for 5 min Fig. 2 Stress-strain curves of Cr19+V and Cr19+W DSSs aged at 750 °C for 0, 0.5, 2, 4, and 6 h after annealing at 1050 °C for 5 min, (a) Cr19+V, (b) Cr19+W Fig. 3 Charpy impact toughness of hot-forged Cr19 series DSSs at 25, 0, and -40 ºC Fig. 4 SEM fracture surface morphologies of Cr19 series DSSs at 0 °C of (a) Cr19, (b) Cr19+V, (c) Cr19+W, and at -40 °C of (d) Cr19, (e) Cr19+V, (f) Cr19+W 15
Fig. 5 Polarization curves of Cr19 series DSSs in 1 mol/L NaCl at 25 °C Fig. 6 Optical micrographs of hot-forged (a) Cr19, (b) Cr19+V, (c) Cr19+W DSSs, and cold-rolled (d) Cr19, (e) Cr19+V, (f) Cr19+W DSSs after heat treatment at 1050 °C Fig. 7 TEM micrographs of Cr19+V near the fracture surface of tensile samples solution heat treated at 1050 °C for 5 min: (a) austenite, (b) αʹ -martensite, (c) TEM-SADP, (d) VN and dislocation lines, (e) TEM-SADP, and (f) TEM-EDS of VN Fig. 8 TEM micrographs of hot-forged Cr19+W after tensile tests: (a) bright field TEM image, (b) TEM-EDS, (c) TEM-SADP of “Region 1”, (d) TEM-SADP of “Region 2” Fig. 9 TEM micrographs near fracture surface of tensile samples after aging at 750 °C for 2 h and tensile tests: (a) bright field TEM image of Cr19+V, corresponding (b) TEM-EDS of VN, and (c) TEM-SADP, (d) bright field TEM image of Cr19+W, corresponding (e) TEM-EDS of Cr23C6, (f) TEM-SADP of Cr23C6 in “Region 1”, and (g) TEM-SADP of αʹ -martensite in “Region 2”
Fig. 1 Stress-strain curves of Cr19, Cr19+V and Cr19+W DSSs after hot forging and solution heat treatment at 1050 °C for 5 min
16
Fig. 2 Stress-strain curves of Cr19+V and Cr19+W DSSs aged at 750 °C for 0, 0.5, 2, 4, and 6 h after annealing at 1050 °C for 5 min, (a) Cr19+V, (b) Cr19+W
Fig. 3 Charpy impact toughness of hot-forged Cr19 series DSSs at 25, 0, and -40 ºC
17
Fig. 4 SEM fracture surface morphologies of Cr19 series DSSs at 0 °C of (a) Cr19, (b) Cr19+V, (c) Cr19+W, and at -40 °C of (d) Cr19, (e) Cr19+V, (f) Cr19+W
Fig. 5 Polarization curves of Cr19 series DSSs in 1 mol/L NaCl at 25 °
18
Fig. 6 Optical micrographs of hot-forged (a) Cr19, (b) Cr19+V, (c) Cr19+W DSSs, and cold-rolled (d) Cr19, (e) Cr19+V, (f) Cr19+W DSSs after heat treatment at 1050 °C
Fig. 7 TEM micrographs of Cr19+V near the fracture surface of tensile samples solution heat treated at 1050 °C for 5 min: (a) austenite, (b) αʹ -martensite, (c) TEM-SADP, (d) VN and dislocation lines, (e) TEM-SADP, and (f) TEM-EDS of VN
19
Fig. 8 TEM micrographs of hot-forged Cr19+W after tensile tests: (a) bright field TEM image, (b) TEM-EDS, (c) TEM-SADP of “Region 1”, (d) TEM-SADP of “Region 2”
20
Fig. 9 TEM micrographs near fracture surface of tensile samples after aging at 750 °C for 2 h and tensile tests: (a) bright field TEM image of Cr19+V, corresponding (b) TEM-EDS of VN, and (c) TEM-SADP, (d) bright field TEM image of Cr19+W, corresponding (e) TEM-EDS of Cr23C6, (f) TEM-SADP of Cr23C6 in “Region 1”, and (g) TEM-SADP of αʹ -martensite in “Region 2”
Table 1 Chemical composition of V and W modified Cr19 series DSSs DSSs
Cr
Ni
Mo
C
N
S
V
W
Fe
Cr19
19.21
3.46
1.75
0.026
0.19
0.012
-
-
Bal.
Cr19+V
19.17
3.48
1.81
0.026
0.19
0.013
0.18
-
Bal.
Cr19+W
19.37
3.40
1.83
0.025
0.19
0.012
-
0.65
Bal.
21
PII: DOI: Reference:
S0921-5093(17)31084-5 http://dx.doi.org/10.1016/j.msea.2017.08.071 MSA35419
To appear in: Materials Science & Engineering A Received date: 8 May 2017 Revised date: 16 August 2017 Accepted date: 17 August 2017 Cite this article as: Huawei Zhang, Yulai Xu, Pengfei Hu, Wanjian Xu, Jun Li, Kunfang Li and Xueshan Xiao, Strengthening behaviors of V and W modified Cr19 series duplex stainless steels with transformation induced plasticity, Materials Science & Engineering A, http://dx.doi.org/10.1016/j.msea.2017.08.071 This is a PDF file of an unedited manuscript that has been accepted for publication. As a service to our customers we are providing this early version of the manuscript. The manuscript will undergo copyediting, typesetting, and review of the resulting galley proof before it is published in its final citable form. Please note that during the production process errors may be discovered which could affect the content, and all legal disclaimers that apply to the journal pertain.
Strengthening behaviors of V and W modified Cr19 series duplex stainless steels with transformation induced plasticity Huawei Zhang1, Yulai Xu1, 2, *, Pengfei Hu1, Wanjian Xu1, Jun Li1, 2, Kunfang Li1, Xueshan Xiao1, 2 1. Institute of Materials, Key Laboratory for Microstructures, Shanghai University, Shanghai 200072, China 2. Shanghai University Xinghua Institute of Special Stainless Steels, Jiangsu 225721, China
Abstract: The V and W modified Cr19 series duplex stainless steels with transformation induced plasticity (TRIP) have been newly developed and the effects of microstructures on room temperature mechanical properties have been investigated. The addition of ferrite forming elements V and W slightly increases the volume fraction of ferrite phase from about 55 to 57-60%, however, the cold rolling process hardly influence the fraction. When the duplex phases exhibit a coarse banding morphology before cold rolling, the ultimate tensile strength and elongation are improved significantly with the addition of V and W elements mainly due to the occurrence of TRIP effect. The ultimate tensile strength and yield strength of Cr19+V increase very slightly, but the elongation increases to about 61% after cold rolling. Both the ultimate tensile strength and elongation of Cr19+W significantly increase to about 800 MPa and 67%, respectively. The improvement is attributed to TRIP effect and refinement of banded ferrite and austenite phases. After aging heat treatment, the precipitate of VN and Cr23C6 particles contributes to the further increase of ultimate tensile strength to about 1000 and 920 MPa of Cr19+V and Cr19+W DSSs, respectively. Strain induced αʹ -martensite is transformed from austenite directly with an orientation relationship. The strengthening behaviors have been discussed based on the microstructural evolutions. Key words: Duplex stainless steels, Transformation induced plasticity, Mechanical properties, Microstructures, Strengthening behaviors.
*
Corresponding author E-mail address: [email protected] 1
1. Introduction Duplex stainless steels (DSSs) composed of body-centered cubic ferrite (α) and face-centered cubic austenite (γ) exhibit good mechanical properties and corrosion resistance [1-3]. In order to reduce weight and improve safety, the TRIP steels show great potential for the application in automotive industry due to the good combination of strength and ductility [4]. Considerable research efforts have been carried out to improve both mechanical and corrosion properties, particularly by controlling alloying elements. The substitution of the costly alloying elements such as Ni and Mo with N and Mn makes the austenite phase more prone to the occurrence of the TRIP effect [5]. Ran et al. [6] had investigated the influence of heat treatment on TRIP effect of Fe-19.1Cr-1.3Ni-6.2Mn-1.0Mo-0.173N-0.027C-0.37Si and found that the ultimate tensile strength and elongation are about 1100 MPa and 50%, respectively. Herrera et al. [7] had designed a ductile
economical
Mn-based
duplex
stainless
TRIP
steel
Fe-19.9Cr-0.42Ni-0.16N-4.79Mn-0.11C-0.46Cu-0.35Si with ultimate tensile strength of 1000 MPa and an elongation value higher than 60%. The Mn element is often added into the DSSs to increase the solubility of N which gives potent solid solution strengthening and improves the resistance to pitting corrosion [8]. However, Jang et al. [9] had investigated the tensile and corrosion behavior of CD4MCU cast DSSs with different Mn contents, and found that the addition of 0.8 wt. % Mn was detrimental to both pitting corrosion and stress corrosion cracking properties. Yang et al. [10] had studied the effect of Mn on microstructure, mechanical and pitting corrosion resistant properties of economical Cr19 DSS with solution temperatures ranging from 1040 to 1220 °C, which indicated that the impact energy at 20 °C increased to more than 180 J when the solution temperature decreased from 1220 to 1040 and 1120 °C, and that the impact energy could be improved by Mn addition due to increased amount of austenite phase. But it also exhibited that pitting potential decreased with the increased Mn content from 3.6 to 8.1 wt. %. Interstitial elements such as carbon and nitrogen are always employed to form precipitates which pin the movement of dislocations to achieve good mechanical properties. Choi et al. [11] had investigated the effect of nitrogen content on TRIP effect of the new series of nearly Ni-free DSSs and found that the TRIP effect led to the enhancement of ultimate tensile strength over 1000 MPa and 2
ductility of about 60% for 0.35 wt. % N DSSs. Fraction of strain induced martensite increased with the N content resulting in a better strain hardening. Under severe corrosion conditions, sufficient amount of chromium, molybdenum and nitrogen elements are needed to obtain a high pitting resistance equivalent value. Zhang et al. [12] had designed a new series of Mo-free 21.5Cr-3.5Ni-xW-0.2N economical DSSs and shown that the ultimate tensile strength decreased from about 900 to 790 MPa, but the elongation increased from 35 to 61% and the pitting corrosion potential increased from 760 to 1011 mV with the increase of W content, which indicated that the ductility and corrosion resistance property can be improved with the addition of W element. Tensile properties of DSSs also have great relationship with the volume fraction of ferrite phase and the shape of austenite phase. The chemical compositions, heat treatment steps, and pre-treatment processes are key factors in influence the phase ratios. The DSSs show high strength and resistance to corrosion when the ferrite to austenite phase ratio is approximately 1:1 [13-15]. Solution heat treatment promotes dissolution of carbides and other precipitates after rolling process, which is an important process for DSSs to control the suitable phase ratio and hence obtain a good combination of mechanical properties and corrosion resistance. Lai et al. [16] studied the effect of solution heat treatment on the transformation kinetics of γ to α in DSSs at the temperature range of 1050-1250 °C and found that higher temperature increased the amount of ferrite and modified the morphology of ferrite phase in the matrix. The effects of precipitate strengthening which caused by dislocation and interstitial carbon or nitrogen atoms may be reduced during this period. The mechanical properties of DSSs have great relationship with deformation modes of ferrite and austenite phases. Plastic deformation of ferrite is mainly dominated by movement of dislocation due to high stacking fault energy [17]. The deformation of austenite includes transformation induced plasticity, twinning induced plasticity and planar glide of dislocations with the increase in stacking fault energy [18, 19]. The influence of V and W elements on mechanical properties of Cr19 series DSSs has been hardly investigated, and the new Cr19 series DSSs with additions of V and W elements have been designed to obtain a possible good combination of strength × elongation value and corrosion resistance property. The newly developed Cr19 series DSSs with 0.2 wt. % N without the addition of Mn element have been prepared, and this paper aims to study the effects of V and W elements on microstructure 3
evolutions and mechanical properties, and the strengthening behaviors after different deformation and heat treatments processes have also been discussed. 2. Experimental The Cr19 series DSSs modified with V and W elements were melted in a 50 kW induction furnace in an argon atmosphere. The chemical compositions of the castings designated as Cr19, Cr19+V, and Cr19+W are shown in Table 1. The casting ingots were hot forged into Φ 30 mm sticks, and the as-forged samples were solution heat treated at 1050 °C for 5 min followed by water quenching. The plates with the thickness of about 3.2 mm were cut from the sticks and then cold rolled to the sheets with the thickness of about 1.2 mm followed by solution heat treatment at 1050 °C for 5 min and water quenching. After that the samples are aged at 750 °C for 0, 0.5, 2, 4, and 6 h. The Charpy V-notch specimens with 10 mm × 10 mm × 55 mm were cut from the solution heat treated sticks along the forging direction according to P.R.C., GB/T229-2007 [20]. Charpy impact tests were carried out at 25, 0, and -40 ºC by using AHC3000/2-AT impact test machine with maximum capacity of 450 J. Polarization tests of the steels after hot-forging and solution heat treatment at 1050 °C were performed at least three times to ensure the reproducibility
of
the
results
by
using
an
EG&G
Princeton
Applied
Research
Potentiostat/Galvanostat Model 273A. Pitting corrosion samples were progressively wet ground up to 1200 grit SiC papers and then well rinsed in water, and the edges of the exposed samples were mounted with an epoxy resin. Then the samples with an area of 1 cm2 were tested in 1 mol/L NaCl solution at 25 ºC. The saturated calomel electrode (SCE) was used as a reference electrode, a platinum foil was served as the counter electrode, and open circuit potential (VOC) measurements were carried out for 2500 s. The polarization scan started at -0.4VSCE below VOC until 1.0VSCE with a scan rate of 1 mV/s. The specimens were electrochemically etched by 10 wt. % KOH solution and the optical micrographs were observed by KEYENCE VHX-100 microscopy. The volume fraction of ferrite was measured by quantitative metallographic analysis systems and the average value of ten measurements on each sample was taken as the fraction of ferrite. The fracture surface morphologies were analyzed by using HITACHI SU-1510 scanning electron microscope (SEM). 4
The tensile tests were carried out at room temperature according to National Standard of the P.R.C. GB/T228.1-2010 with the specimens having a gauge length of 30 mm, width of 10 mm, and thickness of 1 mm [21]. The microstructures of the samples after tensile tests were analyzed by JEM-2010F transmission electron microscope (TEM). The specimens were sliced from the tensile plates along the stress direction near the fracture surface, polished with abrasive papers from no. 800 to 2000, and thinned using a twin-jet electro-polishing in the solution composed of 10 vol. % HClO4 and 90 vol. % C2H5OH at -30 ºC and 45 V. 3. Results and discussions The stress-strain curves of hot-forged Cr19 series DSSs after solution heat treatment at 1050 °C for 5 min are shown in Fig. 1. The ultimate tensile strength significantly increases from about 720 MPa of Cr19 to 805 MPa of Cr19+V and 750 MPa of Cr19+W, and the corresponding elongation value also increases from about 43 to 52 and 57%. Addition of V element increases the yield strength from about 545 MPa of Cr19 to 590 MPa of Cr19+V. Although the yield strength decreases from about 590 MPa of Cr19+V to 520 MPa of Cr19+W, the corresponding elongation value increased from about 52 to 57%. It is evident that the addition of V and W elements is beneficial to room temperature tensile mechanical properties. Fig. 2 shows the stress-strain curves of cold-rolled Cr19+V and Cr19+W after solution heat treatment at 1050 °C for 5 min and aging at 750 °C for up to 6 h. It is clear that both the ultimate tensile strength and yield strength of Cr19+V slightly increase by about 10 MPa compared with those before cold rolling, but both the ultimate tensile strength and yield strength of Cr19+W increase by about 70 MPa. The ultimate tensile strength increases while the elongation decreases significantly with the increased aging time of Cr19+V and Cr19+W DSSs. The ultimate tensile strength remains at about 1000 MPa with the corresponding elongation value about 23% after aging over 2 h for Cr19+V. However, the elongation keeps the value higher than 38% even after aging for 6 h for Cr19+W with the improved ultimate tensile strength of about 920 MPa. It can be found that mechanical properties of Cr19+V and Cr19+W get better significantly after cold rolling, because the elongation value increases from about 52% to 61% for Cr19+V although the ultimate tensile strength increases very slightly from about 805 to 815 MPa, while both the ultimate tensile 5
strength and elongation respectively increase from about 750 and 57% to 800 MPa and 67% of Cr19+W. So the cold rolling process improves both the tensile strength and ductility at the same time. The strength × elongation (MPa%) values of cold rolled V and W modified Cr19 DSS after solution treatment at 1050 °C for 5 min are respectively about 49715 and 53600 MPa%, which indicates that the modified Cr19 series DSSs obtain a very good combination of strength × elongation value. The
combination
of
strength
Fe-19.1Cr-1.3Ni-6.2Mn-1.0Mo-0.173N-0.027C-0.37Si
×
elongation
value
of and
Fe-19.9Cr-0.42Ni-0.16N-4.79Mn-0.11C-0.46Cu-0.35Si DSSs are respectively about 55000 MPa% and 66300 MPa% according to Ran et al. [6] and Herrera et al. [7], and these values are very high at present. However, the addition of Si and high content of C may be harmful to the impact toughness and pitting corrosion resistant properties. At the same time, the high Mn contents may also do harm to the pitting corrosion resistant property. Fig. 3 shows the impact energy at 25, 0, and -40 ºC of the hot-forged Cr19, Cr19+V, Cr19+W DSSs after solution heat treatment at 1050 °C for 5 min. It can be obtained that addition of V reduces the impact energy of Cr19 DSS especially at -40 °C. Although W reduces the energy at 25 °C, the impact energy hardly varies at 0 °C and the value even increases from about 128 to 145 J at -40 °C. Fig. 4 shows the SEM fracture surface morphologies after impact test at 0 °C and -40 °C. Large amounts of ductile dimples uniformly distribute on the fracture surface of Cr19 DSS, which indicates a good toughness (Fig. 4(a)). The typical SEM morphologies vary slightly of Cr19+V and Cr19+W DSSs (Fig. 4(b) and (c)). The amount of ductile dimples significantly reduces at -40 °C for both the Cr19 and Cr19+V DSSs (Fig. 4(d) and (e)). However, it should be noticed that the fracture surface morphology of Cr19+W shows a typical ductile fracture at -40 °C because both the amount and size of the dimples increases and decreases respectively compared with that of Cr19 DSS at 0 °C. Fig. 5 shows the representative polarization curves of Cr19 series DSSs after hot-forging and solution heat treatment at 1050 °C. The pitting corrosion potential (Epit) is about 700 mV for Cr19 DSS, and Epit respectively increases to about 820 and 930 mV for Cr19+V and Cr19+W DSSs. It has been suggested that pitting corrosion resistance depend basically on Cr, Mo and N contents, and resistance to pitting corrosion can be characterized in terms of a pitting resistance equivalent 6
(PRE) value, which is often defined as [22, 23]: PRE = wt. % Cr + 3.3 (wt. % Mo + 1/2 wt. % W) + 30 wt. % N
(1)
The PRE values of the experimental alloys calculated according to Eq. (1) are respectively about 30.7, 30.8, and 32.1. It can be seen that the PRE values increase slightly due to the chemical composition vibration and addition of W element, which is in good agreement with the results of the pitting corrosion test. So the yield strength and elongation values of the newly developed Cr19 series DSSs are high, the impact toughness, and pitting corrosion resistance is also very good especially for Cr19+W steel. Fig. 6 shows the typical optical micrographs of the hot forged and cold-rolled specimens after solution heat treatment at 1050 °C. The typical duplex phases of ferrite and austenite have been identified. The volume fraction of ferrite phase which has been measured is respectively about 55, 57, and 60% for Cr19, Cr19+V, and Cr19+W DSSs (Fig. 6(a)-(c)). So it confirms that both the V and W are ferrite forming elements. Fig. 6(d)-(f) shows that both the thickness of the banded ferrite and austenite has been refined dramatically along the rolling direction, but the volume fraction of ferrite phase hardly changes compared with that before cold rolling. The TEM micrographs of hot-forged Cr19+V DSS after solution heat treatment at 1050 °C for 5 min and tensile tests are shown in Fig. 7. The elongated austenite phase has been observed in Fig. 7(a) and the αʹ -martensite has also been identified among the austenite phase (Fig. 7(b)). The corresponding selected area diffraction pattern (SADP) in Fig. 7(c) shows that the face centered cubic austenite phase and the near cubic αʹ -martensite phase co-exist following the orientation relationship of <220>γ // <211>αʹ , which confirms that formation of αʹ -martensite is due to the deformation of austenite along tensile direction. Lots of precipitates marked with dotted circles in Fig. 7(d) have been identified in the matrix, and dislocation lines distribute around these particles. The corresponding TEM-SADP (Fig. 7(e)) and TEM-EDS (Fig. 7(f)) indicate that these precipitates belong to VN particles. Saenarjhan et al. [24] also pointed out that formation of the ɛ -martensite wasn’t a pre-requisite for transforming austenite into αʹ -martensite and suggested that austenite directly transformed to αʹ -martensite during deformation rather than via the deformation bands due to a relatively high driving force for αʹ -martensite transformation in Fe-0.17N-0.02C-20.5Cr-2.0Ni-1.8Mn-0.7Cu-0.6Mo-0.5Si. At the tensile strain of 30%, the 7
αʹ -martensite appeared in austenite with <110>γ // <001>αʹ orientation relationship. For specimen deformed up to failure, a large number of dislocations and stacking faults were observed and both of Nishiyama-Wasserman orientation relationship and Kurdjumov-Sachs orientation relationship of <110>γ // <111>αʹ existed between austenite and αʹ -martensite. In addition, Zhang and Hu [25] demonstrated that deformation induced martensite transformation (DIMT) followed γ →
hexagonally
indexed
deformation
bands
→
αʹ
in
0.14N-0.02C-20Cr-2.1Ni-5.1Mn-0.2Cu-0.3Mo-0.4Si DSS which had a higher concentration of Mn. Deformation bands appeared at the early stage of deformation and αʹ -martensite formed at the intersection of the deformation bands later. Calculation showed that the transformation driving force of austenite into ɛ -martensite was always positive and much bigger than that of austenite into αʹ -martensite because higher Mn content significantly reduced driving force of αʹ -martensite transformation. Therefore, the DIMT sequence seemed to be determined by the value of driving force. But no ɛ -martensite has been identified in Cr19+V steel according to TEM observations. Fig. 8 shows the TEM micrographs of hot-forged Cr19+W after solution heat treatment at 1050 °C for 5 min and tensile tests. The typical TEM-EDS of “Region 1” and “Region 2” in Fig. 8(a) is shown in Fig. 8(b), and W element which has been identified confirms that addition of W mainly solutes in the matrix. The SADPs of “Region 1” and “Region 2” in Fig. 8(a) are respectively shown in Fig. 8(c) and 8(d). Only αʹ -martensite exists at “Region 1” while both γ-austenite and αʹ -martensite are found at “Region 2” following the orientation relationship of <110>γ // <110>αʹ , which also proves the formation of strain induced martensite due to the TRIP effect. It is known that the austenite with different stability has two different transformation paths during deformation controlled by elemental partitioning and austenite grain size distribution. One way is a direct transformation from austenite to αʹ -martensite in medium Mn-TRIP steels [26-28] and quenching and partitioning steels [29, 30], and another way through an intermediate ε-martensite and subsequently the formation of αʹ -martensite in Mn-N DSS [31, 32] and high-Mn austenitic steels [33]. Now that ɛ -martensite wasn’t found in the Cr19 series DSSs, αʹ -martensite was transformed from austenite directly. Fig. 9 shows the typical TEM micrographs of Cr19+V and Cr19+W near the fracture surface 8
after aging at 750 °C for 2 h and tensile tests. The approximately spherical precipitate with the size about 70 nm had been observed in the grain interior in Fig. 9(a), both the corresponding TEM-EDS and TEM-SADP in Fig. 9(b) and (c) confirm that this precipitate belongs to VN. The size of VN in the aged specimen is bigger than that found in Fig. 7(d) due to the coarsening of VN particles to minimize the interfacial energy [34]. A large precipitate had been observed in “Region 1” from the bright filed TEM image of aged Cr19+W in Fig. 9(d), the length and width are respectively about 1200 and 630 nm. The corresponding TEM-EDS pattern shown in Fig. 9(e) indicates that the large particle is enrich of Cr element which can be identified as Cr23C6 carbide by TEM-SADP in Fig. 9(f). Dong et al. [35] showed that a considerable amount of M7C3 carbides transformed into M23C6 carbides while M23C6 continued to grow up when the carbon partitioning time was extended to 15 min at 600 °C in Nb-V-Ti microalloyed ultra-high strength steel. In the studies of Kucharova and Shankar et al. [36, 37], W mostly enriched in M6C and played an important role in suppressing the coarsening of Cr23C6. But no peak of W element had been observed from the EDS patterns, which confirms that W element still solutes in the matrix during aging heat treatment. The SADP of “Region 2” is shown in Fig. 9(g), which proves the existence of αʹ -martensite because of TRIP effect during tensile tests. The above results indicate that addition of V and W elements has great influence on room temperature tensile mechanical properties (Fig. 1), impact toughness (Fig. 3), and pitting corrosion resistant property (Fig. 5), because the V and W elements change the microstructures of the TRIP-assisted Cr19 series DSSs. After hot forging and solution heat treatment, the morphologies of the duplex phase vary slightly except that the volume fraction of ferrite phase increases from about 55 to 57-60% (Fig. 6). The TEM observations also indicate that the αʹ -martensite has been identified in austenite in both the V and W modified DSSs and the αʹ -martensite exhibit an orientation relationship with γ, and VN has been observed in the matrix in Cr19+V steel (Figs. 7 and 9). So the significantly increased elongation from about 43% of hot-forged Cr19 to higher than 52% of hot-forged Cr19+V and Cr19+W DSSs is primarily due to the easier occurrence of TRIP effect during deformation, because addition of ferrite forming elements V and W reduces the stability of austenite. Among the three alloys, Cr19+W shows the highest ferrite fraction and corresponding the least stability of austenite, which leads to the best elongation. Increased tensile 9
strength of Cr19+V results from the TRIP effect, precipitate strengthening of VN particles which hinders the dislocation moving, and solid solution strengthening of V element. However, no carbide particles have been observed in the tensile samples of Cr19+W after solution heat treatment, and the increased tensile strength of Cr19+W is attributed to the TRIP effect because of the phase transformation of αʹ -martensite from austenite, and to the solid solution strengthening of W because the W element mainly solute in the matrix according to the TEM-EDS patterns. The Charpy impact test shows that all the DSSs exhibit good toughness because the impact energy is 200 ± 15 J at 25 °C, the impact energy hardly varies with the addition of W element, and the most important one is that the low temperature impact toughness of Cr19+W is the best among the three steels. Wan et al. [38] showed that impact toughness of 15Cr DSSs decreased with Al content at the same temperature mainly resulting from the ferrite-austenite ratio in duplex phase structure. Austenite phase has a better toughness than that of ferrite phase, so ferrite phase is the phase that exhibits ductile-brittle transition (DBT) phenomenon in DSSs. Pecker et al. [39] also suggested that higher impact energy at 0 °C and -40 °C of 22Cr-2.0Ni was because of the increased stacking fault energy. Fig. 3 shows that the impact energy of Cr19+V is slightly lower than that of Cr19, especially at low temperatures. The reduced toughness of Cr19+V steel may be primary due to the slightly increased volume fraction of ferrite phase and precipitate of VN particles. As to the Cr19+W steel, no precipitate has been identified in the matrix which may benefit the improved toughness at low temperatures. After cold rolling and heat treatment, the ultimate tensile strength, yield strength, and elongation further increase to a very high level. The optical micrographs indicate that the volume fraction of ferrite hardly varies compared with that before cold rolling, but the duplex phases exhibit a banded morphology and the thickness dramatically reduces (Fig. 6). Hence, the enhancement of room temperature tensile mechanical properties after cold rolling is primarily due to the refinement of both ferrite and austenite phases compared with those after hot forging. The reduced thickness of the phases significantly contributes to the increased elongation of Cr19+V and the improved tensile strength and elongation of Cr19+W DSS. So the reduced thickness of the duplex phase does not play an important role in improving the tensile strength because the tensile strength is in a relatively high level due to the precipitate strengthening of VN nano-sized particles. 10
However, the reduced thickness great influence the tensile strength of Cr19+W steel because the strengthening mainly comprised of solid solution strengthening and TRIP effect. After aging heat treatment, the tensile strength and ductility of Cr19+V is respectively stable at about 1000 MPa and 23%. The improved tensile strength of Cr19+V and Cr19+W after aging heat treatment is primary due to the precipitate of VN and Cr23C6 carbides, respectively. Conclusion The microstructures, room temperature tensile mechanical properties, impact toughness, and pitting corrosion resistance of TRIP-assisted V and W modified Cr19 series DSSs have been investigated and the corresponding strengthening behaviors have also been discussed. Addition of ferrite forming elements V and W increase the volume fraction of ferrite from 55 to 57-60%, but the fraction hardly varies after cold rolling and following heat treatment. The V and W elements reduce the stability of austenite and result in an easier occurrence of TRIP effect of Cr19 series DSSs because the ultimate tensile strength and elongation can be improved obviously when the duplex phases exhibit a coarse banding morphology: the V element increases the ultimate tensile strength to 805 MPa with elongation of 52%, and W improves the ultimate tensile strength of 750 MPa with elongation about 57% after solution heat treatment at 1050 °C of the hot-forged Cr19 series DSSs. Addition of W increases the impact energy at -40 °C of hot-forged steels while V reduces impact energy of Cr19 DSSs due to the precipitate of VN particles. The improvement of room temperature mechanical properties before cold rolling is mainly due to the phase transformation strengthening. After cold rolling and heat treatment at 1050 °C, the improvement of tensile strength and elongation is mainly due to the TRIP effect and refinement of the banded duplex phases especially for the Cr19+W, because the ultimate tensile strength and yield strength of Cr19+V increase very slightly, while both the ultimate tensile strength and elongation of Cr19 +W increases to about 800 MPa and 67%, respectively. The significantly increased value of strength × elongation is primarily due to the phase refining strengthening and transformation strengthening. Strain induced αʹ -martensite is transformed from austenite directly without the formation of ɛ -martensite with an orientation relationship existing between austenite phase and strain induced martensite. 11
The precipitate strengthening plays an important role after aging heat treatment, the ultimate tensile strength of Cr19+V is promoted notably to about 1000 MPa due to the precipitation of VN particles, but both the yield strength and elongation dramatically decrease to 475-525 MPa and about 23%, respectively. The ultimate tensile strength of Cr19+W increases to as high as 920 MPa, but the yield strength and elongation respectively decrease by about 10-45 MPa and 22-29%, which is primary due to the precipitate of Cr23C6 carbide.
Acknowledgements Sponsored by Shanghai Sailing Program [17YF1405800] of Shanghai Municipality. Prospective Joint Research Project of Department of Science and Technology of Jiangsu Province [BY2015068-01].
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Table and Figure Captions Table 1 Chemical composition of V and W modified Cr19 series DSSs Fig. 1 Stress-strain curves of Cr19, Cr19+V and Cr19+W DSSs after hot forging and solution heat treatment at 1050 °C for 5 min Fig. 2 Stress-strain curves of Cr19+V and Cr19+W DSSs aged at 750 °C for 0, 0.5, 2, 4, and 6 h after annealing at 1050 °C for 5 min, (a) Cr19+V, (b) Cr19+W Fig. 3 Charpy impact toughness of hot-forged Cr19 series DSSs at 25, 0, and -40 ºC Fig. 4 SEM fracture surface morphologies of Cr19 series DSSs at 0 °C of (a) Cr19, (b) Cr19+V, (c) Cr19+W, and at -40 °C of (d) Cr19, (e) Cr19+V, (f) Cr19+W 15
Fig. 5 Polarization curves of Cr19 series DSSs in 1 mol/L NaCl at 25 °C Fig. 6 Optical micrographs of hot-forged (a) Cr19, (b) Cr19+V, (c) Cr19+W DSSs, and cold-rolled (d) Cr19, (e) Cr19+V, (f) Cr19+W DSSs after heat treatment at 1050 °C Fig. 7 TEM micrographs of Cr19+V near the fracture surface of tensile samples solution heat treated at 1050 °C for 5 min: (a) austenite, (b) αʹ -martensite, (c) TEM-SADP, (d) VN and dislocation lines, (e) TEM-SADP, and (f) TEM-EDS of VN Fig. 8 TEM micrographs of hot-forged Cr19+W after tensile tests: (a) bright field TEM image, (b) TEM-EDS, (c) TEM-SADP of “Region 1”, (d) TEM-SADP of “Region 2” Fig. 9 TEM micrographs near fracture surface of tensile samples after aging at 750 °C for 2 h and tensile tests: (a) bright field TEM image of Cr19+V, corresponding (b) TEM-EDS of VN, and (c) TEM-SADP, (d) bright field TEM image of Cr19+W, corresponding (e) TEM-EDS of Cr23C6, (f) TEM-SADP of Cr23C6 in “Region 1”, and (g) TEM-SADP of αʹ -martensite in “Region 2”
Fig. 1 Stress-strain curves of Cr19, Cr19+V and Cr19+W DSSs after hot forging and solution heat treatment at 1050 °C for 5 min
16
Fig. 2 Stress-strain curves of Cr19+V and Cr19+W DSSs aged at 750 °C for 0, 0.5, 2, 4, and 6 h after annealing at 1050 °C for 5 min, (a) Cr19+V, (b) Cr19+W
Fig. 3 Charpy impact toughness of hot-forged Cr19 series DSSs at 25, 0, and -40 ºC
17
Fig. 4 SEM fracture surface morphologies of Cr19 series DSSs at 0 °C of (a) Cr19, (b) Cr19+V, (c) Cr19+W, and at -40 °C of (d) Cr19, (e) Cr19+V, (f) Cr19+W
Fig. 5 Polarization curves of Cr19 series DSSs in 1 mol/L NaCl at 25 °
18
Fig. 6 Optical micrographs of hot-forged (a) Cr19, (b) Cr19+V, (c) Cr19+W DSSs, and cold-rolled (d) Cr19, (e) Cr19+V, (f) Cr19+W DSSs after heat treatment at 1050 °C
Fig. 7 TEM micrographs of Cr19+V near the fracture surface of tensile samples solution heat treated at 1050 °C for 5 min: (a) austenite, (b) αʹ -martensite, (c) TEM-SADP, (d) VN and dislocation lines, (e) TEM-SADP, and (f) TEM-EDS of VN
19
Fig. 8 TEM micrographs of hot-forged Cr19+W after tensile tests: (a) bright field TEM image, (b) TEM-EDS, (c) TEM-SADP of “Region 1”, (d) TEM-SADP of “Region 2”
20
Fig. 9 TEM micrographs near fracture surface of tensile samples after aging at 750 °C for 2 h and tensile tests: (a) bright field TEM image of Cr19+V, corresponding (b) TEM-EDS of VN, and (c) TEM-SADP, (d) bright field TEM image of Cr19+W, corresponding (e) TEM-EDS of Cr23C6, (f) TEM-SADP of Cr23C6 in “Region 1”, and (g) TEM-SADP of αʹ -martensite in “Region 2”
Table 1 Chemical composition of V and W modified Cr19 series DSSs DSSs
Cr
Ni
Mo
C
N
S
V
W
Fe
Cr19
19.21
3.46
1.75
0.026
0.19
0.012
-
-
Bal.
Cr19+V
19.17
3.48
1.81
0.026
0.19
0.013
0.18
-
Bal.
Cr19+W
19.37
3.40
1.83
0.025
0.19
0.012
-
0.65
Bal.
21