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Journal of Materials Processing Technology 213 (2013) 229–237

Contents lists available at SciVerse ScienceDirect

Journal of Materials Processing Technology

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ighly efficient TIG welding of Cr13Ni5Mo martensitic stainless steel

.P. Lu ∗, M.P. Qin, W.C. Donghenyang National Laboratory for Materials Science, Institute of Metal Research, Chinese Academy of Sciences, Shenyang 110016, China

r t i c l e i n f o

rticle history:eceived 30 June 2012eceived in revised form8 September 2012ccepted 29 September 2012vailable online 8 October 2012

a b s t r a c t

A double-shielded TIG welding process using pure He gas as the inner shielding layer and He and CO2

mixed gas as the outer shielding layer was proposed for the welding of Cr13Ni5Mo martensitic stain-less steels. This proposed welding process can successfully address the problem of electrode oxidationwith mixed-gas TIG welding and the issue of low weld depth and welding efficiency of traditional TIGwelding. A change in the direction of the surface tension convection mode was the primary mecha-nism that affected the fusion zone profile. When the oxygen content in the weld pool was in the range of

eywords:artensitic stainless steel

IG weldingxygen contenturface tension convection

80–120 ppm, the surface tension convection direction changed from outward to inward, resulting in botha larger weld depth and a larger weld depth/width ratio. This process not only allows for a high weldingefficiency comparing with traditional TIG welding but also produces better weld impact properties thanthose of MAG welding (metal active gas welding).

© 2012 Elsevier B.V. All rights reserved.

usion zone profile

. Introduction

Traditional TIG welding processes yield good weld qualitynd are one of the most frequently used welding processesor Cr13Ni5Mo martensitic stainless steels. However, the weldepth created by traditional TIG welding via a single pass is nor-ally restricted to 3 mm (Lucas and Howse, 1996), and the weld

epth/width (D/W) ratio is restricted to approximately 0.2–0.3or stainless steel under argon shielding (Lu et al., 2003), whichesults in low welding efficiencies. To overcome these disadvan-ages, Gurevich et al. (1965) of the E.U. Paton Institute of Electric

elding first proposed the A-TIG welding process in the 1960s, arocess in which a fine layer of activating flux (halides or oxides)

s smeared on the surface of a workpiece prior to welding. Lecontet al. (2006) have found that the weld depth of A-TIG welding viasingle pass can be increased by a factor of 1.5–4 compared with

hat of traditional TIG welding. However, the fusion zone profilef A-TIG welding is sensitive to the quantity of the activating fluxmeared on the plate surface, and slag is easily retained on theeld surface. Another improvement to traditional TIG welding isixed-gas TIG welding, a process that utilizes the addition of an

ctive gas to the shielding gas and has been explored with H2 byisieh et al. (1999), with SO2 by Heiple and Burgardt (1985), with

2 by Lu et al. (2007), with CO2 by Lu et al. (2008) and with Ar-

Ar + CO2) double shielding by Lu et al. (2010). The addition of H2n He-Ar shielding can increase the weld depth, but this addition

∗ Corresponding author. Tel.: +86 24 23971429.E-mail address: shplu@imr.ac.cn (S.P. Lu).

924-0136/$ – see front matter © 2012 Elsevier B.V. All rights reserved.ttp://dx.doi.org/10.1016/j.jmatprotec.2012.09.025

also results in variable penetration arising from cast-to-cast mate-rial differences; additionally, this process is sensitive to the parentmaterial composition (Lucas and Rodwell, 1987). The active gas SO2is toxic, and the limitations of O2 or CO2 as active gas additives aredue to electrode oxidation, which shortens the application life ofthe electrode.

Limmaneevichitr and Kou (2000a,b) investigated the NaNO3simulated weld pool (also known as the Marangoni Convec-tion) variation by defocused laser heating with and without theC2H5COOH surface active agent and showed the visualization resultof the weld pool convection and heat transfer. The results showedthat the pool surface convection and heat transfer mode are relatedto the pool surface tension variation, which is dependent on thepool temperature. Heiple and Roper (1982) proposed that oxygenacts as a liquid ferrous-based alloy surface-active element. Whenthe oxygen content in a weld pool is low, the temperature coef-ficient of the surface tension, ∂�/∂T, is negative. A temperaturegradient exists on the weld pool surface, with high temperatures atthe pool center and relatively low temperatures at the pool edges.Because the temperature coefficient of the surface tension is nega-tive, the higher the temperature, the lower the surface tension. Inthis situation, the surface tension at the pool center is lower thanthat at the pool edges, and an outward convection from the centerto the edge of the weld pool thus occurs. This convection leads toa heat flux from the arc that is easily transferred to the edge of theweld pool and results in a low weld depth and low weld D/W ratio,

as shown in Fig. 1(a). When the oxygen content of the weld poolis high, the temperature coefficient of the surface tension, ∂�/∂T,is positive, and the surface tension at the pool center is larger thanthat at the pool edges and thus induces an inward convection from

230 S.P. Lu et al. / Journal of Materials Processing Technology 213 (2013) 229–237

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Table 1Chemical composition of Cr13Ni5Mo martensitic stainless steel (wt.%).

C Cr Ni Mo Si Mn P S O Fe

Fig. 1. Convection in the weld pool driven by surface tension.

he edge to the center of the weld pool, leading to a heat flux thats easily transferred to the bottom of the weld pool. This weld poolorms a high weld depth and high weld D/W ratio, as shown inig. 1(b) (Lu et al., 2010).

A double-shielded TIG welding process was proposed with He ashe inner layer shielding gas and He + O2 as the outer layer shieldingas, and the weld efficiency from this process is increased com-ared with that of the common TIG welding process (Li et al., 2010).heaper than oxygen, CO2 gas is widely used in industry and canlso decompose into oxygen in the welding process. In this paper,O2 was used instead of O2, and He-(He + CO2) double-shieldingIG welding of Cr13Ni5Mo material was investigated under dif-erent welding parameters, as shown in Fig. 2. The pure inert Heas in the inner layer avoids direct contact of the electrode withhe CO2 active gas in the outer layer. The addition of CO2 to theuter layer provides oxygen as a surface-active element, which dis-olves in the weld pool and increases the weld depth and weldingfficiency. The variation in the fusion zone profile under differentonditions and the mechanisms affecting the fusion zone profileere studied. Additionally, the welded joint mechanical propertiesere investigated.

. Experimental procedures

Bead-on-plate welding was performed on Cr13Ni5Mo marten-itic stainless steel, the chemical composition of which is shown in

able 1. A W-2%ThO2 tungsten electrode (2.4 mm in diameter, 60◦

ertex angle) and a direct current power source with the electrodes the cathode were used in these experiments.

Fig. 2. Schematic diagram of the double-shielded TIG welding process.

0.011 12.52 4.45 0.64 0.25 0.70 0.031 0.0023 0.0042 Bal.

The welding plates were machined to a size of100 mm × 50 mm × 12 mm. Four different TIG welding condi-tions were chosen to study the feasibility of this double-shieldedTIG welding process. The welding parameters consisted of a 160 Awelding current, a 2 mm/s welding speed, and a 3 mm electrodegap. With the double-shielded TIG welding, additional experi-ments were conducted to investigate the effect of the weldingparameters on the fusion zone profile. In these experiments, theinner layer gas flow rate was 10 L/min (the inner diameter of thechannel for inner shielding gas is 5.0–6.0 mm), the outer layer gasflow rate was varied from 0 L/min to 20 L/min, the welding speedwas in the range of 1.5–5 mm/s, the welding current was in therange of 120–240 A, and the electrode gap was in the range of1–7 mm.

According to the experimental results, the double-shielded TIGwelding process (with the set of parameters that ensures a deepdepth) was adopted to weld 9 mm thick plates in a single pass.The welded joint was tempered at 590 ◦C (with a heating rate of2 ◦C/min and a 4 h holding time followed by furnace cooling). Thetensile and impact properties of the welded joint were analyzed forboth the as-welded joints and the tempered joints.

After welding, both the electrode and the weld surface mor-phologies were imaged. Cross-sections of the welds were etchedwith an HCl + CuSO4 solution prior to imaging. The weld D/W ratiowas calculated by measuring the weld depth and width. The oxygencontent in the weld was measured by an oxygen/nitrogen analyzer(TC-436). The composition of the weld metal was also analyzed.A scanning electron microscope (SEM) was used to observe thefracture of the Charpy V-notch impact specimens.

3. Results

3.1. Feasibility of the double-shielded TIG welding process

Four different TIG welding conditions were chosen to study thefeasibility of the double-shielded TIG welding process, as shownin Table 2. The morphologies of the electrode and the weld sur-faces are shown in Figs. 3 and 4, respectively, for the traditional TIGwelding, mixed-gas TIG welding, double-shielded TIG welding (2%CO2) and double-shielded TIG welding (20% CO2) processes. Theelectrode and weld surface oxidation are associated with the arcoxidizability (when the CO2 in the shielding gas is decomposed orionized into oxygen atoms or ions by the arc in the welding process,the arc can oxidize the electrode and weld pool; in this situation,

the arc is said to have oxidizability). Even if the CO2 concentration isquite low (∼2%) under mixed-gas TIG welding, the arc oxidizabilityis strong, and the oxidation of the electrode and weld surface can be

Table 2Four different welding processes.

Welding process Shielding gas (vol.%) and gas flowrate

Traditional TIG welding He, 10 L/minMixed-gas TIG welding He + 2% CO2, 10 L/min

Double-shielded TIG welding, 2% CO2 Inner layer He, 10 L/minOuter layer He + 2% CO2, 10 L/min

Double-shielded TIG welding. 20% CO2 Inner layer He, 10 L/minOuter layer He + 20% CO2, 10 L/min

S.P. Lu et al. / Journal of Materials Processing Technology 213 (2013) 229–237 231

Fig. 3. Weld surface morphology: (a) traditional TIG welding, (b) mixed-gas shielded TIG welding; (c) double-shielded TIG welding, 2% CO2; (d) double-shielded TIG welding,20% CO2.

Fig. 4. Electrode morphology: (a) traditional TIG welding; (b) mixed-gas shielded TIG welding; (c) double-shielded TIG welding, 2% CO2; (d) double-shielded TIG welding,20% CO2.

F weldiC

siactw

iTTwidTdw

TW

ig. 5. Fusion zone profile: (a) traditional TIG welding; (b) mixed-gas-shielded TIGO2.

ignificant. However, even with CO2 concentrations as high as 20%n double-shielded TIG welding, the arc oxidizability is not strong,nd the oxidation of the electrode and weld surfaces are not signifi-ant. Therefore, double-shielded TIG welding can successfully solvehe problem of electrode oxidation that occurs with mixed-gas TIGelding.

The shape and dimension of the welds are shownn Fig. 5 and Table 3, respectively, for the traditionalIG welding, mixed-gas TIG welding, double-shieldedIG welding (2% CO2) and double-shielded TIGelding (20% CO2) processes. Both the mixed-gas TIG weld-

ng and the double-shielded TIG welding can increase the weldepth and reduce the weld width relative to that of traditional

IG welding. However, when the CO2 concentration is 20% underouble-shielded TIG welding, the weld depth is the deepest, theeld width is the narrowest, and the weld D/W ratio is 0.63.

able 3eld dimensions with four different welding processes.

Welding process Weld depth Weld width Weld D/Wratio

Traditional TIG welding 2.6 mm 13.2 mm 0.20Mixed-gas TIG welding 4.00 mm 9.50 mm 0.42Double-shielded TIG welding, 2% CO2 3.89 mm 11.63 mm 0.33Double-shielded TIG welding, 20% CO2 5.89 mm 9.37 mm 0.63

ng; (c) double-shielded TIG welding, 2% CO2; (d) double-shielded TIG welding, 20%

The addition to the shielding gas of CO2 as the active gashas important effects on the arc. The arc voltages are 24.0 V,29.2 V, 23.8 V and 24.1 V for the traditional TIG welding, mixed-gasTIG welding, double-shielded TIG welding (2% CO2) and double-shielded TIG welding (20% CO2) processes, respectively.

3.2. Effect of the welding parameters on the fusion zone profile

3.2.1. Effect of the gas flow rate on the fusion zone profileThe weld shapes with outer layer gas flow rates from 2 L/min

to 20 L/min are shown in Fig. 6. As the outer layer gas flow rateincreases, the fusion zone profile changes from shallow and wideto deep and narrow. The weld oxygen content and weld D/W ratiofor the different outer layer gas flow rates are shown in Fig. 7. Theweld oxygen content and weld D/W ratio initially increase withincreasing outer layer gas flow rate. However, when the outer layergas flow rate is greater than 10 L/min, both the weld oxygen con-tent and the weld D/W ratio tend to decrease. The magnitude ofthe inner/outer layer gas flow rate ratios represents the effectivestrength of the outer layer shielding gas. If the inner/outer layergas flow rate ratio is less than 1, the CO2 in the peripheral region ofthe arc can be easily carried to the center region of the arc, resulting

in an increase in the weld oxygen content and weld D/W ratio. Ifthe inner/outer layer gas flow rate ratio is greater than 1, the CO2in the peripheral region of the arc is not easily carried to the cen-ter region of the arc, resulting in a decrease in the weld oxygen

232 S.P. Lu et al. / Journal of Materials Processing Technology 213 (2013) 229–237

Fig. 6. Effect of the outer layer gas flow rate on the fusion zone profile: (a) 2 L/min, (b) 5 L/min, (c) 10 L/min, (d) 20 L/min.

Fc

ct

3

tbfioailptoa

3

twwflrodrof

ig. 7. Effect of the outer layer gas flow rate on the weld D/W ratio and oxygenontent.

ontent and weld D/W ratio. When the oxygen content is greaterhan 120 ppm, the weld D/W ratio is greater than 0.5.

.2.2. Effect of welding speed on the fusion zone profileThe fusion zone profiles at different welding speeds from 5 mm/s

o 1.5 mm/s are shown in Fig. 8. As the welding speed decreases,oth the weld depth and width increase, and the fusion zone pro-le changes from shallow and wide to deep and narrow. The weldxygen content and weld D/W ratio at different welding speedsre shown in Fig. 9. The weld oxygen content and weld D/W rationcrease with decreasing welding speed. When the welding speed isess than 1.5 mm/s, the 10 mm thick plate can be welded in a singleass. However, if the welding speed is too low, the arc has enoughime to dissolve oxygen in the weld pool, and an oxide tends to formn the weld surface and reduces the surface tension convection tocertain extent.

.2.3. Effect of welding current on the fusion zone profileThe fusion zone profiles at different welding currents from 120 A

o 240 A are shown in Fig. 10. As the welding current increases, theeld depth first increases and subsequently decreases, the weldidth consistently increases, and the fusion zone profile changes

rom shallow and wide to deep and narrow, then changes to shal-ow and wide once again. The weld oxygen content and weld D/Watio at different welding currents are shown in Fig. 11. The weldxygen content and weld D/W ratio first increase and subsequently

ecrease with increasing welding current. When the welding cur-ent is greater than 200 A, the arc oxidizability is strong, and anxide is easily formed on the weld surface and weakens the sur-ace tension convection. In addition, an increase in the welding

Fig. 8. Effect of the welding speed on the fusion zone profi

Fig. 9. Effect of the welding speed on the weld D/W ratio and the oxygen content.

current results in an increase in the plasma drag force. Tanaka et al.(2004) have shown that the plasma force induces outward con-vection. Therefore, the weld D/W ratio is low for a 240 A weldingcurrent.

3.2.4. Effect of the electrode gap on the fusion zone profileThe fusion zone profiles for different electrode gaps from 1 mm

to 7 mm are shown in Fig. 12. As the electrode gap increases, theweld depth first increases and subsequently decreases, the weldwidth first decreases and subsequently increases, and the fusionzone profile changes from shallow and wide to deep and nar-row before returning to shallow and wide once again. Data forthe weld oxygen content and weld D/W ratio produced by differ-ent electrode gaps are shown in Fig. 13. The weld oxygen contentconsistently increases, but the weld D/W ratio increases and subse-quently decreases with decreasing electrode gap. As the electrodegap is increased, the protection of the inner shielding gas providedto the welding pool is weakened, and the carbon dioxide in theouter layer shielding easily enters the arc, which enhances the arcoxidizability. Hence, the weld metal oxygen content increases withthe increasing electrode gap, as shown in Fig. 13. In this situation,the weld pool is easily oxidized and forms an oxide on the pool sur-face, which weakens the surface-tension-induced convection anddecreases the weld D/W ratio, as shown in Fig. 13.

3.3. Mechanical properties of the welded joint

A square butt weld joint without a gap (closed joint) was pre-pared with two Cr13Ni5Mo plates of size 100 mm × 50 mm × 9 mm.Prior to welding, the two plates were spot-welded at the mid-dle and on two edges of the weld seam. The double-shielded TIG

le: (a) 5 mm/s, (b) 3 mm/s, (c) 2 mm/s, (d) 1.5 mm/s.

S.P. Lu et al. / Journal of Materials Processing Technology 213 (2013) 229–237 233

Fig. 10. Effect of the welding current on the fusion zone profile: (a) 120 A, (b) 160 A, (c) 200 A, (d) 240 A.

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tsTtt

Fig. 13. Effect of the electrode gap on the weld D/W ratio and the oxygen content.

Fig. 14. Macrograph of the cross-section of a welded joint.

ig. 11. Effect of the welding current on the weld D/W ratio and the oxygen content.

elding process was used to penetrate the welding plates in a sin-le pass. The welding parameters were set to 10% CO2, a 10 L/minas flow rate for both the outer and inner layer shielding gases, a60 A welding current, a 1.5 mm/s welding speed, and a 3 mm elec-rode gap. The cross-sectional shape of the welded joint is shownn Fig. 14, and the microstructure of the welded metal is shownn Fig. 15. Because the cooling speed of the weld pool during the

elding process is fast compared with the cooling rate of the largeasting process, the � ferrite transformation into austenite is notomplete, and the microstructure of the weld metal is thus com-osed of martensite and small amounts of retained � ferrite.

MAG welding is often used for Cr13Ni5Mo heavy castingsn industry. In order to evaluate the welding joint property byouble-shielded TIG welding, MAG welding was also used to weldr13Ni5Mo martensitic stainless steel. A 50–60 degree double-Vroove butt weld joint was prepared with two Cr13Ni5Mo plates ofize 300 mm × 200 mm × 50 mm. The wire electrode diameter was.2 mm, and the chemical composition of the welding wire used inhe MAG welding is shown in Table 4. The MAG welding param-ters were set to a Ar + 5% CO2 shielding gas with a gas flow ratef 12–15 L/min, a welding current of 220–280 A, an arc voltage of4–32 V, a welding speed of 30–45 cm/min and a wire feed speedf 5–7 m/min. A macrosection of the MAG weld joint is shown inig. 16.

The tensile and impact properties of the welded joints in bothhe as-welded state and in the 590 ◦C tempered state for the double-

hielded TIG welding and MAG welding processes are shown inable 5. The tensile strengths of the weld joints decreased afterempering at 590 ◦C. A larger difference was not observed in theensile strengths of the two welding methods. The dilatometric

Fig. 15. Microstructure of the weld metal.

Fig. 12. Effect of the electrode gap on the fusion zone profile: (a) 1 mm, (b) 3 mm, (c) 5 mm, (d) 7 mm.

234 S.P. Lu et al. / Journal of Materials Processing Technology 213 (2013) 229–237

Table 4Chemical composition of the welding wire (wt.%).

Element C Cr Ni Mo Si Mn P S Fe

Requirement ≤0.03 11.5–13.5 4.0–6.0 0.4–0.8 ≤0.90 ≤1.0 ≤0.02 ≤0.01 Bal.Measurement 0.024 12.86 5.17 0.50 0.42 0.60 0.012 0.004 Bal.

Table 5Tensile and impact properties of the welded joint at 20 ◦C.

Welding process Condition Tensile property, �b (MPa) Impact property, Akv (J)

5 mm × 10 mm × 55 mm

Individual value Even value

Double TIG As-welded 910 42 43 43 43Tempered at 590 ◦C 872 46 52 52 50

MAG As-welded 945 13 14 14 14Tempered at 590 ◦C 890 20 20 22 21

Table 6Chemical composition of the weld metal (wt.%).

Welding process C Cr Ni Mo Si Mn P S O Fe

Double TIG 0.011 12.20 4.60 0.66 0.29 0.69 0.029 0.003 0.0148 Bal.MAG 0.034 13.58 5.08 0.55 0.49 0.78 0.025 0.004 0.053 Bal.

Fig. 16. Macrosection of the MAG weld joint.

Fig. 18. Weld defects: (a

Fig. 17. Dilatometric curve of the completely austenitized sample of Cr13Ni5Momartensitic stainless steel (heating rate 0.05 ◦C/s, cooling rate 100 ◦C/s).

curve of a completely austenitized sample of Cr13Ni5Mo marten-sitic stainless steel is shown in Fig. 17, where Ms is 578 ◦C. Wanget al. (2008) proposed that when the tempering temperature is inthe range of 578–600 ◦C, the formation of stable reversed austenite

will improve the impact properties. Therefore, tempering at 590 ◦Ccan increase the weld impact energy for both of these welding pro-cesses due to the formation of stable reversed austenite. However,the impact energy of the double-shielded TIG weld is much larger

) slag, (b) porosity.

S.P. Lu et al. / Journal of Materials Processing Technology 213 (2013) 229–237 235

F d TIG w5

teetdmcitClboofpAiBttfi

ig. 19. Scanning electron micrographs of fractures: (a) As-welded, double-shielde90 ◦C, MAG welding.

han that of the MAG weld. The first reason for the low impactnergy of the MAG weld is that a greater amount of slag and porosityxists in the weld metal, as shown in Fig. 18. The second reason ishat the carbon content of the MAG weld is higher than that of theouble-shielded TIG weld. The chemical compositions of the weldetal for the double-shielded TIG welding and MAG welding pro-

esses are shown in Table 6. The oxygen content of the MAG welds much higher than that of the double-shielded TIG weld. Becausehe shielding gas of the MAG welding process consists of Ar + 5%O2, the dissociation of CO2 is strong enough to produce not only a

arge number of oxygen atoms but also a quantity of carbon atoms,oth of which dissolve in the weld pool and increase the carbon andxygen content of the weld metal. Scanning electron micrographsf the weld impact fractures are shown in Fig. 19. It is clear that theractures from these two welding processes display typical dim-les, at the bottom of which exist fine oxide inclusions (1–2 �m).n EDS analysis of the oxide inclusions is shown in Fig. 20. Oxide

nclusions are often the sources of cracks in the impact process.

ecause the size of the oxide inclusions is nearly the same underhese two welding processes, greater oxygen content creates addi-ional sources for cracking. Therefore, the growth and joining ofne holes in the double-shielded TIG weld is more complete than

Fig. 20. EDS analysis of inclusions: (a) double-

elding; (b) 590 ◦C, double-shielded TIG welding; (c) as-welded, MAG welding; (d)

that of the MAG weld. Consequently, the dimples in the double-shielded TIG weld are larger than those of the MAG weld. MAGwelding process is quite different from the TIG welding process. ForMAG welding process, there is consumable transfer in the weldingprocess and the weld pool convection is more complex. Since cer-tain content CO2 is added to the argon shielding gas directly, theweld pool is easily oxidized, and more oxide (slag) forms on theweld pool. In this situation, the liquid pool contacts with the slagdirectly, and pool/slag interface formed instead of pool/arc surface.Therefore, the surface induced Marangoni convection will not existin MAG welding when the oxygen content in the welding pool ishigh.

4. Discussion

4.1. Role of CO2 in the shielding gas

Under mixed-gas TIG welding, the dissociation of CO2 is strongand arc constriction occurs, as first proposed by Simonik et al.(1973). In the cooler peripheral region of the arc, the dissociation ofCO2 exists in the form of molecules and dissociated atoms, which

shielded TIG welding, (b) MAG welding.

236 S.P. Lu et al. / Journal of Materials Processing Technology 213 (2013) 229–237

funct

antoeeHiCiai

ikeectt

�

wimmf(htsso1FtdiiItft∂Ivwpc

Fig. 21. Surface tension of Fe-O system as a

re large enough for the electrons to become attached and formegatively charged particles. Consequently, the number of elec-rons that are the main charge carriers in the peripheral regionf the arc is reduced. The dissociation and ionization of CO2 arendothermic reactions. Accordingly, the arc constricts to a newquilibrium state with a higher arc voltage and current density.owever, the arc is mainly controlled by the amount of He, the

nner layer shielding gas in double-shielded TIG welding. The activeO2 gas in the outer layer serves only to increase the arc oxidizabil-

ty to a certain extent but has little effect on the constriction of therc. Therefore, both the arc voltage and the current density do notncrease remarkably with double-shielded TIG welding.

The addition of the active CO2 gas to the shielding gas also hasmportant effects on the heat transfer of the weld pool. It is wellnown that oxygen is a liquid ferrous-based alloy surface-activelement. The oxygen content in the weld pool has an importantffect on the fusion zone profile by changing the surface tensiononvection mode. Sahoo et al. (1988) proposed the following equa-ion for the surface tension, �, of a Fe–O system as a function ofemperature, T, and oxygen content, ai:

= �o,m − A(T − Tm) − RT�s ln[

1 + K1ai exp(−Q

RT

)](1)

here �o,m is the surface tension of the pure metal at its melt-ng point (1.943 N m−1), A is the negative of d�/dT for the pure

etal (4.3 × 10−4 N m−1 K), Tm is the melting point for the pureetal (1811 K), R is the gas constant (8.314 J/mol K), � s is the sur-

ace excess at saturation (2.03 × 10−8 kg mol/m2), k1 is a constant1.38 × 10−2), ai is the activity in wt.% [O], and Q is the standardeat of adsorption (−1.463 × 105 kJ/kg mol). The predicted surfaceension of the Fe–O system as a function of the oxygen content ishown in Fig. 21(a). The surface tension of the Fe–O system is sen-itive to the oxygen content and decreases quickly with increasingxygen content, especially when the oxygen content is less than000 ppm. Fig. 21(b) shows the predicted surface tension of thee–O system as a function of temperature. When the oxygen con-ent is less than 20 ppm, the surface tension of the Fe–O systemecreases with increasing temperature. When the oxygen content

s greater than 40 ppm, the surface tension of the Fe–O system firstncreases and subsequently decreases with increasing temperature.n other words, when the oxygen content is greater than 40 ppm,he temperature coefficient of the surface tension, ∂�/∂T, changesrom positive to negative at a particular temperature. At higheremperatures, the temperature coefficient of the surface tension,�/∂T, is negative, and outward surface tension convection occurs.

n the relatively low temperature area, inward surface tension con-

ection occurs. A temperature gradient exists on the pool surface,ith high temperatures at the pool center and relatively lower tem-erature at the pool edges. Consequently, inward surface tensiononvection easily occurs in the peripheral region of the weld pool,

ion of: (a) oxygen content, (b) temperature.

while outward surface convection occurs easily in the center regionof the weld pool. When the area of the inward surface tension con-vection in the peripheral region of the weld pool is much larger thanthat of the outward surface tension convection in the center regionof the weld pool, the surface tension induces an inward convectionin the weld pool, and a heat flux is easily transferred to the bottomof the weld pool, forming a high weld depth and weld D/W ratio,as shown in Fig. 1(b).

The weld oxygen contents are 56 ppm, 400 ppm, 76 ppm and168 ppm for the traditional TIG welding, mixed-gas TIG welding,double-shielded TIG welding (2% CO2) and double-shielded TIGwelding (20% CO2) processes, respectively. Because the oxygencontent of the traditional TIG welding is low, the surface tensioninduces an outward convection, forming a shallow and wide weld.With mixed-gas TIG welding, because the oxygen content is notablyhigh, large amounts of oxide are easily formed on the weld sur-face, which reduces the effect of the surface tension convectionto a large extent. However, because arc constriction occurs withmixed-gas TIG welding, the weld depth and weld D/W ratio of thisprocess are both larger than those of the traditional TIG welding.When the weld oxygen content increases from 76 ppm to 168 ppmwith double-shielded TIG welding, the surface tension convectionchanges from outward to inward, and the welds become character-ized by a deep and narrow fusion zone profile, as shown in Fig. 5d.

4.2. Critical oxygen content affecting the weld pool convection

According to the results presented above regarding the effectof the welding parameters on the fusion zone profile, the oxygencontent in the weld pool can be effectively adjusted using double-shielded TIG welding by changing the welding parameters, and thefusion zone profile can be controlled by changing the surface ten-sion convection direction from outward to inward. From Fig. 21(b),it can been observed that when the oxygen content in the weldpool is greater than 20 ppm, the inward surface tension convectionin the peripheral region of the weld pool coexists with the out-ward surface tension convection in the center region of the weldpool. As the oxygen content in the weld pool increases, the criti-cal temperature point where the temperature coefficient of surfacetension changes from position to negative is increased, and the tem-perature range, where inward surface tension convection occurs,expanded. Therefore, the inward surface tension convection in theperipheral region of the weld pool is expanded, and the outwardsurface tension convection in the center region of the weld pool is

narrowed. The statistical relationship between the weld D/W ratioand the oxygen content, as determined by the results from Fig. 7,Fig. 9, Fig. 11 and Fig. 13, is shown in Fig. 22 for double-shieldedTIG welding. This diagram can be divided into four regions.

S.P. Lu et al. / Journal of Materials Processi

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ig. 22. Statistical relationship between the weld D/W and the oxygen content.

Region A: The oxygen content is less than 80 ppm and the weldD/W ratio is approximately 0.2. In this case, the surface tensionconvection is dominated by outward convection, and the fusionzone profile is shallow and wide.Region B: The oxygen content is in the range of 80–120 ppm,and the weld D/W ratio is in the range of 0.2–0.5. In this case,the inward surface tension convection competes with the out-ward surface tension convection. When the strength of the inwardsurface tension convection is much larger than that of the out-ward surface tension convection, the fusion zone profile changesfrom shallow and wide to deep and narrow. Therefore, the criticaloxygen content of the surface tension convection changes fromoutward to inward in the range of 80–120 ppm.Region C: The oxygen content is in the range of 120–300 ppm, andthe weld D/W ratio is greater than 0.5. In this case, the surfacetension convection is dominated by the inward convection, andthe fusion zone profile is shallow and wide.Region D: On the left side of this region, the oxygen content is highenough to induce inward surface tension convection, based theo-retically on Fig. 21(b). However, because the amount of oxygen ishigh, SiO2 and Cr2O3 oxides (slag) can be formed on the periph-eral areas of the pool surface for stainless steel (Lu et al., 2004),and the weld pool/arc surface changes to a pool/slag interface onthe periphery area of the pool. The surface-induced inward con-vection on the pool periphery is largely reduced or disappears, andthis results in a lower weld D/W ratio.

Liquid pool movement and its fusion zone profile variation areomplex phenomena influenced by different physical forces in theelding process. The explanation and discussion in this investiga-

ion are only based on the experimental results qualitatively, andurther more detailed quantitative investigation on the subject bothy simulation and experiment are necessary to clarify and under-tand the complex process of the liquid weld pool movement in theelding process in the future.

. Conclusions

The He-(He + 10%CO2) double-shielded TIG welding process canuccessfully address both the problem of electrode oxidation withe(Ar) + O2(CO2) mixed-gas TIG welding and the issue of low weldepth and welding productivity observed with the use of tradi-ional TIG welding for stainless steel. The change in the surfaceension convection mode is the main mechanism that affects the

usion zone profile. The critical oxygen content at which the sur-ace tension convection changes from outward to inward lies inhe range of 80–120 ppm for Cr13Ni5Mo martensitic stainlessteel.

ng Technology 213 (2013) 229–237 237

With the double-shielded TIG welding process, adjustment ofthe welding parameters can effectively control the weld oxygencontent and provide further control of the surface tension con-vection and fusion zone profile. With suitable welding parameters,consisting of a 1.5–2.0 mm/s welding speed, a 120–200 A weldingcurrent and a 3 mm electrode gap, a weld depth of 6–12 mm and aweld D/W ratio of 0.6–1.0 can be achieved using He-(He + 10%CO2)double shielding in a single pass with a 10 L/min flow rate for boththe outer and inner layer shielding gas. The double-shielded TIGwelding process displays an efficiency 2–4 times greater than thatof traditional TIG welding.

The weld joint strength of the Cr13Ni5Mo stainless steel fromHe-(He + 10%CO2) double shielding is similar to the strength of aMAG-welded joint under Ar + 5%CO2 shielding, and the weld impactproperty is superior to that of MAG welding as well.

Acknowledgments

The authors are grateful for financial support from the NationalScience Foundation of China under grant No. 50874101 and No.51104142.

References

Gurevich, S.M., Zamokov, V.N., Kushirenko, N.A., 1965. Improving the penetrationof Titanium alloys when they are welded by Tungsten arc process. AutomaticWelding USSR 18, 1–4.

Heiple, C.R., Burgardt, P., 1985. Effect of SO2 shielding gas additions on GTA weldshape. Welding Journal 64, 159–162.

Heiple, C.R., Roper, J., 1982. Mechanism for minor element effect on TIG fusion zonegeometry. Welding Journal 61, 92–102.

Hisieh, R.L., Pan, Y.T., Liou, H.Y., 1999. The study of minor elements and shielding gason penetration in TIG welding of type 304 stainless steel. Journal of MaterialsEngineering and Performance 8, 68–74.

Leconte, S., Paillard, P., Saindrenan, J., 2006. Effect of fluxes containing oxides ontungsten inert gas welding process. Science and Technology of Welding andJoining 11, 43–47.

Li, D.J., Lu, S.P., Li, D.Z., Li, Y.Y., 2010. Weld pool shape variations and electrode pro-tection in double shielded TIG welding. Advanced Materials Research 97–101,3978–3981.

Limmaneevichitr, C., Kou, S., 2000a. Visualization of Marangoni convection in sim-ulated weld pools. Welding Journal 79, 126–135.

Limmaneevichitr, C., Kou, S., 2000b. Visualization of Marangoni convection insimulated weld pools containing a surface-active agent. Welding Journal 79,324–330.

Lu, S.P., Fujii, H., Sugiyama, H., Tanaka, M., Nogi, K., 2003. Effect of oxygenadditions to argon shielding gas on GTA weld shape. ISIJ International 43,1590–1595.

Lu, S.P., Fujii, H., Nogi, K., 2010. Weld shape variation and electrode oxidationbehavior under Ar-(Ar-CO2) double shielded GTA welding. Journal of MaterialsScienece 26, 170–176.

Lu, S.P., Fujii, H., Nogi, K., Sato, T., 2007. Effect of oxygen content in He-O2 shieldinggas on weld shape in ultra deep penetration TIG. Science and Technology ofWelding and Joining 12, 689–695.

Lu, S.P., Fujii, H., Nogi, K., 2008. Marangoni convection and weld shape variations inHe-CO2 shielded gas tungsten arc welding on SUS304 stainless steel. Journal ofMaterials Science 43, 4583–4591.

Lu, S.P., Fujii, H., Nogi, K., 2004. Marangoni convection and weld shape variationsin Ar-O2 and Ar-CO2 shielded GTA welding. Materials Science and EngineeringA380, 290–297.

Lucas, W., Howse, D., 1996. Activating flux-increasing the performance and pro-ductivity of TIG and plasma process. Welding and Metal Fabrication (Journal),11–17.

Lucas, W., Rodwell, M.H., 1987. Process techniques to improve control of penetrationin TIG welding. Welding Review 8, 185–196.

Simonik, A.G., Petviashvili, V.I., Ivanov, A.A., 1973. The effect of contraction of thearc discharge upon the introduction of the electro-negative elements. WeldingProduction 23, 68–71.

Sahoo, P., Debboy, T., Mcnallan, M.J., 1988. Surface tension of binary metal-surfaceactive solute system under conditions relevant to welding metallurgy. Metal-lurgical Transactions B 19, 381–389.

Tanaka, M., Ushio, M., Lowke, J.J., 2004. Numerical study of gas tungsten arc plasma

Wang, P., Lu, S.P., Li, D.Z., Kang, X.H., Li, Y.Y., 2008. Investigation on phase transfor-mation of low carbon martensitic stainless steel ZG06Cr13Ni4Mo in temperingprocess with low heating rate. Acta Metallurgica Sinica (Engl. Lett.) 44,681–685.