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AA 6082-T6 aluminium alloys: Influence of material properties and tool geometry on weld strength
Materials and Design 87 (2015) 721–731
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Dissimilar friction stir lap welding of AA 5754-H22/AA 6082-T6 aluminium alloys: Influence of material properties and tool geometry on weld strength M.I. Costa a, D. Verdera c, C. Leitão a, D.M. Rodrigues a,b,⁎ a b c
CEMUC, Department of Mechanical Engineering, University of Coimbra, Rua Luis Reis Santos, 3030-788 Coimbra, Portugal ISISE, Department of Civil Engineering, University of Coimbra, Rua Luis Reis Santos, 3030-788 Coimbra, Portugal AIMEN, Relva 27A Torneiros, 36410 Porriño, Spain
a r t i c l e
i n f o
Article history: Received 25 May 2015 Received in revised form 27 July 2015 Accepted 16 August 2015 Available online 23 August 2015 Keywords: FSW Microstructure Mechanical behaviour Aluminium alloys
a b s t r a c t Dissimilar friction stir welding (FSW) of heat (AA 6082-T6) and non-heat (AA 5754-H22) treatable aluminium alloys, in lap joint configuration, was performed in this work. The base material plates were 1 mm thick. Welds were performed combining different plates positioning, relative to the tool shoulder, in order to assess the influence of base materials properties on welds strength. Three different tools were tested, one cylindrical and two conical, with different taper angles. Welds strength was characterized by performing transverse and tensile–shear tests. Strain data acquisition by Digital Image Correlation (DIC) was used to determine local weld properties. The results obtained enabled to conclude that the dissimilar welds strength is strongly dependent on the presence of the well-known hooking defect and that the hooking characteristics are strongly conditioned by base materials properties/positioning. By placing the AA 6082-T6 alloy, as top plate, in contact with the tool shoulder, superior weld properties are achieved independently of the tool geometry. It is also concluded that the use of unthreaded conical pin tools, with a low shoulder/pin diameter relation, is the most suitable solution for the production of welds with similar strengths for advancing and retreating sides. © 2015 Elsevier Ltd. All rights reserved.
1. Introduction Friction Stir Lap Welding (FSLW) is a Friction Stir Welding (FSW) related technique which is being studied as an alternative to riveting [1–3] and fusion welding technologies [3]. Due to its importance in the chemical, nuclear, aerospace, transportation, power generation and electronics industries, the lap joining of dissimilar alloys become one of the most important topics under analysis on FSLW research and development. The joining of alloys of titanium and aluminium (e.g. [4]), magnesium and aluminium (e.g. [5]), magnesium and steel (e.g. [6]), copper–nickel and steel (e.g. [7]), aluminium and steel (e.g. [8]), titanium and steel (e.g. [9]), aluminium and copper (e.g. [10]), dissimilar steels (e.g. [11]) and dissimilar aluminium alloys, was already addressed by several authors. In Fig. 1 are provided the references on dissimilar aluminium alloys welding, highlighting the scarcity of works on this subject and the limited number of base materials combinations tested. As shown in Fig. 1, Dubourg et al. [1], Ciliberto et al. [3], Cederqvist and Reynolds [12], and Song et al. [13] analysed the joining of AA 7075 and AA 2024 alloys. Despite the different tool geometries and ⁎ Corresponding author at: CEMUC, Department of Mechanical Engineering, University of Coimbra, Rua Luis Reis Santos, 3030-788 Coimbra, Portugal. E-mail address: [email protected] (D.M. Rodrigues).
http://dx.doi.org/10.1016/j.matdes.2015.08.066 0264-1275/© 2015 Elsevier Ltd. All rights reserved.
process parameters tested by the different authors (see Fig. 2), hooking and top sheet thinning defects were reported in all the works. In order to suppress the hooking defect, Cederqvist and Reynolds [12] and Dubourg et al. [1] performed double pass welding. The presence of kissing bond was also detected by Dubourg et al. [1] and Ciliberto et al. [3]. The authors argue that this defect, which is not currently reported in FSLW literature, has an important influence on the mechanical and electrochemical behaviours of the welds. Velotti et al. [14] also reported the presence of kissing bond in dissimilar AA 7075–AA 2198 alloys lap welds. In addition to the presence of the above described defects, the lap joints strength also depends on the base materials relative positioning. This issue was already addressed by some authors which analysed different material combinations and welding conditions, as described in Figs. 1 and 2. Soundararajan et al. [16] reported superior performance for the welds performed with the AA 6022 alloy on the top of the joint, in contact with the FSLW tool. These authors highlight the importance of the base materials thermal properties on welds quality. Lee et al. [17], in accordance with Soundararajan et al. [16], also reported superior mechanical strength for welds performed with the AA 6061 alloy placed on the top of AA 6061–AA 5052 alloys lap joints. However, these authors justify the mechanical performance of the welds based on the different thicknesses of the plates. Finally, Song et al. [13] also analysed the relative positioning of the base materials, in dissimilar lap joining of AA
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Fig. 1. Dissimilar aluminium alloys FSLW: base materials combinations [15].
Fig. 2. Dissimilar aluminium alloys FSLW: tool geometries and welding parameters [15].
2024 and AA 7075 aluminium alloys, concluding that the hook size was increased when the AA 7075 alloy was placed on the top of the joint. In current work dissimilar lap welding of 1 mm thick plates of AA 6082-T6 and AA5754-H22 aluminium alloys is analysed. The influence of base materials relative positioning and tool geometry on welds strength will be explained based on weld morphology and base materials plastic properties. 2. Materials and methods In this work dissimilar friction stir lap welding of 1 mm thick plates of AA 6082-T6 and AA 5754-H22 aluminium alloys was carried out in a
MTS I-Stir Pds equipment. From Fig. 3, where are compared tensile stress–strain curves for both base materials, it is possible to conclude that, at room temperature, the AA 6082-T6 alloy displays much higher yield and tensile strength than the AA 5754-H22 aluminium alloy. The welds were performed using three different tools, with a 12 mm diameter conical shoulder, and 1.3 mm long pins, but different pin geometries, as summarized in Table 1. In the text, the tools will be labelled according to the pin geometry (CL and CN for the cylindrical and conical, respectively) and top diameter, e.g., the cylindrical pin tool will be labelled CL6 tool. The welds were performed in position control, at the constant welding (υ) and rotational (ω) speeds of 350 mm/min and 600 rpm, respectively, tool tilt angel of 2° and plunge depth of 1.2 mm. In order to analyse the influence of the base materials properties on weld strength, the positioning of the alloys relatively to the tool shoulder
Table 1 Tools geometries and dimensions.
Fig. 3. Base materials tensile stress–strain curves.
Tool designation
Pin geometry
Pin top diameter [mm]
Pin bottom diameter [mm]
CL6
Cylindrical
6
6
CN6
Conical
6
3
CN8
Conical
8
4
Illustration
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Fig. 4. Scheme of the transverse tensile samples (a) and of advancing (b) and retreating (c) tensile–shear samples.
was alternated, i.e., welds were performed by placing alternately the AA 6082-T6 and the AA 5754-H22 aluminium alloys at the top of the lap joint. In this text the welds performed with AA 6082-T6 and AA 5754H22 aluminium alloys as top plates will be named as D65 and D56, respectively. Metallographic examinations were done in transverse, longitudinal and horizontal sections of the welds. The metallographic specimens were polished and etched with two different reagents (modified Poulton's and Hatch) and observed using the Zeiss Stemi 2000-C and Leica DM4000M LED microscopes. Scanning electron microscopy/ energy dispersive X-ray spectroscopy (SEM/EDS) was performed using a Philips XL30 SE. Hardness measurements were performed in a Struers Duramin equipment, applying a 200 g load, for 15 s. Welds strength was assessed by performing tensile–shear tests as well as uniaxial tensile tests of transverse weld samples (Fig. 4.a). Attending to the well-known strength asymmetry between the advancing (AS) and retreating (RS) sides of the lap welds, two different types of tensile–shear samples were tested, as schematized in Fig. 4.b) and c). At least three samples of each type were tested. All tests were performed in quasi-static loading conditions (5 mm/min), using a 5kN universal testing machine (Shimadzu Autograph AG-X). Local strain fields were acquired by Digital Image Correlation (DIC) using the GOM Aramis 5 M system. Before testing, the specimens were prepared by applying a random black speckle pattern, over the previously mat white painted surface, in order to enable strain data acquisition by DIC. All the procedures adopted for plotting local TMAZ stress–strain curves are explained in [18].
In Fig. 5.a and b are compared, respectively, the NYL and NML values obtained for all the D56 and D65 welds performed. From Fig. 5 it is possible to conclude that, for most of the welds, the
3. Results and discussion 3.1. Tensile–shear tests As described in the experimental procedure, the strength of the dissimilar joints was assessed by performing tensile–shear tests of advancing (Fig. 4.b) and retreating (Fig. 4.c) side samples. Two parameters were calculated in order to characterize the welds strength, the Normalized Yield Load (NYL) and the Normalized Maximum Load (NML) parameters, determined according to the equations: NYL ¼
NML ¼
F weld yield F 5754 yield
;
F weld max F 5754 max
ð1Þ
:
ð2Þ
weld In Eqs. (1) and (2), F weld yield and F max correspond to the yield and maximum loads registered in the tensile–shear tests of the welds, 5754 respectively, and F 5754 yield and F max correspond, respectively, to the yield and maximum loads registered for the weakest base material, the AA 5754-H22 aluminium alloy, in uniaxial tension of samples of the same width of the tensile–shear weld samples.
Fig. 5. NML and NYL versus weld speed for retreating and advancing side tensile–shear samples. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)
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Fig. 6. DIC strain distribution maps for advancing side tensile–shear samples of CN8 and CL6 welds. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)
strength of the retreating side (RS) was higher than that of the advancing side (AS). It is also possible to conclude that the difference between the strengths of the AS and RS sides was higher for the D56 than for the D65 welds. 3.1.1. Strength of the advancing side According to the results in Fig. 5, the strength of the advancing side of the D56 welds was the lowest among all the samples tested, being around 65% and 40% of the yield and maximum strength, respectively, of the weakest base material. Another interesting finding is that the strength of the advancing side of the D56 welds was independent of the tool geometry, displaying very low values for all the samples tested. On the other hand, for the welds performed using the D65 base materials combination, the strength of the advancing side varied among the welds performed with the different tool geometries, being in any case higher than that registered for the welds performed using the D56 materials configuration. For the D65 welds, the strength of the advancing side varied between minimum and maximum values of 85% and 100%, of the weakest base material yield strength, and 60% and 90%, of the weakest base material maximum strength, for the welds performed with the CL6 and CN8 tools, respectively. 3.1.2. Strength of the retreating side Regarding the strength of the retreating side, it is possible to conclude, from Fig. 5, that the differences in strength between the welds performed with the D56 and D65 material configurations were much less significant than that reported for the advancing
side. The retreating side yield and maximum strengths were always higher than 90% and 75% of the yield and maximum strengths, respectively, of the weakest base material. It is also possible to conclude that the maximum NYL and NML values, for the retreating side samples, were always higher for the welds performed with the CN8 tool. Actually, the best mechanical efficiency, in both retreating and advancing side strength, was always obtained for the welds performed with the CN8 tool adopting the D65 material combination. For these welds no differences between retreating and advancing side strengths were reported as highlighted in the graphs using a red circle. In Figs. 6 and 7 are compared the strain distribution maps, corresponding to the maximum load, for advancing and retreating side tensile–shear samples, respectively, of CN8 and CL6 welds. As shown by the sketches provided in both figures, the strain maps correspond always to the strain distribution in the top plate, i.e., to the AA 6082-T6 alloy, in the case of D65 welds, and to the AA 5754-H22 alloy, in the case of D56 welds. Fig. 6 shows that almost all the advancing side samples, except that of the D65 CN8 welds, failed almost without plastic deformation, which is in accordance with the very small NYL values reported in Fig. 5 for almost all the welds. For the D65 CN8 samples, which displayed NYL and NML values close to 100%, it is possible to see a larger plastic deformation area spreading from the necking region (red colour area). Analysing now the strain distribution maps in Fig. 7, corresponding to the retreating side samples of the same welds of Fig. 6, it is possible to conclude that these samples failed for plastic deformation values much larger than that registered for the advancing side samples.
Fig. 7. DIC strain distribution maps for retreating side tensile–shear samples of CN8 and CL6 welds. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)
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Actually, for the D56 welds it is possible to see that plastic deformation spread across the top plate before necking took place at the retreating side of the weld, which proves the excellent performance of these welds. For the D65 welds important plastic deformation was only registered in the proximity of the necking area, as it was registered for the D65 CN8 advancing side sample. In order to enable a better understanding of the D65 welds strain distribution maps, hardness and strain profiles for the D65 welds performed with the CL6 and CN8 tools, are compared in Fig. 8. The strain profiles in the graphs correspond to the strain distribution, at maximum load, in the top plate, along the longitudinal axis of the samples, and the hardness profiles correspond to the hardness measurements performed in the transverse cross-section of the welds, in the top plate. Analysing the figure it is possible to conclude that the hardness profiles are similar, for the CN8 and CL6 welds, displaying the traditional W shape characteristic of the AA 6xxx alloys friction stir welds [19]. The strain profiles
Fig. 8. Hardness profiles, for the AA 6082-T6 top plate, of the D65 welds performed with the CN8 (a) and CL6 (b) tools.
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corresponding to the retreating side samples are also similar, for the CN8 and CL6 welds, displaying very high strain values in the same location where minimum hardness values were registered, i.e., in the heat affected zone (HAZ). For the advancing side samples, the strain profiles for the CN8 and CL6 welds are no longer similar, being possible to observe higher plastic deformation values in the HAZ for the CN8 sample than for the CL6. This result indicates that plastic deformation took place in the lower plate, which corresponds to the lower strength AA 5754 alloy, before necking started in the AA 6082 alloy top plate, at the advancing side of the weld. Plastic deformation of the lower plate base material, together with the plastic deformation of the top plate HAZ, conducted to the higher strength values reported in Fig. 5 for the D65 CN8 welds. 3.2. Transverse tensile tests The tensile–shear tests revealed not only the superior performance of the D65 welds, performed with the CN8 tool, in terms of advancing and retreating side mechanical efficiency, but also important differences in strength between the welds performed using D56 and D65 material combinations. In order to determine any possible influence of local weld properties, on the dissimilar joints performance, transverse tensile tests were also performed (Fig. 4.a). These tests enabled to determine the local strength of the TMAZ material, for each weld, as well as to better assess the importance of base materials positioning on welds performance. In Fig. 9.a and c are shown the load–displacement curves obtained in the transverse tensile test of D56 and D65 samples, respectively, and in Fig. 9.b and d are shown the strain distribution maps corresponding to points 1 to 4 in the graphs of Fig. 9.a and c. The strain maps represent the plastic deformation in the lower plate of CN6 welds, i.e., the AA 6082-T6 alloy, in the case of the D56 samples, and the AA 5754-H22 alloy, in the case of the D65 samples. From the load–displacement curves in Fig. 9.a it is possible to conclude that all the transverse D56 samples failed for similar values of load and displacement. As is exemplified in Fig. 9.b, all the D56 welds failed in the top plate, the AA 5754-H22 alloy plate, at the advancing side, after very small plastic deformation took place in the weld (image 1). After that, the load decreased almost instantaneously, due to the load transfer to a single plate, the AA 6082-T6 alloy lower plate. Strain localization in the HAZ was almost immediate (image 2), and the final rupture occurred before the plastic deformation had spread across the lower plate. These results confirm the extreme fragility of the advancing side of the D56 welds already reported when analysing Fig. 5. Analysing now the load–displacement results in Fig. 9.c, corresponding to the transverse D65 welds, it is possible to conclude that the welds produced with the different tools displayed different mechanical behaviour. Actually, despite all the welds failed at the advancing side, in the top plate, like the D56 welds, the CN6 transverse samples failed for load–displacement values much lower than that registered for the CN8 samples. As shown by the strain maps in Fig. 9.d, for the D65 welds, top plate failure occurred after some plastic deformation took place in the weld. After top plate failure, the load dropped drastically and the plastic deformation proceeded in the lower plate, the AA 5754-H22 alloy, until its maximum strength was reached. For the welds produced with the CL6 and CN8 tools, plastic deformation in the weld, before failure at the advancing side, was even higher than that registered for the CN6 weld, as can be inferred from the load– displacement curves in Fig. 9.c. Using the strain data acquired by DIC, local stress–strain curves corresponding to the tensile behaviour of the TMAZ material, of each weld, were calculated being plotted in Fig. 10. In the same graphs are also plotted the stress–strain curves corresponding to the uniaxial tensile behaviour of the weakest base material, the AA 5754-H22 alloy. Analysing the stress–strain curves it is possible to
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Fig. 9. Load–displacement curves and strain distribution maps, at maximum load, from transverse tensile tests of D56 (a and b) and D65 (c and d) samples.
conclude that all the welds were in slight under-match relative to the weaker base material yield strength, which justifies the plastic localization in the welds, before the failure of the top plate took place, exemplified by the strain maps in Fig. 9.b and d. The results in Fig. 10 also show that all the TMAZs had similar local strength, independently of the base materials combination adopted or of the tool used to produce the welds. This way, the differences obtained in terms of global strength, between the different welds, have to be related to the presence of important defects at the advancing side. This assumption was confirmed through the metallographic analysis results provided in the next item. 3.3. Metallographic analysis The mechanical characterization work showed that welds with excellent mechanical performance can be produced by adopting the D65 base materials combination and using a conical tool with suitable dimensions/geometry, i.e., the CN8 tool. In order to understand the differences in mechanical performance between the welds tested, a systematic microstructural analysis was performed. In Figs. 11 and 12 are shown transverse and longitudinal crosssections of the D56 and D65 welds, respectively. In the transverse cross-sections it is possible to see a clear and almost straight interface between both base materials, suggesting that, independently of the base material combination adopted (D56 or D65), no important transport of material took place across the faying interface during the lap welding. However, the longitudinal cross-sections in the same figures,
which sample the shear layer around the pin, formed at the end of the weld, clearly shows that there is a flow of material ascending from the lower plate. The ascendant flow is easier to visualize in the longitudinal section of the welds since the tool performed several rotations, in the same location, before stopping, promoting an intense and localized dragging and mixing of both base materials. The amount of material being dragged into the shear layer, at each tool revolution, during the welding operation, is necessarily much smaller and for this reason no evidence of material mixing can be apperceived in the transverse cross-sections. The first conclusion that can be drawn, after comparing the images in Figs. 11 and 12, is that the amount of lower plate material dragged upwards depends on the pin geometry and on the relative positioning of the base materials. 3.3.1. Influence of pin geometry Comparing the longitudinal sections in Figs. 11.a and b, and 12.a and b, for example, it is possible to see that, independently of the base materials positioning, the amount of ascendant flow is much higher when welding with the CN6 tool than when welding with the CN8 tool. In the same way, analysing Figs. 11.c and 12.c, it is possible to conclude that the volume of upward material flow was also important for the CL6 tool. The micrographs in Figs. 11 and 12 only enable a comparative/ qualitative analysis of the material flow patterns, during welding, for the different tools. In order to have quantitative data on the amount of material dragged by the tools, the axial force exerted on the tool during welding was recorded: The higher the force exerted
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Fig. 10. Local stress–strain curves for the TMAZ material of D56 (a) and D65 (b) welds.
on the tools, the larger the amount of base materials being dragged by it. In Fig. 13 is shown the average axial force (Fz) as a function of the shoulder plunge depth (SPD) for each tool. The force Fz correspond to an average of the axial force values registered by the machine, at each instant, during the welding operation. The SPD was calculated, for each tool, considering the pin plunge depth of 1.2 mm (Fig. 13.a). The graph shows that, for a similar pin plunge depth, the CN6 pin geometry determines the highest shoulder penetration and the CN8 tool the lowest one (60% lower than the CN6 SPD). The SPD and the axial force, for the CL6 tool, are both higher and lower than that registered for the CN8 and CN6 tools, respectively. From the figure it is possible to conclude that the axial force increases linearly with increasing SPD and that the SPD depends on the pin geometry. No strong non-linearity in axial force evolution, which could be directly related with the varying tool geometry, can be apperceived from the graph, independently of the base materials combination adopted. The higher force values were registered for the welds for which an intense tool dragging action was noticed when analysing the Figs. 11 and 12. As it is well known from previous works on FSLW, it is the ascendant material flow that promotes the interface pull-up phenomenon
Fig. 11. Transverse and longitudinal cross-sections of the D56 welds.
usually called hooking defect, with important consequences on welds strength [12]. In current work, the highest strength values were reported for the CN8 tool which, according to Figs. 11, 12 and 13, was the one promoting the lowest amount of upward material flow. Naturally, by diminishing the pin plunge depth would be possible to reduce the upward material flow for the CL6 and CN6 tools, due to the reduction in SPD. However, in industrial production, using a deep pin penetration into the lower plate is preferable due to the difficulties in controlling the process for a very small tolerance in tool positioning and/or variations in plate's thickness.
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cross-sections, for the D65 configuration, an ascendant flow (darker material) reaching the top of the weld is only noticeable in the longitudinal cross-section of Fig. 12.c. These results are easily explained by the differences in plastic flow, at high temperatures, of both alloys. According to Leitão et al. [20], the AA 6082-T6 alloy, which displays intense flow softening at high temperatures, easily flows upwards. The upward flow is particularly intense at the advancing side of the weld, where an “open space” exists before the shear layer material is extruded against the non-deformed base material [21]. In order to better illustrate the previous assumption, horizontal cross-sections, sampling the ending of the welds, where the final hole left by the pin is located, were analysed. Pictures of D56 and D65 horizontal cross-sections, of CN6 welds, are shown in Fig. 14. The image of the D56 weld (Fig. 14.a) clearly shows a white line all along the advancing side of the weld. The white line, which corresponds to AA 6082 alloy base material, extends from the advancing side of the weld, to the front of the tool, where it is incorporated into the shear layer. Contrary to this, in the horizontal cross-section of the D65 weld, only a very thin discontinuity can be observed in the same location. This discontinuity, which corresponds to the hooking usually detected at the advancing side of the transverse samples, is not filled with lower plate base material, as for the D56 welds. Actually, no important upward flow of the lower plate base material is noticeable in the transverse cross-sections of any of the D65 welds (Fig. 12). According to Leitão et al. [20], at high temperatures, the AA 5XXX alloys, which displays perfect plastic behaviour, doesn't flow so easily upward as the AA 6082 alloy. This way, the amount of material stirred by the tool during welding, is higher when welding with the D56 base material configuration, than when welding with the D65 base material configuration. This can be confirmed by comparing the tool axial force results in Fig. 13, which enable to confirm that, independently of the tool geometry, the axial force is always higher when welding with D56 material configuration.
Fig. 12. Transverse and longitudinal cross-sections of the D65 welds.
3.3.2. Influence of base materials positioning By comparing the cross-sections in Figs. 11 and 12, it is now possible to analyse the influence of base materials positioning/ properties on welds morphology/strength. In fact, it is evident from the figures that the upward material flow is much more significant when the AA 6082-T6 alloy is the lower plate (D56 weld configuration) than when the AA 5754-H22 alloy is the lower plate (D65 weld configuration). Actually, meanwhile for the D56 material combination, the upward flow (white material) always reaches the top of the weld, as it can be seen in all transverse and longitudinal
3.3.3. Hooking morphology As shown in Figs. 11 and 14, the upward flow of the AA 6082 alloy, when adopting the D56 base materials combination, induces the formation of important hooking at the advancing side of all the welds. The upward transport of material is noticeable, in Fig. 11, for all the welds, independently of the tool geometry, and is the reason for the low mechanical strength of the D56 advancing side tensile–shear (Fig. 5) and transverse tensile (Fig. 9) samples. On the other hand, the strength of the advancing side of the D65 welds, for which no important upward material flow can be noticed in Fig. 12, is much higher than that of the D56 welds. Actually, the D65 CN8 welds, for which no important upward material transport can be noticed in any of the cross sections of Fig. 12.a, were the only ones with yield and maximum strength efficiencies close to 100%. In order to understand the mechanical response of the welds performed with the CN8 tool, the hooking at the AS of the D56 and D65 welds is compared in Figs. 15 and 16, respectively. Analysing the macroscopic images, it is possible to conclude that the non-bonded part of the hooking, displayed in the SEM images of Figs. 15.a and 16.a, is similar for both types of welds. The non-bonded part is followed by a dark line, clearly visible in the optical microscopy images in Figs. 15.b and 16.b, which according to the SEM EDS analysis performed, corresponds to remnants of the oxide layer from the plates interface, i.e., to the presence of the kissing bond defect [1,3]. Meanwhile for the D56 weld, the kissing bond is continuous, delimiting a well-defined hooking, that extends to the plates surface, for the D65 welds, the kissing bond shifts to the right, delimiting a small hooking embedded in the top plate lower thickness. This way it is possible to conclude that the presence of kissing bond alone has no influence on the weld strength. The mechanical efficiency of
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Fig. 13. Evolution of the tool axial force (Fz) with shoulder plunge depth (SPD).
the welds depends on the combined effect of kissing bond and hooking shape and size. 4. Conclusions The mechanical and microstructural characterization of the dissimilar aluminium welds enabled to conclude that – the hooking defect is the principal factor in determining the strength of the lap joints in both tensile–shear and transverse tensile loading; – dissimilar base materials positioning has strong influence on material flow during welding and, in this way, on hooking formation; – the tool shoulder plunge depth, which depends on the tool geometry and penetration depth, has a strong influence on hooking formation; – the hooking defect size and shape may be conditioned, in order to diminish/eliminate its deleterious effect, by suppressing the upward
flow of the lower plate material, through an appropriate choice of tool geometry and base materials relative positioning; – in AA6xxx − AA5xxx dissimilar welding, excellent quality welds may be obtained by placing the AA 6xxx alloy in the top of the lap joints and welding with conical tools with low shoulder/pin diameters ratio.
Acknowledgements This research work is sponsored by national funds from the Portuguese Foundation for Science and Technology (FCT), via the project PEst-C/EME/UI0285/2013, and Quadro de Referência Estratégica Nacional (QREN), under contract grant no. 2012/24804. The authors, C. Leitão and M.I. Costa, are supported by the Portuguese Foundation for Science and Technology through SFRH/BPD/93685/2013 and SFRH/ BD/104073/2014 fellowships, respectively. All supports are gratefully acknowledged.
Fig. 14. Transverse and Horizontal cross-sections of CN6 D56 (a) and D65 (b) welds.
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Fig. 15. Microstructural analysis of the hooking defect in a D56 CN8 weld: a) and b) SEM and OM images of the hooking, respectively, c) and d) SEM details of the kissing bond and e) EDS analysis.
Fig. 16. Microstructural analysis of the hooking defect in a D65 CN8 weld: a) and b) SEM and OM images of the hooking, respectively, c) and d) SEM details of the kissing bond and e) EDS analysis.
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M.I. Costa et al. / Materials and Design 87 (2015) 721–731 [9] Y. Gao, K. Nakata, K. Nagatsuka, F.C. Liu, J. Liao, Interface microstructural control by probe length adjustment in friction stir welding of titanium and steel lap joint, Mater. Des. 65 (2015) 17–23. [10] B. Kuang, Y. Shen, W. Chen, X. Yao, H. Xu, J. Gao, et al., The dissimilar friction stir lap welding of 1A99 Al to pure Cu using Zn as filler metal with “pinless” tool configuration, Mater. Des. 68 (2015) 54–62. [11] H. Fujii, Y.D. Chung, Y.F. Sun, H. Tanigawa, Interface shape and microstructure controlled dissimilar friction stir lap welded steels, Sci. Technol. Weld. Join. 18 (2013) 279–286. [12] L. Cederqvist, A.P. Reynolds, Factors affecting the properties of friction stir welded aluminum lap joints, Weld. J. 80 (2001) 281–287. [13] Y. Song, X. Yang, L. Cui, X. Hou, Z. Shen, Y. Xu, Defect features and mechanical properties of friction stir lap welded dissimilar AA2024–AA7075 aluminum alloy sheets, Mater. Des. 55 (2014) 9–18. [14] C. Velotti, A. Astarita, A. Squillace, S. Ciliberto, M.G. Villano, M. Giuliani, et al., On the critical technological issues of friction stir welding lap joints of dissimilar aluminum alloys, Surf. Interface Anal. 45 (2013) 1643–1648. [15] H. Izadi, J. Fallu, A. Abdel-Gwad, T. Liyanage, A.P. Gerlich, Analysis of tool geometry in dissimilar Al alloy friction stir welds using optical microscopy and serial sectioning, Sci. Technol. Weld. Join. 18 (2013) 307–313.
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[16] V. Soundararajan, E. Yarrapareddy, R. Kovacevic, Investigation of the friction stir lap welding of aluminum alloys AA 5182 and AA 6022, J. Mater. Eng. Perform. 16 (2007) 477–484. [17] C.-Y. Lee, W.-B. Lee, J.-W. Kim, D.-H. Choi, Y.-M. Yeon, S.-B. Jung, Lap joint properties of FSWed dissimilar formed 5052 Al and 6061 Al alloys with different thickness, J. Mater. Sci. 43 (2008) 3296–3304. [18] C. Leitão, I. Galvão, R.M. Leal, D.M. Rodrigues, Determination of local constitutive properties of aluminium friction stir welds using digital image correlation, Mater. Des. 33 (2012) 69–74. [19] A. Simar, Y. Bréchet, B. de Meester, A. Denquin, C. Gallais, T. Pardoen, Integrated modeling of friction stir welding of 6xxx series Al alloys: process, microstructure and properties, Prog. Mater. Sci. 57 (2012) 95–183. [20] C. Leitão, R. Louro, D.M. Rodrigues, Analysis of high temperature plastic behaviour and its relation with weldability in friction stir welding for aluminium alloys AA5083-H111 and AA6082-T6, Mater. Des. 37 (2012) 402–409. [21] Z.W. Chen, T. Pasang, Y. Qi, Shear flow and formation of Nugget zone during friction stir welding of aluminium alloy 5083-O, Mater. Sci. Eng. A 474 (2008) 312–316.
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Materials and Design journal homepage: www.elsevier.com/locate/jmad
Dissimilar friction stir lap welding of AA 5754-H22/AA 6082-T6 aluminium alloys: Influence of material properties and tool geometry on weld strength M.I. Costa a, D. Verdera c, C. Leitão a, D.M. Rodrigues a,b,⁎ a b c
CEMUC, Department of Mechanical Engineering, University of Coimbra, Rua Luis Reis Santos, 3030-788 Coimbra, Portugal ISISE, Department of Civil Engineering, University of Coimbra, Rua Luis Reis Santos, 3030-788 Coimbra, Portugal AIMEN, Relva 27A Torneiros, 36410 Porriño, Spain
a r t i c l e
i n f o
Article history: Received 25 May 2015 Received in revised form 27 July 2015 Accepted 16 August 2015 Available online 23 August 2015 Keywords: FSW Microstructure Mechanical behaviour Aluminium alloys
a b s t r a c t Dissimilar friction stir welding (FSW) of heat (AA 6082-T6) and non-heat (AA 5754-H22) treatable aluminium alloys, in lap joint configuration, was performed in this work. The base material plates were 1 mm thick. Welds were performed combining different plates positioning, relative to the tool shoulder, in order to assess the influence of base materials properties on welds strength. Three different tools were tested, one cylindrical and two conical, with different taper angles. Welds strength was characterized by performing transverse and tensile–shear tests. Strain data acquisition by Digital Image Correlation (DIC) was used to determine local weld properties. The results obtained enabled to conclude that the dissimilar welds strength is strongly dependent on the presence of the well-known hooking defect and that the hooking characteristics are strongly conditioned by base materials properties/positioning. By placing the AA 6082-T6 alloy, as top plate, in contact with the tool shoulder, superior weld properties are achieved independently of the tool geometry. It is also concluded that the use of unthreaded conical pin tools, with a low shoulder/pin diameter relation, is the most suitable solution for the production of welds with similar strengths for advancing and retreating sides. © 2015 Elsevier Ltd. All rights reserved.
1. Introduction Friction Stir Lap Welding (FSLW) is a Friction Stir Welding (FSW) related technique which is being studied as an alternative to riveting [1–3] and fusion welding technologies [3]. Due to its importance in the chemical, nuclear, aerospace, transportation, power generation and electronics industries, the lap joining of dissimilar alloys become one of the most important topics under analysis on FSLW research and development. The joining of alloys of titanium and aluminium (e.g. [4]), magnesium and aluminium (e.g. [5]), magnesium and steel (e.g. [6]), copper–nickel and steel (e.g. [7]), aluminium and steel (e.g. [8]), titanium and steel (e.g. [9]), aluminium and copper (e.g. [10]), dissimilar steels (e.g. [11]) and dissimilar aluminium alloys, was already addressed by several authors. In Fig. 1 are provided the references on dissimilar aluminium alloys welding, highlighting the scarcity of works on this subject and the limited number of base materials combinations tested. As shown in Fig. 1, Dubourg et al. [1], Ciliberto et al. [3], Cederqvist and Reynolds [12], and Song et al. [13] analysed the joining of AA 7075 and AA 2024 alloys. Despite the different tool geometries and ⁎ Corresponding author at: CEMUC, Department of Mechanical Engineering, University of Coimbra, Rua Luis Reis Santos, 3030-788 Coimbra, Portugal. E-mail address: [email protected] (D.M. Rodrigues).
http://dx.doi.org/10.1016/j.matdes.2015.08.066 0264-1275/© 2015 Elsevier Ltd. All rights reserved.
process parameters tested by the different authors (see Fig. 2), hooking and top sheet thinning defects were reported in all the works. In order to suppress the hooking defect, Cederqvist and Reynolds [12] and Dubourg et al. [1] performed double pass welding. The presence of kissing bond was also detected by Dubourg et al. [1] and Ciliberto et al. [3]. The authors argue that this defect, which is not currently reported in FSLW literature, has an important influence on the mechanical and electrochemical behaviours of the welds. Velotti et al. [14] also reported the presence of kissing bond in dissimilar AA 7075–AA 2198 alloys lap welds. In addition to the presence of the above described defects, the lap joints strength also depends on the base materials relative positioning. This issue was already addressed by some authors which analysed different material combinations and welding conditions, as described in Figs. 1 and 2. Soundararajan et al. [16] reported superior performance for the welds performed with the AA 6022 alloy on the top of the joint, in contact with the FSLW tool. These authors highlight the importance of the base materials thermal properties on welds quality. Lee et al. [17], in accordance with Soundararajan et al. [16], also reported superior mechanical strength for welds performed with the AA 6061 alloy placed on the top of AA 6061–AA 5052 alloys lap joints. However, these authors justify the mechanical performance of the welds based on the different thicknesses of the plates. Finally, Song et al. [13] also analysed the relative positioning of the base materials, in dissimilar lap joining of AA
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Fig. 1. Dissimilar aluminium alloys FSLW: base materials combinations [15].
Fig. 2. Dissimilar aluminium alloys FSLW: tool geometries and welding parameters [15].
2024 and AA 7075 aluminium alloys, concluding that the hook size was increased when the AA 7075 alloy was placed on the top of the joint. In current work dissimilar lap welding of 1 mm thick plates of AA 6082-T6 and AA5754-H22 aluminium alloys is analysed. The influence of base materials relative positioning and tool geometry on welds strength will be explained based on weld morphology and base materials plastic properties. 2. Materials and methods In this work dissimilar friction stir lap welding of 1 mm thick plates of AA 6082-T6 and AA 5754-H22 aluminium alloys was carried out in a
MTS I-Stir Pds equipment. From Fig. 3, where are compared tensile stress–strain curves for both base materials, it is possible to conclude that, at room temperature, the AA 6082-T6 alloy displays much higher yield and tensile strength than the AA 5754-H22 aluminium alloy. The welds were performed using three different tools, with a 12 mm diameter conical shoulder, and 1.3 mm long pins, but different pin geometries, as summarized in Table 1. In the text, the tools will be labelled according to the pin geometry (CL and CN for the cylindrical and conical, respectively) and top diameter, e.g., the cylindrical pin tool will be labelled CL6 tool. The welds were performed in position control, at the constant welding (υ) and rotational (ω) speeds of 350 mm/min and 600 rpm, respectively, tool tilt angel of 2° and plunge depth of 1.2 mm. In order to analyse the influence of the base materials properties on weld strength, the positioning of the alloys relatively to the tool shoulder
Table 1 Tools geometries and dimensions.
Fig. 3. Base materials tensile stress–strain curves.
Tool designation
Pin geometry
Pin top diameter [mm]
Pin bottom diameter [mm]
CL6
Cylindrical
6
6
CN6
Conical
6
3
CN8
Conical
8
4
Illustration
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Fig. 4. Scheme of the transverse tensile samples (a) and of advancing (b) and retreating (c) tensile–shear samples.
was alternated, i.e., welds were performed by placing alternately the AA 6082-T6 and the AA 5754-H22 aluminium alloys at the top of the lap joint. In this text the welds performed with AA 6082-T6 and AA 5754H22 aluminium alloys as top plates will be named as D65 and D56, respectively. Metallographic examinations were done in transverse, longitudinal and horizontal sections of the welds. The metallographic specimens were polished and etched with two different reagents (modified Poulton's and Hatch) and observed using the Zeiss Stemi 2000-C and Leica DM4000M LED microscopes. Scanning electron microscopy/ energy dispersive X-ray spectroscopy (SEM/EDS) was performed using a Philips XL30 SE. Hardness measurements were performed in a Struers Duramin equipment, applying a 200 g load, for 15 s. Welds strength was assessed by performing tensile–shear tests as well as uniaxial tensile tests of transverse weld samples (Fig. 4.a). Attending to the well-known strength asymmetry between the advancing (AS) and retreating (RS) sides of the lap welds, two different types of tensile–shear samples were tested, as schematized in Fig. 4.b) and c). At least three samples of each type were tested. All tests were performed in quasi-static loading conditions (5 mm/min), using a 5kN universal testing machine (Shimadzu Autograph AG-X). Local strain fields were acquired by Digital Image Correlation (DIC) using the GOM Aramis 5 M system. Before testing, the specimens were prepared by applying a random black speckle pattern, over the previously mat white painted surface, in order to enable strain data acquisition by DIC. All the procedures adopted for plotting local TMAZ stress–strain curves are explained in [18].
In Fig. 5.a and b are compared, respectively, the NYL and NML values obtained for all the D56 and D65 welds performed. From Fig. 5 it is possible to conclude that, for most of the welds, the
3. Results and discussion 3.1. Tensile–shear tests As described in the experimental procedure, the strength of the dissimilar joints was assessed by performing tensile–shear tests of advancing (Fig. 4.b) and retreating (Fig. 4.c) side samples. Two parameters were calculated in order to characterize the welds strength, the Normalized Yield Load (NYL) and the Normalized Maximum Load (NML) parameters, determined according to the equations: NYL ¼
NML ¼
F weld yield F 5754 yield
;
F weld max F 5754 max
ð1Þ
:
ð2Þ
weld In Eqs. (1) and (2), F weld yield and F max correspond to the yield and maximum loads registered in the tensile–shear tests of the welds, 5754 respectively, and F 5754 yield and F max correspond, respectively, to the yield and maximum loads registered for the weakest base material, the AA 5754-H22 aluminium alloy, in uniaxial tension of samples of the same width of the tensile–shear weld samples.
Fig. 5. NML and NYL versus weld speed for retreating and advancing side tensile–shear samples. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)
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Fig. 6. DIC strain distribution maps for advancing side tensile–shear samples of CN8 and CL6 welds. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)
strength of the retreating side (RS) was higher than that of the advancing side (AS). It is also possible to conclude that the difference between the strengths of the AS and RS sides was higher for the D56 than for the D65 welds. 3.1.1. Strength of the advancing side According to the results in Fig. 5, the strength of the advancing side of the D56 welds was the lowest among all the samples tested, being around 65% and 40% of the yield and maximum strength, respectively, of the weakest base material. Another interesting finding is that the strength of the advancing side of the D56 welds was independent of the tool geometry, displaying very low values for all the samples tested. On the other hand, for the welds performed using the D65 base materials combination, the strength of the advancing side varied among the welds performed with the different tool geometries, being in any case higher than that registered for the welds performed using the D56 materials configuration. For the D65 welds, the strength of the advancing side varied between minimum and maximum values of 85% and 100%, of the weakest base material yield strength, and 60% and 90%, of the weakest base material maximum strength, for the welds performed with the CL6 and CN8 tools, respectively. 3.1.2. Strength of the retreating side Regarding the strength of the retreating side, it is possible to conclude, from Fig. 5, that the differences in strength between the welds performed with the D56 and D65 material configurations were much less significant than that reported for the advancing
side. The retreating side yield and maximum strengths were always higher than 90% and 75% of the yield and maximum strengths, respectively, of the weakest base material. It is also possible to conclude that the maximum NYL and NML values, for the retreating side samples, were always higher for the welds performed with the CN8 tool. Actually, the best mechanical efficiency, in both retreating and advancing side strength, was always obtained for the welds performed with the CN8 tool adopting the D65 material combination. For these welds no differences between retreating and advancing side strengths were reported as highlighted in the graphs using a red circle. In Figs. 6 and 7 are compared the strain distribution maps, corresponding to the maximum load, for advancing and retreating side tensile–shear samples, respectively, of CN8 and CL6 welds. As shown by the sketches provided in both figures, the strain maps correspond always to the strain distribution in the top plate, i.e., to the AA 6082-T6 alloy, in the case of D65 welds, and to the AA 5754-H22 alloy, in the case of D56 welds. Fig. 6 shows that almost all the advancing side samples, except that of the D65 CN8 welds, failed almost without plastic deformation, which is in accordance with the very small NYL values reported in Fig. 5 for almost all the welds. For the D65 CN8 samples, which displayed NYL and NML values close to 100%, it is possible to see a larger plastic deformation area spreading from the necking region (red colour area). Analysing now the strain distribution maps in Fig. 7, corresponding to the retreating side samples of the same welds of Fig. 6, it is possible to conclude that these samples failed for plastic deformation values much larger than that registered for the advancing side samples.
Fig. 7. DIC strain distribution maps for retreating side tensile–shear samples of CN8 and CL6 welds. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)
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Actually, for the D56 welds it is possible to see that plastic deformation spread across the top plate before necking took place at the retreating side of the weld, which proves the excellent performance of these welds. For the D65 welds important plastic deformation was only registered in the proximity of the necking area, as it was registered for the D65 CN8 advancing side sample. In order to enable a better understanding of the D65 welds strain distribution maps, hardness and strain profiles for the D65 welds performed with the CL6 and CN8 tools, are compared in Fig. 8. The strain profiles in the graphs correspond to the strain distribution, at maximum load, in the top plate, along the longitudinal axis of the samples, and the hardness profiles correspond to the hardness measurements performed in the transverse cross-section of the welds, in the top plate. Analysing the figure it is possible to conclude that the hardness profiles are similar, for the CN8 and CL6 welds, displaying the traditional W shape characteristic of the AA 6xxx alloys friction stir welds [19]. The strain profiles
Fig. 8. Hardness profiles, for the AA 6082-T6 top plate, of the D65 welds performed with the CN8 (a) and CL6 (b) tools.
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corresponding to the retreating side samples are also similar, for the CN8 and CL6 welds, displaying very high strain values in the same location where minimum hardness values were registered, i.e., in the heat affected zone (HAZ). For the advancing side samples, the strain profiles for the CN8 and CL6 welds are no longer similar, being possible to observe higher plastic deformation values in the HAZ for the CN8 sample than for the CL6. This result indicates that plastic deformation took place in the lower plate, which corresponds to the lower strength AA 5754 alloy, before necking started in the AA 6082 alloy top plate, at the advancing side of the weld. Plastic deformation of the lower plate base material, together with the plastic deformation of the top plate HAZ, conducted to the higher strength values reported in Fig. 5 for the D65 CN8 welds. 3.2. Transverse tensile tests The tensile–shear tests revealed not only the superior performance of the D65 welds, performed with the CN8 tool, in terms of advancing and retreating side mechanical efficiency, but also important differences in strength between the welds performed using D56 and D65 material combinations. In order to determine any possible influence of local weld properties, on the dissimilar joints performance, transverse tensile tests were also performed (Fig. 4.a). These tests enabled to determine the local strength of the TMAZ material, for each weld, as well as to better assess the importance of base materials positioning on welds performance. In Fig. 9.a and c are shown the load–displacement curves obtained in the transverse tensile test of D56 and D65 samples, respectively, and in Fig. 9.b and d are shown the strain distribution maps corresponding to points 1 to 4 in the graphs of Fig. 9.a and c. The strain maps represent the plastic deformation in the lower plate of CN6 welds, i.e., the AA 6082-T6 alloy, in the case of the D56 samples, and the AA 5754-H22 alloy, in the case of the D65 samples. From the load–displacement curves in Fig. 9.a it is possible to conclude that all the transverse D56 samples failed for similar values of load and displacement. As is exemplified in Fig. 9.b, all the D56 welds failed in the top plate, the AA 5754-H22 alloy plate, at the advancing side, after very small plastic deformation took place in the weld (image 1). After that, the load decreased almost instantaneously, due to the load transfer to a single plate, the AA 6082-T6 alloy lower plate. Strain localization in the HAZ was almost immediate (image 2), and the final rupture occurred before the plastic deformation had spread across the lower plate. These results confirm the extreme fragility of the advancing side of the D56 welds already reported when analysing Fig. 5. Analysing now the load–displacement results in Fig. 9.c, corresponding to the transverse D65 welds, it is possible to conclude that the welds produced with the different tools displayed different mechanical behaviour. Actually, despite all the welds failed at the advancing side, in the top plate, like the D56 welds, the CN6 transverse samples failed for load–displacement values much lower than that registered for the CN8 samples. As shown by the strain maps in Fig. 9.d, for the D65 welds, top plate failure occurred after some plastic deformation took place in the weld. After top plate failure, the load dropped drastically and the plastic deformation proceeded in the lower plate, the AA 5754-H22 alloy, until its maximum strength was reached. For the welds produced with the CL6 and CN8 tools, plastic deformation in the weld, before failure at the advancing side, was even higher than that registered for the CN6 weld, as can be inferred from the load– displacement curves in Fig. 9.c. Using the strain data acquired by DIC, local stress–strain curves corresponding to the tensile behaviour of the TMAZ material, of each weld, were calculated being plotted in Fig. 10. In the same graphs are also plotted the stress–strain curves corresponding to the uniaxial tensile behaviour of the weakest base material, the AA 5754-H22 alloy. Analysing the stress–strain curves it is possible to
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Fig. 9. Load–displacement curves and strain distribution maps, at maximum load, from transverse tensile tests of D56 (a and b) and D65 (c and d) samples.
conclude that all the welds were in slight under-match relative to the weaker base material yield strength, which justifies the plastic localization in the welds, before the failure of the top plate took place, exemplified by the strain maps in Fig. 9.b and d. The results in Fig. 10 also show that all the TMAZs had similar local strength, independently of the base materials combination adopted or of the tool used to produce the welds. This way, the differences obtained in terms of global strength, between the different welds, have to be related to the presence of important defects at the advancing side. This assumption was confirmed through the metallographic analysis results provided in the next item. 3.3. Metallographic analysis The mechanical characterization work showed that welds with excellent mechanical performance can be produced by adopting the D65 base materials combination and using a conical tool with suitable dimensions/geometry, i.e., the CN8 tool. In order to understand the differences in mechanical performance between the welds tested, a systematic microstructural analysis was performed. In Figs. 11 and 12 are shown transverse and longitudinal crosssections of the D56 and D65 welds, respectively. In the transverse cross-sections it is possible to see a clear and almost straight interface between both base materials, suggesting that, independently of the base material combination adopted (D56 or D65), no important transport of material took place across the faying interface during the lap welding. However, the longitudinal cross-sections in the same figures,
which sample the shear layer around the pin, formed at the end of the weld, clearly shows that there is a flow of material ascending from the lower plate. The ascendant flow is easier to visualize in the longitudinal section of the welds since the tool performed several rotations, in the same location, before stopping, promoting an intense and localized dragging and mixing of both base materials. The amount of material being dragged into the shear layer, at each tool revolution, during the welding operation, is necessarily much smaller and for this reason no evidence of material mixing can be apperceived in the transverse cross-sections. The first conclusion that can be drawn, after comparing the images in Figs. 11 and 12, is that the amount of lower plate material dragged upwards depends on the pin geometry and on the relative positioning of the base materials. 3.3.1. Influence of pin geometry Comparing the longitudinal sections in Figs. 11.a and b, and 12.a and b, for example, it is possible to see that, independently of the base materials positioning, the amount of ascendant flow is much higher when welding with the CN6 tool than when welding with the CN8 tool. In the same way, analysing Figs. 11.c and 12.c, it is possible to conclude that the volume of upward material flow was also important for the CL6 tool. The micrographs in Figs. 11 and 12 only enable a comparative/ qualitative analysis of the material flow patterns, during welding, for the different tools. In order to have quantitative data on the amount of material dragged by the tools, the axial force exerted on the tool during welding was recorded: The higher the force exerted
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Fig. 10. Local stress–strain curves for the TMAZ material of D56 (a) and D65 (b) welds.
on the tools, the larger the amount of base materials being dragged by it. In Fig. 13 is shown the average axial force (Fz) as a function of the shoulder plunge depth (SPD) for each tool. The force Fz correspond to an average of the axial force values registered by the machine, at each instant, during the welding operation. The SPD was calculated, for each tool, considering the pin plunge depth of 1.2 mm (Fig. 13.a). The graph shows that, for a similar pin plunge depth, the CN6 pin geometry determines the highest shoulder penetration and the CN8 tool the lowest one (60% lower than the CN6 SPD). The SPD and the axial force, for the CL6 tool, are both higher and lower than that registered for the CN8 and CN6 tools, respectively. From the figure it is possible to conclude that the axial force increases linearly with increasing SPD and that the SPD depends on the pin geometry. No strong non-linearity in axial force evolution, which could be directly related with the varying tool geometry, can be apperceived from the graph, independently of the base materials combination adopted. The higher force values were registered for the welds for which an intense tool dragging action was noticed when analysing the Figs. 11 and 12. As it is well known from previous works on FSLW, it is the ascendant material flow that promotes the interface pull-up phenomenon
Fig. 11. Transverse and longitudinal cross-sections of the D56 welds.
usually called hooking defect, with important consequences on welds strength [12]. In current work, the highest strength values were reported for the CN8 tool which, according to Figs. 11, 12 and 13, was the one promoting the lowest amount of upward material flow. Naturally, by diminishing the pin plunge depth would be possible to reduce the upward material flow for the CL6 and CN6 tools, due to the reduction in SPD. However, in industrial production, using a deep pin penetration into the lower plate is preferable due to the difficulties in controlling the process for a very small tolerance in tool positioning and/or variations in plate's thickness.
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cross-sections, for the D65 configuration, an ascendant flow (darker material) reaching the top of the weld is only noticeable in the longitudinal cross-section of Fig. 12.c. These results are easily explained by the differences in plastic flow, at high temperatures, of both alloys. According to Leitão et al. [20], the AA 6082-T6 alloy, which displays intense flow softening at high temperatures, easily flows upwards. The upward flow is particularly intense at the advancing side of the weld, where an “open space” exists before the shear layer material is extruded against the non-deformed base material [21]. In order to better illustrate the previous assumption, horizontal cross-sections, sampling the ending of the welds, where the final hole left by the pin is located, were analysed. Pictures of D56 and D65 horizontal cross-sections, of CN6 welds, are shown in Fig. 14. The image of the D56 weld (Fig. 14.a) clearly shows a white line all along the advancing side of the weld. The white line, which corresponds to AA 6082 alloy base material, extends from the advancing side of the weld, to the front of the tool, where it is incorporated into the shear layer. Contrary to this, in the horizontal cross-section of the D65 weld, only a very thin discontinuity can be observed in the same location. This discontinuity, which corresponds to the hooking usually detected at the advancing side of the transverse samples, is not filled with lower plate base material, as for the D56 welds. Actually, no important upward flow of the lower plate base material is noticeable in the transverse cross-sections of any of the D65 welds (Fig. 12). According to Leitão et al. [20], at high temperatures, the AA 5XXX alloys, which displays perfect plastic behaviour, doesn't flow so easily upward as the AA 6082 alloy. This way, the amount of material stirred by the tool during welding, is higher when welding with the D56 base material configuration, than when welding with the D65 base material configuration. This can be confirmed by comparing the tool axial force results in Fig. 13, which enable to confirm that, independently of the tool geometry, the axial force is always higher when welding with D56 material configuration.
Fig. 12. Transverse and longitudinal cross-sections of the D65 welds.
3.3.2. Influence of base materials positioning By comparing the cross-sections in Figs. 11 and 12, it is now possible to analyse the influence of base materials positioning/ properties on welds morphology/strength. In fact, it is evident from the figures that the upward material flow is much more significant when the AA 6082-T6 alloy is the lower plate (D56 weld configuration) than when the AA 5754-H22 alloy is the lower plate (D65 weld configuration). Actually, meanwhile for the D56 material combination, the upward flow (white material) always reaches the top of the weld, as it can be seen in all transverse and longitudinal
3.3.3. Hooking morphology As shown in Figs. 11 and 14, the upward flow of the AA 6082 alloy, when adopting the D56 base materials combination, induces the formation of important hooking at the advancing side of all the welds. The upward transport of material is noticeable, in Fig. 11, for all the welds, independently of the tool geometry, and is the reason for the low mechanical strength of the D56 advancing side tensile–shear (Fig. 5) and transverse tensile (Fig. 9) samples. On the other hand, the strength of the advancing side of the D65 welds, for which no important upward material flow can be noticed in Fig. 12, is much higher than that of the D56 welds. Actually, the D65 CN8 welds, for which no important upward material transport can be noticed in any of the cross sections of Fig. 12.a, were the only ones with yield and maximum strength efficiencies close to 100%. In order to understand the mechanical response of the welds performed with the CN8 tool, the hooking at the AS of the D56 and D65 welds is compared in Figs. 15 and 16, respectively. Analysing the macroscopic images, it is possible to conclude that the non-bonded part of the hooking, displayed in the SEM images of Figs. 15.a and 16.a, is similar for both types of welds. The non-bonded part is followed by a dark line, clearly visible in the optical microscopy images in Figs. 15.b and 16.b, which according to the SEM EDS analysis performed, corresponds to remnants of the oxide layer from the plates interface, i.e., to the presence of the kissing bond defect [1,3]. Meanwhile for the D56 weld, the kissing bond is continuous, delimiting a well-defined hooking, that extends to the plates surface, for the D65 welds, the kissing bond shifts to the right, delimiting a small hooking embedded in the top plate lower thickness. This way it is possible to conclude that the presence of kissing bond alone has no influence on the weld strength. The mechanical efficiency of
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Fig. 13. Evolution of the tool axial force (Fz) with shoulder plunge depth (SPD).
the welds depends on the combined effect of kissing bond and hooking shape and size. 4. Conclusions The mechanical and microstructural characterization of the dissimilar aluminium welds enabled to conclude that – the hooking defect is the principal factor in determining the strength of the lap joints in both tensile–shear and transverse tensile loading; – dissimilar base materials positioning has strong influence on material flow during welding and, in this way, on hooking formation; – the tool shoulder plunge depth, which depends on the tool geometry and penetration depth, has a strong influence on hooking formation; – the hooking defect size and shape may be conditioned, in order to diminish/eliminate its deleterious effect, by suppressing the upward
flow of the lower plate material, through an appropriate choice of tool geometry and base materials relative positioning; – in AA6xxx − AA5xxx dissimilar welding, excellent quality welds may be obtained by placing the AA 6xxx alloy in the top of the lap joints and welding with conical tools with low shoulder/pin diameters ratio.
Acknowledgements This research work is sponsored by national funds from the Portuguese Foundation for Science and Technology (FCT), via the project PEst-C/EME/UI0285/2013, and Quadro de Referência Estratégica Nacional (QREN), under contract grant no. 2012/24804. The authors, C. Leitão and M.I. Costa, are supported by the Portuguese Foundation for Science and Technology through SFRH/BPD/93685/2013 and SFRH/ BD/104073/2014 fellowships, respectively. All supports are gratefully acknowledged.
Fig. 14. Transverse and Horizontal cross-sections of CN6 D56 (a) and D65 (b) welds.
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Fig. 15. Microstructural analysis of the hooking defect in a D56 CN8 weld: a) and b) SEM and OM images of the hooking, respectively, c) and d) SEM details of the kissing bond and e) EDS analysis.
Fig. 16. Microstructural analysis of the hooking defect in a D65 CN8 weld: a) and b) SEM and OM images of the hooking, respectively, c) and d) SEM details of the kissing bond and e) EDS analysis.
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