See discussions, stats, and author profiles for this publication at: http://www.researchgate.net/publication/222424537Weldability of a 2205 duplex stainless steelusing plasma arc weldingARTICLEinJOURNAL OF MATERIALS PROCESSING TECHNOLOGY FEBRUARY 2007Impact Factor: 2.04 DOI: 10.1016/j.jmatprotec.2006.08.030CITATIONS25DOWNLOADS132VIEWS5504 AUTHORS, INCLUDING:Alejandro UreaKing Juan Carlos University138 PUBLICATIONS 1,038 CITATIONS SEE PROFILEM. V. UtrillaKing Juan Carlos University30 PUBLICATIONS 175 CITATIONS SEE PROFILEAvailable from: Alejandro UreaRetrieved on: 08 August 2015Journal of Materials Processing Technology 182 (2007) 624631Weldability of a 2205 duplex stainless steel using plasma arc weldingA. Ure na, E. Otero, M.V. Utrilla, C.J. M unezDepartamento de Ciencia e Ingeniera de Materiales, Escuela Superior de Ciencias, Experimentales y Tecnologa,Universidad Rey Juan Carlos, 28933 M ostoles, Madrid, SpainReceived 22 June 2005; received in revised form 28 July 2006; accepted 15 August 2006AbstractThis paper reports the determination of optimum welding conditions (welding intensity and travel speed) for butt joints of 2205 duplex stainlesssteel sheets using plasma-arc welding (PAW). Minimum net energy input for proper operative and metallurgical weldabilities is studied using twodifferent welding modes: the melt-in or conduction mode and the keyhole mode. The inuence of the welding parameter for each mode on thedimensions and shape of the welds and on their ferrite contents is investigated. 2006 Elsevier B.V. All rights reserved.Keywords: Plasma arc welding; Duplex stainless steels; Keyhole; Weldability1. IntroductionUse of duplex stainless steels is one the increase thanks totheir combinationof excellent anti-corrosionpropertieswithgood mechanical behaviour, especially in temperature-sensitivecomponents suchas heat exchangers andchemical reactorsused in the chemical and petrochemical industries [13]. Goodmechanical properties (high strength combined with high tough-ness) are associated with the presence of a duplex structure witha good balance in the proportion of austenite/ferrite, which isusually 50/50.Given this increased use of duplex stainless steels, we needto gain a better understanding of those metallurgical factors thatinuenceweldability. Conventional fusionweldingprocessesrequired for construction assembly have a considerable impacton duplex structure, both in the fusion zone (FZ) and in the heataffected zone (HAZ). It is well known that impact toughness ofthe welds in duplex stainless steels decreases with the increaseof-ferrite in the HAZ [4], since the local duplex structure isseverely ferritized by the high peak temperature and by the fastcooling rate of the thermal cycle. Another problem associatedwith fusion welding of these materials is their susceptibility tosolidication cracking, which is greater than that of the 304 Laustenitic stainless steels [5]. The precipitation of undesirableCorresponding author. Tel.: +34 914887015; fax: +34 914888143.E-mail address: alejandro.urena@urjc.es (A. Ure na).phases such as intermetallic compounds, carbides and nitridescancauseadrasticdeteriorationintoughnessandcorrosionresistance, for instance in the case of phase, which has veryfast formation kinetics [6].Therefore it is necessary to assure the continuity of duplexstructure properties across the weld by controlling the phase bal-ance both in the FZ and in the HAZ. For practical applicationof this kind of welded joints, an adequate proportion of ferritein the FZ would be in the range of 3070%. This ferrite contentdepends on the chemical composition of the FZ and the coolingrates of the weld, which are related to the input energy appliedduringwelding[7,8]. Forthisreason, thepresentresearchisintended to determine the optimal welding conditions for auto-genous welding (without ller) of duplex stainless steels whilecontrolling the input energy.Theweldingtechniqueusedinthisresearchisplasmaarcwelding (PAW), in which the electric arc generated between anon-consumabletungstenelectrodeandtheworkingpieceisconstrained using a copper nozzle with a small opening at thetip. By forcing the plasma gas and arc through a constricted ori-ce, the torch delivers a high concentration of energy to a smallarea, giving higher welding speeds and producing welds withhighpenetration/widthratios, thuslimitingtheHAZdimen-sions. ForthesereasonsPAWisaveryuseful techniqueforwelding austenitic steels and can also be applied to duplex stain-less steels [911].Inadditiontothemelt-inweldingmode,whichisusuallyadopted in conventional welding processes (such as gas tungsten0924-0136/$ see front matter 2006 Elsevier B.V. All rights reserved.doi:10.1016/j.jmatprotec.2006.08.030A. Ure na et al. / Journal of Materials Processing Technology 182 (2007) 624631 625arcwelding-GTAW), thekeyholingmodecanalsobeusedinPAWincertainrangesofmetalthickness(e.g.2.56 mm).With a proper combination of orice gas ow, travel speed andwelding current, keyhole forming is possible, allowing higherwelding speeds than GTAW with full penetrations. The presentwork addresses the use of both the melt-in or conductionmode and the keyhole mode of welding, for each of which itdenes the minimum net energy input needed to achieve properoperation and metallurgical weldability in a 3 mm thick 2205duplex stainless steel sheet.2. Experimental procedureParent material used in this research was a 3 mm thick rolled sheet of 2005commercial duplex stainless steel. Its chemical composition in weight percent-agesisgiveninTable1. Priortowelding, sheetsweresolutionannealedat1000C for 30 min and quenched to achieve a homogenized microstructure (asindicated by ASTM A923-A) and eliminate any intermetallic phases [12].Autogenous butt joints were made by transferred plasma arc welding(TPAW). Standard coupons were obtained by welding 25 mm wide 100 mmlong blanks with a square-groove joint conguration, without ller. Work pieceswere welded with the weld bead perpendicular to the rolling direction, so thatweldability would not be affected by a change of grain orientation. Sides to bewelded were wire brushed and degreased with acetone.Welds were produced using plasma welding equipment (Plasmaweld 202)inwhichthetorchisxedtoanautomaticmobilesystem(Miggytrac2000)to control both the travel speed and the nozzle/parent sheet distance. Weldingcurrent (I) and welding speed () were changed to control the welding mode(melt-in or keyhole modes), and the inuence of energy input (H) on duplex steelweldabililty was investigated. Arc voltage (E) was measured during welding, andH was calculated from its average value. All joints were made using a ceramicbackingattachedtotherootsidetocontrolpenetrationandprotection;pureargon was used both as shielding and backing gas. The orice gas was a mixtureof 98% Ar2% H2. Welding conditions, including gas ows and electrode andorice nozzle diameters, are given in Table 2.From the welded coupons, specimens were machined for both macroscopicand microstructural studies. The rst group of coupons was also used to measurehardness proles across the welded joints (according to standards EN 1043-2andEN-ISO6507-1)[13,14]andtocalculatetheferritecontents.Theselastmeasurements were carried out on polished and etched specimens, using a FisherFerriteScope calibrated to IIW secondary standards, although for comparativepurposes prior calculations of austenite content were conducted on the parentsheet using X-ray diffraction (XDR) [15]. X-ray diffraction (XRD) patterns ofparent sheets were obtained using a Philips Xpert PW3040/00 diffractometer(Philips, Netherlands) equipped with a Cu source ( =1.5406A), operating at40 kV and 50 mA and scanning rate of 0.04/s from 2 10 through 120. Datawere processed using the XPert Organizer software.Table 1Composition of 2205 (UNS S32205) duplex stainless steels (in wt.%)Composition (wt.%)C 0.020Si 0.40P 0.021Cr 22.37Ni 5.74Mo 3.20Cu 0.17N 0.171Nb 0.05Mn 1.52S 0.001Fe content has been balanced.Table 2Welding conditions applied for PAW of duplex stainless steel sheetsWelding conditions Melt-inmodeKeyholemodeWelding speed (cm/min)15 3520 4045Welding current (A)100 75125150Shielding gas nozzle/workpiece distance (mm)3 34Orice gas nozzle diameter (mm)1.75 1.152.25Tungsten electrode diameter (mm) 2.40 2.40Orice gas ow: Ar2% H2 (L/min) 0.4 1.0Shielding gas ow: Ar (L/min) 13.0 14.0Backing gas ow: Ar (L/min) 7.0 7.0The different weld zones were examined by light microscopy. Marbles met-allographic reactive solution was used to develop the weld macrostructure, andthe dimensions of the welding zones (width and penetration of weld pools andHAZextension)weremeasuredfromthem. Electrolyticetchingwithoxalicacid (10%) was used to develop the weld microstructures. Transmission elec-tron microscopy (TEM) using a 200 kV Philips Tecnai 20 was used to detect thepresence of phases other than the majority ones.3. Results3.1. Inuence of the net input energy on operativeweldabilityThe inuence of the net input energy (Hnet), dened as theproportion of heat input per unit of length reaching the work-piece, onthepenetration, shapeandsizeof thewelds wasevaluated as indicated in Eq. (1). To that end the values of Hnetwere calculated considering the plasma welding conditions I,E and together with the values of energy transfer efciencies() for both groups of welds (melt-in and keyhole). The valuesof melt-in and keyhole used were typical of plasma arc welding asreported by other authors [16] given transfer efciency ranges of0.700.85 for melt-in mode and 0.850.95 for keyhole mode. Inboth cases, we chose an intermediate value within these ranges(melt-in =0.8 and keyhole =0.9).Hnet = IE(J/cm) (1)For purposes of comparison, Eq. (1) is valid assuming thatmost of the parameters that could inuence the energy lossesoccurring between the welding source and the work piece duringweldingarexed. Therefore, themodeofenergydeposition(melt-inorkeyhole)will bethemainfactordeterminingthedifference in transfer energy between the two groups of welds.Table 3 gives the Hnet values calculated for melt-in and keyholewelds, respectively, and also the dimensions of the fusion poolsobtained for PAW joints with complete penetration.Theratiooffusionpoolwidthtonetinputenergyisplot-ted in Fig. 1 for the two welding modes applied, showing theinuence of other variables, such as welding speed (for melt-626 A. Ure na et al. / Journal of Materials Processing Technology 182 (2007) 624631Table 3Net input energies and fusion pools dimensions of duplex stainless steel weldsmade with PAWI (A) (cm/min) Nozzleworkpiecedistance (mm)Hnet(J/cm)Fusion poolwidth (mm)Melt-in welds100 20 3 5350 6.7100 15 3 6450 8.3125 20 3 7400 8.2150 20 3 9000 9.5125 15 3 9850 9.4150 15 3 12600 11.0I (A) (cm/min) Nozzleworkpiecedistance (mm)Hnet(J/cm)L1L2Key-hole welds75 45 3 2500 3.0 1.275 45 4 2550 3.0 1.275 40 3 3750 4.2 1.275 40 4 2850 3.7 1.675 35 3 3100 3.9 1.375 35 4 3200 4.4 1.2L1 and L2 are dened in Fig. 3.in joints) or the distance between the shielding nozzle and theworkpiece (for keyhole joints). The increase of Hnetgenerallyproduced joints with wider fusion pools, and there was a closelinear dependence between the two parameters in both weldingmodes. Only in the case of keyhole welds with a short arc dis-tance (3 mm) was any anomalous behaviour observed, mainlyarising from the difculty of keeping the keyhole stable duringwelding.Macroscopic studies of the cross-sections of the melt-inwelded joints (Fig. 2) showed that the lowest net input energycondition (5350 J/cm) was insufcient to produce complete pen-etration, while Hnet values in the range of 9000 J/cm or higherproduced welds with excessive penetration and concavity. Opti-malenergyconditionsforoperativeweldabilityinthemet-inmodeareintherangeof 65007500 J/cm, althougheveninFig. 1. Ratio between fusion pool width and net input energy for PAW joints ina 3 mm duplex stainless steel sheet.these conditions fusion pools are wider than parent sheet thick-nessesandweldingfaultssuchasmisalignment aredifcultto avoid.Keyholewelds,whichwereallcarriedoutusingthesameorice gas nozzle with a diameter of 1.15 mm, presented lowerwidth/penetration ratios, with fusion pools ranging from 3.0 to4.4 mm in width at the top (L1) and from 1.2 to 1.6 mm at therootweld(L2). ThemacrographsinFig. 3revealthatopera-tive weldability is possible working with Hnet values lower than3000 J/cmandweldingspeedsdownto40 cm/min.However,some welding faults were detected, such as root concavity andsome undercuts, generallylocatedat one side of the fusionwelds.All welds were made under constriction conditions by clamp-ing the weld coupons on the backing plate, but there were nosigns of cracking in either melt-in or keyhole mode welds, orwithin the molten pool or the HAZs. For this reason, the stresslevelsinbothweldingzoneswerenot consideredsignicantwith the given work-piece thickness and welding conditions.Fig. 2. Cross-section of melt-in PAW welds in duplex stainless steel sheets.A. Ure na et al. / Journal of Materials Processing Technology 182 (2007) 624631 627Fig. 3. Cross-section of keyhole PAW welds in duplex stainless steel sheets.Fig. 4. Microstructure of the parent duplex stainless steel. (a) lamination plane (LTL); (b) transverse plane (TLTC). (c) TEM image of the parent sheet.628 A. Ure na et al. / Journal of Materials Processing Technology 182 (2007) 624631Table 4Average compositions of ferrite and austenite phase determined by quantitativeEDS microanalysisPhases Cr Ni Mo Mn Si FeFerrite () 23.80 4.59 3.90 1.58 0.85 BalanceAustenite () 22.08 7.19 2.98 2.02 0.76 Balance3.2. Inuence of net input energy on metallurgicalweldability3.2.1. Microstructural characteristics of parent materialInthesolutionannealedcondition,parentduplexstainlesssteel presentedthetypical biphasicmicrostructurecomposedof alternating bands of ferrite and austenite, showing partiallyrecrystallized grains elongated in the direction of roll (Fig. 4).Compositions of both phases were determined by quantitativeenergy dispersive X-ray microanalysis (EDS) and are shown inTable 4. Ni and Mn contents were higher in the austenite grains,and the proportion of Cr and Mo in the ferrite was greater. Nosigns of other majority phases, such as sigma, were detected ineitherXDRtestsortransmissionelectronmicroscopy(TEM)observations on the parent sheets. Fig. 4c shows a TEM imageof the parent alloy: ferrite and austenite grains are visible, theirgrain boundaries and interiors completely free of precipitates.The distributionof bothphases inthe sheets was alsoanalysedusing X-ray diffraction (XRD) and ferritoscope measurements.Fig. 5 shows the quantitative values from the XRD of the twometallographic sections shown in Fig. 4a and b. In the case of theLTL plane, diffractions were done at the original sheet surfaceand 0.1 and 2 mm deeper.With this method, the proportion of austenite phase isdeduced from the ratio between the integrated intensity (I) ofthe (1 1 1)reection (d111 =2.075A) and the integrated inten-sity of the same reection (Ip) obtained from a pure austeniticphase of an austenitic stainless steel (AISI 316). The proportionFig. 5. Ferrite/austenite ratios measured on the parent sheet by XRD.Fig. 6. Microstructure of the melt-in weld (input energy =7380 J/cm). (a) HAZand (b) fusion pool.of ferritic phase is calculated from the relationw+w = 1.The ferrite values calculated in this way ranged from 30% to45%; however, these were inuenced by the duplex texture, andhigher ferrite contents were determined at the lamination sheetsurface.Ferritoscope measurements were used to calculate the volu-metric percentage of-phase, which is less inuenced by thesheet texture. These measurements determined average ferritecontents of 45.9 5.5%. In this case, the volumetric percentageof phase was determined by the difference between 100% andthe measured ferrite content.3.2.2. Microstructure of the melt-in PAW weldsMelt-in plasma arc welded joints with mediumnet input ener-gies (65007500 J/cm) were characterized by narrow HAZs inthe range 400500 m, but they exhibited considerable ferritegrain growth (Fig. 6a). This grain growth zone (marked A) inu-enced subsequent epitaxial growth of the columnar ferrite grainsinside the fusion pool (marked B) (Fig. 6b). The precipitation ofsecondary austenite in the form of Widmanst atten needles fromthe ferrite grain boundaries (arrowed in Fig. 6b) was enhanced inthe fusion pool, and the proportion and needle width increasedwhen more energetic welding conditions were used.The proportion of volumetric ferrite content was determinedin the different zones of the welds for the various input ener-giesusedformelt-inPAW(Fig.7aandb).TheresultsshowA. Ure na et al. / Journal of Materials Processing Technology 182 (2007) 624631 629Fig. 7. Variation of ferrite content (a and b) and of microhardness (c and d) in different zones of melt-in PAW welds in the different welding conditions (weldingspeed and current) applied.an increase in ferrite content in the HAZ and particularly in thefusion pool, where proportions in excess of 60% were recordedfor themost energeticconditions. Theproportionof ferritedepends on the net input energy and the maintenance of weldingpower; increasing welding speed from15 to 20 cm/s reduced theproportion of ferrite inside the melting pool.It was observed that the increases in ferrite in the differentmelt-inweldsproducedahardeningeffectinsidethemeltingpools, and hardnesses ranged from 262 to 270 HV. These valuesincreasedinproportiontothe weldinginput energy. The oppositeeffect was observed in the HAZ, where the growth of the ferriticgrains produced a softening effect and average values of 250 HVwere recorded in the grain growth zone (Fig. 7c and d).3.2.3. Microstructure of the key-hole PAW weldsFig. 8a shows the microstructure inthe proximityof the fusionline of a keyhole plasma arc weld made with a welding current of75 Aat a speed of 45 cm/min. The working distance for this jointwas 4 mm. In these welding conditions there was very little ther-mal damage to the duplex stainless steel sheet; a narrow HAZwasformedwithwidthsoflessthan150250 mwherethegrain growth was very limited although there was some recrys-tallization of the banded texture. A detail of this zone at highermagnication (Fig. 8b) shows a microstructure formed by quasi-equiaxial ferrite grains, with allotriomorphic austenite at grainboundariesandalowproportionofintracrystallineausteniteneedles. The formation of Widsmanst atten austenite in the HAZwas suppressed. Moreover, the limited grain growth in the HAZreduced the sizes of the columnar ferrite grains formed by epi-taxial solidication from the base grain at the fusion line. Thesmall grain size of the fusion pool also limited the formation ofsecondary Widsmanst atten austenite in this zone, increasing theproportion of intracrystalline acicular austenite.The formation of intermetallic phases (e.g., sigma phase) wasnot detected in these low-energy conditions. However, when theinput energy was increased to 3100 J/cm during keyhole weld-ing, and especially when short arc distances were used (3 mm),the formation of partial melting zones was detected at base metalgrain boundaries along with precipitation of intermetallic phases(Fig. 9). This phenomenon could be a result of the higher inputenergyusedontheseweldscombinedwiththehighthermalgradient produced during application of a high density energywelding process like PAW in keyhole mode.The ferritometry measurements conrm the microstructuralstudies (Fig. 10a and b). Ferrite contents in the fusion pools ofkeyhole welds were generally lower than in melt-in welds, witha proportion of consistently around 50%. No great differenceswhere detected in the ferrite proportion for the different inputenergy conditions used, although reduction of the working dis-tance between plasma torch and piece produced a slight increaseof the ferrite content in the central zone of the welds.Vickers hardness measurements (Fig. 10c and d) conrmedthat levels of hardening inside the fusion pools were higher thanthose measured in melt-in welds, with maximum values in therange of 280300 HV, and were higher in less energetic weldingconditions. This is explainedbythe fact that the grainstructure of630 A. Ure na et al. / Journal of Materials Processing Technology 182 (2007) 624631Fig. 8. Microstructure of a keyhole weld (Hnet =2550 J/cm). (a) Fusion line, (b)detail at higher magnication.Fig.9. FormationofintergranularintermetalliccompoundsintheHAZofakeyhole PAW weld made with a Hnet of 3100 J/cm (I =75 A, =35 cm/min andwork distance =3 mm).these welds was ner andtemperature gradients were higher; thisincreased the cooling rates and also produced ner secondaryaustenitic aggregates. Moreover, the softening effects detectedin the HAZ of melt-in welds and associated with excessive graingrowth did not occur in this case; here, hardness in HAZs wasintermediate between the fusion pool and the parent material.Fig. 11 compares the maximum Vickers hardness measuredat the centre of the welds in relation to input energy, weldingmode and welding parameters. In the case of melt-in welds, thevariation of hardness was as expected because the increase ofinput-energy is generally associated with an increase of weldingFig. 10. Variation of ferrite content (a and b) and of microhardness (c and d) in different zones of keyhole PAW welds in the different welding conditions (weldingspeed and work distance) applied.A. Ure na et al. / Journal of Materials Processing Technology 182 (2007) 624631 631Fig. 11. Comparison of the maximumhardness attained in fusion pools as relat-ing to the net input energy and the welding mode.current or a decrease of welding speed, producing an almost lin-ear hardening effect in the fusion zones in correlation with theirhigher ferrite contents. In the case of keyhole welds, the reduc-tionof ferrite content didnot produce a reductionof hardness. Onthe contrary, the ner microstructure of the melting pools result-ing fromthe lowgrowth in the fusion line and the higher coolingrates determined by the steep temperature gradients increasedthe hardening levels, which were generally higher because ofthe low net input energy and high density energy used to weld.4. Conclusions(1) Good operative weldability by PAW in 3 mm thick sheet ofa 2205 duplex stainless steel is achieved by welding with anet input energy in the range 25003200 J/cm, if the keyholemode is used.(2) Welds producedbykeyhole PAWhave higher penetra-tion/width ratios than welds produced in the melt-in mode.(3) Welds produced in duplex stainless steels under high energyconditions inconductionmode are characterizedbyanincrease of ferrite contents inside the fusion pools to over45% more than in the parent material. Ferrite enrichment islimited to less than 20% when the keyhole mode is used.(4) HAZs of melt-in PAWjoints undergo a softening effect asso-ciated with excessive grain growth near the fusion line. Thiseffect was not observed in keyhole welding.(5) AlthoughbothmodesofPAWenhancehardeninginthefusion pool, this effect is more pronounced in keyhole weldsbecause they have a ner microstructure resulting from thelowerinputenergiesandthehighertemperaturegradientused.(6) If the net input energy is increased to 3000 J/cm in keyholewelds, there is a risk of melting fusion near the fusion lineand formation of brittle intermetallic phases.AcknowledgementsTheauthorswishtothanktheSpanishMinisteriodeEdu-caci on y Ciencia for the nancial support given to the presentresearch (MAT 2001-1123-C03-03). Also are grateful with Mr.Gilberto del Rosario from Centro de Apoyo Tecnol ogico (CAT-URJC) for his contribution to the quantitative X-ray diffractionstudy of parent sheets.References[1]K.M. Lee, H.S. Cho, D.C. Choi, Effect of isothermal treatment of SAF2205duplex stainless steel on migration of/interface boundary and growthof austenite, J. Alloys Compd. 285 (1999) 156161.[2]Z.L. Jiang, X.Y. Chen, X.Y. 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