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23 pages, 17563 KiB  
Article
Creep Resistance and Microstructure Evolution in P23/P91 Welds
by Vlastimil Vodárek, Jan Holešinský, Zdeněk Kuboň, Renáta Palupčíková, Petra Váňová and Jitka Malcharcziková
Materials 2025, 18(1), 194; https://doi.org/10.3390/ma18010194 (registering DOI) - 5 Jan 2025
Viewed by 192
Abstract
This paper summarizes the results of investigations into heterogeneous P23/P91 welds after long-term creep exposure at temperatures of 500, 550 and 600 °C. Two variants of welds were studied: In Weld A, the filler material corresponded to P91 steel, while in Weld B, [...] Read more.
This paper summarizes the results of investigations into heterogeneous P23/P91 welds after long-term creep exposure at temperatures of 500, 550 and 600 °C. Two variants of welds were studied: In Weld A, the filler material corresponded to P91 steel, while in Weld B, the chemical composition of the consumable material matched P23 steel. The creep rupture strength values of Weld A exceeded those of Weld B at all testing temperatures. Most failures in the cross-weld samples occurred in the partially decarburized zones of P23 or WM23 steel. The results of the investigations on the minor phases were in good agreement with kinetic simulations that considered a 0.1 mm fusion zone. Microstructural studies proved that carburization occurred in the P23/P91 weld fusion zones. The partial decarburization of P23 steel or WM23 was accompanied by the dissolution of M7C3 and M23C6 particles, and detailed studies revealed the precipitation of the Fe2 (W, Mo) Laves phase in decarburized areas. Thermodynamic simulations proved that the appearance of this phase in partially decarburized P23 steel or WM23 is related to a reduction in the carbon content in these areas. According to the results of creep tests, the EBSD investigations revealed a better microstructural stability of the partially decarburized P23 steel in Weld A. Full article
(This article belongs to the Special Issue Advanced Materials Joining and Manufacturing Techniques)
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Figure 1
<p>Design of Weld A filler material. Thermanit E CrMo 9 1B matches P91 steel.</p>
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<p>Design of Weld B filler material. Thyssen Cr2WV matches P23 steel.</p>
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<p>Sketch of the cross-weld creep test sample.</p>
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<p>Results of creep rupture tests at 500, 550 and 600 °C and failure locations. Open symbols represent interrupted tests—Weld A.</p>
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<p>Results of the creep rupture tests at 500, 550 and 600 °C and failure locations—Weld B.</p>
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<p>Longitudinal section of the creep-ruptured A1 sample.</p>
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<p>Cavities in the IC/FGHAZ of P91 steel—sample A1.</p>
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<p>Precipitation of minor phases in the base material P23—sample A2.</p>
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<p>Fracture line in the partially decarburized CGHAZ of P23 steel close to the P23/WM91 fusion boundary—sample A2.</p>
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<p>Temperature dependence of carbon activity in P23 and P91 steels, calculated using Thermo-calc software.</p>
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<p>Temperature dependence of nitrogen activity in P23 and P91 steels, calculated using Thermo-calc software.</p>
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<p>Kinetic simulation of minor phase profiles across the P23/P91interface, considering the fusion zone with a thickness of 0.1 mm, calculated using Dictra software. NPM(*) - mole fractions of individual phases present in the graph.</p>
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<p>M<sub>6</sub>X and Fe<sub>2</sub> (W, Mo) Laves-phase particles in the CGHAZ of P23 steel. Inserts: [<math display="inline"><semantics> <mrow> <mover accent="true"> <mrow> <mn>4</mn> </mrow> <mo>¯</mo> </mover> <mover accent="true"> <mrow> <mn>3</mn> </mrow> <mo>¯</mo> </mover> </mrow> </semantics></math>1]<sub>Laves</sub> and [11<math display="inline"><semantics> <mrow> <mover accent="true"> <mrow> <mn>1</mn> </mrow> <mo>¯</mo> </mover> </mrow> </semantics></math>]<sub>M6X</sub>. Sample A2.</p>
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<p>Distribution of M<sub>6</sub>X and Fe<sub>2</sub> (W, Mo). Laves-phase particles in the CGHAZ of P23 steel—BSE image. Sample A2.</p>
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<p>Precipitation in base material P91—sample A2.</p>
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<p>Heavy precipitation in the carburized fusion zone of P23/WM91. Insert: Ring diffraction pattern of the M<sub>23</sub>C<sub>6</sub> phase. Sample A2.</p>
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<p>Microhardness (HV 0.01) and chromium gradients across the P23/WM91 fusion zone—sample A2.</p>
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<p>Longitudinal section of the ruptured sample B1.</p>
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<p>Microhardness (HV 0.01) and chromium profiles across the WM23/P91 interface—sample B1.</p>
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<p>Precipitation in the WM23/P91 fusion zone (mainly M<sub>23</sub>C<sub>6</sub> and some Mo- and W-rich M<sub>6</sub>X particles): (<b>a</b>) bright field; (<b>b</b>–<b>d</b>) X-ray maps of Cr, Mo and W. Sample B1.</p>
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<p>Precipitation in the WM23/P91 fusion zone (mainly M<sub>23</sub>C<sub>6</sub> and some Mo- and W-rich M<sub>6</sub>X particles): (<b>a</b>) bright field; (<b>b</b>–<b>d</b>) X-ray maps of Cr, Mo and W. Sample B1.</p>
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<p>Partly decarburized zone of WM23: (<b>a</b>) bright field (insert: spot diffraction pattern of Fe<sub>2</sub> (W, Mo) Laves phase with zone axis <math display="inline"><semantics> <mrow> <mrow> <mo stretchy="false">[</mo> <msub> <mrow> <mn>11</mn> <mover accent="true"> <mrow> <mn>1</mn> </mrow> <mo>¯</mo> </mover> <mo stretchy="false">]</mo> </mrow> <mrow> <mi mathvariant="normal">L</mi> <mi mathvariant="normal">a</mi> <mi mathvariant="normal">v</mi> <mi mathvariant="normal">e</mi> <mi mathvariant="normal">s</mi> </mrow> </msub> </mrow> </mrow> </semantics></math>); (<b>b</b>,<b>c</b>) X-ray maps of Mo and W. Sample B1.</p>
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<p>Effect of carbon content on equilibrium minor phases in P23 steel. The red line shows the carbon content in the heat investigated, calculated using Thermo-Calc software.</p>
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<p>IPF (normal direction) + image quality maps across the fusion zones P23/WM91 and WM23/P91: (<b>a</b>) sample A2, Weld A; (<b>b</b>) sample B1, Weld B; (<b>c</b>) color coding.</p>
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28 pages, 4725 KiB  
Review
High Energy Density Welding of Ni-Based Superalloys: An Overview
by Riccardo Donnini, Alessandra Varone, Alessandra Palombi, Saveria Spiller, Paolo Ferro and Giuliano Angella
Metals 2025, 15(1), 30; https://doi.org/10.3390/met15010030 - 1 Jan 2025
Viewed by 291
Abstract
High energy density technologies for welding processes provide opportune solutions to joint metal materials and repair components in several industrial applications. Their high-performance levels are related to the high penetration depth and welding speed achievable. Moreover, the localized thermal input helps in reducing [...] Read more.
High energy density technologies for welding processes provide opportune solutions to joint metal materials and repair components in several industrial applications. Their high-performance levels are related to the high penetration depth and welding speed achievable. Moreover, the localized thermal input helps in reducing distortion and residual stresses in the welds, minimizing the extension of the fusion zone and heat-affected zone. The use of these welding technologies can be decisive in the employment of sophisticated alloys such as Ni-based superalloys, which are notoriously excellent candidates for industrial components subjected to high temperatures and corrosive work conditions. Nonetheless, the peculiar crystallographic and chemical complexity of Ni-based superalloys (whether characterized by polycrystalline, directionally solidified, or single-crystal microstructure) leads to high susceptibility to welding processes and, in general, challenging issues related to the microstructural features of the welded joints. The present review highlights the advantages and drawbacks of high energy density (Laser Beam and Electron Beam) welding techniques applied to Ni-based superalloy. The effects of process parameters on cracking susceptibility have been analyzed to better understand the correlation between them and the microstructure-mechanical properties of the welds. The weldability of three different polycrystalline Ni superalloys, one solid solution-strengthened alloy, Inconel 625, and two precipitation-strengthen alloys, Nimonic 263 and Inconel 718, is reviewed in detail. In addition, a variant of the latter, the AF955 alloy, is also presented for its great potential in terms of weldability. Full article
(This article belongs to the Special Issue Advanced Welding Technology in Metals III)
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Figure 1
<p>Weldability reference diagram for Ni-based superalloys depending on chemical composition (%wt.) Reprinted from ref. [<a href="#B33-metals-15-00030" class="html-bibr">33</a>]. Copyright 2002 Sage Publications.</p>
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<p>Schematic of the weld width with different welding techniques. Reprinted from Ref. [<a href="#B48-metals-15-00030" class="html-bibr">48</a>].</p>
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<p>(<b>a</b>) Sketches of Laser Beam Welding, and (<b>b</b>) Electron Beam Welding processes. Reprinted from Ref. [<a href="#B54-metals-15-00030" class="html-bibr">54</a>].</p>
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<p>Influence of G and R parameters on the possible resulting solidified microstructure. Reprinted from Refs. [<a href="#B52-metals-15-00030" class="html-bibr">52</a>,<a href="#B53-metals-15-00030" class="html-bibr">53</a>].</p>
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<p>Representation of the solidification processes and constituents formation in FZ (equivalent to melted zone MZ) and HAZ of GH909 Ni-based superalloy. Reprinted from Ref. [<a href="#B86-metals-15-00030" class="html-bibr">86</a>].</p>
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<p>Constituents formation in the fusion zone in an IN738 Ni-based superalloy by Nd:YAG pulsed laser welding process. Reprinted from Ref. [<a href="#B31-metals-15-00030" class="html-bibr">31</a>]. HAZ liquation cracks associated with: (<b>a</b>) Ni-Zr intermetallic, (<b>b</b>) γ-γ′ eutectic and Cr–Mo boride (<b>c</b>) Cr rich Carbides and (<b>d</b>) Ta rich MC carbides.</p>
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<p>Details by scanning electron micrograph for the microstructure of the FZ in a Haynes282 welded by CO<sub>2</sub> laser beam. Reprinted from Ref. [<a href="#B89-metals-15-00030" class="html-bibr">89</a>].</p>
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<p>Schematic of solidification cracking formation. Reprinted from Ref. [<a href="#B58-metals-15-00030" class="html-bibr">58</a>].</p>
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<p>Pulsed laser weld on PWA 1480 single crystal Ni-based superalloy showing large regions of epitaxial growth as well as abundant stray grains and cracking along stray-grain high angle boundaries. Reprinted from Ref. [<a href="#B94-metals-15-00030" class="html-bibr">94</a>]. Copyright 2004 The Minerals, Metals &amp; Materials Society.</p>
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<p>Comparison between the cross-sections obtained by GTAW (first row) and EBW (second row) varying the process parameters. Reprinted from Ref. [<a href="#B53-metals-15-00030" class="html-bibr">53</a>].</p>
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<p>Typical microstructure of an EB weld. Reprinted from Ref. [<a href="#B107-metals-15-00030" class="html-bibr">107</a>].</p>
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<p>Schematic of the solidification process of GDT-111 nickel-based superalloy. Reprinted from Ref. [<a href="#B112-metals-15-00030" class="html-bibr">112</a>]. Schematic of GTD-111 superalloy solidification steps, (<b>a</b>) Liquid metal, (<b>b</b>) Formed dendrite structure (γ1), (<b>c</b>) The residual liquid in IDR, (<b>e</b>) Precipitation of MC carbides, (<b>f</b>) Formation of γ-γ′ eutectic, (<b>g</b>) precipitation of ordered γ′ phase.</p>
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21 pages, 17144 KiB  
Article
Failure and Degradation Mechanisms of Steel Pipelines: Analysis and Development of Effective Preventive Strategies
by Marcin Kowalczyk, Jakub Andruszko, Paweł Stefanek and Robert Mazur
Materials 2025, 18(1), 134; https://doi.org/10.3390/ma18010134 - 31 Dec 2024
Viewed by 387
Abstract
The increasing challenges related to the reliability and durability of steel pipeline infrastructure necessitate a detailed understanding of degradation and failure mechanisms. This study focuses on selective corrosion and erosion as critical factors, analyzing their impact on pipeline integrity using advanced methods, including [...] Read more.
The increasing challenges related to the reliability and durability of steel pipeline infrastructure necessitate a detailed understanding of degradation and failure mechanisms. This study focuses on selective corrosion and erosion as critical factors, analyzing their impact on pipeline integrity using advanced methods, including macroscopic analysis, corrosion testing, microscopic examination, tensile strength testing, and finite element method (FEM) modeling. Selective corrosion in the heat-affected zones (HAZs) of longitudinal welds was identified as the dominant degradation mechanism, with pit depths reaching up to 6 mm, leading to tensile strength reductions of 30%. FEM analysis showed that material loss exceeding 8 mm in weld areas under standard operating pressure (16 bar) induces critical stress levels, risking pipeline failure. Erosion was found to exacerbate selective corrosion, accelerating degradation in high-stress zones. Practical recommendations include the use of corrosion-resistant materials, such as duplex steels, and implementing integrated monitoring strategies combining non-destructive testing with FEM-based predictive modeling. These insights contribute to developing robust preventive measures to ensure the safety and longevity of pipeline infrastructure. Full article
(This article belongs to the Special Issue Advances in Corrosion and Protection of Metallic Materials)
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<p>Example segments of the analyzed pipes.</p>
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<p>Example of a damaged water pipeline segment—longitudinal seam rupture: (<b>a</b>) General view along the crack and (<b>b</b>) close-up view of the seam rupture location.</p>
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<p>Weld from the inside of the pipe on the unruptured section after cleaning.</p>
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<p>Hard, compact, subsurface permeable corrosion product coatings that separate the pipe material from the water flow inside the pipe: (<b>a</b>) Start of the pipe seam and (<b>b</b>) mechanical separation of the crust.</p>
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<p>Pipe after the removal of corrosion products, showing characteristic pits that degrade the longitudinal weld of the pipe.</p>
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<p>Pipe with a spiral SAWH seam: Longitudinal seam rupture in pipe P1, with visible defects in the weld: (<b>a</b>) Seam with visible corrosion products and (<b>b</b>) seam after cleaning off the corrosion products.</p>
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<p>Water pipe with longitudinal seam—mechanism of selective corrosion formation: 1—steel pipe shell of the pipeline; 2—trapped water layer between the pipe material and the layer of corrosion products; 3—compact, semi-permeable layer of corrosion products and contaminants; 4—external protective layer; 5—areas of selective corrosion development along the HAZ of the longitudinal weld; 6—local pits as a result of selective corrosion. σr—circumferential stresses in the pipe; p—network pressure.</p>
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<p>View of the remaining weld remnants at the edge of the joint with the longitudinal weld of the fractured pipe.</p>
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<p>View of the internal surface of the pipe segment illustrating the degree of weld degradation for water transport pipes.</p>
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<p>Division of the pipe segment into samples for macroscopic analysis (1–8) and mechanical testing (SP1–SP3).</p>
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<p>Macrostuctures of welded joint samples: (<b>a</b>) Sample 1, (<b>b</b>) Sample 2, (<b>c</b>) Sample 6, (<b>d</b>) Sample 8 from <a href="#materials-18-00134-f010" class="html-fig">Figure 10</a>.</p>
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<p>Anodic areas (<b>blue</b>) and cathodic areas (<b>pink</b>) in the welded joints of the pipe: (<b>a</b>) Sample 6, (<b>b</b>) Sample 8 from <a href="#materials-18-00134-f010" class="html-fig">Figure 10</a>.</p>
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<p>Potential profile: (<b>a</b>) Base material; (<b>b</b>) heat affected zone (HAZ) area; (<b>c</b>) weld area.</p>
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<p>Microscopic analysis—Pits in HAZ: (<b>a</b>) Sample 1, (<b>b</b>) Sample 2, (<b>c</b>) Sample 6, (<b>d</b>) Sample 8 from <a href="#materials-18-00134-f010" class="html-fig">Figure 10</a>.</p>
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<p>Samples after strength testing: (<b>a</b>)—SP1; (<b>b</b>)—SP2; (<b>c</b>)—SP3.</p>
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<p>Tensile curves: (<b>a</b>)—SP1; (<b>b</b>)—SP2; (<b>c</b>)—SP3 where: green—Test 1, blue—Test 2, red—Test 3.</p>
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<p>Schematics for modeling the pipe section and example model dimensions: (<b>a</b>) base weld cross-section dimensions; (<b>b</b>) schematic of weld material loss.</p>
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<p>Circumferential normal stresses [MPa] at a pressure of 16 bar at various levels of longitudinal weld degradation.</p>
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<p>Equivalent plastic strains in the material at a pressure of 16 bar at various levels of longitudinal weld degradation.</p>
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<p>Sharp termination of the notch bottom in the pit caused by selective corrosion: (<b>a</b>) Dimensions of the model in the second stage of numerical analysis and (<b>b</b>) discrete model in the pit area during the second stage of numerical analysis.</p>
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<p>Results of the numerical analyses for the second stage: (<b>a</b>) Circumferential normal stress in [MPa] at a pressure of 30 bar and a degradation depth of 6 mm and (<b>b</b>) equivalent plastic strain at a pressure of 30 bar and a degradation depth of 6 mm.</p>
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<p>Plastic deformation as a function of pressure at the bottom of the hole in the longitudinal weld of the pipe depending on the depth of loss of material.</p>
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17 pages, 10666 KiB  
Article
Prediction of Mechanical Properties and Fracture Behavior of TC17 Linear Friction Welded Joint Based on Finite Element Simulation
by Xuan Xiao, Yue Mao and Li Fu
Materials 2025, 18(1), 128; https://doi.org/10.3390/ma18010128 - 31 Dec 2024
Viewed by 301
Abstract
TC17 titanium alloy is widely used in the aviation industry for dual-performance blades, and linear friction welding (LFW) is a key technology for its manufacturing and repair. However, accurate evaluation of the mechanical properties of TC17−LFW joints and research on their joint fracture [...] Read more.
TC17 titanium alloy is widely used in the aviation industry for dual-performance blades, and linear friction welding (LFW) is a key technology for its manufacturing and repair. However, accurate evaluation of the mechanical properties of TC17−LFW joints and research on their joint fracture behavior are still not clear. Therefore, this paper used the finite element numerical simulation method (FEM) to investigate the mechanical behavior of the TC17−LFW joint with a complex micro−structure during the tensile processing, and predicted its mechanical properties and fracture behavior. The results indicate that the simulated elastic modulus of the joint is 108.5 GPa, the yield strength is 1023.2 MPa, the tensile strength is 1067.5 MPa, and the elongation is 1.98%. The deviations from measured results between simulated results are less than 2%. The stress and strain field studies during the processing show that the material located at the upper and lower edges of the joint in the WZ experiences stress and strain concentration, followed by the extending of the stress and strain concentration zone toward the center of the WZ. And finally, the strain concentration zone covered the entire WZ. The fracture behavior studies show that the material necking occurs in the TMAZ of TC17(α + β) and WZ, while cracks first appear in the WZ. Subsequently, joint cracks propagate along the TC17(α + β) side of the WZ until fracture occurs. There are obvious tearing edges formed by the partial tearing of the WZ structure in the simulated fracture surface, and there are fracture surfaces with different height differences at the center of the joint crack, indicating that the joint has mixed fracture characteristics. Full article
(This article belongs to the Special Issue Advanced Materials Joining and Manufacturing Techniques)
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<p>Diagram flow of the methodology and investigation steps.</p>
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<p>LFW machine setup and joint sample: (<b>a</b>) machine; (<b>b</b>) operation mode; (<b>c</b>) TC17−LFW joint.</p>
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<p>Morphology of TC17−LFW Joint.</p>
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<p>Schematic diagram of tensile specimen.</p>
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<p>Geometric model and mesh of tensile processing in TC17−LFW joint: (<b>a</b>) geometric model; (<b>b</b>) mesh.</p>
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<p>Stress–strain curves of different areas of TC17−LFW joint obtained by inverse material parameters using nanoindentation load displacement curves.</p>
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<p>Material parameter settings of tensile processing model in TC17−LFW joint.</p>
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<p>Boundary conditions of tensile processing model in TC17−LFW joint.</p>
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<p>Boundary conditions of tensile processing model in TC17−LFW joint.</p>
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<p>Stress field during tensile processing of TC17−LFW joint: (<b>a</b>) after yielding; (<b>b</b>) deformed 45%; (<b>c</b>) deformed 60%; (<b>d</b>) Deformed 90%.</p>
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<p>Strain field during tensile processing of TC17−LFW joint: (<b>a</b>) after yielding; (<b>b</b>) deformed 45%; (<b>c</b>) deformed 60%; (<b>d</b>) deformed 90%.</p>
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<p>Stress field during tensile processing of surface XOY of TC17−LFW joint: (<b>a</b>) deformed 45%; (<b>b</b>) deformed 60%; (<b>c</b>) deformed 90%.</p>
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<p>Strain field during tensile processing of surface XOY of TC17−LFW joint: (<b>a</b>) deformed 45%; (<b>b</b>) deformed 60%; (<b>c</b>) deformed 90%.</p>
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<p>DIC full-field strain during tensile processing of TC17−LFW joint: (<b>a</b>) deformed 45%; (<b>b</b>) deformed 60%; (<b>c</b>) deformed 90%.</p>
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<p>Neck shrinkage deformation and fracture behavior during tensile processing of TC17−LFW joint: (<b>a</b>) initiation; (<b>b</b>) neck shrinkage; (<b>c</b>) cracking; (<b>d</b>) fracture.</p>
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<p>Simulation fracture morphology of TC17−LFW joint.</p>
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<p>Fracture morphology of TC17−LFW joint: (<b>a</b>) overall morphology; (<b>b</b>–<b>d</b>) local fracture morphology.</p>
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14 pages, 6153 KiB  
Article
Forging-Submerged Arc Additive Hybrid Manufacturing of the Mn-Mo-Ni Component: In Situ Reheat Cycles Inducing the Homogenization of the HAZ Microstructure
by Qiang Chi, Meijuan Hu, Jun Wang, Shuai Yan, Manye Xue, Shaojie Wu and Fangjie Cheng
Materials 2025, 18(1), 20; https://doi.org/10.3390/ma18010020 - 25 Dec 2024
Viewed by 377
Abstract
Forging additive hybrid manufacturing integrated the high efficiency of forging and the great flexibility of additive manufacturing, which has significant potential in the construction of reactor pressure vessels (RPVs). In the components, the heat-affected zone (HAZ, also called as bonding zone) between the [...] Read more.
Forging additive hybrid manufacturing integrated the high efficiency of forging and the great flexibility of additive manufacturing, which has significant potential in the construction of reactor pressure vessels (RPVs). In the components, the heat-affected zone (HAZ, also called as bonding zone) between the forged substrate zone and the arc deposition zone was key to the final performance of the components. In this study, the Mn-Mo-Ni welding wire was deposited on the 16MnD5 substrate with a submerged arc heat source. The in situ reheat cycle effect of the submerged arc heat source on the microstructure and mechanical properties of the HAZ were studied. The results showed that the HAZ underwent four heat treatment processes, including two full austenitizing stages, one high-temperature stage, and continuous low-temperature tempering, which formed a homogenized microstructure in the HAZ and was mainly composed of tempered sorbite (Tempered-S). The HAZ microhardness is around 278.7 HV, which is about 150 HV lower than the microhardness only conducted by one thermal cycle. Furthermore, the effects of preheating the substrate and adjusting the heat inputs on the HAZ were studied. The results indicated that the clustered cementite was precipitated, which destroys the low-temperature impact toughness of the HAZ after preheating. A suitable heat input not only homogenized the microstructure within the HAZ but also promoted the transformation of grains into equiaxed grains. The −60 °C impact toughness of the HAZ was significantly increased from 96.7 J to 113 J. Full article
(This article belongs to the Special Issue Microstructure Engineering of Metals and Alloys, 3rd Edition)
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Figure 1
<p>(<b>a</b>) The forging-submerged arc additive hybrid manufacturing device, (<b>b</b>) deposition strategy and representation of the extraction location of the tested samples.</p>
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<p>The evolution process of the microstructure of the HAZ with different thermal cycling conditions: (<b>a</b>) one thermal cycle, (<b>b</b>) two thermal cycles, (<b>c</b>) three thermal cycles, (<b>d</b>) four thermal cycles, (<b>e</b>) reheating cycle curves.</p>
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<p>The HAZ of the components by forging-submerged arc additive hybrid manufacturing after undergoing four thermal cycles. (<b>a</b>) the microstructure of the HAZ, (<b>b</b>) a partial enlarged view of (<b>a</b>).</p>
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<p>EBSD characterization of microstructures in the HAZ after undergoing four thermal cycles: (<b>a</b>) IPF map, (<b>b</b>) GB map, (<b>c</b>) phase map, (<b>d</b>) misorientation angle distribution map.</p>
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<p>Microhardness of the HAZ after undergoing different thermal cycles: (<b>a</b>) under different thermal cycles, (<b>b</b>) under different processing techniques [<a href="#B32-materials-18-00020" class="html-bibr">32</a>,<a href="#B33-materials-18-00020" class="html-bibr">33</a>,<a href="#B34-materials-18-00020" class="html-bibr">34</a>]. (SAW [<a href="#B32-materials-18-00020" class="html-bibr">32</a>], submerged arc welding; TAAW [<a href="#B33-materials-18-00020" class="html-bibr">33</a>], tungsten argon arc welding; LTTA [<a href="#B34-materials-18-00020" class="html-bibr">34</a>], long-term thermal aging.)</p>
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<p>Impact fracture. (<b>a</b>). SEM of the fracture surface of the impact specimen, enlarged SEM of the selected areas marked as “b” and “c” in (<b>a</b>), (<b>b</b>). and (<b>c</b>), respectively.</p>
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<p>Continuous cooling transformation curve.</p>
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<p>Impact toughness of the Mn-Ni-Mo alloy by SAAM, SAW, and forging [<a href="#B25-materials-18-00020" class="html-bibr">25</a>,<a href="#B36-materials-18-00020" class="html-bibr">36</a>,<a href="#B37-materials-18-00020" class="html-bibr">37</a>,<a href="#B38-materials-18-00020" class="html-bibr">38</a>,<a href="#B39-materials-18-00020" class="html-bibr">39</a>].</p>
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<p>Microstructure and crystallographic characteristics of the HAZ with different heat parameters: (<b>a</b>–<b>d</b>) preheating at 150 °C of 1800 J/mm, (<b>e</b>–<b>h</b>) 2250 J/mm, (<b>i</b>–<b>l</b>) 2700 J/mm.</p>
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16 pages, 6797 KiB  
Article
Improving the Metal Inert Gas Welding Efficiency and Microstructural Stability in the Butt and Lap Joints of Aluminum Automotive Components Using Sc- and Zr-Enhanced Filler Wires
by Hansol Ko, Hye-Jin Kim, Dong-Yoon Kim and Jiyoung Yu
Metals 2025, 15(1), 1; https://doi.org/10.3390/met15010001 - 24 Dec 2024
Viewed by 291
Abstract
The grain growth in the fusion zone (FZ) and heat-affected zone (HAZ) of metal inert gas (MIG) welding processes negatively affect the mechanical properties of aluminum alloy MIG welds used in automotive components. Although the addition of Sc- and Zr-based filler wires can [...] Read more.
The grain growth in the fusion zone (FZ) and heat-affected zone (HAZ) of metal inert gas (MIG) welding processes negatively affect the mechanical properties of aluminum alloy MIG welds used in automotive components. Although the addition of Sc- and Zr-based filler wires can refine weld microstructures and enhance the mechanical properties, conditions resembling actual automotive component joints have not been sufficiently investigated. In this study, 5083-O aluminum alloy base material was welded into butt and lap joints using conventional 5000-series aluminum alloy filler wires (Al-5.0Mg) and wires containing Sc and Zr (Al-4.8Mg-0.7Sc-0.3Zr) under various heat input conditions. The mechanical properties of the welds were evaluated via tensile tests, and the microstructures in the FZ and HAZ were analyzed. In butt joints, Al-4.8Mg-0.7Sc-0.3Zr exhibited a finer and more uniform grain structure with increased tensile strength compared with those welded using Al-5.0Mg. The microstructure became coarser with the increased heat input, and the tensile strength tended to decrease. In lap joints, the tensile-shear strength of Al-4.8Mg-0.7Sc-0.3Zr was higher than that of Al-5.0Mg; it further increased with the increase in the amount of deposited metal. The coarsening of the microstructure with the increased heat input was disadvantageous for the tensile-shear strength, and the increased weld size offset the adverse effects of the coarse microstructure. These results indicate that the heat input and the amount of deposited metal must be optimized to ensure stiffness in various joints of automotive components. Full article
(This article belongs to the Special Issue Welding and Joining of Advanced High-Strength Steels (2nd Edition))
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<p>Weld joint preparation: (<b>a</b>) butt joint; (<b>b</b>) lap joint.</p>
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<p>Weld joint preparation: (<b>a</b>) butt joint; (<b>b</b>) lap joint.</p>
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<p>Configuration of the tensile test specimens: (<b>a</b>) top view; (<b>b</b>) side view of the butt weld; (<b>c</b>) side view of the lap weld.</p>
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<p>Cross-sectional image of a butt joint weld with a filler wire: (<b>a</b>) Al-5.0Mg wire at a feed rate (WFR) of 5.0 m/min; (<b>b</b>) Al-4.8Mg-0.7Sc-0.3Zr at a WFR of 5.0 m/min.</p>
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<p>Relationship between WFR changes and the tensile strength of the butt weld joints.</p>
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<p>Hardness of butt-welded joints with each wire (WFR 5.0 m/min).</p>
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<p>Microstructural images of butt joints in AA5083 with a filler metal: (<b>a</b>) Al-5.0Mg, WFR = 5.0 m/min; (<b>b</b>) Al-5.0Mg, WFR = 7.0 m/min; (<b>c</b>) Al-5.0Mg, WFR = 9.0 m/min; (<b>d</b>) Al-4.8Mg-0.7Sc-0.3Zr, WFR = 5.0 m/min; (<b>e</b>) Al-4.8Mg-0.7Sc-0.3Zr, WFR = 7.0 m/min; (<b>f</b>) Al-4.8Mg-0.7Sc-0.3Zr, WFR = 9.0 m/min.</p>
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<p>Electron backscatter diffractometer (EBSD) images of butt joints in AA5083 with a filler metal (WFR = 5.0 m/min, white box: grain size measurement location): (<b>a</b>) Al-5.0Mg; (<b>b</b>) Al-4.8Mg-0.7Sc-0.3Zr.</p>
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<p>Cross-sectional images of the lap weld based on variations in the WFR.</p>
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<p>Tensile-shear strength and fracture morphology in the lap weld according to variations in the WFR.</p>
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<p>Relationship between the weld size of a lap joint and the tensile-shear strength.</p>
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<p>Relationship between the hardness of the fusion zone (FZ) in lap joints and variations in the WFR.</p>
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<p>Welded joint efficiency in the butt and lap joints based on variations in the WFR.</p>
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16 pages, 5953 KiB  
Article
Microstructural and Electrochemical Analysis of the Physically Simulated Heat-Affected Zone of Super-Duplex Stainless Steel UNS S32750
by Francisco Magalhães dos Santos, Leonardo Oliveira Passos da Silva, Ygor Tadeu Bispo dos Santos, Bruna Callegari, Tiago Nunes Lima and Rodrigo Santiago Coelho
Metals 2025, 15(1), 2; https://doi.org/10.3390/met15010002 - 24 Dec 2024
Viewed by 457
Abstract
Super-duplex stainless steels (SDSSs) were introduced in the oil and gas industry due to their high resistance to pitting corrosion, promoted by the high content of alloying elements. The welding process can cause an unbalanced ferrite/austenite microstructure and, consequently, the possibility of deleterious [...] Read more.
Super-duplex stainless steels (SDSSs) were introduced in the oil and gas industry due to their high resistance to pitting corrosion, promoted by the high content of alloying elements. The welding process can cause an unbalanced ferrite/austenite microstructure and, consequently, the possibility of deleterious phases, increasing the risk of failure. The aim of this work is to investigate the behavior of the heat-affected zone (HAZ) of SDSS UNS S32750 steel produced with different thermal inputs simulated in a Gleeble® welding simulator and correlate these findings with its corrosion properties. The pitting resistance was investigated by electrochemical techniques in sodium chloride solution, and the critical pitting temperature (CPT) was calculated for each evaluated microstructure. The material as received presents 46.19 vol% ferrite and a high corrosion resistance, with a CPT of 71.54 °C. HAZ-simulated cycles resulted in similar ferrite percentages, between 54.09 vol% and 57.25 vol%. A relationship was found between heat input, ferrite content, and CPT: increasing the heat input results in greater ferrite content and lowers the CPT, which may favor the pitting corrosion process. Therefore, it is concluded that the ferrite content directly influences the pitting behavior of the material. Full article
(This article belongs to the Special Issue Welding and Joining of Advanced High-Strength Steels (2nd Edition))
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<p>Flowchart of tests and specimens.</p>
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<p>Thermal curves used for the three different conditions.</p>
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<p>Experimental curves obtained after simulation of HTHAZ along with the reference (calculated) curve for all conditions.</p>
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<p>Optical microscopy images showing the microstructure of the super-duplex steel in all studied conditions, along with respective ferrite fractions obtained by image analysis. The rolling direction is parallel to the width of the images.</p>
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<p>EBSD phase maps showing ferrite (red) and austenite (green). The rolling direction is parallel to the height of the images.</p>
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<p>EDS maps showing Cr, Ni, and Mo distribution in the as-received condition. The rolling direction is parallel to the width of the images.</p>
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<p>PREN evolution in ferrite and austenite after the application of different cycles.</p>
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<p>Potentiodynamic polarization curves in 3.5% (m/V) NaCl solution, for as-received and heat-treated conditions.</p>
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<p>Correlation between CPT and ferrite amount in all conditions.</p>
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<p>SEM images showing the microstructures of the as-received condition after polarization tests at 70 °C and of conditions I–III after polarization tests at 50 °C. The black regions correspond to the pits formed after polarization.</p>
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16 pages, 24622 KiB  
Article
Welding Pores Evolution in the Detector Bottom-Locking Structure Fabricated Using the Hybrid Pulsed Arc–Laser Method
by Yonglong Yu, Jianzhou Xu, Xiaoquan Yu, Liang Guo, Tongyu Zhu and Ding Fan
Metals 2024, 14(12), 1469; https://doi.org/10.3390/met14121469 - 23 Dec 2024
Viewed by 349
Abstract
The welding of the bottom-locking structure in a detector receptacle plays an essential role in ensuring the safety of nuclear equipment. A pulsed TIG–laser hybrid welding method is proposed to address the problem of welding pores in locking structural parts. The effects of [...] Read more.
The welding of the bottom-locking structure in a detector receptacle plays an essential role in ensuring the safety of nuclear equipment. A pulsed TIG–laser hybrid welding method is proposed to address the problem of welding pores in locking structural parts. The effects of the pulse frequency on the escape of porosity and of porosity on the mechanical properties of the hybrid welding joint were investigated. The results were compared to those of direct current (0 Hz), showing that the pulse frequency affects the stability of the arc. With an increase in pulse frequency, the grain size of the fusion zone gradually decreases, and the flow in the middle area of the molten pool increases. This subjects bubbles in the molten pool to a thrust force, which causes the bubbles to escape to the surface of the molten pool. Compared with 0 Hz, the tensile strength of the joint increased by 67%. This provides a new solution for obtaining reliable welded joints for the bottom-locking structure of detector storage tanks. Full article
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<p>(<b>a</b>) Schematic of the laser–TIG hybrid welding system; (<b>b</b>) schematic of the positional relationship between the laser and the arc; (<b>c</b>) dimensions of the butt plate and the bottom pad plate.</p>
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<p>Schematic of metallographic and tensile specimens.</p>
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<p>High-speed photographic images at different pulse frequencies: (<b>a</b>) 50 Hz; (<b>b</b>) 250 Hz; (<b>c</b>) 500 Hz.</p>
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<p>Cross-sectional morphology of welded joints.</p>
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<p>Cross-section of the joint at different pulse frequencies.</p>
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<p>Cross-sectional histomorphology of joints at different pulse frequencies: (<b>a</b>) 0 Hz; (<b>b</b>) 50 Hz; (<b>c</b>) 250 Hz; (<b>d</b>) 500 Hz.</p>
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<p>The enlarged views of the fusion line (the white box area) in <a href="#metals-14-01469-f006" class="html-fig">Figure 6</a>: (<b>a</b>) 0 Hz; (<b>b</b>) 50 Hz; (<b>c</b>) 250 Hz; (<b>d</b>) 500 Hz.</p>
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<p>Hardness distribution of joints at different pulse frequencies.</p>
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<p>Stress–strain curve of the joint.</p>
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<p>SEM fracture microstructure: (<b>a</b>) base material; (<b>b</b>) joint (250 Hz); (<b>c</b>) joint (0 Hz); (<b>d</b>) 5000× magnification (base material); (<b>e</b>) 5000× magnification (250 Hz); (<b>f</b>) 5000× magnification (250 Hz).</p>
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<p>Three-dimensional model.</p>
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<p>Heat source model: (<b>a</b>) Gaussian surface heat source; (<b>b</b>) cylinder heat source; (<b>c</b>) combined heat source.</p>
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<p>Simulation results of the weld cross-section.</p>
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<p>Comparison score of joint cross-section morphology.</p>
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<p>Flow field of the molten pool at different pulse frequencies: (<b>a</b>) 0 Hz; (<b>b</b>) 250 Hz.</p>
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<p>Schematic of flow field at different pulse frequencies: (<b>a</b>) 0 Hz; (<b>b</b>) 50 Hz; (<b>c</b>) 250 Hz; (<b>d</b>) 500 Hz.</p>
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<p>Schematic of laser–TIG hybrid weld pool.</p>
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<p>Schematic of the bubble forces.</p>
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27 pages, 17020 KiB  
Article
Evaluation of the Wear of Ni 200 Alloy After Long-Term Carbon Capture in Molten Salts Process
by Piotr Palimąka, Stanisław Pietrzyk, Maciej Balcerzak, Krzysztof Żaba, Beata Leszczyńska-Madej and Justyna Jaskowska-Lemańska
Materials 2024, 17(24), 6302; https://doi.org/10.3390/ma17246302 - 23 Dec 2024
Viewed by 332
Abstract
Reducing CO2 emissions is one of the major challenges facing the modern world. The overall goal is to limit global warming and prevent catastrophic climate change. One of the many methods for reducing carbon dioxide emissions involves capturing, utilizing, and storing it [...] Read more.
Reducing CO2 emissions is one of the major challenges facing the modern world. The overall goal is to limit global warming and prevent catastrophic climate change. One of the many methods for reducing carbon dioxide emissions involves capturing, utilizing, and storing it at the source. The Carbon Capture in Molten Salts (CCMS) technique is considered potentially attractive and promising, although it has so far only been tested at the laboratory scale. This study evaluates the wear of the main structural components of a prototype for CO2 capture in molten salts—a device designed and tested in the laboratories of AGH University of Kraków. The evaluation focused on a gas barbotage lance and a reactor chamber (made from Nickel 200 Alloy), which were in continuous, long-term (800 h) contact with molten salts CaCl2-CaF2-CaO-CaCO3 at temperatures of 700–940 °C in an atmosphere of N2-CO2. The research used light microscopy, SEM, X-ray, computed tomography (CT), and 3D scanning. The results indicate the greatest wear on the part of the lance submerged in the molten salts (3.9 mm/year). The most likely wear mechanism involves grain growth and intergranular corrosion. Nickel reactions with the aggressive salt environment and its components cannot be ruled out. Additionally, the applied research methods enabled the identification of material discontinuities in the reactor chamber (mainly in welded areas), pitting on its surface, and uneven wear in different zones. Full article
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<p>The concept of the carbon capture in molten salts process.</p>
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<p>Diagram of the CCMS process using two chambers and an intermediate tank: TC—thermocouple, MS—molten salts CaCl<sub>2</sub>-CaF<sub>2</sub>-CaO-CaCO<sub>3</sub>.</p>
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<p>The 3D appearance of a reactor chamber with inlet and outlet transport pipes for molten salts placed within a heating module.</p>
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<p>View of the prototype reactor for carbon capture at the AGH laboratory, (<b>a</b>) overall view: 1, 2—power supplies; 3—absorber; 4—desorber; 5—intermediate tank, (<b>b</b>) view of the system without additional heating modules: 6—high temperature valve; 7—transporting pipe, (<b>c</b>) view of the system with additional heating modules: 8—heating module; (<b>d</b>) view of the system with valves insulation: 9—ceramic wool insulation.</p>
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<p>Changes in carbon dioxide concentration during a single, complete operating cycle of the CO<sub>2</sub> capture reactor chamber (green line) and process temperature (red line).</p>
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<p>View: (<b>a</b>) cross-sectional view of the reactor chamber with marked zones with different operating conditions, (<b>b</b>) lance after 800 h tests with marked zones subjected to analysis, (<b>c</b>) reactor chamber after 800 h tests with marked areas and cut-out sections subjected to analysis.</p>
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<p>Reaction chamber: (<b>a</b>) main initial dimensions (before experiments), in mm, (<b>b</b>) appearance before testing.</p>
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<p>Appearance of samples L1, L2, and L3 and wall thickness measurement points (light microscopy—macro image).</p>
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<p>Images of the L3 lance section after testing and before testing (L0). A 50% image transparency was applied for L0.</p>
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<p>Microstructure images of sample L3, SEM: (<b>a</b>) magnification 50×, (<b>b</b>) inner section, magnification 200×, (<b>c</b>) outer section, magnification 500×.</p>
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<p>Ni, Ca, Cl distribution maps of sample L3; SEM/EDS.</p>
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<p>Microstructure images of sample L2, SEM: (<b>a</b>) magnification 50×, (<b>b</b>) inner section, magnification 500×, (<b>c</b>) outer section, magnification 500×.</p>
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<p>Ni, Ca, Cl distribution maps of sample L2; SEM/EDS.</p>
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<p>Microstructure images of sample L1, SEM: (<b>a</b>) magnification 50×, (<b>b</b>) inner section, magnification 500×, (<b>c</b>) outer section, magnification 500×.</p>
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<p>Ni, O distribution maps of sample L1; SEM/EDS.</p>
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<p>Comparison of grain size in Ni Alloy 200 before testing (L0) and after testing (L1, L2, L3).</p>
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<p>Selected inspection photos of the reactor chamber after testing.</p>
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<p>Analysis of the lower part of the reactor chamber after reconstruction to the 3D model. Yellow arrows indicate voids and discontinuities in the material.</p>
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<p>Cross-sectional images of the element R3.</p>
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<p>Element R3: (<b>a</b>) 3D view and (<b>b</b>) wall thickness distribution.</p>
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<p>Cross-sectional images of the element R2.</p>
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<p>Element R2: (<b>a</b>) 3D view and (<b>b</b>) wall thickness distribution.</p>
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<p>Cross-sectional images of the element R1.</p>
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<p>Element R1: (<b>a</b>) 3D view and (<b>b</b>) wall thickness distribution.</p>
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<p>Thickness measurement of element R3 using the 3D scanning method.</p>
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<p>Lower part of reactor chamber (R3) shown from two views—dimensional comparison with CAD model.</p>
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<p>SEM image of the L3 surface (<b>a</b>) and schematic illustrating the formation of a pore-salt network for Ni Alloy 200 and molten salts (<b>b</b>).</p>
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11 pages, 6808 KiB  
Article
Finite Element and Experimental Analysis of Microstructural and Hardness Variations in Plasma Arc Welding of AISI 304 Stainless Steel
by Serafino Caruso, Francesco Borda, Michela Sanguedolce and Luigino Filice
J. Manuf. Mater. Process. 2024, 8(6), 299; https://doi.org/10.3390/jmmp8060299 - 23 Dec 2024
Viewed by 366
Abstract
AISI 304 is widely regarded as the most common austenitic stainless steel and is utilized in various household and industrial applications, including food handling equipment, machinery components, and heat exchangers. Its popularity stems from its excellent mechanical properties, corrosion resistance, and ease of [...] Read more.
AISI 304 is widely regarded as the most common austenitic stainless steel and is utilized in various household and industrial applications, including food handling equipment, machinery components, and heat exchangers. Its popularity stems from its excellent mechanical properties, corrosion resistance, and ease of manufacturing. Given its diverse applications, it is crucial to study the microstructural evolution and mechanical properties of the welded zone, especially considering the potential for weld decay during fusion welding. In this context, two critical thermal-dependent factors for ensuring high-quality welds are grain growth and hardness variation in the heat-affected zone (HAZ) during the welding process. This paper presents an innovative finite element (FE) model developed to analyze the grain growth and hardness reduction that occur in the HAZ during plasma arc welding (PAW) of AISI 304 steel for solid expansion tube (SET) technology. Using the commercial FE software SFTC DEFORM-3D™, a user subroutine was created that integrates a physics-based model with the Hall–Petch (H-P) equation to predict changes in grain size and hardness. This study introduces a comprehensive numerical model, encompassing the user subroutine, heat source fitting, and geometry, which accurately predicts the thermal phenomena associated with grain coarsening and hardness reduction in the HAZ during the welding of austenitic stainless steel. The results from the numerical model, including the customized user routines, show good agreement with experimental data, leading to a maximum error prediction of 10 HV in hardness, 30 µm in grain size, and 10% in HAZ extension. Full article
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<p>(<b>a</b>) Experimental set up and (<b>b</b>) sample geometry.</p>
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<p>Transverse section metallographic analysis of the welded samples.</p>
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<p>Transverse section grain size: (<b>a</b>) base metal; (<b>b</b>) HAZ; and (<b>c</b>) weld metal.</p>
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<p>(<b>a</b>) Hardness indentation matrix and (<b>b</b>) transverse section indentation measurement.</p>
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<p>Finite element modeling.</p>
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<p>(<b>a</b>) Three-dimensional conical heat source model; (<b>b</b>) measurements of the Gaussian parameters; and (<b>c</b>) DEFORM heat exchange windows for heat source modeling.</p>
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<p>(<b>a</b>) Calibration procedure for the temperature and convection coefficient of the heat exchange window; (<b>b</b>) numerical and experimental comparison of the temperature.</p>
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<p>Validation of the predicted HAZ along with the shape and size of the welding zones.</p>
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<p>Numerical predicted (<b>a</b>) grain size in mm, (<b>b</b>) hardness [HV0.1] evolution, and (<b>c</b>) comparison between measured and predicted hardness variation.</p>
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11 pages, 25539 KiB  
Article
Generation of Pre-Caldera Qixiangzhan and Syn-Caldera Millennium Rhyolites from Changbaishan Volcano by Shallow Remelting: Evidence from Zircon Hf–O Isotopes
by Haibo Zou and Jie Tong
Minerals 2024, 14(12), 1297; https://doi.org/10.3390/min14121297 - 22 Dec 2024
Viewed by 346
Abstract
The Changbaishan volcano is well known for its major caldera-forming Millennium Eruption (ME) in 946 CE (Common Era). We report Hf–O isotopes of zircon grains from pre-caldera Qixiangzhan (QXZ) and syn-caldera eruptions of the Changbaishan (Baitoushan) volcano to constrain magma chamber processes. Zircon [...] Read more.
The Changbaishan volcano is well known for its major caldera-forming Millennium Eruption (ME) in 946 CE (Common Era). We report Hf–O isotopes of zircon grains from pre-caldera Qixiangzhan (QXZ) and syn-caldera eruptions of the Changbaishan (Baitoushan) volcano to constrain magma chamber processes. Zircon grains from the pre-caldera QXZ comendite lavas have δ18O ranging from 4.46 to 5.16 (lower than mantle values) and εHf ranging from −4.47 to +4.37. Zircon grains from the syn-caldera ME1 charcoal-bearing non-welded comendite pyroclastic flow deposits have δ18O ranging from 2.25 (lower than mantle values) to 5.51 and εHf from −3.75 to +3.31. By comparison, zircon grains from the ME2 welded trachytes have δ18O ranging from 5.66 to 6.20 (higher than mantle zircon values) and εHf from −1.97 to +6.23. There are no correlations between O and Hf isotopes for all zircon grains in QXZ and ME1 comendites and ME2 trachyte. The ubiquitous occurrence of low-δ18O zircon grains in QXZ and ME1 comendites indicates shallow remelting of hydrothermally altered low-δ18O juvenile rocks. By contrast, ME2 trachyte zircons (except for two zircon grains) have normal δ18O (5.66 to 6.10) values, indicating a lack of remelting processes. Similar zircon Hf–O isotopes between pre-caldera QXZ comendites and syn-caldera ME1 comendites indicate tapping of the upper portion of a zoned magma chamber. Higher δ18O in ME2 trachyte zircons indicate tapping of the deeper portion of a zoned magma chamber free from shallow remelting. The lack of significant correlations between zircon O and Hf isotopes, and the relatively high εHf values for all Changbai zircon grains, argue against partial melting of ancient continental crust or significant contaminations by ancient crustal rocks as an origin for these felsic magmas. The QXZ and ME1 comendites were formed by shallow remelting of hydrothermally altered juvenile volcanic rocks, and ME2 trachytes were formed by evolution of mantle-derived basaltic magmas free of hydrothermal assimilations. A proto-caldera likely formed prior to the generation of QXZ lavas at 10 ka. Full article
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Graphical abstract

Graphical abstract
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<p>(<b>a</b>) Location of the Changbaishan volcano between the China –North Korea border. The 5 cm and 10 cm dashed lines represent the thickness of volcanic ash deposits [<a href="#B21-minerals-14-01297" class="html-bibr">21</a>]. (<b>b</b>) Sample locations from the Changbaishan volcano. ME2 sample 2017.7 (N 41°53′38.66″, E 128°05′47.22″) and ME1 sample 2017.9 (N 41°48 49.56, E 128°06′38.80″) are located on the south slope of the Chanbaishan volcano. 02QXZ is located on the north slope of the Changbaishan volcano. Revised after Zou et al. [<a href="#B12-minerals-14-01297" class="html-bibr">12</a>].</p>
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<p>Zircon Hf–O isotopes for pre-caldera QXZ lavas, ME1 non-welded charcoal bearing pyroclastic flow deposits, and ME2 welded trachyte pyroclastic flow deposits. Note the low δ<sup>18</sup>O for zircons from QXZ comendites and the ME1 comenditic non-welded pyroclastic flow deposit.</p>
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<p>Pre-caldera QXZ zircon CL images (<b>top</b>), transmitted light images (<b>middle</b>), and reflected light images (<b>bottom</b>). The white real numbers are oxygen isotope compositions in δ<sup>18</sup>O; the yellow real numbers are Hf isotopic compositions in ε<sub>Hf</sub>. Circles represent analysis spots. The integers (e.g., 1, 2) represent spot numbers.</p>
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<p>ME1 (sample 2017.9) zircon CL images (<b>top</b>), transmitted light images (<b>middle</b>), and reflected light images (<b>bottom</b>). The white real numbers are oxygen isotope compositions in δ<sup>18</sup>O; the yellow real numbers are Hf isotopic compositions in ε<sub>Hf</sub>. Circles represent analysis spots. The integers (e.g., 1, 2) represent spot numbers.</p>
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<p>ME2 (sample 2017.7) zircon CL images (<b>top</b>), transmitted light images (<b>middle</b>), and reflected light images (<b>bottom</b>). The white real numbers are oxygen isotope compositions in δ<sup>18</sup>O; the yellow real numbers are Hf isotopic compositions in ε<sub>Hf</sub>. Circles represent analysis spots. The integers (e.g., 1, 2) represent spot numbers.</p>
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<p>Comparison of ME zircons from this study (open square) and previous data (x, ME Cheong [<a href="#B13-minerals-14-01297" class="html-bibr">13</a>]; +, ME Liu [<a href="#B20-minerals-14-01297" class="html-bibr">20</a>]). All three studies yield similar ME zircon Hf–O isotope data.</p>
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<p>(<b>a</b>) Proto-caldera prior to 10 ka and (<b>b</b>) mature caldera in 946 CE for the Changbaishan volcano. Low-δ<sup>18</sup>O zircons formed in the shallow magma chamber rather than the deep magma chamber.</p>
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<p>Comparison of Changbaishan zircon Hf–O with Tengchong zircon Hf–O isotopes. Tengchong zircons [<a href="#B14-minerals-14-01297" class="html-bibr">14</a>,<a href="#B15-minerals-14-01297" class="html-bibr">15</a>] show negative Hf–O isotope correlations whereas Changbaishan zircons show no significant correlations.</p>
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9 pages, 4197 KiB  
Communication
Study on Properties of Additive Manufacturing Ta10W Alloy Laser-Welded Joints
by Rui Zhen, Liqun Li, Yunhao Gong, Jianfeng Gong, Yichen Huang and Shuai Chang
Materials 2024, 17(24), 6268; https://doi.org/10.3390/ma17246268 - 22 Dec 2024
Viewed by 456
Abstract
This investigation focuses on Selective Laser Melting (SLM)-fabricated thin-walled Ta10W alloy components. Given the inherent limitations of SLM in producing large-scale, complex components in a single operation, laser welding was investigated as a viable secondary processing method for component integration. The study addresses [...] Read more.
This investigation focuses on Selective Laser Melting (SLM)-fabricated thin-walled Ta10W alloy components. Given the inherent limitations of SLM in producing large-scale, complex components in a single operation, laser welding was investigated as a viable secondary processing method for component integration. The study addresses the critical issue of weldability in additively manufactured tantalum-tungsten alloys, which frequently exhibit internal defects due to process imperfections. Comprehensive analyses were conducted on weldability, microstructural evolution, texture intensity, and mechanical properties for welds oriented along both traveling and building directions. Results demonstrate that welds oriented along the traveling direction exhibit superior performance characteristics, including enhanced tensile strength, increased yield strength, improved elongation, and reduced texture intensity compared to building direction welds. Notably, grain orientation alignment between the weld zone and base material was observed consistently in both directional configurations. Full article
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<p>Welding equipment, weld morphology, and the diagram of tensile specimen size. (<b>a</b>) Welding assembly drawing. (<b>b</b>) Macro morphology of joints. (<b>c</b>) Schematic diagram of tensile specimen.</p>
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<p>Welding results under different thermal inputs: (<b>a</b>) Surface morphology of the weld seam. (<b>b</b>) Relationship between weld dimensions and heat input. (<b>c</b>)-1 Weld in traveling direction at 60 J/mm; (<b>c</b>)-2 75 J/mm; (<b>c</b>)-3 90 J/mm; (<b>c</b>)-4 108 J/mm; (<b>c</b>)-5 120 J/mm. (<b>d</b>)-1 Weld in building direction at 60 J/mm; (<b>d</b>)-2 75 J/mm; (<b>d</b>)-3 90 J/mm; (<b>d</b>)-4 108 J/mm; (<b>d</b>)-5 120 J/mm.</p>
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<p>Metallographic diagram and EBSD results: (<b>a</b>) Building direction microstructural change. (<b>b</b>) Traveling direction microstructural change. (<b>c</b>) Building direction base metal EBSD results. (<b>d</b>) Traveling direction base metal EBSD results. (<b>e</b>) Welding zone of building direction EBSD results. (<b>f</b>) Welding zone of traveling direction EBSD results.</p>
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<p>Tensile test results and fracture morphology: (<b>a</b>) Tensile specimen of the weld in the building direction. (<b>b</b>) Tensile specimen of the weld in the traveling direction. (<b>c</b>) Stress–strain curve. (<b>d</b>) Bar chart of tensile strength and elongation. (<b>e</b>) Fracture morphology of the base material. (<b>f</b>) Fracture morphology of the weld in the building direction. (<b>g</b>) Fracture morphology of the weld in the traveling direction.</p>
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9 pages, 3257 KiB  
Article
High Corrosion Resistance of Aluminum Alloy Friction Stir Welding Joints via In Situ Rolling
by Wei Wang, Xiangchen Meng, Yuming Xie, Naijie Wang, Xiaotian Ma, Jiaze Gao and Yongxian Huang
Coatings 2024, 14(12), 1604; https://doi.org/10.3390/coatings14121604 - 21 Dec 2024
Viewed by 461
Abstract
Despite the extensive use of 7xxx aluminum alloys in aerospace, intergranular corrosion is yet to be appropriately addressed. In this work, in situ rolling friction stir welding (IRFSW) was proposed to improve the corrosion resistance of joints via microstructural design. A gradient-structured layer [...] Read more.
Despite the extensive use of 7xxx aluminum alloys in aerospace, intergranular corrosion is yet to be appropriately addressed. In this work, in situ rolling friction stir welding (IRFSW) was proposed to improve the corrosion resistance of joints via microstructural design. A gradient-structured layer was successfully constructed on the surface of the joint, and the corrosion resistance was improved by in situ rolling. The intergranular corrosion depth of the IRFSW joint was reduced by 59.8% compared with conventional joints. The improved corrosion resistance was attributed to the redissolved precipitates, the disappearance of precipitate-free zones, and the discontinuous distribution of grain boundary precipitates. This study offers new insights for enhancing the corrosion resistance of aluminum alloy FSW joints. Full article
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<p>Schematic diagram of IRFSW tool.</p>
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<p>Surface morphologies of the specimens: (<b>a</b>) conventional FSW and (<b>b</b>) IRFSW.</p>
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<p>Microstructure of the cross-sections under different processes: (<b>a</b>) FSW, (<b>b</b>) IRFSW, (<b>c</b>) HAZ, and (<b>d</b>) HAZ by in situ rolling.</p>
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<p>Typical EBSD images of (<b>a</b>) HAZ by in situ rolling, (<b>b</b>) HAZ, and (<b>c</b>) deformed grain layer; grain diameter of (<b>d</b>) HAZ and (<b>e</b>) deformed grain layer.</p>
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<p>TEM images of the specimens: (<b>a</b>,<b>c</b>) bright-field and HAADF-STEM images in the HAZ; (<b>b</b>,<b>d</b>) bright-field and HAADF-STEM images in the HAZ by in situ rolling.</p>
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<p>Corrosion morphologies of specimens after intergranular corrosion: (<b>a</b>) BM, (<b>b</b>) HAZ, (<b>c</b>) WNZ of conventional FSW; (<b>d</b>–<b>f</b>) BM, HAZ, and WNZ by in situ rolling.</p>
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<p>PDP curves of microzones of the joint.</p>
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<p>Electrochemical results of microzones: (<b>a</b>) Nyquist plot, (<b>b</b>) Bode plot, (<b>c</b>) phase plot, and (<b>d</b>) corresponding equivalent circuit.</p>
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19 pages, 18317 KiB  
Article
A Study on the Impact Toughness of the Simulated Heat-Affected Zone in Multi-Layer and Multi-Pass Welds of 1000 MPa Grade Steel for Hydroelectric Applications
by Yuwei Li, Yuanbo Li and Jianxiu Chang
Metals 2024, 14(12), 1455; https://doi.org/10.3390/met14121455 - 19 Dec 2024
Viewed by 397
Abstract
The microstructure and impact toughness of a steel material subjected to multi-layer and multi-pass welding with varying secondary peak temperatures were investigated using welding thermal simulation. The detailed microstructures and fracture morphologies were examined by SEM, TEM, and EBSD. When the secondary peak [...] Read more.
The microstructure and impact toughness of a steel material subjected to multi-layer and multi-pass welding with varying secondary peak temperatures were investigated using welding thermal simulation. The detailed microstructures and fracture morphologies were examined by SEM, TEM, and EBSD. When the secondary peak temperature reaches 650 °C, the microstructure resembles that of a primary thermal cycle at 1300 °C, characterized by coarse grains and straight grain boundaries. As the temperature increases to 750 °C, chain-like structures of bulky M/A (martensite/austenite) constituents form at grain boundaries, widening them significantly. At 850 °C, grain boundaries become discontinuous, and large bulky M/A constituents disappear. At 1000 °C, smaller austenitic grains form granular bainite during cooling. However, at 1200 °C, grain coarsening occurs due to the significant increase in peak temperature, accompanied by a lath martensite structure at higher cooling rates. In terms of toughness, the steel exhibits better toughness at 850 °C and 1000 °C, with ductile fracture characteristics. Conversely, at 650 °C, 750 °C, and 1200 °C, the steel shows brittle fracture features. Microscopically, the fracture surfaces at these temperatures exhibit quasi-cleavage fracture characteristics. Notably, chain-like M/A constituents at grain boundaries significantly affect impact toughness and are the primary cause of toughness deterioration in the intercritical coarse-grained heat-affected zone. Full article
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<p>The schematic of the sample orientation.</p>
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<p>Thermal simulation curves at different secondary peak temperatures.</p>
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<p>Microstructure of the base material: (<b>a</b>) OM; (<b>b</b>) SEM; (<b>c</b>) TEM.</p>
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<p>The F<sub>m</sub> values of the test steel at −40 °C under different secondary thermal cycle peak temperatures.</p>
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<p>The impact absorption energy of the test steel at −40 °C under different secondary thermal cycle peak temperatures.</p>
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<p>The load–displacement curves of the test steel at −40 °C under different secondary thermal cycle peak temperatures.</p>
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<p>Impact fracture morphology of test steel at different peak temperatures of one thermal cycle: (<b>a</b>) 650 °C; (<b>b</b>) 750 °C; (<b>c</b>) 850 °C; (<b>d</b>) 1000 °C; (<b>e</b>) 1200 °C.</p>
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<p>OM and SEM microstructure of steel at different peak temperatures of secondary thermal cycles: (<b>a</b>,<b>a’</b>,<b>a”</b>): 650 °C; (<b>b</b>,<b>b’</b>,<b>b”</b>): 750 °C; (<b>c</b>,<b>c’</b>,<b>c”</b>): 850 °C; (<b>d</b>,<b>d’</b>,<b>d”</b>): 1000 °C; (<b>e</b>,<b>e’</b>,<b>e”</b>): 1200 °C. The grain boundaries were marked with yellow lines.</p>
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<p>TEM of the tested steel after a secondary thermal cycle with a peak temperature of 650 °C: (<b>a</b>) elongated lath bundles; (<b>b</b>) high-density dislocations; (<b>c</b>) carbides.</p>
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<p>TEM of the tested steel after a secondary thermal cycle with a peak temperature of 750 °C: (<b>a</b>) elongated lath bundles; (<b>b</b>) blocky M/A at grain boundary. The outline of M/A constituent was marked with yellow line; (<b>c</b>) carbides.</p>
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<p>TEM of the tested steel after a secondary thermal cycle with a peak temperature of 850 °C: (<b>a</b>) grain morphology; (<b>b</b>) blurred lath edges; (<b>c</b>) carbides.</p>
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<p>TEM of the tested steel after a secondary thermal cycle with a peak temperature of 1000 °C. (<b>a</b>) Lath bundles arranged in multiple orientations; (<b>b</b>) blurred lath edges; (<b>c</b>) carbides.</p>
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<p>TEM of the tested steel after a secondary thermal cycle with a peak temperature of 1200 °C: (<b>a</b>) lath morphologies; (<b>b</b>) carbides; (<b>c</b>) blurred lath edges.</p>
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<p>EBSD results (GB, KAM, IPF) at different peak temperatures: (<b>a</b>,<b>a’</b>,<b>a”</b>): 650 °C; (<b>b</b>,<b>b’</b>,<b>b”</b>): 750 °C; (<b>c</b>,<b>c’</b>,<b>c”</b>): 850 °C; (<b>d</b>,<b>d’</b>,<b>d”</b>): 1000 °C; (<b>e</b>,<b>e’</b>,<b>e”</b>): 1200 °C.</p>
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<p>EBSD results (GB, KAM, IPF) at different peak temperatures: (<b>a</b>,<b>a’</b>,<b>a”</b>): 650 °C; (<b>b</b>,<b>b’</b>,<b>b”</b>): 750 °C; (<b>c</b>,<b>c’</b>,<b>c”</b>): 850 °C; (<b>d</b>,<b>d’</b>,<b>d”</b>): 1000 °C; (<b>e</b>,<b>e’</b>,<b>e”</b>): 1200 °C.</p>
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<p>Results of parent phase grain reconstruction for the tested steel at different peak temperatures of secondary thermal cycles: (<b>a</b>) 650 °C; (<b>b</b>) 750 °C; (<b>c</b>) 850 °C; (<b>d</b>) 1000 °C; (<b>e</b>) 1200 °C.</p>
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<p>Crystallographic analysis results of the tested steel at different peak temperatures of secondary thermal cycles: (<b>a</b>) grain size; (<b>b</b>) grain boundary misorientation; (<b>c</b>) grain boundary misorientation; (<b>d</b>) local misorientations.</p>
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23 pages, 8531 KiB  
Article
Investigation of Residual Stress Variation in Sequential Butt Welding and Pocket Material Removal Machining Processes Utilizing Pre-Stress Method: A 3D Simulation Approach
by Isik Cetintav, Yilmaz Can and Nihat Akkus
Metals 2024, 14(12), 1454; https://doi.org/10.3390/met14121454 - 18 Dec 2024
Viewed by 817
Abstract
This study investigates the residual stresses arising from welding and machining processes, recognizing their adverse implications in manufacturing. Employing experimental analysis and simulation techniques, the research scrutinizes residual stress alterations resulting from sequential butt welding and subsequent machining. Utilizing MSC Marc Mentat software(version [...] Read more.
This study investigates the residual stresses arising from welding and machining processes, recognizing their adverse implications in manufacturing. Employing experimental analysis and simulation techniques, the research scrutinizes residual stress alterations resulting from sequential butt welding and subsequent machining. Utilizing MSC Marc Mentat software(version 2016), three-dimensional models are developed to simulate these processes. The finite element model from welding simulation seamlessly integrates into cutting simulations via the pre-state option. The experimental procedures involve 100 × 100 × 10 mm AISI 304 steel plates subjected to sequential welding and machining, with residual stresses measured at each stage. A comparative analysis between experimental and simulation results elucidates variations in residual stresses induced by sequential processes. The study focuses on examining the initial stress state post-welding and numerically assessing stress modifications due to milling. The results suggest minimal material removal insignificantly affects stress distribution and magnitude at the weld centerline. However, increased material removal leads to noticeable changes in through-thickness transverse stress within the weld zone, contrasting with marginal alterations in through-thickness longitudinal stress. Regions distanced from the weld seam show substantial increases in through-thickness longitudinal stress compared to marginal changes in through-thickness transverse stress. Full article
(This article belongs to the Special Issue Recent Advances in Welding Technology of Alloys and Metals)
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<p>Three different types of residual stresses at different scales. Adapted from Ref. [<a href="#B9-metals-14-01454" class="html-bibr">9</a>].</p>
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<p>Schematic representation of the main and side effect factors representing the residual stress generation processes.</p>
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<p>Representation of milling cutters used—(<b>a</b>) Cutter, (<b>b</b>) Drawing of Cutter, (<b>c</b>) Cutter-related features.</p>
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<p>Technical illustration of component, (<b>a</b>) Drawing of welded part, (<b>b</b>) After welding, (<b>c</b>) During machining, (<b>d</b>) After machining.</p>
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<p>Thermomechanical analysis procedures for welding and material removal.</p>
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<p>Representation of the (<b>a</b>) node and elements of the meshed part and (<b>b</b>) welding (red arrow) and clamp boundary condition (pink arrow).</p>
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<p>Goldak’s Double Ellipsoidal Source Flux showing Local Coordinate Dimensions.</p>
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<p>Definition of tool paths and tool geometry through CATIA.</p>
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<p>Definition of cutting tool geometry in MSC.Marc software.</p>
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<p>Flowchart of the whole procedure.</p>
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<p>Residual stress distribution in welded and machined components in a 2D simulation: Longitudinal stresses (<b>a</b>) Graphical representation of residual stress in the z-direction (longitudinal), (<b>b</b>) Distribution of σ<sub>zz</sub> post-welding and machining process.</p>
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<p>Residual stress distribution in welded and machined components in a 3D simulation: Longitudinal stresses (<b>a</b>) Graphical representation of residual stress in the z-direction (longitudinal), (<b>b</b>) Distribution of σ<sub>zz</sub> post-welding and machining process.</p>
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<p>Residual stress distribution in welded and machined components in a 3D simulation: Longitudinal stresses (<b>a</b>) Graphical representation of residual stress in the z-direction (longitudinal), (<b>b</b>) Distribution of σ<sub>zz</sub> post-welding and machining process.</p>
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<p>Residual stress distribution in welded and machined components in a 3D simulation: Transverse stresses (<b>a</b>) Graphical representation of residual stress in the x-direction (transverse), (<b>b</b>) Distribution of σ<sub>xx</sub> post-welding and machining process.</p>
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<p>Residual stress distribution in welded and machined components in a 3D simulation: Transverse stresses (<b>a</b>) Graphical representation of residual stress in the x-direction (transverse), (<b>b</b>) Distribution of σ<sub>xx</sub> post-welding and machining process.</p>
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<p>Longitudinal (z-direction) stress variations: (<b>a</b>) after welding, (<b>b</b>) normal view, and (<b>c</b>) bottom view of three different chip-removal conditions.</p>
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<p>Equivalent von Mises stress variations: (<b>a</b>) after welding, (<b>b</b>) normal view, and (<b>c</b>) bottom view of three different chip-removal conditions.</p>
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<p>Overview of stress distribution after welding and machining, (<b>a</b>) Longitudinal (z-direction) stress, (<b>b</b>) Transverse stress (x-direction), (<b>c</b>) Equivalent von Mises stress.</p>
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