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Analysis of the temperature evolution during lined pipe welding
Obeid Obeid1, a, Giulio Alfano1,b, Hamid Bahai1,c
1School of Engineering and Design, Brunel University, UB8 3PH Uxbridge, UK
aobeid.obeid@brunel.ac.uk, bgiulio.alfano@brunel.ac.uk, hamid.bahai@brunel.ac.uk
Abstract
A numerical analysis of thermal phenomena occurring during lined-pipe welding is presented in this
paper. Numerical models of surfaces and volumetric heat sources were used to predict the time
evolution of the temperature field both in a corrosion-resistance-alloy (CRA) liner, made of SUS304
stainless steel (SS), and for the single-pass girth welding of a carbon-manganese (C-Mn) steel pipe.
Using the finite-element code ABAQUS, three-dimensional non-liner heat-transfer analyses was
carried out to simulate the gas-tungsten-arc (GTA) welding process used in liner welding and the
metal-inert-gas (MIG) welding process consumed in C-Mn steel backing welding. FORTRAN user
subroutines were coded to implement the movable welding heat source and heat transfer coefficient
models. The thermal history was numerically computed at locations where circumferential angles
from the welding start/atop position are 90°, 180° and 270° with respect to axial distances from the
weld centerline (WC).
Keywords: Finite element analysis FEA, CRA Liner, C-Mn steel backing, Heat source, Thermal
history.
1. Introduction
Pipelines used in the Oil and Gas industry often face cyclic loading generated by fluid pressure
changes and, in the case of offshore pipelines, also as a result of waves and currents, the latter possibly
leading to vortex-induced vibrations. Harsh operating conditions and corrosive production fluids
make the use of C-Mn steel pipes for flow line impossible, whereby alternatives are required. One
option is provided by lined pipes, in which a CRA liner is inserted mechanically into a C-Mn steel
backing pipe through a thermo-hydraulic manufacturing process to provide an appropriate fit between
the liner and C-Mn steel. Economically, there is a massive demand for lined pipes because of the
associated reduced maintenance cost and increased life cycle as per recent marketing studies [1].
To fix the liner with carbon steel backing, a weld overlay is applied to the internal surface of the pipe
and a girth weld can then be produced between joints of the lined pipe. After that, a weld bevel and
subsequent welds can be located in the weld overlaid region. Thermal stresses occur in the weld and
adjacent areas producing significant residual stress fields due to the no-uniform expansion and
contraction of the metal.
Due to the high temperature gradients experienced in the vicinity of both the weld overlay and the
girth weld, incompatible plastic strain are produced leading to increase in the susceptibility of a weld
to fatigue damage. Furthermore, fracture and stress corrosion cracking increases due to detrimental
residual stresses [2].
Lined pipes are a relatively new technology for which knowledge and experimental testing are still
limited, particularly in terms of the temperature history in the welding regions which is still not well
understood. As a result, the aims of this research are to numerically simulate, validate and better
understand the thermal response of lined pipelines during welding.
To this aim, finite element modelling FEM is a valid tool to predict the thermal activity during liner
welding and backing steel welding because of the complex nature of the phenomena involved.
Models for butt-welding have developed by Goldak and co-workers [3], who proposed a double
ellipsoidal geometry so that the size and shape of the heat source can be easily changed to model both
the shallow penetration arc welding processes and the deeper penetration laser and electron beam
processes. Brickstad and Josefson [4] used the so-called ”element birth” technique to represent the
sequence of weld beads. Karlsson and Josefson [5] used the code ADINA to study temperatures,
stress and deformation in a single-pass butt welding with full three-dimensional FEM. Indeed, to our
knowledge, up to now, neither two-dimensional nor three-dimensional thermal and mechanical study
of the lined pipe welding process has been reported in the literature because of the complexity of
lined-pipe welding modelling.
In this work, accurate prediction and analysis of the thermal cycle in the welding joint is the first step
required in the welding simulation to predict the residual stresses. To this end, a non-linear 3D
analysis is required to predict the temperature distribution, by consideringa moving heat source
(torch) to generate the temperature field during the welding process. The knowledge of the
temperature history is necessary to determine the thermal stresses, displacement and strain fields
during and after the welding process (cooling) including, in particular, the plastic deformation as a
result of thermal expansion.
2. Finite Element Modelling
This study aims to calculate three-dimensional thermal field during single-pass overlay welding for
CRA liner followed by the single-pass girth welding of a backing steel pipe. The pipe has an outer
diameter of 114.3 mm and a wall thickness of 6 mm, of which 4.5 mm is the thickness of the C-Mn
outer pipe and 1.5 mm is the liner thickness. Only one-half of the pipe, which is 200 mm long, is
analyzed due to symmetry about the weld line. The outer pipe material is C-Mn steel with a
composition of 0.18%C, 1.3%Mn, 0.3% Si, 0.3%Cr, 0.4%Cu (Swedish standard steel SIS2172) and
the temperature-dependent thermal material properties such as density, specific heat, latent
temperature and conductivity used for the outer pipe are taken from the work of Karlsson and
Josefson [5]. The thermal properties for the SUS304 SS liner are obtained from the study of Deng and
Murakawa [2].
The 3-D finite element model is shown in Figure 1 with a total of 51840 nodes associated with 10560
fully-integrated 20 noded heat-transfer brick (i.e. solid, continuum) elements, named DC3D20 in
ABAQUS, which is the code used for the analyses. Among these, 2400 elements belong to the liner
whereas the remaining of 10560 elements belongs to the outer pipe. A fine mesh is used in the fusion
zone (FZ) and the heat affected zone (HAZ) where structural transformation is occurring.
Figure 1: 3-D finite element model with welding direction.
2.1 Thermal Analysis
To simulate the moving heat source it is assumed that the electrode heat power is distributed on a
hemispherical volume with a Gaussain distribution. The Goldak [3] ellipsoidal heat source model
used here is given by :
Herein, is the weld efficiency multiplyed by energy input rate which is 670 and 2900 W for liner
and backing steel welding respectively; a is the half bead length (along the x direction), b is the bead
height (along the y direction) and c is the half bead width (along the z direction) as illustrated in
Figure 2; τ is the current time corresponding to the new position of weld bead whereas t0 is the initial
time to start moving torch. V is the welding speed which is 6.25 and 1.33 (mm/s) for carbon steel and
SS welding speed, respectively.
Figure 2: Ellipsoidal weld bead with semi-axes a, b and c.
Eq. (1), representing the the body flux, was implemented in the FORTRAN user-subroutine DFLUX.
During the thermal cycle, radiation heat losses are dominant in and nearby the welding pool while the
convection ones are dominant in the areas far away from the welding pool, which can be modelled
using non-uniform film coefficients. In this study, two different heat-transfer coefficients have been
implemented in the FORTRAN user-subroutine FILM. The heat-transfer coefficient for the exposed
surfaces of the C-Mn backing steel, hcarbon is taken equal to [6]:
(2)
where hcon is the convective heat transfer (hcon = 7 W/m2C°) and is the effective radiation
emissivity ( ). T is the current temperature at the pipe whereas Tamb is the ambient
temperature. For the SS surfaces the temperature-dependant heat-transfer coefficient is given by [2]:
(3)
2.2 Solution Strategy:
The element-birth technique is used to add sections of the weld bead incrementally over the moving
heat source to represent the transient nature of weld deposition. This approach consists of the
following three stages: (1) all beads elements sets are de-activated; (2) sets of the weld bead are
re-activated in successive and corresponding steps to simulate the deposition of weld beads when the
torch travels around the pipe circumferentially; (3) weld deposition is complete and the pipe is
allowed to cool down. The procedure is applied for the liner welding first and then for the
backing-pipe welding. The MODEL CHANGE option in ABAQUS is used to remove and then add
the weld bead element sets.
3. Results and Discussion:
Figure 3 shows the transient temperature distribution at 270° from the welding start/stop position on
the CRA during liner welding and C-Mn welding respectively. It can be seen that the backing steel
weld material is removed during the liner welding. Different points along axial direction are selected
as shown in Figure 4 where points 1, 2 and 3 are located on the CRA liner and 4, 5 and 6 are located on
backing steel. Figures 5 and 6 show the temperature history during welding at three different
circumferential angles 90°, 180° and 270° from the start/stop position for the previous axial points
from 1 to 6.
(a)
(b)
Figure 3: Temperature distribution at 270° in (a) liner ; (b) C-Mn backing.
Figure 4: Illustration of axial locations (mm).
Figure 5: temperature histories in the liner at (a) 90°, (b) 180° and (c) 270°.
Figure 6: temperature histories in the C-Mn backing steel pipe at (a) 90°, (b) 180° and (c) 270°.
4. Conclusions
It is clear from Figures 5-6 the maximum temperature located in the middle of the welding pool where
the temperature starts out at the ambient temperature of the environment, 20 °C and then rises very
rapidly once the heat source reaches point 1 in SS welding and point 4 in C-Mn welding.
As expected, all the nodes in liner and girth welding region exceed the melting point of their materials
SUS304 and SIS2172 which are 1365°C and 1500°C, respectively, in order to ensure that the melting
is complete and sufficient fluidity is achieved to allow material to flow into the weld groove properly.
The boundary of the HAZ extends from the FZ boundary to reach temperatures between 800-900°C
(A1-A3 in the Fe-C phase diagram). As a result, the length of the HAZ in the SS is equal to the distance
from points 2 to 3, around 2 mm, while the length of the HAZ in the C-Mn is 2.55 mm, from point 5 to
6.
After the heat source has passed a point, the cooling process starts in this point. It is noticeable there is
a rapid drop for points reaching higher heating temperature because radiation is the dominant heat
loss mechanism at higher temperatures while convection comes into action at somewhat lower
temperature.
Welding parameters such as welding current, voltage, efficiency, initial temperature of the pipe and
welding speed are constant values. Consequently, the points on the same circumferential path should
have the same temperature history and peak temperature, as proved in our FE model. Furthermore, the
temperature history is not sensitive to the variations in the circumferential angle from the welding
start/stop position as expected. Experimental validation by comparison of the FE results with
experimental data from the literature is in progress and will presented in a forthcoming article, but
preliminary results are very promising.
References
[1] De Koning, A. C., and H. Nakasugi. "Li Ping,“TFP and TFT Back in Town (Tight Fit CRA Lined
Pipe and Tubing)”." Stainless Steel World (2004): 53-61.
[2] Deng, Dean, and Hidekazu Murakawa. "Numerical simulation of temperature field and residual
stress in multi-pass welds in stainless steel pipe and comparison with experimental measurements."
Computational materials science 37.3 (2006): 269-277.
[3] Goldak, John, Aditya Chakravarti, and Malcolm Bibby. "A new finite element model for welding
heat sources." Metallurgical transactions B 15.2 (1984): 299-305.
[4] Brickstad, Bjorn, and B. L. Josefson. "A parametric study of residual stresses in multi-pass
butt-welded stainless steel pipes." International Journal of Pressure Vessels and Piping 75.1 (1998):
11-25.
[5] Karlsson, R. I., and B. L. Josefson. "Three-dimensional finite element analysis of temperatures
and stresses in a single-pass butt-welded pipe." Journal of pressure vessel technology 112.1 (1990):
76-84.
[6] Abid, M., M. Siddique, and R. A. Mufti. "Prediction of welding distortions and residual stresses in
a pipe–flange joint using the finite element technique." Modelling and Simulation in Materials
Science and Engineering 13.3 (2005): 455.
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