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Search Results (3)

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Keywords = 6056 aluminum alloy

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19 pages, 9493 KiB  
Article
Numerical Simulation and Process Optimization of Laser Welding in 6056 Aluminum Alloy T-Joints
by Jin Peng, Shihua Xie, Tiejun Chen, Xingxing Wang, Xiaokai Yu, Luqiang Yang, Zenglei Ni, Zicheng Ling, Zhipeng Yuan, Jianjun Shi and Zhibin Yang
Crystals 2025, 15(1), 35; https://doi.org/10.3390/cryst15010035 - 30 Dec 2024
Viewed by 529
Abstract
This paper conducts a numerical simulation of the laser welding process for 6056 aluminum alloy stringers and skin T-joints using Simufact Welding. Initially, the accuracy of the finite element simulation is validated, followed by an exploration of the impact of bilateral asynchronous and [...] Read more.
This paper conducts a numerical simulation of the laser welding process for 6056 aluminum alloy stringers and skin T-joints using Simufact Welding. Initially, the accuracy of the finite element simulation is validated, followed by an exploration of the impact of bilateral asynchronous and bilateral synchronous laser welding on molten pool stability. Process parameters, including laser power, welding speed, fixture clamping force, and preheat temperature, are optimized through orthogonal testing. Furthermore, the influence of welding sequences on post-weld equivalent stress and deformation in three stringers’ T-joints is analyzed. The numerical simulation results indicate that the stability of the molten pool is superior in bilateral synchronous welding compared to asynchronous welding. Optimized process parameters were obtained through orthogonal testing, and subsequent experiments demonstrated that the welding sequence of welding both sides first, followed by the middle, produced lower post-weld equivalent stress and reduced overall joint deformation. Full article
(This article belongs to the Special Issue Design, Development and Processing of Aluminium Alloys and Their Composite Materials)
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Figure 1

Figure 1
<p>Dimensions of the stringer T-joints. (<b>a</b>) Stringer; (<b>b</b>) base plate; (<b>c</b>) joint.</p> Full article ">Figure 2
<p>Model of three stringer T-joints. (<b>a</b>) Base plate; (<b>b</b>) joint.</p> Full article ">Figure 3
<p>Mesh generation. (<b>a</b>) Stringer; (<b>b</b>) base plate; (<b>c</b>) base plate for the three stringer T-joints.</p> Full article ">Figure 4
<p>Schematic of the laser welding heat source model.</p> Full article ">Figure 5
<p>Comparison between experimental and simulated weld seam cross-section.</p> Full article ">Figure 6
<p>Comparison between experimental and simulated melt pool.</p> Full article ">Figure 7
<p>Melt pool at different time instances during double-sided synchronous laser welding. (<b>a</b>) t = 0.1078 s; (<b>b</b>) t = 0.5286 s; (<b>c</b>) t = 0.8451 s; (<b>d</b>) t = 1.266 s.</p> Full article ">Figure 8
<p>Double-sided synchronous laser welding. (<b>a</b>) Schematic of weld penetration, (<b>b</b>) penetration depth variation curve.</p> Full article ">Figure 8 Cont.
<p>Double-sided synchronous laser welding. (<b>a</b>) Schematic of weld penetration, (<b>b</b>) penetration depth variation curve.</p> Full article ">Figure 9
<p>Melt pool at different time instances during double-sided asynchronous laser welding. (<b>a</b>) t = 0.1078 s; (<b>b</b>) t = 0.4243 s; (<b>c</b>) t = 1.266 s; (<b>d</b>) t = 1.475 s.</p> Full article ">Figure 10
<p>Penetration depth variation curve of double-sided asynchronous laser welding.</p> Full article ">Figure 11
<p>Trends of various factors with equivalent stress.</p> Full article ">Figure 12
<p>Trends of various factors with deformation.</p> Full article ">Figure 13
<p>Melt pool morphology of orthogonal test schemes. Figures (<b>a</b>–<b>i</b>) correspond to case 1 to case 9.</p> Full article ">Figure 13 Cont.
<p>Melt pool morphology of orthogonal test schemes. Figures (<b>a</b>–<b>i</b>) correspond to case 1 to case 9.</p> Full article ">Figure 14
<p>Cross-sectional morphology of the molten pool with optimized parameter combinations.</p> Full article ">Figure 15
<p>Variation curves of equivalent force and deformation at welded joints. (<b>a</b>) Equivalent force (<b>b</b>) deformation.</p> Full article ">Figure 16
<p>Simulation model of three stringer T-joints.</p> Full article ">Figure 17
<p>Cloud plot of equivalent stress distribution.</p> Full article ">Figure 18
<p>Cloud plot of total deformation distribution.</p> Full article ">
16 pages, 8358 KiB  
Article
Influence of Pre-Aging on the Hardness and Formability of a Thread Rolled 6056 Aluminum Alloy after Conventional Extrusion and Artificial Aging
by Lisa Winter, Ralph Jörg Hellmig, Kristin Hockauf and Thomas Lampke
J. Manuf. Mater. Process. 2021, 5(4), 116; https://doi.org/10.3390/jmmp5040116 - 29 Oct 2021
Viewed by 2696
Abstract
For the production of aluminum screws, an effective thermomechanical treatment is necessary for enabling high strength combined with good formability. In this study, the influence of pre-aging as initial heat treatment prior to following processing steps was investigated for the precipitation hardenable 6056 [...] Read more.
For the production of aluminum screws, an effective thermomechanical treatment is necessary for enabling high strength combined with good formability. In this study, the influence of pre-aging as initial heat treatment prior to following processing steps was investigated for the precipitation hardenable 6056 aluminum alloy. The short-term low temperature pre-aged condition was compared to a naturally aged one representing storage time in manufacturing. As reference, a solution-annealed condition was used. After these initial heat treatments, conventional extrusion and artificial aging followed prior to final thread rolling. The distribution of strain introduced by these forming processes was numerically investigated using finite element simulation. The initial heat treatment had a significant influence on the mechanical properties achievable after the complete thermomechanical processing route. After extrusion and artificial aging, the highest hardness was achieved by the pre-aged condition. Despite its high initial hardness, this condition exhibited the best formability indicated by well-formed threads combined with the highest hardness achieved after thread rolling. Therefore, pre-aging seems to be an advantageous heat treatment for integration in the manufacturing process of screws due to its beneficial effect on the mechanical properties. Full article
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Figure 1

Figure 1
<p>Finite element simulation of the (<b>a</b>) strain distribution after extrusion (70% area reduction) and (<b>b</b>) using these results for early stage thread rolling process. The strain introduced by extrusion is highest close to the outer surface and decreases with increasing distance to the surface.</p> Full article ">Figure 2
<p>Finite element simulation of the strain distribution after thread rolling: (<b>a</b>) 3d view and (<b>b</b>) cut through area. The area at the thread root shows the highest introduced strain, whereas the strain introduced is lowest at the thread tip.</p> Full article ">Figure 3
<p>Vickers hardness as a function of the artificial aging time dependent on the processing route. The first and second hardness peak for the investigated maximum aging time is marked for each condition by a larger symbol. By pre-aging prior to linear extrusion, the highest hardness is achieved, when compared to the other processing routes.</p> Full article ">Figure 4
<p>Vickers Hardness across half of the wire section measured for the three different processing routes dependent on the artificial aging time prior to thread rolling as a function of the surface distance: initial heat treatment condition (<b>a</b>) solution-annealed, (<b>b</b>) naturally aged and (<b>c</b>) pre-aged. For all processing routes and artificial aging times, the hardness decreases with the increasing distance from the surface.</p> Full article ">Figure 5
<p>Vickers Hardness achieved for the three different processing routes dependent on the artificial aging time prior to and after thread rolling: (<b>a</b>) comparison between thread root and thread flank for each condition and (<b>b</b>) percentage increase in hardness by thread rolling. The highest hardness and percentage increase after thread rolling is achieved by processing route C.</p> Full article ">Figure 6
<p>Distribution of Martens hardness over the metallographically polished longitudinal section of the differently processed threaded parts after thread rolling. Schematic contours of thread teeth marked with dashed line in pink. Processing route and initial heat treatment: (<b>a</b>,<b>b</b>) processing route A (initially solution-annealed), (<b>c</b>,<b>d</b>) processing route B (initially naturally aged), (<b>e</b>,<b>f</b>) processing route C (initially pre-aged). Artificial aging time: (<b>a</b>–<b>c</b>) local hardness maximum and (<b>d</b>–<b>f</b>) global peak-hardness achieved. In general, the hardness of the thread root is significantly higher than of the thread flank. The highest overall hardness is achieved by processing route C (pre-aging as initial heat treatment).</p> Full article ">Figure 7
<p>Stereo micrographs of the different processed and thread rolled parts. Processing route and initial heat treatment: (<b>a</b>,<b>b</b>) processing route A (initially solution-annealed), (<b>c</b>,<b>d</b>) processing route B (initially naturally aged), (<b>e</b>,<b>f</b>) processing route C (initially pre-aged). Artificial aging time: (<b>a</b>–<b>c</b>) local hardness maximum and (<b>d</b>–<b>f</b>) global peak-hardness achieved. Independent of the processing route, for all threaded parts, the threads are not fully formed and the closing fold near the top of the teeth is still open.</p> Full article ">Figure 8
<p>Stereo micrographs of the different processed and thread rolled parts. Processing route and initial heat treatment: (<b>a</b>,<b>b</b>) processing route A (initially solution-annealed), (<b>c</b>,<b>d</b>) processing route B (initially naturally aged), (<b>e</b>,<b>f</b>) processing route C (initially pre-aged). Artificial aging time: (<b>a</b>–<b>c</b>) local hardness maximum and (<b>d</b>–<b>f</b>) global peak-hardness achieved. All thread rolled studs show scaling at the thread root, but this effect is significantly less pronounced for the studs processed by route C, when compared to the other processing routes.</p> Full article ">Figure 9
<p>Optical micrograph of a longitudinal section of a threaded part processed by route A. The thread teeth are not fully formed and as a result, the closing fold at the tip of the tooth is clearly visible and has a crack-like appearance. Further, the thread root exhibits distinct scaling.</p> Full article ">
13 pages, 3678 KiB  
Article
Influence of Pre-Aging on the Artificial Aging Behavior of a 6056 Aluminum Alloy after Conventional Extrusion
by Lisa Winter, Kristin Hockauf, Mario Scholze, Ralph Jörg Hellmig and Thomas Lampke
Metals 2021, 11(3), 385; https://doi.org/10.3390/met11030385 - 26 Feb 2021
Cited by 9 | Viewed by 3043
Abstract
In the present study, the influence of the initial heat-treatment conditions on the artificial aging behavior after conventional linear extrusion at room temperature was investigated for the precipitation hardening of a 6056 aluminum alloy. A solution-annealed condition was systematically compared to naturally-aged and [...] Read more.
In the present study, the influence of the initial heat-treatment conditions on the artificial aging behavior after conventional linear extrusion at room temperature was investigated for the precipitation hardening of a 6056 aluminum alloy. A solution-annealed condition was systematically compared to naturally-aged and pre-aged conditions. Differential scanning calorimetry was used for analyzing the precipitation sequence and its dependence on the initial heat treatment. The natural aging behavior prior to extrusion and the artificial aging behavior after extrusion were determined by microhardness measurements as a function of the aging time. Furthermore, the microstructure, dependent on the induced strain, was investigated using optical microscopy and transmission electron microscopy. As a result of pre-aging, following a solid-solution treatment, the formation of stable room-temperature clusters was suppressed and natural aging was inhibited. The artificial aging response after extrusion was significantly enhanced by pre-aging, and the achieved hardness and strength were significantly higher when compared with the equally processed solution-annealed or naturally-aged conditions. Full article
(This article belongs to the Special Issue Heat Treatment and Mechanical Properties of Metals and Alloys)
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Figure 1

Figure 1
<p>Differential scanning calorimetry (DSC) thermograms for the 6056 aluminum alloy in solution-annealed, naturally-aged and pre-aged conditions obtained for the following three heating rates: (<b>a</b>) 5 K min<sup>−1</sup>, (<b>b</b>) 10 K min<sup>−1</sup> and (<b>c</b>) 20 K min<sup>−1</sup>. With an increased heating rate, the peaks shift towards higher temperatures. After pre-aging at 80 °C, the formation of Mg-Si co-clusters (peak I) is not observed in the thermograms. The naturally-aged condition shows a pronounced endothermic peak, which corresponds to the dissolution of the Mg-Si co-clusters.</p> Full article ">Figure 2
<p>Microhardness of the 6056 aluminum alloy in solution-annealed and pre-aged conditions as a function of the natural aging time. Depicted are the mean, minimum and maximum values of the microhardness. The hardness of the solution-annealed sample increases continuously along with the increasing aging time, whereas for the pre-aged condition, the hardness remains almost constant during the first weeks.</p> Full article ">Figure 3
<p>Microhardness of the 6056 aluminum alloy in solution-annealed, pre-aged and naturally-aged conditions as a function of the artificial aging time at 120 °C: (<b>a</b>) without previous plastic deformation and (<b>b</b>,<b>c</b>) after linear extrusion with induced strains of (<b>b</b>) <span class="html-italic">φ</span> = 0.8 and (<b>c</b>) <span class="html-italic">φ</span> = 1.2. Depicted are the mean, minimum and maximum values of each microhardness. Plastic deformation and pre-aging increase the hardness during artificial aging.</p> Full article ">Figure 4
<p>Ultimate tensile strengths of the 6056 aluminum alloy in solution-annealed, pre-aged and naturally-aged conditions as a function of the artificial aging time after linear extrusion with <span class="html-italic">φ</span> = 1.2. Depicted here are the mean, minimum and maximum values of each tensile strength. The strength of the pre-aged condition is significantly higher when compared with the solution-annealed or naturally-aged conditions.</p> Full article ">Figure 5
<p>Optical micrographs of the 6056 aluminum alloy in the naturally-aged condition: (<b>a</b>) without previous plastic deformation and (<b>b</b>,<b>c</b>) after further linear extrusion with induced strains of (<b>b</b>) <span class="html-italic">φ</span> = 0.8 and (<b>c</b>) <span class="html-italic">φ</span> = 1.2. Because of the linear extrusion, the grains are highly elongated. This effect is more pronounced for the sample with the higher induced strain.</p> Full article ">Figure 6
<p>TEM micrographs of the 6056 aluminum alloy in (<b>a</b>,<b>b</b>) solid-solution treated, (<b>c</b>,<b>d</b>) naturally-aged and (<b>e</b>,<b>f</b>) pre-aged condition after linear extrusion with an induced strain of <span class="html-italic">φ</span> = 1.2 and subsequent artificial aging at 120 °C for 10 min. Regardless of the initial heat treatment, all conditions exhibit a high dislocation density along with the beginning arrangement in dislocation cells.</p> Full article ">
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