Numerical simulation of the effect of constraints on welding deformations and residual stresses in a...

INSTITUTE OF PHYSICS PUBLISHING MODELLING AND SIMULATION IN MATERIALS SCIENCE AND ENGINEERING

Modelling Simul. Mater. Sci. Eng. 13 (2005) 919–933 doi:10.1088/0965-0393/13/6/010

Numerical simulation of the effect of constraints onwelding deformations and residual stresses in apipe–flange joint

Muhammad Abid and Muhammad Siddique1

Faculty of Mechanical Engineering, GIK Institute of Engineering Sciences and Technology, Topi,NWFP, Pakistan

E-mail: [email protected], [email protected] (M Siddique) and siddique [email protected]

Received 11 May 2005, in final form 1 July 2005Published 29 July 2005Online at stacks.iop.org/MSMSE/13/919

AbstractThis paper presents a detailed three-dimensional finite element (FE) study toinvestigate the effect of mechanical constraints on welding distortions andresidual stresses in a pipe–flange joint. The FE model of a pipe–flange joint issubjected to sequentially couple nonlinear transient thermo-mechanical analysisto simulate complex welding phenomena. Single-pass gas metal arc welding forsingle ‘V’ butt-weld joint geometry of a 100 mm diameter pipe with compatibleweld-neck ANSI flange class # 300 of low carbon steel is simulated. Two tack-welds at 90˚ and 270˚ from the weld start position are modelled. Four differentconstraint conditions representing the welding of unassembled joints, weldingof assembled joints, welding of assembled joints with reflective symmetry andwelding of perfectly constrained joints are analysed. To model the constraintsand boundary conditions more realistically contact pairs are used between thematching surfaces of different structural components. Basic FE models arevalidated with experimental data for temperature distribution and deformations.Predicted welding distortions and residual stresses are compared and discussedin detail. From the results, axial displacement and tilt of the flange face arefound to be strongly dependant on the constraint conditions. Minimum axialdistortion on the flange face is found for rigidly clamped flanges. However,residual stresses have a weak dependence on the constraints set.

1. Introduction

Welded pipe–flange joints are widely used in a variety of engineering applications such as oiland gas pipelines, nuclear and thermal power plants and chemical plants etc. A non-uniform

1 Author to whom any correspondence should be addressed.

0965-0393/05/060919+15$30.00 © 2005 IOP Publishing Ltd Printed in the UK 919

http://dx.doi.org/10.1088/0965-0393/13/6/010

mailto: [email protected]

mailto:[email protected]

mailto: [email protected]

http://stacks.iop.org/ms/13/919

920 M Abid and M Siddique

temperature field, applied during welding, produces deformation and residual stresses in thewelded structures. For flange joints, any tilt or out-of-plane deformation in the flange faceresults in gasket damage [1–3]. In addition, uneven bolt-up load consequently producesadverse effects on the service life of the joint [4]. Residual stresses in the piping systemmay have a larger contribution to the total stress field as compared to the applied loadingswhile assessing the risks for defect growth and static fracture in piping systems with brittlefracture behaviour [5]. Therefore, realistic prediction of contributing factors is of vitalimportance. The extent of deformations and residual stresses in welded components dependson several factors such as geometrical size, welding parameters, weld pass sequence andapplied thermal and structural boundary conditions.

Finite element (FE) simulation has become a proven and reliable technique for predictionof welding distortions and residual stresses. The accuracy of predicted results depends on theanalysis technique, FE model and assumptions considered. A substantial amount of simulationand experimental work focusing on circumferential welding with emphasis on pipe weldingis available in the [5–9]. To reduce computational power requirements, assumptions suchas rotational symmetry [5–7] and lateral symmetry [6–9] have been employed in numericalsimulations. These assumptions, though, reduce the computational demand but make theproblem over simplified by limiting the analysis to a portion of the complete geometry.Therefore, these models are not capable of predicting the effects of weld start/stop location,root gap, tack-welds and structural constraints. In the available FE studies of pipe–flangejoints Troive et al [10–12] presented experimental and FE studies for deformation andresidual stresses in multi-pass welded pipe–flange joints. In all these studies two-dimensionalmodels (with assumption of rotational symmetry) have been used to model double-J multi-passmanual metal arc welding of plate flanges in horizontal position (without structural restraints)for investigation of radial shrinkage, flange twist and resulting residual stresses. Later onSiddique et al [13] and Abid et al [14], though, used three-dimensional models for simulationof an automatic gas metal arc welding (GMAW) process for single-pass butt-girth welding ofweld-neck ANSI flanges, but again no welding fixture (for structural constraint) is modelled.

During welding of the flange with a pipe, undesirable flange twist and residual stressesare produced which subsequently need remedial action [2]. In critical applications, post-weldmachining of the flange face is sometimes essential [2,11]. Since flange twist severely affectsthe sealing ability of the joint, it is important to control flange deformation during welding sothat post-weld machining is minimized, even better avoided. The objective of the present studyis to investigate the role of structural boundary conditions in minimizing flange deformationand to study their impact on the extent of residual stress distribution.

2. Present study

The aim of the current paper is to define a methodology for the welding of a pipe–flange jointto minimize flange deformation. Four cases with different structural boundary conditions areanalysed and compared for their relative effects, in the following cases:

Case-I: Welding of joint without assembling and without any welding fixture.Case-II: Welding of joint in assembled position.Case-III: Welding of joint in assembled position and with reflective symmetry.Case-IV: Welding of joint without assembling and in rigid welding fixture.

A three-dimensional FE simulation of single-pass butt-weld joint geometry is performedusing ANSYS. A low carbon steel pipe of 100 mm nominal diameter, 200 mm length and6 mm wall thickness is modelled with a same-sized weld-neck type ANSI class # 300 flange.

Effect of constraints in welded pipe–flange joints 921

Figure 1. Experimental setup (a) welding of joint in free condition. (b) Welding of joint inassembled condition.

(This figure is in colour only in the electronic version)

In FE analysis, pipe, flange and filler metal are supposed to be of the the same chemicalcomposition, i.e. 0.18% C, 1.3% Mn, 0.3% Si, 0.3% Cr, 0.4% Cu (Swedish standard steelSIS 2172). Temperature dependent material properties are taken from [15] and changessuggested in [14] are also incorporated in the material model. FE models for Case-I andCase-II are validated with experimentally measured weld-pool size and resulting deformations.Manufacturing stresses of components and initial effect of tack-welds on distortion and residualstresses are ignored.

3. Experimental setup

For automatic circumferential welding of a pipe–flange joint, a dc powered conventional lathewith an open loop continuous-speed controller is synchronized with a Lincoln Power SourcePowerwave-355. Synchronization is achieved through an interface controller operated by limitswitches, mounted on a lathe chuck, indicating weld start and end positions. A tacked sampleof a pipe–flange joint is rotated in the chuck of the lathe, while the welding torch is heldstationary by mounting it in a special fixture. This torch mounting fixture, in combination withmachine carriage, provides 4-Axis (3 translational and 1 rotational) adjustments to the torch.Automatic circumferential welding facility, used in the present study, is shown in figure 1.

The GMAW welding process with gross heat input of 792 KJ m−1 is used. In the absence ofa weaving facility a forehand-welding technique (having torch angle 17.5˚ with the normal) isused to control penetration and avoid blow off. The dimensions of the physical specimenare in accordance with the size and geometry described in section 2. The material forpipe and flange is carbon–manganese steel with chemical composition 0.2% C, 0.85% Mn,0.22% Si, 0.15% Cr, <0.05% S and <0.05% P. Filler metal ER70S-6 Carbon Steel wire ofdiameter 1.14 mm (0.045′′) is used. Though the chemical composition of the base metalused here is slightly different from the one used in the simulation, such a difference isbelieved to have no significant effect on mechanical properties. Thus it permits adoptionof material properties for numerical welding simulation especially when structural responseis validated experimentally. A similar case for material properties adoption can also be foundin, e.g. [16].


Single-pass butt-weld joint geometry with a 6 mm-deep single V-groove (60˚ includedangle) and 1.2 mm root gap is used. Two initial tack-welds at angular positions of 90˚ and270˚ from the weld start position were deposited. Each tack-weld was machined to a lengthof 10 mm and thickness of 3.0 mm. Subsequently samples were stress relieved to removemanufacturing and tack-welding stresses from pipe and flange, as existing residual stressesaffect thermal expansion behaviour [16]. Mechanical dial indicators of 1 µm resolution and2 µm accuracy are used for in situ measurement of axial displacement on flange face and radialshrinkage of pipe due to welding. Measurements on the flange face are taken at a radius of117 mm. For radial shrinkage the surface of the pipe at four angular positions (at 0˚, 90˚, 180˚and 270˚ from weld start position) is scanned before and after welding, and radial shrinkageis found from the difference of measurements.

4. Analysis procedure

Taking advantage of the weak structure to thermal filed couplings, the problem is formulated asa sequentially coupled thermal stress analysis. At first, a nonlinear transient thermal analysisis performed to predict temperature history of the whole domain. Subsequently, results ofthermal analysis are applied as thermal body load in a nonlinear transient structural analysis tocalculate deformation and stresses. The FE model for both thermal and structural analysis iskept the same, except element type. During analysis a ‘Full Newton–Raphson’ iterative solutiontechnique with a direct sparse matrix solver is employed for obtaining the solution. Duringthe thermal cycle, temperature, and consequently temperature-dependent material properties,change very rapidly; thus, a Full Newton–Raphson scheme which uses a modified materialproperties table and reformulated stiffness matrix at every equilibrium iteration is believedto give more accurate results than other options such as Modified or Initial Newton–Raphsonschemes. Line search option of the FE code ANSYS [17] is set to ON to improve convergence.A single point reduced integration scheme with hour-glass control is implemented to facilitateconvergence and to avoid excessive locking during structural analysis.

The conventional quiet elements technique, as described in [18], is chosen for themodelling of filler material and is implemented by using an element birth and death feature.A complete FE model is generated at the start; however, all elements representing filler metalexcept elements for the tack-welds are deactivated by assigning them a low stiffness andlow thermal conductivity. During thermal analysis all the nodes of deactivated elements(excluding those shared with the base metal) are also fixed at room temperature till the birth ofthe respective elements. Deactivated elements are re-activated sequentially when they comeunder the influence of the welding torch. For the subsequent structural analysis, birth of anelement takes place at the solidification temperature. Melting and ambient temperatures are setas reference temperatures (temperature at which thermal strain is zero) for thermal expansioncoefficients of filler and base metals, respectively. To avoid excessive distortion initial strainin the elements is set to zero at the time of element reactivation.

For thermal analysis, total welding time of the complete circumferential weld, i.e. 58 s isdivided into 144 equally spaced solution steps, further divided into two sub-steps. A steppedload option is used for realistic application of thermal load. After extinguishing the arc another56 load steps of different time duration are meant for cooling of the weldment. It took about52 min to return to the ambient temperature of 27˚C. Load step time in structural analysis iskept equal to thermal analysis; however each load step is solved in a single sub-step exceptin the cases of numerical non-convergence. The restart option of software with correctedsub-step setting is effectively used to handle non-convergences. Total CPU time remainedapproximately 3.5 h and 56.4 h for thermal and structural analysis (of Case-II), respectively,on an IBM compatible P-IV 3.2 GHz PC with 2 GB RAM.


Figure 2. FE model used. (a) FE model-I (for Case-I and IV), (b) Mesh refinement, V-groove,tack-weld and root gap, (c) FE model-II (for Case-II), (d) FE model-III (for Case-III).

5. Finite element models

Three FE models for different studies, developed in ANSYS, are shown in figure 2. AnFE model-I in figure 2(a) is used for analysis of Case-I and Case-IV, described above. Thismodel consists of total 26 630 nodes and associated 22 304 linear elements, out of which 13 952elements are used for flange and the other 8352 elements represent pipe. Due to anticipated hightemperature and stress gradients near the weld, a relatively fine mesh is used up to a distanceof 10 mm on both sides of the weld centreline. The element size increases progressively withdistance from the weld centreline. At the weld centreline the whole circumference is dividedinto 144 elements giving an average element length of 2.4 mm. In axial direction elementsize varies from 0.2–1.5 mm and in radial direction pipe thickness is divided into four, 1.5 mmlong, elements. To model each individual tack, four elements in hoop, two in radial and twoin axial direction are used. A mesh refinement scheme with approximate V-groove formationand tack-weld is shown in figure 2(b).

Mostly eight-node brick elements with linear shape functions are used in the model. Lin-ear elements are preferred because one, in general, favours more lower-order elements thanfewer higher-order elements in nonlinear problems [19]. Though not preferred a few six-nodetetrahedrons are also used to match coarse and dense meshes in the model. In order to facil-itate data mapping between thermal and structural analysis, the same FE model is used with


respective element types. For thermal analysis the element type is SOLID70 which has eightnodes with a single degree of freedom, and temperature, at each node. For structural analysisthe element type is SOLID45 with three translational degrees of freedom at each node.

FE model-II (figure 2(c)) represents Case-II where the pipe–flange joint, to be welded, isbolted with another flange and the whole model is used to simulate welding of the assembledjoint. To develop this model, FE model-I is extended by adding the bolts and a second flange.A meshing scheme near the weld zone is kept the same but due to addition of other partsthe total number of solid elements increased to 34 060. Further, to model the bolting effectmore realistically, contact element pairs are defined between the matching raised faces of theflanges and also between the flange webs and nut/bolt-heads. A three-dimensional contactelement CONTA173, 4-node surface-surface contact elements, with matching target elementsTARGE170 are used for this purpose. Iterative numerical testing was carried out for optimalselection of key options for the contact pair to improve convergence.

FE model-III (figure 2(d)) represents an assembled joint with translational symmetryi.e. two identical pipe–flange joints are assembled and welded simultaneously. A section ofsymmetry is taken at a distance of 1.0E-8 mm from the raised face and bolts are also truncated atthis section. The bolts are modelled as a part of the flange by joining bolt head with the flangebecause excessive non-convergence was observed during the trial run with contact elementpairs between the bolt head and the flange web. However, to model the opposing surface ofthe symmetric flange face, three-dimensional node–node contact elements CONTAC52 aregenerated between the nodes of flange face and offset nodes (by a distance of 1.0E-8 mm fromflange face).

6. Boundary conditions

In thermal analysis, both radiation and surface convection are considered for modelling ofheat loss from the surface. During the thermal cycle, radiation dominates over the surfaceconvection in weld pool and surrounding areas; whereas, away from the weld pool convectionis the primary mechanism of heat loss from the body. Instead of modelling convection andradiation separately, a combined heat transfer coefficient, as used in [14], is calculated andapplied on all areas exposed to the ambient air. The ambient temperature (27˚C) is taken asthe initial condition for the entire mass involved.

In structural analysis, the boundary conditions vary from case to case. In the FE model-I theonly constraint applied is representation of clamping in machine chuck, as shown in figure 1(a).For this purpose all the nodes at the far end of the pipe, on a Cartesian coordinate axis, areconstrained in the axial and radial directions. For Case-II, welding of joints in assembledposition (extreme ends of the model are constrained in all directions) as schematically shownin figure 3(a). During the course of experimental investigation we tried to achieve similarboundary conditions. The arranged shown in figure 3(b) is used for this purpose. Figures 3(c)and (d) show constraint conditions for Case-III and Case-IV, respectively.

7. Heat source modelling

Proper modelling of heat flow from the welding torch to the weldment is quite crucial as itcontrols the application of thermal load which consequently produces distortion and residualstresses in the weldment. For the determination of weld pool size and shape, a section of weldis cut, polished, chemically etched and scanned. The so-called double ellipsoidal heat sourcemodel by Goldak et al [20] is slightly modified to achieve the shape of measured isotherms


Weld Bead Constraint

Contact Surface

Contact Surface

Constraint

Flange APipe Flange B

(a) Contact Surface

Contact Surface

Symmetry B.C

(c)

ConstraintConstraint

(d)

Solid BarZ

Side Disc

(b)

Weld Bead

Figure 3. Schematic of structural boundary conditions in different cases. (a) For Case-II (FE),(b) for Case-II (experimental), (c) for Case-III, (d) for Case-IV.

for fusion boundary and outer HAZ. The governing equations for power density distributionin front and rear ellipsoids of the three-dimensional model are as under:

qf = 6√

3M(r,z)Qff

π√

πafbcexp

[−3

((rθ)2

a2f

+z2

b2+

(Ro − r)2

c2

)], (1)

qr = 6√

3M(r,z)Qfr

π√

πarbcexp

[−3

((rθ)2

a2r

+z2

b2+

(Ro − r)2

c2

)], (2)

where

Q = V I,

ff + fr = 2,

η =∑n

i=1 qiVi

Q.

The description and numerical values for different variables in power density distributionequations are given in table 1. Since the same model and welding parameters were used,further details about the heat source and resulting weld pool shape can be found in Abid andSiddique [21].

8. Model validation

In order to validate FE models and the technique used, the results obtained from FE analysiswere experimentally validated for Case-I and Case-II. For Case-I, only axial displacement offlange face is recorded while for Case-II axial displacement of flange face as well as radialcompression of the pipe (at four positions—0˚, 90˚, 180˚ and 270˚ from the weld start position)are measured. For Case-I, when welding is carried out without any constraints on flange,calculated axial deformation showed good agreement with the experimental data, figure 4(a).

For Case-II in figure 4(b), when welding is performed under a constraint set as shownin figure 3(a), calculated axial deformation in bolted condition showed a relatively betteragreement with the experiment. However FE simulation overestimated axial deformation


Table 1. Heat source parameters.

Symbol Description Value

af Front Ellipsoidal semi-axes length (m) 0.012 9ar Rare Ellipsoidal semi-axes length (m) 0.010 3b Half width of arc (m) 0.005c Depth of arc (m) 0.006ff Fraction of heat deposited in front of arc centre 1.55fr Fraction of heat deposited in rare of arc centre 0.45I Welding current (Amp) 225M(y,z) Scalar multiplier —n Total number of element under torch influence —qi Power density for ith element (W m−3) —qf Power density in front ellipsoid (W m−3) —qr Power density in rare ellipsoid (W m−3) —Q Gross heat input per second (W) 4950r Radius of pipe (m) —Ro Pipe outer radius (m) 0.057 5v Welding speed (m s−1) 0.006 25V Voltage (V) 22Vi Volume of ith element (m3) —z Distances from the torch centre in axial direction (m) —θ Angle from instantaneous arc position (Radian) —η Arc efficiency 85%

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 Coordinates (Deg)

Axi

al D

ispl

acem

ent (

mm

)

Exp-Case-I

FE-Case-I

(a)

0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0 60 120 180 240 300 360


Axi

al D

ispl

acem

ent (

mm

)

Exp-Case-II-Bolted

FE-Case-II-Bolted

Exp-Case-II-Unbolted

FE-Case-II-Unbolted

(b)

Figure 4. Comparison of experimental and FE results for flange face deformation at a radius of117 mm. (a) For Case-I, (b) for Case-II.

on unbolting as compared to experimental measurements. The probable reason for thisoverestimation is the pre-tensioning of bolts. Since no initial bolt-up load is applied insimulation thus it ignores the initial flange face rotation and resulting axial and/or radialdeformations. Numerical simulation of bolting behaviour is quite complicated due to thecomplexities involved in experimentally determined bolt behaviour such as the effects ofthe bolt tightening sequence [22], bolt embedding and bolt load relaxation under applied


-0.08

-0.04

0

0.04

0.08

0.12

0.16

5 15 25 35 45 55 6

Distance from Weld Centerline (mm)R

adia

l Dis

plac

emen

t (m

m)

Exp-B-180 Deg

Exp-UB-180 Deg

FE-B-180 Deg

FE-UB-180 Deg

B - BoltedUB - Unbolted

(b)

-0.14

-0.12

-0.1

-0.08

-0.06

-0.04

-0.02

0

0.02

0.04

0.06

5 15 25 35 45 55 65 75

Distance from Weld Centerline (mm)

Rad

ial D

ispl

acem

ent (

mm

)Exp-B-0 Deg

Exp-UB-0 Deg

FE-B-0 Deg

FE-UB-0 Deg


(a)

Figure 5. Comparison of experimental and FE results for radial shrinkage of pipe. (a) At 0˚ fromweld start position. (b) At 180˚ from weld start position.

axial loads [3, 4] etc. On the numerical side, contact element behaviour further aggravatesthe situation. For the sake of simplicity bolt pre-tensioning is ignored, which consequentlyeliminated bolt embedding, the effect of the bolt tightening sequence and bolt relaxation underload. However, contact elements showed ill behaviour during cooling at about 28˚C above theambient temperature where rigid body motion of the bolts was experienced due to abruptlychanged contact conditions of these elements. This ill-conditioned behaviour persisted forseveral variations in the contact modelling; thus the joint was unbolted at this temperature bydeactivating contact elements (also reducing their stiffness to zero). Subsequent cooling wasthen allowed to take place in an unbolted condition. This might be the other reason for slightlylower bolted but higher unbolted deformations.

In figure 5, numerically calculated radial shrinkage, in comparison with the experimentalmeasurements, is found slightly higher at 0˚ while slightly lower at 180˚ from the weld startposition. Though the radial shrinkage quantitatively differs it qualitatively matches well withthe experimental measurements. However, radial shrinkage at 90˚ and 270˚ has different trendsas compared to the measurements, shown in figure 6. This is attributed to the flange face rotationduring initial bolt-up. Figure 7(a) indicates the schematic of flange and pipe distortion underinitial bolt-up. The effect of flange distortion shifts to the pipe through tack-weld (at 90˚ and270˚). This effect vanishes from the pipe as the tacks come under the influence of welding torchand become soft. Since the FE model does not contain the effect of initial bolt-up, thereforeexperimentally measured radial shrinkage differs from the numerically calculated shrinkageand the difference becomes more evident at the tack locations.

9. Results and discussion

9.1. Deformations

Flange face deformation, resulting from transverse shrinkage of the weld bead, is of vitalimportance for functional performance and the sealing ability of the pipe–flange joint. From


-0.1

-0.05

0

0.05

0.1

0.15

5 15 25 35 45 55 65 75


Rad

ial D

ispl

acem

ent (

mm

)Exp-B-90 Deg

Exp-UB-90 Deg

FE-B- 90 DegFE-UB-90 Deg


(a)

-0.09

-0.06

-0.03

0

0.03

0.06

5 15 25 35 45 55 6

Distance from Weld Centerline(mm)R

adia

l Dis

plac

emen

t (m

m)

Exp-B-270 Deg

Exp-UB-270 DegFE-B- 270 Deg

FE-UB-270 Deg

(b)


Figure 6. Comparison of experimental and FE results for radial shrinkage of pipe. (a) At 90˚ fromweld start position. (b) At 270˚ from weld start position.

Bolting Force

Pivot Point Initial Geometry

Distorted Geometry

0

0.1

0.2

0.3

0.4

0.5

0 60 120 180 240 300 360


Axi

al D

ispl

acem

ent (

mm

)

Bolted (15 N-m)

Unbolted (15 N-m)

Bolted (108 N-m)

Unbolted (108 N-m)

(a) (b)

Figure 7. (a) Effect of bolt tightening on initial geometry. (b) Effect of bolt pre-tension on resultingwelding deformation (for Case-II).

the application view-point, axial deformation of the flange face is more critical when comparedto the radial shrinkage of the pipe. Therefore, only flange face deformations are focused on inthe following discussion.

9.1.1. Effect of bolt-up torque. The initial bolt-up load in a pipe–flange joint is believedto have a significant effect on the resulting deformation during welding since it serves asinitial pre-tensioning of bolts and the joint. The bolt load relaxation is a widely experienced


0.15

0.20

0.25

0.30

0.35

0.40

0.45

0.50

0.55

0.60

0.65

0 60 120 180 240 300 360


Axi

al D

ispl

acem

ent (

mm

)

Case-I

Case-II

Case-III

Case-IV

Exp- Case-II

(a)

0.00

0.02

0.04

0.06

0.08

0.10

Case-I Case-II Case-III Case-IV

Fla

nge

Fac

e T

ilt (D

eg)

(b)

Figure 8. (a) Comparison of flange face displacement for different boundary conditions.(b) Resulting flange face tilt.

phenomenon (reported relaxation up to 61% of initial bolt-up load); thus bolting conditionand sequence can contribute significantly to flange face rotation [3]. In an experimental studyby Kumakura et al [22] bolt tension reduction was attributed to bolt embedding and wasalso found dependent on nut thickness. The dependence of bolt tension loss on nut thicknessindicates that it is probably due to localized yielding of threads.

During the present experimental study (Case-II) flange deformation during welding fortwo different bolt tightening torques (with all the other parameters unchanged) are examined.In the first case feel tight (15 N m) torque is applied on all the bolts; whereas, in the secondcase 108 N m torque is applied. In figure 7(b), post-weld axial displacements on the flangeface (at a radius of 117 mm) for the two cases shows that after cooling to ambient temperaturethe flange profile in bolted condition is more or less identical to the final profile of the flangeafter unbolting. It further revealed that higher bolt-up torque produced a drastic reduction inflange face tilt and some reduction in overall axial displacement as well.

Though experimental studies showed a significant effect of initial bolt-up torque due to thecombined effect of bolt embedding, localized yielding in threads and elastic bending of flangeunder bolt up, the bolt embedding and localized yielding in the bolt threads are too difficult tomodel, therefore bolt pre-tensioning is not used in the numerical simulation for simplification.Therefore in the following discussion comparison will be made between simulation results andexperimental data for the 108 N m bolt tightening torque case.

9.1.2. Axial displacement of flange face. In figure 8(a), numerical results for axialdisplacement under different boundary conditions are shown. Final flange face profilesobtained for all the cases, except Case-I, are found to be identical. The results show thataxial displacement, due to lateral shrinkage of weld, reduces for more rigid constraints. Axialdisplacement as well as out-of-plane deformation are found to be maximum for Case-I whenwelding is supposed to be carried out without bolting and are minimum when the flange faceis assumed to be rigidly clamped in a rigid fixture (Case-IV). Figure 8(b) shows that the flange


-350

-250

-150

-50

50

150

250

350

450

0 45 90 135 180 225 270 315 360


Axi

al S

tres

ses

(MP

a)

Un-Bolted Bolted

Un-Bolted Bolted

For Inner Surface

For Outer Surface

(a)

-50

0

50

100

150

200

250

300

350

0 45 90 135 180 225 270 315 360

Hoop Coordinates (Deg)H

oop

Stre

ss (M

Pa)

Un-Bolted Bolted

Un-Bolted Bolted

For Inner Surface

For Outer Surface

(b)

Figure 9. Effect of unbolting on residual stresses at weld centreline (Case-IV). (a) Axial stresses,(b) Hoop stresses.

face tilt also reduces for more rigid constraints. From the results it can be concluded thatthough axial displacement and tilt of flange face cannot be eliminated completely they can bereduced significantly with the increase in axial constraint.

9.2. Residual stresses

9.2.1. Bolted versus unbolted stresses. External constraints during welding usually reducedeformations significantly but give higher stresses. Stress relaxation with simultaneousincrease in deformations is reported when these constraints are removed after welding.Logically the effect should be more pronounced in highly constrained structures. In orderto evaluate the post-weld boundary condition removal effects in pipe–flange joints, stresses forbolted and unbolted conditions in the most rigidly constrained case, i.e. Case-IV are compared.A change in stress profile at the weld centreline (for both the inner and outer surfaces) is shownin figure 9, which shows stress relaxation everywhere except in compressive axial stresses on theouter surface. The results indicate that axial stresses are mainly affected by the unbolting andthe effect is even more pronounced on the inner surface. Since the axial unbolting strains ‘springback’ are usually much higher than radial unbolting strains (for reference see figures 4(b) and 5)more pronounced axial stress relaxation is quite understandable. Axial stress variation in axialdirection on the inner surface, figure 10(a), indicates that stress in the bolted condition isgenerally higher on the pipe side and can be attributed to a lesser cross-sectional area of thepipe when compared to the taper hub of the flange. However, the effect of unbolting is notsignificant for hoop stresses, figure 10(b).

9.2.2. Stress variations in the hoop direction. The effect of different boundary conditionson final stress distribution (unbolted) in the hoop direction at the weld centreline is shownin figure 11(a) and (b). The results indicate that the higher the constraint the lower is theresidual stress field (except hoop stresses on outer surface). It further reveals that the effect ofconstraint is more significant on the inner surface as compared to the outer surface and even


-300

-200

-100

0

100

200

300

400

-65 -55 -45 -35 -25 -15 -5 5 15 25 35 45


Axi

al S

tres

s (M

Pa)

Un-Bolted

Bolted

Pipe Side Flange Side

(a)

-200

-100

0

100

200

300

400

-50 -40 -30 -20 -10 0 10 20 30 40 50

Distance from Weld Centerline (mm)H

oop

Stre

ss (M

Pa)

Un-Bolted

Bolted

(b)


Figure 10. Effect of unbolting on residual stress variation in axial direction on inner surface at180˚ from weld start position (Case-IV). (a) Axial stresses, (b) Hoop stresses.

-300

-200

-100

0

100

200

300

400

0 45 90 135 180 225 270 315 360


Axi

al S

tres

ses

(MP

a)

Case-ICase-IICase-IIICase-IV

For Inner Surface

For Outer Surface

(a)

-50

0

50

100

150

200

250

300

350

0 45 90 135 180 225 270 315 360Hoop Coordinates (Deg)

Hoo

p St

ress

(MP

a)

Case-ICase-IICase-IIICase-IV

For Inner Surface

For Outer Surface

(b)

Figure 11. Residual stress variation in hoop direction at weld centreline on outer and inner surfaces.(a) Axial stresses, (b) Hoop stresses.

more pronounced for axial stresses. The most affected place is the stress concentration nearthe weld end position.

9.2.3. Stress variations in the axial direction. Since stresses on only the inner surface aremainly influenced by the boundary conditions, comparison of stress variation in the axialdirection is kept limited to the inner surface. Figure 12 shows that both axial and hoop residualstresses exhibit very low sensitivity for different axial constraints. In general, the stress profileremained unaffected. A very short zone on the pipe side (near the weld centreline) is influenced


-300

-200

-100

0

100

200

300

400

-50 -40 -30 -20 -10 0 10 20 30 40 50Distance from Weld Centerline (mm)

Axi

al S

tres

s (M

Pa)

Case-I

Case-II

Case-III

Case-IV


(a)

-200

-100

0

100

200

300

400

-50 -40 -30 -20 -10 0 10 20 30 40 50

Distance from Weld Centerline (mm)H

oop

Stre

ss (M

Pa)

Case-I

Case-II

Case-III

Case-IV

(b)


Figure 12. Residual stress variation in axial direction on inner surface at a section 180˚ from weldstart position. (a) Axial stresses, (b) Hoop stresses.

by the boundary conditions, where the highly constrained model showed relatively lowerstresses. It is however worth mentioning that excessively high constraint (ideally rigid, suchas in Case-IV) has not shown any significant additional benefit in terms of residual stressreduction.

10. Conclusion

From the comparison of results for different boundary conditions, the following are concluded:

(a) Both the flange face tilt and extent of lateral shrinkage of weld can be reduced significantlywith the application of bolting constraints during welding of pipe flange joints. Therefore,for better service life and improved leak tight operations, the joints should be preferablywelded in assembled condition.

(b) Pre-tensioning of bolts has a positive effect on the out-of-plane deformation of theflange face.

(c) Axial constraint on the flange has no significant effect on the residual stress profile. Eventhe minor effects observed favour the use of axial constraint (bolting).

(d) Moderate axial constraint, such as welding in assembled condition with pre-tensioningof bolt can give an effective control over the final flange geometry and can eliminate theneed of post-weld machining. Designing of excessively rigid welding fixtures may notcontribute a lot in addition.

(e) For improved approximate solutions, more work is needed for more appropriate numericalmodelling of bolts.

References

[1] Abid M and Nash D H 2002 Risk assessment studies for gasketed and non-gasketed bolted pipe joints IPC2002:International Pipeline Conf. (Proc. IPC2002/IPC-27386) (Calgary, Canada, 29 September–3 October 2002)pp 1–11


[2] Nash D H and Abid M 2004 Surface sensitivity study of non-gasketed flange joint J. Proc. Mech. Eng. 218205–12

[3] Abid M 2000 Experimental and analytical studies of conventional (gasketed) and unconventional (non-gasketed)flange pipe joints (with special emphasis on the engineering of joint strength and sealing) PhD ThesisUniversity of Strathclyde, Glasgow

[4] Abid M and Nash D H 2004 Comparative study of the behavior of conventional gasketed and compact non-gasketed flangeed pipe joints under boltup and operating conditions Int. J. Pressure Vessels Piping 80 831–41

[5] Brickstad B and Josefson B L 1998 A parametric study of residual stresses in multi-pass butt-welded stainlesssteel pipes Int. J. Pressure Vessels Piping 75 11–25

[6] Rybicki E F, Schmueser D W, Stonesifer R W, Groom J J and Mishaler H W 1978 A finite element model forresidual stresses and deflections in girth-butt welded pipes ASME J. Pressure Vessel Technol. 100 256–62

[7] Rybicki E F and Stonesifer R B 1979 Computation of residual stresses due to multipass welds in piping systemASME J. Pressure Vessel Technol. 101 149–54

[8] Karlsson L, Jonsson M, Lindgren L E, Nasstrom M and Troive L 1989 Residual stresses and deformations in awelded thin-walled pipe ASME Pressure Vessels Piping Conf. ed E Rybicki et al Weld Residual Stresses andPlastic Deformation (Honolulu, Hawai, 1989), vol 173, pp 7–14

[9] Dong Y, Hong J, Tasi C and Dong P 1997 Finite element modeling of residual stresses in austenitic stainlesssteel pipe girth welds AWS Weld J. (Welding Res. Suppl.) 442-s–9-s 76 (10)

[10] Troive L and Jonsson M 1994 Numerical and experimental study of residual deformations to double-j multi-passbutt-welding of a pipe-flange joint Proc. IEMS ’94: 1994 Annual International Conf. on Industry, Engineeringand Management Systems (Cocoa Beach, Florida USA, 1994) pp 107–14

[11] Troive L, Nasstrom M and Jonsson M 1998 Experimental and numerical study of multi-pass welding processof pipe-flange joints ASME J. Vessels Technol. 120 244–51

[12] Troive L, Nasstrom M and Lin R 1998 Numerical simulation of multi-pass butt-welding of pipe-flange jointswith neutron diffraction determination of residual stresses Proc. NUMIFORM 98 (Enschede, Nederlanderna,22–25 June 1998) ed Huetink and Baaijens Simulation of Material Processing: Theory, Methods andApplications

[13] Siddique M, Abid M, Junejo H F and Mufti R A 2005 3-D finite element simulation of welding residual stressesin pipe-flange joints: effect of welding parameters J. Mater. Sci. Forum 490–491 79–84

[14] Abid M, Siddique M and Mufti R A 2005 Prediction of welding distortions and residual stress in pipe-flangejoint using finite element technique Modelling Simul. Mater. Sci. Eng. 13 455–70

[15] Karlsson R I and Josefson B L 1990 Three dimensional finite element analysis of temperature and stresses insingle-pass butt-welded pipe ASME J. Pressure Vessels Technol. 112 76–84

[16] Wang X, Hoffmann C, Hsueh C, Sarma G and Hubbard C 1999 Influence of residual stresses on thermal expansionbehaviour Appl. Phys. Lett. 75 3294–6

[17] ANSYS User’s Manual 1998 ANSYS User’s Manual SAS IP inc.[18] Lindgren L E, Runnemalm H and Nasstrom M 1999 Simulation of multipass welding of a thick plate Int. J.

Numer. Methods Eng. 44 1301–16[19] Lindgren L E 2001 Finite element modeling and simulation. Part 1. Increased complexity J. Therm. Stresses 24

141–92[20] Goldak J, Chakravarti A and Bibby M A 1984 A finite element model for welding heat sources Metall. Trans. B

5 299–305[21] Abid M and Siddique M 2005 Finite element simulation of tack welds in girth welding of pipe-flange joint Acta

Mech. submitted (DOI:10.1007/s00707-005-0241-3)[22] Kumakura S, Hosokawa S and Sawa T 1998 Bolt load reduction of bolted joints under axial load Proc. 1998

ASME/JSME Joint Pressure Vessels and Piping Conf. (San Diago, California) PVP-vol 367, pp 103–8

Numerical simulation of the effect of constraints on welding deformations and residual stresses in a...

Documents

Transcript of Numerical simulation of the effect of constraints on welding deformations and residual stresses in a...