Measurement of residual stresses in T-plate weldments

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introduced into a sample. It has been suggested by. Leggatt [18] that the extent of the plastic zone …r0† determines the position from the weld toe at which the.
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Measurement of residual stresses in T-plate weldments R. C Wimpory, P. S May, N. P O'Dowd, G. A Webster, D J Smith and E Kingston The Journal of Strain Analysis for Engineering Design 2003 38: 349 DOI: 10.1243/03093240360692931 The online version of this article can be found at: http://sdj.sagepub.com/content/38/4/349

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349

Measurement of residual stresses in T-plate weldments R C Wimpory1 , P S May1** , N P O’Dowd1* , G A Webster1 , D J Smith2 and E Kingston2 Department of Mechanical Engineering, Imperial College London, UK 2 Department of Mechanical Engineering, University of Bristol, UK

1

Abstract: Tensile welding residual stresses can, in combination with operating stresses, lead to premature failure of components by fatigue and/or fracture. It is therefore important that welding residual stresses are accounted for in design and assessment of engineering components and structures. In this work residual stress distributions, obtained from measurements on a number of ferritic steel T-plate weldments using the neutron diffraction technique and the deep-hole drilling method, are presented. It has been found that the residual stress distributions for three different plate sizes are of similar shape when distances are normalized by plate thickness. It has also been found that the conservatisms in residual stress pro®les recommended in current fracture mechanics-based safety assessment procedures can be signi®cantÐof yield strength magnitude in certain cases. Based on the data presented here a new, less-conservative transverse residual stress upper bound distribution is proposed for the T-plate weldment geometry. The extent of the plastic zone developed during the welding process has also been estimated by use of Vickers hardness and neutron diffraction measurements. It has been found that the measured plastic zone sizes are considerably smaller than those predicted by existing methods. The implications of the use of the plastic zone size as an indicator of the residual stress distributions are discussed. Keywords: residual stresses, T-plate weldments, fatigue, fracture, ferritic steel, neutron diffraction method, deep-hole drilling method

NOTATION C dhkl ¯ dhkl DHD E E110 E211 FE FWHM HAZ HV ILL ND

constant used in the de®nition of weld plastic zone size spacing of the atomic plane hkl initial spacing of the atomic plane hkl deep-hole drilling Young’s modulus Young’s modulus for the (110) plane Young’s modulus for the (211) plane ®nite element full-width half-maximum heat-affected zone Vickers hardness Institute Laue Lagevin neutron diffraction

The MS was received on 26 November 2002 and was accepted after revision for publication on 19 May 2003. * Corresponding author: Department of Mechanical Engineering, Imperial College London, South Kensington Campus, London SW7 2AZ, UK. ** Present address: Atkins Consultants, London, UK.

PWHT q r0 t w wa x, y, z

post-weld heat treatment weld arc power weld plastic zone size weld plate thickness weld plate width width of attachment coordinate directions

Ddhkl

change in spacing of the atomic plane hkl due to a residual stress ®eld change in Bragg peak angle due to a residual stress ®eld strain Cartesian components of strain weld process ef®ciency Bragg peak angle corresponding to the plane hkl neutron wavelength Poisson’s ratio Poisson’s ratio for the (110) plane Poisson’s ratio for the (211) plane Cartesian components of stress yield stress principal stresses weld travel speed

Dyhkl e ex , ey , ez Z yhkl l n n110 n211 sx , sy , sz sYP s1 , s2 , s3 u

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1

R C WIMPORY, P S MAY, N P O’DOWD, G A WEBSTER, D J SMITH AND E KINGSTON

INTRODUCTION

Welding residual stresses are considered in safety assessment procedures, such as British Energy code R6 [1] and BS 7910 [2]. In these procedures, the residual (secondary) stresses are superimposed on the applied (primary) stress in conjunction with a plasticity interaction factor, which is dependent on the extent of plasticity generated in the component. Recommendations are provided in the procedures with regard to the magnitude and distribution of the residual stress ®elds that should be assumed for a range of welded geometries when measurements are not available. Also, allowances for post-weld heat treatment (PWHT) are indicated. Only limited residual stress data are available for welded T-plate joints [3] and there is therefore a strong need for further measurements on this weld geometry in order to provide a range of distributions for use in failure assessments. Allen et al. [4] carried out what is believed to be the earliest neutron diffraction investigation on Tplate weldments. However, the complete stress tensor was calculated only at two positions for each specimen. Perhaps the most relevant measurements are those of Holden et al. [5] who measured residual stress distributions in a 25 mm thick ferritic steel T-plate weld of 350 mm length using the neutron diffraction method and of Cheng and Finnie [6] who used the crack compliance method to measure the residual stress distribution in a 166 mm thick A533-B steel plate welded to a 51 mm attachment. In this work, residual stress distributions in

Fig. 1

T-plate weldments made from the steel BS EN 10025 Grade S355 J2G3 [7] have been obtained using the neutron diffraction (ND) and deep hole drilling (DHD) methods [8, 9]. The neutron diffraction measurements were carried out at three institutions, the Institute Laue Langevin (ILL), Grenoble, France, the NFL neutron facility of the University of Uppsala, Studsvik, Sweden, and the ISIS neutron source at the Rutherford Appleton Laboratory, UK. Measurements have been made on three specimen sizes in the as-welded and heat-treated conditions, in order to allow a comprehensive picture of residual stress distributions in T-plate joints to be developed.

2

MATERIAL AND SPECIMEN SPECIFICATIONS

T-plate weldments of 25, 50 and 100 mm thickness were manufactured from the steel BS EN 10025 Grade S355 J2G3. This is a low-carbon ferritic steel and represents a group of materials commonly used in the offshore and power generation industries. Uniaxial tensile tests were carried out on 25 and 50 mm plate material to verify the material properties prescribed by B3 EN 10025. The specimens were machined from the plate material with the centre-line of the sample parallel to the rolling direction of the plate. It is seen in Fig. 1 that the material shows an

Uniaxial stress±strain curves obtained from 25 and 50 mm plate material, BS EN 10025 S355 J3G3 Downloaded from sdj.sagepub.com at University of Limerick on July 31, 2014

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Table 1

Speci®c tensile properties for BS EN 10025 S355 J2G3

Source 25 mm plate 50 mm plate BS EN 10025

sYP (MPa)

su (MPa)

E (GPa)

358 348 345/335

510 515 490±630

212 212 Ð

upper yield strength of approximately 400 MPa with an associated LuÈders plateau (up to approximately 2 per cent strain). Measured material properties are provided in Table 1. The quoted values from BS EN 10025 [7] are also included for comparison. The values for the yield stress, sYP, 345/335 MPa from BS EN 10025, are the minimum values speci®ed for 25 and 50 mm plate respectively, and the two values for the ultimate tensile stress …su † indicate the range within which the ultimate stress is required to lie in the standard.

3

WELD DETAILS

Manual metal arc (MMA) welding was performed at TWI to manufacture the T-plate weld samples of base plate thickness w ˆ 25, 50 and 100 mm from the steel (see Fig. 2). In all cases the thickness of the attachment, wa ˆ w. Full details of the welding procedure are provided in reference [10], but a short summary of the most relevant aspects is given here. The welding was carried out with a welding consumable (electrode) Oerlikon Tenacito 38R, which is in common use in the offshore industry. The yield stress and ultimate tensile strength of the weld material are 492

Fig. 2

351

and 590 MPa respectively [10]. The 25 mm T-®llet weld, shown in Fig. 2b, was manufactured from 300 mm length plate with eight weld passes (four on each side of the weld). Partial penetration T-plate welds were made from the 50 and 100 mm plate thickness (see Figs 2a and c). These were manufactured with 18 and 90 weld passes respectively. A summary of the weld parameters for the different sample geometries is presented in Table 2; Fig. 3 shows the welding sequence. An alternating depositioning sequence was used for all the samples to minimize distortion during the welding process. The 25 and 50 mm plates were restrained by being clamped during welding to rigid strongbacks. The use of strongbacks was not deemed necessary for the 100 mm weld because of its intrinsic rigidity. To inhibit cracking, heating blankets were used to ensure that the temperature of the plates never dropped below 250 8C during the welding process. The majority of the T-plate welds were subsequently cut into approximately 12.5 mm thick slices to provide specimens for the non-destructive neutron diffraction measurements. A further 100 mm thick slice of the 100 mm T-plate weld was prepared for the semidestructive deep-hole drilling method.

4

RESIDUAL STRESS MEASUREMENT METHODS

Residual stress measurements have been carried out using the neutron diffraction method and the deep-hole drilling method. These methods are described brie¯y below.

Cross-sections of T-plate weld samples used in this investigation (all dimensions in mm) Downloaded from sdj.sagepub.com at University of Limerick on July 31, 2014

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R C WIMPORY, P S MAY, N P O’DOWD, G A WEBSTER, D J SMITH AND E KINGSTON

Table 2

4.1

Weld parameters for different sample geometries

Plate thickness (mm)

Weld type

Restraint

Passes

Current (A)

Voltage (V)

Heat input (kJ/mm)

Weld area/(wwa )

25 50 100

T-®llet T-butt T-butt

Yes Yes No

8 18 90

250 170±240 170±240

24 21±23 21±23

2.0±2.5 2.0±2.5 1.8±2.2

0.18 0.11 0.13

The neutron diffraction (ND) technique

Diffraction methods for measuring residual stress can be used to determine nondestructively the stress state inside a sample, by measuring changes in lattice spacing from the `unstressed’ state. Neutrons have a penetration depth of several cm in most metals, allowing the stress state deep inside a sample to be determined [11]. When illuminated by radiation of wavelength similar to the interplanar spacing, crystalline materials diffract this radiation as distinctive Bragg peaks. The angle, y, at which any given peak occurs can be calculated using

Fig. 3

Bragg’s law of diffraction: 2dhkl sin yhkl ˆ l

…1a†

where l is the wavelength of the radiation, dhkl is the spacing of the lattice plane responsible for the Bragg peak for a given …hkl† plane and yhkl is the corresponding Bragg angle. The peak is observed at an angle of 2yhkl from the incident beam. In practice the angle, 2yhkl , is obtained through the use of `peak ®tting’ routines, which generally also supply the associated uncertainty in the value of 2yhkl related to the goodness of ®t. Thus if

Geometry (not to scale) and welding sequence for (a) 25 mm, (b) 50 mm and (c) 100 mm T-plate specimens. (From reference [10]) Downloaded from sdj.sagepub.com at University of Limerick on July 31, 2014

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the neutron wavelength l is known and the diffraction angle yhkl is measured, the lattice spacing dhkl may be 0 obtained from equation (1a). Strain …Ddhkl =dhkl , where 0 is the unstressed lattice spacing) can subsequently be dhkl calculated from the differentiated form of the Bragg equation:  eˆ

Ddhkl 0 dhkl

´ ˆ ¡ Dyhkl cot yhkl

…1b†

and the stress ®elds obtained directly from the linear elastic properties of the material and the measured strains in the relevant directions. Neutrons can be generated in a nuclear ®ssion reactor or at a spallation source. Generally at reactor sources, a monochromatic beam of neutrons is obtained whereas at spallation sources a polychromatic beam is employed. Therefore at a reactor source a single lattice plane is used to obtain the strain in a material, while for a spallation source, strain is normally determined from many planes using the Rietveld re®nement method [11]. In general, measurements in six directions at a point are required to de®ne completely the stress state. However, when the principal stress directions are known, three orientations suf®ce and when the principal directions coincide with the coordinate measurement directions x, y and z, the principal stresses s1 , s2 and s3 in terms of the strains ex , ey , ez become E ‰…1 ¡ n†ex ‡ n…ey ‡ ez †Š …1 ‡ n†…1 ¡ 2n† E ‰…1 ¡ n†ey ‡ n…ex ‡ ez †Š s2 ˆ sy ˆ …1 ‡ n†…1 ¡ 2n† E ‰…1 ¡ n†ez ‡ n…ex ‡ ey †Š s3 ˆ sz ˆ …1 ‡ n†…1 ¡ 2n†

s1 ˆ sx ˆ

…2† where E is the elastic modulus and n is Poisson’s ratio. Note that equations (2) can be used to obtain sx , sy and sz , regardless of whether these are the principal stresses. The values of E and n used in equations (2) generally depend on the type of source used. In this work the bulk properties E ˆ 212 MPa and n ˆ 0:3 have been used for the data from the spallation source, as strains are averaged over a number of lattice planes. For the monochromatic measurements, the values used were those for the (211) plane (which is the recommended plane for ferritic steels [11]), E211 ˆ 224 MPa, n211 ˆ 0:28. These values have been obtained by taking the average of the Voigt and Reuss solutions for the linear elastic constants [12]. [One measurement was made using the (110) plane rather than the (211) plane but E110 ˆ E211 and n110 ˆ n211 so the same values are used for all the measurements.]

353

4.1.1 Experimental procedure for ND measurements Neutron diffraction measurements were carried out on the 25 mm T-plate weld using two monochromatic neutron sources and one polychromatic source. The monochromatic sources were the Institut Laue Langevin (ILL), Grenoble, France, and the NFL neutron facility of the University of Uppsala, Studsvik, Sweden, whereas the polychromatic source used was that at the ISIS neutron facility at the Rutherford Appleton Laboratory, UK. Neutron diffraction measurements were also made on the 50 and 100 mm T-plate welds using the monochromatic neutron source at the ILL. In all neutron experiments 12.5 mm weld slices were used in order to reduce the neutron path length (which controls the amount of neutron beam time required) in the experiments. Note that slicing the specimens in this way may relax in-plane as well as out-of-plane stresses. However, provided in-plane shear stresses are not signi®cant, the specimen is unrestrained and no reverse yielding occurs during cutting; any in-plane stress redistribution is expected to be small [13]. Figure 4 shows a schematic of a typical T-plate sample with the measuring line of main interest indicated. The majority of the measurements were carried out on the centre-line of the sample, i.e. z ˆ t=2 where t ˆ 12:5 mm. In order to identify the region of interest, the neutron beam was masked to provide a small sampling volume at each experimental point. Typically a 26262 mm3 sampling volume was used and measurements were made at about 12 locations across a specimen width. Strains were measured in three directions, i.e. the transverse …x†, normal …y† and longitudinal …z† directions (see Fig. 4), which allows the three components of stress in these directions to be determined via equations (2). The orientation of the sample in the instrument for the longitudinal and transverse measurements was as shown in Fig. 4. The stresses in the normal direction were measured with the sample rotated by 908 about the longitudinal axis in Fig. 4. Further details of the measurements are given in Table 3. Reference measurements were made in the parent material at an extremity of the sample to obtain the lattice spacing for the unstressed material. The strain at a point is then measured relative to this `strain-free’ lattice spacing. Note that this approach does not take into account local changes in strain-free lattice spacing due to intergranular/ interphase strains at the micro level [13, 14].

4.2

The deep-hole drilling (DHD) method

The deep-hole drilling method was developed as an extension of the standard hole drilling technique to allow the full through-thickness stress ®eld of a specimen to be obtained [8, 9]. This procedure can be divided into four steps: (a) a smooth reference hole is drilled into

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R C WIMPORY, P S MAY, N P O’DOWD, G A WEBSTER, D J SMITH AND E KINGSTON

Fig. 4 Table 3

T-plate weld with measurement directions indicated, where the origin of the axes is at the weld toe

Experimental parameters for measurement on T-plate welds. The sampling volumes are de®ned as input width 6 input height 6 output width

Plate width (mm)

Site

Plane hkl

Sampling volume (mm3 )

2y [equation (1)] (deg)

Wavelength AÊ

E (GPa)

v

25 25 25 25 25 50 100

ISIS ILL Studsvik ILL ILL ILL ILL

Rietveld 211 211 211 110 211 211

26261:4 26261:2 26262 26161 2:56161 26261:2 16161:2

90 109.4 93.5 109.4 119.95 109.4 109.4

Polychromatic 1.91 1.71 1.91 3.50 1.91 1.91

212 224 224 224 224 224 224

0.3 0.28 0.28 0.28 0.28 0.28 0.28

the sample, (b) the diameter of the reference hole is measured at different depths and angles, (c) a column of material containing the reference hole (the `core’) is extracted [usually by electric discharge machining (EDM)] and (d) the diameter of the reference hole is measured again at the same locations to obtain the change in diameter as a function of angle and depth. This change in diameter of the reference hole is then related to the in-plane residual stress in the sample. It is assumed that the stress relieved by the introduction of the initial reference hole is negligible and the ®nal cylinder containing the reference hole is completely stress free after it is removed. The out-of-plane (through-thickness) stresses in the sample can also be calculated if the axial distortion of the core is recorded.

A reference hole was drilled through the bushes and specimen by a single pass with a gundrill of 3.175 mm diameter and the diameter of the reference hole was measured using an air probe at 0.2 mm steps along its whole length. The diameter measurement was then repeated at 108 intervals. Having trepanned the core the hole diameter was re-measured along its whole length, again at 0.2 mm steps and at 108 intervals. No measurements of the axial distortion of the core were made. The measured change in diameter of the hole provides longitudinal and transverse stress distributions, which are reported in the next section. For DHD measurements, a typical error in the measured residual stress for a material with a Young’s modulus of 212 GPa is + 33 MPa.

5 4.2.1

Experimental procedure for DHD measurements

The geometry of the specimen and measurement location is illustrated in Fig. 5. For the measurement carried out here the reference hole was introduced into the sample by gun drilling, which allows for accurate drilling of straight deep holes. Bushes were attached to the sample to allow for any bell mouthing associated with the gun drilling and to provide an accurate reference point for the measurement of the hole.

RESIDUAL STRESS PROFILES (AS-WELDED)

Here the measured residual stress distributions of the w ˆ 25, 50 and 100 mm T-plate welds in the as-welded condition are presented.

5.1

Neutron measurements

The multiple measurements taken on the 25 mm weld enabled an average of ®ve sets of data to be calculated.

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355

Fig. 5 Location of deep-hole drilling measurements

The average distributions from the ®ve sets of measurement are presented in Fig. 6. The individual data sets are shown in Fig. 7. A modi®ed Bayesian average of the data was used [15, 16], which reduces the in¯uence of outlying data points within a data set and therefore provides a more realistic estimate of the true distribution. An estimate of the uncertainty of each data set (residual stress distribution) can be obtained by comparing the data set with this average distribution to provide an estimate of the overall uncertainty in the data. In Table 4 the standard deviations of each individual data set about the average are presented. The standard deviations obtained from the neutron Bragg peak ®tting routines are also provided for each data set (this is the average for all points in each data set). It appears that the uncertainty obtained from the peak ®tting routines is

Fig. 6

greater than that from the modi®ed Bayesian analysis. The overall uncertainty in the measurements indicated in Fig. 6 is approximately 20 MPa, which is slightly less than the average of the uncertainties of the peak ®tting routines in Table 4. From Fig. 6 the peak transverse stress in the 25 mm Tplate weld is estimated to be approximately 120 + 20 MPa at the weld toe …&34 per cent of the material yield stress). It appears that the stress ®eld is almost hydrostatic tension …s x &sy &sz † close to the toe of the weld. The transverse stress distribution …sx † shown in Fig. 6 simultaneously satis®es the force and bending moment equilibrium balance to within + 5 MPa (assuming that the transverse stress is independent of z). Measurements on 12.5 mm thick slices of 50 and 100 mm T-plate welds were carried out at ILL using the

Bayesian averages of measured stress distributions in the 25 mm T-plate weld Downloaded from sdj.sagepub.com at University of Limerick on July 31, 2014

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R C WIMPORY, P S MAY, N P O’DOWD, G A WEBSTER, D J SMITH AND E KINGSTON

Fig. 7

Individual data sets of 25 mm T-plate weld residual stresses in (a) the transverse direction, (b) the longitudinal direction and (c) the normal direction

Table 4

Experimental uncertainties in ND measurements on the 25 mm specimen

Laboratory

Crystal plane …hkl†

Standard deviation of data set about average (MPa)

ISIS ILL Studsvik ILL ILL

Rietveld 211 211 211 110

+14 +20 +16 +18 +26

experimental parameters quoted in Table 3. Analyses have indicated that a ‡ 33 MPa shift in transverse stress was required to simultaneously satisfy the force and bending moment equilibrium conditions for the 50 mm weld (indicating a possible inaccuracy in the reference `strain-free’ lattice spacing estimation) and ‡ 5 MPa for the 100 mm weld. Figures 8a and b show the three stress components for the 50 and 100 mm specimens respec-

Standard deviation obtained from peak ®tting routines (MPa) +25 +22 +23 +22 +25

tively. The transverse distributions have been shifted to satisfy equilibrium in each case. A peak stress of 230+25 MPa …&64 per cent of the material yield stress) was measured at 4 mm from the weld toe in the 50 mm weld and 270 + 40 MPa …&75 per cent of the material yield stress) at the weld toe for the 100 mm weld. In Fig. 9 the stresses are replotted with distances normalized by the plate width, w. It may be noted that for all

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Fig. 8

Fig. 9

Residual stress distributions: (a) 50 mm and (b) 100 mm specimens

Comparison of measured transverse stress distributions in all welds

specimens, the transverse stress distributions are similar and become compressive at approximately the same normalized distance, y=w ˆ 0:2, from the weld toe. The similarity in the transverse residual stresses may be due to the geometrical similarity of the welds (though the smallest weld is a ®llet weld, while the other two are partially penetrating welds) and to the cross-sectional area of the weld metal being in approximately the same proportion for all three weld sizes (see Table 2). All three stress components in the 50 mm specimen show a drop near the weld toe, which is not observed in the 25 and 100 mm specimens. This may be due to small differences between the weld geometries local to the toe.

5.2

357

Deep-hole drilling measurements

Deep-hole drilling measurements were carried out on the 100 mm T-plate specimen. Two sets of measurements

were performed, the ®rst set being incomplete due to dif®culties with the experimental apparatus. The residual stress distributions in the transverse and longitudinal directions obtained from these measurements are shown in Figs 10 and 11 respectively. Equilibrium calculations, based on the assumption that the transverse stress is uniform in the longitudinal …z† direction, have shown that the transverse stress distribution obtained from the second deep-hole drilling measurement requires a ‡ 38 MPa offset to satisfy force and bending moment equilibrium and this shifted curve is also included in Fig. 10a. (The magnitude of the shift is close to the + 33 MPa experimental uncertainty for these measurements.) No equilibrium calculation can be performed on the ®rst data set, as the data are incomplete. It may be noted that the two hole drilling measurements agree very well when the second data set is shifted. In Fig. 10b the deep-hole drilling measurements are compared with those obtained from neutron diffraction. Excellent agreement between the two methods is clear, with a peak stress of approximately 240 MPa (67 per cent of the material yield stress) at a distance of 3 mm from the weld toe. The close agreement between the ND and DHD measurements seen in Fig. 10b provides con®dence in the accuracy of both measurement techniques. It also suggests that the in-plane residual stress ®eld has not been appreciably affected by the slicing of the sample for the ND measurements (note that the DHD measurements were on a 100 mm long plate). A similar result has been observed in ND measurements on butt-welded plates of different thickness [13]; i.e. in-plane stresses were unaffected by slicing a complete cross-section parallel to the transverse direction (perpendicular to the welding direction). Figure 11, however, shows that the out-of-plane stress (longitudinal direction) ®eld is strongly affected by slicing the specimen. The ND results exhibit approximately zero stress except near to the weld toe. Also it

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R C WIMPORY, P S MAY, N P O’DOWD, G A WEBSTER, D J SMITH AND E KINGSTON

Fig. 10

Transverse stresses in the 100 mm T-plate (a) obtained by DHD, (b) ND measurement on a 12.5 mm slice and DHD measurement

can be seen that the transverse and longitudinal residual stress pro®les obtained using DHD exhibit a similar form and magnitude (compare Figs 10 and 11), a trend also observed for the ND measurements on the 25 mm specimen (see Fig. 6).

6

6.1

RESIDUAL STRESS PROFILES AFTER POSTWELD HEAT TREATMENT

properties in the heat-affected zone (HAZ) of the weld. To assess the reduction in residual stress due to PWHT, measurements have been made on a 12.5 mm thick slice of the 25 mm T-plate weld in the PWHT condition. The PWHT essentially involved holding the specimen at a temperature of 600 8C for 1 hour, as recommended in BS 5500 [17]. Note that the PWHT period depends on the specimen size; larger or thicker specimens would be heated over longer periods of time.

Post-weld heat treatment

The integrity of a welded joint can often be enhanced through the application of post-weld heat treatment (PWHT), as it may reduce the magnitude of the residual stresses in the component and improve the material

Fig. 11

Longitudinal stresses in the 100 mm T-plate obtained by ND on a 12.5 mm slice and by DHD

6.2

Neutron diffraction results

The neutron diffraction measurements were made at ISIS on the PWHT specimen using similar experimental parameters to those given in Table 3 for the as-received T-plate welds. Due to time restrictions, measurements were only possible in the transverse and longitudinal directions. To estimate the residual stresses in the samples, plane stress conditions were assumed (i.e. longitudinal stress, sz ˆ 0 in accordance, approximately, with the ®ndings in Figs 6 and 8 away from the weld toe). Figure 12 shows a comparison of the residual stresses from the neutron measurements in the transverse direction of the as-welded and PWHT samples. It can be seen that the stresses after PWHT have reduced across the entire cross-section to zero, to within an experimental accuracy of + 20 MPa. BS 7910 [2] recommends that the residual stress distribution after PWHT be assumed to be constant and equal to 20 per cent of the lesser of the parent or weld material yield stress in the transverse direction. This implies that these stresses should be taken to be equal to about 70 MPa for this steel. This value is somewhat higher than the result shown in Fig. 12.

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7.1

Fig. 12

7

Transverse residual stress in 25 mm as-welded and after PWHT

ESTIMATION OF THE PLASTIC ZONE SIZE

During the welding procedure, plastic strains are introduced into a sample. It has been suggested by Leggatt [18] that the extent of the plastic zone …r0 † determines the position from the weld toe at which the residual stress ®rst becomes compressive and estimates of plastic zone size are therefore useful in predicting residual stress distributions for welded joints. In this section, the plastic zone sizes in the welded specimens are determined using Vickers hardness measurements [19] and the full-width half-maximum (FWHM) of the measured neutron diffraction peaks [20]. The values obtained are compared with estimates obtained in references [1] and [2].

Fig. 13

359

Hardness investigation on the welded samples

It is well known that there is an approximate relationship between yield strength (or ¯ow stress) and Vickers hardness (see, for example, reference [21]). By comparing the hardness (HV) and yield strength …sYP † of a wide range of carbon steels (see, for example, reference [22]), it was found that the ratio sYP =HV was in the range 2.4 + 0.4 …sYP in MPa and HV in kgf/mm2 ). Vickers hardness tests on the as-received plate indicate that the appropriate sYP =HV ratio for the steel under examination in this work is 2.0. To assess the variation in material ¯ow stress around the weld toe as a result of the welding procedure, Vickers hardness tests have been carried out. Microstructural examination of the 25 and 50 mm specimens indicates that the size of the heat-affected zone (HAZ), where microstructural changes are most signi®cant, is approximately 5 mm. Outside this region it has been assumed that changes in ¯ow stress are associated with plastic deformation only and therefore any increase in hardness may be interpreted as due to plastic deformation. Vickers hardness measurements were taken on the surface of the T-plate specimens along line YY in Fig. 4. The measured hardness values were scaled by 2.0 to give the apparent material yield strength and will be compared with the results from the ND estimates of plastic zone sizes, described in the next section.

7.2

FWHM investigation on the welded samples

In a neutron diffraction measurement, the full-width half-maximum (FWHM) of the Bragg peak can be related to the number of dislocations in a material [20]

Comparison of FWHM (6 1120) measurements with a typical tensile stress±strain curve of the material Downloaded from sdj.sagepub.com at University of Limerick on July 31, 2014

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and hence the plastic strain. Therefore any change in the FWHM from the reference distribution (on an unstressed sample) may be linked to an increase in plastic deformation and thus the plastic zone may be determined as the region where the FWHM is above that for the undeformed material. Figure 13 shows scaled FWHM results for specially prepared pre-strained coupon samples measured at ILL. Further details of these measurements are available in reference [23]. The scaling factor 1120 was chosen so that the resultant stress±strain curve from the FWHM analysis closely approximated the measured uniaxial tensile response of the material. The good agreement seen in Fig. 13 between the scaled FWHM data and the material stress±strain curve supports its use as a predictor of plastic strain. Figure 14 shows plots of scaled FWHM pro®les for the three specimens measured at ILL compared with the results obtained from the Vickers hardness tests. At each

Fig. 14

position the average FWHM of the normal, longitudinal and transverse directions was used. The close agreement between the results from the two independent methods is noted and, given the uncertainties inherent in both techniques in precisely locating the onset of plastic deformations in Fig. 14, the plastic zone size may be estimated as 9+2 mm for the 25 mm weld, 23+5 mm for the 50 mm weld and 15+3 mm for the 100 mm weld. The differences may be associated with the increased number of weld passes between the 50 and 25 mm welds (Fig. 3) and the lower heat input used for the 100 mm weld (see Table 2). Note that these values are not consistent with the size of the tensile transverse residual stress region for the specimens, which are &5, 10 and 20 mm for the 25, 50 and 100 mm specimens respectively (see Figs 6 and 8). Thus it seems the suggestion in reference [18] that the size of the tensile region corresponds with the plastic zone size is not valid here. However, both the 25 and 50 mm plates were globally

Comparison of variation in ¯ow stress (yield point) estimated from FWHM and Vickers hardness measurements: (a) 25 mm weld, (b) 50 mm weld and (c) 100 mm weld Downloaded from sdj.sagepub.com at University of Limerick on July 31, 2014

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restrained during welding and so may not satisfy the assumption in reference [18] that the restraint is very localized to the weld. Futhermore, it should be pointed out that the assumption that the extent of the zone of residual tension at the weld toe is equal to the plastic zone size will lead to conservative predictions of residual stress (see section 9).

8

COMPARISON OF MEASURED RESIDUAL STRESS DISTRIBUTIONS WITH LITERATURE

In this section a comparison is made between the measured residual stress distributions and relevant distributions found in the literature, for a range of steels with ¯ow stress of 356±421 MPa. Allen et al. [4] carried out ND investigations on 25 mm thick T-plate weldments of a BS 4360 steel. The complete strain tensor was measured at two points on two similar welds. Holden et al. [5] also used the ND method to obtain residual stress distributions in a 25 mm thick ferritic steel T-plate weld while Cheng and Finnie [6] used the crack compliance method to measure the residual stress distribution in a 166 mm thick A533-B steel plate welded to a 51 mm attachment. Finite element (FE) predictions of the transverse residual stress ®eld in a Lloyds LT60 steel 26 mm T-plate weld have been obtained by Mok and Pick [24]. Figure 15 illustrates the residual stress distributions measured in the 25, 50 and 100 mm T-plate welds and the distributions obtained from the literature as described above. Here stresses have been normalized by sYP and distances by w. It is seen that the measured residual stress distribution for the 25 mm weld is

Fig. 15

361

considerably lower in peak stress magnitude than the distributions obtained from the literature, but that the shape of the distribution is similar. As discussed previously, it is believed that slicing the weldment does not signi®cantly alter the in-plane stresses and the observed differences are believed to be due to differences in specimen size and welding conditions, i.e. the use of strongbacks, different heat inputs and interpass heating. The transverse residual stress distributions measured on the 50 and 100 mm T-plate joints show better agreement with those in the literatureÐin particular the close agreement between the 166 mm sample [6] and the 100 mm T-plate sample is noted. The stresses from the Allen et al. [4] specimens (calculated from the strain data provided in reference [4]) are somewhat lower than the other measurements near the weld toe. However, it should be pointed out that for these investigations a relatively large sampling volume …46464 mm3 † was used (compared to a weld width of 25 mm), which will tend to reduce the peak stress magnitude. As discussed previously, if it is assumed that the transverse stress is uniform in the longitudinal …z† direction, then an equilibrium condition can be imposed on the distributions in Fig. 15a. (Note that the FE prediction of Mok and Pick satis®es equilibrium.) Figure 15b shows a comparison of the transverse residual stress pro®les with the same pro®les after equilibrium shift (where necessary). For the Allen et al. data an equilibrium shift is not possible as the stress pro®les are incomplete. Shifts in the pro®les were found to be no more than +0:11sYP to achieve force balance equilibrium. It is seen that the shifted curves now agree more closely (differences are now on the order of +0:25sYP ).

Transverse residual stress distributions measured in 25, 50 and 100 mm T-plate welds compared to data from the literature: (a) as-measured and (b) after satisfying force balance equilibrium Downloaded from sdj.sagepub.com at University of Limerick on July 31, 2014

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9

R C WIMPORY, P S MAY, N P O’DOWD, G A WEBSTER, D J SMITH AND E KINGSTON

COMPARISON WITH RECOMMENDED STRESS DISTRIBUTIONS IN R6 AND BS 7910

British Energy R6 [1] and BS 7910 [2] provide compendia containing representative residual stress distributions for a variety of weld geometries. To assess the conservatism of the recommended residual stress distributions, they are compared to the measured distributions in this section.

R6). However, equation (5) is not used in BS 7910. If equation (4) results in a plastic zone greater than the base plate width, the stress is taken to be equal to the yield strength across the whole specimen thickness. This distribution is referred to here as BS 7910(1). The second transverse residual stress distribution is a polynomial equation representing an upper bound ®t to experimental data and is given by sres ˆ sYP ‰0:97 ‡ 2:3267…y=w† ¡ 24:125…y=w†2 ‡ 42:485…y=w†3 ¡ 21:087…y=w†4 Š

9.1

R6 distribution for T-plate geometries [1]

In the R6 procedure, the recommended throughthickness transverse residual stress distribution in Tplate welded joints consists of an upper bound bilinear function. The peak stress is at the weld toe and equal to the parent material yield stress and reduces linearly to zero at a distance r0 from the weld toe. The distance r0 represents the size of the yielded zone as recommended by Leggatt [18] and can be estimated from  ´ C Zq 1=2 r0 ˆ …3† sYP u where sYP is the parent material yield stress (or 0.2 per cent proof strength of the parent material), Z is a process ef®ciency parameter, q is the arc power (in J/s) and u is the weld travel speed (in mm/s). C is a constant that depends on the coef®cient of thermal expansion, Young’s modulus, density and speci®c heat of a material. Typical values of C and Z are included in the R6 document for a range of materials. For ferritic steels the values provided are C ˆ 153 N mm=J and Z ˆ 0:8, resulting in a simpli®ed version of equation (3):  r0 ˆ

´ 122q 1=2 sYP u

…4†

If equation (3) or (4) results in a plastic zone greater than the base plate width …r0 > w†, r0 must be recalculated using r0 ˆ

9.2

1:033C Zq sYP u…w ‡ 0:5wa †

…5†

…6†

This distribution is referred to here as BS 7910(2). R6 stipulates a validity range for the use of their residual stress distributions for T-plate welds as w ˆ 25± 100 mm, sYP ˆ 375±420 MPa and q=v ˆ 1:4 kJ= mm (a rather limited range). However, validity ranges for Tplate welds do not appear to be provided in BS 7910.

9.3

Comparison of measured plastic zone sizes with BS 7910 and R6

Table 5 provides the calculated plastic zone sizes using equations (4). and (5) from R6 and BS 7910. The values here are identical for w > 25 mm and are both given by equation (4). For W ˆ 25 mm, in R6 equation (5) is used to de®ne the plastic zone size while following BS 7910 gives r0 ˆ w. The measured plastic zone sizes are also included in Table 5 (see section 9.2) for comparison. It may be seen that equations (4) and (5) overestimate the plastic zone sizes in all the specimens (conservative assumption). This may be explained by the use of estimates for C and Z in equation (3), which may not be appropriate for this particular steel. Figure 16 shows the measured transverse residual stress distributions in the 25, 50 and 100 mm T-plates respectively compared to the distributions recommended by R6 and BS 7910. The recommended distributions provide a rather poor estimation of the residual stresses in the 25 mm weld (Fig. 16a), both in terms of magnitude and shape, though they are conservative (conservatisms can be as high as yield stress in magnitude at some locations). For the 50 mm weld (Fig. 16b), the R6 and BS 7910(1) distributions provide a better estimate of the residual stress ®eld than BS

BS 7910 distributions for T-plate geometries [2]

BS 7910 provides two transverse residual stress distributions for T-plate joints. The ®rst follows the approach in R6, with the distribution dependent on the size of the plastic zone. When the plastic zone, calculated via equation (3) or (4) is less than the base plate thickness the residual stress is taken to be that of the parent material yield stress level at the weld toe, reducing linearly to zero over the size of the yielded zone (as in

Table 5

Calculated and measured plastic zone sizes in T-plate welds

Base plate thickness w (mm)

R6 (mm)

BS 7910 (mm)

Measured (mm)

25 50 100

23.5 29.2 27.2

25 29.2 27.2

12 + 2 20 + 5 15 + 3

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7910(2), although the former distributions are nonconservative in the region y=w > 0:8. Finally, for the 100 mm weld (Fig. 16c), the R6 and BS 7910(1) distributions provide a good conservative estimate for y=w 4 0:8 with the BS 7910(2) distribution being again somewhat overconservative. It should be pointed out that the recommended residual stress distributions in R6 and BS 7910 are based on limited experimental data and indeed the heat input, q=n, for the T-plate welds in this study is in the range 1.8 kJ/mm 4 q=n 4 2:5 kJ= mm (see Table 2), which is greater than the value of 1.4 kJ/mm quoted in R6, although no such stipulation is given in BS 7910. Furthermore, equation (6) is based on a range of through-thickness transverse residual stress data for a number of joints, including pipe-on-plate joints, tubular joints and only one T-plate joint [3]. The additional measurements presented here imply that the conservatism, in R6 and BS 7910, can be reduced for T-plate welds. Using a modi®ed Bayesian approach [15, 16], the averages of the available data before and after equilibrium balancing are provided in Fig. 17a. Note that the Allen et al. [4] data were not included in the averaging procedure, as they did not follow the trends of the other data. The magnitudes of the error bars shown in Fig. 17a are given by +2 standard deviations about the mean. As shown in Fig. 17a the average experimental data can be represented by a bilinear plot starting from a stress of 0:75sYP at y=w ˆ 0, decreasing to ¡0:3sYP at y=w ˆ 0:275 and increasing to 0:25sYP at y=w ˆ 1:0 (this line captures the average distributions both before and after equilibrium balancing and satis®es force and moment balance to within +0:05sYP ). Also as shown in Fig. 17a the upper limit of the data can be represented by this mean curve displaced by 0:25sYP . Comparisons of all the available T-plate residual stress data with the parametric BS 7901(2) equation and the new upper bound ®t of Fig. 17a are provided in Fig. 17b. The R6 and BS 7910(1) distributions are not included in this ®gure as they do not collapse to a single curve when normalized in this manner (see Fig.16). It may be seen that the BS 7910(2) distribution is conservative in all cases but that a more accurate estimate is provided by the upper bound line.

10 DISCUSSION AND CONCLUSIONS

Fig. 16

Transverse residual stress distributions measured in (a) 25 mm, (b) 50 mm and (c) 100 mm T-plate welds compared with recommendations in R6 [1] and BS 7910 [2]

Residual stress measurements have been carried out using the neutron diffraction and deep-hole drilling methods on ferritic steel weldments, representative of components used in the offshore and nuclear industries. Neutron diffraction measurements have been made at a number of facilities throughout Europe. The measured distributions have been compared with data on similar

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Fig. 17

Transverse residual stress distributions: (a) bilinear estimations of averages and upper bounds of data and (b) transverse residual stress distributions after satisfying force balance equilibrium compared to upper bounds

geometries in the literature and with the recommended distributions in the R6 [1] and BS 7910 [2] procedures. The residual stress distribution below the toe of a Tplate weld has a maximum tensile value close to the weld toe. In the 25 mm T-plate weld the value is estimated to be approximately 120 MPa at the weld toe (34 per cent of the material yield stress), approximately 230 MPa (64 per cent of the material yield stress) at 4 mm from the weld toe in the 50 mm weld and approximately 270 MPa (75 per cent of the material yield stress) in the 100 mm weld at the weld toe. The transverse through-thickness residual stress pro®les of T-plate geometry welds show good agreement in shape when stresses are normalized with respect to yield strength and distances with respect to the plate width. The average experimental data can be represented by a bilinear plot starting from a normalized stress of 0.75 at y=w ˆ 0, decreasing to ¡ 0:3 at y=w ˆ 0:275 and increasing to 0:25 at y=w ˆ 1:0. The distributions given in R6 and BS 7910 provide conservative predications of the residual stress in the Tplates. A more accurate conservative estimate is obtained by adopting the upper bound bilinear plot of Fig. 17. The similarity of form observed in all transverse residual stress data for T-plate weldments is re¯ected in the BS 7910(2) distribution [equation (6)] and the value of sx ˆ syp at y=w ˆ 0 provides a reasonable upper bound. However, equation (6) overestimates the residual stress quite considerably over most of the cross-section. The assumption inherent in the R6 and BS 7910(1) distributions that the residual stress distribution may be based on the plastic zone size does not seem to apply to the T-plate geometries examined here. Indeed, neither the equation used to estimate the plastic zone size nor

the assumption that stress becomes compressive close to the boundary of the plastic zone appear to be valid. However, it should be pointed out that the estimates of the plastic zone size presented here rely to a certain extent on subjective observations (see Fig. 14). Measurements on a PWHT specimen indicate that the transverse residual stress relaxes to approximately zero to within experimental error. BS 7910 recommends that the residual stress distribution after PWHT be assumed to be constant and equal to 20 per cent of the lesser of the parent or weld material yield stress in the transverse direction. This is again conservative, based on the results presented here.

ACKNOWLEDGEMENTS The authors would like to acknowledge Dr M. Daymond at ISIS (UK), Dr T. Pirling at ILL and Dr R. L. Peng at Studsvik (Sweden) for assistance with the neutron diffraction measurements. Financial support for the work was provided by the IMC, HSE, EPSRC and DERA. Helpful input of the industrial sponsors, in particular Dr R.A. Ainsworth, Dr A Stacey and Dr S. Birley, is gratefully acknowledged.

REFERENCES 1 Milne, I., Ainsworth, R.A., Dowling, A.R. and Stewart, A.T. Assessment of the integrity of structures containing defects. CEGB Report R/H/R6-Rev. 4, 2001.

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13 Webster, G.A. and Wimpory, R.C. Residual stresses in weldments. J. Neutron Res., 2001, 9, 281±287. 14 Bonner, N.W., Modlen, G.F., Webster, P.J., Root, J.H. and Holden, T.M. Load sharing and interphase residual microstress in a high carbon pearlitic steel. In Proceedings of Fifth International Conference on Residual Stresses (ICRS-5), LinkoÈping, Sweden, 1997. 15 Webster, G.A. (Ed.) Neutron diffraction measurements of residual stress in a shrink-®t ring and plug. VAMAS Report 38, ISSN 1016±2186, NPL, 2000. 16 Daymond, M.R., Johnson, M.W. and Sivia, D.S. Analysis of neutron diffraction strain measurement data from a round robin sample. J. Strain Analysis, 2002, 37(1), 73±85. 17 BS 5500:1994 Speci®cation for Un®red Fusion Welded Pressure Vessels, 1994 (British Standards Institution, London). 18 Leggatt, R.H. Welding residual stresses. In Proceedings of Fifth International Conference on Residual Stress (ICRS-5), LinkoÈping, Sweden, 1997, pp. 12±25. 19 Ashby, M.F. and Jones, D.R.H. Engineering Materials, Vol. 1, 1980 (Butterworth and Heinemann, London). 20 Swallowe, G.M., Osborn, J.C., Lukas, P. and Vrana, M. Peak pro®le analysis as a measure of substructural evolution with plastic strain in metals. In Proceedings of Fourth European Conference on Residual Stresses, Cluny, France, 1996, pp. 61±67. 21 Tunvisut, K., Busso, E.P., O’Dowd, N.P. and Brantner, H.P. Determination of the mechanical properties of metallic thin ®lms and substrates from indentation tests. Phil. Mag. A, 2002, 82, 2013±2029. 22 MATWEB material property data, http://www.matweb. com. 23 Wimpory, R.C. The development of microstresses in offshore structural steel. ILL Experimental Report 5±2667, 2000. 24 Mok, D. and Pick, R. Finite element study of residual stresses in a plate T-joint fatigue specimen. Proc. Instn Mech. Engrs, Part C: J. Mechanical Engineering Science, 1990, 204 (C2), 127±134.

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