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16 pages, 8305 KiB  
Article
Investigating Fracture Behavior in Titanium Aluminides: Surface Roughness as an Indicator of Fracture Mechanisms in Ti-48Al-2Cr-2Nb Alloys
by Alessia Serena Perna, Lorenzo Savio, Michele Coppola and Fabio Scherillo
Metals 2025, 15(1), 49; https://doi.org/10.3390/met15010049 - 7 Jan 2025
Viewed by 178
Abstract
Titanium aluminides, particularly the Ti-48Al-2Cr-2Nb alloy, have drawn significant attention for their potential in high-temperature aerospace and automotive applications due to their exceptional performances and reduced density compared to nickel-based superalloys. However, their intermetallic nature poses challenges such as limited room-temperature ductility and [...] Read more.
Titanium aluminides, particularly the Ti-48Al-2Cr-2Nb alloy, have drawn significant attention for their potential in high-temperature aerospace and automotive applications due to their exceptional performances and reduced density compared to nickel-based superalloys. However, their intermetallic nature poses challenges such as limited room-temperature ductility and fracture toughness, limiting their widespread application. Additive manufacturing, specifically Electron Beam Melting (EBM), has emerged as a promising method for producing complex-shaped components of titanium aluminides, overcoming challenges associated with conventional production methods. This work investigates the fracture behavior of Ti-48Al-2Cr-2Nb specimens with different microstructures, including duplex and equiaxed, under tensile and high-cycle fatigue at elevated temperatures. Fracture surfaces were analyzed to distinguish between static and dynamic fracture modes. A novel method, employing confocal microscopy acquisitions, is proposed to correlate surface roughness parameters with the causes of failure, offering new insights into the fracture mechanisms of titanium aluminides. The results reveal significant differences in roughness values between the propagation and fracture zones for all the temperatures and microstructure tested. At 650 °C, the crack propagation zone exhibits lower Sq values than the fracture zone, with the fracture zone showing more pronounced roughness, particularly for the equiaxed microstructure. However, at 760 °C, the difference in Sq values between the propagation and fracture zones becomes more pronounced, with a more substantial increase in Sq values in the fracture zone. These findings contribute to understanding fracture behavior in titanium aluminides and provide a predictive framework for assessing structural integrity based on surface characteristics. Full article
(This article belongs to the Special Issue Research on Fatigue Behavior of Additively Manufactured Materials)
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<p>SEM images of the microstructure of specimens obtained via EBM technology and subjected to HIP treatment characterized by of (<b>a</b>) equiaxed microstructure (<b>b</b>) duplex microstructure.</p>
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<p>Stereomicroscope images of the specimens subjected to tensile tests characterized by: (<b>a</b>) duplex microstructure tested at 650 °C; (<b>b</b>) duplex microstructure tested at 760 °C; (<b>c</b>) equiaxial microstructure tested at 650 °C; (<b>d</b>) equiaxed microstructure tested at 760 °C.</p>
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<p>Comparison of fracture surfaces at 500× magnification, following tensile testing for specimens characterized by: (<b>a</b>) duplex microstructure tested at 650 °C; (<b>b</b>) duplex microstructure tested at 760 °C; (<b>c</b>) equiaxed microstructure tested at 650 °C; (<b>d</b>) equiaxed microstructure tested at 760 °C.</p>
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<p>Comparison of fracture surfaces at 1500× magnification, following tensile testing for specimens characterized by: (<b>a</b>) duplex microstructure tested at 650 °C; (<b>b</b>) duplex microstructure tested at 760 °C; (<b>c</b>) equiaxed microstructure tested at 650 °C; (<b>d</b>) equiaxed microstructure tested at 760 °C.</p>
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<p>Confocal acquisition of fracture surfaces of specimens after tensile testing: (<b>a</b>) Waviness surface of a duplex specimen at 650 °C; (<b>b</b>) Roughness surface of a duplex specimen at 650 °C; (<b>c</b>) Waviness surface of a duplex specimen at 760 °C; (<b>d</b>) Roughness surface of a duplex specimen at 760 °C.</p>
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<p>Stereomicroscope images of the specimens subjected to HCF tests characterized by: (<b>a</b>) duplex microstructure tested at 650 °C; (<b>b</b>) duplex microstructure tested at 760 °C; (<b>c</b>) equiaxed microstructure tested at 650 °C; (<b>d</b>) equiaxed microstructure tested at 760 °C.</p>
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<p>Comparison of fracture surfaces at 500× magnification, following HCF tests for specimens characterized by: (<b>a</b>) duplex microstructure tested at 650 °C; (<b>b</b>) duplex microstructure tested at 760 °C; (<b>c</b>) equiaxed microstructure tested at 650 °C; (<b>d</b>) equiaxed microstructure tested at 760 °C.</p>
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<p>Comparison of fracture surfaces at 1500× magnification, following HCF tests for specimens characterized by: (<b>a</b>) duplex microstructure tested at 650 °C; (<b>b</b>) duplex microstructure tested at 760 °C; (<b>c</b>) equiaxed microstructure tested at 650 °C; (<b>d</b>) equiaxed microstructure tested at 760 °C.</p>
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<p>Confocal acquisition of fracture surfaces of specimens after HCF testing: (<b>a</b>) Waviness surface of a duplex specimen at 650 °C; (<b>b</b>) Roughness surface of a duplex specimen at 650 °C; (<b>c</b>) Waviness surface of a duplex specimen at 760 °C; (<b>d</b>) Roughness surface of a duplex specimen at 760 °C.</p>
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<p>Comparison of surface roughness values (Sq) for the propagation and fracture zones of specimens tested for HCF at 650 °C and 760 °C: (<b>a</b>) specimens with a duplex microstructure; (<b>b</b>) specimens with an equiaxed microstructure.</p>
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19 pages, 9026 KiB  
Article
Fatigue Life Analysis of Titanium Torsion Spring Based on Continuous Damage Mechanics
by Dehai Meng, Changming Zhang, Fan Yang and Feixiang Duan
Materials 2025, 18(2), 221; https://doi.org/10.3390/ma18020221 - 7 Jan 2025
Viewed by 185
Abstract
In this study, a titanium alloy torsional spring used in aviation was taken as the research subject. Aiming at the fatigue life prediction problem of this spring, the life analysis of the titanium alloy torsional spring was performed using a customized UMAT subroutine [...] Read more.
In this study, a titanium alloy torsional spring used in aviation was taken as the research subject. Aiming at the fatigue life prediction problem of this spring, the life analysis of the titanium alloy torsional spring was performed using a customized UMAT subroutine based on the theory of continuous damage mechanics. Several sets of life prediction models and tests were compared. The fatigue lives of the springs at 60, 80, 100, and 120 degrees were 45,070, 65,067, 99,677, and 181,322 cycles, respectively. Compared with other fatigue life prediction methods, the fatigue life calculated by the customized subroutine was the most consistent with the fatigue life of the titanium alloy torsion spring tests. The average relative error between the measured experimental life value and the predicted value was 2.04%, which is less than 5%, meeting engineering measurement requirements. The effectiveness and applicability of the proposed model and method were verified, and the time and economic cost caused by excessively long experimental cycles were reduced. This helps improve the accuracy of fatigue life prediction for this titanium alloy torsional spring and provides analysis support for subsequent structural optimization and improvement. Full article
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<p>Torsion spring structure diagram.</p>
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<p>Schematic diagram of load and boundary conditions.</p>
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<p>Maximum stress of torsion spring.</p>
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<p>Fatigue strength test bench of titanium alloy torsion spring.</p>
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<p>Spring break test stand stopped rotating.</p>
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<p>Representative volume element.</p>
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<p>Uniaxial monotonic tension stress–strain curve for Ti-6Al-4V.</p>
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<p>Test fatigue data and prediction curves of standard specimens.</p>
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<p>Simplified UMAT subroutine block diagram.</p>
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<p>Stress distribution of the section at the maximum stress.</p>
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<p>Experimental fracture torsion spring.</p>
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<p>(<b>a</b>) Relationship between different torsion angles and damage values; (<b>b</b>) relationship between different torsion angles and damage rate.</p>
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<p>Comparison of test lives from different prediction models.</p>
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<p>Relative errors of different prediction models for four torsion angles.</p>
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15 pages, 27525 KiB  
Article
Microstructure Evolution and Mechanical Properties of B4C-Reinforced TC11 + xFe Composites Fabricated by HIP
by Shenwei Qian, Nan Wang, Feng Chen, Yangyang Sun, Jiong Zhao, Hui Chang, Liang Feng and Lian Zhou
Metals 2025, 15(1), 37; https://doi.org/10.3390/met15010037 - 3 Jan 2025
Viewed by 403
Abstract
The present study involved (TiB + TiC)/TC11 (Ti-6.5Al-3.5Mo-1.2Zr-0.3Si) + xFe titanium matrix composites (TMCs) reinforced by in situ TiB whiskers and TiC particles fabricated by hot isostatic pressing. Microstructure observation reveals a substantial distribution of in situ reinforcements, which form a network-reinforced structure [...] Read more.
The present study involved (TiB + TiC)/TC11 (Ti-6.5Al-3.5Mo-1.2Zr-0.3Si) + xFe titanium matrix composites (TMCs) reinforced by in situ TiB whiskers and TiC particles fabricated by hot isostatic pressing. Microstructure observation reveals a substantial distribution of in situ reinforcements, which form a network-reinforced structure at the prior particle boundaries of the TC11 matrix. The micro–nanoscale TiB whiskers and TiC particles within and surrounding this network serve as effective dislocation pinning. The enhancement of mechanical properties can be attributed to load-bearing strengthening, fine-grain strengthening, and dislocation strengthening. The hardness and compressive strengths were investigated through mechanical properties testing. The hardness increased by 19.4% (2 wt% B4C-reinforced composites) compared with TC11 alloy. However, the addition of 2 wt% Fe at the same B4C level (2 wt% B4C + 2 wt% Fe-reinforced composites) resulted in a significant increase in hardness by 37.5% and 15.2% in compressive strengths of TMC and can be attributed to the solid solution strengthening effect and higher dislocation density provided by the addition of Fe. In addition, the optimal overall properties can be achieved by strictly regulating the addition ratio of 2 wt% Fe and 1 wt% B4C, allowing for a compressive strength of 2301 MPa while still maintaining a compressive strain of 24.6%. Full article
(This article belongs to the Special Issue Design, Processing and Characterization of Metals and Alloys)
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<p>Flow chart of (TiB + TiC)/TC11 composites fabricated by hot isostatic pressing.</p>
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<p>Sample preparation for SEM, XRD, EBSD, and room-temperature compression tests.</p>
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<p>Microstructure characteristics of powders: (<b>a</b>) TC11 powder; (<b>b</b>) Fe powder; (<b>c</b>) B<sub>4</sub>C powder; (<b>d</b>) mixed powder and amplified single particle.</p>
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<p>(<b>a</b>) XRD patterns of TC11 and TMCs; (<b>b</b>) amplified spectrum marked as A; (<b>c</b>) amplified spectrum marked as B.</p>
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<p>SEM micrographs of TC11 alloy and B<sub>4</sub>C-reinforced TC11 composites: (<b>a</b>) TC11; (<b>b</b>) TMC1; (<b>c</b>) TMC2.</p>
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<p>SEM micrographs of B<sub>4</sub>C-reinforced TC11 composites with Fe addition: (<b>a</b>) TMC3-F; (<b>b</b>) TMC4-F; (<b>c</b>) TMC5-F; (<b>d</b>) TMC6-F.</p>
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<p>EDS results of TC11-2wt% B<sub>4</sub>C with added 2wt% Fe (TMC5-F): (<b>a</b>) EDS mapping of TMC5-F; (<b>b</b>) EDS point scanning of (TiB + TiC)-rich region in TMC5-F.</p>
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<p>EBSD characterization of TMC1 (<b>a</b>–<b>c</b>) and TMC5-F (<b>d</b>–<b>f</b>): (<b>a</b>,<b>d</b>) BC maps; (<b>b</b>,<b>e</b>) IPB maps; (<b>c</b>,<b>f</b>) phase maps.</p>
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<p>Pole figure of α phase: (<b>a</b>) TMC1; (<b>b</b>) TMC5−F.</p>
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<p>The KAM distribution, along with the mean GND density computed from the KAM maps: (<b>a</b>) TMC1; (<b>b</b>) TMC5−F.</p>
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<p>(<b>a</b>) The stress–strain curves of 2 wt% and 5 wt% reinforced TC11 composites; (<b>b</b>) The stress–strain curves of 2 wt% B<sub>4</sub>C-reinforced TC11 composites with the addition of 0, 1, 2 wt% Fe; (<b>c</b>) The stress–strain curves of 1, 2, 3 wt% B<sub>4</sub>C-reinforced TC11 composites with the addition of 2 wt% Fe; (<b>d</b>) The stress–strain curves of TMCs and TC11 alloy formed by HIP; (<b>e</b>) The Vickers hardness comparison of TMCs and TC11 alloy.</p>
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<p>Schematic diagram of microstructural evolution of the HIPed (TiB + TiC)/TC11.</p>
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<p>Schematic diagram of microstructural evolution of the HIPed (TiB + TiC)/TC11 + xFe.</p>
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19 pages, 8417 KiB  
Article
Effect of Nb and Si Content on Phase Stability, Microstructure and Mechanical Properties of Sintered Ti–Nb–Si Alloys
by Derek Manoel Luup Carvalho, Deivison Daros Paim, Isadora Schramm Deschamps, Claudio Aguilar, Aloísio Nelmo Klein, Francisco Cavilha Neto, Guilherme Oliveira Neves and Cristiano Binder
Metals 2025, 15(1), 34; https://doi.org/10.3390/met15010034 - 3 Jan 2025
Viewed by 294
Abstract
The development of beta titanium alloys with biocompatible elements to replace Al and V is a subject of significant interest in the biomedical industry. This approach aims to enhance biocompatibility and mitigate potential cytotoxic effects associated with traditional alloying elements. In this work, [...] Read more.
The development of beta titanium alloys with biocompatible elements to replace Al and V is a subject of significant interest in the biomedical industry. This approach aims to enhance biocompatibility and mitigate potential cytotoxic effects associated with traditional alloying elements. In this work, Ti–xNb–ySi alloys were produced using powder metallurgy, with x of 35, 40, and 45 wt.%, and y of 0.10, 0.35, and 0.60% wt.%, using a 32 experimental design. Milling was used to mix and disperse the powders, followed by cold pressing, sintering, and heat treatment. Nb was the main element used to stabilize the β phase, and Si was used to form Si precipitates, although Si also exhibits a β-stabilizing effect. It was found that an increase from 0.10 to 0.35 wt.% of Si improved relative density, with no benefits observed at 0.60 wt.% Si. Electron microscopy showed the presence of β phase grains, and grains with β + α intragranular structures and precipitates. Increasing Nb content resulted in a decrease in ultimate tensile strength while increasing Si content from 0.10% to 0.35 wt.% exhibited the opposite effect. Full article
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<p>Dimensions of pressed samples (<b>top</b>), and after sintering and heat treatment (<b>bottom</b>).</p>
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<p>Equilibrium diagram for the alloy with 0.35 wt.% Si, calculated with the Thermo-Calc<sup>®</sup> software, demonstrating the predicted phase formation. The model predicts the homogenization of the β phase at 1200 °C for the studied Nb concentration, the formation of Ti<sub>3</sub>Si precipitates around 850 °C, and a predominantly α phase at room temperature.</p>
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<p>Scanning electron microscopy (SEM) analysis of the powders was conducted in secondary electron (SE) mode. The SEM images of Ti (<b>a</b>), Nb (<b>b</b>), and Si (<b>c</b>) prior to milling reveal an angular morphology. By contrast, image (<b>d</b>) illustrates the powders after milling, displaying an irregular and granular morphology. Also, image (<b>d</b>) is in the backscattered electron (BSE) mode, revealing Ti (darker color) and Nb (lighter color) particles.</p>
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<p>X-ray diffraction pattern of Ti0.60Si45Nb after grinding. It shows the typical XRD pattern observed after grinding, with mainly Ti and Nb peaks observed, and a small peak of Si.</p>
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<p>XRD analysis of Ti0.10Si35Nb and Ti0.10Si45Nb (<b>a</b>) and Ti0.60Si35Nb and Ti0.60Si45Nb (<b>b</b>) samples.</p>
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<p>SEM images of samples with 45 wt.% Nb. Image (<b>a</b>) contains 0.10 wt.% Si and image (<b>b</b>) contains 0.60.</p>
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<p>Energy-dispersive X-ray spectroscopy (EDS) analysis of the Ti–0.60Si–45Nb sample after sintering and heat treatment revealed distinct compositional variations between the β and α regions. The β regions were enriched in niobium (Nb), while the α lamellae exhibited a higher concentration of titanium (Ti). The analyzed region is indicated by a yellow line, as shown in the top section, which crosses both β and α regions, providing a detailed compositional profile across the phases.</p>
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<p>Contour plot of relative density as % of theoretical values (<b>a</b>), where colors closer to red indicate higher relative density and colors closer to green indicate lower relative density. SEM image of typical surface after sintering and heat treatment showing the porous structure (<b>b</b>). Plots of marginal means with confidence limits of 95%, for a 3<sup>2</sup> design with 81 runs, showing the effect of each alloying element on relative density, Si (<b>c</b>), and Nb (<b>d</b>).</p>
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<p>The stress–strain curve of the Ti0.35Si40Nb alloy illustrates the typical behavior observed across all test runs. A change in the curve’s inclination occurs due to the removal of the clip-on extensometer during the test.</p>
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<p>Results after tensile strength tests. Contour plot of elastic modulus (GPa) at 0.2% Ɛ as a function of Nb and Si (<b>a</b>) with the darker green showing lower E, and plots of marginal means for Si (<b>b</b>) and Nb (<b>c</b>) showing the effect of each alloying element, with confidence limits of 95%. Contour plot for ultimate tensile strength (MPa) as a function of Nb and Si (<b>d</b>) with the darker red showing higher UTS, and plots of marginal means for Si (<b>e</b>) and Nb (<b>f</b>), with confidence limits of 95%.</p>
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<p>Fracture SEM after tensile strength test for Ti0.10Si35Nb (<b>a</b>) and Ti0.60Si45Nb (<b>b</b>). Plots of marginal means for elongation for Si (<b>c</b>) and Nb (<b>d</b>) showing the individual effect of each alloying element on elongation, with confidence limits of 95%.</p>
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14 pages, 1169 KiB  
Article
Antimicrobial Efficacy of Five Wound Irrigation Solutions in the Periprosthetic Joint Infection Microenvironment In Vitro and Ex Vivo
by Anja L. Honegger, Tiziano A. Schweizer, Yvonne Achermann and Philipp P. Bosshard
Antibiotics 2025, 14(1), 25; https://doi.org/10.3390/antibiotics14010025 - 3 Jan 2025
Viewed by 320
Abstract
Background/Objectives: Periprosthetic joint infections (PJI) are difficult to treat due to biofilm formation on implant surfaces and the surrounding tissue, often requiring removal or exchange of prostheses along with long-lasting antibiotic treatment. Antiseptic irrigation during revision surgery might decrease bacterial biofilm load and [...] Read more.
Background/Objectives: Periprosthetic joint infections (PJI) are difficult to treat due to biofilm formation on implant surfaces and the surrounding tissue, often requiring removal or exchange of prostheses along with long-lasting antibiotic treatment. Antiseptic irrigation during revision surgery might decrease bacterial biofilm load and thereby improve treatment success. This in vitro study investigated and compared the effect of five advanced wound irrigation solutions to reduce bacterial burden in the PJI microenvironment. Methods: We treated in vitro biofilms grown on titanium alloy implant discs with clinical bacterial strains isolated from patients with PJIs, as well as abscess communities in a plasma-supplemented collagen matrix. The biofilms were exposed for 1 min to the following wound irrigation solutions: Preventia®, Prontosan®, Granudacyn®, ActiMaris® forte (‘Actimaris’), and Octenilin®. We measured the bacterial reduction of these irrigation solutions compared to Ringer–Lactate and to the strong bactericidal but not approved Betaseptic solution. Additionally, ex vivo free-floating bacteria isolated directly from clinical sonication fluids were treated in the same way, and regrowth or lack of regrowth was recorded as the outcome. Results: Irrigation solutions demonstrated variable efficacy. The mean CFU log10 reduction was as follows: Octenilin, 3.07, Preventia, 1.17, Actimaris, 1.11, Prontosan, 1.03, and Granudacyn, 0.61. For SACs, the reduction was: Actimaris, 8.27, Octenilin, 0.58, Prontosan, 0.56, Preventia, 0.35, and Granudacyn, 0.24. Conclusions: All solutions achieved complete bacterial eradication in all tested ex vivo sonication fluids, except Granudacyn, which was ineffective in 33% of the samples (2 out of 6). Advanced wound irrigation solutions have the potential to reduce bacterial burden in the PJI microenvironment during revision surgery. However, their efficacy varies depending on bacterial species, growth state, and the composition of the irrigation solution. This underscores the importance of considering these factors when developing future PJI-specific irrigation solutions. Full article
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<p>Antimicrobial efficacy of the irrigation solutions on biofilms formed on titanium alloy (TAV) discs. (<b>A</b>) Log<sub>10</sub> CFUs/mL reduction of the irrigation solutions on 6-day-old <span class="html-italic">S. aureus</span>, <span class="html-italic">S. epidermidis</span>, and <span class="html-italic">E. coli</span>, as well as 8-day-old <span class="html-italic">C. acnes</span> biofilms formed on TAV discs. Results are depicted as a relative reduction to Ringer’s lactate solution (negative control) and are the mean (±SEM) of three independent experiments performed in duplicates. Betaseptic (red) is not a wound irrigation solution but was used as a positive control with strong bactericidal activity. (<b>B</b>) Heat map indicating the relative log<sub>10</sub> CFUs/mL reductions across all irrigation solutions and bacterial species. (<b>C</b>) Heat map indicating the mean log<sub>10</sub> CFUs/mL reduction of each irrigation solution across all bacteria. Statistical analyses in A and C were performed by two-way ANOVA. Statistically significant reductions are indicated with asterisks: *, <span class="html-italic">p</span> &lt; 0.5; **, <span class="html-italic">p</span> &lt; 0.1; ****, <span class="html-italic">p</span> &lt; 0.0001.</p>
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<p>Antimicrobial efficacy of the irrigation solutions in the <span class="html-italic">S. aureus</span> abscess communities (SAC) model. (<b>A</b>) Representative image of SACs grown in collagen. (<b>B</b>) Log<sub>10</sub> CFUs/mL reduction of the irrigation solutions on 24 h old SACs formed in plasma-supplemented collagen matrices. Results are depicted as a relative reduction to Ringer’s lactate solution (negative control) and are the mean (±SEM) of three independent experiments performed in duplicates. Betaseptic (red) is not a wound irrigation solution. Statistically significant reductions are indicated with asterisks: ****, <span class="html-italic">p</span> &lt; 0.0001.</p>
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14 pages, 6317 KiB  
Article
Improving Anti-Corrosion Property of Aluminium Alloy by Fabrication MAO Coating via Mixing Titanium Potassium Oxalate into Electrolyte
by Wei Song, Yasheng Xing and Zhen Song
Molecules 2025, 30(1), 153; https://doi.org/10.3390/molecules30010153 - 2 Jan 2025
Viewed by 383
Abstract
Titanium potassium oxalate had been mixed into the electrolyte to improve the anti-corrosion property of the micro arc oxidation coating on the surface of the aluminium alloy. The surface and cross-section of the coating at different titanium potassium oxalate concentrations had been observed [...] Read more.
Titanium potassium oxalate had been mixed into the electrolyte to improve the anti-corrosion property of the micro arc oxidation coating on the surface of the aluminium alloy. The surface and cross-section of the coating at different titanium potassium oxalate concentrations had been observed by scanning electron microscopy, showing that when the titanium potassium oxalate concentration was 10 g/L, the coating compactness was better. Additionally, the element content of the coating had been studied by the energy dispersive spectrometer, and results proved that the coating consisted of Al, O, Ti, Si, and P. The Ti increased with the increase in titanium potassium oxalate concentration. The X-ray diffractometer had been employed to analyze the crystalline structure of the coating. Then, it was found that after the micro arc oxidation, an alumina oxide coating was prepared on the surface of the aluminium alloy, and the Al2O3 and TiO2 characteristic peak was observed. Furthermore, the electrochemical workstation was used to test the anti-corrosion of the coating. It was proved that when the titanium potassium oxalate concentration was 10 g/L, the open circuit voltage, corrosion current, corrosion potential, and impedance of the coating were improved, and the anti-corrosion property of the aluminium alloy had been strengthened. Full article
(This article belongs to the Section Electrochemistry)
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<p>The SEM photos of the surface coatings at titanium potassium oxalate concentration of (<b>a</b>) 0 g/L, (<b>b</b>) 5 g/L, (<b>c</b>) 8 g/L, (<b>d</b>) 10 g/L, (<b>e</b>) 12 g/L, and (<b>f</b>) 15 g/L.</p>
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<p>The Image J photos of the surface coatings at titanium potassium oxalate concentration of (<b>a</b>) 0 g/L, (<b>b</b>) 5 g/L, (<b>c</b>) 8 g/L, (<b>d</b>) 10 g/L, (<b>e</b>) 12 g/L, and (<b>f</b>) 15 g/L.</p>
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<p>The cross-section of the coatings at titanium potassium oxalate concentrations of (<b>a</b>) 0 g/L, (<b>b</b>) 5 g/L, (<b>c</b>) 8 g/L, (<b>d</b>) 10 g/L, (<b>e</b>) 12 g/L, and (<b>f</b>) 15 g/L.</p>
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<p>(<b>a</b>) The micro surface morphology of the coating, the element distribution of (<b>b</b>) Al, (<b>c</b>) O, and (<b>d</b>) Ti, and the (<b>e</b>) element content in the coating.</p>
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<p>The XRD spectrum of the coatings.</p>
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<p>The open circuit voltage of the coatings.</p>
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<p>The corrosion current and corrosion potential of the coatings.</p>
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<p>The impedance of the coatings.</p>
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<p>The equivalent circuit of the micro arc coatings.</p>
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13 pages, 4387 KiB  
Article
Electromagnetic Interference Shielding Effectiveness of Pure SiC–Ti3SiC2 Composites Fabricated by Reactive Melt Infiltration
by Mingjun Zhang, Zhijun Ma, Xueqin Pan, Yun Li, Nanlong Zhang, Jiaxiang Xue, Jianfeng Yang and Bo Wang
Materials 2025, 18(1), 157; https://doi.org/10.3390/ma18010157 - 2 Jan 2025
Viewed by 384
Abstract
Silicon carbide-based titanium silicon carbide (SiC–Ti3SiC2) composites with low free alloy content and varying Ti3SiC2 contents are fabricated by two-step reactive melt infiltration (RMI) thorough complete reactions between carbon and TiSi2 alloy in SiC-C preforms [...] Read more.
Silicon carbide-based titanium silicon carbide (SiC–Ti3SiC2) composites with low free alloy content and varying Ti3SiC2 contents are fabricated by two-step reactive melt infiltration (RMI) thorough complete reactions between carbon and TiSi2 alloy in SiC-C preforms obtained. The densities of SiC-C preform are tailored by the carbon morphology and volumetric shrinkage of slurry during the gel-casting process, and pure composites with variable Ti3SiC2 volume contents are successfully fabricated with different carbon contents of the preforms. Due to the increased Ti3SiC2 content in the obtained composites, both electrical conductivity and electromagnetic interference (EMI) shielding effectiveness improved progressively, while skin depth exhibited decreased consistently. The improvement in the EMI shielding effectiveness of the composite is due to the free electrons being bound to move in the conductive network formed by the Ti3SiC2 phase, converting electrical energy into thermal energy and reducing the energy of electromagnetic waves. Notably, at a Ti3SiC2 content of 31 vol.%, the EMI shielding effectiveness of the SiC–Ti3SiC2 composites in the X-band reached an impressive 62.1 dB, confirming that SiC–Ti3SiC2 composites can be treated as high-performance EMI shielding materials with extensive application prospects. Full article
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<p>Change of volumetric shrinkage of the SiC-C preforms obtained as in <a href="#materials-18-00157-t002" class="html-table">Table 2</a> with different CB contents in the total carbon powder in the slurries.</p>
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<p>XRD pattern for obtained SiC–Ti<sub>3</sub>SiC<sub>2</sub> composites.</p>
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<p>BSE images of polished surface of SiC–Ti<sub>3</sub>SiC<sub>2</sub> composites (<b>a</b>) ST20, (<b>b</b>) ST25, (<b>c</b>) ST33; (<b>d</b>,<b>e</b>) EDS analysis at points A, B for sample ST25.</p>
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<p>Electrical conductivity of the SiC–Ti<sub>3</sub>SiC<sub>2</sub> composites with different Ti<sub>3</sub>SiC<sub>2</sub> contents.</p>
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<p>Variation of SE<sub>T</sub> as a function of frequency for the SiC–Ti<sub>3</sub>SiC<sub>2</sub> composites with different Ti<sub>3</sub>SiC<sub>2</sub> contents in X-band.</p>
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<p>Average SE<sub>T</sub>, SE<sub>A</sub>, SE<sub>R,</sub> and SE<sub>A</sub>/SE<sub>T</sub> of SiC–Ti<sub>3</sub>SiC<sub>2</sub> composites with different Ti<sub>3</sub>SiC<sub>2</sub> contents.</p>
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<p>Skin depth of the SiC–Ti<sub>3</sub>SiC<sub>2</sub> composites with different Ti<sub>3</sub>SiC<sub>2</sub> contents as a function of frequency in X-band.</p>
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<p>Comparison of EMI shielding effectiveness of various EMI shielding materials(Ti<sub>3</sub>SiC<sub>2</sub> [<a href="#B4-materials-18-00157" class="html-bibr">4</a>], C-Mxene [<a href="#B20-materials-18-00157" class="html-bibr">20</a>], SiC<sub>f</sub>-SiC [<a href="#B21-materials-18-00157" class="html-bibr">21</a>], C-SiC [<a href="#B22-materials-18-00157" class="html-bibr">22</a>], GNPS-SiC [<a href="#B23-materials-18-00157" class="html-bibr">23</a>] and SiC-Si<sub>3</sub>N<sub>4</sub> [<a href="#B24-materials-18-00157" class="html-bibr">24</a>]) reported in the literature.</p>
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19 pages, 12807 KiB  
Article
Modification of Mechanical Properties of Ti–6Al–4V Using L-PBF for Anatomical Plates
by Soumyabrata Basak, Sang-Hun Lee, Jeong-Rim Lee, Dong-Hyun Kim, Jeong Hun Lee, Myunghwan Byun and Dong-Hyun Kim
Metals 2025, 15(1), 32; https://doi.org/10.3390/met15010032 - 2 Jan 2025
Viewed by 458
Abstract
In this research, as-built Ti–6Al–4V anatomical plates were successfully fabricated using laser powder bed fusion (LPBF). This study thoroughly examines the microstructural evolution and its role in enhancing the mechanical properties of clavicle bone plates under sub-β-transus heat treatment for medical application. Scanning [...] Read more.
In this research, as-built Ti–6Al–4V anatomical plates were successfully fabricated using laser powder bed fusion (LPBF). This study thoroughly examines the microstructural evolution and its role in enhancing the mechanical properties of clavicle bone plates under sub-β-transus heat treatment for medical application. Scanning electron microscope (SEM) images of the as-built specimens reveal a dense formation of a hard α’ hcp martensite structure, which decomposes during annealing at 650 °C and ultimately transforms into an α + β lamellar structure at 950 °C. Additionally, coarse grains resulting from recrystallization and reduced dislocation density were observed through electron backscatter diffraction (EBSD) following heat treatment. Due to these microstructural evolutions, the desired mechanical properties of as-built Ti64 parts for surgical applications were achieved. Heat treatment of the anatomical plates at 950 °C demonstrated an excellent strength–ductility synergy under tensile deformation and the highest energy absorption capability under bending deformation, indicating sufficient durability for medical implantation applications. Full article
(This article belongs to the Section Additive Manufacturing)
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<p>Schematic representation of (<b>a</b>) LPBF process for Ti–6Al–4V (SEM image of raw powder particles at inset); (<b>b</b>) Z-built clavicle bone plate, cuboid, and tensile specimen; (<b>c</b>) a 4-point bending test for clavicle bone plate.</p>
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<p>(<b>a</b>) 3D X-ray computed tomography analysis of pore characteristics within the as-built Ti-–6Al–4V cube specimen; (<b>b</b>) a bar chart showing the volume distribution and count of the interior pores.</p>
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<p>(<b>a</b>) XRD analyses of as-built and heat-treated Ti64 alloy fabricated by LPBF. (<b>b</b>,<b>c</b>) Magnified images from the marked by purple and green dashed line reveal details of peak profiling and the presence of phase constituents. β-phase peak intensity increased at 950 HT and 800 HT thermal annealing conditions.</p>
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<p>Dislocation density estimated by XRD peak analysis for as-built, 650 HT, 800 HT, and 950 HT.</p>
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<p>SEM images through build direction (<span class="html-italic">Z</span> axis) of (<b>a</b>) as-built showing α’ martensitic structures; (<b>b</b>) 650 HT showing decomposition of martensitic α’ and stable α; (<b>c</b>) 800 HT showing α + β phase; and (<b>d</b>) 950 HT showing primary α, α + β lamella, and rod-shaped β. All the conditions present the nanosized β particles.</p>
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<p>Inverse pole Figure (IPF) maps of the (<b>a</b>) as-built, (<b>b</b>) 650 HT, (<b>c</b>) 800 HT, and (<b>d</b>) 950 HT. The prior β grain boundaries are marked with black dashed lines.</p>
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<p>Kernel average misorientation (KAM) maps of the (<b>a</b>) as-built, (<b>b</b>) 650 HT, (<b>c</b>) 800 HT, and (<b>d</b>) 950 HT. The prior β grain boundaries are indicated by white dashed lines.</p>
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<p>Surface hardness maps sequentially for as-built, 650 HT, 800 HT, and 950 HT, respectively. Hardness decreased as the thermal annealing temperature increased to 950 HT condition.</p>
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<p>Engineering stress–strain curves for as-built, 650 HT, 800 HT, and 950 HT specimens.</p>
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<p>SEM fractography and magnified tensile fracture surface images for (<b>a,e</b>) as-built; (<b>b</b>,<b>f</b>) 650 HT; (<b>c</b>,<b>g</b>) 800 HT; (<b>d</b>,<b>h</b>) 950 HT. The fracture mode change from quasi-cleavage failure of as-built to ductile failure of 950 HT specimen.</p>
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<p>Plots of 4-point bending behavior (force–displacement curve) for various annealing treatment conditions: as-built, 650 HT, 800 HT, and 950 HT.</p>
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<p>4-point bending properties of clavicle bone plate under different conditions: (<b>a</b>) maximum load and rigidity; (<b>b</b>) energy absorption capacity in the elastic and plastic regions for as-built and heat-treated specimens.</p>
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<p>(<b>a</b>) Pictographs of the broken specimens after bending tests and corresponding SEM images reveal morphologies of fracture zone for (<b>b</b>–<b>b2</b>) as-built, (<b>c</b>–<b>c2</b>) 800 HT, (<b>d</b>–<b>d2</b>) 950 HT. The yellow and red arrows indicate the tearing edges and shear dimples, relatively.</p>
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14 pages, 4044 KiB  
Article
Influence of Aging Processes on the Characteristics of Power LEDs Soldered Using Composite Solder Pastes
by Krzysztof Górecki, Przemysław Ptak and Agata Skwarek
Appl. Sci. 2025, 15(1), 324; https://doi.org/10.3390/app15010324 - 31 Dec 2024
Viewed by 414
Abstract
In this paper, the issue of an aging process’s influence on power LEDs’ properties is considered. Some measured DC characteristics of these devices and their thermal and optical parameters obtained are presented after considering different values of the aging process’ duration. Components soldered [...] Read more.
In this paper, the issue of an aging process’s influence on power LEDs’ properties is considered. Some measured DC characteristics of these devices and their thermal and optical parameters obtained are presented after considering different values of the aging process’ duration. Components soldered using different metal–ceramic composite pastes, e.g., with TiO2, were tested. The tested devices and the used measurement setup are described. The measurement procedure is described in detail. The obtained measurement results are discussed. It is shown that after the aging process at elevated temperatures, worse properties were observed for the power LEDs soldered using classical SACX0307 alloy. Most of the samples soldered with reference alloy (not composite) were damaged during the test. The best properties were obtained for the samples soldered with solder paste with the addition of titanium oxide. Full article
(This article belongs to the Special Issue New Technologies for Power Electronic Converters and Inverters)
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<p>Temperature profile of convection reflow oven SMT 460C.</p>
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<p>LED mounted on MCPCB (<b>a</b>) and assembled on heat sink (<b>b</b>).</p>
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<p>Solder paste stencil for SACX0307-REF (<b>a</b>), SACX0307-nanoTiO<sub>2</sub> (<b>b</b>), and SACX0307-TiO<sub>2</sub> (<b>c</b>).</p>
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<p>Measurement setup of the LED test system.</p>
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<p>Views of lens and chips of tested LEDs after aging test.</p>
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<p>I-V characteristics of tested devices obtained at t = 0 (solid lines) and at t = 1000 h (dashed lines).</p>
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<p>Thermal resistance of LEDs during the aging process.</p>
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<p>Luminous flux changes in LEDs during the aging process.</p>
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<p>Thermal resistance changes in LEDs during the aging process.</p>
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<p>Luminous efficacy LEDs during the aging process.</p>
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<p>Optical power of LEDs during the aging process.</p>
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<p>Luminous flux of LEDs during the aging process.</p>
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<p>Waveforms of the transient thermal impedance of LEDs during the aging process.</p>
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21 pages, 11713 KiB  
Article
Superplastic Forming/Diffusion Bonding of TA15 Titanium Alloy for Manufacturing Integrated Solid/Hollow Four-Layer Grid Lightweight Structure Components
by Zheng Han, Yuhan Xing, Taiying Liu, Ning Zhang, Shaosong Jiang and Zhen Lu
Metals 2025, 15(1), 28; https://doi.org/10.3390/met15010028 - 31 Dec 2024
Viewed by 362
Abstract
In recent years, the excellent mechanical properties and lightweight characteristics of multi-layer hollow components have led to a surge in research focused on their forming processes. This growing interest has greatly advanced technological progress in aerospace and other related fields. In this paper, [...] Read more.
In recent years, the excellent mechanical properties and lightweight characteristics of multi-layer hollow components have led to a surge in research focused on their forming processes. This growing interest has greatly advanced technological progress in aerospace and other related fields. In this paper, the metal flow behavior of TA15 titanium alloy at different temperatures from 840 °C to 940 °C and different strain rates from 0.001 s−1 to 0.1 s−1 was studied. Utilizing the finite element method, this study examined the local stress concentration, total strain distribution, thickness thinning characteristics, and pressure loading control during the superplastic forming process of the component. The integrated solid/hollow four-layer grid lightweight structural parts were successfully fabricated using the superplastic forming/diffusion bonding (SPF/DB) process. The quality of the components was evaluated using X-ray and ultrasonic C-scan detection methods. The results show that the maximum elongation of the alloy is 1340% at 900 °C/0.001 s−1. When the temperature is too high, the grain size increases remarkably, and the elongation decreases. Based on the finite element simulation results, 900 °C is the best superplastic forming temperature. Under this temperature parameter, the maximum thinning rate of the core sheet is 39.7%, the SPF time is 10,000 s, the maximum thinning rate of the face sheet is 9.8%, and the SPF time is 2400 s. In addition, the solid block has a minimal effect on the thinning of the core sheet. The grid exhibits obvious stress concentration and thinning at its rounded corners, while the thickness distribution in other areas remains relatively uniform. The nondestructive testing results confirmed that the ribs of the component are fully formed, with no missing or broken ribs. The grid exhibits good geometry and high-quality diffusion bonding. The average thickness at key positions of the component is 1.84 mm, with the minimum thickness being 1.7 mm. As the size of the grid cavity decreases, the thickness of the component tends to increase gradually. The maximum error between the simulated and measured values is 4.47%, indicating good accuracy in the simulation. Additionally, the thickness distribution of the component is relatively uniform. Full article
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<p>(<b>a</b>) The original microstructure of TA15; (<b>b</b>) the shape and dimensions of the high-temperature tensile test specimen; and (<b>c</b>) the sampling position diagram of the deformation region.</p>
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<p>Three-dimensional model of the four-layer hollow grid structure. (<b>a</b>) The diagram of the spatial position relationship of the four-layer hollow grid structure; (<b>b</b>) a sectional view of the outer structure along Path<sub>1</sub>; (<b>c</b>) the shape and dimensions of the solid block; and (<b>d</b>) a sectional view of the inner structure along Path<sub>2</sub>.</p>
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<p>True stress–strain curves of TA15 alloy at different temperatures. (<b>a</b>) 840 °C; (<b>b</b>) 860 °C; (<b>c</b>) 880 °C; (<b>d</b>) 900 °C; (<b>e</b>) 920 °C; and (<b>f</b>) 940 °C.</p>
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<p>Mechanical properties of TA15 titanium alloy under different parameters. (<b>a</b>) Variation of elongation under diverse parameter conditions; (<b>b</b>) variation of tensile strength under diverse parameter conditions.</p>
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<p><math display="inline"><semantics> <mrow> <mi>ln</mi> <mi>σ</mi> <mo>−</mo> <mi>ln</mi> <mover accent="true"> <mi>ε</mi> <mo>˙</mo> </mover> </mrow> </semantics></math> relationship diagram.</p>
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<p>High temperature tensile microstructure of TA15 alloy at different temperatures. (<b>a</b>) 840 °C; (<b>b</b>) 860 °C; (<b>c</b>) 880 °C; (<b>d</b>) 900 °C; (<b>e</b>) 920 °C; and (<b>f</b>) 940 °C.</p>
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<p>SPF time and thickness distribution results of the core plate and face plate at different temperatures. (<b>a</b>) The thickness distribution of the core sheet at 880 °C; (<b>b</b>) the thickness distribution of the core sheet at 900 °C; (<b>c</b>) the thickness distribution of the core sheet at 920 °C; (<b>d</b>) the thickness distribution of the face sheet at 880 °C; (<b>e</b>) the thickness distribution of the face sheet at 900 °C; (<b>f</b>) the thickness distribution of the face sheet at 920 °C; (<b>g</b>) the effect of temperature on the minimum thickness of the core sheet and SPF time; and (<b>h</b>) the effect of temperature on the minimum thickness of the face sheet and SPF time.</p>
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<p>The results of the equivalent Mises stress and total equivalent plastic strain of the face sheet at 900 °C. (<b>a</b>) The distribution of the equivalent Mises stress; (<b>b</b>) the curve of the equivalent Mises stress at point A changing with time; (<b>c</b>) the distribution of the total equivalent plastic strain; and (<b>d</b>) the curve of the total equivalent plastic strain along path A.</p>
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<p>The results of the equivalent Mises stress and total equivalent plastic strain of the core sheet at 900 °C. (<b>a</b>) The distribution of the equivalent Mises stress; (<b>b</b>) the curve of the equivalent Mises stress at point B with time; (<b>c</b>) the distribution of the total equivalent plastic strain; and (<b>d</b>) the curve of the total equivalent plastic strain along path B.</p>
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<p>The simulation results of the thickness of the four-layer hollow grid structure in the superplastic forming process at 900 °C. (<b>a</b>) The core sheet thickness change diagram; (<b>b</b>) the A point thickness change curve; (<b>c</b>) the core sheet along the path A thickness change curve; (<b>d</b>) the face sheet thickness change diagram; (<b>e</b>) the B point thickness change curve; and (<b>f</b>) the face sheet along the path B thickness change curve.</p>
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<p>(<b>a</b>) The superplastic forming pressure loading path of the core sheet; (<b>b</b>) the superplastic forming pressure loading path of the face sheet; (<b>c</b>) the contour plots of the core sheet at various times; and (<b>d</b>) the contour plots of the face sheet at various times.</p>
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<p>The SPF/DB process of the integrated solid/hollow four-layer grid lightweight component. (<b>a</b>) Seal welding; (<b>b</b>) The pressure loading curve; (<b>c</b>) Diffusion bonding; (<b>d</b>,<b>e</b>) Front and side views of the integrated solid/hollow four-layer grid lightweight structural parts; (<b>f</b>) X-ray detection images of structural parts; (<b>g</b>) Ultrasonic C-scan test image.</p>
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<p>The mold for the SPF/DB stage. (<b>a</b>) A 3D model diagram; (<b>b</b>) a sectional view of the mold; and (<b>c</b>) photos of the mold.</p>
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<p>Thickness detection results of integrated solid/hollow four-layer grid lightweight structural parts. (<b>a</b>) Schematic diagram of cutting path; (<b>b</b>) schematic diagram of thickness measurement points; (<b>c</b>) thickness measurement curve of characteristic points; and (<b>d</b>) curve of thickness simulation value and true value of measurement points.</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 319
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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17 pages, 25948 KiB  
Article
The Evolution of the Tensile Properties of MoS2-Coated Titanium Alloy Bolts Under the Synergistic Damage of NaCl Corrosion and Preloading
by Derong Feng, Maoyang Xie, Weilin Yu, Chao Li, Raolong Guo, Yunpeng Hu, Quanyuan Ming and Qiang Wan
Materials 2025, 18(1), 123; https://doi.org/10.3390/ma18010123 - 31 Dec 2024
Viewed by 292
Abstract
MoS2 coating is a newly developed method to prevent bolt corrosion and the seizure of bolts used in equipment in sea areas. It is of great significance to investigate the evolution of the tensile properties and intact coatings for the maintenance of [...] Read more.
MoS2 coating is a newly developed method to prevent bolt corrosion and the seizure of bolts used in equipment in sea areas. It is of great significance to investigate the evolution of the tensile properties and intact coatings for the maintenance of coated bolts. To evaluate the tensile properties of MoS2-coated titanium alloy bolts, titanium alloy bolts coated with MoS2 (TC4+MoS2) and bolts treated with a composite treatment of anodizing oxidation and MoS2 coating (TC4+AO+MoS2) were corroded in salt spray tests for 4300 h. The MoS2 coating significantly enhanced the bolts’ corrosion resistance, demonstrating exceptional protective performance by only experiencing minor peeling due to oxidation-induced cracking of the coating during the extensive 4300 h salt spray test. The tensile strengths of the TC4+MoS2 and TC4+AO+MoS2 bolts both decreased as compared with the original bolts. The bolts pretreated with anodic oxidation revealed lighter coating peeling and maintained a higher tensile strength after corrosion. Therefore, it can be concluded that the coatings provided excellent corrosion resistance, leading to a minor impact on the bolts’ tensile strength and fracture behavior under the synergistic damage of sea water corrosion and preloading. Full article
(This article belongs to the Special Issue Corrosion Behavior and Mechanical Properties of Metallic Materials)
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<p>SCHATZ threaded fastener analysis system.</p>
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<p>Cross-sectional morphology of TC4+MoS<sub>2</sub> bolts: (<b>a</b>) cross-sectional morphology, (<b>b</b>) EDS mapping profiles, (<b>c</b>) EDS line scanning.</p>
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<p>Cross-sectional morphology of TC4+AO+MoS<sub>2</sub> bolts: (<b>a</b>) cross-sectional morphology, (<b>b</b>) EDS mapping profiles, (<b>c</b>) EDS line scanning.</p>
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<p>Friction curves of (<b>a</b>) TC4+MoS<sub>2</sub> bolts and (<b>b</b>) TC4+AO+MoS<sub>2</sub> bolts.</p>
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<p>Macroscopic morphology of TC4+MoS<sub>2</sub> bolts with different corrosion times.</p>
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<p>Macroscopic morphology of TC4+AO+MoS<sub>2</sub> bolts with different corrosion times.</p>
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<p>Microscopic morphology and elemental distribution of two bolts after 192 h of salt spray corrosion.</p>
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<p>Microscopic morphology and elemental distribution of two bolts after 2100 h of salt spray corrosion.</p>
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<p>Microscopic morphology and elemental distribution of two bolts after 4300 h of salt spray corrosion.</p>
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<p>Tensile curves of TC4+MoS<sub>2</sub> and TC4+AO+MoS<sub>2</sub> bolts after different times of salt spray corrosion.</p>
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<p>Tensile curves of TC4+MoS<sub>2</sub> and TC4+AO+MoS<sub>2</sub> bolts after different times of coastal hanging corrosion test.</p>
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<p>Yield strength and tensile strength of TC4+MoS<sub>2</sub> and TC4+AO+MoS<sub>2</sub> bolts after different times of salt spray corrosion and coastal hanging corrosion.</p>
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<p>Fracture morphologies of TC4+MoS<sub>2</sub> bolts: (<b>a1</b>–<b>a3</b>) after 192 h of salt spray corrosion; (<b>b1</b>–<b>b3</b>) after 2100 h of salt spray corrosion; and (<b>c1</b>–<b>c3</b>) after 4300 h of salt spray corrosion.</p>
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<p>Fracture morphologies of TC4+ac+MoS<sub>2</sub> bolts: (<b>a1</b>–<b>a3</b>) after 192 h of salt spray corrosion; (<b>b1</b>–<b>b3</b>) after 2100 h of salt spray corrosion; and (<b>c1</b>–<b>c3</b>) after 4300 h of salt spray corrosion.</p>
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21 pages, 7674 KiB  
Article
Fatigue Experiment and Failure Mechanism Analysis of Aircraft Titanium Alloy Wing–Body Connection Joint
by Xianmin Chen, Shanshan Li, Yuanbo Liang, Shuo Wang, Liang Yan and Shichang Du
Sensors 2025, 25(1), 150; https://doi.org/10.3390/s25010150 - 30 Dec 2024
Viewed by 346
Abstract
Taking the titanium alloy wing–body connection joint at the rear beam of a certain type of aircraft as the research object, this study analyzed the failure mechanism and verified the structural safety of the wing–body connection joint under actual flight loads. Firstly, this [...] Read more.
Taking the titanium alloy wing–body connection joint at the rear beam of a certain type of aircraft as the research object, this study analyzed the failure mechanism and verified the structural safety of the wing–body connection joint under actual flight loads. Firstly, this study verified the validity of the loading system and the measuring system in the test system through the pre-test, and the repeatability of the test was analyzed for error to ensure the accuracy of the experimental data. Then, the test piece was subjected to 400,000 random load tests of flight takeoffs and landings, 100,000 Class A load tests, and ground–air–ground load tests, and the test piece fractured under the ground–air–ground load tests. Lastly, the mechanism analysis and structural safety verification of the fatigue fracture of the joints were carried out by using a stereo microscope and scanning electron microscope. The results show that fretting fatigue is the main driving force for crack initiation, and the crack shows significant fatigue damage characteristics in the stable growth stage and follows Paris’ law. Entering the final fracture region, the joint mainly experienced ductile fracture, with typical plastic deformation features such as dimples and tear ridges before fracture. The fatigue crack growth behavior of the joint was quantitatively analyzed using Paris’ law, and the calculated crack growth period life was 207,374 loadings. This result proves that the crack initiation life accounts for 95.19% of the full life cycle, which is much higher than the design requirement of 400,000 landings and takeoffs, indicating that the structural design of this test piece is on the conservative side and meets the requirements of aircraft operational safety. This research is of great significance in improving the safety and reliability of aircraft structures. Full article
(This article belongs to the Special Issue Applications of Manufacturing and Measurement Sensors: 2nd Edition)
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<p>The diagram of the experimental piece configuration and the main dimensions.</p>
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<p>The diagram of the typical load spectrum.</p>
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<p>The diagram of the experimental support scheme.</p>
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<p>Schematic diagram of attaching strain gauges on pull rod.</p>
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<p>Fatigue experiment scheme and inspection interval flow chart.</p>
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<p>Symmetry analysis of strain measurement data.</p>
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<p>Repeatability analysis of strain measurement data.</p>
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<p>Diagram of the joint fatigue fracture.</p>
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<p>Diagram of crack location and macroscopic shape of the experimental piece. (<b>a</b>) a side view of the wing-fuselage connector, with a stereoscopic microscope diagram showing the fracture location on the right; (<b>b</b>) another side view of the wing-fuselage connector (the back of (<b>a</b>)), with a stereoscopic microscope diagram showing the fracture location on the right; (<b>c</b>) a top view of the wing-fuselage connector.</p>
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<p>Diagram of the direction of loading force and crack direction. (<b>a</b>) physical picture of the experimental piece; (<b>b</b>) plan view of the experimental piece.</p>
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<p>Macroscopic morphology of the fracture of the experimental piece.</p>
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<p>SEM electron micrograph of the fracture.</p>
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<p>SEM electron micrograph of the fatigue source region. (<b>a</b>) SEM electron microscope photo at 100 times; (<b>b</b>) SEM electron microscope photo at 200 times; (<b>c</b>) SEM electron microscope photo at 500 times; (<b>d</b>) SEM electron microscope photo at 1000 times.</p>
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<p>SEM electron micrograph of fatigue growth region. (<b>a</b>) SEM electron microscope photo at 50 times; (<b>b</b>) SEM electron microscope photo at 100 times; (<b>c</b>) SEM electron microscope photo at 200 times; (<b>d</b>) SEM electron microscope photo at 1000 times.</p>
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<p>SEM electron micrograph of fatigue final fracture region. (<b>a</b>) SEM electron microscope photo at 50 times; (<b>b</b>) SEM electron microscope photo at 100 times; (<b>c</b>) SEM electron microscope photo at 500 times; (<b>d</b>) SEM electron microscope photo at 1000 times.</p>
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<p>Stereo microscopic results of the fracture.</p>
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<p>Fatigue stripes at different crack lengths in the crack growth region.</p>
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<p>Relationship between crack growth rate da/dN and crack length <span class="html-italic">a</span>.</p>
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14 pages, 11102 KiB  
Article
Wear and Optical Properties of MoSi2 Nanoparticles Incorporated into Black PEO Coating on TC4 Alloy
by Hao Zhang, Jiayi Zhu, Jingpeng Xia, Shang Sun and Jiaping Han
Coatings 2025, 15(1), 21; https://doi.org/10.3390/coatings15010021 - 29 Dec 2024
Viewed by 359
Abstract
Wear resistance and optical properties are the key point for the application of titanium alloys as structural materials in the aerospace field. To enhance the wear resistance and optical properties of titanium alloys, a black plasma electrolytic oxidation (PEO) coating incorporating MoSi2 [...] Read more.
Wear resistance and optical properties are the key point for the application of titanium alloys as structural materials in the aerospace field. To enhance the wear resistance and optical properties of titanium alloys, a black plasma electrolytic oxidation (PEO) coating incorporating MoSi2 nanoparticles was fabricated on the TC4 alloy via the PEO process, with the MoSi2 nanoparticles being in situ doped into the coating. The doping of MoSi2 nano-particles can effectively reduce the pore size of the PEO layer. The nPEO coating exhibited lower surface roughness than that of the PEO layer. The surface hardness of the nPEO coating increased to 42.5 HRC, significantly enhancing the wear resistance of the PEO layer (40.7 HRC). Furthermore, the PEO coatings exhibited better optical property compared to TC alloy, and the incorporation of MoSi2 particles further improved the performance in most wavelength ranges. The infrared emissivity of the nPEO coating was 0.87, a dramatic increase from the 0.38 value of the TC4 alloy. This coating strategy effectively enhances the wear resistance and optical performance of TC4 alloy, which is critical for the surface design of titanium alloys used in aerospace applications. Full article
(This article belongs to the Special Issue Advanced Alloy Degradation and Implants, 2nd Edition)
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<p>XRD patterns of TC4 alloy and the PEO coatings after 20 min of deposition.</p>
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<p>Surface morphology and elemental mappings of the PEO coatings after 20 min of deposition. (<b>a</b>) PEO and (<b>b</b>) nPEO, the numbers was the selected region for elemental analysis.</p>
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<p>Surface characteristics of TC4 alloy and the PEO coatings given by AFM and LSCM. (<b>a</b>) TC4, (<b>b</b>) PEO coating, (<b>c</b>) nPEO coating.</p>
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<p>Porosity and average pore size of the PEO coatings after 20 min of deposition.</p>
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<p>Cross-section morphology of the PEO coatings after 20 min of deposition. (<b>a</b>) PEO coating and (<b>b</b>) nPEO coating, the numbers was the selected region for elemental analysis.</p>
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<p>Friction coefficient of the specimens and the characteristics of the wear tracks. (<b>a</b>) Friction coefficient and (<b>b</b>) the characteristics of the wear tracks.</p>
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<p>Wear depth and width of the wear tracks and the surface hardness. (<b>a</b>) Wear depth and width and (<b>b</b>) surface hardness.</p>
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<p>Three-dimensional wear tracks of TC4 alloy and the PEO coatings. (<b>a</b>,<b>b</b>) TC4 alloy, (<b>c</b>,<b>d</b>) PEO coating, and (<b>e</b>,<b>f</b>) nPEO coating.</p>
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<p>Optical properties of TC4 alloy and the PEO coatings. (<b>a</b>) Absorptivity and (<b>b</b>) emissivity.</p>
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17 pages, 12008 KiB  
Article
Analysis of Tool Wear in Finish Turning of Titanium Alloy Ti-6Al-4V Under Minimum Quantity Lubrication Conditions Observed with Recurrence Quantification Analysis
by Joanna Lisowicz, Krzysztof Krupa, Kamil Leksycki, Rafał Rusinek and Szymon Wojciechowski
Materials 2025, 18(1), 79; https://doi.org/10.3390/ma18010079 - 27 Dec 2024
Viewed by 356
Abstract
Titanium alloys, particularly Ti-6Al-4V, are widely used in many industries due to their high strength, low density, and corrosion resistance. However, machining these materials is challenging due to high strength at elevated temperatures, low thermal conductivity, and high chemical reactivity. This study investigates [...] Read more.
Titanium alloys, particularly Ti-6Al-4V, are widely used in many industries due to their high strength, low density, and corrosion resistance. However, machining these materials is challenging due to high strength at elevated temperatures, low thermal conductivity, and high chemical reactivity. This study investigates Recurrence Plot (RP) and Recurrence Quantification Analysis (RQA) to analyze tool wear during the finish turning of Ti-6Al-4V. The tests were conducted under Minimum Quantity Lubrication (MQL). Three inserts (two coated, one uncoated) were tested, and tool life was evaluated based on material removal volume. The issue of tool exploitation and process reliability is crucial, as it directly impacts machining performance. Results show that the uncoated insert outperformed the coated ones. RQA parameters indicated a stable-to-unstable transition in coated inserts but not in the uncoated insert. This suggests that recurrence analysis can monitor cutting dynamics in coated insert machining, but further research is needed for uncoated tools. This paper’s novelty lies in applying RP and RQA to diagnose tool wear in titanium alloy machining under MQL conditions, a method not previously explored in this context. Full article
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<p>Number of journal articles containing the keyword “recurrence plot”, divided into scientific fields, published from 1980 to 2024 (based on data from Lens.org).</p>
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<p>Number of journal articles using recurrence techniques for the analysis of the cutting process, divided by machined material, published from 2000 to 2024 (based on data from Lens.org).</p>
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<p>Workstation—NEF 600 lathe with MQL system.</p>
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<p>Record of cutting force components during turning on the example of cutting insert C.</p>
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<p>Mean value of cutting force components in the function of the volume of removed material for (<b>a</b>) cutting insert A, (<b>b</b>) cutting insert B, and (<b>c</b>) cutting insert C.</p>
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<p>Images of (<b>a</b>) cutting insert A, (<b>b</b>) cutting insert B, and (<b>c</b>) cutting insert C after achieving the tool wear criterion (point 3 in <a href="#materials-18-00079-f005" class="html-fig">Figure 5</a>).</p>
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<p>Volume of removed material for each cutting insert.</p>
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<p>Unthreshold recurrence plots for subsequent stages of tool wear.</p>
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<p>RQA parameters on the basis of analysis of the records of main cutting force component <span class="html-italic">F<sub>c</sub></span> for (<b>a</b>) cutting insert A, (<b>b</b>) cutting insert B, and (<b>c</b>) cutting insert C.</p>
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<p>RQA parameters on the basis of analysis of the records of main cutting force component <span class="html-italic">F<sub>p</sub></span> for (<b>a</b>) cutting insert A, (<b>b</b>) cutting insert B, and (<b>c</b>) cutting insert C.</p>
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<p>RQA parameters on the basis of analysis of the records of main cutting force component <span class="html-italic">F<sub>f</sub></span> for (<b>a</b>) cutting insert A, (<b>b</b>) cutting insert B, and (<b>c</b>) cutting insert C.</p>
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<p>Comparison of selected RQA parameters determined on the basis of the recording of passive force for stable machining (point 1 in <a href="#materials-18-00079-f005" class="html-fig">Figure 5</a>) and for the end of machining (point 3 in <a href="#materials-18-00079-f005" class="html-fig">Figure 5</a>).</p>
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<p>Comparison of selected RQA parameters determined on the basis of the recording of feed force for stable machining (point 1 in <a href="#materials-18-00079-f005" class="html-fig">Figure 5</a>) and for the end of machining (point 3 in <a href="#materials-18-00079-f005" class="html-fig">Figure 5</a>).</p>
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