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Journal of Minerals & Materials Characterization & Engineering, Vol. 7, No.3, pp 247-263, 2008
jmmce.org Printed in the USA. All rights reserved
Duplex Stainless Steel Welds and their Susceptibility to Intergranular Corrosion
, L. Ward
and G. Biddle
ARV Offshore, 199/62 LandCo Building, Vipavadee-Rangsit Road Samsaen-Nai, Phayathai,
Bangkok 10400, Thailand.
School of Civil and Chemical Engineering, RMIT University, GPO Box 2476V, Melbourne, Vic.
Surface Materials Analysis Research & Training Group, Department of Applied Physics RMIT
University, GPO Box 2476V, Melbourne, Vic. 3001, Australia
*Corresponding Author: email@example.com
Duplex stainless steels (DSS) as alternatives to conventional austenitic stainless steels for the
construction of pipelines is becoming more wide-spread, particularly for sour service applications
where corrosion resistance / stress corrosion cracking resistance is required in aggressive chloride /
sulphide environments. While these steels show many superior characteristics, limitations are
associated with the welding of these steels, particularly controlling the weld structure and properties
and understanding how the weld metallurgy may influence the susceptibility to intergranular
The focus of this paper is to report on the findings from a detailed study of the various weld
sections within a DSS pipeline, as a function of heat input and type of weld, in terms of the
metallurgical structure, composition and mechanical properties and to assess the
susceptibility to IGC. Welding was performed using the manual Gas Tungsten Arc Welding
(GTAW) technique at both high and low heat input conditions. Two different join configurations
(double bevel single V bevel and double bevel single U joint configuration) were adopted.
Structural analysis consisted of (i) optical microscopy of welded specimens; (ii) ferrite content
determination; (iii) Vickers hardness measurements; (iv) Charpy impact studies and (v) transverse
tensile testing. Two test methods namely a modified ASTM A262 and a modified Double Loop
Electrochemical Potentiokinetic Reactivation (DL-ERP) test was used to determine susceptibility to
IGC. The electrolyte solution used was 0.5M H
+ 0.001 M TA (thioacetamide). The test was
conducted at 60 °C. The potential was scanned from -500 mV (SCE) to +200 mV (SCE) and back to
-500 mV (SCE) at a rate of 1.67 mV/s. The ratio of the reactivation charge to the passivation charge
was calculated. From the results obtained it can be shown that the fill region was most susceptible to
IGC for all weld conditions.
248 B. Gideon, L. Ward and G. Biddle
DSS with a microstructure comprising of proportionate amount of ferrite and austenite, have both
the properties of austenitic and ferritic stainless steels such as high tensile strength, good toughness
at low temperatures and excellent resistance to stress corrosion cracking (1). They are used
increasingly as an alternative to austenitic stainless steels, especially in chloride or sulphide
environments, e.g., in the oil, gas and petrochemical industries, as such is a prime candidate for sour
service application (2). In the welding of DSS, it is essential to maintain a ferrite–austenite ratio
close to 50:50. This phase balance, may however, be upset due to rapid cooling involved in most
weld thermal cycles resulting in weld metal ferrite contents in excess of 50%. In order to restore the
phase balance, weld filler materials are usually overalloyed with Ni than in the base material (3). The
resultant phase ratio is dependent on the energy input during welding, as this determines the cooling
rate and the extent of the phase transformation which is diffusion based. If high heat inputs are used,
coarser grains are produced in the weld region, wide heat-affected zones and possibly, precipitation
of brittle intermetallic phases may develop (4, 5). However, due to the resultant slow cooling rates,
the transformation may yields a more favorable phase balance. It is thus desirable to control welding
conditions such that cooling is slow enough for adequate formation of austenite, but fast enough to
prevent the formation of precipitation (6).
It has been reported (7, 8, 9) that various intermetallic phases (χ, σ
and R) can be present in the
weld metal under different aging conditions. These phases are predominantly associated with
slow cooling rates through specific temperature regions from 600
C to 1000
It has been
shown that the effect of intermetallic phases on the majority of the weld metal properties may
only be significant at approximately 2 % volume (10). Secondary austenite (γ’) has also been
identified in the reheated weld beads of DSS. The corrosion resistances in these welds are
reduced due to the lower Mo, Cr, and N contents in primary austenite, and as a result, welds
have poorer corrosion resistance (11). One way to limit the formation of γ’ in the
weld region is to reduce the heat input in the successive passes. This will reduce the reheating
and cooling cyclic effect on the structure of the root pass, which will also be in contact with
corrosive liquids medium in linepipe. Nilsson et al. (11)
determined that the as welded
microstructure was almost devoid of intermetallic phases, but a small amounts of γ’ was present.
This observation was based on multipass welds with heat inputs of either 0
8 kJ or 1
4 kJ and
interpass temperatures not exceeding 100°C.
Heat input and interpass temperature are the two main welding parameters often cited as
influencing the cooling rate. It has been determined that increasing the heat input increases the
width of the HAZ (12), while the use of high heat input increases the corrosion resistance has
also been reported (13). A simulated multipass welds was developed to show that the
toughness of the HAZ exhibited a maximum when the interpass temperature was maintained
below 100°C. In addition, it was established that the toughness of the HAZ increased with an
Vol.7, No.3 Duplex Stainless Steel Welds 249
increase in cooling rate (14). It has also been demonstrated that as the interpass temperature
increases, the cooling rate decreases (15).
Numerous papers have been published on the microstructures of duplex stainless steel weld
metal and HAZs, but detailed studies of the as welded structures and their susceptibility to
sensitization are limited.
The aim of this study is to conduct a detailed analysis of the various weld sections within a
DSS pipeline, as a function of heat input and type of weld, in terms of the metallurgical
structure, composition and mechanical properties and to assess the susceptibility to IGC.
2. EXPERIMENTAL PROCEDURES
2.1. Welding Conditions
The parent material chosen for the investigation was a 10mm wall thickness, 250mm diameter DSS
linepipe corresponding to UNS 31803. The filler material used was the conventional ER2209 AWS
A5.9-93 classification. Full details of the chemical composition of both the parent material and filler
material are listed in Table 1, confirming that the primary solidification mode was ferrite.
Table 1. Chemical composition of pipe and filler material
C Mn P S Si Ni Cr Mo N Cu Pren Cr
Min - - - - - 5.00 21.50 3.00 0.15 - 35 - -
Max 0.030 2.0 0.025 0.015 1.0 6.50 23.00 5.50 0.20 0.16 - 32.04 10.78
Material Max 0.016 1.69 - - 0.42 8.60 23.07 3.20 0.160 0.16 ≥35 28.09 11.90
= Cr+1.37Mo+1.5Si+2Nb+3Ti and Ni
Welding was performed using the manual Gas Tungsten Arc Welding (GTAW) technique. Two
different join configurations were adopted, namely double bevel single V bevel and double bevel single
U joint configuration. Details are given in Table 2.
Upon completion of welding, all test conditions were visually inspected for surface defects both
internally and externally. Liquid dye penetrent tests were performed 4 hours after completion of
welding, and radiography (X-ray) was performed 24 hours later to determine the integrity of the girth
250 B. Gideon, L. Ward and G. Biddle
Table 2 Weld Conditions
Pass Travel Speed Heat Input
1 (weld root) 51.00 1474.71
2 (weld fill) 123.00 883.12
3 (weld fill) 66.00 1745.45
4 (weld fill) 64.00 1788.00
5 (weld cap) 64.00 1685.63
1 (weld root) 45.00 1591.20
2 (weld fill) 105.00 1440.46
3 (weld fill) 79.00 2756.05
4 (weld cap) 94.00 2216.17
1 (weld root) 110.00 419.78
2 (weld fill) 62.00 757.55
3 (weld fill) 38.00 1733.05
4 (weld fill) 40.00 2194.80
5 (weld cap) 43.00 2041.67
1 (weld root) 115.00 420.31
2 (weld fill) 66.00 942.55
3 (weld fill) 73.00 2219.18
4 (weld cap) 57.00 2135.37
2.2. Mechanical Testing
Ferrite contents of the four weld conditions were measured using the Magna-Gauge, Fischer ferrite-
scope and as a comparison, determined metallographically by point count method (16).
Vickers hardness measurements were made with a 10 kg load in the parent material, HAZ, weld cap,
weld fill and weld root regions. Transverse tensile test specimens were used to determine the tensile
values whilst the more restrictive transverse side bend specimens was used. Charpy impact tests were
performed to assess the notch toughness of samples extracted from the weldments as described in
ASTM A 370 (17).
2.3. Intergranular Corrosion tests (IGC)
Two test methods namely a modified ASTM A262 (25) and a modified Double Loop Electrochemical
Potentiokinetic Reactivation (DL-ERP ) test was employed to determine susceptibility to IGC.
Vol.7, No.3 Duplex Stainless Steel Welds 251
Modified ASTM A262 Standard Practices E—copper–copper sulfate sulfuric acid test for detecting
susceptibility to intergranular attack was used. The specimen was covered with copper shot and
grindings and immersed in a solution of 16 wt% sulfuric acid with 6 wt% copper sulfate. The solution
was then heated to its boiling point and maintained at this temperature for 48 hours. On removal from
solution, the specimen was bent through 180° over a rod with a diameter equivalent to twice the
thickness of the specimen instead of four times the thickness to ensure if cracks appeared they would be
apparent by the more restrictive bending radius. The bent surface of the specimen was then examined for
cracks at low magnifications in the range X5 to X20.
A modified double loop test was used as conducted by Schultz et al (19-21) as described in Fig.1. The
solution used was 0.5M H2SO4 + 0.001M TA (thioacetamide). TA is added to reduce the extent of
ferrite dissolution. The test was conducted at 60 °C. The polarization scan was started 5 minutes after
immersion of the specimen. The potential was scanned from -500 mV (SCE) to +200 mV (SCE) and
back to -500 mV (SCE) at a rate of 1.67 mV/s. The ratio of the reactivation charge to the passivation
charge was calculated and is shown in the results and discussion section. Schultze et al also evaluated
the single loop test described in ASTM G108 for DSS on the basis that the extent of ferrite dissolution
would be less in a single loop test than in a double loop test. It was also suggested that the EPR method
could be applied to DSS welds but was not conducted.
(%) <1 ≥1
(%) <1 ≥1
Sensitization criteria based on (28-33, 38)
Fig. 1. Principle of DL-ERP on Degree of Sensitization (DOS).
Although the EPR test is well established for austenitic stainless steels and has been used for DSS (22-
25) the test parameters are dependent on surface preparation and material quality.
Samples for DL-ERP tests were cut from areas as close as possible to where macro sections were taken
from for metallographic examinations and hardness measurements. Specimens were cut along the
252 B. Gideon, L. Ward and G. Biddle
centerline of the welds to accommodate for an exposed area of 1cm
required for the DL-ERP testes and
to reveal the relevant weld layers (Root, Fill and Cap).
3. RESULTS AND DISCUSSION
The microstructure, resulting phase transformation, mechanical properties and degree of susceptibility to
IGC are discussed in detail in this section. Microstructural characterization was performed using optical
metallography, macrohardness and energy dispersive X-ray spectroscopy (EDS). Table 3 summarises
the results of the mechanical properties of the welded duplex stainless steels carried out in this study,
while Fig. 2 shows optical micrographs of the welded regions according to the various welding
conditions, as indicated in Table 3.
Table 3 Summary of the results of the mechanical properties of the welded duplex stainless steels.
Scope Point Count DL-ERP Test
J % % % Qr/Qa Ir/Ia
1 (weld root)
49.50 35.00 41.02 0.04 0.06
2 (weld fill) - - - - -
3 (weld fill) - 36.00 38.67 0.09 0.12
4 (weld fill) - - - - -
5 (weld cap) 48.25 37.00 43.67 0.04 0.07
48.88 36.00 41.12 0.06 0.08
1 (weld root)
39.75 35.00 35.31 0.05 0.07
2 (weld fill) - 35.00 36.48 0.08 0.10
3 (weld fill) - - - - -
4 (weld cap) 48.00 45.00 46.33 0.02 0.04
43.88 38.33 39.37 0.05 0.07
1 (weld root)
43.00 40.00 42.58 0.05 0.08
2 (weld fill) - - - - -
3 (weld fill) - 35.00 36.95 0.10 0.12
4 (weld fill) - - - - -
5 (weld cap) 43.75 36.00 40.03 0.03 0.06
43.38 37.00 39.85 0.06 0.08
1 (weld root)
54.50 42.00 45.47 0.07 0.10
2 (weld fill) - 35.00 33.75 0.08 0.09
3 (weld fill) - - - - -
4 (weld cap) 43.75 36.00 37.19 0.07 0.10
49.13 37.67 38.80 0.07 0.10
Vol.7, No.3 Duplex Stainless Steel Welds 253
Fig. 2. Optical micrographs of the duplex stainless steels at various weld conditions; a) Root region
associated with weld condition 1, b) Fill region associated with weld condition 2, c) Fill region
associated with weld condition 3, d) Cap region associated with weld condition 4 (X1000
254 B. Gideon, L. Ward and G. Biddle
3.1. Microstructural Evaluation
Microstructural analysis for all four GTAW weld conditions as shown by the optical micrographs in Fig
2 reveals the presence of a two-phase banded structure, typical of such materials. In general, the
austenite regions observed in the DSS weld metal is formed from ferrite in three modes, viz., as
allotriomorphs at the prior-ferrite grain boundaries, as Widmanstätten side-plates growing into the grains
from these allotriomorphs and as intragranular precipitates. In the micrographs, the grain boundary
allotriomorphs and Widmanstätten austenite are clearly seen. However, the austenite seen within the
grain could be either intragranular precipitates or Widmanstätten austenite intercepted transverse to long
axis. Fig. 2a shows that the grain boundary austenite layer is not continuous. It is further known that the
formation of grain boundary and side-plate fractions requires a relatively smaller driving force (22) and
therefore can occur at higher temperatures with little undercooling. The formation of intragranular
acicular ferrite, on the other hand, requires a greater degree of undercooling and therefore occurs at
lower transformation temperatures. It is likely that a similar transformation sequence is adopted during
the microstructural evolution of the regions associated with the DSS weld. Thus the grain boundary
austenite and Widmanstätten side-plates form early at higher temperatures, while the intragranular
austenite particles require a greater driving force and precipitate later at a lower temperature. This can
also be seen in Fig. 2d where the intragranular precipitates are seen to form in the regions partitioned by
the Widmanstätten plates. When cooling occurs rapidly in the cap region of the welds, it is expected that
the transformation product requiring a higher degree of undercooling is formed.
3.2. Mechanical Properties
A graphical representation of the Vickers hardness measurements (10 kg load) on the various welds,
taken transversally over the weld regions, as depicted below each graph, are given in; Fig. 3a) for
condition 1 and 2 and Fig. 3b) for condition 3 and 4. In the as-welded condition, all the four fusion
zones exhibit hardness values in the range of 235-285HV
. The average value in the root run for
Condition 1 was 274 HV
, with a maximum value of 279HV
. The results taken from condition 4
generally showed values higher than 274HV
Typical values for the different weld regions were: for
the root (281-287HV
), for the fill (274-276HV
) and for the cap (262-276HV
), with the following
average values of root (280HV
), fill (273HV
) and cap (268HV
). As for Condition 2 and 3 all weld
passes had values lower than 247HV
The findings here are significant because there is little difference in the ferrite–austenite proportion in
these fusion zones (Root, Fill and Cap). It has been reported that the ferrite and austenite phases do not
differ much in composition because substitution elements do not have time to partition significantly
during DSS welding (23). However, this was found not to be the case in all Weld Conditions. A
comparison of the hardness values and varying amounts of ferrite in the welds show a correlation to the
hardness level except for condition 4 where the reverse was observed as shown in Fig.4.
Vol.7, No.3 Duplex Stainless Steel Welds 255
0 1 2 3 4 5 6 78 91011
Cond.1 (Cap)Cond.1 (Fill)Cond.1 (Root)
Cond.2 (Cap)Cond.2 (Fill)Cond.2 (Root)
Fig. 3 a) - Hardness profile for Condition 1 and Condition 2
0 1 2 3 4 5 6 7891011
Cond.3 (Cap)Cond.3 (Fill)Cond.3 (Root)
Cond.4 (Cap)Cond.4 (Fill)Cond.4 (Root)
Fig. 3b) - Hardness profile for Condition 3 and Condition 4
Cond. 1 Root
Cond. 1 Fill
Cond. 1 Cap
Cond. 2 Root
Cond. 2 Fill
Cond. 2 Cap
Cond. 3 Root
Cond. 3 Fill
Cond. 3 Cap
Cond. 4 Root
Cond. 4 Fill
Cond. 4 Cap
Fig. 4. Correlation of Hardness values and % ferrite in all conditions.
The notch toughness values obtained by Charpy impact testing at -43 °C are listed in Table 3. The
values show that, although slight differences exist in the ferrite content among the four test conditions,
condition 1 and 2 had impact values of 116 joules while condition 3 and 4 have impact energy values of
97 joules and 100 joules, indicating a reduction of 16.4% and 13.8% respectively. This suggests that the
differences in the toughness cannot be explained simply in terms of the changes in ferrite content. The
fact that all four weld conditions retain much of their toughness even at -43 °C may not be due to a
reduction in ferrite–austenite ratio alone. Tensile values for all 4 conditions were within the range of
to 794 N/mm
256 B. Gideon, L. Ward and G. Biddle
3.3. Intergranular Corrosion Tests
The results for the modified A262 Standard Practice E test for all conditions were found to be
acceptable, with no intermetallic phases observed during metallographic examination. Such findings
can also be substantiated by the high impact energy values obtained (97-116 joules). No evidence of
sensitization could be observed in all four weld conditions even under a restricted and reduced bending
radius as shown in Fig.5.
Fig. 5. ASTM A262 Standard Practices E test results, a) Condition 1, b) Condition 2, c) Condition 3, d)
3.4. Modified DL-ERP test
The sensitivity of the DSS welds was evaluated using the DL-EPR test, which consisted of subjecting
the weld regions to potentiokinetic scanning in a solution of 0.5M H
+ 0.001M TA (thioacetamide),
from an active to a passive domain (activation scan), followed by a return to the initial potential
(reactivation scan). The test efficiency was measured by means of a response test, which was
characterized by weak values of the current density ratio (Ir/Ia <1%) and the charge ratio (Qr/Qa < 1%)
for nonsensitized materials, and relatively high ratio values (Ir/Ia ≥1%) and (Qr/Qa ≥ 1%) for high-
sensitized materials. The reverse polarization from the passive to the active region gave rise to a
reactivation peak, the magnitude of which is sensitive to the degree of alloy element depletion. The
susceptibility to corrosion was characterized in terms of both the ratio of the reactivation-current peak to
the activation current peak as well as the ratio of the reactivation charge to the activation charge (24).
Both Ir/Ia and Qr/Qa results are represented graphically in Fig. 6. Analysis of the results shows that the
fill region for all four weld conditions had a higher degree of sensitization (DOS) compared to the root
and cap region of the welds. It was also observed that collectively, condition 4 had the highest DOS for
the root and cap region.
Vol.7, No.3 Duplex Stainless Steel Welds 257
Degree of Sensitization
Fig. 6. Results of DL-ERP test showing DOS for all conditions in terms of Qr/Qa and Ir/Ia.
The Ir/Ia and Qr/Qa results were then compared to the percent of Ferrite transformed in each layer
(Root, Fill and Cap). A trend was observed where the peaks in the graph were indicative of the highest
levels of ferrite in the various conditions correlating to troughs on the Ir/Ia and Qr/Qa lines, and peaks in
the Ir/Ia and Qr/Qa lines (Fig. 7).
Cond. 1 Root
Cond. 1 Fill
Cond. 1 Cap
Cond. 2 Root
Cond. 2 Fill
Cond. 2 Cap
Cond. 3 Root
Cond. 3 Fill
Cond. 3 Cap
Cond. 4 Root
Cond. 4 Fill
Cond. 4 Cap
Fig. 7. Correlation of Qr/Qa, Ir/Ia, DOS and %ferrite in all Conditions.
The results of this study confirm the usefulness of the DL-EPR method in evaluating quantitatively the
sensitization of DSS in accordance with previous studies (19-21). The DL-EPR method has the
advantage of being a fast and quantitative test that can easily be incorporated into the monitoring of
equipment to identify IGC in particular (24).
3.5. Structural Analysis of Corroded Weld Regions
EDS Analysis was performed on the cap region of condition 1 and all weld locations of condition 4, to
determine the distribution of elements present in the austenite and ferrite phases and to provide
confirmation of the mechanisms by which phase transformation were occurring. The various DL-ERP
test results are shown in Fig 8 ,10 and 11. Little information is available in the literature regarding the
258 B. Gideon, L. Ward and G. Biddle
width and depth of depleted zones. Since precipitates were not observed, therefore this could not be
determined. Never the less a clear indication of elements, e.g. Cr and Ni, preferentially nucleating in the
ferrite and austenite regions in relative % is evident in the results obtained (Fig. 11). A correlation
therefore seems to exist between precipitate size and corrosion resistance. This view is supported by
results from a comparative study showing that smaller precipitates formed by welding were less
detrimental than larger produced by isothermal aging (1-7, 22, 23).
Fig. 8. a) SEM image of Cap in Condition 1 etched with 15% NaOH, b1) EDS analysis of Austenite in
Cap region of Condition 1, b2) EDS analysis of Ferrite in Cap region of Condition 1, c) DL-ERP plot of
Root, Fill and Cap region in Condition 1 showing difference in DOS in 0.5M H
+ 0.001 M TA, d)
Anodic Polarization curve of Root, Fill and Cap region in Condition 1, 0.5M H
+ 0.001 M TA.
Vol.7, No.3 Duplex Stainless Steel Welds 259
Fig. 9. a) DL-ERP plot of Root, Fill and Cap region in Condition 2 showing difference in DOS in 0.5M
+ 0.001 M TA, b) Anodic Polarization curve of Root, Fill and Cap region in Condition 2, 0.5M
+ 0.001 M TA. c) DL-ERP plot of Root, Fill and Cap region in Condition 3 showing difference
in DOS in 0.5M H
+ 0.001 M TA, d) Anodic Polarization curve of Root, Fill and Cap region in
Condition 3, 0.5M H
+ 0.001 M TA.
260 B. Gideon, L. Ward and G. Biddle
Fig. 10. a) Secondary electron ( SE) image, showing surface features of Fill region of Condition 4 etched
with 15% NaOH, b) Back-Scattered Electron (BSE) yielding images where different phases of differing
mean atomic number stand out sharply in the Fill region of Condition 4 etched with 15% NaOH c1)
EDS analysis of Ferrite in Fill region of Condition 4, c2) EDS analysis of Austenite region in Fill region
of Condition 4, d) EDS analysis entire Fill region of Condition 4, , e) DL-ERP plot of Root, Fill and Cap
region in Condition 4 showing difference in DOS in 0.5M H
+ 0.001 M TA, f) Anodic Polarization
curve of Root, Fill and Cap region in Condition 4, 0.5M H
+ 0.001 M TA.
Vol.7, No.3 Duplex Stainless Steel Welds 261
0.59 0.4753 0.5675 0.4647 0.4864 0.6569
2.179 1.986 1.792 1.901 2.129 1.985
1.38 1.11 1.187
406B, Fill ((Fer)406C, Fill (Aust)406E, Cap (Fer)406F, Cap (Aust)406H, Root (Fer)406I, Root (Aust)
Fig. 11. EDS results on relative distribution of elements in austenite and ferrite regions of the root, fill
and cap regions associated with Condition 4.
As a consequence of the varying heat inputs from 1.5 kJ to 2.0 kJ, for the 4 different weld conditions,
there was a move towards the development of a structure containing 35-47% ferrite. Analysis of
data given in Tables 2, suggests that control of heat input was not a significant factor in
influencing the specific weld metal mechanical properties within the conditions specified. It was
established that there was no relationship between the percentage ferrite content in the notch region
and the toughness although results of impact testing differ by approximately 16% for Condition 3
and 14% for Condition 4.
The optical microstructure of the weld metal as detailed in Fig. 1 shows a formation of a finer and
coarse structure within the weld metal dependent on the level of undercooling which would be expected
in a welding process, since the weld metal in the cap region would cool down more quickly as a
result of the lower ambient temperature of the weld surface and adjacent parent material at the
start of welding. There was no evidence of secondary austenite (γ’), being present in any of the
weld metal conditions examined. In addition, no areas were observed to contain any of the normally
expected intermetallic phases or carbides, even at higher magnifications (X 80 000 times) used in the
The resultant hardness values in Condition 1 and 3 exhibited higher hardness values. There was a
significant correlation between the hardness values and the % transformation of Ferrite in all
conditions except for Condition 4 where the reverse was observed in the root, fill and cap region
as shown in Fig. 4.
All 4 test conditions passed the qualitative type IGC tests even under restricted and modified
262 B. Gideon, L. Ward and G. Biddle
The DL-ERP test results revealed that the fill area for all 4 test conditions and the base material
had the highest values for Ir/Ia and Qr/Qa. It was also determined that Condition 4 had the highest
DOS based on DL-ERP test conducted even though Condition 4 had the lowest average heat input.
EDS was used for qualitative elemental analysis to determine which elements were present and their
relative abundance. In general, depletion was expected to be the most pronounced in ferrite since (i)
intermetallic particles mainly grow into the ferrite regions and (ii) diffusion is faster in ferrite than in
1. V. Muthupandi, P. Bala Srinavasan, S.K. Seshadri and S. Sunderasen; Materials Science and
Engineering A, Volume 358, issue 1-2, 15 October 2003.
2. B.D. Craig, Sour Gas Design Consideration, SPE Monograph No. 15, 1993.
3. M. Liljas, Proceedings of the Fourth International Conference on Duplex Stainless Steels, Glasgow,
Scotland, Keynote Paper V, vol. 2, 13–16 Nov 1994.
4. A.J. Ramiraz, J.C. Lippold and S.D. Brandi, Metallurgical and Materials Transaction A, Vol. 34A,
Aug 2003, pg. 1575-1597.
5. N. Stephenson, Welding and Metal Fabrication, 55, 5 and 6, 1987, 159-164.
6. R.N. Gun, Editor, Duplex Stainless Steels,: Microstructure, Properties and Applications, Abington
Publishing, Cambridge, England, 1997.
7. L. KARLSSON, L. RYEN, and S. PAK: Weld J, 1995
8. J.O. NILSSON and P. Lru: Mater. Sci. Technol., 1991
(9), 853 - 862.
9. A. RADJARVIIA, G. METuR, and M. GANTOIS: Proc. Conf. Duplex Stainless Steels '91, Beaune,
France, October 1991, Societe Frangaise de Metallurgie et de Mat6riaux, 119-126.
10. L. KARLSSON: Proc. Conf. Duplex Stainless Steels '97, Maastricht, The Netherlands, October
1997; Stainl. Steel World, 1997, 43-58.
11. J.-O. NILSSON, L. KARLSSON, and J.-O. ANDERSSON: Mater. SCL. Technol., 1995, 11
12. S. JANA: `Innovation stainless steel', 3
348; 1993, Florence, Associazione Italiana di
13. S. ATAMERT and J. E. KING: ACta. Metall, 1991, 39
14. H. J. Yu: Proc 4th Int. Conf. on Trends in Welding Research, 87-92; 1995, Materials Park, OH,
15. N.A. Mc Pherson, Y. Li and T.N. Baker, Science and Technology of Welding Journal, 2000,, Vol.
5, No. 4, pg 235-244.
16. ASTM E562, Standard Test Method for Determining Volume Fraction by Systematic Manual Point
17. ASTM A370, Standard Test Methods and Definitions for Mechanical Testing of Steel Products
18. ASTM G108, Standard Test Method for Electrochemical Reactivation (EPR) for
Detecting Sensitization of AISI Type 304 and 304L Stainless Steels
Vol.7, No.3 Duplex Stainless Steel Welds 263
19. S. Schultze, J. Gollner, K. Eick, P. Veit, I. Garz, “ The modified EPR test: A new tool for
examination of corrosion susceptibility of duplex stainless steel”, Duplex ’97, Paper D97-067, KCI
Publishing, Zutphen, The Netherlands (1997), p 639.
20. A. Turnbull, P.E. Francis, A.J. Griffiths, E. Bennett, W. Nimmo, “Measurement of Corrosion
Resistance of Super-Duplex Stainless Steel Welds by Electrochemical Techniques,” Eurocorr” 2000
(London, U.K.: Institute of Materials, 2000).
21. E. Otero, C. Merino, C. Fosca, P. Fernandez, “Electrochemical Characterization of Secondary
Phases in a Duplex Stainless Steel by EPR Test,” Duplex ’94, paper no. 56 (Cambridge, U.K.: TWI,
22. B.J. Ginn and T.G. Gooch: `Effect of Intermetallic Content on Pitting Resistance of Ferritic-
Austenitic Stainless Steels', proc. conf. Stainless Steels'91 Science and market, Chia Laguna
Sardinia, Italy, 1999, vol. 3, p. 81-89.
23. L. Karlsson: 'Intermetallic Phase Precipitation in Duplex Stainless Steels and Weld Metals: Metal-
lurgy, Influence on Properties and Testing Aspects', Welding in the World, vol. 43, no. 5, 1999.
24. T. Amadou, C. Braham, and H. Sidhom; Metallurgical and Materials Transactions A Vol. 35A,
25. ASTM A262, Standard Practices for Detecting Susceptibility to Intergranular Attack in Austenitic
26. Verneau, M., Lojewski, C. and Charles, J., “ Modified EPR test for duplex stainless steels surface
contamination and microstructural investigation, Duplex ‘91, Les Editions de Physique, 1991, pp
27. Merino, P., Novoa, X.R., Pena, G., Porto, E. and Espada, L., “Study of the intergranular corrosion
susceptibility of duplex stainless steels using potentiokinetic reactivation tests (PRT), Duplex ’91,
Les Editions de Physique, 1991, pp 871-875.
28. Scully, J. R. and Kelly, R.G., Corrosion, 42 (1986), 537-542.
29. Otero, E., Merino, C. Fosca, C. and Fernadez, P., “Electrochemical characterisation of secondary
phases in a duplex stainless steel by EPR test”, Duplex ’94, TWI, 1994, Paper 56.