Numerical simulation to study the effect of tack welds and root gap on welding deformations and...

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

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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: [email protected] (M. Abid), mabid00@hotmail.

com (M. Abid), [email protected] (M. Siddique).
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2/12

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

M. Abid, M. Siddique / International Journal of Pressure Vessels and Piping 82 (2005) 860871 861


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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.

M. Abid, M. Siddique / International Journal of Pressure Vessels and Piping 82 (2005) 860871862


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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.



10/12

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.



11/12

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.

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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.



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