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Numerical simulation to study the effect of tack welds and root gap
on welding deformations and residual stresses of a pipe-flange joint
M. Abida,*, M. Siddiqueb
aAssistant Professor, Faculty of Mechanical Engineering, GIK Institute of Engineering Sciences and Technology, Topi, NWFP, PakistanbGraduate Student, Faculty of Mechanical Engineering, GIK Institute of Engineering Sciences and Technology, Topi, NWFP, Pakistan
Received 23 February 2005; revised 23 June 2005; accepted 23 June 2005
Abstract
This paper presents a three dimensional sequentially coupled non-linear transient thermo-mechanical analysis to investigate the effect of
tack weld positions and root gap on welding distortions and residual stresses in a pipe-flange joint. Single-pass MIG welding for a single V
butt-weld joint geometry of a 100 mm diameter pipe with compatible weld-neck ANSI flange class # 300 of low carbon steel is simulated.
Two tack welds at circumferentially opposite locations, with the crucial effect of the tack welds orientation from the weld start position is the
focus in this study. Four different angular positions of tack welds (0 and 1808, 45 and 2258, 90 and 2708, 135 and 3158) are analyzed. In
addition, four cases for root gaps (0.8, 1.2, 1.6, 2.0 mm) are considered and computational results are compared. A basic FE model is also
validated with experimental data for temperature distribution and deformations. From the results, the axial displacement and tilt of the flange
face are found to be strongly dependent on the tack weld orientation and weakly dependent on the root gap.
q 2005 Elsevier Ltd. All rights reserved.
Keywords: FEA Simulation; Tack welds; Root gap; Welding deformation; Pipe-flange joint; Residual stresses
1. Introduction
Welded pipe-flange joints are widely used in a variety
of engineering applications such as oil and gas pipelines,
nuclear and thermal power plants and chemical plants.
A non-uniform temperature field, applied during welding,
produces deformation and residual stresses in welded
structures. For flange joints, any tilt or out-of-plane
deformation in the flange face results in gasket damage
[1,2]. In addition, uneven bolt-up loads consequently
produce adverse effects on the service life of the joint.
Residual stresses in a piping system may have a largercontribution to the total stress field compared to the
applied loadings while assessing the risks for defect
growth and static fracture in piping systems with brittle
fracture behaviour [3]. Therefore, realistic prediction of
contributing factors is of vital importance. The extent of
deformations and residual stresses in welded components
depends on several factors such as geometrical size,
welding parameters, weld pass sequence and applied
structural boundary conditions.
Finite Element (FE) simulation has become a popular
tool for the prediction of welding distortions and residual
stresses. A substantial amount of simulation and experi-
mental work focusing on circumferential welding
with emphasis on pipe welding is available in the
literature [312]. To reduce computational power require-
ments, assumptions such as rotational symmetry and
lateral symmetry have been employed in numerical
simulations [46]. These assumptions reduce the compu-
tational demand but make the problem over simplified by
limiting the analysis to one section of the complete
geometry and eliminate modeling of root gap and tack
welds. Therefore, these models are not capable of
predicting the effects of weld start/stop location, root
gap and tack welds. Brickstad and Josefson [3] presented
a parametric study of multi-pass butt welded pipes in
which both sides of the weldment are modeled but due to
the assumption of rotational symmetry the tack welds are
ignored. In the available 3D FE studies of pipe welding,
International Journal of Pressure Vessels and Piping 82 (2005) 860871
www.elsevier.com/locate/ijpvp
0308-0161/$ - see front matter q 2005 Elsevier Ltd. All rights reserved.
doi:10.1016/j.ijpvp.2005.06.008
*Corresponding author. Tel.: C92 938 71858x2293; fax: C92 938
71889.
E-mail addresses: abid@giki.edu.pk (M. Abid), mabid00@hotmail.
com (M. Abid), gme0102@giki.edu.pk (M. Siddique).
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only half models (with assumption of lateral symmetry)
without tack welds are analyzed. Fricke et al. [10]
investigated multi-pass welding on a complete 3D model
for pipe weld but nothing is mentioned about tack welds.
Siddique et al. [11] used a 3D model for welding of
pipe-flange joints with initial tacks but no further detail
about modeling of tacks is provided.
The issue of tack welds is addressed in FE simulations
of butt welded plates. Jonsson et al. [12,13], using a
plane stress simplification, described the influence of tack
welding sequence and subsequently compared platemotion and thermal stresses of root-bead and single-
pass butt-welding of tacked plates. Shibahara et al. [14]
examined the effect of tack welds and root gap in a butt-
welded plate by using temperature dependent interface
elements. In these studies [1214] it is concluded that
tack welding sequence, their interspacing and subsequent
butt-welding have a significant effect on root opening
and transverse shrinkage. Jang et al. [15], by using a
plane strain assumption, concluded that root gap has
some effect on symmetrically distributed residual stresses
across the weld.
2. Present study
The effect of tack orientation in girth welding of pipe
flange joints, especially for small diameter joints such as
100 mm nominal diameter pipe, is believed to be
significant because there are only two tacks and the
time interval between reheating/remelting of successive
tacks is very small. This paper presents a parametric
study to determine the effect of tack weld locations with
respect to weld start position and the effect of root
opening on welding deformations and residual stresses.
3D FE simulation of a single pass butt weld joint
geometry is performed using ANSYS [16]. A low carbon
steel pipe of 115 mm outer diameter, 6 mm wall
thickness (Ri/tZ8.583) and 200 mm length is welded
with a 100 mm nominal diameter weld-neck type ANSI
class # 300 flange. The joint configuration is shown
schematically in Fig. 1. A total of seven cases has been
formulated and analyzed, see Table 1. The basic FE
model, with 1.2 mm root gap and two tack welds at 90 8
and 2708 from the weld start position, for the single-pass
single-V butt-weld joint geometry is validated exper-imentally. The manufacturing stress of components and
the initial effect of tack welds on distortion and residual
stresses are neglected.
3. Experimental setup
For automatic circumferential welding of a pipe flange
joint, a DC powered conventional lathe with open loop
continuous speed controller is synchronized with a welding
power source. Synchronization is achieved through an
Fig. 1. Pipe-flange joint configuration.
Table 1
Details of FE studies performed
Sr. No Identification Tack weld
Location (8)
Root Gap (mm)
1 Tack 0-180 0, 180 1.2
2 Tack 45-225 45, 225 1.2
3 Tack 90-270
(or) Root 1.2
90, 270 1.2
4 Tack 135-315 135, 315 1.2
5 Root 0.8 90, 270 0.8
6 Root 1.6 90, 270 1.6
7 Root 2.0 90, 270 2.0
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interface controller operated by limit switches, mounted on
the lathe chuck, indicating weld start and end positions. The
tacked sample of the pipe-flange joint is rotated in the chuck
of the lathe, while the welding torch is held stationary by
mounting it in a special fixture. This torch mounting fixture,
in combination with machine carriage, provides 4-axis
(3 translational and one rotational) adjustments to the torch.
The automatic circumferential welding facility, used in the
present study, is shown inFig. 2.
A metal inert gas (MIG) welding process with gross heat
input of 792 KJ/m is used. In the absence of a weaving
facility, a forehand welding technique (having torch angle
17.58 with the normal) is used to control penetration and
avoid blow off. Dimensions of the physical specimen are in
accordance with the size and geometry described in Section
2. The material for pipe and flange is carbon-manganese
steel with chemical composition 0.18w0.22% C,
0.6w1.05% Mn, 0.2w0.26% Si, 0.1w0.2% Cr,! 0.05%
S and!0.05% P. Filler metal is ER70S-6 Carbon Steel wire
of diameter 1.14 mm (0.045). A single pass butt-weld joint
geometry with a 6 mm deep single V-groove (608
includedangle) and a 1.2 mm root gap is used. The weld joint
contains two initial tack-welds at angular positions of 908
and 2708 from the weld start position. Each tack weld is
machined to a length of 10 mm and thickness of 3.0 mm.
Subsequently, the sample is stress relieved to remove
manufacturing and tack-welding stresses from the pipe and
flange, as existing residual stresses affects thermal expan-
sion behaviour [17]. A mechanical dial indicator of 1mm
resolution and 2mm accuracy is used for in-situ measure-
ment of axial displacement on the flange face, introduced
during welding.
4. Material model
Material modelling has always been a critical issue in the
simulation of welding because of the scarcity of material
data at elevated temperatures. Some simplifications and
approximations are usually introduced to cope with this
problem. These simplifications are necessary due to both
lack of data and numerical problems when trying to model
the actual high-temperature behaviour of the material [18].
The detailed material model for the material described
above is not available in the literature; therefore material
data available for a similar composition, i.e. 0.18% C, 1.3%
Mn, 0.3% Si, 0.3% Cr, 0.4% Cu (Swedish standard steel SIS
2172), is used from Karlsson and Josefson[9]. Though there
is a minor difference in the chemical composition of the two
materials, however, such a difference may not have
significant effect on the thermal and mechanical material
properties. This approximation seems justified for para-
metric comparative studies because material behaviour
contributes equally in the results of all cases and differences
in structural response can be attributed to changes in theparameters.
The pipe, flange and filler metal are supposed to be of the
same chemical composition. Karlsson and Josefson
collected temperature dependent material properties from
previously published literature. They used specific heat
formulation and accounted for latent heat for solid-state and
solid-liquid phase transformations. In the mechanical
material properties, microstructural evolution is accounted
for by defining different thermal dilatations and yield
strengths for different zones in the domain depending on the
peak temperatures reached in a particular point during
Fig. 2. Experimental setup.
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the thermal cycle. Most of the plastic strains are formed at
high temperature and the low alloy steel shows nearly
ideally plastic behaviour at temperatures above 1073 K.
Furthermore, it is argued that plastic strains accumulated
before the final solid state phase transformation to a large
extent are relieved during the transformation. For these
reasons, the material in the model was assumed to be elastic-ideally plastic (without hardening).
In the present work, an enthalpy formulation is used
instead of specific heat and latent heats are evenly
distributed over the respective temperature ranges. Further-
more, the thermal conductivity is given an artificial increase
to 230 W/mK above the melting temperature to incorporate
stirring effects in the weld pool. As suggested in [7],
volumetric changes associated with solid sate phase
transformation are ignored in the absence of transformation
induced plasticity effects because they produce compressive
hoop stresses near the weld centreline which is contrary to
the experimental measurements given in [19]. The
suggested changes in the material model are discussed in
more detail in[20].
5. Analysis procedure
Taking advantage of the weak structure to thermal field
couplings, the problem is formulated as a sequentially
coupled thermal stress analysis. Firstly a non-linear
transient thermal analysis is performed to predict the
temperature history of the whole domain. Subsequently
the results of the thermal analysis are applied as a thermal
body load in a non-linear transient structural analysis tocalculate deformation and stresses. The finite element
model for both thermal and structural analysis is the same
except for element type. During the analysis a full Newton-
Raphson (NR) iterative solution technique with direct
sparse matrix solver is employed for obtaining a solution.
During the thermal cycle, temperature and consequently
temperature dependent material properties change very
rapidly. Thus, full NR, which uses a modified material
properties table and reformulates the stiffness matrix at
every iteration is believed to give more accurate results. The
line search option of the FE code ANSYS [16]is set to ON
to improve convergence. A single point reduced integration
scheme with hour glass control is implemented to facilitate
convergence, and to avoid excessive locking during
structural analysis.
A conventional quiet element technique named element
birth and death [21], is used for modeling of the filler
material. A complete FE model is generated in the start;
however, all elements representing filler metal except
elements for the tack welds are deactivated by assigning
them very low stiffness. During the thermal analysis, all the
nodes of deactivated elements (excluding those shared with
the base metal) are also fixed at room temperature till the
birth of the respective element. Deactivated elements are re-
activated sequentially when they come under the influence
of the welding torch. For the subsequent structural analysis,
birth of an element takes place at the solidification
temperature. Melting and ambient temperatures are set as
reference temperatures (temperature at which thermal strain
is zero) for thermal expansion coefficients of filler and base
metals, respectively. To avoid excessive distortion, initialstrain in the elements is set to zero at the time of element
reactivation.
For thermal analysis, the total welding time of the
complete circumferential weld, i.e. 58 s, is divided into 144
equally spaced solution steps. Each step is further divided
into two sub-steps, which effectively reduces the load
application time to 0.201 s. A stepped load option is used for
realistic application of the thermal load. After extinguishing
the arc, another 56 load steps of different time duration are
used for cooling of the weldment. It took about 52 min. to
return to the ambient temperature of 27 8C. Load step time
in the structural analysis is kept equal to the thermal
analysis. However, each load step is solved in a single sub-
step except for cases of numerical non-convergence. The
restart option of the software with corrected sub-step setting
is effectively used to handle non-convergences. Total CPU
time remained approximately 5 hrs and 100 hrs for the
thermal and structural analysis respectively on an IBM
compatible P-IV 2 GHz PC with 1 GB RAM.
6. FE model
Four finite element models, with minor changes for
different studies, representing the same physical geometryare developed in ANSYS. Being away from the zone of
interest, bolt holes are not included in the models and it is
assumed that this geometrical simplification will have no
significant effect on distortions and residual stresses. Eight-
node brick elements with linear shape functions are mostly
used in the model. Linear elements are preferred because, in
general, favours more lower-order elements than fewer
higher-order elements in non-linear problems [22]. The
basic FE model, used for all the cases of tack weld positions,
is shown inFig. 3(a). This model consists of 25488 nodes
and associated 21456 linear elements, out of which 12960
elements are used for the flange and the other 8496 elements
represent the pipe. The other three models, used for root gap
studies, are similar to the above described model except for
the element sizes at the root gap position.
In order to facilitate data mapping between thermal and
structural analysis, the same FE model is used with
respective element types. For the thermal analysis the
element type is SOLID70 which has single degree of
freedom, temperature, on its each node. For structural
analysis the element type is SOLID45 with three transla-
tional degrees of freedom at each node. Due to anticipated
high temperature and stress gradients near the weld, a
relatively fine mesh is used in a distance of 10 mm on both
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sides of the weld centreline. Element size increases
progressively with distance from the weld centreline. The
mesh refinement scheme with approximate V-groove
formation and tack weld is shown in Fig. 3(b).
7. Boundary conditions
In the thermal analysis, both radiation and surface
convection are considered for realistic modeling of heat
loss from the surface. During the thermal cycle, radiation
dominates over surface convection in areas adjacent to the
weld pool; whereas, away from the weld pool convection is
the primary mechanism of heat loss from the body. Instead
of modeling convection and radiation separately, a
combined heat transfer coefficient, as used in [23], is
calculated by using.
~hZ3emsbolTC2734KTambC2734
TKTambChcon (1)
where ~h,3em, hcon,sbol, T and Tambrepresent combined heat
transfer coefficient, emissivity, convective heat transfer
coefficient, Stefan-Boltzmann constant, instantaneous body
temperature and ambient temperature, respectively. The
calculated combined heat transfer coefficient was applied on
all areas exposed to the ambient air, as shown in a sectioned
view inFig. 4.The ambient temperature (27 8C) is taken as
the initial condition for the entire mass involved. During
structural analysis, the only constraint applied is represen-tation of clamping in the machine chuck, as shown in Fig. 2.
For this purpose, all the nodes of the far end of the pipe, in
Cartesian coordinate axes, are constrained in the axial and
radial directions.
8. Heat source modeling
Proper modeling of heat flow from the welding torch to
the weldment is quite crucial as it controls the application of
thermal load which consequently produces distortion and
residual stresses in the weldment. For the determination of
the weld pool size and shape, a section of the weld is cut,
polished, chemically etched and scanned. This cross-
sectional metallographic data revealed the so called hot
top nail head configuration of the weld pool. Thisconfiguration is difficult to achieve by using a conventional
double ellipsoidal heat source model by Goldak et al. [24].
However, for such cases Goldak et al.[25]suggest the use of
superimposed four ellipsoid quadrants (compound double
ellipsoid model) for better results. In the present study, the
authors used a modified double ellipsoidal scheme.
The governing equations for power density distribution in
the front and rear ellipsoids of a 3D model are as follows:
qfZ6ffiffiffi
3p
My;zQffp ffiffiffipp
afbceK3
rq2a2f
Cz2
b2C
RoKr2c2
(2)
qrZ6ffiffiffi
3p
My;zQfrpffiffiffip
p arbc
eK3
rq2a2r
Cz2
b2C
RoKr2c2
h i (3)
where,
QZVI; ffCfrZ2; hZ
PniZ1
qiVi
Q
The description and numerical values for different
variables in the power density distribution equations are
Fig. 3. (a). 3D FE model (b). Mesh refinement, V-groove, tack weld and root gap.
FlangePipe Weld Bead
Fig. 4. A sectioned view of pipe flange joint with combined convection and
radiation (indicated with arrows) from the surfaces exposed to air.
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given in Table 2. M(y,z) in the above equations is a scalar
multiplier which is used to modify the shape of the weld
pool and is a function of spatial location in the axial and
radial directions. Its initial values are selected arbitrarily and
readjusted iteratively to match the weld pool shape. Final
values ofM(y,z)are shown inFig. 5. Numerical values usedfor other variables in the power density distribution
equations are given above.
For calculation of spatial heat distribution using
equations (2) and (3), the origin of the coordinate system
is located at the centre of the moving arc and movement of
the heat source is achieved through a user sub-routine.
Another subroutine is used to calculate instantaneous
centroidal distances of elements from the moving arc
centre. To describe the heat source size, five elements in the
front and four elements in the rear of the heat source are
taken in the direction of weld torch motion. Across the weld
line, heat is given to five elements on each side. The heatinput from the moving arc to the elements is modeled as
volumetric heat generation, as this has an additional
advantage that surface convection can be applied to the
same elements without defining 2D-elements, required
otherwise. It is also assumed that the intensity of the heat
source is independent of time. In order to validate the
thermal model, the etched sample is used to reveal liquidus
isotherms at 17888K, representing the fusion zone (FZ), and
outer HAZ isotherms at 10838K. Comparison of measured
and simulation isotherms, at a section 1808 from the weld
start position, shows good agreement, Fig. 6.
9. Results and discussion
9.1. Effect of tack position
9.1.1. Effect on welding distortions
Tack welds are used to restrain excessive transverse
shrinkage and to maintain the root gap. The size and
location of tacks with respect to the weld start point can alter
the resistance offered by the tacks. This can have a dominant
effect on transverse shrinkage and resultant flange face
displacement. In the present work, only the effect of tacklocation is analyzed by keeping the tack size unchanged.
Immediately after the initiation of the arc, thermal
expansion of metal beneath the moving arc is the source of
structural distortions. As the arc proceeds, contraction of the
solidifying weld bead behind the arc becomes another
Axial Distance from Weld Centerline (mm)
RadialDistancefromO
uterSurface(mm)
Fig. 5. Values of scalar multiplier M(y,z)as functions of spatial location in axial and radial directions.
Table 2
Description and numerical values for different variables used in power
density distribution equations for heat source modeling
Symbol Description Value
af Front Ellipsoidal semi-axes length (mm) 12.9
ar Rare Ellipsoidal semi-axes length (mm) 10.3
b Half width of arc (mm) 5.0
C Depth of arc (mm) 6.0
ff Fraction of heat deposited in front 1.55
fr Fraction of heat deposited in rare 0.45
I Welding current (Amp) 225
M (y,z) Scalar Multiplier
N Total number of element under torch
influence
qi Power density for ith element (W/mm3)
qf Power density in front ellipsoid (W/mm3)
qr Power density in rear ellipsoid (W/mm3)
Q Total Arc heat (W) 4950
R Radius of pipe (mm)
Ro Pipe outer radius (mm) 57.5
v Welding speed (mm/s) 6.25
V Voltage (Volt) 22Vi Volume of ith element (mm3)
Z Distances from the torch centre in axial
direction (mm)
q Angle from instantaneous arc position
(Radian)
h Arc efficiency 85%
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dominant source of distortions. Until the welding torch
reaches the first tack, both the tacks collectively restrain
flange motion thus minimizing change in root gap. When the
first tack is heated by the arc, its resistance against thetransient forces gradually vanishes with increase in
temperature. Thereafter the second tack and solidified
weld metal behind the welding torch, if it has cooled to a
substantially low temperature, resist the transient forces.
The time of first tack reheating after arc initiation is critical
since if it is too short, weld metal behind the moving arc will
not contribute significantly and thus the second tack alone
may not effectively resist these forces. Consequently there
will be higher axial displacement on the flange face. Results
of axial displacement (AD) on the flange face at a radius of
117.3 mm and resulting face tilt, calculated by using
equation (4), are shown inFig. 7a, 7b respectively.
TiltZ tanK1 Max:ADKMin:AD
2!117:3
(4)
A maximum axial displacement of 1.156 mm with a face
tilt of 0.398 (with the initial plane) is observed for Tack
0-180. Being the first tack at zero degree, it is directly
exposed to the welding torch on arc initiation (no weld seam
exists behind the arc yet) and hence results in large axial
displacement. The next highest axial displacement of
0.78 mm with a face tilt of 0.1678
is found in the case ofTack 45-225, whereas, in the other two cases i.e. Tack 90-
270 and Tack 135-315 axial displacements are 0.66 and
0.64 mm respectively with corresponding face tilts of 0.085
and 0.0998. Minimum face tilt is found for Tack 90270
which indicates that the time taken by the arc to travel
through 908(from weld start position), i.e. 14 s, is sufficient
for solidification of the preceding weld bead and the
solidified weld bead is stiff enough to attenuate the effect of
reheating/re-melting of the tack. Changing tack weld
position from 908 to 1358has not contributed significantly
in the axial displacement or tilt. However, an inverted
displacement pattern is produced. Therefore, the mostappropriate location of the first tack, for the joint size
under discussion, is concluded between 908and 1358.
Comparison of transient axial displacement of two nodes
on the flange face at a radius of 117.3 mm and angular
positions of 90 and 2708 respectively for Tack 0180 and
Tack 135315,(Fig. 8a) shows that transient displacements
(a) (b)
0.6
0.4
0.2
0.0
0.2
0.4
0.60.8
1.0
1.2
1.4
0 60 120 180 240 300 360
Hoop Coordinate (Deg)
AxialDisplacement(mm)
Tack 0-180Tack 45-225
Tack 90-270Tack 135-315Exp
0.00
0.10
0.20
0.30
0.40
0.50
Tack
0-180
Tack
45-225
Tack
90-270
Tack
135-315
FlangeFaceTilt(Deg)
Fig. 7. (a) Comparison of flange face axial displacement, representing lateral shrinkage (b). Resulting flange face tilt.
0
1.5
3
4.5
6
10 9 8 7 6 5 4 3 2 1 0 1 2 3 4 5 6 7 8 9 10
Distance from Weld CL (mm)
PipeThic
kness(mm)
Measured FZ Measured HAZ FEFZ FEHAZ
Fig. 6. Comparison of measured and simulation isotherms.
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forTack 0180are much larger thanTack 135315and areconcluded to be due to immediate reheating of the first tack
in the former case. In addition, almost reflective nodal
motion (in opposite directions) in both cases has been
observed which indicates flange face tilt. The largest
contribution to the flange face tilt is found between 020
sec after the arc initiation (about 1208of the arc travel), after
which the increase in tilt is slow (Fig. 8b).
9.1.2. Effect on residual stresses
Variation of axial residual stresses in the hoop direction
at the weld centreline on both inner and outer surfaces is
shown inFig. 9.In general, axial residual stresses are tensile
on the inner surface and compressive on the outer surface
with very strong influence of weld start/stop positions.
Away from the weld start/stop position, a slight decreasing
trend in tensile stress on the inner surface and a slight
increasing trend in compressive stresses on the outer
surface, in the welding direction, are observed. The stressprofile is almost identical in all the four cases except at the
positions of tacks. In every case, significant localized stress
reduction on the inner surface is found at corresponding tack
positions, whilst localized stress increase in compressive
residual stresses on the outer surface is observed.
In order to investigate the mechanism for stress
variations at tacked locations, transient stress variation at
two points; at angular positions of 1358(point on weld bead-
Node 8198) and 1808 (on tack-Node 8171) at the inner
surface forTack 0180is presented inFig. 10a. Being on the
weld bead, node 8198 remains deactivated and stress free in
structural analysis during heating and subsequent cooling tothe solidus temperature, 1738 K (Fig. 10b). Stress
accumulation is not significant at elevated temperature
above 1052 K due to the very low yield strength. A tensile
transient axial stress of 100 MPa is observed at
a temperature of 663 K, below which stress increases
300
200
100
0
100
200
300
400
0 45 90 135 180 225 270 315 360
Hoop Coordinates (Deg)
AxialStresses(MPa)
Tack 0-180 Tack 45-225
Tack 90-270 Tack 135-315
Inner Surface
Outer Surface
Fig. 9. Axial residual stress variation in hoop direction at weld centreline on outer and inner surfaces.
(a) (b)
2
1.5
1
0.5
0
0.5
1
0 15 30 45 60
Time(sec)
AxialDisplacem
ent(mm)
Tack 0-180 (90)
Tack 0-180 (270)
Tack 135-315 (90)
Tack 135-315 (270)
135
180
315
90
0
0.1
0.2
0.3
0.4
0.5
0.6
0 15 30 45 60
Time (sec)
FlangeFaceT
ilt(Deg)
Tack 0-180
Tack 135-315
Fig. 8. (a) Comparison of transient axial displacement of nodes on flange face at 90 and 2708from weld start position. (b) Transient flange face tilt.
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rapidly due to the rapid increase of yield strength.
Accumulation of tensile residual stresses on cooling is
analogous to the modified Wells model for thermal/stress
cycle, described in Lin and Chou[26]. Though this model is
primarily for an element of material near the fusion zone, it
is found to be suitable to describe structural response of the
cooling weld bead.
On the other hand, node 8171 belongs to the tack and by
virtue of the peak temperature in thermal cycle it belongs to
the heat affected zone. As the torch proceeds after initiation
of the arc, cooling weldbead behind the torch causes stress
accumulation on the tack. Axial stress is initially tensile
which turns to compressive as the torch approaches the tack
and a stress ofK66 MPa is found just prior to heating. On
heating, the stress increases to K166 MPa at 700 K, beyond
which it starts decreasing and becomes zero at about 1050 K
due to decrease in yield strength. In the beginning of the
subsequent cooling the structural response is quite different
from the prediction using the modified Wells model.
Cooling below 1027 K causes generation of compressive
stresses instead of tensile and is concluded to be a major
cause of stress reduction at the tack location. The unusual
differential temperature distribution on the tack produces
positive thermal strain which when restrained by surround-
ing material causes negative elastic strain and then a
dominant negative plastic strain, as shown in Fig. 11a,
Fig. 11b. By comparing Fig. 10a and Fig. 11b, negative
elastic axial strain is found as the basic reason for these
compressive stresses. With further decrease in temperature,
the response is once again in an opposite direction
0.08
0.06
0.04
0.02
0
0.02
0.04
0.06
0.08
0.1
0.12
20 25 30 35 40 45 50
Time (sec)
Strain
Plastic 135
Thermal 135
Plastic 180
Thermal 180
1
0.5
0
0.5
1
1.5
0 25 50 75 100 125 150
Time (sec)
Strain(x
103)
Elastic 135
Elastic 180
Fig. 11. Transient strain on the inner surface at angular positions of 135 and 1808from the weld start position forTack 0-180(a) Thermal and plastic strain (b)
Elastic strain.
200
150
100
50
0
50
100
150
200
250
300
350
0 25 50 75 100 125 150
Time(sec)
AxialStress(MPa)
135 Deg (8198)
180 Deg (8171)
200
150
100
50
0
50
100150
200
250
300
350
400
300 700 1100 1500 1900
Temperature (Deg K)
AxialStress(
MPa)
135 Deg (8198)
180 Deg (8171)
Fig. 10. Axial stress variation on the inner surface at angular positions of 135 and 180 8from weld start position forTack 0-180(a) Transient response (b) As a
function of temperature.
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producing tensile stress and follows the material character-istic tensile yield strength curve after the elastic limit.
Axial stress variation in the axial direction on the inner
surface at an angle of 908 from the weld start position is
found identical in all cases except for case Tack 90-270
(Fig. 12). As the stresses are relatively lower for the case in
which the tack exists at 908 (on the section under
observation) therefore, it is concluded that the tack serves
as a stress reducer in its close proximity. Variation of hoop
residual stresses in the hoop direction at the weld centreline
on both inner and outer surfaces is shown inFig. 13. A weld
start/end effect is pronounced in hoop residual stresses and
is dominant in the start side as compared to the end.Similarly the tacks serve as stress raisers though the effect is
not as significant. Hoop stresses are tensile on both inner
and outer surfaces and can fairly well be approximated as
axisymmetric, if the weld start effect and the effect of tack
welds are ignored.
9.2. Effect of root gap
Four cases for different root gaps are analyzed to study
the effect of root gap on welding distortions and residual
stress distributions. All the other parameters including tack
weld positions, heat inputs, thermal and structural boundary
conditions etc. are kept the same in all the four cases. Axial
displacements on the flange face at a radius of 117.3 mm for
all the cases are compared inFig. 14. It is concluded that
root gap less than 1.2 mm does not have any significant
effect on axial deformation and flange face tilt. On the other
hand, axial deformation and flange face tilt increase
significantly with increase in root gap from 1.2 mm to2.0 mm. The stiffness of a column can be described with
the relation:
KZAE
L (5)
300
200
100
0
100
200
300
400
50 40 30 20 10 0 10 20 30 40 50
Distance from Weld Centerline (mm)
AxialStre
ss(MPa)
Tack 0-180
Tack 45-225
Tack 90-270
Tack 135-315
Fig. 12. Axial residual stress variation in axial direction at a section 908from weld start position.
50
0
50
100
150
200
250
300
350
0 45 90 135 180 225 270 315 360
Hoop Coordinates (Deg)
HoopStress(MPa)
Tack 0-180 Tack 45-225
Tack 90-270 Tack 135-315
For Inner Surface
For Outer Surface
Fig. 13. Hoop residual stress variation in hoop direction at weld centerline on inner and outer surfaces.
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where,Kis the stiffness,Eis Youngs modulus andAand Lare the cross sectional area and length of the column
respectively. Treating the tack as a column, the stiffness of
the tack weld is found to be inversely proportional to the
axial length of the tack. Axial length of the tack increases
with root gap and thus tack stiffness decreases. A tack with
lower stiffness gives higher deformations under the same set
of transient forces. On the other hand change in root gap
does not have any impact on the residual stress distribution,
provided the other parameters such as heat input etc. are
kept unchanged.
10. Conclusion
From the results it is concluded that a change in tack weld
location alters the axial displacement and tilt of the flange
face. Furthermore it is concluded that the first tack weld
should at least be at some distance from the weld start point
and for 100 mm nominal diameter pipe most appropriate
positions for tack welds are 90 and 2708from the weld start
point. Tack weld location has no significant effect on overall
residual stress distribution, but a localized effect is
experienced in terms of a stress raiser for axial and hoop
stresses on both inner and outer surfaces except axial residual
stresses on the inner surface, where it serves as a stressreducer. Regarding root gap opening it is concluded that root
gap should be a minimum, just to meet the need of weld
penetration. A large root gap increases the lateral shrinkage
and results in large axial displacement and flange face tilt.
References
[1] Abid, M., and Nash, D.H., Risk Assessment studies for Gasketed and
Non-gasketed Bolted Pipe Joints, International Pipeline Conference
(IPC2002), Calgary Canada, Sep 29 w Oct 3 2002. Proceedings of
IPC2002/IPC-27386 2002; 1-11.
[2] Nash DH, Abid M. Surface sensitivity study of non-gasketed flangejoint. J Process Mech Eng; Part-E 2004;E4:218.
[3] Brickstad B, Josefson BL. A parametric Study of Residual Stresses in
Multi-pass Butt-Welded stainless steel pipes. Int J Pres Vessels Piping
1998;75:1125.
[4] Rybicki EF, Schmueser DW, Stonesifer RW, Groom JJ,
Mishaler HW. A Finite Element Model for Residual Stresses and
Deflections in Girth-Butt Welded Pipes. Trans ASME J Press Vessel
Technol 1978;100:25662.
[5] Rybicki EF, McGuire PA, Merrick E, Wert J. The Effect of Pipe
Thickness on Residual Stresses due to Girth Welds. ASME J Press
Vessel Technol 1982;104:2049.
[6] Rybicki EF, Stonesifer RB. Computation of residual stresses due to
multipass welds in piping system. Trans ASME J Press Vessel
Technol 1979;101:14954.
[7] Karlsson L, Jonsson M, Lindgren LE, Nasstrom M, Troive L.Residual stresses and deformations in a welded thin-walled pipe.
In: Rybicki E, Shiratori E, Widera GEO, Miyoshi T, editors.
ASME pressure wessels and piping conference 1989. Weld
residual stresses and plastic deformation, PVP-vol. 173. Hawai:
Honolulu; 1989. p. 714.
[8] Dong Y, Hong J, Tasi C, Dong P. Finite element modeling of residual
stresses in Austenitic stainless steel pipe girth welds. AWS Weld J,
Weld Res Suppl 1997;442:449444.
[9] Karlsson RI, Josefson BL. Three dimensional finite element analysis
of temperature and stresses in single-pass butt-welded pipe. Trans
ASME J Press Vessel Technol 1990;112:7684.
[10] Fricke S, Keim E, Schmidt J. Numerical weld modeling-a method for
calculating weld-induced residual stresses. Nucl Eng Des 2001;206:
13950.
[11] Siddique M, Abid M, Junejo HF, Mufti RA. 3-D finite element
simulation of welding residual stresses in pipe-flange joints: effect of
welding parameters. Mater Sci Forum 2005;490491:7984.
[12] Jonsson M, Karlsson L, Lindgren LE. Simulation of tack welding
procedures in Butt joint welding of plates. AWS Weld J, Weld Res
Suppl 1995;296s3301.
[13] Jonsson M, Karlsson L, Lindgren LE. Plate motion and thermal
stresses in root-bead Butt-welding of plates. In: Lewis RW,
Morgan K, editors. Numerical methods in heat transfer, vol. III.
New York: Wiley; 1995.
[14] Shibahara M, Serizawa H, Murakawa H. Finite element analysis
using interface elements for predicting deformation during Butt
welding considering root gap and tack welds. Trans JWRI 2002;
31(1):6370.
0.0
0.1
0.2
0.3
0.4
0.5
0.6
0.7
0.8
0 60 120 180 240 300 360
Hoop Coordinate (Deg)
AxialDisplacement(mm)
Root 0.8 Root 1.2
Root 1.6 Root 2.0
Fig. 14. Comparison of flange face axial displacements, representing lateral shrinkage.
M. Abid, M. Siddique / International Journal of Pressure Vessels and Piping 82 (2005) 860871870
7/27/2019 Numerical simulation to study the effect of tack welds and root gap on welding deformations and residual stresses of a pipe-flange joint
12/12
[15] Jang GB, Kim HK, Kang SS. The effects of root opening on
mechanical properties, deformation and residual stresses of weld-
ments. AWS Weld J, Weld Res Suppl 2001;80s88.
[16] ANSYS Users Manual, ANSYS Users Manual, SAS IP inc., 1998.
[17] Wang X, Hoffmann C, Hsueh C, Sarma G, Hubbard C. Influence of
residual stresses on thermal expansion behavior. Appl Phys Lett 1999;
75(21):32946.
[18] Lindgren LE. Finite element modeling and simulation of welding part
2: improved material modeling. J Thermal Stress 2001;24:195231.
[19] Jonsson M, Josefson BL. Experimentally determined transient and
residual stresses in the Butt-welded pipes. J Strain Anal 1988;
23(1):2531.
[20] Siddique M, Abid M, Mufti R. Simulation of welding distortions and
residual stresses in pipe-flange joint using finite element technique:
comparison of 2D and 3D models. Proceedings of IMEC2004
(International Mechanical Engineering Conference and Expo),
Kuwait, December 58 2004.
[21] Lindgren LE, Hedblom R. Modelling of addition of filler material in
large deformation analysis of multipass welding. Commun Numer
Methods Eng 2001;17:64757.
[22] Lindgren LE. Finite Element Modeling and Simulation of
Welding Part 1: Increased complexity. J Therm Stress 2001;24:
14192.
[23] Andersen, L., Residual Stresses and Deformations in Steel Structures,
PhD. thesis, Technical University of Denmark, 2000.
[24] Goldak J, Chakravarti A, Bibby M. A new finite element
model for welding heat sources. Metall Trans B 1984;15B:
299305.
[25] Goldak J, Bibby M, Moore J, House R, Patel B. Computer
modeling of heat flow in welds. Metall Trans B 1985;17B:
587600.
[26] Lin YC, Chou CP. A new technique for reduction of residual stress
induced by welding in type 304 stainless steel. J Mater Process
Technol 1995;48:6938.
M. Abid, M. Siddique / International Journal of Pressure Vessels and Piping 82 (2005) 860871 871