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Comparative study on girth weld-induced residual stresses between austenitic and duplex stainless steel pipe welds
Applied Thermal Engineering 63 (2014) 140e150
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Comparative study on girth weld-induced residual stresses between austenitic and duplex stainless steel pipe welds Chin-Hyung Lee, Kyong-Ho Chang* Department of Civil and Environmental & Plant Engineering, Chung-Ang University, 84, Huksuk-ro, Dongjak-ku, Seoul 156-756, Republic of Korea
h i g h l i g h t s Residual stresses in austenitic and duplex stainless steel pipes were investigated. Comparative study on the girth weld-induced residual stresses was performed. Sequentially coupled 3-D thermo-mechanical FE analysis was conducted. Considerably different welding residual stress distributions could be demonstrated.
a r t i c l e i n f o
a b s t r a c t
Article history: Received 3 June 2013 Accepted 1 November 2013 Available online 13 November 2013
Duplex stainless steel pipes find increasing use as an alternative to austenitic stainless steel pipes, particularly where chloride or sulphide stress corrosion cracking is of primary concern, due to the excellent combination of strength and corrosion resistance. During welding, duplex stainless steel pipes do not create the same magnitude or distribution of weld-induced residual stresses as those in girthwelded austenitic stainless steel pipes due to the different physical and mechanical properties between them. In this work, a comparison of the residual stresses between girth-welded austenitic and duplex stainless steel pipes was performed utilizing sequentially coupled three-dimensional thermomechanical finite element analysis method to accurately predict temperature fields and residual stress states in pipe girth welds. The results have shown that girth-welded austenitic stainless steel pipes produce much higher axial and hoop residual stresses normalized by the yield stress at the weld and its vicinity in which the welding start/stop effects are more significant and they have wider regions subjected to tensile or compressive stresses adjacent to the weld area. Ó 2013 Elsevier Ltd. All rights reserved.
Keywords: Austenitic stainless steel pipe welds Duplex stainless steel pipe welds Welding residual stresses Three-dimensional finite element simulation Thermal and mechanical properties
1. Introduction Duplex stainless steels, with a microstructure comprised of nearly equal proportions of ferrite and austenite, combine the attractive properties of austenitic and ferritic steels: high tensile strength and fatigue strength, good toughness even at low temperatures, adequate formability and weldability and excellent resistance to stress corrosion cracking, pitting and general corrosion [1e3]. They, especially in the pipe form, find increasing use as an alternative to austenitic stainless steels, particularly where chloride or sulphide stress corrosion cracking is of primary concern, e.g., in the chemical, oil and gas, paper and pulp, marine and petrochemical industries and pollution control equipment. In
* Corresponding author. Tel.: þ82 2 820 5337; fax: þ82 2 823 5339. E-mail address: ifi[email protected] (K.-H. Chang). 1359-4311/$ e see front matter Ó 2013 Elsevier Ltd. All rights reserved. http://dx.doi.org/10.1016/j.applthermaleng.2013.11.001
practical situations, girth welding of duplex stainless steel pipes is often needed owing to the long geometry relative to the diameter and the wall-thickness. Welding is a crucial manufacturing process and widely used in industries to assemble various engineering and structural components. The advantage of welding as joining process includes high joint efficiency, simple set up and low fabrication cost. In welded structures, the inevitable existence of welding residual stresses is well known, which are produced as a result of plastic deformation caused by non-uniform thermal expansion and contraction imposed during welding process. The residual stresses are always self-equilibrated and the magnitude of tensile residual stresses within and near the weld area is great enough to have deleterious effects on the structural integrity of the welded joints, increasing the susceptibility to fatigue damage, stress corrosion cracking and brittle fracture [4]. These stresses when combined with the applied stresses can reduce the fatigue life, accelerate growth rates of pre-
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Nomenclature A c h k q_ q(r) Q1 Q2 r r0 T T_ V Vp dT
h r
arc current specific heat temperature-dependent heat transfer coefficient thermal conductivity rate of moving heat generation per unit volume heat flux distribution heat input from the welding arc energy induced by high temperature melt droplets radial coordinate with the origin at the arc center arc beam radius temperature rate of temperature change arc voltage considered weld pool volume temperature increment arc efficiency factor density
existing or service-induced defects in structural systems. Knowledge of the distribution and magnitude of the residual stresses, therefore, would be of big help for the production of an efficient and economic design and safety of the structures. However, accurate prediction of welding residual stresses is very difficult because of the complexity of welding process which includes localized heating, temperature dependence of material properties and moving heat source, etc. Nowadays, numerical modeling based on finite element (FE) method is used to predict welding residual stresses due to the expense and impracticalities of generating comprehensive structural performance data through experiments [5e13]. Numerical techniques have become an important part of most structural research communities, since they can be employed as a useful tool for analyzing the behavior of structures provided that suitable care is taken to ensure that the modeling is appropriate for the analysis [14]. Until now, a large number of FE simulations focusing on the circumferential welding of austenitic stainless steel pipes have been performed to predict the girth weld-induced residual stresses through the axisymmetric models [15e20] or the threedimensional (3-D) models [7,21e27] by using commercially available FE-codes such as ABAQUS, ANSYS and SYSWELD, etc. Thus, welding residual stresses in girth-welded stainless steel pipe components have been thoroughly investigated. As for the duplex stainless steel pipe welds, to the knowledge of the authors, very few works have been published on the FE analysis of welding residual stresses. Jin et al. [28] evaluated the axial and hoop residual stresses in a circumferentially butt-welded 2205 (EN 1.4462) duplex stainless steel pipe through the numerical simulation based on the nonlinear thermo-mechanical FE analysis. Nevertheless, their work was confined to the axisymmetric model which was not capable of predicting the 3-D effects induced by the girth welding process. Díaz et al. [2] conducted thermo-mechanical FE analyses in order to compare weld-induced distortion modes and magnitudes between austenitic and duplex stainless steel butt welds. However, in their study, comparison of welding residual stress distributions between the dissimilar steel butt welds was not reported, i.e. they only provided limited information on the weld deformations, and thus the residual stress distributions in duplex stainless steel butt welds could not be presented. Therefore, 3-D FE analysis of the residual stresses in girth-welded duplex stainless steel pipes is needed to
[B] {c}
[Dd] {dF} {dR} {dw} [K] {L} {dε} {ε} {εO} {εL} {ds} V
141
displacement-strain matrix parameter to reflect stress increment due to the dependence of physical and mechanical properties of a material on the temperature constitutive matrix increment of equivalent nodal force due to the external force increment of equivalent nodal force due to the temperature change increment of nodal displacement element stiffness matrix load correction term incremental form of strain strain small strain large strain incremental form of stress spatial gradient operator
clearly identify the magnitude and distribution of the weldinduced residual stresses. Moreover, studies on the comparison of welding residual stresses in austenitic and duplex stainless steel pipe welds seem to be very scarce in the open literature. It cannot simply be assumed that welding residual stresses in duplex stainless steel welds are of the same magnitude or distribution as those in austenitic stainless steel welds, due to the different physical and mechanical properties. Differences include lower thermal expansion rate, higher thermal conductivity and higher strength with duplex stainless steels, and a rounded stress-strain curve with significant strain hardening (work hardening) with austenitic stainless steels. In this paper, at first, 3-D thermo-mechanical FE analysis method was developed in order to establish exact numerical model which can accurately capture the 3-D features of welding residual stress distributions in austenitic and duplex stainless steel pipe welds. In the FE analysis method, temperature-dependent material properties, work hardening behavior of the material welded, and weld filler variation with time are taken into account. Verification of the FE method was also implemented through the previously published experimental works. Then, based on the FE analysis method, the residual stresses in girth-welded austenitic stainless steel pipes were investigated and compared with those in duplex stainless steel pipe welds. 2. FE simulation of the girth welding process It is generally understood that welding involves highly complex phenomena, arising from the interactions between heat transfer, metallurgical transformation and mechanical fields. Complex numerical approaches are then needed to accurately model the welding process. However, as far as welding residual stress modeling is concerned, numerical procedures can be significantly simplified, as discussed in Ref. [29]. The welding process is essentially a coupled thermo-mechanical process. Because the thermal field has a strong influence on the stress field with little inverse influence, sequentially coupled analysis works very well. In this study, the girth welding process was simulated using a sequentially coupled 3-D thermomechanical FE formulation based on the in-house FE code written by Fortran language [30] in order to accurately capture the residual stress distributions in girth-welded stainless steel pipes. The
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procedure for welding residual stress analysis can be split into two solution steps: a transient thermal analysis followed by a transient thermalemechanical analysis. At the first step, a transient heat transfer analysis solves for the temperature distribution and its history associated with the heat flow of welding. The thermal analysis is based on the heat conduction formulation with the moving heat source. Then, the resulting temperature history solutions are fed into the thermalemechanical analysis as the thermal loading for thermal stress evolution. Welding residual stresses are the final state of the thermal stresses after all welding passes are over and the work piece is cooled down to the ambient temperature. The FE mesh refinements as well as the time increments for both the thermal analysis and the subsequent mechanical analysis are identical.
( h ¼
0 C < T < 500 C 0:0668 T W=m2 C 2 0:231 T 82:1 W=m C T > 500 C
(5)
To account for the heat effects relevant to the molten metal of the respective weld pool, two methodologies were used: (1) the liquidto-solid phase transformation effects of the weld pool were modeled by taking into account the latent heat of fusion, and (2) an artificially increased thermal conductivity, which is three times larger than the value at room temperature, was assumed for temperatures above the melting point to allow for its convective stirring effect, as suggested in Ref. [20]. The latent heat and melting temperature for austenitic stainless steels are 330 J/Kg K and 1663 K, respectively, and 500 J/Kg K and 1773 K for duplex stainless steels, respectively [2]. 2.2. Mechanical (stress) analysis
2.1. Thermal analysis The energy balance equation for the thermal analysis is given by:
VðkVTÞ þ q_ rcT_ ¼ 0
(1)
The most important parameter to determine the temperature distribution in a welded component is the heat input. This heat quantity is the output from a particular heat source used to fabricate welded joints. In all the welding processes, heat source provides the required energy and causes localized high temperature spot. In this study, the combined heat source model was employed to simulate the heat of the welding arc and the melt droplets. The heat input to the work piece is divided into two portions in the combined heat source model [31]. One is the heat of the welding arc, and the other is that of the melt droplets. The heat of the welding arc is modeled by a surface heat source with a Gaussian distribution, and that of the melt droplets is modeled by a volumetric heat source with uniform density. Heat flux distribution at the surface of the work piece within r0 is defined by the following equation:
qðrÞ ¼
3Q1 ðr=r0 Þ2 e pr02
Q1 ¼ hAV Q2
(2)
(3)
In Eq. (3), h accounts for radiative and other losses from the arc to the ambient environment. On the other hand, the heat from the melt droplets is applied as a volumetric heat source with a uniform heat flux, which is represented by the distributed heat flux (DFLUX) working on individual elements in the fusion zone. The DFLUX is calculated by
DFLUX ¼
Q2 Vp
(4)
where Vp can be obtained by calculating the volume fraction of the elements in currently being welded zone. The heat of the welding arc was assumed to be 40% of the total heat input, and that of the melt droplets 60% of the total heat input [32]. The arc efficiency factor was assumed as 0.7 for the gas tungsten arc (GTA) welding process used for both the stainless steel pipes [3,24]. The heat flux was applied during the time variation that corresponded to the approach and passing of the welding torch. As for the boundary conditions during the thermal analysis, convection and radiation are both taken into consideration and their combined effect is represented by h. In this work, the heat transfer coefficient of duplex stainless steels was assumed to be the same as that for austenitic stainless steels given by Ref. [17]
The second step of the current analysis involves the use of the thermal histories predicted by the previous thermal analysis for each time increment as an input (thermal loading) for the calculation of transient and residual thermal stress distributions. During the welding process, solid-state phase transformation is not considered because the metallurgical phase transformation does not occur in austenitic stainless steels and has little influence on the evolution of welding residual stresses in duplex stainless steels [2,28]. The modeling procedure for the FE simulation of welding incorporating solid-state phase transformation can be found in Ref. [33]. The following continuum mechanics can then be used to estimate the weld-induced residual stresses. 2.2.1. Displacementestrain relationship The strain of element can be written as follows
fεg ¼
n
εx εy εz gyz gzx gxy
oT
where {εO} and {εL} are given by
8 εx > > > > ε > y > < εz fεO g ¼ gyz > > > > g > > : zx
9 > > > > > > =
;
fεg ¼ fεO g þ fεL g
8 vu=vx > > > vv=vy > > < vw=vz ¼ > > > > ðvw=vyÞ þ ðvv=vzÞ > > > > > > > : ðvu=vzÞ þ ðvw=vxÞ gxy ; ðvv=vxÞ þ ðvu=vyÞ 8 vu2 vv2 vw2 9 þ vx þ vx > > vx > > > > > > > > > > > 2 2 2 > > > vu þ vv þ vw > > > > > > vy vy vy > > > > > > > > > > > > > > vu 2 > 2 2 > vv vw < = þ þ 1 vz vz vz ¼ 2> > 2 vu vu þ vv vv þ vw vw > > > > > > vy vz vy vz vy vz > > > > > > > > > > > > > > 2 vu vu þ vv vv þ vw vw > > > > > > vz vx vz vx vz vx > > > > > > > > > > : ; vu þ vv vv þ vw vw 2 vu vx vy vx vy vx vy
9 > > > > > = > > > > > ;
(6)
fεL g
It has been known that {dε} can be written as
fdεg ¼ ½Bfdwg
(7)
2.2.2. Stressestrain relationship The incremental form of stressestrain relation can be written as
fdsg ¼ ½Dd fdεg fcgdT
(8)
where [Dd] is divided into ½Ded for the elastic range and ½Dpd for the plastic range.
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2.2.3. Equilibrium equation Appling the virtual work theorem to the equilibrium equation between nodal force and nodal displacement yields
fdFg þ fLg fdRg ¼ ½Kfdwg
(9)
Composing Eq. (9) for the whole structure and solving the linear simultaneous equations considering boundary conditions, {dw} can be obtained. With the increment and using the displacemente strain relationship, the element strain can be gotten, thereby the element stress can be acquired using the constitutive law with the Von Mises yield criterion, temperature-dependent mechanical properties and linear isotropic hardening rule. In the thermal and mechanical analyses, a consistent filler activation/deactivation scheme was used to simulate the weld filler variation with time. This scheme keeps track of the movement of the welding torch and updates the status of weld filler (deposited or not). The thermal aspect of the scheme is to change the thermal conductivity of the weld filler. For the weld filler that is not yet deposited at a given time, a value for thermal conductivity equivalent to that of air is assigned. This process is called filler deactivation. After the weld filler is deposited, it is reactivated and the thermal conductivity is made to change from air value to that of the material used. The mechanical aspect of the scheme is to change the stiffness of the weld filler. For the weld filler which the welding torch has not yet approached, a severely reduced material stiffness is assigned [9]. At the time of application of the weld filler
Fig. 1. Comparison of the FE analysis results with the experimental measurements (austenitic stainless steel pipe weld): (a) axial residual stresses, (b) hoop residual stresses.
143
deposition, it is reactivated and the temperature-dependent mechanical properties of the material are assigned with no record of strain history to bring it into existence without incurring strain incompatibilities. During the analysis, the full NewtoneRaphson (NR) iterative solution technique [34] was employed for obtaining a solution. During the thermal cycle, temperature and temperaturedependent material properties change very rapidly. Thus, the full NR, which uses modified material properties and reformulates the stiffness matrix at every iteration step, was believed to give more accurate results [23]. 2.3. Verification In order to confirm the validity of the FE analysis method adopted in the present investigation, the experimental work by Deng and Murakawa [24] in which the residual stress distributions in a girth-welded austenitic stainless steel pipe (SUS304) were measured by the classical sectioning method was simulated. Furthermore, the experimental investigation by Um and Yoo [35] where the axial and hoop residual stresses in a ferritic carbon steel pipe weld (KS SPPS42) were evaluated by the hole-drilling method was reproduced due to the scarcity of the experimental data on the residual stress distributions in a girth-welded duplex (austenitic-ferritic) stainless steel pipe. The numerical results were compared with the corresponding experimental measurements. The specific details on the materials used and the preparation of the specimens are given in elsewhere [24,35]. The modeling procedure
Fig. 2. Comparison of the FE analysis results with the experimental measurements (ferritic carbon steel pipe weld): (a) axial residual stresses, (b) hoop residual stresses.
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Fig. 3. Analysis model: (a) dimensions of the analysis model and the welding direction, (b) 3-D FE model.
for the respective experiment is similar to that given in the forthcoming section except for the temperature-dependent thermal and mechanical properties for the girth-welded ferritic steel pipe [35]. The axial and hoop residual stresses calculated by the FE simulation of the stainless steel pipe weld at those locations where the circumferential angle from the welding start/stop position is 180 on the inner surface with respect to axial distance from the weld centerline are shown in Fig. 1(a) and (b), respectively. On the other hand, Fig. 2(a) and (b) depicts the axial and hoop residual stress distributions computed by the numerical reproduction of the girthwelded ferritic carbon steel pipe at those locations where the circumferential angle is 120 on the inner surface. Superimposed, in the figures, are the experimental measurements performed on the respective weld specimen. It can be seen that the residual stress distributions predicted by the FE analysis show very good
agreement with those determined by the experiment. Therefore, the FE analysis method used here can be considered appropriate for analyzing the residual stresses in girth-welded austenitic and duplex stainless steel pipes. 2.4. FE model FE thermal simulation of the girth butt-joint welding is carried out to compare girth-weld induced residual stresses in austenitic and duplex stainless steel pipe welds with the same geometry and groove shape using the aforementioned FE technique. Two 120 mm (outer diameter, D) 120 mm (length, L) 6 mm (thickness, t) stainless steel pipes with a single V-groove joint between them, as shown in Fig. 3(a), are assumed to be welded by single-pass. The figure also illustrates the welding arc travel direction by the arrow
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Fig. 4. Temperature-dependent thermo-physical constants of the materials: (a) SUS304 austenitic stainless steel pipe, (b) S32750 super duplex stainless steel pipe.
and the welding start/stop position (q ¼ 0 ). The same welding parameters for both the pipe welds were chosen here, which are similar to industrial practice [36] and are as follows: welding method, GTA welding process; welding current, 230 A; welding voltage, 22 V; and welding speed, 1.3 mm/s, respectively. The mesh refinement scheme for the 3-D FE model is shown in Fig. 3(b), which consists of 5152 elements and 6580 nodes. In the FE simulation of the girth welding process, symmetry conditions with respect to the weld centerline can be employed. Hence, only half of the girth-welded stainless steel pipe was modeled with 8-noded isoparametric solid elements and four layers were employed through the thickness. Generally, temperature around the arc is higher than the melting temperature of the material and drops sharply in regions away from weld pool. In high temperature gradient regions in the weld and its vicinity, more refined mesh is required to be adopted for obtaining an accurate temperature field. Element size increases progressively with distance from the weld centerline. Mesh sensitivity study has been conducted to examine
145
Fig. 5. Temperature-dependent thermo-mechanical properties of SUS304 austenitic stainless steel pipe.
the dependence of FE mesh size on the accuracy of the analysis results. As a result, the present FE mesh with the smallest element size of 0.9 mm (axial) 1.5 mm (thickness) 12.1 mm (circumference) is considered to give sufficiently accurate results using a reasonable amount of computer time and memory. In order to facilitate nodal data mapping between thermal and mechanical models, the same FE mesh model was used except for the element type and applied boundary conditions. For the thermal model, the
Table 1 Values of the thermal properties at the melting temperature. Material
Density (g/mm3)
Specific heat (J/g C)
Thermal conductivity (J/mm s C)
SUS304 S32750
0.00732 0.00732
0.708 0.724
0.1029 0.1131
Fig. 6. Temperature-dependent thermo-mechanical properties of S32750 super duplex stainless steel pipe.
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Fig. 7. Thermal cycles at different locations from the weld centerline on the outer surface during the welding: (a) SUS304 austenitic stainless steel pipe weld, (b) S32750 super duplex stainless steel pipe weld.
element type is one which has single degree of freedom, temperature, on its each node. For the structural model, the element type is the other with three translational degrees of freedom at each node. As the pipe was assumed not to be clamped during welding, the mechanical boundary conditions were prescribed for preventing rigid body motion of the weld piece. The detailed boundary conditions used in the FE model are shown in Fig. 3(b) by the arrows. The base materials chosen for this study are SUS304 austenitic stainless steel pipe and S32750 super duplex stainless steel pipe. Detailed information on the base materials is described in Refs. [1,24]. Temperature-dependent thermophysical (e.g., thermal conductivity, the heat transfer coefficient, specific heat and density) and mechanical properties (e.g., Young’s modulus, thermal expansion coefficient, Poisson’s ratio and yield stress) of the base metals were incorporated into the FE simulation. Fig. 4(a) and (b) shows the physical constants at high temperatures of SUS304 austenitic stainless steel pipe [24] and S32750 super duplex stainless steel pipe [2,37], respectively. It should be noted that the units in these figures are organized so that they can be shown on one graph for clarity. Note that only the temperaturedependent thermal conductivity and specific heat of the duplex stainless steel pipe are dissimilar to those of the austenitic stainless steel pipe and the other properties are similar to each other. The specific values of the thermal properties at the melting temperature are tabulated in Table 1. Temperature-dependent thermo-
mechanical properties of SUS304 austenitic stainless steel pipe are presented in Fig. 5 [24]. For the mechanical properties at high temperatures of S32750 super duplex stainless steel pipe, the elevated temperature tensile coupon tests were conducted in accordance with Korean standards [38]. An universal testing machine equipped with a specially made electrical furnace heated by thermal rays was used for the elevated temperature tensile test. Test specimens were machined as per the specifications [38] and tests were carried out in the elevated temperature range from 20 C (room temperature) to 900 C at intervals of 100 C with a strain rate of 1 mm/min, and the temperature was controlled to be within 2 C. In the experiment, thermal expansion was allowed by maintaining zero tension load during the heating process. Each specimen was held for approximately 20 min at the testing temperature before testing began to make sure the temperatures evenly distributed throughout the specimens. The dependence of the mechanical properties on temperatures based on the elevated temperature tensile test results are represented in Fig. 6. Both the yield stress and the elastic modulus are reduced to 5.0 MPa and 5.0 GPa, respectively, at the melting temperature to simulate low strength at high temperatures [39]. For the weld metal, autogenous weldment was assumed. This means that weld metal, heat affected zone and base metal share the same thermal and mechanical properties [40]. During welding process, because the weld metal and the base metal adjacent to the weld region are subject to cyclic thermal loading, the materials in these zones undergo plastic deformation to some extent; hence the weld area and its neighborhood work harden during welding. Austenitic stainless steel used in this study works harden very rapidly and, therefore, has a significant influence on the yield stress at the region near the welded zone. Higher yield stress due to the strain hardening induces high residual stresses there. As such, the work hardening behavior should be carefully taken into account when a numerical method is utilized to accurately predict welding residual stresses. In this work, temperature-dependent strain hardening rule was used. The strain hardening rates of the austenitic stainless steel pipe at high temperatures were measured and given in Ref. [20]. On the other hand, temperature-dependent strain hardening rates of the duplex stainless steel pipe were assumed to be the same as those of ferritic carbon steels [37,41]. 3. Results and discussion Results are first prepared for the transient thermal cycles associated with the girth welding of the stainless steel pipes. Fig. 7(a) and (b) compares the thermal cycles at different locations from the weld centerline on the outer surface during the girth welding of SUS304 austenitic stainless steel pipes and S32750 duplex stainless steel pipes, respectively, which are at the position where the circumferential angle from the weld start/stop position is 90 . The solid curve without any mark represents the temperature history at the weld region, the broken curve with solid marks the temperature history at the location which is 5.0 mm away from the weld centerline, the solid curve with hollow marks the temperature history at the location which is 10.0 mm away, and the dashed curve with hollow marks the temperature history at the location whose distance is 20.0 from the weld centerline. It can be found that the maximum temperature at the weld pool is 2100 C or so in both the pipe welds. This result agrees with the welding process in practice. It is worth noting that the peak temperature at the weld area of the duplex stainless steel pipe weld is higher compared to that of the austenitic stainless steel pipe weld in spite of the higher specific heat at the melting temperature (see Table 1), which is attributed to the fact that the difference between the melting
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147
Fig. 8. Residual stresses at locations with different circumferential angle q in the austenitic stainless steel girth-welded pipe: (a) axial stresses, inner surface, (b) axial stresses, outer surface, (c) hoop stresses, inner surface, (d) hoop stresses, outer surface.
temperatures of the stainless steel pipes is larger than that between the temperature increases due to the specific heats. Moreover, the cooling rate of the duplex stainless steel pipe weld is higher than that of the austenitic stainless steel pipe weld, which originates from the higher thermal conductivity of the duplex stainless steel pipe. Results are next presented for the comparison of the residual stress distributions induced by the girth welding of the austenitic and the duplex stainless steel pipe. Four different positions along the circumference are selected to portray welding residual stresses in order to examine the 3-D effects, i.e. circumferential variations of the residual stresses. The four positions have different circumferential angles from the welding start/stop position, which are 0 , 90 , 180 and 270 , respectively. From the stress analysis, all the stress and strain elements can be obtainable. Here, we will discuss only the relevant data. In this discussion, two words “axial” and “hoop” are used to denote the residual stress components. Axial residual stresses stand for the stress components which act normal to the weld line, whilst hoop residual stresses represent the stress components acting parallel to the weld line. The final residual stress distributions at locations with different circumferential angle in the austenitic stainless steel pipe weld are given in Fig. 8, with respect to axial distance from the weld centerline. Comparisons of the residual stresses in the girth-welded super duplex stainless steel pipe are also provided in Fig. 9. Fig. 8(a) and (b) depicts the axial residual stresses at the four locations on the inner and outer surfaces, respectively. From the simulated results, it can be observed that
within and near the weld area, the predicted axial residual stresses are tensile on the inner surface and compressive on the outer surface. For the case of thin-walled pipes, the heat deposited during girth welding is high enough to result in a uniform temperature increase through the pipe thickness at the weld region. Thus, the only deformation that will create thermal stresses is the circumferential strain due to the radial expansion and subsequent contraction. The circumferential shrinkage causes a local inward deformation in the vicinity of the welded zone, and thus generating a bending moment through the thickness. This leads to tensile axial residual stresses on the inner surface balanced by compressive stresses on the outer surface. Compressive axial residual stresses are formed on the inner surface away from the weld centerline, and tensile axial residual stresses on the outer surface for the self-equilibrating purpose. Moreover, it can also be found that even though the stress profiles along the four locations on both the inner and outer surfaces are similar to some extent, the magnitudes are different among them. It indicates that the axial residual stresses change with the circumferential angle. This is because the internal restraint during the pipe welding changes spatially due to the sequential deposition of the weld filler material as the welding arc moves along the circumference. Furthermore, the welding start/stop effects at the overlapping region make the variation severe. Similar trends can be noticed in the axial residual stress distributions of the duplex stainless steel girth-welded pipe. Fig. 8(c) and (d) portrays the hoop residual stresses at the four locations on the inner and outer surfaces, respectively. Regarding
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Fig. 9. Residual stresses at locations with different circumferential angle q in the duplex stainless steel girth-welded pipe: (a) axial stresses, inner surface, (b) axial stresses, outer surface, (c) hoop stresses, inner surface, (d) hoop stresses, outer surface.
the hoop residual stresses, their magnitude is influenced by the axial residual stresses. This explains why on the outer surface, which is experiencing axial compression, the hoop residual stresses are less tensile at the weld region and its vicinity compared to those on the inner surface. Stress reversal from tensile to compressive on the inner and outer surfaces away from the weld centerline is also identified. Furthermore, through careful observation of the results, it can be known that like the preceding axial residual stresses, spatial variations are present along the circumference. A rapid change of the residual stresses is also seen at the weld start/stop position. The hoop residual stress distributions in the girth-welded duplex stainless steel pipe shares similar characteristics. Therefore, it can be concluded that the axial and hoop residual stress distributions along the circumference are by no means axisymmetric, from which it is apparent that 3-D FE model is necessary to accurately capture the characteristics of welding residual stresses in the girth-welded stainless steel pipes. The results also unveil that the axial and hoop residual stresses at the welded zone in the duplex stainless steel pipe weld are higher compared to those in the girth-welded austenitic stainless steel pipe owing to the considerably higher yield stress of the duplex steel pipe. On the contrary, the stresses normalized by the yield stress are much higher in the austenitic stainless steel pipe weld in which the welding start/stop effects are more pronounced. It is also clear that the range subjected to tensile or compressive stress adjacent to the weld region is wider in the girth-welded austenitic stainless steel pipe. The variations in
outcomes between the dissimilar stainless steel pipe welds mainly arise from the differences in thermal conductivity, coefficient of thermal expansion and work hardening property. The larger coefficient of thermal expansion together with the much higher strain hardening rate of the austenitic stainless steel pipe gives rise to the higher stress ratios balanced by the higher reverse stress ratios away from the weld region. They also induce the steeper stress change at the overlapping area. In addition, the lower thermal conductivity in conjunction with the larger coefficient of thermal expansion of the austenitic stainless steel pipe contributes to the larger stress distribution region in and around the weld area. Results are finally given for the residual effective stresses for the girth-welded stainless steel pipes. Fig. 10(a) and (b) portrays the residual effective stresses in the austenitic stainless steel pipe weld at the four locations on the inner and outer surfaces with respect to axial distance from the weld centerline. It is worth noting that the maximum residual effective stress produced at the girth weld is 380 MPa or so, which is much below the ultimate strength of the austenitic stainless steel pipe, though the maximum axial and hoop residual stresses at the weld start/stop position are nearly 800 MPa. The effective stresses also demonstrate the 3-D effects within and near the welded zone. The higher effective stresses decrease as the distance from the weld centerline increases and finally converge to zero at the region far from the girth weld. The residual effective stresses in the girth-welded duplex stainless steel pipe are also presented in Fig. 11, which share similar characteristics.
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Fig. 10. Residual effective stresses at the four locations in the austenitic stainless steel pipe weld: (a) inner surface, (b) outer surface.
149
Fig. 11. Residual effective stresses at the four locations in the duplex stainless steel pipe weld: (a) inner surface, (b) outer surface.
Acknowledgements 4. Conclusions In the present work, FE thermal simulation of the girth welding process was performed using a sequentially coupled 3-D thermomechanical FE analysis method whose effectiveness was verified by the comparisons with the experimental measurements to identify and compare temperature fields and welding residual stress distributions in girth-welded austenitic and duplex stainless steel pipes which have different thermal and mechanical properties. Based on the results, the following observations and conclusions can be drawn: a) 3-D FE model is apparently necessary to accurately predict the residual stress distributions in girth-welded austenitic and duplex stainless steel pipes which can incorporate the 3-D effects. b) Girth-welded duplex stainless steel pipes have higher peak temperature at the weld region and higher cooling rate compared to those developed during the girth welding of austenitic stainless steel pipes. c) The axial and hoop residual stresses normalized by the yield stress at the weld and its neighborhood are much higher in austenitic stainless steel pipe welds in which the welding start/stop effects are more significant, compared to those in duplex stainless steel pipe welds. Moreover, girth-welded austenitic stainless steel pipes have wider regions subjected to tensile or compressive stresses adjacent to the weld area.
This research was supported by Basic Science Research Program through the National Research Foundation of Korea (NRF) funded by the Ministry of Education (2012R1A1A2044895). References [1] A.-A. Iris, Duplex stainless steels: brief history and some recent alloys, Recent Pat. Mech. Eng. 1 (2008) 51e57. [2] J.J. del Coz Díaz, P. Menendez Rodriguez, P.J. Garcia Nieto, D. Castro-Fresno, Comparative analysis of TIG welding distortions between austenitic and duplex stainless steels by FEM, Appl. Therm. Eng. 30 (2010) 2448e2459. [3] A.-H.I. Mourad, A. Khourshid, T. Sharef, Gas tungsten arc and laser beam welding processes effects on duplex stainless steel 2205 properties, Mater. Sci. Eng. A 549 (2012) 105e113. [4] P.J. Withers, Residual stress and its role in failure, Rep. Prog. Phys. 70 (2007) 2211e2264. [5] H.D. Hibbitt, P.V. Marcal, A numerical thermo-mechanical model for the welding and subsequent loading of fabrication structure, Comput. Struct. 3 (1973) 1145e1174. [6] J. Goldak, M. Bibby, J. Moore, J.R. House, B. Patel, Computer modeling of heat flow in welds, Metall. Mater. Trans. B 17 (1986) 587e600. [7] R.I. Karlsson, B.L. Josefson, Three-dimensional finite element analysis of temperatures and stresses in a single-pass butt-welded pipe, J. Press. Vessel Technol. 112 (1990) 76e84. [8] L.-E. Lindgren, Finite element modelling and simulation of welding, part 1 increased complexity, J. Therm. Stresses 24 (2001) 141e192. [9] L.-E. Lindgren, Finite element modelling and simulation of welding, part 2 improved material modeling, J. Therm. Stresses 24 (2001) 195e231. [10] L.-E. Lindgren, Finite element modelling and simulation of welding, part 3 efficiency and integration, J. Therm. Stresses 24 (2001) 305e334. [11] J. Goldak, M. Akhlagi, Computational Welding Mechanics, Springer, 2005. [12] L.-E. Lindgren, Numerical modelling of welding, Comput. Meth. Appl. Mech. Eng. 195 (2006) 6710e6736.
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[13] L.-E. Lindgren, Computational Welding Mechanics (Thermomechanical and Microstructural Simulations), Woodhead Publishing, 2007. [14] M. Ashraf, L. Gardner, D.A. Nethercot, Finite element modelling of structural stainless steel cross-sections, Thin-Walled Struct. 44 (2006) 1048e1062. [15] E.F. Rybicki, D.W. Schmueser, R.W. Stonesifer, J.J. Groom, H.W. Mishaler, A finite element model for residual stresses and deflections in girth-butt welded pipes, J. Press. Vessel Technol. 100 (1978) 256e262. [16] E.F. Rybicki, P.A. McGuire, E. Merrick, J. Wert, The effect of pipe thickness on residual stresses due to girth welds, J. Press. Vessel Technol. 104 (1982) 204e209. [17] B. Brickstad, B.L. Josefson, A parametric study of residual stresses in multi-pass butt-welded stainless steel pipes, Int. J. Press. Vessels Pip. 75 (1998) 11e25. [18] M. Mochizuki, M. Hayashi, T. Hattori, Residual stress distribution depending on welding sequence in multi-pass welded joints with X-shaped groove, J. Press. Vessel Technol. 122 (2000) 27e32. [19] A. Yaghi, T.H. Hyde, A.A. Becker, W. Sun, J.A. Williams, Residual stress simulation in thin and thick-walled stainless steel pipe welds including pipe diameter effects, Int. J. Press. Vessels Pip. 83 (2006) 864e874. [20] D. Deng, H. Murakawa, W. Liang, Numerical and experimental investigations on welding residual stress in multi-pass butt-welded austenitic stainless steel pipe, Comput. Mater. Sci. 42 (2008) 234e244. [21] S. Fricke, E. Keim, J. Schmidt, Numerical weld modeling - a method for calculating weld-induced residual stresses, Nucl. Eng. Des. 206 (2001) 139e150. [22] P. Duranton, J. Devaux, V. Robin, P. Gilles, J.M. Bergheau, 3D modeling of multipass welding of a 316L stainless steel pipe, J. Mater. Process. Technol. 153e154 (2004) 457e463. [23] M. Abid, M. Siddique, Numerical simulation to study the effect of tack welds and root gap on welding deformations and residual stresses of a pipe-flange joint, Int. J. Press. Vessels Pip. 82 (2005) 860e871. [24] D. Deng, H. Murakawa, Numerical simulation of temperature field and residual stress in multi-pass welds in stainless steel pipe and comparison with experimental measurements, Comput. Mater. Sci. 37 (2006) 269e277. [25] I. Sattari-Far, M.R. Farahani, Effect of the weld groove shape and pass number on residual stresses in butt-welded pipes, Int. J. Press. Vessels Pip. 86 (2009) 723e731. [26] D. Deng, S. Kiyoshima, FEM prediction of welding residual stresses in a SUS304 girth-welded pipe with emphasis on stress distribution near weld start/end location, Comput. Mater. Sci. 50 (2010) 612e621.
[27] C.H. Lee, K.H. Chang, Influence of the residual stresses and distortions on the structural behavior of girth-welded cylindrical steel members, Constr. Build. Mater. 41 (2013) 766e776. [28] X. Jin, L. Huo, H. Yu, B. Bai, X. Li, J. Cao, Numerical simulation of parameter of residual stresses on butt-welded duplex stainless steel pipes, J. Tianjin Univ. 37 (2004) 116e125. [29] P. Dong, B.W. Brust, Welding residual stresses and effects on fracture in pressure vessel and piping components: a millennium review and beyond, J. Press. Vessel Technol. 122 (2000) 329e338. [30] C.H. Lee, A Study on the Mechanical Characteristics of High Strength Steel for the Application to the Steel Bridge (Ph.D. Thesis), Chung-Ang University, Korea, 2005. [31] D. Deng, W. Liang, H. Murakawa, Determination of welding deformation in fillet-welded joint by means of numerical simulation and comparison with experimental measurements, J. Mater. Process. Technol. 186 (2007) 219e225. [32] E. Pardo, D.C. Weckman, Prediction of weld pool and reinforcement dimensions of GMA welds using a finite element model, Metall. Mater. Trans. B 20 (1989) 937e947. [33] C.H. Lee, K.H. Chang, Prediction of residual stresses in high strength carbon steel pipe weld considering solid-state phase transformation effects, Comput. Struct. 89 (2011) 256e265. [34] K.J. Bathe, Finite Element Procedures, Prentice Hall, 1996. [35] D.S. Um, K.Y. Yoo, The experimental studies on residual stresses due to circumferential welds in thin steel cylinder, J. Korean Weld. Soc. 15 (1997) 149e156. [36] I. Sattari-Far, Y. Javadi, Influence of welding sequence on welding distortions in pipes, Int. J. Press. Vessels Pip. 85 (2008) 265e274. [37] Stainless Steels and Their Properties, 2013. http://www.hazmetal.com/f/kutu/ 1236776229.pdf. [38] KS D 0026, Method of Elevated Temperature Tensile Test for Steels and Heatresisting Alloys, Korean Standards, 2002. [39] Z. Barsoum, Residual stress analysis and fatigue of multi-pass welded tubular structures, Eng. Fail. Anal. 15 (2008) 863e874. [40] T.L. Teng, P.H. Chang, W.C. Tseng, Effect of welding sequence on residual stresses, Comput. Struct. 81 (2003) 273e286. [41] B. Taljat, B. Radhakrishnan, T. Zacharia, Numerical analysis of GTA welding process with emphasis on post-solidification phase transformation effects on the residual stresses, Mater. Sci. Eng. A 246 (1998) 45e54.
Contents lists available at ScienceDirect
Applied Thermal Engineering journal homepage: www.elsevier.com/locate/apthermeng
Comparative study on girth weld-induced residual stresses between austenitic and duplex stainless steel pipe welds Chin-Hyung Lee, Kyong-Ho Chang* Department of Civil and Environmental & Plant Engineering, Chung-Ang University, 84, Huksuk-ro, Dongjak-ku, Seoul 156-756, Republic of Korea
h i g h l i g h t s Residual stresses in austenitic and duplex stainless steel pipes were investigated. Comparative study on the girth weld-induced residual stresses was performed. Sequentially coupled 3-D thermo-mechanical FE analysis was conducted. Considerably different welding residual stress distributions could be demonstrated.
a r t i c l e i n f o
a b s t r a c t
Article history: Received 3 June 2013 Accepted 1 November 2013 Available online 13 November 2013
Duplex stainless steel pipes find increasing use as an alternative to austenitic stainless steel pipes, particularly where chloride or sulphide stress corrosion cracking is of primary concern, due to the excellent combination of strength and corrosion resistance. During welding, duplex stainless steel pipes do not create the same magnitude or distribution of weld-induced residual stresses as those in girthwelded austenitic stainless steel pipes due to the different physical and mechanical properties between them. In this work, a comparison of the residual stresses between girth-welded austenitic and duplex stainless steel pipes was performed utilizing sequentially coupled three-dimensional thermomechanical finite element analysis method to accurately predict temperature fields and residual stress states in pipe girth welds. The results have shown that girth-welded austenitic stainless steel pipes produce much higher axial and hoop residual stresses normalized by the yield stress at the weld and its vicinity in which the welding start/stop effects are more significant and they have wider regions subjected to tensile or compressive stresses adjacent to the weld area. Ó 2013 Elsevier Ltd. All rights reserved.
Keywords: Austenitic stainless steel pipe welds Duplex stainless steel pipe welds Welding residual stresses Three-dimensional finite element simulation Thermal and mechanical properties
1. Introduction Duplex stainless steels, with a microstructure comprised of nearly equal proportions of ferrite and austenite, combine the attractive properties of austenitic and ferritic steels: high tensile strength and fatigue strength, good toughness even at low temperatures, adequate formability and weldability and excellent resistance to stress corrosion cracking, pitting and general corrosion [1e3]. They, especially in the pipe form, find increasing use as an alternative to austenitic stainless steels, particularly where chloride or sulphide stress corrosion cracking is of primary concern, e.g., in the chemical, oil and gas, paper and pulp, marine and petrochemical industries and pollution control equipment. In
* Corresponding author. Tel.: þ82 2 820 5337; fax: þ82 2 823 5339. E-mail address: ifi[email protected] (K.-H. Chang). 1359-4311/$ e see front matter Ó 2013 Elsevier Ltd. All rights reserved. http://dx.doi.org/10.1016/j.applthermaleng.2013.11.001
practical situations, girth welding of duplex stainless steel pipes is often needed owing to the long geometry relative to the diameter and the wall-thickness. Welding is a crucial manufacturing process and widely used in industries to assemble various engineering and structural components. The advantage of welding as joining process includes high joint efficiency, simple set up and low fabrication cost. In welded structures, the inevitable existence of welding residual stresses is well known, which are produced as a result of plastic deformation caused by non-uniform thermal expansion and contraction imposed during welding process. The residual stresses are always self-equilibrated and the magnitude of tensile residual stresses within and near the weld area is great enough to have deleterious effects on the structural integrity of the welded joints, increasing the susceptibility to fatigue damage, stress corrosion cracking and brittle fracture [4]. These stresses when combined with the applied stresses can reduce the fatigue life, accelerate growth rates of pre-
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Nomenclature A c h k q_ q(r) Q1 Q2 r r0 T T_ V Vp dT
h r
arc current specific heat temperature-dependent heat transfer coefficient thermal conductivity rate of moving heat generation per unit volume heat flux distribution heat input from the welding arc energy induced by high temperature melt droplets radial coordinate with the origin at the arc center arc beam radius temperature rate of temperature change arc voltage considered weld pool volume temperature increment arc efficiency factor density
existing or service-induced defects in structural systems. Knowledge of the distribution and magnitude of the residual stresses, therefore, would be of big help for the production of an efficient and economic design and safety of the structures. However, accurate prediction of welding residual stresses is very difficult because of the complexity of welding process which includes localized heating, temperature dependence of material properties and moving heat source, etc. Nowadays, numerical modeling based on finite element (FE) method is used to predict welding residual stresses due to the expense and impracticalities of generating comprehensive structural performance data through experiments [5e13]. Numerical techniques have become an important part of most structural research communities, since they can be employed as a useful tool for analyzing the behavior of structures provided that suitable care is taken to ensure that the modeling is appropriate for the analysis [14]. Until now, a large number of FE simulations focusing on the circumferential welding of austenitic stainless steel pipes have been performed to predict the girth weld-induced residual stresses through the axisymmetric models [15e20] or the threedimensional (3-D) models [7,21e27] by using commercially available FE-codes such as ABAQUS, ANSYS and SYSWELD, etc. Thus, welding residual stresses in girth-welded stainless steel pipe components have been thoroughly investigated. As for the duplex stainless steel pipe welds, to the knowledge of the authors, very few works have been published on the FE analysis of welding residual stresses. Jin et al. [28] evaluated the axial and hoop residual stresses in a circumferentially butt-welded 2205 (EN 1.4462) duplex stainless steel pipe through the numerical simulation based on the nonlinear thermo-mechanical FE analysis. Nevertheless, their work was confined to the axisymmetric model which was not capable of predicting the 3-D effects induced by the girth welding process. Díaz et al. [2] conducted thermo-mechanical FE analyses in order to compare weld-induced distortion modes and magnitudes between austenitic and duplex stainless steel butt welds. However, in their study, comparison of welding residual stress distributions between the dissimilar steel butt welds was not reported, i.e. they only provided limited information on the weld deformations, and thus the residual stress distributions in duplex stainless steel butt welds could not be presented. Therefore, 3-D FE analysis of the residual stresses in girth-welded duplex stainless steel pipes is needed to
[B] {c}
[Dd] {dF} {dR} {dw} [K] {L} {dε} {ε} {εO} {εL} {ds} V
141
displacement-strain matrix parameter to reflect stress increment due to the dependence of physical and mechanical properties of a material on the temperature constitutive matrix increment of equivalent nodal force due to the external force increment of equivalent nodal force due to the temperature change increment of nodal displacement element stiffness matrix load correction term incremental form of strain strain small strain large strain incremental form of stress spatial gradient operator
clearly identify the magnitude and distribution of the weldinduced residual stresses. Moreover, studies on the comparison of welding residual stresses in austenitic and duplex stainless steel pipe welds seem to be very scarce in the open literature. It cannot simply be assumed that welding residual stresses in duplex stainless steel welds are of the same magnitude or distribution as those in austenitic stainless steel welds, due to the different physical and mechanical properties. Differences include lower thermal expansion rate, higher thermal conductivity and higher strength with duplex stainless steels, and a rounded stress-strain curve with significant strain hardening (work hardening) with austenitic stainless steels. In this paper, at first, 3-D thermo-mechanical FE analysis method was developed in order to establish exact numerical model which can accurately capture the 3-D features of welding residual stress distributions in austenitic and duplex stainless steel pipe welds. In the FE analysis method, temperature-dependent material properties, work hardening behavior of the material welded, and weld filler variation with time are taken into account. Verification of the FE method was also implemented through the previously published experimental works. Then, based on the FE analysis method, the residual stresses in girth-welded austenitic stainless steel pipes were investigated and compared with those in duplex stainless steel pipe welds. 2. FE simulation of the girth welding process It is generally understood that welding involves highly complex phenomena, arising from the interactions between heat transfer, metallurgical transformation and mechanical fields. Complex numerical approaches are then needed to accurately model the welding process. However, as far as welding residual stress modeling is concerned, numerical procedures can be significantly simplified, as discussed in Ref. [29]. The welding process is essentially a coupled thermo-mechanical process. Because the thermal field has a strong influence on the stress field with little inverse influence, sequentially coupled analysis works very well. In this study, the girth welding process was simulated using a sequentially coupled 3-D thermomechanical FE formulation based on the in-house FE code written by Fortran language [30] in order to accurately capture the residual stress distributions in girth-welded stainless steel pipes. The
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procedure for welding residual stress analysis can be split into two solution steps: a transient thermal analysis followed by a transient thermalemechanical analysis. At the first step, a transient heat transfer analysis solves for the temperature distribution and its history associated with the heat flow of welding. The thermal analysis is based on the heat conduction formulation with the moving heat source. Then, the resulting temperature history solutions are fed into the thermalemechanical analysis as the thermal loading for thermal stress evolution. Welding residual stresses are the final state of the thermal stresses after all welding passes are over and the work piece is cooled down to the ambient temperature. The FE mesh refinements as well as the time increments for both the thermal analysis and the subsequent mechanical analysis are identical.
( h ¼
0 C < T < 500 C 0:0668 T W=m2 C 2 0:231 T 82:1 W=m C T > 500 C
(5)
To account for the heat effects relevant to the molten metal of the respective weld pool, two methodologies were used: (1) the liquidto-solid phase transformation effects of the weld pool were modeled by taking into account the latent heat of fusion, and (2) an artificially increased thermal conductivity, which is three times larger than the value at room temperature, was assumed for temperatures above the melting point to allow for its convective stirring effect, as suggested in Ref. [20]. The latent heat and melting temperature for austenitic stainless steels are 330 J/Kg K and 1663 K, respectively, and 500 J/Kg K and 1773 K for duplex stainless steels, respectively [2]. 2.2. Mechanical (stress) analysis
2.1. Thermal analysis The energy balance equation for the thermal analysis is given by:
VðkVTÞ þ q_ rcT_ ¼ 0
(1)
The most important parameter to determine the temperature distribution in a welded component is the heat input. This heat quantity is the output from a particular heat source used to fabricate welded joints. In all the welding processes, heat source provides the required energy and causes localized high temperature spot. In this study, the combined heat source model was employed to simulate the heat of the welding arc and the melt droplets. The heat input to the work piece is divided into two portions in the combined heat source model [31]. One is the heat of the welding arc, and the other is that of the melt droplets. The heat of the welding arc is modeled by a surface heat source with a Gaussian distribution, and that of the melt droplets is modeled by a volumetric heat source with uniform density. Heat flux distribution at the surface of the work piece within r0 is defined by the following equation:
qðrÞ ¼
3Q1 ðr=r0 Þ2 e pr02
Q1 ¼ hAV Q2
(2)
(3)
In Eq. (3), h accounts for radiative and other losses from the arc to the ambient environment. On the other hand, the heat from the melt droplets is applied as a volumetric heat source with a uniform heat flux, which is represented by the distributed heat flux (DFLUX) working on individual elements in the fusion zone. The DFLUX is calculated by
DFLUX ¼
Q2 Vp
(4)
where Vp can be obtained by calculating the volume fraction of the elements in currently being welded zone. The heat of the welding arc was assumed to be 40% of the total heat input, and that of the melt droplets 60% of the total heat input [32]. The arc efficiency factor was assumed as 0.7 for the gas tungsten arc (GTA) welding process used for both the stainless steel pipes [3,24]. The heat flux was applied during the time variation that corresponded to the approach and passing of the welding torch. As for the boundary conditions during the thermal analysis, convection and radiation are both taken into consideration and their combined effect is represented by h. In this work, the heat transfer coefficient of duplex stainless steels was assumed to be the same as that for austenitic stainless steels given by Ref. [17]
The second step of the current analysis involves the use of the thermal histories predicted by the previous thermal analysis for each time increment as an input (thermal loading) for the calculation of transient and residual thermal stress distributions. During the welding process, solid-state phase transformation is not considered because the metallurgical phase transformation does not occur in austenitic stainless steels and has little influence on the evolution of welding residual stresses in duplex stainless steels [2,28]. The modeling procedure for the FE simulation of welding incorporating solid-state phase transformation can be found in Ref. [33]. The following continuum mechanics can then be used to estimate the weld-induced residual stresses. 2.2.1. Displacementestrain relationship The strain of element can be written as follows
fεg ¼
n
εx εy εz gyz gzx gxy
oT
where {εO} and {εL} are given by
8 εx > > > > ε > y > < εz fεO g ¼ gyz > > > > g > > : zx
9 > > > > > > =
;
fεg ¼ fεO g þ fεL g
8 vu=vx > > > vv=vy > > < vw=vz ¼ > > > > ðvw=vyÞ þ ðvv=vzÞ > > > > > > > : ðvu=vzÞ þ ðvw=vxÞ gxy ; ðvv=vxÞ þ ðvu=vyÞ 8 vu2 vv2 vw2 9 þ vx þ vx > > vx > > > > > > > > > > > 2 2 2 > > > vu þ vv þ vw > > > > > > vy vy vy > > > > > > > > > > > > > > vu 2 > 2 2 > vv vw < = þ þ 1 vz vz vz ¼ 2> > 2 vu vu þ vv vv þ vw vw > > > > > > vy vz vy vz vy vz > > > > > > > > > > > > > > 2 vu vu þ vv vv þ vw vw > > > > > > vz vx vz vx vz vx > > > > > > > > > > : ; vu þ vv vv þ vw vw 2 vu vx vy vx vy vx vy
9 > > > > > = > > > > > ;
(6)
fεL g
It has been known that {dε} can be written as
fdεg ¼ ½Bfdwg
(7)
2.2.2. Stressestrain relationship The incremental form of stressestrain relation can be written as
fdsg ¼ ½Dd fdεg fcgdT
(8)
where [Dd] is divided into ½Ded for the elastic range and ½Dpd for the plastic range.
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2.2.3. Equilibrium equation Appling the virtual work theorem to the equilibrium equation between nodal force and nodal displacement yields
fdFg þ fLg fdRg ¼ ½Kfdwg
(9)
Composing Eq. (9) for the whole structure and solving the linear simultaneous equations considering boundary conditions, {dw} can be obtained. With the increment and using the displacemente strain relationship, the element strain can be gotten, thereby the element stress can be acquired using the constitutive law with the Von Mises yield criterion, temperature-dependent mechanical properties and linear isotropic hardening rule. In the thermal and mechanical analyses, a consistent filler activation/deactivation scheme was used to simulate the weld filler variation with time. This scheme keeps track of the movement of the welding torch and updates the status of weld filler (deposited or not). The thermal aspect of the scheme is to change the thermal conductivity of the weld filler. For the weld filler that is not yet deposited at a given time, a value for thermal conductivity equivalent to that of air is assigned. This process is called filler deactivation. After the weld filler is deposited, it is reactivated and the thermal conductivity is made to change from air value to that of the material used. The mechanical aspect of the scheme is to change the stiffness of the weld filler. For the weld filler which the welding torch has not yet approached, a severely reduced material stiffness is assigned [9]. At the time of application of the weld filler
Fig. 1. Comparison of the FE analysis results with the experimental measurements (austenitic stainless steel pipe weld): (a) axial residual stresses, (b) hoop residual stresses.
143
deposition, it is reactivated and the temperature-dependent mechanical properties of the material are assigned with no record of strain history to bring it into existence without incurring strain incompatibilities. During the analysis, the full NewtoneRaphson (NR) iterative solution technique [34] was employed for obtaining a solution. During the thermal cycle, temperature and temperaturedependent material properties change very rapidly. Thus, the full NR, which uses modified material properties and reformulates the stiffness matrix at every iteration step, was believed to give more accurate results [23]. 2.3. Verification In order to confirm the validity of the FE analysis method adopted in the present investigation, the experimental work by Deng and Murakawa [24] in which the residual stress distributions in a girth-welded austenitic stainless steel pipe (SUS304) were measured by the classical sectioning method was simulated. Furthermore, the experimental investigation by Um and Yoo [35] where the axial and hoop residual stresses in a ferritic carbon steel pipe weld (KS SPPS42) were evaluated by the hole-drilling method was reproduced due to the scarcity of the experimental data on the residual stress distributions in a girth-welded duplex (austenitic-ferritic) stainless steel pipe. The numerical results were compared with the corresponding experimental measurements. The specific details on the materials used and the preparation of the specimens are given in elsewhere [24,35]. The modeling procedure
Fig. 2. Comparison of the FE analysis results with the experimental measurements (ferritic carbon steel pipe weld): (a) axial residual stresses, (b) hoop residual stresses.
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Fig. 3. Analysis model: (a) dimensions of the analysis model and the welding direction, (b) 3-D FE model.
for the respective experiment is similar to that given in the forthcoming section except for the temperature-dependent thermal and mechanical properties for the girth-welded ferritic steel pipe [35]. The axial and hoop residual stresses calculated by the FE simulation of the stainless steel pipe weld at those locations where the circumferential angle from the welding start/stop position is 180 on the inner surface with respect to axial distance from the weld centerline are shown in Fig. 1(a) and (b), respectively. On the other hand, Fig. 2(a) and (b) depicts the axial and hoop residual stress distributions computed by the numerical reproduction of the girthwelded ferritic carbon steel pipe at those locations where the circumferential angle is 120 on the inner surface. Superimposed, in the figures, are the experimental measurements performed on the respective weld specimen. It can be seen that the residual stress distributions predicted by the FE analysis show very good
agreement with those determined by the experiment. Therefore, the FE analysis method used here can be considered appropriate for analyzing the residual stresses in girth-welded austenitic and duplex stainless steel pipes. 2.4. FE model FE thermal simulation of the girth butt-joint welding is carried out to compare girth-weld induced residual stresses in austenitic and duplex stainless steel pipe welds with the same geometry and groove shape using the aforementioned FE technique. Two 120 mm (outer diameter, D) 120 mm (length, L) 6 mm (thickness, t) stainless steel pipes with a single V-groove joint between them, as shown in Fig. 3(a), are assumed to be welded by single-pass. The figure also illustrates the welding arc travel direction by the arrow
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Fig. 4. Temperature-dependent thermo-physical constants of the materials: (a) SUS304 austenitic stainless steel pipe, (b) S32750 super duplex stainless steel pipe.
and the welding start/stop position (q ¼ 0 ). The same welding parameters for both the pipe welds were chosen here, which are similar to industrial practice [36] and are as follows: welding method, GTA welding process; welding current, 230 A; welding voltage, 22 V; and welding speed, 1.3 mm/s, respectively. The mesh refinement scheme for the 3-D FE model is shown in Fig. 3(b), which consists of 5152 elements and 6580 nodes. In the FE simulation of the girth welding process, symmetry conditions with respect to the weld centerline can be employed. Hence, only half of the girth-welded stainless steel pipe was modeled with 8-noded isoparametric solid elements and four layers were employed through the thickness. Generally, temperature around the arc is higher than the melting temperature of the material and drops sharply in regions away from weld pool. In high temperature gradient regions in the weld and its vicinity, more refined mesh is required to be adopted for obtaining an accurate temperature field. Element size increases progressively with distance from the weld centerline. Mesh sensitivity study has been conducted to examine
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Fig. 5. Temperature-dependent thermo-mechanical properties of SUS304 austenitic stainless steel pipe.
the dependence of FE mesh size on the accuracy of the analysis results. As a result, the present FE mesh with the smallest element size of 0.9 mm (axial) 1.5 mm (thickness) 12.1 mm (circumference) is considered to give sufficiently accurate results using a reasonable amount of computer time and memory. In order to facilitate nodal data mapping between thermal and mechanical models, the same FE mesh model was used except for the element type and applied boundary conditions. For the thermal model, the
Table 1 Values of the thermal properties at the melting temperature. Material
Density (g/mm3)
Specific heat (J/g C)
Thermal conductivity (J/mm s C)
SUS304 S32750
0.00732 0.00732
0.708 0.724
0.1029 0.1131
Fig. 6. Temperature-dependent thermo-mechanical properties of S32750 super duplex stainless steel pipe.
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Fig. 7. Thermal cycles at different locations from the weld centerline on the outer surface during the welding: (a) SUS304 austenitic stainless steel pipe weld, (b) S32750 super duplex stainless steel pipe weld.
element type is one which has single degree of freedom, temperature, on its each node. For the structural model, the element type is the other with three translational degrees of freedom at each node. As the pipe was assumed not to be clamped during welding, the mechanical boundary conditions were prescribed for preventing rigid body motion of the weld piece. The detailed boundary conditions used in the FE model are shown in Fig. 3(b) by the arrows. The base materials chosen for this study are SUS304 austenitic stainless steel pipe and S32750 super duplex stainless steel pipe. Detailed information on the base materials is described in Refs. [1,24]. Temperature-dependent thermophysical (e.g., thermal conductivity, the heat transfer coefficient, specific heat and density) and mechanical properties (e.g., Young’s modulus, thermal expansion coefficient, Poisson’s ratio and yield stress) of the base metals were incorporated into the FE simulation. Fig. 4(a) and (b) shows the physical constants at high temperatures of SUS304 austenitic stainless steel pipe [24] and S32750 super duplex stainless steel pipe [2,37], respectively. It should be noted that the units in these figures are organized so that they can be shown on one graph for clarity. Note that only the temperaturedependent thermal conductivity and specific heat of the duplex stainless steel pipe are dissimilar to those of the austenitic stainless steel pipe and the other properties are similar to each other. The specific values of the thermal properties at the melting temperature are tabulated in Table 1. Temperature-dependent thermo-
mechanical properties of SUS304 austenitic stainless steel pipe are presented in Fig. 5 [24]. For the mechanical properties at high temperatures of S32750 super duplex stainless steel pipe, the elevated temperature tensile coupon tests were conducted in accordance with Korean standards [38]. An universal testing machine equipped with a specially made electrical furnace heated by thermal rays was used for the elevated temperature tensile test. Test specimens were machined as per the specifications [38] and tests were carried out in the elevated temperature range from 20 C (room temperature) to 900 C at intervals of 100 C with a strain rate of 1 mm/min, and the temperature was controlled to be within 2 C. In the experiment, thermal expansion was allowed by maintaining zero tension load during the heating process. Each specimen was held for approximately 20 min at the testing temperature before testing began to make sure the temperatures evenly distributed throughout the specimens. The dependence of the mechanical properties on temperatures based on the elevated temperature tensile test results are represented in Fig. 6. Both the yield stress and the elastic modulus are reduced to 5.0 MPa and 5.0 GPa, respectively, at the melting temperature to simulate low strength at high temperatures [39]. For the weld metal, autogenous weldment was assumed. This means that weld metal, heat affected zone and base metal share the same thermal and mechanical properties [40]. During welding process, because the weld metal and the base metal adjacent to the weld region are subject to cyclic thermal loading, the materials in these zones undergo plastic deformation to some extent; hence the weld area and its neighborhood work harden during welding. Austenitic stainless steel used in this study works harden very rapidly and, therefore, has a significant influence on the yield stress at the region near the welded zone. Higher yield stress due to the strain hardening induces high residual stresses there. As such, the work hardening behavior should be carefully taken into account when a numerical method is utilized to accurately predict welding residual stresses. In this work, temperature-dependent strain hardening rule was used. The strain hardening rates of the austenitic stainless steel pipe at high temperatures were measured and given in Ref. [20]. On the other hand, temperature-dependent strain hardening rates of the duplex stainless steel pipe were assumed to be the same as those of ferritic carbon steels [37,41]. 3. Results and discussion Results are first prepared for the transient thermal cycles associated with the girth welding of the stainless steel pipes. Fig. 7(a) and (b) compares the thermal cycles at different locations from the weld centerline on the outer surface during the girth welding of SUS304 austenitic stainless steel pipes and S32750 duplex stainless steel pipes, respectively, which are at the position where the circumferential angle from the weld start/stop position is 90 . The solid curve without any mark represents the temperature history at the weld region, the broken curve with solid marks the temperature history at the location which is 5.0 mm away from the weld centerline, the solid curve with hollow marks the temperature history at the location which is 10.0 mm away, and the dashed curve with hollow marks the temperature history at the location whose distance is 20.0 from the weld centerline. It can be found that the maximum temperature at the weld pool is 2100 C or so in both the pipe welds. This result agrees with the welding process in practice. It is worth noting that the peak temperature at the weld area of the duplex stainless steel pipe weld is higher compared to that of the austenitic stainless steel pipe weld in spite of the higher specific heat at the melting temperature (see Table 1), which is attributed to the fact that the difference between the melting
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Fig. 8. Residual stresses at locations with different circumferential angle q in the austenitic stainless steel girth-welded pipe: (a) axial stresses, inner surface, (b) axial stresses, outer surface, (c) hoop stresses, inner surface, (d) hoop stresses, outer surface.
temperatures of the stainless steel pipes is larger than that between the temperature increases due to the specific heats. Moreover, the cooling rate of the duplex stainless steel pipe weld is higher than that of the austenitic stainless steel pipe weld, which originates from the higher thermal conductivity of the duplex stainless steel pipe. Results are next presented for the comparison of the residual stress distributions induced by the girth welding of the austenitic and the duplex stainless steel pipe. Four different positions along the circumference are selected to portray welding residual stresses in order to examine the 3-D effects, i.e. circumferential variations of the residual stresses. The four positions have different circumferential angles from the welding start/stop position, which are 0 , 90 , 180 and 270 , respectively. From the stress analysis, all the stress and strain elements can be obtainable. Here, we will discuss only the relevant data. In this discussion, two words “axial” and “hoop” are used to denote the residual stress components. Axial residual stresses stand for the stress components which act normal to the weld line, whilst hoop residual stresses represent the stress components acting parallel to the weld line. The final residual stress distributions at locations with different circumferential angle in the austenitic stainless steel pipe weld are given in Fig. 8, with respect to axial distance from the weld centerline. Comparisons of the residual stresses in the girth-welded super duplex stainless steel pipe are also provided in Fig. 9. Fig. 8(a) and (b) depicts the axial residual stresses at the four locations on the inner and outer surfaces, respectively. From the simulated results, it can be observed that
within and near the weld area, the predicted axial residual stresses are tensile on the inner surface and compressive on the outer surface. For the case of thin-walled pipes, the heat deposited during girth welding is high enough to result in a uniform temperature increase through the pipe thickness at the weld region. Thus, the only deformation that will create thermal stresses is the circumferential strain due to the radial expansion and subsequent contraction. The circumferential shrinkage causes a local inward deformation in the vicinity of the welded zone, and thus generating a bending moment through the thickness. This leads to tensile axial residual stresses on the inner surface balanced by compressive stresses on the outer surface. Compressive axial residual stresses are formed on the inner surface away from the weld centerline, and tensile axial residual stresses on the outer surface for the self-equilibrating purpose. Moreover, it can also be found that even though the stress profiles along the four locations on both the inner and outer surfaces are similar to some extent, the magnitudes are different among them. It indicates that the axial residual stresses change with the circumferential angle. This is because the internal restraint during the pipe welding changes spatially due to the sequential deposition of the weld filler material as the welding arc moves along the circumference. Furthermore, the welding start/stop effects at the overlapping region make the variation severe. Similar trends can be noticed in the axial residual stress distributions of the duplex stainless steel girth-welded pipe. Fig. 8(c) and (d) portrays the hoop residual stresses at the four locations on the inner and outer surfaces, respectively. Regarding
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Fig. 9. Residual stresses at locations with different circumferential angle q in the duplex stainless steel girth-welded pipe: (a) axial stresses, inner surface, (b) axial stresses, outer surface, (c) hoop stresses, inner surface, (d) hoop stresses, outer surface.
the hoop residual stresses, their magnitude is influenced by the axial residual stresses. This explains why on the outer surface, which is experiencing axial compression, the hoop residual stresses are less tensile at the weld region and its vicinity compared to those on the inner surface. Stress reversal from tensile to compressive on the inner and outer surfaces away from the weld centerline is also identified. Furthermore, through careful observation of the results, it can be known that like the preceding axial residual stresses, spatial variations are present along the circumference. A rapid change of the residual stresses is also seen at the weld start/stop position. The hoop residual stress distributions in the girth-welded duplex stainless steel pipe shares similar characteristics. Therefore, it can be concluded that the axial and hoop residual stress distributions along the circumference are by no means axisymmetric, from which it is apparent that 3-D FE model is necessary to accurately capture the characteristics of welding residual stresses in the girth-welded stainless steel pipes. The results also unveil that the axial and hoop residual stresses at the welded zone in the duplex stainless steel pipe weld are higher compared to those in the girth-welded austenitic stainless steel pipe owing to the considerably higher yield stress of the duplex steel pipe. On the contrary, the stresses normalized by the yield stress are much higher in the austenitic stainless steel pipe weld in which the welding start/stop effects are more pronounced. It is also clear that the range subjected to tensile or compressive stress adjacent to the weld region is wider in the girth-welded austenitic stainless steel pipe. The variations in
outcomes between the dissimilar stainless steel pipe welds mainly arise from the differences in thermal conductivity, coefficient of thermal expansion and work hardening property. The larger coefficient of thermal expansion together with the much higher strain hardening rate of the austenitic stainless steel pipe gives rise to the higher stress ratios balanced by the higher reverse stress ratios away from the weld region. They also induce the steeper stress change at the overlapping area. In addition, the lower thermal conductivity in conjunction with the larger coefficient of thermal expansion of the austenitic stainless steel pipe contributes to the larger stress distribution region in and around the weld area. Results are finally given for the residual effective stresses for the girth-welded stainless steel pipes. Fig. 10(a) and (b) portrays the residual effective stresses in the austenitic stainless steel pipe weld at the four locations on the inner and outer surfaces with respect to axial distance from the weld centerline. It is worth noting that the maximum residual effective stress produced at the girth weld is 380 MPa or so, which is much below the ultimate strength of the austenitic stainless steel pipe, though the maximum axial and hoop residual stresses at the weld start/stop position are nearly 800 MPa. The effective stresses also demonstrate the 3-D effects within and near the welded zone. The higher effective stresses decrease as the distance from the weld centerline increases and finally converge to zero at the region far from the girth weld. The residual effective stresses in the girth-welded duplex stainless steel pipe are also presented in Fig. 11, which share similar characteristics.
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Fig. 10. Residual effective stresses at the four locations in the austenitic stainless steel pipe weld: (a) inner surface, (b) outer surface.
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Fig. 11. Residual effective stresses at the four locations in the duplex stainless steel pipe weld: (a) inner surface, (b) outer surface.
Acknowledgements 4. Conclusions In the present work, FE thermal simulation of the girth welding process was performed using a sequentially coupled 3-D thermomechanical FE analysis method whose effectiveness was verified by the comparisons with the experimental measurements to identify and compare temperature fields and welding residual stress distributions in girth-welded austenitic and duplex stainless steel pipes which have different thermal and mechanical properties. Based on the results, the following observations and conclusions can be drawn: a) 3-D FE model is apparently necessary to accurately predict the residual stress distributions in girth-welded austenitic and duplex stainless steel pipes which can incorporate the 3-D effects. b) Girth-welded duplex stainless steel pipes have higher peak temperature at the weld region and higher cooling rate compared to those developed during the girth welding of austenitic stainless steel pipes. c) The axial and hoop residual stresses normalized by the yield stress at the weld and its neighborhood are much higher in austenitic stainless steel pipe welds in which the welding start/stop effects are more significant, compared to those in duplex stainless steel pipe welds. Moreover, girth-welded austenitic stainless steel pipes have wider regions subjected to tensile or compressive stresses adjacent to the weld area.
This research was supported by Basic Science Research Program through the National Research Foundation of Korea (NRF) funded by the Ministry of Education (2012R1A1A2044895). References [1] A.-A. Iris, Duplex stainless steels: brief history and some recent alloys, Recent Pat. Mech. Eng. 1 (2008) 51e57. [2] J.J. del Coz Díaz, P. Menendez Rodriguez, P.J. Garcia Nieto, D. Castro-Fresno, Comparative analysis of TIG welding distortions between austenitic and duplex stainless steels by FEM, Appl. Therm. Eng. 30 (2010) 2448e2459. [3] A.-H.I. Mourad, A. Khourshid, T. Sharef, Gas tungsten arc and laser beam welding processes effects on duplex stainless steel 2205 properties, Mater. Sci. Eng. A 549 (2012) 105e113. [4] P.J. Withers, Residual stress and its role in failure, Rep. Prog. Phys. 70 (2007) 2211e2264. [5] H.D. Hibbitt, P.V. Marcal, A numerical thermo-mechanical model for the welding and subsequent loading of fabrication structure, Comput. Struct. 3 (1973) 1145e1174. [6] J. Goldak, M. Bibby, J. Moore, J.R. House, B. Patel, Computer modeling of heat flow in welds, Metall. Mater. Trans. B 17 (1986) 587e600. [7] R.I. Karlsson, B.L. Josefson, Three-dimensional finite element analysis of temperatures and stresses in a single-pass butt-welded pipe, J. Press. Vessel Technol. 112 (1990) 76e84. [8] L.-E. Lindgren, Finite element modelling and simulation of welding, part 1 increased complexity, J. Therm. Stresses 24 (2001) 141e192. [9] L.-E. Lindgren, Finite element modelling and simulation of welding, part 2 improved material modeling, J. Therm. Stresses 24 (2001) 195e231. [10] L.-E. Lindgren, Finite element modelling and simulation of welding, part 3 efficiency and integration, J. Therm. Stresses 24 (2001) 305e334. [11] J. Goldak, M. Akhlagi, Computational Welding Mechanics, Springer, 2005. [12] L.-E. Lindgren, Numerical modelling of welding, Comput. Meth. Appl. Mech. Eng. 195 (2006) 6710e6736.
150
C.-H. Lee, K.-H. Chang / Applied Thermal Engineering 63 (2014) 140e150
[13] L.-E. Lindgren, Computational Welding Mechanics (Thermomechanical and Microstructural Simulations), Woodhead Publishing, 2007. [14] M. Ashraf, L. Gardner, D.A. Nethercot, Finite element modelling of structural stainless steel cross-sections, Thin-Walled Struct. 44 (2006) 1048e1062. [15] E.F. Rybicki, D.W. Schmueser, R.W. Stonesifer, J.J. Groom, H.W. Mishaler, A finite element model for residual stresses and deflections in girth-butt welded pipes, J. Press. Vessel Technol. 100 (1978) 256e262. [16] E.F. Rybicki, P.A. McGuire, E. Merrick, J. Wert, The effect of pipe thickness on residual stresses due to girth welds, J. Press. Vessel Technol. 104 (1982) 204e209. [17] B. Brickstad, B.L. Josefson, A parametric study of residual stresses in multi-pass butt-welded stainless steel pipes, Int. J. Press. Vessels Pip. 75 (1998) 11e25. [18] M. Mochizuki, M. Hayashi, T. Hattori, Residual stress distribution depending on welding sequence in multi-pass welded joints with X-shaped groove, J. Press. Vessel Technol. 122 (2000) 27e32. [19] A. Yaghi, T.H. Hyde, A.A. Becker, W. Sun, J.A. Williams, Residual stress simulation in thin and thick-walled stainless steel pipe welds including pipe diameter effects, Int. J. Press. Vessels Pip. 83 (2006) 864e874. [20] D. Deng, H. Murakawa, W. Liang, Numerical and experimental investigations on welding residual stress in multi-pass butt-welded austenitic stainless steel pipe, Comput. Mater. Sci. 42 (2008) 234e244. [21] S. Fricke, E. Keim, J. Schmidt, Numerical weld modeling - a method for calculating weld-induced residual stresses, Nucl. Eng. Des. 206 (2001) 139e150. [22] P. Duranton, J. Devaux, V. Robin, P. Gilles, J.M. Bergheau, 3D modeling of multipass welding of a 316L stainless steel pipe, J. Mater. Process. Technol. 153e154 (2004) 457e463. [23] M. Abid, M. Siddique, Numerical simulation to study the effect of tack welds and root gap on welding deformations and residual stresses of a pipe-flange joint, Int. J. Press. Vessels Pip. 82 (2005) 860e871. [24] D. Deng, H. Murakawa, Numerical simulation of temperature field and residual stress in multi-pass welds in stainless steel pipe and comparison with experimental measurements, Comput. Mater. Sci. 37 (2006) 269e277. [25] I. Sattari-Far, M.R. Farahani, Effect of the weld groove shape and pass number on residual stresses in butt-welded pipes, Int. J. Press. Vessels Pip. 86 (2009) 723e731. [26] D. Deng, S. Kiyoshima, FEM prediction of welding residual stresses in a SUS304 girth-welded pipe with emphasis on stress distribution near weld start/end location, Comput. Mater. Sci. 50 (2010) 612e621.
[27] C.H. Lee, K.H. Chang, Influence of the residual stresses and distortions on the structural behavior of girth-welded cylindrical steel members, Constr. Build. Mater. 41 (2013) 766e776. [28] X. Jin, L. Huo, H. Yu, B. Bai, X. Li, J. Cao, Numerical simulation of parameter of residual stresses on butt-welded duplex stainless steel pipes, J. Tianjin Univ. 37 (2004) 116e125. [29] P. Dong, B.W. Brust, Welding residual stresses and effects on fracture in pressure vessel and piping components: a millennium review and beyond, J. Press. Vessel Technol. 122 (2000) 329e338. [30] C.H. Lee, A Study on the Mechanical Characteristics of High Strength Steel for the Application to the Steel Bridge (Ph.D. Thesis), Chung-Ang University, Korea, 2005. [31] D. Deng, W. Liang, H. Murakawa, Determination of welding deformation in fillet-welded joint by means of numerical simulation and comparison with experimental measurements, J. Mater. Process. Technol. 186 (2007) 219e225. [32] E. Pardo, D.C. Weckman, Prediction of weld pool and reinforcement dimensions of GMA welds using a finite element model, Metall. Mater. Trans. B 20 (1989) 937e947. [33] C.H. Lee, K.H. Chang, Prediction of residual stresses in high strength carbon steel pipe weld considering solid-state phase transformation effects, Comput. Struct. 89 (2011) 256e265. [34] K.J. Bathe, Finite Element Procedures, Prentice Hall, 1996. [35] D.S. Um, K.Y. Yoo, The experimental studies on residual stresses due to circumferential welds in thin steel cylinder, J. Korean Weld. Soc. 15 (1997) 149e156. [36] I. Sattari-Far, Y. Javadi, Influence of welding sequence on welding distortions in pipes, Int. J. Press. Vessels Pip. 85 (2008) 265e274. [37] Stainless Steels and Their Properties, 2013. http://www.hazmetal.com/f/kutu/ 1236776229.pdf. [38] KS D 0026, Method of Elevated Temperature Tensile Test for Steels and Heatresisting Alloys, Korean Standards, 2002. [39] Z. Barsoum, Residual stress analysis and fatigue of multi-pass welded tubular structures, Eng. Fail. Anal. 15 (2008) 863e874. [40] T.L. Teng, P.H. Chang, W.C. Tseng, Effect of welding sequence on residual stresses, Comput. Struct. 81 (2003) 273e286. [41] B. Taljat, B. Radhakrishnan, T. Zacharia, Numerical analysis of GTA welding process with emphasis on post-solidification phase transformation effects on the residual stresses, Mater. Sci. Eng. A 246 (1998) 45e54.