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Assessment of sensitization resistance of AISI 409M grade ferritic stainless steel joints using Modified Strauss test
Materials and Design 39 (2012) 175–185
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Assessment of sensitization resistance of AISI 409M grade ferritic stainless steel joints using Modified Strauss test A.K. Lakshminarayanan ⇑, V. Balasubramanian Centre for Materials Joining and Research (CEMAJOR), Department of Manufacturing Engineering, Annamalai University, Annamalai Nagar 608 002, Tamil Nadu, India
a r t i c l e
i n f o
Article history: Received 14 December 2011 Accepted 18 February 2012 Available online 28 February 2012 Keywords: D. Welding E. Corrosion F. Microstructure
a b s t r a c t The paper reports on the findings from a detailed study of the joints fabricated by four different welding processes namely, gas tungsten arc welding, friction stir welding, laser beam welding and electron beam welding to assess the susceptibility to intergranular corrosion. Bend test and weight loss methods were used to evaluate the degree of sensitization after Modified Strauss test. The results are correlated with the microstructure by optical, scanning and transmission microscopy and the chemical composition by energy dispersive X-ray spectrometry. Friction stir welded joint is less prone to sensitization, but suffered from general corrosion by dissolution of martensite phase. Laser beam and electron beam welded joints exhibited lowest corrosion rate as compared to other joints. Gas tungsten arc welded joint is prone to sensitization and showed lower corrosion resistance compared to its counterparts. Ó 2012 Elsevier Ltd. All rights reserved.
1. Introduction Binary Iron–Chromium (Fe–Cr) alloys with a chromium content of 10–15% are found to exhibit an excellent corrosion resistance under ambient conditions and, because being less costly, they may substitute FeCrNi stainless steels for many applications [1]. Welding is an important method of fabricating complex plant, and the stability of welds in stainless steel is closely related to the composition of the steel, the thermal conditions associated with welding, and any stress that may be present in service [2]. The thermal cycle produced during welding can cause carbide precipitation and various phases in the heat-affected zone (HAZ). Corrosion resistance can also be altered mainly when phenomenon such as sensitization is likely to occur [3]. Some studies were reported on sensitization behaviour of ferritic stainless steel (FSS) joints fabricated using conventional welding processes such as gas tungsten arc welding, gas metal arc welding and shielded metal arc welding processes. Tomlinson and Matthews [4] investigated the sensitization behaviour of unwelded and gas tungsten arc welded flat and bent AISI 405 stainless steel specimens by chemical, electrochemical techniques and they reported that the HAZ was susceptible to intergranular attack and can be eliminated by proper annealing treatment. Greef and Du Toit [5] made an attempt to determine the extent to which thick gauge 3CR12 material was susceptible to sensitisation, to ⇑ Corresponding author. Tel.: +91 4144 239734, mobile: +91 9865431106; fax: +91 4144 238080/238275. E-mail addresses: [email protected] (A.K. Lakshminarayanan), balasubrama [email protected] (V. Balasubramanian). 0261-3069/$ - see front matter Ó 2012 Elsevier Ltd. All rights reserved. doi:10.1016/j.matdes.2012.02.038
characterize its occurrence in multiple pass welds, and to determine to what extent it could be minimized by appropriate control of the gas tungsten arc welding parameters. It was concluded that, the low heat inputs and fast cooling rates reduces the level of sensitisation. The susceptibility of 11–12% chromium type EN1.4003 FSS to sensitization during continuous cooling after welding at low heat input levels was investigated by Van Warmelo et al. [6] and it was found that welding at low heat inputs can suppress the transformation of ferrite to austenite as the HAZ cools through the (austenite + ferrite) dual phase region during welding. Sensitization is prevented by the presence of enough austenite to eliminate continuous ferrite–ferrite grain boundaries. Du Toit and Naude [7] studied the sensitization behaviour of unstabilized and titanium stabilized FSS and reported that the unstabilized 1.4003 steel contained considerably more grain boundary martensite in the high temperature heat affected zone (HTHAZ) after low heat input welding than the titanium stabilized grade, despite having very similar kaltenhauser ferrite factor (KFF). The presence of grain boundary martensite in the unstabilized grade increased the resistance to sensitization under low heat input conditions. Using AISI 409M grade ferritic stainless steel for structural application in corrosive environments will be economical as it imparts acceptable corrosion resistance using minimum Cr content. However, sensitization of stainless steel during thermo mechanical processing and welding has been a major challenge to face. Stainless steels like AISI 409M grade, which are having low chromium content (with 11.4%), are susceptible to intergranular corrosion when improper welding conditions are used. The susceptibility of
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up. The experimental setup is displayed in Fig. 1. Trials were carried out to find the optimum immersion conditions to compare the corrosion attack of joints fabricated by four different welding processes. Improper testing condition (i.e., acid concentration, immersion time and amount of copper shots) resulted in aggressive attack as displayed in Fig. 2. Initially, the immersion time was varied from 10 to 25 h and the corrosion rate was calculated. However, the specimens (shown in Fig. 3) boiled for 20 h were further analyzed for comparison among four different welding processes. Both weight loss and bend test were used to evaluate the degree of sensitization. Evaluation by bending is preferred for welds since mass loss of different welds at a range of heat inputs would not give the real effect of intergranular corrosion. Bending reveals embrittlement of the weld and heat affected zone due to chromium depletion at the grain boundaries. The specified bend test requires the sample to be bent through 180 over a radius equal to twice the thickness of the specimen. The appearance of fissures or cracks indicates the presence of intergranular attack. In this study, the welded joints were grounded to 3 mm and bent to 180 with a bent radius of 6 mm at constant deformation. The bent specimens are displayed in Fig. 4. The U bend prepared from the welded joints were subjected to boiling immersion testing for 24 h. After immersion testing, the specimens were cleaned with soft brush to remove the copper deposits and polished with different grades of fine emery papers and then analyzed using optical and scanning electron microscopy.
these alloys is strongly affected by welding processes, which change the microstructure of the alloy in order to have optimum mechanical properties. However, a detailed comparison of different welding processes on the sensitization behaviour of 409M grade FSS welds are hitherto not reported and hence the present investigation was carried out to understand the effect of welding processes namely, friction stir welding (FSW), laser beam welding (LBW), electron beam welding (EBW) and gas tungsten arc welding (GTAW) on sensitization behaviour of AISI 409M grade stainless steel. 2. Experimental procedure 2.1. Base metal The as-received base metal (BM) used in this study was 4 mm thick cold rolled, annealed and pickled AISI 409M grade FSS plates. The chemical composition of the base metal obtained using vaccum spectrometer is presented in Table 1. 2.2. Welding Square butt joints were prepared to fabricate GTAW, LBW, EBW and FSW processes without filler metal additions. GTAW joints were fabricated using Lincoln welding machine (USA) with a capacity of 400 A. LBW joints were fabricated using a CO2 laser beam welding machine (Rofin Slab: CO2 laser). EBW joints were fabricated using an electron beam welding machine (Techmeta, France) with a capacity of 100 kV. FSW joints were fabricated using an indigenously designed and developed FSW machine (22 kW; 4000 rpm; 6 Tonne) using a tungsten based alloy tool. Few welding trials were carried out and specimens were extracted from various locations of the joint and subjected to macrostructural analysis. The specimen free of volumetric defect and lack of penetration was considered as the optimized welding condition. The welding conditions and optimized process parameters presented in Table 2 were used to fabricate the joints for further investigation.
2.5. Microstructural analysis Microstructure of the as-welded specimens prior to boiling immersion test was examined using transmission electron microscopy (TEM). Thin foils with a thickness of about 1 mm were cross sectioned from the welded joints to prepare the TEM samples. After being mechanically ground to approximately 80 lm, the foils were further ground to a thickness of 15 lm by a dimple grinding machine. Then the electron transparent thin sections were electrolytically prepared by the twin-jet polisher. An electrolyte of 10% perchloric acid +90% ethanol solution ( 25 °C) was used during thinning. These thin foils were observed under a 200 kV potential using a Transmission Electron Microscope (Make: Philips, UK: Model: CM20). Microstructure of the flat and U bent specimens prepared from the welded joint after boiling immersion test was analyzed using optical and scanning electron microscopy. Energy dispersive X-ray Spectrometry (EDAX) analysis was performed to determine the distribution of elements in the ferrite and martensite phases of the weld metal region of all the joints.
2.3. Specimen preparation The welded joints were sliced using abrasive cutting and then machined to the required dimensions for preparing Strauss test specimens as per ASTM A763-93 [8] guidelines. The slices derived from the welded joints were polished using standard metallographic procedures to obtain flat and scratch free surface. 2.4. Modified Strauss test The Modified Strauss test is a variation on ASTM A763-93 Practice Z [8] for detecting susceptibility to sensitization in ferritic stainless steels. Samples were placed in a copper sulphate solution (60 g CuSO4 in 1 l of water) which was acidified by adding 0.5% sulphuric acid (3 ml H2SO4 in 1 l of water). A layer of copper shot was placed on the bottom of the vessel and the samples were placed in the solution in such a way to prevent direct contact with the welded region. The polished specimen were placed in the Erlenmeyer flask, so that the polished surface was horizontal and facing
3. Results 3.1. Macrostructure The cross sectional macrostructure of the welded joints are shown in Fig. 5. No volumetric defect is observed. Laser and electron beam weldments exhibited very thin weld beads (1.5 and 2.5 mm respectively) as compared with much wider beads of GTAW and FSW (10 and 20 mm respectively).
Table 1 Chemical composition (wt.%) of base metal (spectrometry results). Material
C
Mn
P
S
Si
Cr
Ni
Ti
Fe
Base metal (AISI 409M)
0.028
1.10
0.030
0.010
0.40
11.40
0.39
0.004
Bal.
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A.K. Lakshminarayanan, V. Balasubramanian / Materials and Design 39 (2012) 175–185 Table 2 Welding conditions and process parameters. Parameters
Process
Welding machine Tungsten electrode diameter (mm) Power (kW) Current Voltage Focal length (mm) Focal spot diameter (lm) Shielding gas Gas flow rate (l/min) Gun to work distance (mm) Vacuum (bar) Welding speed (mm min-1) Tool rotational speed (rpm) Axial force (kN) Tool material composition Tool pin profile Shoulder diameter (mm) Pin diameter (mm) Pin length (mm) Heat input (J/mm)
FSW
LBW
EBW
GTAW
RV machine tools, India – – – – – – Argon 8 – – 110 850 24 W 90%, Ni 3%, Mo 1%, Co 1%, Fe bal Taper cylindrical 18 Tapered from 8 to 5 3.7 828
Rofin, Germany – 3.5 – – 300 180 Helium 10 – – 3000 – – – –
Techmeta, France – – 51 mA 55 kV – – – – 298 10 4 1000 – – – – – – – 224
Lincoln, USA 3 – 120 A 20 V – – Argon 16 – – 80 – – – – – – – 1800
Fig. 1. Setup for Modified Strauss test.
3.2. Microstructure of as welded specimen SEM image (Fig. 6a) and TEM image (Fig. 6b) of base metal shows the characteristics of grain boundaries with ferrite grains and the presence of Cr23C6 carbides. TEM image of stir zone (Fig. 7a) comprises the banded microstructure of ferrite and martensite. Stirzone (SZ) consisting of a mixed deformation structure with equiaxed ultrafine grains and the cellular substructure between martensite laths. Severe deformation due to friction stirring resulted in the formation of fine subgrained structure, which propagates throughout the ferrite grains. This strain-induced sub microcrystalline structure involves highly misoriented equiaxed grains with a relatively low dislocation density in their interiors as compared with high dislocation densities in cell substructures that evolve at smaller strains. The presence of high dislocation density in the substructure discloses that SZ had undergone strain hardening during welding. TEM image of FZ of EBW joint (Fig. 7b) clearly shows the presence of prior d ferrite grain boundary, substructure and subgrain structure. The presence of chromium carbides and grain boundary martensite are also evidenced from Fig. 7b. This suggests that the equiaxed coarse grain structure base metal undergone a refinement during weld thermal cycle. Similar microstructure is observed in the FZ of LBW joints. The difference in FZ microstruc-
– 84
ture between EBW joint and LBW joint is the degree of fineness in the subgrain structure. In LBW joint (Fig. 7c), the ferrite lath width is comparatively smaller than the width of the ferrite lath evolved in EBW joint. Coarse lath ferrite structure with partially covered needle shape martensite and with chromium carbides decorated the grain boundaries as observed in case of GTAW joint (Fig. 7d). The HTHAZ of all four joints was analyzed using TEM and the resultant images are presented in Fig. 8. The HTHAZ of FSW joint (Fig. 8a) also shows a duplex structure of ferrite and martensite. The martensite was formed along many sub-grain boundaries of the ferrite as well as in grain boundaries. Chromium carbides are also observed in some of the grain boundaries. Dislocations are observed both in ferrite and martensite of this region. It indicates that this region undergone partial plastic deformation during the processes. The HTHAZ of GTAW joint (Fig. 8b) shows a mixed microstructure of ferritic grains with both intra and intergranular chromium carbides. The growth of ferrite was accompanied by the precipitation of carbides at moving interface and this result in the aligned precipitation which follows the shape of ferrite grain. Coarse plate like martensite and needle shape martensite along the grain boundaries of ferrite are also observed in the HTHAZ. The HTHAZ of LBW and EBW joints (Fig. 8c and d) shows mostly of ferritic structures enriched with grain boundary chromium carbides. Low carbon lath martensite is also observed at the grain boundaries. 3.3. Composition analysis of the weld regions from EDAX analysis Liquid–solid and solid-state phase transformations during welding are generally associated with the partitioning of alloying elements between the phases. The extent of this partitioning is dependent on the transformation temperature and the weld cooling rate. A greater extent of partitioning will occur at high temperatures where diffusion of alloying elements occurs most rapidly. Since such elemental partitioning strongly influences the final properties of each phase and the alloy in general, particularly the mechanical and corrosion properties, it is important to determine the extent of any such changes due to welding. The sole purpose of the EDAX analysis was to verify subtle differences in elemental composition between the two phases namely ferrite and martensite. The results of EDAX analysis of the microstructures developed during the d ferrite to austenite transformation in fusion welded
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(a) Copper deposit over LBW specimen (b) Cross section of LBW joint after the removal of copper deposit
(c) Cross section of GTAW joint after the removal of copper deposit
(d) FSW joint after the removal of copper deposit
Fig. 2. Aggressive corrosion attack due to improper testing condition.
(a) GTAW
(b) LBW
(c) FSW
(d) EBW
Fig. 3. Top and cross section of welded joints under optimum testing condition.
(EBW, LBW and GTAW) joints indicated element partitioning between the ferrite and austenite phases during the transformation. It is found that the main ferrite stabiliser, chromium, partitioned to the ferrite phase, whereas the austenite stabilizing elements, nickel and manganese, partitioned to the transformed austenite phase. Though the temperature during FSW did not reached the d ferrite range, the welding was taken place at approximately 1150 °C, where the maximum amount austenite forms in the steel used in the present study. Despite of the absence of phase transformation from d ferrite to austenite, the EDAX analysis showed the redistribution of chromium and nickel in the ferrite and martensite phases. The compositions obtained for the ferrite and martensite phases in the welded joints are presented in Table 3. For all conditions, higher levels of chromium and lower levels of nickel were observed in the ferrite regions of all the welded joints. However, there were no significant differences between the welded joints.
Fig. 4. Bend test specimens after boiling immersion test.
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179
(i) GTAW
(ii) LBW
(iii) EBW (a – basemetal, b – fusion zone (FZ), c – high temperature heat affected zone (HTHAZ), d – low temperature heat affected zone (LTHAZ))
(iv) FSW (a – basemetal (BM), b – flow arm: shoulder influenced region, c – banded structure : pin influenced region, d, e – thermo mechanical affected zone, f- high temperature heat affected zone (HTHAZ), AS – advancing side, RS – retreating side). Fig. 5. Cross-sectional macrostructure of 409M ferritic stainless steel welded joints.
Fig. 6. SEM and TEM image of the basemetal.
3.4. Immersion test
3.5. Microstructure of flat specimens after boiling immersion testing
The immersion tests were conducted for all the four welded joints by varying the immersion time from 10 to 25 h. The solution attacks both chromium carbides and matrix. When the carbides are in the grain boundaries and are attacked by the corrosion medium, the alloy suffers a weight loss. The weight loss method was used to calculate the corrosion rate as per ASTM A763-93 [8] guidelines. The relationship between the immersion time and the corrosion rate is plotted in Fig. 9. It can be inferred that the corrosion rate of LBW specimen is much lower when compared to the other joints. The GTAW joints exhibited higher corrosion rate when compared to its counterpart.
For comparison, among the welded joints, the weld metal (WM) region and HTHAZ of the specimens boiled at an immersion time of 20 h was analyzed using optical and scanning electron microscopy and are presented in Figs. 10–12. The WM region of all the four joints after boiling immersion test is compared in Fig. 10. It can be seen that the FZ of the GTAW joint (Fig. 10a) contains ditches in many regions (mostly ferrite–martensite boundaries) and severe pit holes of various sizes in the WM region (mostly in ferrite phase). Some of pits show that the chromium carbides are removed during boiling and some pits still shows the presence of chromium carbides. The pits are formed due
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(a) FSW-SZ
(b) EBW-FZ
(c) LBW-FZ
(d) GTAW-FZ
Fig. 7. Transmission electron micrographs of weld metal region.
(a) FSW-HTHAZ
(b) GTAW-HTHAZ
(c) EBW-HTHAZ
(d) LBW-HTHAZ
Fig. 8. Transmission electron micrographs of high temperature heat affected zone.
A.K. Lakshminarayanan, V. Balasubramanian / Materials and Design 39 (2012) 175–185 Table 3 Microanalysis of phases obtained from EDAX (wt.%). Joint/phase
GTAW FSW LBW EBW
Ferrrite
Martensite
Cr
Ni
Mn
Cr
Ni
Mn
10.1 11.8 11.4 11.1
0.3 0.4 0.4 0.4
0.8 1.0 1.0 0.8
9.2 9.8 10.0 10.2
0.4 0.7 0.6 0.6
1.0 1.4 1.2 1.4
181
ilar to LBW joint, very fine pits are observed mainly on the martensitic phase and selective dissolution of martensite phase are also observed in few regions (Fig. 10d). Fig. 12 shows the optical and scanning electron microstructure of the HTHAZ of all four welded joints after Modified Strauss test. It can be inferred from, Fig. 12a and e, that the grain boundaries of GTAW joint exhibit continuous ditches and it is an indication of severely sensitized grain boundaries. The ditches are much wider and deeper in the HTHAZ compared to other joints. Corrosion attack penetrated intergranularly almost on all grain boundaries in the entire HTHAZ. In addition to the selective dissolution of martensite, grain boundary ditches are also observed in the HTHAZ of FSW joint (Fig. 12b and f). However the ditches are much shallower compared to the other joints. The mode of attack in the HTHAZ of LBW and EBW joint are also appeared to be intergranular, associated with severe grain dropping (Fig. 12c, d, g and h). Although EBW joint exhibited slightly wider HTHAZ compared to LBW joint, severity of grain boundary ditching and grain dropping is less. 3.6. Microstructure of bend test specimens
Fig. 9. The corrosion mass loss of the 409M welded joints during Strauss test.
to both intragranular and intergranular chromium carbides as shown in Fig. 11a and b. EDAX results (Fig. 11c and d) confirmed these precipitates are rich in chromium. The similar pits are also observed in the FZ of LBW and EBW joint (Fig. 10c and d). However, very fine and shallower pits are observed in the interdendritic region of LBW and EBW FZ and no severe ditches are observed. The SZ of FSW joint exhibited no sensitized boundaries. However, sim-
The optical microstructure of the bent specimen after immersion test is presented in Fig. 13. Some transgranular and intergranular cracks are found in GTAW joint (Fig. 13a). These cracks were propagated in the direction perpendicular to that of applied stress. However no such cracks are found in other joints. It is difficult with the complex range of structures, stresses and corrosion conditions to identify in detail the processes and mechanisms of failure. In some cases, surface corrosion occurs enhanced by the presence of tempered martensite as in the case of FSW joint. But in most cases the attack seems to be essentially intergranular corrosion aided by the residual stress of welding, and more importantly, by the applied stress of bending since it will influence the dissolution of carbides.
(a) GTAW
(b) LBW
(c) EBW
(d) FSW
Fig. 10. Optical micrographs of weld metal region after Strauss test.
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(a) Intragranular carbide precipitates
(b) Intergranular carbide precipitates
(c) EDS pattern of intragranular carbides (d) EDS pattern of intergranular carbides
(e) Selective Martensite dissolution Fig. 11. Carbide precipitation and martensite dissolution in 409M FSS joints.
Though the microstructural analysis indicates that, the region nearer to the fusion line of LBW and EBW joint were sensitized, no cracks were found in this region after bend test. This is mainly due to the narrower weld bead of LBW and EBW joints. Most of the plastic deformation will be concentrated in the base metal region with minimum in the narrower weld bead resulted in crack free 180 smooth bends. However, it can be inferred that due to the large amount of carbide precipitation and severity of chromium depletion (width and depth of depleted region) in GTAW joint, the cracks were developed in the bent specimens during the boiling immersion testing is due to the sensitization effect, which is also confirmed by the SEM microstructure of the heat affected zone (Fig. 12e). 4. Discussion In the case of FSS, the sensitization takes place above 950 °C where, there is a marked increase in solubility of carbon in ferrite. On cooling, the solubility of carbon is strongly reduced, and carbide precipitation in the temperature range of approximately 400– 950 °C is too rapid for even water quenching to suppress it. At
the lower end of the precipitation range (400–700 °C), chromium depletion occurs unless the material is held for extended periods of time because chromium diffusion rates are low at these temperatures. In the range 700–950 °C, chromium diffuses more rapidly, which decreases the degree of chromium depletion and intergranular susceptibility [9]. Sensitization behaviour of 12 wt.% Cr ferritic stainless steel is complicated by the partial transformation of d-ferrite to austenite on cooling. During slow cooling or annealing below the A1 temperature, this austenite decomposes to form desensitized ferrite and M23C6-type carbide precipitates. The rapid cooling rates associated with welding, however, prevent the transformation of austenite to ferrite at lower temperatures, and any austenite formed on cooling transforms to martensite below the Ms-temperature. Due to its low solubility in ferrite, the majority of the carbon precipitates as chromium-rich carbides or carbonitrides during annealing, but any chromium-depleted zones formed in the ferrite are healed through rapid chromium back-diffusion from the grain interiors [10]. Two primary variables determining the location and width of the sensitization zone are peak temperature distribution and time
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Joint
Optical Micrographs
Scanning Electron Micrographs
(a)
(e)
(b)
(f)
(c)
(g)
(d)
(h)
183
GTAW
FSW
LBW
EBW
Fig. 12. Heat affected zone microstructure after Strauss test.
spent between these temperatures. Although there will always be a region in the heat affected zone of any weldment that reaches these temperatures, it is possible to adjust the time at temperature by appropriate choice of the welding parameters. In welding, it is often assumed that the location of the sensitization zone is determined solely by the peak temperature distribution. Of the four welding processes used in the investigation, the WM and HTHAZ of friction stir welded AISI 409M ferritic stainless steel joint exhibited lower degree of sensitization (DOS) compared to other joints. However, it suffers from general corrosion attack by martensite dissolution. The HTHAZ of GTAW exhibited very high DOS compared to all other joints. The reason for this behaviour can be explained with the individual welding process characteristics and changes in the microstructure of WM and HAZ due to welding.
4.1. Effect of welding processes on sensitization behaviour In general, the heating and cooling rates experienced by a weldment depend on the characteristics of the heat source used in the welding process, the thickness and geometry of the weldment, and the initial/inter-pass temperatures [11,12]. Though the GTAW, EBW and LBW are fusion welding processes, it is well known that there are appreciable differences in the thermal cycles in the FZ and HAZ regions and are the result of a difference in the power density and heat input supplied by these processes. On the other hand, FSW is a relatively new solid state welding process and hence the heat input is lower when compared to the GTAW process. It is reasonable to expect that the different thermal cycles experienced by WM and HAZ of the these weldments can result in significant microstructural differences among
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(a) GTAW
(b) FSW
(c) EBW
(d) LBW
Fig. 13. Microstructure of the bent specimen after Strauss test.
them, and therefore affect their respective corrosion resistance properties. From Table 2, it is evident that the heat input to the LBW joint (84 J/mm) is significantly lower than the heat input to the GTAW (1800 J/mm), EBW (224 J/mm) and FSW (830 J/mm) joints and thus the LBW joint cools far more rapidly. According to the recent study by Greef [13], welding of modified 12 wt.% Cr ferritic stainless steel with too low heat input or too high heat input result in sensitization of grain boundaries. The results obtained from this study by Modified Strauss test showed that the GTAW joint exhibited higher DOS followed by LBW and EBW joints. FSW joints showed minimum DOS due to the intermediate heat input supplied during welding. Low heat input welding such as LBW results in very fast cooling rates during the early stages of the weld thermal cycle. These rapid cooling rates can suppress austenite nucleation as the heataffected zone cools through the dual-phase (d + c) field, resulting in almost fully ferritic HTHAZ microstructures. Due to the low solubility of carbon in ferrite, the ferrite phase becomes supersaturated in carbon in the absence of sufficient austenite, and extensive carbide precipitation occurs at the ferrite–ferrite grain boundaries during cooling. The fast cooling rates also prevent the back-diffusion of chromium to the depleted regions adjacent to the chromium-rich carbides, resulting in a continuous network of sensitized ferrite–ferrite grain boundaries. However, due to maximum peak temperature caused higher heat input supplied by GTAW process, grain growth was occurred in the FZ and HTHAZ with majority of ferrite–ferrite grain boundaries and ferrite with partially covered martensite decorated by chromium carbides. Also the steel used in this investigation is an unstabilized grade and sensitization of austenite took place under very slow cooling condition with welding at excessively high heat input levels as in case of GTAW joint. In the study by Lu et al. [14], a low heat input and high cooling rate of LBW is likely to reduce the micro segregation of alloying ele-
ments and the formation of Cr-depleted zones, resulting in improvement of the resistance to localized corrosion. Similar studies made by other researchers also showed the same effect of the LBW process [15,16]. Thus, the fine microstructure of the FZ and HTHAZ produced by LBW process is less susceptible to localized corrosion, as compared to the coarse microstructure produced by the GTAW process. The FZ of LBW and EBW joints showed smaller DOS compared to the HTHAZ and this is mainly due to relatively high peak temperature which resulted in the formation of sufficient low carbon martensite in the FZ. If enough austenite forms to absorb excess carbon (austenite has higher carbon solubility than ferrite), a continuous network of chromium-depleted zones does not form and sensitization is minimized. The SZ of FSW joint exhibited the lowest DOS compared to all other joints. Significant corrosion attack was observed in the martensite phase instead of typical intergranular corrosion. However, the ferrite grains also appear to be marginally affected by corrosive attack along what appear to be sub-grain boundaries. This level of corrosive attack might suggest that the material was not sensitized, but rather that preferential corrosion of the martensite was occurring due to the reduced chromium content. The results of EDAX analysis revealed 11.8 wt.% and 9.8 wt.% chromium in the ferrite and martensite, respectively, consistent with the expected partitioning between the two phases as reported by Gooch [17]. Williams and Barbaro [18] stated that martensite introduces intergranular boundaries which are highly favourable for carbide precipitation. However, the presence of martensite increases the phase boundary area due to the inherently fine grain size and due to carbon partioning into austenite (and hence martensite), less carbon is available to precipitate elsewhere in the microstructure. Carbide precipitations occurred within the martensite reduces general corrosion resistance, but is not expected to produce sensitization. The energy associated with grain boundaries makes them favourable sites for solute segregation and precipitation reactions,
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either of which can lead to corrosion at the grain boundary. Grain boundary energy is a function of grain boundary structure and therefore varies among the boundaries of a various regions. This can cause variations in the extent of these processes at the grain boundaries [19]. Finer grain of SZ of FSW joint is also one reason for improved sensitization resistance compared to other joints. The SZ of friction stir welded AISI 409M ferritic stainless steel joint consists of fine dual phase structure of ferrite and martensite of order of 4 lm. On the other hand, the grain size of heat affected zone is almost similar in size of 30 lm with discontinuous network martensite. It is well known that the sensitization resistance of fine grained microstructure is higher compared to the coarse grain microstructure. As inter-granular corrosion is caused by the precipitation of carbides in grain boundaries, the corrosion rate is affected by the volume fraction of precipitated carbides per unit of grain boundary area. As grain size is refined, the grain boundary areas per unit volume increase and the degree of the chromium depletion caused by carbide precipitation will decrease for a given carbon content. Hence, boundaries may not be sensitized in finely grained materials [20]. Rodriguez et al. [21] studied the sensitization behaviour of friction stir welded 304 stainless steel and it was found that, the SZ has lower DOS compared to the base metal. The author stated that the high deformation and temperature during FSW caused the grains with a high density of dislocations to become saturated with more dislocations. Therefore, new free deformation grains are generated, and the grain size is reduced. Due to grain size reduction and the microstructural defects generated in the grains, intergranular corrosion decreased in the weld zone in comparison to the base metal. Therefore, after increasing free deformation grains, the diffusion paths for chromium and carbon atoms decreased. Hence, from the above discussion, FSW joint is less prone to sensitization and suffered from general corrosion by martensite dissolution. LBW joint is prone to sensitization. However, the weight loss method is minimum compared to the other joints due to reduced weld metal and heat affected zone size. GTAW joint showed poorer resistance against intergranular corrosion attack. 5. Conclusions The effect of welding processes on sensitization behaviour was studied using Modified Strauss test and the following important conclusion were derived from this investigation: Of the four joints, the joint fabricated by GTAW process exhibited higher corrosion rate followed by FSW, LBW and EBW joints. Severe grain dropping due to intergranular corrosion and general corrosion attack in the fusion zone are the reasons for the higher corrosion rate of GTAW joints. On the other hand FSW joint suffered from preferential dissolution of martensite phase. The high temperature heat affected zone (HTHAZ) of all the four joints suffered from sensitization. Of the four joints, the joint fabricated using GTAW process showed higher DOS and least sensitization resistance compared to other joints. This is mainly is due to the ferrite grain growth with grain boundary chromium carbides caused by higher heat input supplied by the process. The HTHAZ of LBW joint showed higher DOS compared to the EBW and FSW joints. This is due to very fast cooling rates associated with LBW process during the early stages of the weld thermal cycle. These rapid cooling rates suppress the nucleation of austenite and also prevent the back-diffusion
185
of chromium to the depleted regions adjacent to the chromium-rich carbides, resulting in a continuous network of sensitized ferrite–ferrite grain boundaries. Even though sensitization could not be eliminated by any of these advanced welding processes, it could be seen that reducing the size of the heat-affected zone dramatically reduced the sensitized regions. In this way, the LBW joint offered better overall corrosion resistance compared to other joints.
Acknowledgements The authors wish to place their sincere thanks to the Department of Science and Technology (DST), Government of India for the financial support through Fast Track Scheme for Young Scientists R&D project (SR/FTP/ETA043/2009) to carry out this investigation. Also the authors are to grateful to Dr. G. Madhusudhan Reddy, Scientist-‘G’ and Head, Metal Joining Group, Defence Metallurgical Research Laboratory (DMRL), Hyderabad, for this support and guidance to carry out this investigation. References [1] Lakatos-Varsanyi M, Meisel W. Corrosion studies of chromium steel in imitated seawater. J Radioanal Nucl Chem 2000;251:75–85. [2] Chong PH, Liu Z, Skeldon P, Crouse P. Characterisation and corrosion performance of laser-melted 3CR12 steel. Appl Surf Sci 2005;247:362–8. [3] Silva CC, Farias JP, Miranda HC, Guimaraes RF, Menezes JWA, Neto MAM. Mater Charact 2008;59:528. [4] Tomlinson WJ, Matthews SJ. Intergranular corrosion of welds in type 405 stainless steel. J Mater Sci 1988;23:2064–8. [5] Greef ML, Du Toit M. Looking at the sensitization of 11–12% chromium EN 1.4003 stainless steels during welding. Weld J 2006. 243-s–251-s. [6] Van Warmelo M, Nolan D, Norrish J. Mitigation of sensitisation effects in unstabilized 12% Cr ferritic stainless steel welds. Mater Sci Eng A 2007;464: 157–69. [7] Du Toit M, Naude J, Norrish J. The influence of stabilization with titanium on the heat affected zone sensitization of 11–12% chromium ferritic stainless steels under low heat input welding conditions. Int Inst Weld 2009;IIW Doc. IX-H:705–9. [8] ASTM A763-93 (Reapproved 2009). Standard practices for detecting susceptibility to intergranular attack in ferritic stainless steels. ASTM international; 2009. p. 6–18. [9] Sourmail T, Too CH, Bhadeshia HKDH. Sensitisation and evolution of chromium-depleted zones in Fe–Cr–Ni–C systems. ISIJ Int 2003;43: 1814–20. [10] Irvine KJ, Crowe DJ, Pickering FB. The physical metallurgy of 12% Cr steels. J Iron Steel Inst 1960;8:43–62. [11] Lee HT, Wu JL. Correlation between corrosion resistance properties and thermal cycles experienced by gas tungsten arc welding and laser beam welding Alloy 690 butt weldments. Corr Sci 2009;51:733–43. [12] Zheng H, Ye X, Jiang L, Wang B, Liu Z, Wanga G. Study on microstructure of low carbon 12% chromium stainless steel in high temperature heat-affected zone. Mater Des 2010;31:4836–41. [13] Greef ML. The influence of welding parameters on the sensitisation behaviour of 3CR12. M.S. thesis, University of Pretoria; 2006. [14] Lu BT, Chen ZK, Luo JL, Patchett BM, Xu ZH. Pitting and stress corrosion cracking behavior in welded austenitic stainless steel. Electrochim Acta 2005;50:1391–403. [15] Lee HT, Wu JL. The effects of peak temperature and cooling rate on the susceptibility to intergranular corrosion of alloy 690 by laser beam and gas tungsten arc welding. Corr Sci 2009;51:439–45. [16] Liu Z, Chong PH, Skeldon P, Hilton PA, Spencer JT. Fundamental understanding of the corrosion performance of laser-melted metallic alloys. Surf Coat Technol 2006;200:5514–25. [17] Gooch TG. Corrosion behaviour of welded stainless steels. Weld J 1996;75: 135s–54s. [18] Williams JG, Barbaro FJ. Susceptibility and prevention of HAZ sensitization and intergranular stress corrosion cracking in various 12% Cr steels. In: Proc. of the iron and steel technol conf (AIstech ‘05), Charlotte, NC, USA; 2005. [19] Ogwu AA, Davis TJ. Improving the sensitization resistance of ferritic stainless steel. Scripta Mater 1997;37:259–63. [20] Dischino A, Kenny JM. Effect of grain size on the corrosion resistance of a high nitrogen–low nickel austenitic stainless steel. J Mater Sci Lett 2002;21: 1969–71. [21] Rodriguez NA, Almanza E, Jesus Perez M, Rodrigo Muniz C, Packer S. Steel R analysis of sensitization phenomenon in friction stir welded 304 stainless steel. Front Mater Sci China 2010;4:415–9.
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Assessment of sensitization resistance of AISI 409M grade ferritic stainless steel joints using Modified Strauss test A.K. Lakshminarayanan ⇑, V. Balasubramanian Centre for Materials Joining and Research (CEMAJOR), Department of Manufacturing Engineering, Annamalai University, Annamalai Nagar 608 002, Tamil Nadu, India
a r t i c l e
i n f o
Article history: Received 14 December 2011 Accepted 18 February 2012 Available online 28 February 2012 Keywords: D. Welding E. Corrosion F. Microstructure
a b s t r a c t The paper reports on the findings from a detailed study of the joints fabricated by four different welding processes namely, gas tungsten arc welding, friction stir welding, laser beam welding and electron beam welding to assess the susceptibility to intergranular corrosion. Bend test and weight loss methods were used to evaluate the degree of sensitization after Modified Strauss test. The results are correlated with the microstructure by optical, scanning and transmission microscopy and the chemical composition by energy dispersive X-ray spectrometry. Friction stir welded joint is less prone to sensitization, but suffered from general corrosion by dissolution of martensite phase. Laser beam and electron beam welded joints exhibited lowest corrosion rate as compared to other joints. Gas tungsten arc welded joint is prone to sensitization and showed lower corrosion resistance compared to its counterparts. Ó 2012 Elsevier Ltd. All rights reserved.
1. Introduction Binary Iron–Chromium (Fe–Cr) alloys with a chromium content of 10–15% are found to exhibit an excellent corrosion resistance under ambient conditions and, because being less costly, they may substitute FeCrNi stainless steels for many applications [1]. Welding is an important method of fabricating complex plant, and the stability of welds in stainless steel is closely related to the composition of the steel, the thermal conditions associated with welding, and any stress that may be present in service [2]. The thermal cycle produced during welding can cause carbide precipitation and various phases in the heat-affected zone (HAZ). Corrosion resistance can also be altered mainly when phenomenon such as sensitization is likely to occur [3]. Some studies were reported on sensitization behaviour of ferritic stainless steel (FSS) joints fabricated using conventional welding processes such as gas tungsten arc welding, gas metal arc welding and shielded metal arc welding processes. Tomlinson and Matthews [4] investigated the sensitization behaviour of unwelded and gas tungsten arc welded flat and bent AISI 405 stainless steel specimens by chemical, electrochemical techniques and they reported that the HAZ was susceptible to intergranular attack and can be eliminated by proper annealing treatment. Greef and Du Toit [5] made an attempt to determine the extent to which thick gauge 3CR12 material was susceptible to sensitisation, to ⇑ Corresponding author. Tel.: +91 4144 239734, mobile: +91 9865431106; fax: +91 4144 238080/238275. E-mail addresses: [email protected] (A.K. Lakshminarayanan), balasubrama [email protected] (V. Balasubramanian). 0261-3069/$ - see front matter Ó 2012 Elsevier Ltd. All rights reserved. doi:10.1016/j.matdes.2012.02.038
characterize its occurrence in multiple pass welds, and to determine to what extent it could be minimized by appropriate control of the gas tungsten arc welding parameters. It was concluded that, the low heat inputs and fast cooling rates reduces the level of sensitisation. The susceptibility of 11–12% chromium type EN1.4003 FSS to sensitization during continuous cooling after welding at low heat input levels was investigated by Van Warmelo et al. [6] and it was found that welding at low heat inputs can suppress the transformation of ferrite to austenite as the HAZ cools through the (austenite + ferrite) dual phase region during welding. Sensitization is prevented by the presence of enough austenite to eliminate continuous ferrite–ferrite grain boundaries. Du Toit and Naude [7] studied the sensitization behaviour of unstabilized and titanium stabilized FSS and reported that the unstabilized 1.4003 steel contained considerably more grain boundary martensite in the high temperature heat affected zone (HTHAZ) after low heat input welding than the titanium stabilized grade, despite having very similar kaltenhauser ferrite factor (KFF). The presence of grain boundary martensite in the unstabilized grade increased the resistance to sensitization under low heat input conditions. Using AISI 409M grade ferritic stainless steel for structural application in corrosive environments will be economical as it imparts acceptable corrosion resistance using minimum Cr content. However, sensitization of stainless steel during thermo mechanical processing and welding has been a major challenge to face. Stainless steels like AISI 409M grade, which are having low chromium content (with 11.4%), are susceptible to intergranular corrosion when improper welding conditions are used. The susceptibility of
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up. The experimental setup is displayed in Fig. 1. Trials were carried out to find the optimum immersion conditions to compare the corrosion attack of joints fabricated by four different welding processes. Improper testing condition (i.e., acid concentration, immersion time and amount of copper shots) resulted in aggressive attack as displayed in Fig. 2. Initially, the immersion time was varied from 10 to 25 h and the corrosion rate was calculated. However, the specimens (shown in Fig. 3) boiled for 20 h were further analyzed for comparison among four different welding processes. Both weight loss and bend test were used to evaluate the degree of sensitization. Evaluation by bending is preferred for welds since mass loss of different welds at a range of heat inputs would not give the real effect of intergranular corrosion. Bending reveals embrittlement of the weld and heat affected zone due to chromium depletion at the grain boundaries. The specified bend test requires the sample to be bent through 180 over a radius equal to twice the thickness of the specimen. The appearance of fissures or cracks indicates the presence of intergranular attack. In this study, the welded joints were grounded to 3 mm and bent to 180 with a bent radius of 6 mm at constant deformation. The bent specimens are displayed in Fig. 4. The U bend prepared from the welded joints were subjected to boiling immersion testing for 24 h. After immersion testing, the specimens were cleaned with soft brush to remove the copper deposits and polished with different grades of fine emery papers and then analyzed using optical and scanning electron microscopy.
these alloys is strongly affected by welding processes, which change the microstructure of the alloy in order to have optimum mechanical properties. However, a detailed comparison of different welding processes on the sensitization behaviour of 409M grade FSS welds are hitherto not reported and hence the present investigation was carried out to understand the effect of welding processes namely, friction stir welding (FSW), laser beam welding (LBW), electron beam welding (EBW) and gas tungsten arc welding (GTAW) on sensitization behaviour of AISI 409M grade stainless steel. 2. Experimental procedure 2.1. Base metal The as-received base metal (BM) used in this study was 4 mm thick cold rolled, annealed and pickled AISI 409M grade FSS plates. The chemical composition of the base metal obtained using vaccum spectrometer is presented in Table 1. 2.2. Welding Square butt joints were prepared to fabricate GTAW, LBW, EBW and FSW processes without filler metal additions. GTAW joints were fabricated using Lincoln welding machine (USA) with a capacity of 400 A. LBW joints were fabricated using a CO2 laser beam welding machine (Rofin Slab: CO2 laser). EBW joints were fabricated using an electron beam welding machine (Techmeta, France) with a capacity of 100 kV. FSW joints were fabricated using an indigenously designed and developed FSW machine (22 kW; 4000 rpm; 6 Tonne) using a tungsten based alloy tool. Few welding trials were carried out and specimens were extracted from various locations of the joint and subjected to macrostructural analysis. The specimen free of volumetric defect and lack of penetration was considered as the optimized welding condition. The welding conditions and optimized process parameters presented in Table 2 were used to fabricate the joints for further investigation.
2.5. Microstructural analysis Microstructure of the as-welded specimens prior to boiling immersion test was examined using transmission electron microscopy (TEM). Thin foils with a thickness of about 1 mm were cross sectioned from the welded joints to prepare the TEM samples. After being mechanically ground to approximately 80 lm, the foils were further ground to a thickness of 15 lm by a dimple grinding machine. Then the electron transparent thin sections were electrolytically prepared by the twin-jet polisher. An electrolyte of 10% perchloric acid +90% ethanol solution ( 25 °C) was used during thinning. These thin foils were observed under a 200 kV potential using a Transmission Electron Microscope (Make: Philips, UK: Model: CM20). Microstructure of the flat and U bent specimens prepared from the welded joint after boiling immersion test was analyzed using optical and scanning electron microscopy. Energy dispersive X-ray Spectrometry (EDAX) analysis was performed to determine the distribution of elements in the ferrite and martensite phases of the weld metal region of all the joints.
2.3. Specimen preparation The welded joints were sliced using abrasive cutting and then machined to the required dimensions for preparing Strauss test specimens as per ASTM A763-93 [8] guidelines. The slices derived from the welded joints were polished using standard metallographic procedures to obtain flat and scratch free surface. 2.4. Modified Strauss test The Modified Strauss test is a variation on ASTM A763-93 Practice Z [8] for detecting susceptibility to sensitization in ferritic stainless steels. Samples were placed in a copper sulphate solution (60 g CuSO4 in 1 l of water) which was acidified by adding 0.5% sulphuric acid (3 ml H2SO4 in 1 l of water). A layer of copper shot was placed on the bottom of the vessel and the samples were placed in the solution in such a way to prevent direct contact with the welded region. The polished specimen were placed in the Erlenmeyer flask, so that the polished surface was horizontal and facing
3. Results 3.1. Macrostructure The cross sectional macrostructure of the welded joints are shown in Fig. 5. No volumetric defect is observed. Laser and electron beam weldments exhibited very thin weld beads (1.5 and 2.5 mm respectively) as compared with much wider beads of GTAW and FSW (10 and 20 mm respectively).
Table 1 Chemical composition (wt.%) of base metal (spectrometry results). Material
C
Mn
P
S
Si
Cr
Ni
Ti
Fe
Base metal (AISI 409M)
0.028
1.10
0.030
0.010
0.40
11.40
0.39
0.004
Bal.
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A.K. Lakshminarayanan, V. Balasubramanian / Materials and Design 39 (2012) 175–185 Table 2 Welding conditions and process parameters. Parameters
Process
Welding machine Tungsten electrode diameter (mm) Power (kW) Current Voltage Focal length (mm) Focal spot diameter (lm) Shielding gas Gas flow rate (l/min) Gun to work distance (mm) Vacuum (bar) Welding speed (mm min-1) Tool rotational speed (rpm) Axial force (kN) Tool material composition Tool pin profile Shoulder diameter (mm) Pin diameter (mm) Pin length (mm) Heat input (J/mm)
FSW
LBW
EBW
GTAW
RV machine tools, India – – – – – – Argon 8 – – 110 850 24 W 90%, Ni 3%, Mo 1%, Co 1%, Fe bal Taper cylindrical 18 Tapered from 8 to 5 3.7 828
Rofin, Germany – 3.5 – – 300 180 Helium 10 – – 3000 – – – –
Techmeta, France – – 51 mA 55 kV – – – – 298 10 4 1000 – – – – – – – 224
Lincoln, USA 3 – 120 A 20 V – – Argon 16 – – 80 – – – – – – – 1800
Fig. 1. Setup for Modified Strauss test.
3.2. Microstructure of as welded specimen SEM image (Fig. 6a) and TEM image (Fig. 6b) of base metal shows the characteristics of grain boundaries with ferrite grains and the presence of Cr23C6 carbides. TEM image of stir zone (Fig. 7a) comprises the banded microstructure of ferrite and martensite. Stirzone (SZ) consisting of a mixed deformation structure with equiaxed ultrafine grains and the cellular substructure between martensite laths. Severe deformation due to friction stirring resulted in the formation of fine subgrained structure, which propagates throughout the ferrite grains. This strain-induced sub microcrystalline structure involves highly misoriented equiaxed grains with a relatively low dislocation density in their interiors as compared with high dislocation densities in cell substructures that evolve at smaller strains. The presence of high dislocation density in the substructure discloses that SZ had undergone strain hardening during welding. TEM image of FZ of EBW joint (Fig. 7b) clearly shows the presence of prior d ferrite grain boundary, substructure and subgrain structure. The presence of chromium carbides and grain boundary martensite are also evidenced from Fig. 7b. This suggests that the equiaxed coarse grain structure base metal undergone a refinement during weld thermal cycle. Similar microstructure is observed in the FZ of LBW joints. The difference in FZ microstruc-
– 84
ture between EBW joint and LBW joint is the degree of fineness in the subgrain structure. In LBW joint (Fig. 7c), the ferrite lath width is comparatively smaller than the width of the ferrite lath evolved in EBW joint. Coarse lath ferrite structure with partially covered needle shape martensite and with chromium carbides decorated the grain boundaries as observed in case of GTAW joint (Fig. 7d). The HTHAZ of all four joints was analyzed using TEM and the resultant images are presented in Fig. 8. The HTHAZ of FSW joint (Fig. 8a) also shows a duplex structure of ferrite and martensite. The martensite was formed along many sub-grain boundaries of the ferrite as well as in grain boundaries. Chromium carbides are also observed in some of the grain boundaries. Dislocations are observed both in ferrite and martensite of this region. It indicates that this region undergone partial plastic deformation during the processes. The HTHAZ of GTAW joint (Fig. 8b) shows a mixed microstructure of ferritic grains with both intra and intergranular chromium carbides. The growth of ferrite was accompanied by the precipitation of carbides at moving interface and this result in the aligned precipitation which follows the shape of ferrite grain. Coarse plate like martensite and needle shape martensite along the grain boundaries of ferrite are also observed in the HTHAZ. The HTHAZ of LBW and EBW joints (Fig. 8c and d) shows mostly of ferritic structures enriched with grain boundary chromium carbides. Low carbon lath martensite is also observed at the grain boundaries. 3.3. Composition analysis of the weld regions from EDAX analysis Liquid–solid and solid-state phase transformations during welding are generally associated with the partitioning of alloying elements between the phases. The extent of this partitioning is dependent on the transformation temperature and the weld cooling rate. A greater extent of partitioning will occur at high temperatures where diffusion of alloying elements occurs most rapidly. Since such elemental partitioning strongly influences the final properties of each phase and the alloy in general, particularly the mechanical and corrosion properties, it is important to determine the extent of any such changes due to welding. The sole purpose of the EDAX analysis was to verify subtle differences in elemental composition between the two phases namely ferrite and martensite. The results of EDAX analysis of the microstructures developed during the d ferrite to austenite transformation in fusion welded
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(a) Copper deposit over LBW specimen (b) Cross section of LBW joint after the removal of copper deposit
(c) Cross section of GTAW joint after the removal of copper deposit
(d) FSW joint after the removal of copper deposit
Fig. 2. Aggressive corrosion attack due to improper testing condition.
(a) GTAW
(b) LBW
(c) FSW
(d) EBW
Fig. 3. Top and cross section of welded joints under optimum testing condition.
(EBW, LBW and GTAW) joints indicated element partitioning between the ferrite and austenite phases during the transformation. It is found that the main ferrite stabiliser, chromium, partitioned to the ferrite phase, whereas the austenite stabilizing elements, nickel and manganese, partitioned to the transformed austenite phase. Though the temperature during FSW did not reached the d ferrite range, the welding was taken place at approximately 1150 °C, where the maximum amount austenite forms in the steel used in the present study. Despite of the absence of phase transformation from d ferrite to austenite, the EDAX analysis showed the redistribution of chromium and nickel in the ferrite and martensite phases. The compositions obtained for the ferrite and martensite phases in the welded joints are presented in Table 3. For all conditions, higher levels of chromium and lower levels of nickel were observed in the ferrite regions of all the welded joints. However, there were no significant differences between the welded joints.
Fig. 4. Bend test specimens after boiling immersion test.
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179
(i) GTAW
(ii) LBW
(iii) EBW (a – basemetal, b – fusion zone (FZ), c – high temperature heat affected zone (HTHAZ), d – low temperature heat affected zone (LTHAZ))
(iv) FSW (a – basemetal (BM), b – flow arm: shoulder influenced region, c – banded structure : pin influenced region, d, e – thermo mechanical affected zone, f- high temperature heat affected zone (HTHAZ), AS – advancing side, RS – retreating side). Fig. 5. Cross-sectional macrostructure of 409M ferritic stainless steel welded joints.
Fig. 6. SEM and TEM image of the basemetal.
3.4. Immersion test
3.5. Microstructure of flat specimens after boiling immersion testing
The immersion tests were conducted for all the four welded joints by varying the immersion time from 10 to 25 h. The solution attacks both chromium carbides and matrix. When the carbides are in the grain boundaries and are attacked by the corrosion medium, the alloy suffers a weight loss. The weight loss method was used to calculate the corrosion rate as per ASTM A763-93 [8] guidelines. The relationship between the immersion time and the corrosion rate is plotted in Fig. 9. It can be inferred that the corrosion rate of LBW specimen is much lower when compared to the other joints. The GTAW joints exhibited higher corrosion rate when compared to its counterpart.
For comparison, among the welded joints, the weld metal (WM) region and HTHAZ of the specimens boiled at an immersion time of 20 h was analyzed using optical and scanning electron microscopy and are presented in Figs. 10–12. The WM region of all the four joints after boiling immersion test is compared in Fig. 10. It can be seen that the FZ of the GTAW joint (Fig. 10a) contains ditches in many regions (mostly ferrite–martensite boundaries) and severe pit holes of various sizes in the WM region (mostly in ferrite phase). Some of pits show that the chromium carbides are removed during boiling and some pits still shows the presence of chromium carbides. The pits are formed due
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(a) FSW-SZ
(b) EBW-FZ
(c) LBW-FZ
(d) GTAW-FZ
Fig. 7. Transmission electron micrographs of weld metal region.
(a) FSW-HTHAZ
(b) GTAW-HTHAZ
(c) EBW-HTHAZ
(d) LBW-HTHAZ
Fig. 8. Transmission electron micrographs of high temperature heat affected zone.
A.K. Lakshminarayanan, V. Balasubramanian / Materials and Design 39 (2012) 175–185 Table 3 Microanalysis of phases obtained from EDAX (wt.%). Joint/phase
GTAW FSW LBW EBW
Ferrrite
Martensite
Cr
Ni
Mn
Cr
Ni
Mn
10.1 11.8 11.4 11.1
0.3 0.4 0.4 0.4
0.8 1.0 1.0 0.8
9.2 9.8 10.0 10.2
0.4 0.7 0.6 0.6
1.0 1.4 1.2 1.4
181
ilar to LBW joint, very fine pits are observed mainly on the martensitic phase and selective dissolution of martensite phase are also observed in few regions (Fig. 10d). Fig. 12 shows the optical and scanning electron microstructure of the HTHAZ of all four welded joints after Modified Strauss test. It can be inferred from, Fig. 12a and e, that the grain boundaries of GTAW joint exhibit continuous ditches and it is an indication of severely sensitized grain boundaries. The ditches are much wider and deeper in the HTHAZ compared to other joints. Corrosion attack penetrated intergranularly almost on all grain boundaries in the entire HTHAZ. In addition to the selective dissolution of martensite, grain boundary ditches are also observed in the HTHAZ of FSW joint (Fig. 12b and f). However the ditches are much shallower compared to the other joints. The mode of attack in the HTHAZ of LBW and EBW joint are also appeared to be intergranular, associated with severe grain dropping (Fig. 12c, d, g and h). Although EBW joint exhibited slightly wider HTHAZ compared to LBW joint, severity of grain boundary ditching and grain dropping is less. 3.6. Microstructure of bend test specimens
Fig. 9. The corrosion mass loss of the 409M welded joints during Strauss test.
to both intragranular and intergranular chromium carbides as shown in Fig. 11a and b. EDAX results (Fig. 11c and d) confirmed these precipitates are rich in chromium. The similar pits are also observed in the FZ of LBW and EBW joint (Fig. 10c and d). However, very fine and shallower pits are observed in the interdendritic region of LBW and EBW FZ and no severe ditches are observed. The SZ of FSW joint exhibited no sensitized boundaries. However, sim-
The optical microstructure of the bent specimen after immersion test is presented in Fig. 13. Some transgranular and intergranular cracks are found in GTAW joint (Fig. 13a). These cracks were propagated in the direction perpendicular to that of applied stress. However no such cracks are found in other joints. It is difficult with the complex range of structures, stresses and corrosion conditions to identify in detail the processes and mechanisms of failure. In some cases, surface corrosion occurs enhanced by the presence of tempered martensite as in the case of FSW joint. But in most cases the attack seems to be essentially intergranular corrosion aided by the residual stress of welding, and more importantly, by the applied stress of bending since it will influence the dissolution of carbides.
(a) GTAW
(b) LBW
(c) EBW
(d) FSW
Fig. 10. Optical micrographs of weld metal region after Strauss test.
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(a) Intragranular carbide precipitates
(b) Intergranular carbide precipitates
(c) EDS pattern of intragranular carbides (d) EDS pattern of intergranular carbides
(e) Selective Martensite dissolution Fig. 11. Carbide precipitation and martensite dissolution in 409M FSS joints.
Though the microstructural analysis indicates that, the region nearer to the fusion line of LBW and EBW joint were sensitized, no cracks were found in this region after bend test. This is mainly due to the narrower weld bead of LBW and EBW joints. Most of the plastic deformation will be concentrated in the base metal region with minimum in the narrower weld bead resulted in crack free 180 smooth bends. However, it can be inferred that due to the large amount of carbide precipitation and severity of chromium depletion (width and depth of depleted region) in GTAW joint, the cracks were developed in the bent specimens during the boiling immersion testing is due to the sensitization effect, which is also confirmed by the SEM microstructure of the heat affected zone (Fig. 12e). 4. Discussion In the case of FSS, the sensitization takes place above 950 °C where, there is a marked increase in solubility of carbon in ferrite. On cooling, the solubility of carbon is strongly reduced, and carbide precipitation in the temperature range of approximately 400– 950 °C is too rapid for even water quenching to suppress it. At
the lower end of the precipitation range (400–700 °C), chromium depletion occurs unless the material is held for extended periods of time because chromium diffusion rates are low at these temperatures. In the range 700–950 °C, chromium diffuses more rapidly, which decreases the degree of chromium depletion and intergranular susceptibility [9]. Sensitization behaviour of 12 wt.% Cr ferritic stainless steel is complicated by the partial transformation of d-ferrite to austenite on cooling. During slow cooling or annealing below the A1 temperature, this austenite decomposes to form desensitized ferrite and M23C6-type carbide precipitates. The rapid cooling rates associated with welding, however, prevent the transformation of austenite to ferrite at lower temperatures, and any austenite formed on cooling transforms to martensite below the Ms-temperature. Due to its low solubility in ferrite, the majority of the carbon precipitates as chromium-rich carbides or carbonitrides during annealing, but any chromium-depleted zones formed in the ferrite are healed through rapid chromium back-diffusion from the grain interiors [10]. Two primary variables determining the location and width of the sensitization zone are peak temperature distribution and time
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Joint
Optical Micrographs
Scanning Electron Micrographs
(a)
(e)
(b)
(f)
(c)
(g)
(d)
(h)
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GTAW
FSW
LBW
EBW
Fig. 12. Heat affected zone microstructure after Strauss test.
spent between these temperatures. Although there will always be a region in the heat affected zone of any weldment that reaches these temperatures, it is possible to adjust the time at temperature by appropriate choice of the welding parameters. In welding, it is often assumed that the location of the sensitization zone is determined solely by the peak temperature distribution. Of the four welding processes used in the investigation, the WM and HTHAZ of friction stir welded AISI 409M ferritic stainless steel joint exhibited lower degree of sensitization (DOS) compared to other joints. However, it suffers from general corrosion attack by martensite dissolution. The HTHAZ of GTAW exhibited very high DOS compared to all other joints. The reason for this behaviour can be explained with the individual welding process characteristics and changes in the microstructure of WM and HAZ due to welding.
4.1. Effect of welding processes on sensitization behaviour In general, the heating and cooling rates experienced by a weldment depend on the characteristics of the heat source used in the welding process, the thickness and geometry of the weldment, and the initial/inter-pass temperatures [11,12]. Though the GTAW, EBW and LBW are fusion welding processes, it is well known that there are appreciable differences in the thermal cycles in the FZ and HAZ regions and are the result of a difference in the power density and heat input supplied by these processes. On the other hand, FSW is a relatively new solid state welding process and hence the heat input is lower when compared to the GTAW process. It is reasonable to expect that the different thermal cycles experienced by WM and HAZ of the these weldments can result in significant microstructural differences among
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(a) GTAW
(b) FSW
(c) EBW
(d) LBW
Fig. 13. Microstructure of the bent specimen after Strauss test.
them, and therefore affect their respective corrosion resistance properties. From Table 2, it is evident that the heat input to the LBW joint (84 J/mm) is significantly lower than the heat input to the GTAW (1800 J/mm), EBW (224 J/mm) and FSW (830 J/mm) joints and thus the LBW joint cools far more rapidly. According to the recent study by Greef [13], welding of modified 12 wt.% Cr ferritic stainless steel with too low heat input or too high heat input result in sensitization of grain boundaries. The results obtained from this study by Modified Strauss test showed that the GTAW joint exhibited higher DOS followed by LBW and EBW joints. FSW joints showed minimum DOS due to the intermediate heat input supplied during welding. Low heat input welding such as LBW results in very fast cooling rates during the early stages of the weld thermal cycle. These rapid cooling rates can suppress austenite nucleation as the heataffected zone cools through the dual-phase (d + c) field, resulting in almost fully ferritic HTHAZ microstructures. Due to the low solubility of carbon in ferrite, the ferrite phase becomes supersaturated in carbon in the absence of sufficient austenite, and extensive carbide precipitation occurs at the ferrite–ferrite grain boundaries during cooling. The fast cooling rates also prevent the back-diffusion of chromium to the depleted regions adjacent to the chromium-rich carbides, resulting in a continuous network of sensitized ferrite–ferrite grain boundaries. However, due to maximum peak temperature caused higher heat input supplied by GTAW process, grain growth was occurred in the FZ and HTHAZ with majority of ferrite–ferrite grain boundaries and ferrite with partially covered martensite decorated by chromium carbides. Also the steel used in this investigation is an unstabilized grade and sensitization of austenite took place under very slow cooling condition with welding at excessively high heat input levels as in case of GTAW joint. In the study by Lu et al. [14], a low heat input and high cooling rate of LBW is likely to reduce the micro segregation of alloying ele-
ments and the formation of Cr-depleted zones, resulting in improvement of the resistance to localized corrosion. Similar studies made by other researchers also showed the same effect of the LBW process [15,16]. Thus, the fine microstructure of the FZ and HTHAZ produced by LBW process is less susceptible to localized corrosion, as compared to the coarse microstructure produced by the GTAW process. The FZ of LBW and EBW joints showed smaller DOS compared to the HTHAZ and this is mainly due to relatively high peak temperature which resulted in the formation of sufficient low carbon martensite in the FZ. If enough austenite forms to absorb excess carbon (austenite has higher carbon solubility than ferrite), a continuous network of chromium-depleted zones does not form and sensitization is minimized. The SZ of FSW joint exhibited the lowest DOS compared to all other joints. Significant corrosion attack was observed in the martensite phase instead of typical intergranular corrosion. However, the ferrite grains also appear to be marginally affected by corrosive attack along what appear to be sub-grain boundaries. This level of corrosive attack might suggest that the material was not sensitized, but rather that preferential corrosion of the martensite was occurring due to the reduced chromium content. The results of EDAX analysis revealed 11.8 wt.% and 9.8 wt.% chromium in the ferrite and martensite, respectively, consistent with the expected partitioning between the two phases as reported by Gooch [17]. Williams and Barbaro [18] stated that martensite introduces intergranular boundaries which are highly favourable for carbide precipitation. However, the presence of martensite increases the phase boundary area due to the inherently fine grain size and due to carbon partioning into austenite (and hence martensite), less carbon is available to precipitate elsewhere in the microstructure. Carbide precipitations occurred within the martensite reduces general corrosion resistance, but is not expected to produce sensitization. The energy associated with grain boundaries makes them favourable sites for solute segregation and precipitation reactions,
A.K. Lakshminarayanan, V. Balasubramanian / Materials and Design 39 (2012) 175–185
either of which can lead to corrosion at the grain boundary. Grain boundary energy is a function of grain boundary structure and therefore varies among the boundaries of a various regions. This can cause variations in the extent of these processes at the grain boundaries [19]. Finer grain of SZ of FSW joint is also one reason for improved sensitization resistance compared to other joints. The SZ of friction stir welded AISI 409M ferritic stainless steel joint consists of fine dual phase structure of ferrite and martensite of order of 4 lm. On the other hand, the grain size of heat affected zone is almost similar in size of 30 lm with discontinuous network martensite. It is well known that the sensitization resistance of fine grained microstructure is higher compared to the coarse grain microstructure. As inter-granular corrosion is caused by the precipitation of carbides in grain boundaries, the corrosion rate is affected by the volume fraction of precipitated carbides per unit of grain boundary area. As grain size is refined, the grain boundary areas per unit volume increase and the degree of the chromium depletion caused by carbide precipitation will decrease for a given carbon content. Hence, boundaries may not be sensitized in finely grained materials [20]. Rodriguez et al. [21] studied the sensitization behaviour of friction stir welded 304 stainless steel and it was found that, the SZ has lower DOS compared to the base metal. The author stated that the high deformation and temperature during FSW caused the grains with a high density of dislocations to become saturated with more dislocations. Therefore, new free deformation grains are generated, and the grain size is reduced. Due to grain size reduction and the microstructural defects generated in the grains, intergranular corrosion decreased in the weld zone in comparison to the base metal. Therefore, after increasing free deformation grains, the diffusion paths for chromium and carbon atoms decreased. Hence, from the above discussion, FSW joint is less prone to sensitization and suffered from general corrosion by martensite dissolution. LBW joint is prone to sensitization. However, the weight loss method is minimum compared to the other joints due to reduced weld metal and heat affected zone size. GTAW joint showed poorer resistance against intergranular corrosion attack. 5. Conclusions The effect of welding processes on sensitization behaviour was studied using Modified Strauss test and the following important conclusion were derived from this investigation: Of the four joints, the joint fabricated by GTAW process exhibited higher corrosion rate followed by FSW, LBW and EBW joints. Severe grain dropping due to intergranular corrosion and general corrosion attack in the fusion zone are the reasons for the higher corrosion rate of GTAW joints. On the other hand FSW joint suffered from preferential dissolution of martensite phase. The high temperature heat affected zone (HTHAZ) of all the four joints suffered from sensitization. Of the four joints, the joint fabricated using GTAW process showed higher DOS and least sensitization resistance compared to other joints. This is mainly is due to the ferrite grain growth with grain boundary chromium carbides caused by higher heat input supplied by the process. The HTHAZ of LBW joint showed higher DOS compared to the EBW and FSW joints. This is due to very fast cooling rates associated with LBW process during the early stages of the weld thermal cycle. These rapid cooling rates suppress the nucleation of austenite and also prevent the back-diffusion
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of chromium to the depleted regions adjacent to the chromium-rich carbides, resulting in a continuous network of sensitized ferrite–ferrite grain boundaries. Even though sensitization could not be eliminated by any of these advanced welding processes, it could be seen that reducing the size of the heat-affected zone dramatically reduced the sensitized regions. In this way, the LBW joint offered better overall corrosion resistance compared to other joints.
Acknowledgements The authors wish to place their sincere thanks to the Department of Science and Technology (DST), Government of India for the financial support through Fast Track Scheme for Young Scientists R&D project (SR/FTP/ETA043/2009) to carry out this investigation. Also the authors are to grateful to Dr. G. Madhusudhan Reddy, Scientist-‘G’ and Head, Metal Joining Group, Defence Metallurgical Research Laboratory (DMRL), Hyderabad, for this support and guidance to carry out this investigation. References [1] Lakatos-Varsanyi M, Meisel W. Corrosion studies of chromium steel in imitated seawater. J Radioanal Nucl Chem 2000;251:75–85. [2] Chong PH, Liu Z, Skeldon P, Crouse P. Characterisation and corrosion performance of laser-melted 3CR12 steel. Appl Surf Sci 2005;247:362–8. [3] Silva CC, Farias JP, Miranda HC, Guimaraes RF, Menezes JWA, Neto MAM. Mater Charact 2008;59:528. [4] Tomlinson WJ, Matthews SJ. Intergranular corrosion of welds in type 405 stainless steel. J Mater Sci 1988;23:2064–8. [5] Greef ML, Du Toit M. Looking at the sensitization of 11–12% chromium EN 1.4003 stainless steels during welding. Weld J 2006. 243-s–251-s. [6] Van Warmelo M, Nolan D, Norrish J. Mitigation of sensitisation effects in unstabilized 12% Cr ferritic stainless steel welds. Mater Sci Eng A 2007;464: 157–69. [7] Du Toit M, Naude J, Norrish J. The influence of stabilization with titanium on the heat affected zone sensitization of 11–12% chromium ferritic stainless steels under low heat input welding conditions. Int Inst Weld 2009;IIW Doc. IX-H:705–9. [8] ASTM A763-93 (Reapproved 2009). Standard practices for detecting susceptibility to intergranular attack in ferritic stainless steels. ASTM international; 2009. p. 6–18. [9] Sourmail T, Too CH, Bhadeshia HKDH. Sensitisation and evolution of chromium-depleted zones in Fe–Cr–Ni–C systems. ISIJ Int 2003;43: 1814–20. [10] Irvine KJ, Crowe DJ, Pickering FB. The physical metallurgy of 12% Cr steels. J Iron Steel Inst 1960;8:43–62. [11] Lee HT, Wu JL. Correlation between corrosion resistance properties and thermal cycles experienced by gas tungsten arc welding and laser beam welding Alloy 690 butt weldments. Corr Sci 2009;51:733–43. [12] Zheng H, Ye X, Jiang L, Wang B, Liu Z, Wanga G. Study on microstructure of low carbon 12% chromium stainless steel in high temperature heat-affected zone. Mater Des 2010;31:4836–41. [13] Greef ML. The influence of welding parameters on the sensitisation behaviour of 3CR12. M.S. thesis, University of Pretoria; 2006. [14] Lu BT, Chen ZK, Luo JL, Patchett BM, Xu ZH. Pitting and stress corrosion cracking behavior in welded austenitic stainless steel. Electrochim Acta 2005;50:1391–403. [15] Lee HT, Wu JL. The effects of peak temperature and cooling rate on the susceptibility to intergranular corrosion of alloy 690 by laser beam and gas tungsten arc welding. Corr Sci 2009;51:439–45. [16] Liu Z, Chong PH, Skeldon P, Hilton PA, Spencer JT. Fundamental understanding of the corrosion performance of laser-melted metallic alloys. Surf Coat Technol 2006;200:5514–25. [17] Gooch TG. Corrosion behaviour of welded stainless steels. Weld J 1996;75: 135s–54s. [18] Williams JG, Barbaro FJ. Susceptibility and prevention of HAZ sensitization and intergranular stress corrosion cracking in various 12% Cr steels. In: Proc. of the iron and steel technol conf (AIstech ‘05), Charlotte, NC, USA; 2005. [19] Ogwu AA, Davis TJ. Improving the sensitization resistance of ferritic stainless steel. Scripta Mater 1997;37:259–63. [20] Dischino A, Kenny JM. Effect of grain size on the corrosion resistance of a high nitrogen–low nickel austenitic stainless steel. J Mater Sci Lett 2002;21: 1969–71. [21] Rodriguez NA, Almanza E, Jesus Perez M, Rodrigo Muniz C, Packer S. Steel R analysis of sensitization phenomenon in friction stir welded 304 stainless steel. Front Mater Sci China 2010;4:415–9.