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Paper No. 14169
Prevention of Stress Corrosion Cracking in High
Strength Low Alloy Steel Welds in Boiler Feed Water
Applications
Jacek Kajda
ERA Technology Ltd.
Cleeve Road, Leatherhead,
Surrey KT22 7SA
Tel: +44 (0) 1372 367313
Corresponding Contacts duncan.humphrey@era.co.uk
& leigh.polding@era.co.uk
ABSTRACT
To cope with the demand for increased product capacity in the petrochemical industry traditional carbon
steel (CS) is being replaced with high strength low alloy (HSLA) steel as material for plant construction.
The enhanced material properties come at a price in that some quenched and tempered (Q&T) variations
of HSLA steel cannot be readily subjected to post weld heat treatment (PWHT). While PWHT is not
mandatory, under the ASME Sec. VIII Div. I, for some of the HSLA steel grades in Q&T condition,
these grades could potentially suffer significant loss of strength and toughness should the PWHT be
carried out incorrectly. Examples are presented of severe stress corrosion cracking (SCC) in the heat
affected zone (HAZ) of welds in Q&T HSLA steel which had not had PWHT and were exposed in
service to boiler feed water (BFW) in temperature regime of 250ºC. These examples show that the high
residual stresses in thick section welds and the presence of susceptible material of high hardness in the
HAZ provide a key contributing factor for SCC exposed to hot BFW of high purity. It is demonstrated
that reliance on procedures such as “temper-bead welding” is no guarantee of trouble free welds as they
are prone to human error and should be avoided in large structures. Proper PWHT must be carried out in
the same way as it is done in Carbon Steel thick section welds to reduce the residual stresses and temper
the susceptible material in the HAZ to successfully minimise the potential for SCC. The choice of
appropriate HSLA steel grade for the application without significant concentration of strong carbide
formers and/or furnished in unhardened condition is demonstrated for thick wall vessels exposed to high
temperature BFW.
Abbreviations:
all volatile treatment (AVT); boiler feed water (BFW); carbon equivalent (CE), Carbon Steel (CS); coarse grained
zone (CGZ); ductile to brittle transition temperature (DBTT); Energy dispersive X-Ray analysis (EDX); heat
affected zone (HAZ); high strength low alloy (HSLA); internal diameter (ID); parent metal (PM); post weld heat
treatment (PWHT); pressure vessel reactor (PVR); quenched and tempered (Q&T); root cause assessment (RCA);
scanning electron microscope (SEM); stress corrosion cracking (SCC) ; ultimate tensile strength (UTS); Vickers
hardness (HV); welding electrode (WE); weld metal (WM); yield strength (YS)

Paper No. 14169 15th
Middle East Corrosion Conference & Exhibition Page 2 of 13
February 2-5, 2014
Manama, Kingdom of Bahrain
1. INTRODUCTION
Stress corrosion cracking, (SCC), of metal alloys accounts for approximately 25% of all failures of
petrochemical equipment, Ref. 1
. The occurrence of SCC depends on the simultaneous achievement of
three requirements; a susceptible material, sufficient tensile stress to induce SCC and an environment
that causes SCC for that material. SCC is not an inevitable process, and for most alloys in most
environments, it will not occur. SCC is relatively rare, since all the above three conditions have to be
met simultaneously. Careful control to lessen the effect of one or more of the three conditions will tend
to minimise the potential for SCC of metal alloys.
Welds in ferritic/bainitic steels are particularly prone to SCC in some environments due to the
possibility of enhanced material susceptibility in the heat affected zone (HAZ) due to the formation of
hard phases of martensite and/or bainite and, the presence of high residual stress, as residual stresses as
high as the yield strength (YS) of the parent metal (PM) are well documented in thick section welds, Ref.
2
. Acceptable hardness in the HAZ and low residual stresses in ferritic/bainitic steels welds may be
achieved through: a) careful control of PM composition and use of proper welding procedures
(adherence to proper preheat, inter-pass temperatures, and reliance on techniques such as temper bead
welding); and/or b) use of a suitable PWHT.
Carbon manganese steels with relatively low carbon equivalent (CE), normally, show good weldability
with acceptable HAZ hardness. However, to alleviate the effect of the residual stress in thick section
welds in CS, a PWHT at 6000
C to 6500
C is recommended. For some applications, like caustic service,
all CS welds must be PWHT, Ref.3
. Likewise, for pressure vessel applications all thick section welds in
CS must be PWHT in accordance with ASME (BPVC) Section VIII Divs. I and II, Ref. 4
. The PWHT can
be carried out safely without undue detrimental effect on the mechanical and impact properties of CS,
since the temperature range of the PWHT is considerably below the lower critical transition temperature
of CS, which is the end point transformation of austenite to pro-eutectoid ferrite and the decomposition
product in the microstructure. Weldments in CS pressure vessels designed for boiler feed water (BFW)
applications are mandatorily subjected to routine PWHT.
The development and use of high strength low alloy (HSLA) over conventional carbon steels (CS) has
been driven by the need to reduce costs with the higher strength enabling thinner and lighter structures
to be erected. Whilst, unhardened low alloy steels (micro-alloyed HSLA) are preferred for structural
applications, quenched and tempered (Q&T) HSLA steels are finding increasing use for large capacity
pressure vessel erection in the petrochemical industry. The latter Q&T type HSLA steels are quenched
and subsequently tempered at temperatures of approximately 6000
C to attain the optimum mechanical
and impact strength properties. The obvious drawback of alloying a steel is the reduction of weldability,
which calls for stringent welding control during fabrication. However, from an SCC viewpoint the major
disadvantage of steels which have been Q&T is that the PWHT necessary to relieve the high residual
stresses imparted during welding, cannot be routinely carried out as in the case of CS, as it may
significantly reduce the mechanical and impact properties of the parent metal if carried out incorrectly,
Ref. 5
. The detrimental effect of PWHT on mechanical and impact properties of some HSLA steel grades
is presented in Figure 1, Ref. 5
. To preserve the strength and impact property requirements of Q&T steels,
the PWHT temperature must be kept safely below the tempering temperature specified for the grade of
HSLA steel.
From the time-temperature property relationship Larson-Miller parameter, which is equally applicable to
both the tempering process and the relief of residual stress, the PWHT parameter should be chosen to
not to alter the previous tempering effect. The risk of extending the heat treatment time or raising the

February 2-5, 2014
temperature too close to the tempering conditions or exceeding them exists. Because of these concerns,
ASTM Section VIII code does not specify PWHT as mandatory for the Q&T HSLA steel grades to
prevent the possibility of irreversible material damage in case the PWHT is carried out incorrectly.
Since PWHT cannot be routinely applied, the fabricator of pressure vessels from Q&T HSLA steels is
left with the option of applying specialist welding techniques. Temper bead welding with adequate
preheat and inter-pass procedures may be used to counteract the effect of formation of susceptible
material in the HAZ and the high residual stress required for the occurrence of SCC. Further remedial
measures to minimise the potential for SCC in BFW applications would be to limit the impact of the
environment, through use of all volatile-treated BFW (BFW of high purity). However, there is evidence
that even very pure BFW may not be enough to arrest the SCC process if very high residual stresses are
present in the weldments, Ref. 6
. Figure 2, illustrates, that SCC can occur even in high purity steam when
the residual stresses approach the yield strength of the material, Ref. 6
.
2. INVESTIGATION
2.1. Background
ERA Technology Ltd was invited to investigate apparent cracking in welds of pressure vessel reactors
(PVR) in BFW applications. The PVR normally operate in the temperature range of 2000
C to 2500
C and
pressures of between 3 to 6 MPaG. The BFW was of high purity all volatile treatment.
Widespread cracking in the circumferential tubesheet to shell welds have been reported in 12 PVR units
over a period from 2004 to 2012. The oldest units were installed in 2004 and the most recent in 2010.
Time of initiation of cracks ranged from 1–3 years. Two types of cracks were discovered, namely
transverse cracks and circumferential (longitudinal) cracks.
The transverse cracks, though undesirable, did not pose an imminent integrity concern as they arrested at
the parent metal interface. However, the integrity of the PVR was considered to be seriously
compromised with subsequent discovery of the longitudinal cracks. If allowed to grow to a critical size
in the wall of the PVR, the longitudinal cracks could potentially cause a sudden failure in service with
all its dire consequences.
ERA conducted a root cause assessment (RCA) of the cracking of the PVR.
2.2. Materials of Construction
The PVR were constructed from SA543B Cl.1 Q&T HSLA steel as presented in Table 1. The welds in
the PVR were not subjected to PWHT. The chemical composition of the materials of construction is
presented in Table 2, Ref. 7
.

February 2-5, 2014
Table 1: Material Specifications
Material Specification
Shell Side
Cylinder
SA543 B CL.1 (*
1)
*
1 – Quenched and Tempered Steel
Nozzle Flanges SA508 Gr.4N CL.1
Tube Side
Cylinder SA543 B CL.1 (*
1) + SA240 T.304L CLAD
Nozzle Flanges SA543 B CL.1 + 308L WELD OVERLAY
Tubes DUPLEX 2205 (SA789, S31803) SEAMLESS
Tube Sheet
SA508 Gr.4n CL.1 (*
1) CLADDED with 309L + 308L
OVERLAY
WE Shell Side (120 mm) E11016-G (LB-88LT Kobe Steel ltd.)
Table 2: Chemical composition of specified materials
Type C% Mn% S% P% Si% Ni% Cr% Mo% V%
SA543 B CL.1
0.20
max.
0.40
0.020
max.
0.020
max.
0.15-
0.40
2.25-
4.00
1.00-
1.90
0.2-
0.65
0.03
max.
SA508 Gr.4N CL.1
0.23
max.
0.20-
0.40
0.020
max.
0.020
max.
0.15-
0.40
2.8-
3.9
1.5-
2.00
0.40-
0.60
0.03
max.
E11016-G WE 0.09 1.7
0.03
max.
0.03
max.
0.40 2.2 3.0 1.5 0.2
Duplex 2205
SA789
0.03
max.
2.00
max.
0.020
max.
0.023
max.
1.00
max.
4.50-
6.50
22-23
3.00-
3.50
-
The SA543B Cl.1 is a type 1¼Cr-½Mo alloy steel with significant additions of nickel of between 2% -
4% by weight, Ref. 7
. Addition of nickel is apparently beneficial in reducing proneness to SCC attack. A
YP of minimum 585 MPa and a UTS in the range of 725MPa to 860MPa is specified for SA543B Cl.1
as per ASTM SA543 material specification, Ref. 7
.
3. RESULTS
3.1. Macro Examination
Figures 3 and 4 present macros of the material recovered from the cracked weld, which contains the
longitudinal cracks. The cracks initiate on the ID surface in contact with BFW and propagate along the
grain boundaries of the coarse-grained zone (CGH) in the HAZ. The intergranular path in the CGZ is
clearly shown in Figure 5.
3.2. Microstructural Evaluation
Typical microstructures of the HAZ and the PM are presented in Figure 5. The microstructure in the
HAZ appears to be that of untempered martensitic and/or bainitic phases. In contrast, the microstructure
of the PM is that of tempered martensite.

February 2-5, 2014
3.3. Scanning Electron Microscopy – SEM
Figure 6 shows the SEM image of the tip of the crack that propagated in the grain boundaries in the
HAZ. The crack appears to be filled with a scale deposit. EDX chemical analysis of the scale confirmed
presence of iron oxide free from any contamination. The PM was found to be in conformance with
ASTM SA543B Cl1. material specification.
Table: 3 Chemical analysis of the scale deposit inside the crack and the PM
Spectrum O Si Cr Mn Fe Ni Mo
Scale Deposit 27.4 0.0 0.0 0.2 71.5 0.7 0.0
Scale Deposit 27.0 0.2 0.1 0.7 69.2 2.1 0.7
PM 0.4 0.3 1.2 0.7 94.5 2.6 0.4
3.4. Vickers Hardness testing – HV10
Figure 7 shows the relative size of hardness indentations in the HAZ and the WM. The relatively higher
hardness due to smaller indentations in the HAZ is evident.
The laboratory Vickers micro-hardness and the Vickers field hardness HV10 are presented in Tables 4
and 5, respectively. Abnormally high hardness especially in the HAZ was measured in the laboratory
and in the field.
Table 4: Vickers Hardness Measurements – Laboratory micro-hardness
Position
Laboratory
(Micro HV) UTS (MPa) *1
HAZ next to crack 479 1550
HAZ opposite side 457 1480
Weld Metal 335 1079
Base Metal 258 831
Note: SA543B – (UTS – 725 to 860 MPa); E11016-G (UTS – 960 MPa). (*1) – UTS estimated from HV using ASTM E140,
Ref.8
Table 5: Vickers Hardness Measurements – Field hardness testing HV10
Position
Field hardness
HV10 UTS (MPa) *1
HAZ next to crack 391 1259
Weld Metal 315 1015
Base Metal 243 821
Note: SA543B – (UTS – 725 to 860 MPa); E11016-G (UTS – 960 MPa). (*1) – UTS estimated from HV using ASTM E140,
Ref.8

February 2-5, 2014
3.5. PWHT Simulation Tests
Heat treatments to simulate the effect of a PWHT were carried out on sections of material affected by
cracking. The heat treatments were carried out at 6000
C, 6500
C and 7000
C with a soak time of 30
minutes. Normal heating and cooling ramp rates were used during the simulated PWHT. The heat-
treated samples were subjected to a Vickers hardness testing and mechanical testing.
The results of the simulated PWHT on Vickers hardness HV10 and UTS of the SA543B material are
presented in Tables 9 & 10 and Figures 9 & 10, respectively.
Table 9: Effect of PWHT and Vickers hardness HV10
PWHT (0
C) Parent metal - PM Heat Affected Zone-HAZ Weld Metal - WM
HV UTS (MPa) HV UTS (MPa) HV UTS (MPa)
As Received 253 815 420 1350 292 940
600 250 800 322 1036 272 873
650 231 744 311 999 253 815
700 223 714 281 903 267 856
Table 10: Effect of PWHT on the UTS
Base Metal – SA543B Cl.1
Temp. (0
C) Y.S. (MPa) UTS (MPa) HV YS/UTS % Elongation
Ambient 675.5 784 245 0.86 21
600 641 785 245 0.82 24.5
650 617.5 761.5 236 0.81 23.2
Weld Metal
Temp. (0
C) Y.S. (MPa) UTS (MPa) HV YS/UTS % Elongation
Ambient 835 880 274 0.95 13.9
600 786.5 859 268 0.92 15.75
650 722.5 801.5 250 0.90 24.75
Note: SA543B – (UTS – 725 to 860 MPa); E11016-G (UTS – 960 MPa). Minimum elongation 14-16%
The beneficial effect of reducing the hardness and the UTS to acceptable levels is evident at tempering
temperatures of 6000
C to 6500
C.
4. DISCUSSION
All the experimental evidence gathered in this investigation favours SCC as the mechanism responsible
for the cracking in the SA543B circumferential welds. The high level of hardness measured in the HAZ
of the order of 400HV to 470HV supported by the microstructure of untempered martensite, suggest
presence of “susceptible” material in the HAZ. Existence of the hard zones in the HAZ and the fact that
no PWHT was carried out during fabrication confirms that high residual stress approaching the YP of
the PM of the order of 600 MPa would be present, further favouring the SCC mechanism. The actual
service stress of the order of 800 MPa (200MPa operating stress and 600MPa residual stress) is not an

February 2-5, 2014
overestimate. The combination of the “susceptible” material and the high service stress could lead to
SCC in BFW or high purity as presented in Figure 2, Ref. 6
. Furthermore, the intergranular appearance of
the crack in the CGZ of the HAZ is direct advocate of the SCC mechanism at work.
Presence of “susceptible” material in the HAZ in the absence of PWHT stage confirms that improper
welding procedures were employed during fabrication. Certainly, trouble free weld with acceptable level
of hardness in the HAZ and acceptable level of residual stress were not achieved during fabrication. It is
unlikely that temper bead welding was used during construction of the PVR. The fabricators also did not
follow proper quality control procedures of in situ hardness testing which would have picked up the
excessive hardness in the HAZ.
Chemical analysis of the corrosion product in the base of the intergranular cracks revealed no presence
of contaminants confirming good BFW quality free of contamination.
Examples of SCC in BFW of Q&T HSLA steels that did not receive a PWHT are available in the
literature, Ref. 9, Ref. 10
. Reference 10 discusses cracking of weldments in BFW in PVR made of the same
SA543B material. Similar levels of Vickers hardness in the HAZ of the order of 460HV was measured
in this PVR.
5. CONCLUSIONS
1. The root cause of longitudinal cracking in the circumferential welds was confirmed to be due to
SCC.
2. Evidence suggests that weldments were not temper bead welded. In the absence of a proper
PWHT procedure, the “susceptible” material in the HAZ and the presence of high residual stress
allowed SCC to occur even in high purity BFW. The actual service stress exceeded the yield stress
of the PM and allowed SCC to take place.
3. The BFW was found to be of high purity free of contamination eliminating possibility of
contribution from environmentally assisted cracking.
4. This is a case of improper material selection for the application. A material that could be routinely
PWHT i.e. not in Q&T but normalised condition should have been considered to allow a PWHT
procedure whilst avoiding harmful effects to the material properties.
5. Sole reliance on high purity BFW (high purity environment) cannot be depended upon to suppress
the process of SCC if the material is in “susceptible” condition with high residual stresses.
6. RECOMMENDATIONS
1. The most dependable way of reducing the potential of SCC in BFW is by performing a PWHT to
temper out any “susceptible” material in the HAZ and thermally stress relief the residual stresses
to acceptable level. This should be particularly done to HSLA steels, as their weldability is
impaired due to addition of alloying elements. Specialist welding techniques cannot be solely
relied on, as they are prone to human error. A proper PWHT is the most dependable way to
provide trouble free welds. In Q&T HSLA steels which contain significant additions of strong

February 2-5, 2014
carbide formers such as SA543B, SA517F, A387-22 materials a PWHT is not recommended
because it may degrade materials properties (mechanical and impact) due to further tempering,
Figure 1, Ref. 5
. However, HSLA steels such as the SA302B and SA533B grades, which do not rely
on addition of strong carbide formers and are supplied in normalised condition can be routinely
subjected to PWHT without significant detrimental effect to mechanical and impact properties,
Figure 1, Ref. 5
. The SA302A or the SA533B steel would be a better choice for the PVR application
in BFW conditions where a PWHT can be carried out safely.
2. The designers and fabricators should consider potential use of stainless steel cladding and/or
overlay to provide direct isolation between HSLA material and the BFW. This design was used on
the tube side of the PVR as shown in Table 1.
3. The weldments in the SA543B material may have been subjected to a PWHT at a safe temperature
of 5500
C. This would have tempered the “susceptible” zones and reduced the residual peak
stresses to much lower level. The effect would not be as marked as a PWHT at 6000
C but
nonetheless the potential for SCC in BFW of the weldments would have been reduced
significantly.
7. REFERENCES
1. Graham R. Lobley, “Stress Corrosion Cracking: Case Studies in refinery Equipment”, The 6th
Saudi Engineering Conference, KFUPM, Dhahran, December 2002.
2. R.J Zhou, A.W. Pense, M.L. Basehore and D.H. Lyons, “Study of Residual Stresses in Pressure
Vessel Steels,” Welding Research Council, Bulletin 302, February 1985.
3. NACE SP0403-2008, “Standard Practice – Avoiding Caustic Stress Corrosion Cracking of Carbon
Steel Refinery Equipment and Piping”.
4. ASME Boiler and Pressure Vessel Code (BPVC) Section VIII, “Rules for Construction of Pressure
Vessels,” Division I and Division II (Alternative Rules), 2010.
5. R.D. Stout, “Post Heat Treatment of Pressure Vessels,” Welding Research Council, Bulletin 302,
February 1985.
6. D.A. Rosario, R. Viswanathan, C.H. Wells, and G.J. Licina, “Stress Corrosion Cracking of Steam
Turbine Rotors”, CORROSION–Vol. 54, No. 7, 1997.
7. ASTM SA543-04, “Specification for Pressure Vessel Plates, Alloy Steel, Quenched and
Tempered, Nickel-Chromium-Molybdenum”.
8. ASTM E140 – 07, “Standard Hardness Conversion Tables for Metals Relationship Among Brinell
Hardness, Vickers Hardness, Rockwell Hardness, Superficial Hardness, Knoop Hardness, and
Scleroscope Hardness.”
9. WU Yun-Xiang, “Investigation on Stress Corrosion Cracking Behaviour of Weld Joint of Low
Alloy Steel with High Strength in High Temperature Boiler Feed Water”, Pressure Vessel
Technology, 2010-06.
10. “Environmental Stress Cracking of Shell and Tube Heat Exchanger in High Temperature Boiler
Feed Water Service”, Jim Moen and Glenn Roemer, IPEIA Conference, February 2006.

February 2-5, 2014
8. FIGURES
Figure 1: Effect of PWHT on UTS and DBTT of Q&T HSLA Steels
Figure 2: Crack initiation under operating stress (including the residual stress) vs. BFW purity.

February 2-5, 2014
Figure 3: Longitudinal crack initiated on the ID & propagated along the HAZ
Figure 4: Longitudinal crack initiated on the ID and propagated along the CGZ of the HAZ
WM
PM
HAZ
crack
W
M
WM
WM

February 2-5, 2014
The crack initiated from corrosion pit (arrows) and propagated in the coarse grain
region of the HAZ. Crack propagation is intergranular. (Mag. X50)
Figure 5: Longitudinal crack that initiated on the ID and propagated along the HAZ
A
Untempered bainite/martensite in the
HAZ (more of a needle-like structure).
Mag. X100
B
Tempered bainite/martensite in the base
metal which was Q&T. Mag. X100
Figure 6: Microstructure of the HAZ and the PM, respectively.

February 2-5, 2014
Crack is filled with corrosion product free of any contaminants (longitudinal crack tip).
Figure 7: A crack filled with iron oxide corrosion product free of contaminants
HAZ is a region of relatively high hardness as indicated by smaller indentations.
Figure 8: Hardness comparison between the HAZ and the WM

February 2-5, 2014
Figure 9: Effect of PWHT on Vickers hardness HV10
Figure 10: Effect of PWHT on the UTS

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