j o u r n a l o f m a t e r i a l s p r o c e s s i n g t e c h n o l o g y 2 0 9 ( 2 0 0 9 ) 1231–1239
journal homepage: www.elsevier.com/locate/jmatprotec
Arc ignitability, bead protection and weld
shape variations for He–Ar–O2 shielded
GTA welding on SUS304 stainless steel
Shanping Lu a,b,∗ , Hidetoshi Fujii a , Kiyoshi Nogi a
a
Joining and Welding Research Institute, Osaka University, Osaka 567-0047, Japan
Shenyang National Laboratory for Materials Science, Institute of Metal Research,
Chinese Academy of Science, Shenyang 110016, PR China
b
a r t i c l e
i n f o
a b s t r a c t
Article history:
The influences of argon and oxygen in helium base shielded GTA welding on the arc ignitabil-
Received 11 October 2007
ity, bead protection and weld penetration are systematically investigated by bead-on-plate
Received in revised form
welding on SUS304 stainless steel. Experimental results show that the critical electrode tip
18 March 2008
work distance for arc ignition is increased from 1 mm under pure He shielding to 5 mm
Accepted 24 March 2008
under He–50%Ar shielding. Small addition of oxygen content to the He–Ar mixed shielding
can significantly change the weld shape from a wide shallow type to a narrow deep one, and
the weld depth/width ratio can be doubled due to the change in the Marangoni convection
Keywords:
from an outward to an inward direction.
Arc ignition
© 2008 Elsevier B.V. All rights reserved.
Weld shape
Marangoni convection
Oxygen
Mixed shielding
1.
Introduction
Gas tungsten arc welding (GTAW), also called tungsten inert
gas welding (TIG), is a widely used welding method for stainless steel, titanium alloy and so on in industry for its high weld
quality, good protection and related lower equipment investment. However, the current bearing capacity of the tungsten
electrode is limited in the welding process. Even if a large heat
input by optimizing the welding parameters is applied, the
increase in weld width is always larger than the increase in
weld depth, which decreases the weld depth/width ratio and
welding productivity. Therefore, deep penetration with a narrow weld width cannot be obtained effectively by changing the
welding parameters. Generally, the weld depth of single-pass
GTA welding is below 3 mm and shallow penetration becomes
the main disadvantage.
Improvement in deep GTAW penetration has been a concern for a long time. Experimental research showed that
the GTA weld shape on stainless steel varied with the rawmaterial composition by adding some minor elements, such
as Se (Heiple and Ropper, 1981), Bi (Takeuchi et al., 1992), O
(Pollard, 1988) and S (Heiple and Ropper, 1982), which is of particular interest to steel makers who supply the raw materials.
However, some minor elements such as S and Se are impurities in steel-making and should be limited, so this method is
not widely applied in industry. A novel modification to the TIG
∗
Corresponding author at: Shenyang National Laboratory for Materials Science, IMR, CAS, Shenyang 110016, PR China. Tel.: +86 24 23971973;
fax: +86 24 23971429.
E-mail address: shplu@imr.ac.cn (S. Lu).
0924-0136/$ – see front matter © 2008 Elsevier B.V. All rights reserved.
doi:10.1016/j.jmatprotec.2008.03.043
1232
j o u r n a l o f m a t e r i a l s p r o c e s s i n g t e c h n o l o g y 2 0 9 ( 2 0 0 9 ) 1231–1239
process, smearing a layer of active flux (halides or oxides) on
different plates including SAF2205 duplex stainless steel (Kuo
et al., 2001), mild steel (Fan et al., 2001), AISI316 (Howse and
Lucas, 2000), C–Mn steel (Anderson and Wiktorowicz, 1996),
0Cr18Ni9 stainless steel (Liu et al., 2002a,b) and SUS304 stainless steel with TiO2 (Tanaka et al., 2000), oxides and fluorides
(Modenesi et al., 2000) and SS7 flux (Paskell et al., 1997) before
welding, which was first proposed by the E.O. Paton Institute of
Electric Welding in the 1960s (Gurevich and Zamkov, 1966), can
significantly increase the weld penetration and was brought
to many researchers’ attention since the 1980s. Experimental results showed that the GTA weld shape is sensitive to
the quantity of the oxide flux smeared on the plate (Lu et al.,
2002, 2003). It is difficult for the weld operator to control the
exact quantity used by hand smearing before welding. Adding
a small amount of active gas, such as SO2 (Heiple and Burgardt,
1985), O2 or CO2 (Lu et al., 2004a,b, 2007), to the inert shielding
gas was proposed to transfer the minor element to the weld
pool and to increase the weld penetration. The oxygen content
in the weld pool and the weld penetration can be effectively
adjusted by a very small addition of O2 or CO2 to Ar or He inert
base shielding. This finding is considered to facilitate developing an advanced automatic TIG welding process with deep
penetration because the amount of oxygen content in a weld
pool can be more precisely controlled by adding an active gas
into an inert shielding gas than by smearing an active flux on
the plate.
Even though some types of active fluxes for the A-TIG process have been used in industry, there is still no common
understanding on the mechanism of A-TIG weld penetration.
Four proposed mechanisms have been proposed to explain
the A-TIG phenomena, and the research work on this continues. The first is based on the concept that the surface tension
of a molten pool is lowered and the pool surface is likely to
descend due to the arc pressure, resulting in arc concentration at the descended portion of the pool. This mechanism
is called the TIG keyhole mode (Savitskii and Leskov, 1980).
In the second mechanism, it is considered that vaporized
flux molecules, such as the mixture of metal oxides (Howse
and Lucas, 2000), Fluorides (Leconte et al., 2007) and TiO2
(Rodrigues and Loureiro, 2005), contract the welding arc. The
third one, proposed by Heiple (Heiple and Ropper, 1981, 1982;
Heiple et al., 1983), Lu et al. (2002, 2003) and Leconte et al.
(2006a,b), is based on the following hypothesis: a reverse
Marangoni convection is induced by a change in the temperature coefficient of the surface tension from negative to
positive when the concentration of a surface active element
exceeds the critical level. The fourth mechanism is called the
insulation mode and is proposed by Lowke et al. (2004).
Former experimental researches showed that adding a
small amount of an active gas, CO2 or O2 , to the inert gas Ar (Lu
et al., 2004a,b) or He (Lu et al., 2007), can significantly change
the shallow wide GTA weld shape to a narrow deep one and
can increase the weld penetration. Especially under the He–O2
shielding (Lu et al., 2007), an ultra-deep penetration GTA weld
shape, over 1.0 depth/width ratio and nearly 10 mm weld pool
depth in single pass welding was obtained under 160 A welding current, 0.75 mm/s welding speed and 1 mm electrode tip
work distance. The GTA weld shape is similar to a high-energy
beam-welding bead. However, the arc ignitability and protection of He shielding gas is poor compared with Ar shielding gas
because the ionization energy of He (24.58 eV) is higher than
that of Ar (15.76 eV) and the density of He (0.1667 kg/m3 ) is
much lower than that of Ar (1.656 kg/m3 ) (Chen, 2002). Sometimes, when the electrode tip work distance (arc length) is over
1 mm under pure He shielding, the arc ignition is difficulty. In
this study, Ar gas is mixed into the He gas to investigate its
effect on the arc ignitability and protection of the weld bead
under He–Ar mixed shielding. A small amount of oxygen is
then mixed into the He–Ar mixed shielding gas to study the
influence of the active gas, oxygen, on the GTA weld shape on
SUS304 stainless steel.
2.
Experimental
SUS304 stainless steel plates with a sulfur content of 0.002 wt%
pct and an oxygen content of 0.002 wt% pct were used as the
substrate with dimensions of 100 mm × 50 mm × 10 mm. The
detailed chemical composition of the substrate is shown in
Table 1. Before welding, the plate was ground using 80-grit
flexible abrasive paper and cleaned with acetone. Bead-onplate welding experiments were carried out on the center of
the plate by a direct current electrode negative (DCEN) polarity
power source (YC-300BZ1) with a mechanized system in which
the test piece moves at a constant speed under the torch.
The arc ignition and protection experiments were done under
He–Ar shielding. The effect of oxygen on the GTA weld shape
was carried out under He–Ar–O2 mixed shielding. In the welding process, the plate is in a horizontal position, and the torch
is perpendicular to the plate. The other welding parameters
are given in Table 2.
Table 2 – Welding parameters
Parameters
Value
Electrode type
Diameter of electrode
Vertex angle of electrode
Shield gas
Gas flow rate
Bead length
Spot time
Welding current
Welding speed
DCEN, W–2%ThO2
2.4 mm
60◦
He–Ar, He–Ar–O2
10 L/min
50 mm
3s
160 A
2 mm/s
Table 1 – Chemical composition of SUS304 stainless steel
Alloy element
Content (wt%)
C
Si
Mn
Ni
Cr
P
S
O
0.06
0.44
0.96
8.19
18.22
0.027
0.0020
0.0020
Fe
Bal.
1233
j o u r n a l o f m a t e r i a l s p r o c e s s i n g t e c h n o l o g y 2 0 9 ( 2 0 0 9 ) 1231–1239
After the welding, the weld beads were sectioned
and specimens for weld shape observation were prepared
using standard metallographic techniques and etched with
HCl + Cu2 SO4 solution to reveal the bead shape and size.
The weld metal oxygen content is measured by an oxygen/nitrogen analyzer (Horiba EMGA-520). Samples for the
oxygen measurement were cut directly from the weld metal,
then polished and cleaned in acetone by ultrasonic vibration
before analyzing.
Shielding gas
Results and discussion
3.1.
Arc ignition under He–Ar mixed shielding
Argon gas was added to helium gas to investigate its effect
on the arc ignition. Electrode tip work distance was varied
from 1 mm to 13 mm with a constant welding current of 160 A.
The critical electrode tip work distance for arc ignition under
He–(0–90)%Ar mixed shielding is shown in Table 3. Under pure
He shielding, arc ignition failed when the electrode tip work
distance exceeded 1 mm. With the increase of the Ar content
in the He–Ar shielding gas, the critical electrode tip work distance obviously increased from 1 mm under pure He to 12 mm
under He–90%Ar shielding. Adding argon gas to helium gas
can improve the arc ignitability of helium gas and make the
arc stable because the Ar ionization energy (15.76 eV) is much
lower than the He ionization energy (24.58 eV).
Bead protection under He–Ar shielding
Fig. 1 shows the bead morphology under He–Ar shielding with
an Ar content from 0% to 90%, under a 3 mm electrode tip work
distance with a welding current of 160 A and a welding speed
of 2.0 mm/s. The weld bead surface is clean and well protected
when the argon content in the He–Ar shielding is over 30% as
shown in Fig. 1c–f. Under the pure He or He–10%Ar shielding,
Electrode tip work
distance (mm)
Arc voltage
(V)
Arc
ignition
Pure
He
1
2
12.5
–
Succeed
Failed
He–10%Ar
1
2
12.2
–
Succeed
Failed
2
3
4
4
5
6
13.4
14.5
–
15.1
15.4
–
Succeed
Succeed
Failed
Succeed
Succeed
Failed
He–70%Ar
7
8
9
17.0
17.0
–
Succeed
Succeed
Failed
He–90%Ar
10
11
12
13
17.4
17.8
18.7
–
Succeed
Succeed
Succeed
Failed
He–30%Ar
3.
3.2.
Table 3 – Critical electrode tip work distance for arc
ignition under He–Ar mixed shielding
He–50%Ar
the weld bead surface is dirty and oxidized as shown in Fig. 1a
and b. Since the density of He (0.1667 kg/m3 ) is much lower
than that of the atmosphere (1.29 kg/m3 ) and the density of
Ar (1.656 kg/m3 ) is higher than that of the atmosphere, the
protection of He shielding gas is poor compared with that of
Ar shielding gas. The weld bead can be well protected when
the Ar content in the He-base shielding reaches 30%.
3.3.
Effect of oxygen on the GTA weld shape under
He–Ar–O2 mixed shielding
The experimental results above clearly showed that arc
ignitability, stability and bead protection can be well improved
when a certain quantity of Ar is mixed into He gas, which is of
Fig. 1 – Bead morphology under He–Ar shielding.
1234
j o u r n a l o f m a t e r i a l s p r o c e s s i n g t e c h n o l o g y 2 0 9 ( 2 0 0 9 ) 1231–1239
Fig. 2 – Weld shape variation with He–30%Ar–O2 shielding under 1 mm electrode tip work distance.
benefit to the real applications in industry. Here two kinds of
mixed shielding gases, He–30%Ar and He–50%Ar, are selected
as the base shielding gas, and a small amount of oxygen gas is
added to the He–Ar shielding gas to study its effect on the oxygen solution in welding pool and the weld shape variations.
Figs. 2 and 3 show the weld shape variations with the
different oxygen content from 0% to 1.0% in He–30%Ar–O2
and He–50%Ar–O2 shielding, respectively, at 1 mm electrode
tip work distance. It is clear that the weld shape is shallow and wide under He–30%Ar and He–50%Ar shielding as
shown in Figs. 2a and 3a. When a small amount of oxygen is mixed into the He–Ar shielding gas, the weld shape
changes from the shallow wide type to the deep narrow one
as shown in parts b–f of Figs. 2 and 3. The weld metal oxygen analysis was carried out after welding to show the effect
of torch gas oxygen content on the oxygen absorption in the
weld pool. The weld metal oxygen contents and the weld
depth/width ratio are plotted versus the torch gas oxygen content in Figs. 4 and 5 for He–30%Ar–O2 and He–50%Ar–O2 mixed
shielding, respectively. The weld depth/width ratio is around
0.35 under He–30%Ar and He–50%Ar mixed shielding, and suddenly increases to over 0.8 when the torch gas oxygen content
is over 0.1%. The weld metal oxygen content also increases
with the torch gas oxygen content.
Heat transfer in the welding pool by conduction and convection is the main factor affecting the final weld pool shape
for GTA welding. The heat transfer by conduction is based on
the weld plate thermal properties. Here, all the experiments
are carried out on the same material and the conduction effect
should be nearly same. Therefore, heat transfer by convection
becomes the main factor influencing the weld shape. In the
GTA weld pool, the liquid metal convection is controlled by the
combination of electromagnetic force, Marangoni convection
induced by surface tension, arc plasma drag force and buoyancy. Among the four forces, only the direction of Marangoni
convection induced by surface tension possibly changes in the
welding process.
Generally, the surface tension decreases with the increasing temperature, that is ∂/∂T < 0, for a pure metal and many
alloys. In the weld pool for such materials, the surface tension
is larger in the relatively cooler part of the pool edge than that
in the pool center under the arc. Hence, the fluid flows from
the pool center to the edge, and the heat flux easily transfers to
the edge and forms a wide and shallow weld shape as shown
j o u r n a l o f m a t e r i a l s p r o c e s s i n g t e c h n o l o g y 2 0 9 ( 2 0 0 9 ) 1231–1239
1235
Fig. 3 – Weld shape variation with He–50%Ar–O2 shielding under 1 mm electrode tip work distance.
in Fig. 6a. Heiple and Ropper (1981, 1982) proposed that some
active elements, such as O, S and Se, can change the temperature coefficient of surface tension for iron alloys from negative
to positive, ∂/∂T > 0, when their quantity is over a critical
value. In this case, the Marangoni convection on the pool sur-
face is changed from an outward to an inward direction, and
a relatively deep and narrow weld shape is obtained as shown
in Fig. 6b. Former research results showed that oxygen is an
active element and that the critical value changing the tem-
Fig. 4 – Effect of torch gas oxygen content on the weld
metal oxygen content and weld depth/width ratio with
He–30%Ar–O2 shielding under 1 mm electrode tip work
distance.
Fig. 5 – Effect of torch gas oxygen content on the weld
metal oxygen content and weld depth/width ratio with
He–50%Ar–O2 shielding under 1 mm electrode tip work
distance.
1236
j o u r n a l o f m a t e r i a l s p r o c e s s i n g t e c h n o l o g y 2 0 9 ( 2 0 0 9 ) 1231–1239
Fig. 6 – Marangoni convection mode in the weld pool: (a)
∂/∂T < 0; (b) ∂/∂T > 0.
perature coefficient of surface tension is around 150 ppm and
100 ppm for pure iron (Taimatsu et al., 1992) and SUS304 stainless steel (Lu et al., 2004a), respectively.
Experimental results in Figs. 4 and 5 show that the
oxygen content in weld metal suddenly increases from
19.6 ppm (He–30%Ar shielding) and 18.5 (He–50%Ar shielding) to 134.5 ppm (He–30%Ar–0.1%O2 shielding) and 121.7 ppm
(He–50%Ar–0.1%O2 shielding) when 0.1%O2 is added to the
shielding gas. The outward convection under He–30%Ar and
He–50%Ar shielding changes to an inward direction when a
small oxygen content is mixed into the torch gas, which causes
the weld shape to change from a shallow wide type to a narrow
deep one as shown in Figs. 2 and 3.
The effect of electrode tip work distance on the weld
shape and weld metal oxygen content under He–30%Ar–O2
and He–50%–O2 shielding is also investigated. Figs. 7–10 are
the results at 3 mm electrode tip work distance. Compared
with the results at 1 mm electrode tip work distance as shown
in Figs. 2–5, it is interesting to find that, when the torch gas
oxygen content exceeds 0.2%, the weld shapes change to a
shallow wide type with a concave bottom as shown in parts
d–f of Figs. 7 and 8, which is quite different from the results at
1 mm electrode tip work distance as shown in Figs. 2 and 3.
Fig. 7 – Weld shape variations with He–30%Ar–O2 shielding under 3 mm electrode tip work distance.
j o u r n a l o f m a t e r i a l s p r o c e s s i n g t e c h n o l o g y 2 0 9 ( 2 0 0 9 ) 1231–1239
1237
Fig. 8 – Weld shape variations with He–50%Ar–O2 shielding under 3 mm electrode tip work distance.
Also the weld depth/width ratio suddenly decreases again
when the torch gas oxygen content is over 0.2% as shown in
Figs. 9 and 10.
Thermodynamic calculations for the reactions of oxide
formation have been studied under the assumption that
the weld pool is considered to be an Fe–M–O system (Lu
et al., 2004a), and the results are shown in Table 4. The
weld metal oxygen content is over 180 ppm when the torch
gas oxygen content is 0.1% as shown in Figs. 9 and 10. For
the SiO2 and Cr2 O3 oxide formation reactions, the calcu-
Fig. 9 – Effect of torch gas oxygen content on the weld
metal oxygen content and weld depth/width ratio with
He–30%Ar–O2 shielding under 3 mm electrode tip work
distance.
Fig. 10 – Effect of torch gas oxygen content on the weld
metal oxygen content and weld depth/width ratio with
He–50%Ar–O2 shielding under 3 mm electrode tip work
distance.
1238
j o u r n a l o f m a t e r i a l s p r o c e s s i n g t e c h n o l o g y 2 0 9 ( 2 0 0 9 ) 1231–1239
Table 4 – Equilibrium oxygen content and oxygen partial pressure by thermodynamic calculations welding parameters
Oxides
Equilibrium oxygen content [%O]
1773 K
FeO
SiO2
Cr2 O3
MnO
0.1477
0.0025
0.0023
0.0263
1873 K
0.2289
0.0072
0.0062
0.0570
Equilibrium oxygen partial pressure (atm)
2273 K
1773 K
−9
1.6 × 10
4.7 × 10−13
3.8 × 10−13
5.2 × 10−11
0.8985
0.1922
0.1495
0.5122
lated equilibrium oxygen contents in the liquid iron are very
low, 0.0025% (25 ppm) and 0.0023% (23 ppm) at 1773 K, and
0.0072% (72 ppm) and 0.0062% (62 ppm) at 1873 K as shown in
Table 4. These results are lower than the weld metal oxygen
content as shown in Figs. 9 and 10 when the torch gas oxygen content is over 0.1%. Therefore, SiO2 and Cr2 O3 oxides
possibly form on the liquid pool edges in the welding process. However, when the temperature is increased to 2273 K,
the calculated equilibrium oxygen contents for the SiO2 and
Cr2 O3 formation reactions are 0.1922% (1922 ppm) and 0.1459%
(1459 ppm), respectively, which is much higher than the weld
metal oxygen content shown in Figs. 9 and 10. The SiO2 and
Cr2 O3 may not be generated at high temperature on the pool
center.
Increasing the electrode tip work distance will weaken the
protection of the weld pool from the atmosphere in the welding process. Furthermore, a large electrode tip work distance
will widen the heat flux distribution on the liquid pool and
widen the relatively low temperature pool edge. Therefore,
with a large electrode tip work distance, the pool edge is
easily oxidized in the welding process. Fig. 11 proposed one
model to illustrate the weld pool convection at 3 mm electrode tip work distance with He–Ar–O2 mixed shielding. Under
He–30% and He–50%Ar shielding, the weld metal oxygen contents are 12.3 ppm and 11.8 ppm, respectively, which are below
the critical value, 100 ppm, and the Marangoni convection on
the liquid pool surface is in the outward direction as shown
in Fig. 11a. The weld shape is wide and shallow with a flat
bottom as shown in Figs. 7a and 8a. When the torch gas oxygen content exceeds 0.1% in He–30%Ar–O2 and He–50%Ar–O2
shielding, the weld metal oxygen content suddenly increases
to 188.4 ppm and 198.4 ppm, respectively, which is over the
critical value of 100 ppm, and the Marangoni convection on
the liquid pool changes to the inward direction as illustrated
in Fig. 11b. Deep penetration is obtained as shown in parts b
and c of Figs. 7 and 8. In this case, the Cr2 O3 and SiO2 oxide
film possibly forms on the pool periphery area based on the
thermodynamic calculation in Table 4. Since the torch gas oxygen content is below 0.2%, the oxide film is supposed to be
thin and easily destroyed by the plasma drag force and surface tension force. The oxide layer is discontinuous as shown
in Fig. 11b. With the increasing of the torch gas oxygen content, the oxide layer trends to become thicker and continuous,
covering on the pool periphery as shown in Fig. 11c. The liquid pool/oxide layer interface is present instead of the liquid
pool/gas surface. In this case, the Marangoni convection due
to the liquid pool surface tension at the peripheral area is
no longer the main factor. However, in the pool center area,
an inward Marangoni convection still exists because there
is no oxide formed on the pool center. The inward convec-
1873 K
−9
9.2 × 10
9.0 × 10−12
6.8 × 10−12
5.9 × 10−10
2273 K
2.0 × 10−6
9.1 × 10−8
5.5 × 10−8
6.4 × 10−7
Fig. 11 – Mode for liquid pool convection under 3 mm
electrode tip work distance.
tion only at the pool center area transfers the hot liquid melt
from the center to the bottom. As a result, the weld shape
becomes wide with a concave bottom as shown in parts d–f of
Figs. 7 and 8.
Adding a small amount of oxygen to inert shielding gas will
significantly increase the TIG weld depth. However, the oxygen presence in the shielding gas will dramatically increase
the electrode consumption and cause the drop of efficiency of
TIG welding. How to prevent the electrode from oxidation in
welding process is another research topic for us in the future.
4.
Conclusions
(1) Arc ignitability and stability of He shielded GTA welding
can be significantly improved when Ar gas is mixed into
the He shielding. The critical electrode tip work distance
j o u r n a l o f m a t e r i a l s p r o c e s s i n g t e c h n o l o g y 2 0 9 ( 2 0 0 9 ) 1231–1239
for arc igniting is increased from 1 mm under pure He
shielding to 5 mm under He–50%Ar mixed shielding.
(2) When the argon content in the He–Ar mixed shielding is
below 30%, the shielding gas protection to the weld bead
from the atmosphere is poor, and the weld bead surface is
dirty and oxidized.
(3) By adding a small amount of oxygen to the He–30%Ar
and He–50%Ar shielding, the GTA weld shape significantly
changes from a wide shallow type to a narrow deep one.
The weld depth/width ratio also increases from 0.35 to 0.8
at a 1 mm electrode tip work distance and from 0.20 to 0.50
at a 3 mm electrode tip work distance. Oxygen is an active
element for the SUS304 stainless steel pool. It changes
the temperature coefficient of surface tension from negative to positive when its content in the liquid pool is over
100 ppm. As a result, the Marangoni convection induced
by surface tension changes from an outward to an inward
direction and the weld penetration can be significantly
increased.
(4) The electrode tip work distance is an important parameter affecting the GTA weld shape in He–30%Ar–O2
and He–50%–O2 shielding. At 1 mm electrode tip work
distance, all the weld shapes are narrow and deep
when the torch gas oxygen content is over 0.1%. However, at a 3 mm electrode tip work distance, the weld
shape becomes wide and shallow again when the torch
gas oxygen content is over 0.3% in He–30%Ar–O2 and
He–50%Ar–O2 shielding. The large electrode tip work distance weakens the protection to the weld bead from
the atmosphere in the welding process and widens the
heat distribution on the pool surface. Hence, it causes
the pool periphery area to be easily oxidized and the
Marangoni convection is then not the main force in the
area.
Acknowledgement
This study was supported by New Energy and Industrial Technology Development Organization (NEDO) of Japan, the 21st
Century and global COE Program, the ISIJ research promotion
grant, the JFE 21st Century Foundation, the Creative Fund of
Institute of Metal Research, Chinese Academy of Science (IMR,
CAS) and the Yong Research Fund of Shenyang City.
references
Anderson, P.C.J., Wiktorowicz, R., 1996. Improving productivity
with A-TIG welding. Weld. Met. Fabr. 64, 108–109.
Chen, Z.Y., 2002. Welding Engineer Handbook. Mechanical
Industry Press, Beijing, pp. 357, pp. 911.
Fan, D., Zhang, R., Gu, Y., Ushio, M., 2001. Effect of flux on A-TIG
welding of mild steel. Trans. JWRI 30, 35–40.
Gurevich, S.M., Zamkov, V.N., 1966. Nekotorye osovennosti svarki
titana neplaviachimsia zlektrodom s primeneniem flynsov.
Avtom. Svarka. 12, 13–16.
Heiple, C.R., Ropper, J.R., 1981. Effect of selenium on GTAW
geometry. Weld. J. 60, 143s–145s.
Heiple, C.R., Ropper, J.R., 1982. Mechanism for minor effect on
GTA fusion zone geometry. Weld. J. 61, 97s–102s.
1239
Heiple, C.R., Roper, J.R., Stagner, R.T., 1983. Surface active element
effects on the shape of GTA, laser, and electron beam welds.
Weld. J. 62, 72s–77s.
Heiple, C.R., Burgardt, P., 1985. Effects of SO2 shielding gas
additions on GTA weld shape. Weld. J. 64, 159s–162s.
Howse, D.S., Lucas, W., 2000. Investigation into arc constriction by
active fluxes for tungsten inert gas welding. Sci. Tech. Weld.
Join. 5, 189–193.
Kuo, M., Sun, Z., Pan, D., 2001. Laser welding with activating flux.
Sci. Tech. Weld. Join. 6, 17–22.
Leconte, S., Paillard, P., Chapelle, P., 2006a. Effect of oxide fluxes
on activation mechanisms of tungsten inert gas process. Sci.
Tech. Weld Join. 11, 389–397.
Leconte, S., Paillard, P., Saindrenan, J., 2006b. Effect of fluxes
containing oxides on tungsten inert gas welding process. Sci.
Tech. Weld Join. 11, 43–47.
Leconte, S., Paillard, P., Chapelle, P., 2007. Effects of flux
containing fluorides on TIG welding process. Sci. Tech. Weld
Join. 12, 120–126.
Liu, F., Lin, S., Yang, C., Wu, L., 2002a. Effect of active flux on weld
shape. Trans. Chin. Weld. Inst. 23, 1–4.
Liu, F., Lin, S., Yang, C., Wu, L., 2002b. Effect of activating flux and
welding parameters on the weld depth/width ratio in A-TIG
welding. Trans. Chin. Weld. Inst. 23, 5–8.
Lowke, J.J., Tanake, M., Ushio, M., 2004. Insulation effects of flux
layer in producing greater weld depth. In: The 57th Annual
Assembly of International Institute of Welding, Osaka, Japan,
IIW Doc.212-1053-04.
Lu, S.P., Fujii, H., Sugiyama, H., Tanaka, M., Nogi, K., 2002. Weld
penetration and Marangoni convection with oxide fluxes in
GTA welding. Mater. Trans. 43, 2926–2931.
Lu, S.P., Fujii, H., Sugiyama, H., Nogi, K., 2003. Mechanism and
optimization of oxide fluxes for deep penetration in gas
tungsten arc welding. Metall. Mater. Trans. 34A, 1901–
1907.
Lu, S.P., Fujii, H., Nogi, K., 2004a. Marangoni convection and weld
shape variations in Ar–O2 and Ar–CO2 shielded GTA welding.
Mater. Sci. Eng. A 380, 290–297.
Lu, S.P., Fujii, H., Nogi, K., 2004b. Sensitivity of Marangioni
convection and weld shape variations to weld parameters in
O2 –Ar shielded GTA welding. Scr. Mater. 51, 271–277.
Lu, S.P., Fujii, H., Nogi, K., Sato, T., 2007. Effect of oxygen content
in He–O2 shielding gas on weld shape in ultra-deep
penetration TIG. Sci. Tech. Weld. Join. 12, 689–695.
Modenesi, P.J., Apolinario, E.R., Pereira, I.M., 2000. TIG welding
with single component fluxes. J. Mater. Proc. Tech. 99, 260–265.
Paskell, T., Lundin, C., Castner, H., 1997. GTAW flux increases
weld joint penetration. Weld. J. 76, 57–62.
Pollard, B., 1988. The effects of minor elements on the welding
characteristics of stainless steel. Weld. J. 67, 202s–213s.
Rodrigues, A., Loureiro, A., 2005. Effect of shielding gas and
activating flux on weld bead geometry in tungsten inert gas
welding of austenitic stainless steels. Sci. Tech. Weld Join. 10,
760–765.
Savitskii, M.M., Leskov, G.I., 1980. Mechanizm vliyamia
dzlektrootrida tepvnykh zpementov na proplavlauchyu
sposovnosti luqi s voliframoym katodm Avtom. Svarka 9,
17–22.
Taimatsu, H., Nogi, K., Ogino, K., 1992. Surface tension of liquid
Fe–O alloy. J. High. Temp. Soc. 18, 14–19.
Takeuchi, Y., Takagi, R., Shinoda, T., 1992. Effect of bismuth on
weld joint penetration in austenitic stainless steel. Weld. J. 71,
283s–289s.
Tanaka, M., Shimizu, T., Terasaki, H., Ushio, M., 2000. Effects of
activating flux on arc phenomena in gas tungsten arc welding.
Sci. Tech. Weld. Join. 5, 397–402.