镁合金重叠的搅拌摩擦焊线性焊缝的疲劳寿命预测【中文7900字】
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镁合金重叠的搅拌摩擦焊线性焊缝的疲劳寿命预测【中文7900字】,镁合金,重叠,堆叠,搅拌,摩擦,磨擦,线性,焊缝,疲劳,寿命,预测,中文
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Downloaded From: / on 03/09/2016 Terms of Use: /about-asme/terms-of-useRuijie WangDepartment of Mechanical Engineering, Kunming University of Science and Technology,Kunming 650500, China; Department of Mechanical Engineering, University of Michigan-Dearborn,Dearborn, MI 48128Hong-Tae Kang1Department of Mechanical Engineering, University of Michigan-Dearborn,Dearborn, MI 48128Chonghua (Cindy) JiangAET Integration, Inc., Troy, MI 48084Fatigue Life Prediction for Overlap Friction Stir Linear Welds of Magnesium AlloysThis work was undertaken to analyze the stress/strain state at the critical sites in friction stir welded specimens and, further, to assess the fatigue strength of friction stir welded specimens with conventional fatigue life prediction approaches. Elastoplastic and elastic finite-element stress/strain analyses were carried out for friction-stir-linear-welded (FSLW) specimens made of magnesium alloys. The calculated stress/strain at the periph- ery of the weld nugget was used to evaluate the fatigue life with local life prediction approaches. First, elastoplastic finite-element models were built according to experimen- tal specimen profiles. Fatigue life prediction was conducted with Morrows modified MansonCoffin (MC) damage equation and the SmithWatsonTopper (SWT) damage equation, respectively, for different specimens under different loading cases. Life predic- tion results showed that both equations can to some extent give reasonable results, espe- cially within a low-cycle fatigue life regime, with the SWT damage equation giving more conservative results. As for high-cycle life, predicted results were much longer and scat- tered for both methods. Shell element elastic models were then used to calculate the structural stress at the periphery of the weld nuggets. The correlation between structural stress amplitude and experimental life showed the appropriateness of the structural stress fatigue evaluation for friction stir welds. The effect of the notches at the periphery of thefaying surface on life prediction was further discussed. DOI: 10.1115/1.4032469Keywords: friction stir welding, local stressstrain, elastoplastic finite-element analysis, cyclic stress strain, structural stress, fatigue life prediction1 IntroductionAs utilization of the friction stir welding (FSW) process expands, fatigue assessment of the FSW structures is drawing increasingly intensive attention. Under working applications, the weld is frequently subject to varying load conditions and is vulnerable to fatigue failure. However, standard codes and guide- lines such as Eurocode 9 concern only conventional welding joints and not FSW processes 1. Therefore, proper approaches and pro- cedures need to be identified for fatigue assessment of friction stir welded joints.Different kinds of samples have been considered, including friction stir welded butt joints, friction stir spot welded joints, fric- tion stir welded lap joints, etc. Research on FSW overlap joints is not so exhaustive due to lack of knowledge on the mechanical per- formance of this configuration 2. Many factors affect the fatigue strength of friction stir welded joints, such as different welding parameters and working conditions. Furthermore, during the FSW process, the trapped material originally present on the surfaces of the contacting sheets usually forms two cracklike curve profiles, commonly referred to as hook regions 3 or hooking defects 4. There are relatively fewer references on fatigue life assessment of friction stir welded components, and the hooks were seldom taken into consideration. The hooking defects and the thinning of the overlapping sheets caused by the hooking defects reduce the shear strength and fatigue lifetime 510. Thus, hooking defects need to be taken into consideration for life prediction.In this study, finite-element analyses (FEA) were conducted to assess the stress/strain-fatigue life relationship of friction stir welds made of magnesium alloys. In local elastoplastic stress/strain analysis, the hook region profile in the periphery of the weld nugget root was taken into consideration. For the1Corresponding author.Manuscript received December 22, 2014; final manuscript received December 21, 2015; published online March 9, 2016. Assoc. Editor: Blair E. Carlson.structural stress calculation, the weld was simply represented with shell elements in the FEA model. The adaptability of the two approaches to assess fatigue life of friction stir welds is discussed.2 Fatigue Tests of SpecimensSpecimens considered in this study were made of three different magnesium sheet alloys: AZ31 extruded, AM60 extruded, and AM30 cast. Figure 1 shows the typical shape of the specimens studied, where there is a linear weld seam across the overlap area of the two sheets, here referred to as FSLW. The cross sections perpendicular to the weld line of the welds are shown in Fig. 2. The sheet interface shapes at the weld seam in the advancing side and retreating side are different, where the sheets materials were stirred up by the rotating pin tool and penetrated into each otherFig. 1 FSW specimen configurationJournal of Manufacturing Science and Engineering JUNE 2016, Vol. 138 / 061013-1Copyright V C 2016 by ASMEDownloaded From: / on 03/09/2016 Terms of Use: /about-asme/terms-of-useFig. 2 Specimen weldline cross sections (a) AZ31 to AM60, (b)AZ31 to AZ31, and (c) AM30 to AM60during the welding process; thus, there were two curved shapes on each side of the interfaces of the two sheets, namely, hooks.The configuration of all three groups of specimens is similar, except for the sheet thickness differences. The length of two coupons of each specimen is 75 mm, the width is 30 mm, and the lap size is 30 mm. For the specimens in Fig. 1, the upper sheet is material AZ31, while the lower sheet is AM60. The specimens in the second group are made of two 2-mm thick AZ31 sheets; the weld nugget width of each specimen (the minimum size of theweldment in the length direction) is about 5.38 mm, from the hook ending of the advancing side to the hook top on the retreating side. The third group of specimens consists of joints between AM30 (2.4 mm, upper) and AM60 (2.4 mm, lower) sheets, and the weld width is about 6.27 mm.Load-controlled constant amplitude fatigue tests were carried out on an MTS 810 frame at room temperature, with a loading fre- quency of 50 Hz. Loading levels and experimental life results are shown in Table 1. Fatigue failure is defined as when the specimen is separated into two parts. Failed coupons revealed that fatigue cracks initiated at the weld interface, as shown in Fig. 3. The weld interface has notchlike geometry, where the thickness of the sheets is reduced. The specimens were gripped with shims by each end during fatigue testing so the effects of the thickness of the sheets on the welded joint could be minimized.The correlation between nominal stress and fatigue life is shown in Fig. 4. Here the nominal stress is the ratio of the loading to the sheet cross section area; no notch effect is under considera- tion. This figure shows that the specimens in the AM30-60 group have the shortest lives compared to those in the other two groups.3 Elastoplastic Finite-Element Analysis of FSLW Joint3.1 Finite-Element Model Development. The commercial software ABAQUS was used for finite-element elastoplastic stress/strain analysis. Due to the width and the thickness of the two coupons, the stress state of the weld nugget in the longitude direction can be assumed as plain strain state. Thus, shell element models were built for these specimens and plain strain element CPE4 is used in the finite-element model. The elements near the nugget root for the welded specimens of AZ31AM60 are shown in Fig. 5. In total, there are 973 nodes and 830 elements in this model. The periphery of the weld nugget was densely meshed, where the shortest length of element side is 0.053 mm.3.2 Material Cyclic StressStrain Curves. Figure 1 showed that each specimen consists of two sheets: the upper sheet material is AZ31, and the bottom sheet material is AM60. According to the cross section and material contact lines in Fig. 2, two material models are assigned, respectively, to two different zones in the FEA model, as shown in Fig. 5. Since materials had merged with each other here, there is no way to determine the material interfa- ces within this zone. So, for simplification of defining the material here, the material interfaces were assumed to extend in the hook profile direction. The material mechanical behavior change in the zone of the mixture of the two materials due to the welding pro- cess is not under consideration here; this would have some effects on the stress/strain calculation results. Since the total volume of the material trapped is constant before and after welding, the total area of meshes within the weldment zones of the two coupons here is kept unchanged.Within finite-element analysis, the material stressstrain rela- tionships, namely, Osgood Ramberg curves, need to be deter- mined in advance. The stable cyclic stressplastic strain curves of the three materials are shown in Fig. 6. For AZ31 and AM30, the elastic modulus E 45 GPa and Poissons ratio v 0.35; for AM60, the elastic modulus E 39 GPa and Poissons ratio v 0.35; and for AM30, the elastic modulus E 45 GPa and Pois- sons ratio v 0.3. The stressstrain parameters of AZ31 are from Ref. 11, AM30 from Ref. 12, and AM60 from test results.3.3 Finite-Element Analysis Results. In the FEA models, the left end, which is on the top sheet of the advancing side, was fully constrained with all six degrees-of-freedom, and the right end, on the bottom sheet of the retreating side, was constrained with five degrees-of-freedom, excluding the longitude direction. The tensile cyclic loads were applied along the unstrained061013-2 / Vol. 138, JUNE 2016 Transactions of the ASMEDownloaded From: / on 03/09/2016 Terms of Use: /about-asme/terms-of-useTable 1 Loading and fatigue test resultsMaterials Max load (N) Load ratio (R) Life (cycles)AM30AM60 3200 0.1 3344 1634 AM30AM60 2400 0.1 4737 19,199 AM30AM60 1600 0.1 31,273 AM30AM60 1360 0.1 53,529 AM30AM60 1290 0.1 221,457 138,252 AM30AM60 1280 0.1 116,234 165,269 130,231 104,175AZ31AZ31 5600 0.1 815 AZ31AZ31 4700 0.1 1582 AZ31AZ31 3800 0.1 1815 AZ31AZ31 2800 0.1 3610 AZ31AZ31 1900 0.1 7594 AZ31AZ31 900 0.1 183,110 111,677 108,220 AZ31AZ31 840 0.1 236,948 AZ31AZ31 790 0.1 332,365 318,191 198,308 AZ31AZ31 740 0.1 869,853 531,612 449,423 AZ31AM60 4000 0.1 3336 AZ31AM60 3000 0.1 8063 6669 AZ31AM60 2000 0.1 21,256 AZ31AM60 1500 0.1 76,206 95,402 82,950 AZ31AM60 1200 0.1 258,882 126,056 117,998 AZ31AM60 900 0.1 523,081 AZ31AM60 1000 0.1 400,155 AZ31AM60 580 0.1 568,175 407,115 378,346 AZ31AM60 4000 0.3 6498 6123 8336 AZ31AM60 2000 0.3 64,419 65,535 72,342 AZ31AM60 1500 0.3 202,857 AZ31AM60 1200 0.3 336,949 316,535 AZ31AM60 1100 0.3 312,923 AZ31AM60 1050 0.3 324,039 527,885 1,594,465 Fig. 3 Typical fatigue fracture modeFig. 4 Experimental results in nominal stress amplitudeFig. 5 Finite-element meshes at weld cross sectionFig. 6 Stable cyclic stress plastic strain curves for three materialsdirection on the bottom sheet. In order to get a whole stressstrain loop, at least three load steps were applied on the FEA models.Mises stress obtained from elastoplastic finite-element analysis showed that there was explicit stress concentration at the nuggetperiphery of the two faying surfaces, as shown in Fig. 7. The high- est stress occurred close to the notch root (but a little away from the weld nugget) of the upper coupon on the loading side. Due to the deformation and deflection caused by loading, the site of theJournal of Manufacturing Science and Engineering JUNE 2016, Vol. 138 / 061013-3Downloaded From: / on 03/09/2016 Terms of Use: /about-asme/terms-of-use0ff fffFig. 7 Mises stress (in MPa) distribution near the nugget rootTable 2 The damage equation parametersAZ31 AM30rf MPa 450.0 133.80e0Another frequently used damage model is the SWT equation. Firat 16 used the SWT equation to predict the fatigue test cycles and crack initiation locations. Pereira et al. 17 predicted fatigue lives with Morrows modified MC and the SWT damage equations.Here, two conventionally used local strain methods, Morrows modified MC damage equation and the SWT damage equation, combined with the elastoplastic FEA stress strain results, were chosen for fatigue life prediction. Morrows equation and SWT equations used are shown in Eqs. (1) and (2), respectively,De r 0 mf r b c2 EDe r 022N e0 2N (1)f 2b bcrmax 2 E 2N r0 e0 2N (2)where De=2 is the total strain amplitude, r0 is the fatigue strength coefficient, rm is the mean stress, E is the elastic modulus, b is the fatigue strength exponent, e0 is the fatigue ductility coefficient, c is the fatigue ductility exponent, and rmax is the maximum stress. It is known from Fig. 7 that the stress concentration always exists at the notch on the upper coupon. These fatigue parameters of the two materials concerned, which were determined from fatigue test results, are shown in Table 2.Fig. 8 The strain state near weld root. (a) Maximum Mises stress and three principal stress (MPa) history and (b) three principal strain history.largest stress should be slightly different from the maximum thin- ning site. The typical stresstime curves of the Mises stress and the three principal stresses at the stress concentration site are shown in Fig. 8; the three principal stresses all change with time. The three principal strains are also shown in Fig. 8; here, the strain in the out-of-plane direction is always zero. This indicates the typical plane strain stress state at the critical site.4 Fatigue Life Prediction of Friction Stir Linear Welding JointsIn modern material fatigue studies, the stressstrain curve and the strainlife curve are usually used in combination 13. The strain can usually be obtained from finite-element analysis, and then the strain-life curve is used to calculate the fatigue life. The one strain-life curve typically used is the famous MC equation. Jinu and Sathiya 14 used the MC equation to predict fatigue life with the strain from the FEM results. Heckel and Christ 15 studied the fatigue behavior of the thermomechanical fatigue of the TiAl intermetallic alloy TNB-V2 with the MC equation.4.1 Comparison Between Test Results and Prediction Results. Life prediction results with Morrows modified MC dam- age equation and with the SWT equation are shown in Fig. 9. Here, hollow symbols labeled with MC are the results with Mor- rows modified MC damage equation, while solid symbols labeled with SWT are the results with the SWT equation. The two parallel oblique lines are three-factor lines, implying that the scattering of the predicted results are between three times and one third of the experimental fatigue lives. Prediction results showed reasonable agreement with experimental results. Within the low cycle life regime ( 105 cycles) for the high cycle regime, pre- dicted life values are all longer than experimental fatigue life.Life prediction results with the SWT equation are also shown in Fig. 9; these predicted fatigue lives are much more conservative when compared with experimental fatigue lives than those obtained with Morrows modified MC. In the low cycle regime,Fig. 11 Mises stress distribution obtained from linear elastic FEAlinear elastic FEA. The structural stress used is calculated using Eq. (3) below 18prediction values are almost all shorter than or within three factors f i y 6mi yof the experimental data. For the high cycle regime, the predictedr i y r i y r i y x x (3)results show a tendency of being much longer than the experimen- tal results. where r is m ib t t2m y theThe two damage models used are all local stressstrainapproaches (also called notch approaches). Both of the equations can give reasonable results, especially within the low cycle regime, but the results are too widely scattered in the high cycle region. Furthermore, predicted results for each group in Fig. 9 do not mix with each other, showing the difference between the two methods. This may also mean that including the notch local geometry and local material mechanical behavior (not taking into consideration other factors) does have a salient effect on life pre- diction of friction stir welds. Local stressstrain calculation results depend greatly on the models local configuration of notch root geometry and size, as well as on the material properties, on which the stir welding process greatly depends. So even if very precise profiles of the weld hook can be modeled, life prediction results still would
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