Influence of post weld treatments on the fatigue behaviour of Al-alloy welded joints

Int. J. Fatigue Vol. 20, No. 10, pp. 749–755, 1998  1998 Elsevier Science Ltd. All rights reserved Printed in Great Britain 0142–1123/98/$—see front matter

PII: S0142-1123(98)00045-0

Influence of post weld treatments on the fatigue behaviour of Al-alloy welded joints L. Bertini*‡, V. Fontanari† and G. Straffelini† *Departimento Costruzioni meccaniche e nucleari, University of Pisa, Via Diotisalvi 2, 56126, Pisa, Italy †Dipartimento Ingegneria dei materiali, University of Trento, Via Mesiano, 7738050, Trento, Italy (Received 13 November 1997; revised 5 May 1998 In this paper the influence of different post welding treatments, such as ageing or shot peening, on the fatigue behaviour of Al-alloy welded joints was investigated. The analysed joints were candidates for car structural applications. Several four point bending fatigue tests were conducted on GMAW specimens subjected to different post weld treatments. The residual stress field acting on specimens was measured by the X-ray diffraction (XRD) technique. The results of tests were discussed with the aid of a finite element model of the specimen aimed to calculate the actual fatigue cycle, also taking account of residual stresses and of their redistribution during the test. This allowed to characterize the fatigue resistance of the joints, taking account of the effective stress acting in the region of crack initiation.  1998 Elsevier Science Ltd. All rights reserved (Keywords: Al-alloy; welded joints; residual stress; shot peening; ageing)

heterogeneous microstructure3,4. Indeed, the welding process introduces noticeable metallurgical modifications, residual stress fields and stress concentrations which are likely to significantly affect material behaviour, particularly in the presence of cyclic loading. As regards the base material, it has long been known5–7 that the fatigue strength of Al-alloys (as well as that of other materials, such as steel) can be improved by shot peening, which produces work hardening and roughness modifications of the surface and tends to leave on it a compressive residual stress field. Some authors have also recently observed similar beneficial effects of shot peening in steel welded joints undergoing fatigue loading8–10. It must, however, also be pointed out that some doubt about the effectiveness of shot peening treatments was reported by other authors6,11–15 who observed a relaxation of the residual stress field during cyclic loading. In the present research it was therefore decided to investigate the effects of shot peening on the fatigue performance of Al-alloy welded joints. In addition, the effects of heat treatments aimed to relax residual stresses induced by welding and to increase heat affected zone (HAZ) strength by precipitation hardening phenomena were investigated. Four point bending tests were conducted on GMAW butt welded joints made from 6063 Al-alloy. X-ray diffraction (XRD) technique was employed to measure residual stress fields produced by post weld treatments. Results were analysed with the aid of a finite element (FE) model, operating

NOMENCLATURE ⌬␴nom Nf C m Kt ␴max,nom ⌬␴eff ␴m,eff T␴

Stress amplitude due to external load Number of cycles to failure Coefficient of the Wo¨hler relationship Exponent of the Wo¨hler relationship Stress concentration factor Maximum stress due to external load Effective stress amplitude effective mean stress Scatter in the fatigue results

INTRODUCTION The car industry has undergone, in the last decades, a strong evolution in production technology aimed to reduce vehicle weight in order to limit fuel consumption and to increase performance. In this context the possibility has been considered of developing a vehicle space frame made of Al-alloys. To this purpose the use of cast nodes linked by welded extruded profiles appeared to give the best perspectives1,2. This solution presents two main advantages: a noticeable weight reduction and a simplification in the production process due to the limited number of components. However, the welded joints are often critical points for structural strength because of their high stresses and ‡Corresponding author.

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Figure 1

Geometry of the specimen and configuration of the fatigue tests

in an elasto-plastic regime. This model, which also accounted for the presence of residual stresses, was aimed to give an estimate of the actual stress cycle acting at the weld root. EXPERIMENTAL TECHNIQUE Four point bending fatigue tests were conducted on the specimen shown in Figure 1. Specimens were machined from 6063 Al-alloy sheets and welded in the double V configuration using the GMAW technique. The 4043 Al-alloy was used as the welding metal. Aswelded specimens and specimens subjected to four different post weld treatments were employed, as shown in Table 1, which also reports resulting metallurgical states and yield strength. As regards the HAZ, a careful characterization was performed through microhardness measurements and TEM microstructural analysis. The microhardness variations with distance from the weldment for different material states are compared in Figure 2. The HAZ extension was found to be about 6 mm for all materials with a minimum hardness at about 4 mm. Moreover it appeared that the solution and ageing treatment produced a nearly complete homogenization of material properties. These results were also confirmed by the TEM analysis4,16, which showed, in materials characterized by a T4 structure, the formation of incoherent precipiTable 1

Figure 2 Microhardness profiles [Vickers HV(0.05)] shown by the different materials

tates near the fusion zone, their take up into solution in a region within a 4 mm distance from the weld, and the presence of coherent and semi-coherent precipitates, connected with the preceding heat treatments, in the base material. On the contrary, for the sol. + ag. material, the solubilization followed by an ageing treatment determines homogenization of the microstructure inducing

Conditions of joint production: parameters of the thermal and mechanical post welding treatments

Post welding As-welded (A.W.) Shot peening 1 (S.P.1) Shot peening 2 (S.P.2) Stress relieved (S.R.) Solubilised and aged (sol. + ag.)

No tests

Treatment conditions

Metallurgical state

Yield strength (MPa)

25 15 15 15 15

Intensity 8N Intensity 7A 4 h at 180°C 55 min at 520°C + 4 h at 175°C

T4* T4* T4* T4* T6

100 140† 140† 100 200

*Outside weld and HAZ. †on the surface, estimated from Vickers hardness.

Fatigue of Al-alloy welded joints

Figure 3

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Indeed the as-welded specimens present high values of tensile residual stress in point A (Figure 3), which are almost completely relieved both in the S.R. and in the sol. + ag. material, while shot peening treatments induce considerable compressive residual stresses. Fatigue tests were performed on a electrohydraulic Instron machine at a frequency of 30 Hz, by applying a stress ratio R = ␴min/␴max = 0.05. The fatigue strength at 2 × 106 cycles was investigated by the staircase method17. Some of related run outs were further tested up to rupture. Tests were interrupted after an increase in specimen maximum deflection of about 0.5 mm; at this stage a major crack had usually developed, covering the entire specimen width and about one fourth of the thickness. Cracks were always observed to nucleate at the weld root, on the tensile side of the specimen, most frequently near its lateral edges.

Scheme for the XRD measurement

RESULTS the transition to a metallurgical state (T6) which is characterized by the formation of coherent precipitates, finely dispersed in the metal matrix. Residual stresses were measured using the XRD technique in the positions shown in Figure 3; results for different material states are reported in Table 2. From these data the following observations can be drawn: 쐌 Measures on A.W. material were only aimed at giving a reference value for appreciating the efficiency of heat treatments: they clearly could not allow a complete characterization of the residual stress field, which is 3D with high gradients. 쐌 Tensile ␴yy values were observed on the A.W. specimen axis, near the notch root; due to equilibrium considerations, this indicates that compressive values are likely to be present near the specimen lateral edges.11,20 쐌 Residual stresses appeared to have undergone a nearly complete relieving for S.R. materials. 쐌 S.P. materials presents similar values of residual stresses on the surface both near to and far from the welding. Based on this observation it appeared reasonable to consider the residual stress field uniform on the specimen surface and dependent on the depth only. The characterization of residual stress field can be considered as rather complete for S.R. and S.P. materials, while it was only a qualitative estimate for A.W. specimens, where high gradients have to be expected. Besides, it can be noted that all the post weld treatments strongly modify the residual stress field. Table 2

An example of observed fatigue test results is reported in Figure 4, for the A.W. material, as stress amplitude due to the external applied loads (⌬␴nom) vs number of cycles to failure (Nf). Run outs of the staircase analysis are also indicated. As data fell into two distinct regions of the fatigue curve having different slopes (as typical for this class of materials), the following procedure was applied in order to derive representative analytical relationships relating applied stress with endurance. The break point between the two regions was first fixed by engineering judgement; then, data pertaining to the short life region were fitted with the Wo¨hler curve relationship: ⌬␴nom = C × N−1/m f

(1)

where C and m are coefficients. This allowed to determine the 10, 50 and 90% failure probability curves reported in Figure 4 (the uncertainty range was assumed to be constant and approximated by its centroid value). As data in the long life region were not usually sufficient to allow a satisfactory evaluation of coefficients in Equation (1) (in most cases only staircase data were available), it was decided to arbitrarily fix the value m = 17, based on literature data.18,19 The 50% failure probability curve was then obtained by

Residual stress values Residual stress ␴yy (MPa)

Material

Point A

A.W. S.R. S.P.1 S.P.2 sol. + ag.

60 10 − 115 − 125 0

Point B — — − 100 − 121 — Figure 4 Fatigue curves and data scatter for the A.W. material

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Principal parameters deducible from the fatigue curves Short life region

Material

A.W. S.P.1 S.P.2 S.R. sol. + ag.

C

906.4 334.2 338.9 1439.1 935.4

m

5.54 11 10 4.55 5.48

imposing the 2 × 106 cycle strength to be equal to the estimate given by the staircase analysis. As regards the uncertainty range reported in Figure 4, this was assumed to be equal to the value calculated for the short fatigue life region. The comparison of the resulting curve with a few tests which were produced up to about 5 × 106 cycles showed a fairly good agreement, ensuring that the applied procedure was likely to give a reasonable representation of material behaviour. The coefficients of fitting curves are reported in Table 3. The 50% failure probability curves for the different materials are compared in Figure 5. A noticeable improvement of fatigue strength due to the shot peening can be observed. This improvement is particularly high in the region of long fatigue life (about 25% as compared to A.W. material), whereas it undergoes a progressive reduction for higher loads. Such a convergence can be explained with the onset of plastic deformations at the notch root for high loads, which tend to reduce the compressive residual stresses. No significant difference between the two types of shot peening treatment was observed. A worsening of the fatigue behaviour is observed for the S.R. material. In this case the heat treatment does not have an effect on material microstructure and simply induces a relaxation of residual stresses. As previously observed residual stresses are likely to be slightly compressive in A.W. material in the region of preferential crack initiation (i.e. nearby specimen lateral edges). This can explain the higher fatigue strength exhibited by A.W. in comparison to S.R. specimens.

Figure 5 Comparison between the P50 fatigue curves of the studied materials

Long life region

Scatter

⌬␴l (MPa) (2 × 106 cycles)

T␴ = 1:

71.4 87.3 88.7 61.1 70.2

␴(P = 90%) ␴(P = 10%)

1:1.18 1:1.17 1:1.17 1:1.17 1:1.18

The figure also shows the effect produced on fatigue strength by the solubilization and ageing (sol. + ag.) treatment. As explained before, this heat treatment induces a precipitation hardening of the material, accompanied by an increment in yield strength from 100 to 200 MPa. This increment will presumably produce a similar increase in fatigue strength, which can explain the observed difference between the S.R. and the sol + ag. curves. As regards the comparison of sol. + ag. and A.W. materials it appears that the increase in basic fatigue strength produced in the former by precipitation hardening is substantially equivalent to the beneficial effects of residual stress fields acting in the latter, so that the fatigue curves of the two materials are very close. ANALYSIS OF RESULTS In order to achieve a quantitative interpretation of the observed behaviours, it was decided to conduct a F.E. analysis of the specimen, aimed to characterize actual fatigue cycles experienced by material under different conditions. The study was limited to S.R. and S.P. materials for which a reliable and rather complete characterization of residual stress was available. As a first step, the effective geometry of the welding was characterized, with particular attention to the fillet radius, which was considered the most significant parameter for stress evaluation. The radius was carefully measured for each specimen by an optical stereomicroscope. The great majority (95%) of values ranged between 0.35 and 0.65 mm, with a mean equal to 0.5 mm. No significant effect of surface treatments on this parameter was observed. A F.E. model of the specimen (Figure 6) was then set up by using the Ansys 5.3 FE code and about 1500 plane strain quadrilateral elements. The welded joint was assumed to be symmetric, so that the model was required to represent only one fourth of the specimen. The calculated influence of the fillet radius on the stress concentration factor is reported in Figure 7. A 1.8 value was obtained for a 0.5 mm fillet radius, which implies an excess of yield strength even for the lower load levels applied during fatigue testing. As a consequence, an elasto-plastic analysis was required in order to estimate actual stress cycles at the notch root. Material behaviour was represented by a bilinear kinematic hardening model based on the assumption that yield surface remains constant in size but translates in the stress space with progressive yielding, with

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Figure 8 Microhardness profile obtained from one of the shot peened specimen Figure 6

F.E. mesh of the specimen

However, as the scope of the analysis was to evaluate the stress cycle near the surface, and as the plastic zone was expected to be very small, these approximations appeared acceptable. The F.E. model was employed to simulate a complete fatigue loading–unloading cycle, with special attention to the residual stress distribution acting at the fillet root after load removal. DISCUSSION

Figure 7

Dependence of the form factor from the curvature radius

elastic and elasto-plastic module fixed at 68 and 31 GPa respectively. The elasto-plastic module was estimated from available true stress true strain curves obtained for a material having a microstructure comparable with that of the HAZ. Yield strength was assumed to be equal to 100 MPa for the S.R. material. Considering the well known empirical correlation between hardness and yield strength, this value was raised to 140 MPa in a thin surface layer for the shot peened materials to account for the hardening produced by the treatment. The width of the layer affected by shot peening (about 250 ␮m) was estimated by microhardness measurements (Figure 8). The presence of residual stresses due to shot peening was simulated simply by imposing a hydrostatic uniform initial strain distribution (I.S.D.)21 to the 250 ␮m surface layer. The I.S.D. was fixed so as to reproduce residual stress values measured on the surface. This procedure clearly cannot ensure an accurate modelling of residual stress distribution within thickness, for which more detailed measurements would be required. Moreover the analysis of the interaction between the pre-existent residual stress field and plasticity obtained in this way is clearly approximate.

The results of calculations confirmed that the zone undergoing plastic strain was very small, as compared to specimen thickness. Moreover, an elastic shakedown was predicted after the first fatigue cycle, with no cyclic plasticity. As a consequence, effective stress cycle parameters (⌬␴eff, ␴m,eff) could be estimated by superimposing the residual stress acting after the first load cycle with the stress due to external load and obtained with an elastic analysis. Examples of calculated residual stress distribution after the first cycle are reported in Figure 9 for the

Figure 9 Residual stress in the S.P.1 material before and after the application of first loading cycles with different amplitude

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shot peened materials and in Figure 10 for the S.R. material, for different external load levels. As regards Figure 9 it is interesting to observe the decrease in compressive residual stress as the external load is increased. This effect can be explained by the yielding of material zones immediately underlying the shot peened layer, whose plastic deformation causes a redistribution of residual stresses tending to move the external layer toward tensile stress values. This effect progressively reduces the compressive residual stress field acting on the surface of shot peened specimens as external load is increased and appear to be the responsible of the convergence between the S. R. and the shot peened fatigue curves for high loads (Figure 5). An experimental verification of the distribution given in Figure 9 was carried out by measuring the residual stress on a S.P. run out specimen, which has undergone to 2 × 106 cycles with ⌬␴nom = 89 MPa. The measured residual stress on the surface was − 97 MPa, which is slightly lower than the calculated value of − 110 MPa. The difference can be explained by noting that the model given herein does not account for the slow relaxation of residual stresses which occur throughout the entire fatigue life, due to microplasticity and material hysteresis. From Figure 10 it can be observed that, as expected, the residual stresses originated after the first loading cycle, tend to be more compressive as the external load level is raised. A global comparison of experimental results, based on effective fatigue stress cycles is obtained in Figure 11, where ⌬␴eff is reported vs ␴m,eff for fixed number of cycles to failure and for S.P. and S.R. materials. The curves appear to be in reasonable accordance with the typical effect of mean stress on fatigue life and, in particular, with the prediction of the Gerber model. This observation appears as an indication that the proposed analysis procedure allows to predict the effect of different post welding treatments on the welded joint’s fatigue life by estimating the modification produced on the effective stress cycle by variations of the residual stress state. More extensive applications would however be required for a complete qualification of the procedure.

Figure 11 Effect of mean stress on the fatigue life and parabolic interpolation of the data based on the Gerber model17

CONCLUSIONS The fatigue behaviour of welded joints made from Alalloy was studied with the aim of determining the effect of different post welding treatments, either thermal (stress relieving and solubilization and ageing) or mechanical (shot peening). The results obtained from the tests indicate the potential advantages that can be induced by the shot peening treatment, in particular when the fatigue behaviour at a high number of cycles (106 cycles or more) is considered. On the contrary, post weld heat treatments do not appear to produce similar beneficial effects. Indeed, a simple stress relieving appears to reduce fatigue strength, probably due to a relaxation of compressive residual stresses acting in the region of crack initiation. On the other hand, the solubilization and ageing treatment, which strongly modifies microstructure, tends to produce an increase in yield strength, also increasing fatigue resistance as compared to stress relieved materials. However, this increase is not sufficient to produce an appreciable advantage, as compared to as-welded joints. It would be interesting to analyse the combined effect of solubilization and ageing treatment and of shot peening, which is planned for future work. A simple F.E. analysis, including the effects of plasticity and its interaction with residual stresses was conducted in order to estimate the effective fatigue curves acting in the zone of crack nucleation. The results, even if still incomplete, appear to indicate that the differences between shot peened and S.R. material could be explained by considering differences in effective stress cycles. REFERENCES 1 2 3

Figure 10 Residual stress in the S.R. material after the application of first loading cycles with different amplitude

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