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Bending fatigue behaviour of differently shot peened Al 6082 T5 alloy
International Journal of Fatigue 26 (2004) 889–897 www.elsevier.com/locate/ijfatigue
Bending fatigue behaviour of differently shot peened Al 6082 T5 alloy M. Benedetti a, T. Bortolamedi a, V. Fontanari a,, F. Frendo b a
Dipartimento di Ingegneria dei Materiali e Tecnologie Industriali, University of Trento, Via Mesiano 77, 38050 Trento, Italy Dipartimento di Ingegneria Meccanica, Nucleare e della Produzione, University of Pisa, Via Diotisalvi 2, 56126 Pisa, Italy
b
Received 7 January 2003; received in revised form 5 December 2003; accepted 17 December 2003
Abstract The fatigue behaviour of the Al 6082 T5 alloy subjected to two different shot peening treatments was studied. The fatigue improvements with respect to the unpeened condition and the influence of the peening intensity on fatigue were discussed accounting for the effects of surface modifications (roughness and strain hardening) and of residual stresses. In particular, the extent of the residual stress redistribution during loading was investigated by means of X-ray diffraction (XRD) measurements. An evolution to a stabilised value of the surface stress, depending on the applied loads, was found. The initial and the stabilised residual stress profiles were considered for discussing the improvement in the fatigue behaviour due to peening. To this purpose, a multiaxial fatigue criterion was adopted to account for the biaxial residual stress field. Accordingly, an attempt to separate the effect of residual stresses from that of the microstructural modifications induced by the plastic deformation during peening was carried out and, in particular, the influence of the surface hardening was highlighted. # 2003 Elsevier Ltd. All rights reserved. Keywords: Shot peening; X-ray diffraction; Residual stress; Surface morphology; Surface strain hardening
1. Introduction The fatigue behaviour of mechanical components depends strongly on the surface microstructure as well as on the stress condition in the surface region. The shot peening treatment, due to the intense plastic deformation introduced in the surface layer, affects both the microstructure and the local stress distribution. A significant beneficial effect on fatigue response is usually obtained, which is essentially due to the introduction of compressive residual stresses in the surface region. Moreover, though a sort of microstructural improvement can be sometime observed such as for the stress induced martensitic transformation of the retained austenite in carburised steels [1], the intense plastic deformation induced by shot impacts leads usually to a worsening of the surface morphology [2– Corresponding author. Tel.: +39-461-882430; fax: +39-461881977. E-mail address: [email protected] (V. Fontanari).
0142-1123/$ - see front matter # 2003 Elsevier Ltd. All rights reserved. doi:10.1016/j.ijfatigue.2003.12.003
4], thus exerting a detrimental effect on the fatigue response of the surface layer, that has been directly correlated to a shortening of the crack initiation lives [5]. On the contrary, the effect of shot peening on microstructure is controversial in the literature. In fact, ambiguous is the effect of surface strain hardening, being the build up dislocation structure responsible either for enhanced resistance to crack initiation (high dislocation density) [6–8] or for some material embrittlement induced by cold work hardening reducing the crack growth resistance [9]. The relative importance of these effects and their interaction was extensively discussed in the literature, leading to different conclusions as regards their effectiveness on the damage mechanisms: the crack initiation and subsequent early crack propagation [4,6,8], the crack closure effect [10–12], the residual stress relaxation [7,13–17]. In particular, as regards the evolution of the residual stress field, several investigations showed that residual stresses do relax during fatigue loading. The extent of stress relaxation depends on applied loads and on the tendency of many
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Nomenclature R stress ratio Rm tensile strength ry0,2 yield strength N number of cycles rP50 (resp.10, 90) fatigue endurance with a failure probability of 50% (resp. 10% and 90%) r0P50, k parameters of the S–N curve Tr scatter of fatigue results Kt stress concentration factor Kf fatigue stress concentration factor Rt mean value of the maximum roughness determined on a specific measurement length ra1,2,3 amplitude of principal stress components rm1,2,3 mean principal stress components rRS1,2,3 residual stress (principal components) rV,A equivalent stress amplitude rV,m equivalent mean stress rW alternate fatigue strength ra stress amplitude rm mean stress rmeas measured residual stress a, b parameters of the equation representing the Sines criterion
materials to exhibit cycle-dependent softening in the surface layer because of the intense cold working introduced by the treatment. Experimental evidences both towards a progressive relaxation of the stress field up to material failure and towards an evolution to a stabilised value were reported. In particular, James [13] proposed a model, that, taking into account the dislocation structure, built up during the peening treatment, predicts an exponential decay of the surface stress value to a stabilised asymptotic value. In the present paper, these aspects are considered for studying the fatigue behaviour of the Al 6082 T5 alloy, subjected to two different shot peening treatments. Peening intensity was varied to have different initial residual stress profiles and surface microstructural conditions. An extensive analysis of the residual stress field was carried out by measuring with the X-ray diffraction (XRD) technique the residual stress profile before the fatigue test, the evolution of the surface residual stress during loading and the residual stress profile at a specific fraction of the specimen fatigue life, corresponding to the onset of a stabilised condition. The dependence of the surface stress relaxation on the initial work hardening and on the loading cycle amplitude was studied. The initial and the stabilised residual stress profiles were used to discuss the improvement in the fatigue response due to the peening treatments in the hypothesis of crack initiation and early crack propagation as fatigue controlling parameters. To this
purpose, a multiaxial fatigue criterion was adopted to account for the residual stress field components.
2. Experimental details 2.1. Material and mechanical testing The experimental work was carried out on the Al 6082 alloy, whose chemical composition is reported in Table 1. The material was supplied in the form of extruded plates, in the T5 heat treatment condition. The tensile properties were determined with specimens machined parallel and orthogonal to the rolling direction (Fig. 1) of the plates, in order to evaluate the extent of material anisotropy. The fatigue characterisation was performed on prismatic bars machined from the same metal sheets oriented in the rolling direction and having the following dimensions (in mm): 18 100 4. Part of these specimens was subjected to shot peening, at two different treatment intensities as reported in Table 2. These treatment intensities were chosen to consider two extreme cases, with deep and shallow residual stress Table 1 Materials chemical composition (wt%) Si
Fe
Cu
Mn
Mg
Cr
Zn
Ti
Al
1.3
0.5
0.1
0.4
1.2
0.25
0.2
0.1
Bal.
M. Benedetti et al. / International Journal of Fatigue 26 (2004) 889–897
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nearly 5104 and 1.5106 cycles. The fatigue curves corresponding to 50% of failure probability, represented by the S–N curve: rP50 ¼ r0P50 N 1=k
ð1Þ
were determined, following the standard procedure [18], by fitting the log(N) vs. log(r) results. The uncertainty range was assumed to be constant and approximated by its centroid value. As a representative value of the scatter, the following expression was used: Tr ¼ 1:rP90 =rP10
ð2Þ
P90, P10 denote, respectively, the 90% and 10% levels of failure probability. 2.2. Residual stress measurements Fig. 1. Geometry of the tensile and fatigue specimens (dimensions in mm).
minima, respectively. The shot peening equipment used a motor-driven bladed wheel, rotating at high speed. The shots were propelled by a bladed wheel using a combination of radial and tangential forces to impart the necessary peening speed (v > 100 m=s) to the shots. Peening intensity is governed by the velocity, hardness and size of the shot pellets, and by the angle at which the stream of shot impinges against the surface of the components. The peening intensity is usually expressed in terms of Almen value, which is correlated with the curvature of a so-called Almen strip specimen peened at the saturation coverage (or more) on one side only. For different peening treatments, different thicknesses of the Almen strip are used; this is indicated by the letter accompanying the Almen value. For instance, Almen A and N values refer to thicknesses of 33 0:5 and 20 0:5 mm, respectively. However, the empirical way to assess the Almen intensity enables only a qualitative comparison among peening treatments and it cannot be correlated with microstructural modifications and residual stresses induced in the material. The fatigue tests were carried out, both for the unpeened material and for the two peening conditions, in the four point bending configuration, with a stress ratio R ¼ rmin =rmax ¼ 0:1, at four different stress levels corresponding to fatigue lives in the range between Table 2 Parameters of the shot peening treatments Shot type
Shot size
Shot hardness
Peening intensity
Angle of impingement
Z 850 (ceramic) B 60 (glass)
0.9 mm
>1000 HV
90
0.15 mm
>1000 HV
Almen 10 A Almen 10 N
v
Coverage 100%
The analysis of the residual stress field induced by the two treatments was carried out by measuring the initial stress profile, the evolution of the surface stress during the fatigue tests and finally the stress profile after stabilisation of the surface residual stress. To this purpose, the XRD technique was adopted. The XRD stress measurements are based on the sin2 w method [19] The crystallographic direction h4 2 2i was chosen in order to ensure the optimisation of the material isotropy and the possibility of obtaining high angle measurements with higher stress sensitivity. The elastic properties of the alloy were used in the calculation by means of the Neerfeld–Hill method [19–21]. A specific procedure for the progressive thinning of the specimen was set up, to allow for the evaluation of the residual stress profile. This procedure is based on the controlled chemical etching by immersion in H3PO4–HNO3/water v solution, maintained at 55 C. The material removal speed was 60 lm/h. Electropolishing was discarded because of the difficulty to obtain a homogeneous material removal. The correction accounting for the effect of the removed layer on the stress field was performed following the indications reported in [19] and using an appropriate algorithm defined as follows [22] ðH rmeas ðzÞ rðz1 Þ ¼ rmeas ðz1 Þ þ 2 dz z z1 ðH rmeas ðzÞ 6z1 dz ð3Þ z2 z1 where H is the thickness plate, z1 is the thickness after the layer removal and rmeas is the measured stress. This method works well for shallow depths (up to few percent of the specimen thickness), or where the stress gradient over the total removed depth is neither too steep nor changing in sign. The total thickness of the removed layer was, respectively, 0.25 and 0.4 mm for the B60-10N and the Z850-10A treatments.
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Several measurements were also carried out for comparison by using the blind hole drilling technique obtaining very similar results [23].
3. Results 3.1. Microstructural features As a consequence of the T5 treatment, Mg2Si and (Fe,Mn)3SiAl12 intermetallic compounds can be observed in the microstructure, which is characterised by grains with an average size of nearly 200 lm, only slightly elongated in the rolling direction (Fig. 2). The effect of the two different treatments on the surface morphology can be observed in Fig. 3. The mean value of the final roughness (Rt) is 9.5 and 17 lm for the B60-10N and the Z850-10A treatments, respectively, the initial value being Rt 1 and 1.6 lm in the longitudinal and transverse direction, respectively. The mean distance peak to peak is 142 and 295 lm, respectively. Following [3], on the basis of these values, the notch effect exerted by the surface dimples can be evaluated, giving similar values for the two materials conditions: Kt ¼ 1 þ 4:0ð17=295Þ1:3 ¼ 1:098; Kt ¼ 1 þ 4:0ð9:5=142Þ
1:3
¼ 1:119;
Z850-10A
ð4aÞ
B60-10N
ð4bÞ
It can be observed that both the treatment conditions determine a very limited stress concentration effect. A comparison between the microhardness profiles is reported in Fig. 4 showing a noticeable difference between the two material conditions, which can be directly correlated to the surface strain hardening and to the depth of the surface layer interested by the plastic deformation. A rough estimate of the surface layer interested by a change in the microhardness with respect to the bulk value (115 2 HV) can be 0.5 mm for the B60-10N and 0.7 mm for the Z850-10A materials.
Fig. 2.
Microstructure of the material after the T5 treatment.
Fig. 3.
Surface morphology due to peening.
Fig. 4. Comparison between microhardness profiles for the two peening treatments.
M. Benedetti et al. / International Journal of Fatigue 26 (2004) 889–897
3.2. Tensile monotonic and four point bending fatigue behaviours The tensile curves did not show particular differences between specimens cut along different orientations with respect to the rolling direction, pointing out a substantial isotropy of the material. The mean values of the principal tensile parameters, determined by averaging the results of five tests, are reported in Table 3. A comparison between experimental results for the different material variants is shown in Fig. 5. The parameters representing the fatigue curves corresponding to 50% of failure probability (Eq. (1)) and the results scatter (Eq. (2)) are reported in Table 4. Both the peening treatments determined a noticeable improvement in the fatigue behaviour. This improvement depends on the applied load, being more remarkable for load levels corresponding to longer fatigue lives, leading to lower values of the slope in the P50 fatigue line. It can also be observed that the more intense peening treatment (Z850-10A) does not present a further improvement of the fatigue behaviour as compared to the B60-10N treatment, but, on the contrary, it seems to give a less positive effect, extending the fatigue life. Table 3 Static mechanical properties of the Al 6082 T5 alloy Young’s modulus (E, GPa)
Poisson’s ratio (m)
ry0.2 (MPa)
Rm (MPa)
Fracture strain (A%)
69 (1)
0.33 (0.01)
280 (5)
300 (3)
13 (1)
Fig. 5. Fatigue curves for the untreated and the two peened material variants.
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Table 4 Principal results of fatigue tests
ro (MPa) K Tr ¼ 1: rP10 =rP90
Untreated
B60-10N
Z850-10A
905 5.4 1:1.14
270 15 1:1.12
370 10.4 1:1.15
3.3. Residual stress analysis The XRD measurements were carried out along two orthogonal directions, with both positive and negative w-tilting, giving nearly the same results, thus confirming the equibiaxiality of the stress field, as expected considering the way the surface treatment was carried out on the material. The obtained stress profiles corresponding to the two different peening intensities are reported for comparison in Fig. 6. The specimens subjected to the more intense peening treatment were characterised by deeper compressive residual stress profile and by higher absolute values of the residual stress on the surface and below the surface. For both the material conditions, the peak stress level under the surface reached a considerable value, ranging from 0.65 to 0.8 of the material yield strength. The transition from compressive to tensile stresses took place for the two materials variants at 0.22 mm for B60-10N and 0.4 mm for the Z850-10A. These values are lower than the depths interested by the increase in microhardness (Fig. 4). The surface values of residual stress were measured during the tests with the XRD technique for all the specimens subjected to fatigue testing, in order to follow the residual stress evolution as a function of the number of applied cycles. To this purpose, the fatigue tests were interrupted after blocks of cycles (1000,
Fig. 6. Comparison between initial residual stress profiles for the two peening conditions.
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Fig. 7. Evolution of the surface residual stress for the B60-10N treatment.
Fig. 9. Residual stress profile after 10,000 cycles compared with the initial one for the Z850-10A.
2000, 5000, 10,000, 50,000). The trends reported in Figs. 7 and 8 represent the ratio between the measured value of residual stress (r) and the corresponding initial value (rini) vs. the number of the applied cycles for different loading amplitudes. As it can be observed, an initial decay of the residual stresses takes place in the first 1000 cycles and a stabilised value is reached in dependence on the applied load, at about 10,000 cycles. These measurements were carried out on few specimens up to the specimens failure, showing no variation with respect to the value measured at 10,000 cycles. The relative relaxation is more pronounced for the most intense peening treatment (Z850-10A), especially at the lower loading levels. Finally, the stabilised residual stress profiles, measured after 10,000 cycles for a specific loading level (r ¼ 124 MPa), are illustrated together with the initial profile in Figs. 9 and 10. It can be observed that the
stress relaxation took place in a very thin surface layer, the subsurface compressive peak being nearly unchanged after the application of the aforementioned loading cycles.
Fig. 8. Evolution of the surface residual stress for the Z850-10A treatment.
Fig. 10. Residual stress profile after 10,000 cycles compared with the initial one for the B60-10N.
4. Discussion To account for the effect of residual stresses on the fatigue behaviour, these can be treated as mean stresses [1,7,25]. For the simulation of the fatigue response, the equibiaxiality of the residual stress field must be considered. Therefore, the Sines criterion can be adopted for modelling the stress condition as discussed in [25]. The load leading to fatigue failure can be evaluated when the following relation is satisfied on the specimen surface: rV;a þ a rV;m ¼ b
ð5Þ
M. Benedetti et al. / International Journal of Fatigue 26 (2004) 889–897
where rV,A and rV,m represents, respectively, the equivalent stress amplitude and the equivalent mean stress: qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2 2 2ffi 1 p ffiffi ffi r1;a r2;a þ r2;a r3;a þ r3;a r1;a ð6Þ rV;a ¼ 2 RS RS r1;m þ r2;m þ r3;m þ rRS 1 þ r2 þ r 3 rV;m ¼ ð7Þ 3 The a and b materials parameters refer to the unpeened conditions and can be approximately evaluated considering the four point bending fatigue curve (R ¼ 0:1) and the ultimate tensile strength of the material. To this purpose, the Gerber criterion was considered as the most suitable for representing the dependence on mean stress: 2 ra rm þ ¼1 ð8Þ rW Rm where rW is the alternate fatigue strength and Rm is the ultimate tensile strength. Data regarding some tests carried out at R ¼ 0:5 lie on the Gerber curves derived using the data determined at R ¼ 0:1, thus confirming the effectiveness of the adopted criterion to predict the material behaviour at different R ratios. The initial and the stabilised values of residual stresses on the surface were considered, for evaluating the fatigue response in the hypothesis that the initiation life takes a predominant part of the total fatigue life. It is worth noting that the XRD surface measurements of the residual stresses represent the mean value over a surface layer of 20 lm. It is therefore reasonable the assumption that the fatigue crack is subjected to these values of residual stresses during the initiation and early propagation stages. Though the stress concentration effect due to surface dimples seems to be very small, according to the values calculated using Eqs. (4a) and (4b), its contribution was introduced in the analysis. To this purpose, the notch sensitivity factor was estimated according to [3] using the mean curvature radius of the dimples as geometric parameters. The estimated values of the fatigue notch factor (Kf) ranges for the two peened conditions between 1.05 and 1.06. The comparisons between the experimental and the calculated mean fatigue curves are reported in Figs. 11 and 12. The slope of the curves calculated considering the initial residual stresses (IRS) differs remarkably from the slope of the experimental curve. Surface residual stress relaxation contributes to a consistent correction of the fatigue curves towards a better representation of experimental results up to 5 105 cycles to failure. However, for longer fatigue lives, the calculated curves determined using the stabilised residual stresses (SRS) can be extrapolated to the IRS curves being the residual stresses nearly unchanged at these loading levels. Therefore, the calculated curves seem to
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Fig. 11. Comparison between experimental and calculated P50 fatigue curves (IRS, initial residual stresses; SRS, stabilised residual stresses).
underestimate the fatigue behaviour of the peened materials at the longer fatigue lives. In particular, this is more evident for the B60-10N treatment condition. In order to explain these results, two hypotheses based on the effectiveness of the residual stress profiles and on the surface work hardening can be formulated: 1. Compressive residual stress profiles strongly influence the extent of crack closure during the fatigue cycle. 2. The surface cold working modifies the local microstructure, thus influencing the microcracks behaviour. It can be observed that the closure stress becomes more and more important, the lower the loading level is, the fatigue crack being closed for a more extended
Fig. 12. Comparison between experimental and calculated P50 fatigue curves (IRS, initial residual stresses; SRS, stabilised residual stresses) for Z850-10A material.
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fraction of the loading cycle. This contributes to an increment of the propagation life. However, this hypothesis, though reasonable to explain the fatigue improvement with respect to the unpeened condition, does not completely explain the observed improvement. The crack closure condition depends, indeed, on the value of the compressive peak and on the depth of the compressive residual stresses. On this basis, the best response should be obtained for the Z850-10A material, which, on the contrary, compared to the B60-10N treatment, shows a less evident improvement in the fatigue behaviour and a steeper fatigue curve. Consequently, the first hypothesis can be seen to be inappropriate, since crack closure affects crack growth rather than crack initiation. At long lives, where the predicted and experimental curves differ most markedly, crack growth is known to comprise typically only some 10% of the fatigue life. Therefore, the second hypothesis can give a better interpretation of the observed material behaviour. Material work hardening, determined by the intense plastic deformation on the surface, leads to the build up of a dislocation structure, that can act as a barrier to crack propagation, influencing the conditions of crack arrest for the materials, but can also have an opposite effect determining an easier microcracks initiation, leading to a diffused damage in the material when the lattice distortion is too much intense. The effect of this microstructural modification, which can be assessed as an activation energy barrier, depending on the level of local work hardening, is more effective the lower the external loading. The dislocations structure is more effective at the lower loading levels because not enough energy is introduced in the system to overcome dislocation barriers. At higher loading levels, the amount of energy introduced by external loading is sufficient for redistributing local strain in the microstructure. This can be correlated also to the evolution of the residual stress profile during loading: by comparing the initial and the corresponding stabilised stress profile, it appears that the surface stress do relax, whereas the peak of compressive stress does not change in terms of value and position with respect to the surface. The surface value seems to follow the model of an exponential decay proposed by James [13], which is based on an appropriate hypothesis of the evolution in the dislocation structure induced by peening. The surface stress relaxation is more pronounced for the most intensively peened material (Z850-10A), especially at the lower loading levels, owing to the extent of initial plastic deformation. On the contrary, at the lower loading levels the B60-10N variant does relax in a very limited extent, thus showing that for this material, at these loading levels, the damage mechanisms are little or no active leading to longer fatigue lives. Consequently, the Z850-10A treatment exerts a lower
improvement with respect to the B60-10N variants, especially at the lower loading levels. 5. Conclusion The principal results of this experimental work can be summarised as follows: 1. Initial stress profiles are differentiated for the two peening conditions in terms of surface residual stress, subsurface stress peak and depth of residual stress profile. 2. During the first part of cyclic loading both materials undergo to stress relaxation, which shows a strong dependence on the external applied load and on the initial surface microstructure. 3. Both material conditions reach early a stabilised value of surface residual stress, depending on the loading level. 4. A preliminary evaluation of the stabilised stress profile shows very little modification of the stress profile, except for a very thin surface layer. 5. Accounting for the residual stress distribution, the improvement in fatigue behaviour due to peening can be approximately estimated considering the residual stress field as a mean stress superposed on the stress field due to external loading. 6. The residual stress relaxation, even though limited to a very thin surface layer, seems to have a significant effect on fatigue behaviour especially at the higher loading levels. A better estimate of the fatigue curve can be obtained considering the stabilised residual stresses (SRS) instead of the initial residual stresses (IRS). 7. The fatigue behaviour of the two peening conditions cannot be completely explained considering the effect of residual stresses. In particular, an underestimation of the fatigue behaviour can be obtained at the lower loading levels. 8. The crack closure effect and the surface work hardening effects seem to have an important influence at these loading values. In particular, the extent of strain hardening can explain the observed differences between the different peening conditions. In the long life regime, where initiation events and the associated dislocation network are important, the proposed model should be improved by incorporating the microstructural modifications induced by shot peening. Accordingly, a dislocation-based modelling, thoroughly accounting for residual stress effects (e.g., [10]), would be a more appropriate route to follow. References [1] Benedetti M, Fontanari V, Hohn B, Oster P, Tobie T. Influence of shot peening on bending tooth fatigue limit of case hardened gears. Int J Fatigue 2002;24:1127–36.
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Bending fatigue behaviour of differently shot peened Al 6082 T5 alloy M. Benedetti a, T. Bortolamedi a, V. Fontanari a,, F. Frendo b a
Dipartimento di Ingegneria dei Materiali e Tecnologie Industriali, University of Trento, Via Mesiano 77, 38050 Trento, Italy Dipartimento di Ingegneria Meccanica, Nucleare e della Produzione, University of Pisa, Via Diotisalvi 2, 56126 Pisa, Italy
b
Received 7 January 2003; received in revised form 5 December 2003; accepted 17 December 2003
Abstract The fatigue behaviour of the Al 6082 T5 alloy subjected to two different shot peening treatments was studied. The fatigue improvements with respect to the unpeened condition and the influence of the peening intensity on fatigue were discussed accounting for the effects of surface modifications (roughness and strain hardening) and of residual stresses. In particular, the extent of the residual stress redistribution during loading was investigated by means of X-ray diffraction (XRD) measurements. An evolution to a stabilised value of the surface stress, depending on the applied loads, was found. The initial and the stabilised residual stress profiles were considered for discussing the improvement in the fatigue behaviour due to peening. To this purpose, a multiaxial fatigue criterion was adopted to account for the biaxial residual stress field. Accordingly, an attempt to separate the effect of residual stresses from that of the microstructural modifications induced by the plastic deformation during peening was carried out and, in particular, the influence of the surface hardening was highlighted. # 2003 Elsevier Ltd. All rights reserved. Keywords: Shot peening; X-ray diffraction; Residual stress; Surface morphology; Surface strain hardening
1. Introduction The fatigue behaviour of mechanical components depends strongly on the surface microstructure as well as on the stress condition in the surface region. The shot peening treatment, due to the intense plastic deformation introduced in the surface layer, affects both the microstructure and the local stress distribution. A significant beneficial effect on fatigue response is usually obtained, which is essentially due to the introduction of compressive residual stresses in the surface region. Moreover, though a sort of microstructural improvement can be sometime observed such as for the stress induced martensitic transformation of the retained austenite in carburised steels [1], the intense plastic deformation induced by shot impacts leads usually to a worsening of the surface morphology [2– Corresponding author. Tel.: +39-461-882430; fax: +39-461881977. E-mail address: [email protected] (V. Fontanari).
0142-1123/$ - see front matter # 2003 Elsevier Ltd. All rights reserved. doi:10.1016/j.ijfatigue.2003.12.003
4], thus exerting a detrimental effect on the fatigue response of the surface layer, that has been directly correlated to a shortening of the crack initiation lives [5]. On the contrary, the effect of shot peening on microstructure is controversial in the literature. In fact, ambiguous is the effect of surface strain hardening, being the build up dislocation structure responsible either for enhanced resistance to crack initiation (high dislocation density) [6–8] or for some material embrittlement induced by cold work hardening reducing the crack growth resistance [9]. The relative importance of these effects and their interaction was extensively discussed in the literature, leading to different conclusions as regards their effectiveness on the damage mechanisms: the crack initiation and subsequent early crack propagation [4,6,8], the crack closure effect [10–12], the residual stress relaxation [7,13–17]. In particular, as regards the evolution of the residual stress field, several investigations showed that residual stresses do relax during fatigue loading. The extent of stress relaxation depends on applied loads and on the tendency of many
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Nomenclature R stress ratio Rm tensile strength ry0,2 yield strength N number of cycles rP50 (resp.10, 90) fatigue endurance with a failure probability of 50% (resp. 10% and 90%) r0P50, k parameters of the S–N curve Tr scatter of fatigue results Kt stress concentration factor Kf fatigue stress concentration factor Rt mean value of the maximum roughness determined on a specific measurement length ra1,2,3 amplitude of principal stress components rm1,2,3 mean principal stress components rRS1,2,3 residual stress (principal components) rV,A equivalent stress amplitude rV,m equivalent mean stress rW alternate fatigue strength ra stress amplitude rm mean stress rmeas measured residual stress a, b parameters of the equation representing the Sines criterion
materials to exhibit cycle-dependent softening in the surface layer because of the intense cold working introduced by the treatment. Experimental evidences both towards a progressive relaxation of the stress field up to material failure and towards an evolution to a stabilised value were reported. In particular, James [13] proposed a model, that, taking into account the dislocation structure, built up during the peening treatment, predicts an exponential decay of the surface stress value to a stabilised asymptotic value. In the present paper, these aspects are considered for studying the fatigue behaviour of the Al 6082 T5 alloy, subjected to two different shot peening treatments. Peening intensity was varied to have different initial residual stress profiles and surface microstructural conditions. An extensive analysis of the residual stress field was carried out by measuring with the X-ray diffraction (XRD) technique the residual stress profile before the fatigue test, the evolution of the surface residual stress during loading and the residual stress profile at a specific fraction of the specimen fatigue life, corresponding to the onset of a stabilised condition. The dependence of the surface stress relaxation on the initial work hardening and on the loading cycle amplitude was studied. The initial and the stabilised residual stress profiles were used to discuss the improvement in the fatigue response due to the peening treatments in the hypothesis of crack initiation and early crack propagation as fatigue controlling parameters. To this
purpose, a multiaxial fatigue criterion was adopted to account for the residual stress field components.
2. Experimental details 2.1. Material and mechanical testing The experimental work was carried out on the Al 6082 alloy, whose chemical composition is reported in Table 1. The material was supplied in the form of extruded plates, in the T5 heat treatment condition. The tensile properties were determined with specimens machined parallel and orthogonal to the rolling direction (Fig. 1) of the plates, in order to evaluate the extent of material anisotropy. The fatigue characterisation was performed on prismatic bars machined from the same metal sheets oriented in the rolling direction and having the following dimensions (in mm): 18 100 4. Part of these specimens was subjected to shot peening, at two different treatment intensities as reported in Table 2. These treatment intensities were chosen to consider two extreme cases, with deep and shallow residual stress Table 1 Materials chemical composition (wt%) Si
Fe
Cu
Mn
Mg
Cr
Zn
Ti
Al
1.3
0.5
0.1
0.4
1.2
0.25
0.2
0.1
Bal.
M. Benedetti et al. / International Journal of Fatigue 26 (2004) 889–897
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nearly 5104 and 1.5106 cycles. The fatigue curves corresponding to 50% of failure probability, represented by the S–N curve: rP50 ¼ r0P50 N 1=k
ð1Þ
were determined, following the standard procedure [18], by fitting the log(N) vs. log(r) results. The uncertainty range was assumed to be constant and approximated by its centroid value. As a representative value of the scatter, the following expression was used: Tr ¼ 1:rP90 =rP10
ð2Þ
P90, P10 denote, respectively, the 90% and 10% levels of failure probability. 2.2. Residual stress measurements Fig. 1. Geometry of the tensile and fatigue specimens (dimensions in mm).
minima, respectively. The shot peening equipment used a motor-driven bladed wheel, rotating at high speed. The shots were propelled by a bladed wheel using a combination of radial and tangential forces to impart the necessary peening speed (v > 100 m=s) to the shots. Peening intensity is governed by the velocity, hardness and size of the shot pellets, and by the angle at which the stream of shot impinges against the surface of the components. The peening intensity is usually expressed in terms of Almen value, which is correlated with the curvature of a so-called Almen strip specimen peened at the saturation coverage (or more) on one side only. For different peening treatments, different thicknesses of the Almen strip are used; this is indicated by the letter accompanying the Almen value. For instance, Almen A and N values refer to thicknesses of 33 0:5 and 20 0:5 mm, respectively. However, the empirical way to assess the Almen intensity enables only a qualitative comparison among peening treatments and it cannot be correlated with microstructural modifications and residual stresses induced in the material. The fatigue tests were carried out, both for the unpeened material and for the two peening conditions, in the four point bending configuration, with a stress ratio R ¼ rmin =rmax ¼ 0:1, at four different stress levels corresponding to fatigue lives in the range between Table 2 Parameters of the shot peening treatments Shot type
Shot size
Shot hardness
Peening intensity
Angle of impingement
Z 850 (ceramic) B 60 (glass)
0.9 mm
>1000 HV
90
0.15 mm
>1000 HV
Almen 10 A Almen 10 N
v
Coverage 100%
The analysis of the residual stress field induced by the two treatments was carried out by measuring the initial stress profile, the evolution of the surface stress during the fatigue tests and finally the stress profile after stabilisation of the surface residual stress. To this purpose, the XRD technique was adopted. The XRD stress measurements are based on the sin2 w method [19] The crystallographic direction h4 2 2i was chosen in order to ensure the optimisation of the material isotropy and the possibility of obtaining high angle measurements with higher stress sensitivity. The elastic properties of the alloy were used in the calculation by means of the Neerfeld–Hill method [19–21]. A specific procedure for the progressive thinning of the specimen was set up, to allow for the evaluation of the residual stress profile. This procedure is based on the controlled chemical etching by immersion in H3PO4–HNO3/water v solution, maintained at 55 C. The material removal speed was 60 lm/h. Electropolishing was discarded because of the difficulty to obtain a homogeneous material removal. The correction accounting for the effect of the removed layer on the stress field was performed following the indications reported in [19] and using an appropriate algorithm defined as follows [22] ðH rmeas ðzÞ rðz1 Þ ¼ rmeas ðz1 Þ þ 2 dz z z1 ðH rmeas ðzÞ 6z1 dz ð3Þ z2 z1 where H is the thickness plate, z1 is the thickness after the layer removal and rmeas is the measured stress. This method works well for shallow depths (up to few percent of the specimen thickness), or where the stress gradient over the total removed depth is neither too steep nor changing in sign. The total thickness of the removed layer was, respectively, 0.25 and 0.4 mm for the B60-10N and the Z850-10A treatments.
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Several measurements were also carried out for comparison by using the blind hole drilling technique obtaining very similar results [23].
3. Results 3.1. Microstructural features As a consequence of the T5 treatment, Mg2Si and (Fe,Mn)3SiAl12 intermetallic compounds can be observed in the microstructure, which is characterised by grains with an average size of nearly 200 lm, only slightly elongated in the rolling direction (Fig. 2). The effect of the two different treatments on the surface morphology can be observed in Fig. 3. The mean value of the final roughness (Rt) is 9.5 and 17 lm for the B60-10N and the Z850-10A treatments, respectively, the initial value being Rt 1 and 1.6 lm in the longitudinal and transverse direction, respectively. The mean distance peak to peak is 142 and 295 lm, respectively. Following [3], on the basis of these values, the notch effect exerted by the surface dimples can be evaluated, giving similar values for the two materials conditions: Kt ¼ 1 þ 4:0ð17=295Þ1:3 ¼ 1:098; Kt ¼ 1 þ 4:0ð9:5=142Þ
1:3
¼ 1:119;
Z850-10A
ð4aÞ
B60-10N
ð4bÞ
It can be observed that both the treatment conditions determine a very limited stress concentration effect. A comparison between the microhardness profiles is reported in Fig. 4 showing a noticeable difference between the two material conditions, which can be directly correlated to the surface strain hardening and to the depth of the surface layer interested by the plastic deformation. A rough estimate of the surface layer interested by a change in the microhardness with respect to the bulk value (115 2 HV) can be 0.5 mm for the B60-10N and 0.7 mm for the Z850-10A materials.
Fig. 2.
Microstructure of the material after the T5 treatment.
Fig. 3.
Surface morphology due to peening.
Fig. 4. Comparison between microhardness profiles for the two peening treatments.
M. Benedetti et al. / International Journal of Fatigue 26 (2004) 889–897
3.2. Tensile monotonic and four point bending fatigue behaviours The tensile curves did not show particular differences between specimens cut along different orientations with respect to the rolling direction, pointing out a substantial isotropy of the material. The mean values of the principal tensile parameters, determined by averaging the results of five tests, are reported in Table 3. A comparison between experimental results for the different material variants is shown in Fig. 5. The parameters representing the fatigue curves corresponding to 50% of failure probability (Eq. (1)) and the results scatter (Eq. (2)) are reported in Table 4. Both the peening treatments determined a noticeable improvement in the fatigue behaviour. This improvement depends on the applied load, being more remarkable for load levels corresponding to longer fatigue lives, leading to lower values of the slope in the P50 fatigue line. It can also be observed that the more intense peening treatment (Z850-10A) does not present a further improvement of the fatigue behaviour as compared to the B60-10N treatment, but, on the contrary, it seems to give a less positive effect, extending the fatigue life. Table 3 Static mechanical properties of the Al 6082 T5 alloy Young’s modulus (E, GPa)
Poisson’s ratio (m)
ry0.2 (MPa)
Rm (MPa)
Fracture strain (A%)
69 (1)
0.33 (0.01)
280 (5)
300 (3)
13 (1)
Fig. 5. Fatigue curves for the untreated and the two peened material variants.
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Table 4 Principal results of fatigue tests
ro (MPa) K Tr ¼ 1: rP10 =rP90
Untreated
B60-10N
Z850-10A
905 5.4 1:1.14
270 15 1:1.12
370 10.4 1:1.15
3.3. Residual stress analysis The XRD measurements were carried out along two orthogonal directions, with both positive and negative w-tilting, giving nearly the same results, thus confirming the equibiaxiality of the stress field, as expected considering the way the surface treatment was carried out on the material. The obtained stress profiles corresponding to the two different peening intensities are reported for comparison in Fig. 6. The specimens subjected to the more intense peening treatment were characterised by deeper compressive residual stress profile and by higher absolute values of the residual stress on the surface and below the surface. For both the material conditions, the peak stress level under the surface reached a considerable value, ranging from 0.65 to 0.8 of the material yield strength. The transition from compressive to tensile stresses took place for the two materials variants at 0.22 mm for B60-10N and 0.4 mm for the Z850-10A. These values are lower than the depths interested by the increase in microhardness (Fig. 4). The surface values of residual stress were measured during the tests with the XRD technique for all the specimens subjected to fatigue testing, in order to follow the residual stress evolution as a function of the number of applied cycles. To this purpose, the fatigue tests were interrupted after blocks of cycles (1000,
Fig. 6. Comparison between initial residual stress profiles for the two peening conditions.
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Fig. 7. Evolution of the surface residual stress for the B60-10N treatment.
Fig. 9. Residual stress profile after 10,000 cycles compared with the initial one for the Z850-10A.
2000, 5000, 10,000, 50,000). The trends reported in Figs. 7 and 8 represent the ratio between the measured value of residual stress (r) and the corresponding initial value (rini) vs. the number of the applied cycles for different loading amplitudes. As it can be observed, an initial decay of the residual stresses takes place in the first 1000 cycles and a stabilised value is reached in dependence on the applied load, at about 10,000 cycles. These measurements were carried out on few specimens up to the specimens failure, showing no variation with respect to the value measured at 10,000 cycles. The relative relaxation is more pronounced for the most intense peening treatment (Z850-10A), especially at the lower loading levels. Finally, the stabilised residual stress profiles, measured after 10,000 cycles for a specific loading level (r ¼ 124 MPa), are illustrated together with the initial profile in Figs. 9 and 10. It can be observed that the
stress relaxation took place in a very thin surface layer, the subsurface compressive peak being nearly unchanged after the application of the aforementioned loading cycles.
Fig. 8. Evolution of the surface residual stress for the Z850-10A treatment.
Fig. 10. Residual stress profile after 10,000 cycles compared with the initial one for the B60-10N.
4. Discussion To account for the effect of residual stresses on the fatigue behaviour, these can be treated as mean stresses [1,7,25]. For the simulation of the fatigue response, the equibiaxiality of the residual stress field must be considered. Therefore, the Sines criterion can be adopted for modelling the stress condition as discussed in [25]. The load leading to fatigue failure can be evaluated when the following relation is satisfied on the specimen surface: rV;a þ a rV;m ¼ b
ð5Þ
M. Benedetti et al. / International Journal of Fatigue 26 (2004) 889–897
where rV,A and rV,m represents, respectively, the equivalent stress amplitude and the equivalent mean stress: qffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffiffi 2 2 2ffi 1 p ffiffi ffi r1;a r2;a þ r2;a r3;a þ r3;a r1;a ð6Þ rV;a ¼ 2 RS RS r1;m þ r2;m þ r3;m þ rRS 1 þ r2 þ r 3 rV;m ¼ ð7Þ 3 The a and b materials parameters refer to the unpeened conditions and can be approximately evaluated considering the four point bending fatigue curve (R ¼ 0:1) and the ultimate tensile strength of the material. To this purpose, the Gerber criterion was considered as the most suitable for representing the dependence on mean stress: 2 ra rm þ ¼1 ð8Þ rW Rm where rW is the alternate fatigue strength and Rm is the ultimate tensile strength. Data regarding some tests carried out at R ¼ 0:5 lie on the Gerber curves derived using the data determined at R ¼ 0:1, thus confirming the effectiveness of the adopted criterion to predict the material behaviour at different R ratios. The initial and the stabilised values of residual stresses on the surface were considered, for evaluating the fatigue response in the hypothesis that the initiation life takes a predominant part of the total fatigue life. It is worth noting that the XRD surface measurements of the residual stresses represent the mean value over a surface layer of 20 lm. It is therefore reasonable the assumption that the fatigue crack is subjected to these values of residual stresses during the initiation and early propagation stages. Though the stress concentration effect due to surface dimples seems to be very small, according to the values calculated using Eqs. (4a) and (4b), its contribution was introduced in the analysis. To this purpose, the notch sensitivity factor was estimated according to [3] using the mean curvature radius of the dimples as geometric parameters. The estimated values of the fatigue notch factor (Kf) ranges for the two peened conditions between 1.05 and 1.06. The comparisons between the experimental and the calculated mean fatigue curves are reported in Figs. 11 and 12. The slope of the curves calculated considering the initial residual stresses (IRS) differs remarkably from the slope of the experimental curve. Surface residual stress relaxation contributes to a consistent correction of the fatigue curves towards a better representation of experimental results up to 5 105 cycles to failure. However, for longer fatigue lives, the calculated curves determined using the stabilised residual stresses (SRS) can be extrapolated to the IRS curves being the residual stresses nearly unchanged at these loading levels. Therefore, the calculated curves seem to
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Fig. 11. Comparison between experimental and calculated P50 fatigue curves (IRS, initial residual stresses; SRS, stabilised residual stresses).
underestimate the fatigue behaviour of the peened materials at the longer fatigue lives. In particular, this is more evident for the B60-10N treatment condition. In order to explain these results, two hypotheses based on the effectiveness of the residual stress profiles and on the surface work hardening can be formulated: 1. Compressive residual stress profiles strongly influence the extent of crack closure during the fatigue cycle. 2. The surface cold working modifies the local microstructure, thus influencing the microcracks behaviour. It can be observed that the closure stress becomes more and more important, the lower the loading level is, the fatigue crack being closed for a more extended
Fig. 12. Comparison between experimental and calculated P50 fatigue curves (IRS, initial residual stresses; SRS, stabilised residual stresses) for Z850-10A material.
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fraction of the loading cycle. This contributes to an increment of the propagation life. However, this hypothesis, though reasonable to explain the fatigue improvement with respect to the unpeened condition, does not completely explain the observed improvement. The crack closure condition depends, indeed, on the value of the compressive peak and on the depth of the compressive residual stresses. On this basis, the best response should be obtained for the Z850-10A material, which, on the contrary, compared to the B60-10N treatment, shows a less evident improvement in the fatigue behaviour and a steeper fatigue curve. Consequently, the first hypothesis can be seen to be inappropriate, since crack closure affects crack growth rather than crack initiation. At long lives, where the predicted and experimental curves differ most markedly, crack growth is known to comprise typically only some 10% of the fatigue life. Therefore, the second hypothesis can give a better interpretation of the observed material behaviour. Material work hardening, determined by the intense plastic deformation on the surface, leads to the build up of a dislocation structure, that can act as a barrier to crack propagation, influencing the conditions of crack arrest for the materials, but can also have an opposite effect determining an easier microcracks initiation, leading to a diffused damage in the material when the lattice distortion is too much intense. The effect of this microstructural modification, which can be assessed as an activation energy barrier, depending on the level of local work hardening, is more effective the lower the external loading. The dislocations structure is more effective at the lower loading levels because not enough energy is introduced in the system to overcome dislocation barriers. At higher loading levels, the amount of energy introduced by external loading is sufficient for redistributing local strain in the microstructure. This can be correlated also to the evolution of the residual stress profile during loading: by comparing the initial and the corresponding stabilised stress profile, it appears that the surface stress do relax, whereas the peak of compressive stress does not change in terms of value and position with respect to the surface. The surface value seems to follow the model of an exponential decay proposed by James [13], which is based on an appropriate hypothesis of the evolution in the dislocation structure induced by peening. The surface stress relaxation is more pronounced for the most intensively peened material (Z850-10A), especially at the lower loading levels, owing to the extent of initial plastic deformation. On the contrary, at the lower loading levels the B60-10N variant does relax in a very limited extent, thus showing that for this material, at these loading levels, the damage mechanisms are little or no active leading to longer fatigue lives. Consequently, the Z850-10A treatment exerts a lower
improvement with respect to the B60-10N variants, especially at the lower loading levels. 5. Conclusion The principal results of this experimental work can be summarised as follows: 1. Initial stress profiles are differentiated for the two peening conditions in terms of surface residual stress, subsurface stress peak and depth of residual stress profile. 2. During the first part of cyclic loading both materials undergo to stress relaxation, which shows a strong dependence on the external applied load and on the initial surface microstructure. 3. Both material conditions reach early a stabilised value of surface residual stress, depending on the loading level. 4. A preliminary evaluation of the stabilised stress profile shows very little modification of the stress profile, except for a very thin surface layer. 5. Accounting for the residual stress distribution, the improvement in fatigue behaviour due to peening can be approximately estimated considering the residual stress field as a mean stress superposed on the stress field due to external loading. 6. The residual stress relaxation, even though limited to a very thin surface layer, seems to have a significant effect on fatigue behaviour especially at the higher loading levels. A better estimate of the fatigue curve can be obtained considering the stabilised residual stresses (SRS) instead of the initial residual stresses (IRS). 7. The fatigue behaviour of the two peening conditions cannot be completely explained considering the effect of residual stresses. In particular, an underestimation of the fatigue behaviour can be obtained at the lower loading levels. 8. The crack closure effect and the surface work hardening effects seem to have an important influence at these loading values. In particular, the extent of strain hardening can explain the observed differences between the different peening conditions. In the long life regime, where initiation events and the associated dislocation network are important, the proposed model should be improved by incorporating the microstructural modifications induced by shot peening. Accordingly, a dislocation-based modelling, thoroughly accounting for residual stress effects (e.g., [10]), would be a more appropriate route to follow. References [1] Benedetti M, Fontanari V, Hohn B, Oster P, Tobie T. Influence of shot peening on bending tooth fatigue limit of case hardened gears. Int J Fatigue 2002;24:1127–36.
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