(Sungki Choi)
1†
(Junsang Lee)
1†
(Jae-Yik Lee)
2
(Seung-Kyun Kang)
1
(Young-Cheon Kim)
3*
(Seung-Joon Lee)
4**
(Dongil Kwon)
1
Copyright © 2021 The Korean Institute of Metals and Materials
Key words(Korean)
low transformation temperature, LTT consumable, flux-cored arc welding, mechanical properties, residual stress, indentation
1. Introduction
The integrity of welds in various industrial processes such as building construction,
ships, automobiles, mining is crucial to the overall safety of the structure. In particular,
tensile residual stresses introduced by the welding process can affect in the entire
stress state of the weld. Welding residual stress, which is inevitable during cooling
after welding, can lead to stress-corrosion cracking, distortion, and reduction in
the fatigue strength of the weldment [1-5]. Several thermal and mechanical techniques can reduce the tensile residual stresses
in the joint via introducing compressive residual stresses. These include post-weld
heat treatment (PWHT), shot peening and thermal tensioning techniques [6-11]. However, these methods have limitations such as technical complexity, high cost
and time-consuming process.
Many studies have been carried out on ways to reduce residual stresses and improve
the mechanical properties of weldments using novel welding consumables that can control
the processing temperature, preventing the phase transformation. Alberry and Jones
reported that a reduction in processing temperature can drive the suppression of residual
stresses [12]. Low-transformation-temperature (hereinafter LTT) welding consumables that exploit
this phenomenon have been developed in recent years [13-17]. Conventional consumables have a transformation temperature of about 500 °C, but
in an LTT welding material with Fe-10Ni-10Cr (wt.%) developed by Ohta et al., phase transformation occurred at 180 °C [14]. This welding consumable could reduce tensile residual stresses or generate compressive
residual stresses. Kromm and Kannengiesser reported that the martensite start temperature
(Ms) decreased with increasing Ni content due to the higher austenite (γ) stability
[18]. Murakawa et al. reported that compressive residual stress increased as Ms decreases up to approximately
200 °C [6]. Many researchers have found not only a reduction in residual stress but also an
improvement in fatigue strength [17,19-22] and the suppression of distortions [23-26] in weldments when using LTT welding materials. By tuning the temperature so that
the transformation from γ to α'-martensite occurs at a lower temperature, the volume
expansion accompanied by strain can compensate thermal shrinkage that occurs with
cooling after welding [14,27,28].
Figure 1(a) and (b) show the variation in strain and stress in a conventional consumable and a LTT consumable
during cooling, respectively. The two welding materials differ in the conventional
consumable, and shrinkage is a dominant factor after the expansion due to phase transformation
from γ to α' phase at 500 °C (Figure 1(a)). This results in a tensile residual stress at room temperature (Figure 1(b)). On the other hand, the expansion from 200 °C to room temperature is dominant in
the weld using the LTT consumable, because the phase transformation occurs at a relatively
low temperature of 200 °C (Figure 1(a)), leading to compressive residual stress (Figure 1(b)).
Automating the welding process is essential for efficiency in an industry that requires
in-field and large-scale applicability. Flux-cored arc welding (FCAW) is preferred,
to ensure weld quality and to reduce labor costs. This automatic or semi-automatic
technique is widely used in shipbuilding, manufacturing and repair welding, because
it provides high productivity and industrial applicability compared to other methods
such as shield metal arc welding (SMAW) and friction stir welding. As noted above, using LTT consumables can induce compressive residual stress, so the
PWHT process for relieving residual stresses can be eliminated and productivity can
be increased by utilizing LTT with FCAW.
In this study, the effect of a low-temperature-transformation (LTT) consumable using
FCAW on ASTM A516-70N carbon steel was analyzed quantitatively with respect to microstructure,
mechanical properties and residual stress. Electron backscattered diffractometer and
field-emission scanning electron microscope were used to characterize the microstructure
of the LTT weld metal compared to a conventional weld metal. Micro Vickers hardness
and impact toughness tests evaluated the effect of changes in microstructure on mechanical
properties. The LTT consumable showed a strong advantage in welding residual stress
over conventional weld coupons. The effect of the welding consumable was confirmed
to be superior to the residual stress relaxation effect of PWHT.
2. Experimental Procedure
2.1 Materials and weld fabrication
A schematic layout of the welded coupons is shown in Fig 2. The chemical composition carbon equivalent (Ceq) of welding consumables and base metal is listed in Table 1, and their mechanical properties are presented in Table 2. The carbon equivalent of the welding materials can be calculated with Eq. (1) [29].
Welding plates of ASTM A516-70N carbon steel were prepared for a butt weld size of
300 mmw × 200 mml × 12 mmt (two plates 150 mm wide were joined), which had a 60 ° V-groove geometry. Welding
consumables with 1.2 mm diameter, AWS E71T-1C [30] were used for the conventional welding coupon and Mn-based LTT material for the
LTT welding coupon, were welded in four passes. Flux-cord arc welding (FCAW) was performed
with different welding conditions and two welding consumables, as shown in Table 3. Three FCAW coupons were fabricated, two with conventional consumables and one with
a Mn-based LTT consumable. One of the conventional welding coupons was postweld heat-treated
as specified in ASME section III NX-4622 [31]. PWHT was performed at 635±15 °C for 12 hours after welding, according to ASME section
III NX-4622 [31], for an additional conventional welding coupon.
2.2 Microstructure characterization and mechanical properties
Cross-sectional samples of 10 mmw × 10 mml × 10 mmt containing the fusion line were machined at the center of the welding coupons from
the LTT weld and the conventional weld without PWHT. The samples were mechanically
polished with 400-2400 grit paper and then a suspension with 1-μm-diamond particles
and 0.25-μm-colloidal silica. X-ray diffractometer (D8-Advance, Bruker, Karlsruhe,
Germany) was used to detect the phases of the weld metal with the two welding consumables
using a Cu-Kα target, and a scanning range and speed of 20 ° to 100 ° and 3 °/min,
respectively.
The sample used for X-ray diffraction measurement was chemically etched with 5% nital
for microstructural observation. A field-emission scanning electron microscope (FESEM,
FEI Quanta 250 FEG, Hillsboro, OR, USA) was used to identify the microstructures in
the weld metal. Additional surface preparation was made for electron backscattered
diffractometer (EBSD; EDAX-Hikari) analysis by electropolishing in a mixed solution
of 90% ethanol (C2H5OH) and 10% perchloric acid (HClO4) at 31 V for 30 s. EBSD analysis was operated at an acceleration voltage of 15 kV
and a working distance of 15 mm, and a step size of 2 μm. The phase fraction in the
weld metal with the LTT consumable was analyzed quantitatively under this condition
with a scan area of 200 μm × 200 μm and a step size of 0.5 μm.
Microhardness testing was performed by Micro-AIS (Frontics Inc., Seoul, Korea) with
a Vickers indenter at fixed maximum load of 0.1 kgf. The hardness distribution over
the cross-section was measured from the weld center toward the heat-affected zone
(hereinafter HAZ). Two diagonal lengths of the indent were directly measured by optical
microscope. Three sub-sized Charpy V-notch specimens (5 mm × 10 mm × 55 mm, 45 °)
were extracted from the center of the weld line following ASME section II SA370 [32] in the direction transverse from the welding plates, as shown in Fig 2. The tests were conducted at room temperature
2.3 Residual stress measurement
The residual stress of the LTT welding coupon was measured using X-ray diffraction
prior to indentation testing to compare the results of the two tests. An Xstress3000
(X-ray Stress Analyzer) was used with Cr-Kα radiation (λ = 0.22911 nm), 2θ = 156.4
°, diffracting plane {211} from base metal and weld metal. From Bragg’s Law, the sin2ψ
method was used for measurements. To measure the changes in lattice spacing, the operating
voltage was 30 kV and the current 6.7 mA. The measurements were carried out from the
weld center through the red line (Fig 2) perpendicular to the welding line at 2 mm intervals.
Instrumented indentation testing (IIT) was performed for residual stress measurement
on the top surface of the three welding coupons, carefully polished up to 2000 grit
sandpaper following ISO 14577-1 [33]. A commercial indentation system, AIS 3000 (Frontics Inc., Korea) with a load resolution
of 0.05 N and depth resolution of 0.1 μm was used. A Vickers indenter with an apex
angle of 136 ° [34] was used, and all the tests were controlled with a maximum load of 490 N (50 kgf)
and an indentation loading rate of 0.3 mm/min. The measurements were performed at
4 mm intervals on the red line, as well where the X-ray diffraction was performed.
The normality between the indenter and the specimen surface was carefully controlled
not to exceed 1 degree.
The model developed by Lee and Kwon was used for residual stress evaluation [35]. The indentation load-depth curve is shifted depending on the sign and magnitude
of the residual stress relative to the curve in the stress-free state. The relation
between the force difference as a result of residual stress and the deviatoric stress
of the indenting direction was formulated with the projected area as Eq. (2).
where L0 and Ls are the loads in the stress-free and stressed state, respectively, and As means indentation contact area.
3. Results and Discussion
3.1 Microstructure characterization
The microstructure of the LTT weld metal was characterized by X-ray diffraction, FE-SEM
and EBSD, and then compared to the conventional weld metal without PWHT. The X-ray
diffraction peaks of the two weld metals (except the conventional weld metal with
PWHT, since phase transformation does not generally occur with PWHT) are compared
in Fig 3. Both weldments had phases suspected to be ferrite or bainite/martensite showing
a bcc structure in the weld metal. No retained austenite phase was detected by X-ray
diffraction analysis; it rarely forms when the carbon content is less than 0.4 wt%
[36], as was the cases in these two weldments.
Figure 4 shows the weld metal images obtained from FESEM at ×800, ×1500 and ×3000 magnitude.
Ferrite phases (acicular, allotriomorphic, and Widmanstätten) were observed in the
weld metal with the conventional consumable, as previously reported [37,38]. And, in case of the weldment with the LTT, mixed structures of ferrite, bainite,
and martensite were detected. These results correspond to those of the X-ray diffraction
peak analyses.
For more depth-in-analysis of phases in the conventional and LTT welding, the effective
grain size and phase fraction of the weld metals was measured by EBSD analysis. Figure 5 shows the IQ map of the weld metals. The average effective grain size can be calculated
by defining a misorientation angle larger than 15 ° [39,40]. The average effective grain sizes were 6.70 μm and 5.49 μm for the conventional
and LTT consumable, respectively. It was estimated that the maximum temperature during
welding of the conventional consumable was higher than the LTT consumable, and LTT’s
finer effective grain size affected the hardness and impact toughness.
GOS map and Gaussian fitting graph from IQ analysis using EBSD were performed for
the LTT consumable, since the phases were difficult to distinguish accurately using
only XRD and FE-SEM. The fraction of martensite phase in the weld metal of the Mn-based
LTT consumable was distinguished by analyzing the grain orientation spread (GOS) map
and IQ map from the EBSD results. Low-angle and high-angle boundaries were divided
by ranges into 2 ° < θ < 15 ° and θ > 15 ° [41] and are indicated in Fig 6(a) as red and blue lines, respectively. Ferrite can be identified in the GOS map, having
a tolerance angle of 15 ° and a misorientation angle less than 5 ° [42]; it is represented by the yellow region in Fig 6(a). After eliminating the ferrite phase, bainite and martensite phases can be distinguished
by Gaussian fitting the graph from IQ analysis [36]. In general, the martensite phase has larger lattice imperfections than the bainite
phase, and thus has lower Kikuchi pattern intensity, which in turn results in lower
contrast than the bainite phase [36].
Since the IQ value is affected by many operating factors including image condition,
it should be normalized to minimize error effects. The normalized IQ value can be
represented by Eq. (3) [43]:
where IQN is the normalized IQ value and IQinitial is the value obtained directly from experiment; IQMax and IQMin are the maximum and minimum IQ values.
Figure 6(b), showing the fraction of the normalized IQ value, can be separated into two distributions
by Gaussian fitting. The red distribution of IQ values with relatively low contrast
represents the martensite phase, and the other distribution of IQ values with high
contrast on the right can be identified as the bainite phase.
The phase fractions of ferrite, austenite, martensite and bainite were 50.5%, 0.2%,
40.2% and 9.1%, respectively. The fraction of martensite is considered a major factor
affecting the mechanical properties (strength, hardness and toughness) and residual
stresses.
3.2 Micro Vickers hardness and impact toughness
Micro Vickers hardness was evaluated at 1 mm intervals from the weld center to the
base metal in the transverse direction for a cross section about 1 mm from the top
surfaces of the welding coupons. Similar levels of hardness in the HAZ and base metal
were measured as shown in Fig 7, whereas average hardness values in the weld metal showed a difference, with 216.8
Hv in the conventional consumable and 314.5 Hv in the LTT consumable. Martensite in
the LTT weld metal is reasonably considered a major cause of the higher hardness distribution.
The average values of impact absorption energy for the conventional weld were 174.3±13
J and 53.3±3 J for the LTT weld. The large difference in impact absorption energy
can be explained by microstructure: the greater fraction of acicular ferrite in the
material increases impact toughness [44] and finer effective grain size decreases impact toughness [45].
Martensite or martensite-austenite is the main factor determining absorption energy
at room temperature, because cracks can initiate due to stress concentrations around
a hard phase. The absorption energy can be lower with the higher phase fractions [45,46]. With a conventional consumable, acicular ferrite is distributed and martensite
is not present in the weld metal. However, when the LTT consumable was used, a considerable
martensite phase was distributed. The presence of these phases influences the difference
in impact toughness between the two weldments.
3.3 Residual stress distribution
Figure 8 shows the results of residual stress measurements for three welded coupons. The heat-treated
welding coupon with the conventional consumable is included to compare the residual
stress distribution of the LTT weld with the distribution after relaxation by PWHT.
The stress-relaxing effect was also compared using two coupons welded with a conventional
consumable with and without PWHT.
In addition, for the samples welded with the LTT consumable, the residual stress was
evaluated using X-ray diffraction and indentation testing to compare the accuracy
of the results. The residual stress obtained from indentation is an average value
that reflects the stress in the longitudinal and transverse directions on the welding
line. In X-ray diffraction, the residual stresses were measured in the longitudinal
and transverse directions and the average value was calculated for comparison with
the indentation results.
The maximum average residual stress was 339 MPa in the weld metal of the conventional
consumable without PWHT. This tensile residual stress is a typical stress distribution
that occurs when the thermal shrinkage during the cooling process after welding is
a dominant factor. The stress in the weld metal of the heat-treated coupon was about
200 MPa, clearly showing the stress-relieving effect of PWHT. The effect was greater
in the HAZ: the stress changed from 257 MPa to -29 MPa. Comparison of the two coupons
with and without PWHT emphasize the importance of PWHT on residual stress, especially
for the weld metal and HAZ.
The residual stress in the coupon welded with the LTT consumable showed a similar
distribution in the results of X-ray diffraction and indentation testing, within a
range of 53 MPa. The agreement of the two methods clearly confirms the effect of the
LTT consumable. The difference is believed to arise from the difference in measurement
depth and area. 50 kgf indentation for a depth less than 100 μm senses the residual
stress at depths up to 1 mm, quite different from the X-ray diffraction penetration
depth for carbon steel.
For the weldment welded with the LTT consumable, compressive residual stress was distributed
in the weld metal and HAZ. Strain in volume expansion due to martensitic transformation
at a relatively low temperature was a dominant factor. The martensite phase remaining
in the LTT weld metal, as confirmed by microstructural analysis, resulted in the compressive
residual stress.
The effect of residual stress on fatigue strength has been investigated by many researchers.
The presence of considerable tensile residual stress in welds with conventional consumables
with and without PWHT is expected to reduce the fatigue resistance of the joint. On
the other hand, it was confirmed that compressive residual stress increases fatigue
life [21,27]. Utilizing a Mn-based LTT consumable in the weld can improve fatigue life, like
other LTT consumables, and also eliminate the need for PWHT.
4. Conclusions
In this study, the effect of welding consumables (Mn-based LTT and conventional) and
PWHT for FCAW-welded ASTM A516-70N carbon steel was analyzed in terms of microstructure,
mechanical properties and residual stress. The conclusions are summarized as:
(1) The microstructures containing acicular ferrite, allotriomorphic ferrite, and
Widmanstätten ferrite were distributed in the weld metal with the conventional consumable.
In the Mn-based LTT consumable weld metal, ferrite, austenite, martensite, and bainite
were identified in fractions of 50.5%, 0.2%, 40.2% and 9.1%, respectively, by GOS
mapping and normalized EBSD image quality analysis.
(2) The average hardness in the weld metal was 216.8 HV in the conventional consumable
and 314.5 HV in the LTT consumable; likewise, the absorption energy from Charpy V-notch
testing was 174.3 J and 53.3 J, respectively. For the LTT consumable, the relatively
higher hardness and lower absorption energy than the conventional consumable are attributed
to the finer effective grain size and the martensite phase.
(3) The residual stress distribution with the LTT consumable was double-checked using
two methods, instrumented indentation and X-ray diffraction. The residual stress distributions
obtained with both methods showed a similar trend within a range of 53 MPa difference.
(4) Considerable tensile residual stresses were distributed in the coupons with the
conventional consumable. PWHT considerably decreased the magnitude of residual stress
in the weld metal and the HAZ, especially the HAZ, from 257 MPa to -29 MPa. Compressive
residual stress was identified in the weld metal and HAZ when the LTT consumable was
used, and this can be presented as the main advantage of the low-transformation-temperature
welding consumables.
Acknowledgements
This work was supported by POSCO and the Korea Institute of Energy Technology Evaluation
and Planning (KETEP) and the New Faculty Startup Fund from Seoul National University
for S.-K. K.
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Figures and Tables
Fig. 1.
Variation of strain and stress during cooling process: (a) strain and (b) stress [27].
Fig. 2.
Schematic layout of welding coupons from (a) top view, (b) side view and Charpy impact
test specimen.
Fig. 3.
Results of XRD in weld metal of conventional and LTT consumable.
Fig. 4.
FE-SEM Micrographs in weld metal of sample welded with (a) conventional and (b) LTT
consumables.
Fig. 5.
IQ map of sample welded with (a) conventional and (b) LTT consumables in weld metal.
Fig. 6.
(a) GOS map (Red and blue lines are low angle grain boundary and high
Fig. 7.
Results of micro hardness (0.1 kgf) distribution of weldment in the cross-section.
Fig. 8.
Averaged residual stresses in the longitudinal and transverse directions on the welding
line measured by using indentation and XRD in samples.
Table 1.
Chemical compositions and carbon equivalent (Ceq) of welding consumables and base material (wt-%)
|
|
C
|
Si
|
Mn
|
O
|
P
|
S
|
Fe
|
Ceq |
|
Conventional consumable
|
0.04
|
0.55
|
1.25
|
-
|
0.02
|
0.01
|
Bal.
|
0.25
|
|
Mn- LTT consumable
|
0.03
|
0.38
|
4.30
|
0.07
|
-
|
-
|
Bal.
|
0.75
|
|
Base metal (A516-70N)
|
0.31
|
0.15-0.40
|
0.85-1.20
|
-
|
0.03
|
0.03
|
Bal.
|
0.45-0.51
|
Table 2.
Tensile properties of welding consumables and base material
|
Materials
|
Yield strength (MPa)
|
Tensile strength (MPa)
|
Elongation (%)
|
|
Conventional consumable
|
520
|
580
|
23
|
|
Mn- LTT consumable
|
727
|
820
|
23
|
|
Base metal (A516-70N)
|
404
|
575
|
21
|
Table 3.
Welding conditions for welding consumables
|
|
Welding current (A)
|
Arc voltage (V)
|
Welding speed (cm/min)
|
Heat input (kJ/cm)
|
|
Conventional consumable
|
240
|
30
|
22
|
20
|
|
Mn- LTT consumable
|
260
|
32
|
48
|
10
|