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Keywords = elastic plastic deformation

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24 pages, 10207 KiB  
Article
Integrating Stamping-Induced Material Property Variations into FEM Models for Structural Performance Simulation of Sheet-Metal Components
by Burello Elia, Hamed Rezvanpour, Dario Cimolino, Francesco Capaccioli and Alberto Vergnano
Appl. Sci. 2025, 15(5), 2480; https://doi.org/10.3390/app15052480 - 25 Feb 2025
Viewed by 230
Abstract
The accurate prediction of structural performance in sheet-metal components is critical for optimizing design and ensuring reliability in engineering applications. This study emphasizes the necessity of incorporating non-uniformities induced by stamping processes, such as thickness variation and work-hardening effects, into Finite Element Method [...] Read more.
The accurate prediction of structural performance in sheet-metal components is critical for optimizing design and ensuring reliability in engineering applications. This study emphasizes the necessity of incorporating non-uniformities induced by stamping processes, such as thickness variation and work-hardening effects, into Finite Element Method (FEM) simulations. Experimental and computational analyses reveal that neglecting these variations results in significant discrepancies, particularly in displacement predictions, where deviations exceeding 50% were observed at specific points. While elastic behavior showed reasonable agreement with experimental results, plastic deformation predictions were notably less accurate due to the inherent inhomogeneities of the real work-hardening model compared to the uniform assumptions in standard FEM models. These findings underscore the need for improved methodologies in mapping stamping-induced material properties and validating simulation results. Further refinement of mapping accuracy and validation techniques is essential for enhancing the predictive capabilities of FEM simulations for complex sheet-metal components. Full article
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Figure 1

Figure 1
<p>Workflow for integrating stamping simulation with FEM analysis and experimental validation.</p>
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<p>Stamping simulation, (<b>a</b>) mesh part and (<b>b</b>) stamping direction.</p>
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<p>Stamping simulation, (<b>a</b>) thickness distribution and (<b>b</b>) strain values.</p>
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<p>Effective stress resulting from the stamping simulation.</p>
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<p>Sheet-metal: (<b>a</b>) raw sheet with engraved equidistant points and (<b>b</b>) stamped part.</p>
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<p>Comparison between (<b>a</b>) scanned and (<b>b</b>) simulated strains.</p>
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<p>Workflow for stress calculation, property mapping, and FEM model generation.</p>
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<p>Simulation setup with respect to AFNOR NF F 31-119 standard.</p>
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<p>Free body diagram showing forces and reactions.</p>
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<p>Visualization of the (<b>a</b>) standard and the (<b>b</b>) nonlinear By-Prop models.</p>
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<p>Experiment setup in the tests of the thrust of (<b>a</b>) the backrest along the x-negatives, (<b>b</b>) the backrest along the x-positive and (<b>c</b>) the armrest along the y-positives.</p>
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<p>Positioning of the (<b>a</b>) first and second, and (<b>b</b>) third dial comparators. (<b>c</b>) Measurement points.</p>
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<p>Strain of the sheet-metal seat from the (<b>a</b>) uniform and (<b>b</b>) mapped thickness models.</p>
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<p>Stresses of the sheet-metal seat from the (<b>a</b>) uniform and (<b>b</b>) mapped thickness models.</p>
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<p>Plastic strain of the sheet-metal seat from the (<b>a</b>) uniform and (<b>b</b>) mapped nonlinear material properties.</p>
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<p>Deformations of the sheet-metal from (<b>a</b>) uniform and (<b>b</b>) mapped nonlinear material properties.</p>
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<p>Stresses on the sheet-metal seat in different views from the (<b>a</b>) uniform and (<b>b</b>) mapped nonlinear material properties.</p>
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25 pages, 18928 KiB  
Article
Mechanical, Seepage, and Energy Evolution Properties of Multi-Shaped Fractured Sandstone Under Hydro-Mechanical Coupling: An Experimental Study
by Ying Zhang, Kai He, Jianming Yang, Jiliang Pan, Xun Xi, Xianhui Feng and Leiming Zhang
Minerals 2025, 15(3), 215; https://doi.org/10.3390/min15030215 - 23 Feb 2025
Viewed by 158
Abstract
Rocks with multi-shaped fractures in engineering activities like mining, underground energy storage, and hydropower construction are often exposed to environments where stress and seepage fields interact, which heightens the uncertainty of instability and failure mechanisms. This has long been a long-standing challenge in [...] Read more.
Rocks with multi-shaped fractures in engineering activities like mining, underground energy storage, and hydropower construction are often exposed to environments where stress and seepage fields interact, which heightens the uncertainty of instability and failure mechanisms. This has long been a long-standing challenge in the field of rock mechanics. Current research mainly focuses on the mechanical behavior, seepage, and energy evolution characteristics of single-fractured rocks under hydro-mechanical coupling. However, studies on the effects of multi-shaped fractures (such as T-shaped fractures, Y-shaped fractures, etc.) on these characteristics under hydro-mechanical coupling are relatively scarce. This study aims to provide new insights into this field by conducting hydro-mechanical coupling tests on multi-shaped fractured sandstones (single fractures, T-shaped fractures, Y-shaped fractures) with different inclination angles. The results show that hydro-mechanical coupling significantly reduces the peak strength, damage stress, crack initiation stress, and closure stress of fractured sandstone. The permeability jump factor (ξ) demonstrates the permeability enhancement effects of different fracture shapes. The ξ values for single fractures, T-shaped fractures, and Y-shaped fractures are all less than 2, indicating that fracture shape has a relatively minor impact on permeability enhancement. Fracture inclination and shape significantly affect the energy storage capacity of the rock mass, and the release of energy exhibits a nonlinear relationship with fracture propagation. An in-depth analysis of energy evolution characteristics under the influence of fracture shape and inclination reveals the transition pattern of the dominant role of energy competition in the progressive failure process. Microstructural analysis of fractured sandstone shows that elastic energy primarily drives fracture propagation and the elastic deformation of grains, while dissipative energy promotes particle fragmentation, grain boundary sliding, and plastic deformation, leading to severe grain breakage. The study provides important theoretical support for understanding the failure mechanisms of multi-shaped fractured sandstone under hydro-mechanical coupling. Full article
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<p>Structure observed by polarized light microscopy.</p>
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<p>XRD and XRF test results of sandstone rock.</p>
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<p>Illustration of the hydro-mechanical coupling test procedure.</p>
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<p>Sandstone sample models with prefabricated fractures.</p>
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<p>Stage division of the stress–strain curve of SF30 fractured sandstone sample.</p>
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<p>Variation of normalized characteristic stress with crack inclination: (<b>a</b>) <span class="html-italic">σ</span><sub>c</sub>/<span class="html-italic">σ</span><sub>c-w</sub>, (<b>b</b>) <span class="html-italic">σ</span><sub>cd</sub>/<span class="html-italic">σ</span><sub>cd-w</sub>, (<b>c</b>) <span class="html-italic">σ</span><sub>ci</sub>/<span class="html-italic">σ</span><sub>ci-w</sub>, and (<b>d</b>) <span class="html-italic">σ</span><sub>cc</sub>/<span class="html-italic">σ</span><sub>cc-w</sub>.</p>
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<p>Variation of elastic modulus and Poisson’s ratio of multi-shaped fractured sandstone samples.</p>
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<p>Permeability–time relationship for sandstone samples with different fracture shapes: (<b>a</b>) Single fracture, (<b>b</b>) T-shaped fracture, and (<b>c</b>) Y-shaped fracture.</p>
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<p>Relation between elastic energy and dissipation energy in sandstone samples. (For the convenience of calculation, both the red dashed line and the blue dashed line represent the elastic modulus <span class="html-italic">E</span>).</p>
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<p>Energy evolution pattern of the single-fracture sandstone sample.</p>
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<p>Energy evolution pattern of the T-shaped fracture sandstone sample.</p>
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<p>Energy evolution pattern of the Y-shaped fracture sandstone sample.</p>
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<p>Effect of crack inclination on peak energy distribution of multi-shaped fractured sandstone samples: (<b>a</b>) single fracture, (<b>b</b>) T-shaped fracture, and (<b>c</b>) Y-shaped fracture.</p>
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<p>Effects of multi-shaped fractured sandstone samples on the elastic energy ratio <span class="html-italic">α</span> and dissipation energy ratio <span class="html-italic">β</span>: (<b>a</b>) single fracture, (<b>b</b>) T-shaped fracture, and (<b>c</b>) Y-shaped fracture.</p>
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<p><span class="html-italic">γ</span> evolution curve for the single-fracture sandstone.</p>
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<p><span class="html-italic">γ</span> evolution curve for the T-shaped fracture sandstone.</p>
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<p><span class="html-italic">γ</span> evolution curve for the Y-shaped fracture sandstone.</p>
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<p>Evolution characteristics of <span class="html-italic">γ</span><sub>max</sub> under hydro-mechanical coupling.</p>
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<p>Scanning electron microscope images of single-fracture sandstone: (<b>a</b>) 50× magnification, (<b>b</b>) 100× magnification, (<b>c</b>) 200× magnification, (<b>d</b>) 500× magnification, (<b>e</b>) 1000× magnification, and (<b>f</b>) 5000× magnification.</p>
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<p>Scanning electron microscope images of T-shaped fracture sandstone: (<b>a</b>) 50× magnification, (<b>b</b>) 100× magnification, (<b>c</b>) 200× magnification, (<b>d</b>) 500× magnification, (<b>e</b>) 5000× magnification, and (<b>f</b>) 10,000× magnification.</p>
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<p>Scanning electron microscope images of Y-shaped fracture sandstone: (<b>a</b>) 50× magnification, (<b>b</b>) 100× magnification, (<b>c</b>) 200× magnification, (<b>d</b>) 500× magnification, (<b>e</b>) 1000× magnification, and (<b>f</b>) 5000× magnification.</p>
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30 pages, 11075 KiB  
Article
The Elasto-Plastic Contact Models of Cylinder-Based and Sphere-Based Fractal Rough Surfaces
by Xiaohui Yang, Bo Bai and Shimin Mao
Appl. Sci. 2025, 15(4), 1994; https://doi.org/10.3390/app15041994 - 14 Feb 2025
Viewed by 277
Abstract
The elasto-plastic contact models of cylinder-based and sphere-based fractal rough surfaces are developed. In the two models, the critical contact areas of a single asperity are scale-dependent. With an increase in the contact load and contact area, a transition from elastic, elasto-plastic to [...] Read more.
The elasto-plastic contact models of cylinder-based and sphere-based fractal rough surfaces are developed. In the two models, the critical contact areas of a single asperity are scale-dependent. With an increase in the contact load and contact area, a transition from elastic, elasto-plastic to full plastic deformation takes place in this order. The truncated asperity size distribution functions of different frequency indexes in different contact zones are deduced. The relations between the total real contact area and total contact load for cylinder-based and sphere-based fractal rough surfaces are obtained. The pressure distributions in the contact zone are obtained. The results of elasto-plastic contact models show that the mechanical property of cylinder-based and sphere-based fractal rough surfaces depends on the range of the frequency index of asperities. When the first six frequency indexes are smaller than the elastic critical frequency index, the cylinder-based and sphere-based fractal rough surfaces approximately appear to have an elastic property in the complete contact process. When the minimum frequency index is greater than the elastic critical frequency index, elastic deformation first takes place in the rough surfaces. Then, elasto-plastic deformation takes place with an increase in the total contact load. In elastic deformation, the ratios of the peak pressures of present fractal models to those of Hertzian models are constant for a given range of frequency indexes. In inelastic deformation, the ratios of the peak pressures are inversely proportional to the total contact load. Full article
(This article belongs to the Special Issue Research on Friction and Lubrication: Surfaces, Bearings and Gears)
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<p>Simulated three-dimensional cylinder-based fractal rough surface (<math display="inline"><semantics> <mrow> <mi>M</mi> <mo>=</mo> <mn>1</mn> </mrow> </semantics></math>, <math display="inline"><semantics> <mrow> <mi>G</mi> <mo>=</mo> <mn>2.1</mn> <mo>×</mo> <msup> <mrow> <mn>10</mn> </mrow> <mrow> <mo>−</mo> <mn>3</mn> </mrow> </msup> </mrow> </semantics></math>, <math display="inline"><semantics> <mrow> <mi>D</mi> <mo>=</mo> <mn>2.5</mn> </mrow> </semantics></math>). (<b>a</b>) The three-dimensional cylinder-based fractal surface; (<b>b</b>) the section profile of the cylinder-based fractal surface.</p>
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<p>The cylinder-based fractal rough surface profile contacts with a rigid flat surface.</p>
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<p>Simulated three-dimensional sphere-based fractal rough surface (<math display="inline"><semantics> <mrow> <mi>D</mi> <mo>=</mo> <mn>2.2</mn> </mrow> </semantics></math>, <math display="inline"><semantics> <mrow> <mi>G</mi> <mo>=</mo> <mn>2.4</mn> <mo>×</mo> <msup> <mrow> <mn>10</mn> </mrow> <mrow> <mo>−</mo> <mn>3</mn> </mrow> </msup> </mrow> </semantics></math>, <math display="inline"><semantics> <mrow> <mi>U</mi> <mo>=</mo> <mi>S</mi> <mo>=</mo> <mn>1</mn> </mrow> </semantics></math>). (<b>a</b>) The three-dimensional sphere-based rough surface; (<b>b</b>) the section of the sphere-based rough surface.</p>
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<p>Equivalent sphere-based fractal rough surface contacts with a rigid flat surface.</p>
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<p>A single asperity contact model with a rigid flat surface.</p>
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<p>The contact zone between the cylinder-based fractal rough surface and a rigid flat surface.</p>
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<p>The contact zone between the sphere-based fractal rough surface and a rigid flat surface. (<b>a</b>) The contact zone. (<b>b</b>) The apparent area of each subdivision.</p>
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<p>The relations between the total real contact area and numbers of subdivisions under different interferences. (<b>a</b>) The cylinder-based fractal rough surface; (<b>b</b>) the sphere-based fractal rough surface.</p>
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<p>Comparison between the present fractal model and the Hertzian model on the half contact width under different interferences. (<b>a</b>) The cylinder-based fractal rough surface; (<b>b</b>) the sphere-based fractal rough surface.</p>
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<p>The ratios of contact areas of asperities for which the frequency index is <span class="html-italic">n</span> to the total real contact area. (<b>a</b>) The cylinder-based fractal rough surface with <span class="html-italic">ρ</span> = 0.1 m; (<b>b</b>) the sphere-based fractal rough surface with <span class="html-italic">r</span> = 0.1 m.</p>
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<p>The relation between the real contact area and the total contact load. (<b>a</b>) The cylinder-based fractal rough surface with <span class="html-italic">ρ</span> = 0.1 m; (<b>b</b>) the sphere-based fractal rough surface with <span class="html-italic">r</span> = 0.1 m.</p>
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<p>Comparison between the present fractal model and the Hertzian model based on pressure distributions. (<b>a</b>) The cylinder-based fractal rough surface; (<b>b</b>) the cylinder-based fractal rough surface; (<b>c</b>) the sphere-based fractal rough surface; (<b>d</b>) the sphere-based fractal rough surface.</p>
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<p>The relation between the real contact area and the total contact load. (<b>a</b>) The cylinder-based rough surface with <span class="html-italic">ρ</span> = 0.1 m; (<b>b</b>) the sphere-based rough surface with <span class="html-italic">r</span> = 0.1 m.</p>
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<p>Comparison between the present fractal model and the Hertzian model based on pressure distributions for the cylinder-based fractal rough surface. (<b>a</b>) The pressure distribution with <math display="inline"><semantics> <mrow> <mi>ω</mi> <mo>=</mo> <mn>0.3</mn> <msub> <mi>δ</mi> <mrow> <msub> <mi>n</mi> <mrow> <mi>min</mi> </mrow> </msub> </mrow> </msub> </mrow> </semantics></math>; (<b>b</b>) the pressure distribution with <math display="inline"><semantics> <mrow> <mi>ω</mi> <mo>=</mo> <mn>0.9</mn> <msub> <mi>δ</mi> <mrow> <msub> <mi>n</mi> <mrow> <mi>min</mi> </mrow> </msub> </mrow> </msub> </mrow> </semantics></math>; (<b>c</b>) the ratio of the peak pressure of the present fractal model to the peak pressure of the Hertzian model.</p>
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<p>Comparison between the present fractal model and the Hertzian model based on pressure distributions for the sphere-based fractal rough surface. (<b>a</b>) The pressure distribution with <math display="inline"><semantics> <mrow> <mi>ω</mi> <mo>=</mo> <mn>0.3</mn> <msub> <mi>δ</mi> <mrow> <msub> <mi>n</mi> <mrow> <mi>min</mi> </mrow> </msub> </mrow> </msub> </mrow> </semantics></math>; (<b>b</b>) the pressure distribution with <math display="inline"><semantics> <mrow> <mi>ω</mi> <mo>=</mo> <mn>0.9</mn> <msub> <mi>δ</mi> <mrow> <msub> <mi>n</mi> <mrow> <mi>min</mi> </mrow> </msub> </mrow> </msub> </mrow> </semantics></math>; (<b>c</b>) the ratio of the peak pressure of present fractal model to the peak pressure of the Hertzian model.</p>
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<p>The relation between the real contact area and the total contact load. (<b>a</b>) The cylinder-based fractal rough surface with <span class="html-italic">ρ</span> = 0.1 m; (<b>b</b>) the sphere-based fractal rough surface with <span class="html-italic">r</span> = 0.1 m.</p>
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<p>Comparison between the present fractal model and the elastic foundation model based on pressure distributions for the cylinder-based fractal rough surface. (<b>a</b>) The pressure distribution with <math display="inline"><semantics> <mrow> <mi>ω</mi> <mo>=</mo> <mn>0.9</mn> <msub> <mi>δ</mi> <mrow> <msub> <mi>n</mi> <mrow> <mi>min</mi> </mrow> </msub> </mrow> </msub> </mrow> </semantics></math>; (<b>b</b>) the ratio of the peak pressure.</p>
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<p>Comparison between the present fractal model and the Hertzian model based on pressure distributions for the sphere-based fractal rough surface. (<b>a</b>) The pressure distribution with <math display="inline"><semantics> <mrow> <mi>ω</mi> <mo>=</mo> <mn>0.9</mn> <msub> <mi>δ</mi> <mrow> <msub> <mi>n</mi> <mrow> <mi>min</mi> </mrow> </msub> </mrow> </msub> </mrow> </semantics></math>; (<b>b</b>) the ratio of the peak pressure.</p>
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<p>The relation between the peak pressure and range of frequency indexes of asperities. (<b>a</b>) The cylinder-based fractal rough surface; (<b>b</b>) the sphere-based fractal rough surface.</p>
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20 pages, 7217 KiB  
Article
The Field Monitoring and Numerical Simulation of Spatiotemporal Effects During Deep Excavation in Mucky Soft Soil: A Case Study
by Qiang Wu, Jianxiu Wang, Yanxia Long, Xuezeng Liu, Guanhong Long, Shuang Ding, Li Zhou, Huboqiang Li and Muhammad Akmal Hakim bin Hishammuddin
Appl. Sci. 2025, 15(4), 1992; https://doi.org/10.3390/app15041992 - 14 Feb 2025
Viewed by 384
Abstract
The issue of geotechnical hazards induced by excavation in soft soil areas has become increasingly prominent. However, the retaining structure and surface settlement deformation induced by the creep of soft soil and spatial effect of the excavation sequence are not fully considered where [...] Read more.
The issue of geotechnical hazards induced by excavation in soft soil areas has become increasingly prominent. However, the retaining structure and surface settlement deformation induced by the creep of soft soil and spatial effect of the excavation sequence are not fully considered where only elastic–plastic deformation is used in design. To understand the spatiotemporal effects of excavation-induced deformation in soft soil pits, a case study was performed with the Huaxi Park Station of the Suzhou Metro Line S1, Jiangsu Province, China, as an example. Field monitoring was conducted, and a three-dimensional numerical model was developed, taking into account the creep characteristics of mucky clay and spatiotemporal response of retaining structures induced by excavations. The spatiotemporal effects in retaining structures and ground settlement during excavation processes were analyzed. The results show that as the excavation depth increased, the horizontal displacement of the diaphragm walls increased linearly and tended to exhibit abrupt changes when approaching the bottom of the pit. The maximum horizontal displacement of the wall at the west end well was close to 70 mm, and the maximum displacement of the wall at the standard section reached approximately 80 mm. The ground settlement on both pit sides showed a “trough” distribution pattern, peaking at about 12 m from the pit edge, with a settlement rate of −1.9 mm/m per meter of excavation depth. The excavation process directly led to the lateral deformation of the diaphragm walls, resulting in ground settlement, which prominently reflected the time-dependent deformation characteristics of mucky soft soil during the excavation process. These findings provide critical insights for similar deep excavation projects in mucky soft soil, particularly regarding excavation-induced deformations, by providing guidance on design standards and monitoring strategies for similar geological conditions. Full article
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<p>Soil profile along the standard section of Huaxi Park Station.</p>
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<p>Layout of monitoring points of Huaxi Park Station foundation pit. (CX is the monitoring point number of diaphragm wall horizontal displacement; DB is the monitoring point number of surface subsidence).</p>
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<p>A 3D numerical model of the Huaxi Park Station foundation pit.</p>
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<p>Structure of CVISC model.</p>
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<p>Variation rule of maximum horizontal displacement of enclosure wall: (<b>a</b>) maximum horizontal displacement of west end well wall; (<b>b</b>) maximum horizontal displacement of standard section wall. RW means retaining wall.</p>
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<p>Surface settlement outside the pit: (<b>a</b>) surface settlement outside the west end head pit; (<b>b</b>) surface settlement outside the standard section pit. D is the distance from the pit. Notes−excavation step sequence: 1−arrangement of the first steel support; 2−arrangement of the second steel support; 3−arrangement of the third steel support; 4−arrangement of the fourth steel support; 5−excavation completed.</p>
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<p>The relationship between the surface settlement outside the pit and the maximum horizontal displacement of the retaining wall: (<b>a</b>) the west end well; (<b>b</b>) the standard section.</p>
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<p>Pore water pressure contour with different construction steps: (<b>a</b>) before excavation; (<b>b</b>) excavation of third floors; (<b>c</b>) final excavation.</p>
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<p>Pore water pressure contour with different construction steps: (<b>a</b>) before excavation; (<b>b</b>) excavation of third floors; (<b>c</b>) final excavation.</p>
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<p>Horizontal displacement of diaphragm wall of Huaxi Park Station foundation pit.</p>
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<p>Horizontal displacement of diaphragm wall (CX5): (<b>a</b>) horizontal displacement versus depth curve; (<b>b</b>) variation curve of maximum horizontal displacement with construction sequence. (Construction sequence: 1—arrangement of the first steel support; 2—arrangement of the second steel support; 3—arrangement of the third steel support; 4—arrangement of the fourth steel support; 5—arrangement of the fifth steel support; 6—completion of the footing pouring.)</p>
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<p>Vertical displacement after the completion of excavation: (<b>a</b>) surface settlement outside the pit; (<b>b</b>) vertical displacement.</p>
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<p>Surface settlement outside the pit after the completion of excavation: (<b>a</b>) DB4; (<b>b</b>) DB6; (<b>c</b>) DB1.</p>
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<p>Change in the axial force of the first concrete support.</p>
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<p>Change in the axial force of the second to fifth supports.</p>
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24 pages, 12975 KiB  
Article
Study on the Law of Mine Pressure Manifestation in Three-Soft Coal Seam Isolated Working Face
by Hui Liu, Jiarui Sun, Tao Yang, Jie Zhang, Dong Liu, Haifei Lin, Jiayue Deng and Yiming Zhang
Appl. Sci. 2025, 15(4), 1943; https://doi.org/10.3390/app15041943 - 13 Feb 2025
Viewed by 300
Abstract
The isolated working face is significantly impacted by the adjacent goaf and the mining activities of the working face itself, causing the overlying rock layers above the working face to exhibit far more intense activity compared to an ordinary working face. The stress [...] Read more.
The isolated working face is significantly impacted by the adjacent goaf and the mining activities of the working face itself, causing the overlying rock layers above the working face to exhibit far more intense activity compared to an ordinary working face. The stress levels are high, and the surrounding rock suffers severe damage, posing serious challenges to the safe and efficient extraction of the working face. Improving the service life of the retreating roadway in an isolated working face is a pressing technical issue that coal mining companies must address. Focusing on the characteristics of the strata and mining conditions of the 8213 isolated working face in the Yanjiahe Coal Mine, which features a three-soft coal seam, a combination of field investigation, theoretical analysis, on-site monitoring, and numerical simulation methods was employed. This approach aimed to analyze the fundamental laws of mine pressure behavior in the three-soft coal seam isolated working face as well as the deformation and failure mechanisms of the surrounding rock in the retreating roadway. Using elastic thin plate theory, it was determined that the basic roof periodic fracture step of the 8213 isolated face in the Yanjiahe Coal Mine is approximately 23 m. Field mine pressure monitoring on the 8213 isolated working face revealed that during non-periodic pressure events, the support resistance of the working face generally fluctuated stably below the rated working resistance. When the basic roof collapsed, the average working resistance of the support showed a significant increase with periodic pressure steps ranging from 16 to 27 m and an average of 22 m. Numerical simulations were further used to analyze the changes in stress and the plastic zone of the overlying rock on the 8213 isolated working face, clarifying the mechanism by which instability in the overlying rock structure leads to incidents. This analysis provides theoretical support for the safe mining of isolated working faces. Full article
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Figure 1
<p>The loess plateau landform landscape of Yanjiahe Coal Mine.</p>
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<p>Working face plan.</p>
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<p>Type of overlying rock structure of Gudao working face.</p>
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<p>The basic roof mechanics model with two sides fixed and two sides simply supported.</p>
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<p>The basic roof mechanics model with three sides simply supported and one side fixed supported.</p>
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<p>Face pressure monitoring line layout.</p>
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<p>The curve graph of the average support resistance variation of the 8213 working face during the mining recovery stage. (<b>a</b>) upper monitoring Station 1, (<b>b</b>) middle monitoring Station 2, (<b>c</b>) lower monitoring Station 3.</p>
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<p>The curve graph of the average support resistance variation of the 8213 working face during the mining recovery stage. (<b>a</b>) upper monitoring Station 1, (<b>b</b>) middle monitoring Station 2, (<b>c</b>) lower monitoring Station 3.</p>
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<p>The overall average working resistance variation curve of the working face supports.</p>
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<p>Original computational model. (<b>a</b>) Three-dimensional model stereogram, (<b>b</b>) coal seam cross-section diagram of 3D model.</p>
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<p>Formation process of island working face. (<b>a</b>) Post-mining at face 8212, (<b>b</b>) post-mining at face 8214.</p>
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<p>Vertical stress distribution at different distances of the face advance. (<b>a</b>) Working face advance 40 m, (<b>b</b>) working face advance 80 m, (<b>c</b>) working face advance 120 m, (<b>d</b>) working face advance 160 m, (<b>e</b>) working face advance 200 m, (<b>f</b>) working face advance 240 m.</p>
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<p>Vertical stress distribution at different distances of the face advance. (<b>a</b>) Working face advance 40 m, (<b>b</b>) working face advance 80 m, (<b>c</b>) working face advance 120 m, (<b>d</b>) working face advance 160 m, (<b>e</b>) working face advance 200 m, (<b>f</b>) working face advance 240 m.</p>
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<p>Evolution law of roof plastic zone at different distance of working face advance. (<b>a</b>) Working face advance 40 m, (<b>b</b>) working face advance 80 m, (<b>c</b>) working face advance 120 m, (<b>d</b>) working face advance 160 m, (<b>e</b>) working face advance 200 m, (<b>f</b>) working face advance 240 m.</p>
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<p>Evolution law of roof plastic zone at different distance of working face advance. (<b>a</b>) Working face advance 40 m, (<b>b</b>) working face advance 80 m, (<b>c</b>) working face advance 120 m, (<b>d</b>) working face advance 160 m, (<b>e</b>) working face advance 200 m, (<b>f</b>) working face advance 240 m.</p>
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<p>Evolution law of roof plastic zone at different distance of working face advance. (<b>a</b>) Working face advance 40 m, (<b>b</b>) working face advance 80 m, (<b>c</b>) working face advance 120 m, (<b>d</b>) working face advance 160 m, (<b>e</b>) working face advance 200 m, (<b>f</b>) working face advance 240 m.</p>
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<p>Evolution law of roof plastic zone at different distance of working face advance. (<b>a</b>) Working face advance 40 m, (<b>b</b>) working face advance 80 m, (<b>c</b>) working face advance 120 m, (<b>d</b>) working face advance 160 m, (<b>e</b>) working face advance 200 m, (<b>f</b>) working face advance 240 m.</p>
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19 pages, 4589 KiB  
Article
An Efficient Numerical Model for the Evaluation of the Productivity Considering Depletion-Induced Plastic Behaviors in Weakly Consolidated Reservoirs
by Feifei Luo, Lei Zhong, Zhizhong Wang, Zixuan Li, Bolong Zhu, Xiangyun Zhao, Xuyang Guo and Jiaying Lin
Energies 2025, 18(4), 892; https://doi.org/10.3390/en18040892 - 13 Feb 2025
Viewed by 240
Abstract
Efficient and accurate modeling of rock deformation and well production in weakly consolidated reservoirs requires reliable and accurate reservoir modeling techniques. During hydrocarbon production, the reservoir pressure is dropped, and rock compaction is induced. In such depletion-induced reservoir rock deformation, both elastic and [...] Read more.
Efficient and accurate modeling of rock deformation and well production in weakly consolidated reservoirs requires reliable and accurate reservoir modeling techniques. During hydrocarbon production, the reservoir pressure is dropped, and rock compaction is induced. In such depletion-induced reservoir rock deformation, both elastic and plastic deformation can be generated. The numerical investigation of depletion-induced plasticity in shale oil reservoirs and its impact on coupled reservoir modeling helps provide insights into the optimization of horizontal well productivity. This study introduces a coupled flow and geomechanical model that considers porous media flow, elastoplastic deformation, horizontal well production, and the coupling between the flow and geomechanical processes. Simulation results are then provided along with numerical modeling parameters. Effects of relevant parameters, including depletion magnitude, rock mechanical properties, and hydraulic fracture parameters, jointly affect rock deformation, rock skeleton damage, and horizontal well productivity. Depletion-induced plasticity, stress, pressure, and subsidence are all characterized by the solution strategy. In addition, the implementation of direct and iterative solvers and the usage of full coupling and sequential coupling strategies are investigated, and the associated solver performance is quantified. It helps evaluate the numerical efficiency in the highly nonlinear numerical system. This study provides an efficient coupled flow and elastoplastic model for the simulation of depletion in weakly consolidated reservoirs. Full article
(This article belongs to the Special Issue Development of Unconventional Oil and Gas Fields: 2nd Edition)
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<p>Solution strategy.</p>
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<p>Two−dimensional distributions of (<b>a</b>) pressure, (<b>b</b>) subsidence, (<b>c</b>) plastic strain, and (<b>d</b>) Sx after 1 day of production in the wellbore.</p>
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<p>Two−dimensional distributions of (<b>a</b>) pressure, (<b>b</b>) subsidence, (<b>c</b>) plastic strain, and (<b>d</b>) Sx after 30 days of production in the wellbore.</p>
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<p>The temporal evolutions of displacement and plastic strain at three monitoring points of (0 m, 0 m), (5 m, 0 m), and (10 m, 0 m).</p>
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<p>Cumulative production in the wellbore against depletion-induced plastic strain and displacement in the reservoir.</p>
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<p>Depletion−induced subsidence after 30 days of production with a constant horizontal wellbore depletion pressure of (<b>a</b>) 4 MPa, (<b>b</b>) 3 MPa, and (<b>c</b>) 2 MPa.</p>
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<p>Depletion−induced subsidence during 30 days of production with a constant production pressure difference of 12 MPa, 11 MPa, and 10 MPa.</p>
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<p>Depletion−induced plastic strain after 30 days of production with a constant horizontal wellbore depletion pressure of (<b>a</b>) 4 MPa, (<b>b</b>) 3 MPa, and (<b>c</b>) 2 MPa.</p>
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<p>Depletion−induced Sx after 30 days of production with a constant horizontal wellbore depletion pressure of (<b>a</b>) 4 MPa, (<b>b</b>) 3 MPa, and (<b>c</b>) 2 MPa.</p>
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<p>Cumulative production and average plastic strain in the reservoir induced by horizontal well production with a constant pressure of 2 MPa, 3 MPa, and 4 MPa.</p>
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<p>Comparison of solver performance of four different cases. The four cases are the base case, the case with a finer mesh, the case with a greater depletion magnitude in the well, and the case with stiffer reservoir rocks. Compared solver performance includes the residual evaluation number, the Jacobian evaluation number, the solution evaluation number, and the linear error.</p>
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<p>Comparison between solver performance parameters for the four cases: the base case with a direct solver and full coupling, the case with a direct solver and sequential coupling, the case with an iterative solver and full coupling, and the case with an iterative solver and sequential coupling.</p>
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18 pages, 4595 KiB  
Article
Fracture Mechanism of H13 Steel During Tensile Testing Based on In Situ EBSD
by Yunling Li, Dangshen Ma, Hongxiao Chi, Shulan Zhang, Jian Zhou and Jin Cai
Metals 2025, 15(2), 182; https://doi.org/10.3390/met15020182 - 11 Feb 2025
Viewed by 329
Abstract
This paper employs in situ Electron Backscatter Diffraction (EBSD) tensile technology to thoroughly consider the evolution of microstructure, grain size, grain boundary characteristics, orientation differences, and dislocation density of H13 steel during the elastic and plastic stages of room temperature tensile testing. The [...] Read more.
This paper employs in situ Electron Backscatter Diffraction (EBSD) tensile technology to thoroughly consider the evolution of microstructure, grain size, grain boundary characteristics, orientation differences, and dislocation density of H13 steel during the elastic and plastic stages of room temperature tensile testing. The study unveils the deformation mechanisms of inclusions, carbides, and the matrix in H13 steel during the various stages, providing a comprehensive explanation for the slightly superior tensile properties of H13 steel when refined by Vacuum Induction Melting combined with Vacuum Arc Remelting (VIM + VAR) over those when refined by Electroslag Remelting (ESR). This discrepancy is primarily attributed to the differences in inclusions and carbides present in the two refining processes. The quantity and size of inclusions and carbides are closely related to material fracture. Large-sized carbides and inclusions were shown to be more likely to cause dislocation pile-ups and stress concentration. This, in turn, leads to faster crack initiation and propagation during plastic deformation. Conversely, the formation of micro-pores within these fine inclusions and the matrix is contingent on greater plastic deformation, resulting in a gradual and incremental linkage of these micro-pores to form dimples beneath the influence of slip. Full article
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<p>Illustration of in situ tensile sample: (<b>a</b>,<b>b</b>) sampling position; (<b>c</b>) sample geometry; (<b>d</b>) dimensions of sample (mm).</p>
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<p>In situ EBSD tensile test under SEM: (<b>a</b>) in situ EBSD tensile testing stage; (<b>b</b>) in situ tensile testing sample.</p>
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<p>Types of ≥2 μm carbides and inclusions in ESR and VIM + VAR steel: (<b>a</b>,<b>c</b>,<b>e</b>,<b>g</b>) Morphology and EDS of oxide, silicate, sulfide or carbides in ESR steel; (<b>b</b>,<b>d</b>,<b>f</b>,<b>h</b>) Morphology and EDS of oxide, silicate, sulfide or carbides in VIM + VAR steel.</p>
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<p>Distribution of ≥2 μm carbides and inclusions in ESR and VIM + VAR steel:(<b>a</b>) Inclusion distribution of ESR; (<b>b</b>) Carbide distribution of ESR; (<b>c</b>) Inclusion distribution of VIM + VAR; (<b>d</b>) Carbide distribution of VIM+VAR.</p>
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<p>Metallographic structure and grain size of ESR and VIM + VAR steel.:(<b>a</b>) Metallographic structure of ESR; (<b>b</b>) grain size of ESR; (<b>c</b>) Metallographic structure of VIM + VAR; (<b>d</b>) grain size of VIM + VAR.</p>
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<p>Mechanical properties of ESR and VIM + VAR steel.</p>
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<p>Macro tensile fracture morphology of ESR and VIM + VAR steel at room temperature: (<b>a</b>) macro fracture morphology of ESR; (<b>b</b>) morphology of ESR; (<b>c</b>) macro fracture morphology of VIM + VAR; (<b>d</b>) morphology of VIM + VAR.</p>
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<p>In situ tensile test of ESR sample at different stages: (<b>a</b>) in situ tensile curve; (<b>b</b>) EDS of carbide and inclusions; (<b>c</b>) morphology of elastic stage; (<b>d</b>,<b>f</b>–<b>h</b>) morphology of plastic stage; (<b>e</b>) morphology of fracture.</p>
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<p>In situ tensile fracture morphology of ESR sample: (<b>a</b>) macro fracture morphology; (<b>b</b>) radiation zone; (<b>c</b>) fiber zone; (<b>d</b>–<b>f</b>) carbides or inclusions in the fractured sample; (<b>g</b>–<b>i</b>) EDS of the carbides or inclusions in the (<b>d</b>–<b>f</b>).</p>
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<p>Grain size in elastic stage of in situ tensile test: (<b>a</b>) IPF map before tensile testing; (<b>b</b>) grain size distribution before tensile testing; (<b>c</b>) IPF map with strain 2%; (<b>d</b>) different orientation grain size changes during in situ tensile test.</p>
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<p>Evolution of grain boundary and geometrically necessary dislocation (GND) during in situ tensile test: (<b>a</b>–<b>c</b>) SEM morphology under different strain; (<b>d</b>–<b>f</b>) GND map under different strain; (<b>g</b>–<b>i</b>) grain size distribution; (<b>j</b>) grain boundary map; (<b>k</b>) GND density map.</p>
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<p>Evolution of carbides during in situ tensile test: (<b>a</b>) carbides morphology with 7% strain; (<b>b</b>,<b>c</b>) IPF map of different strains; (<b>d</b>) misorientation of different lines; (<b>e</b>,<b>f</b>) Kernel average misorientation (KAM) image of different strains.</p>
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<p>Evolution of deformation of matrix, carbides, and oxide inclusions during in situ tensile test: (<b>a</b>,<b>c</b>,<b>d</b>,<b>f</b>,<b>g</b>,<b>i</b>) SEM morphology of matrix, carbide, and inclusions under different strain; (<b>b</b>,<b>e</b>,<b>h</b>) KAM image of matrix, carbide under 7% strain.</p>
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11 pages, 3760 KiB  
Article
Analysis of Contact Noise Due to Elastic Recovery of Surface Asperities for Spherical Contact
by Bora Lee, Kyungseob Kim and Taewan Kim
Inventions 2025, 10(1), 17; https://doi.org/10.3390/inventions10010017 - 8 Feb 2025
Viewed by 339
Abstract
Contact noise, often arising from frictional vibrations in mechanical systems, significantly impacts performance and user experience. This study investigates the generation of contact noise due to the elastic recovery of surface asperities during spherical contact with rough surfaces. A numerical algorithm was developed [...] Read more.
Contact noise, often arising from frictional vibrations in mechanical systems, significantly impacts performance and user experience. This study investigates the generation of contact noise due to the elastic recovery of surface asperities during spherical contact with rough surfaces. A numerical algorithm was developed to model the noise produced by the elastic–plastic deformation of asperities, incorporating surface roughness and normal load effects. Gaussian-distributed rough surfaces with varying Ra values (0.01–5 μm) were generated to analyze the interaction between a rigid sphere and the rough surface. Contact pressure, asperity deformation, and the resulting acoustic emissions were calculated. The results indicate that, as surface roughness and applied load increase, noise levels within the audible frequency range also rise, exceeding 70 dB under certain conditions. The transition from elastic to plastic deformation significantly influences the noise characteristics. Surfaces with Ra ≥ 0.1 μm showed a 10–15 dB increase in noise compared to smoother surfaces. These findings offer insights into optimizing surface parameters for noise reduction in rolling contact applications, providing a foundation for designing low-noise mechanical systems. Future experimental validations are expected to enhance the practical applications of this analytical framework. Full article
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<p>Numerically generated Gaussian rough surfaces with (<b>a</b>) Ra 1 μm and (<b>b</b>) Ra 5 μm. The correlation length in the x− and y−directions is 10 μm. The rough surfaces are discretized by 256 × 256 nodes with identical spacing.</p>
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<p>Contact geometry of a sphere and a plane.</p>
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<p>Flowchart of contact analysis.</p>
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<p>Storage and release of elastic energy of the asperity.</p>
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<p>Three-dimensional contact pressures on surfaces with four different Ra values for applied loads of 5 N and 100 N: (<b>a</b>) Ra 0.01 μm; (<b>b</b>) Ra 0.1 μm; (<b>c</b>) Ra 1 μm; and (<b>d</b>) Ra 5 μm.</p>
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<p>Three-dimensional contact deformations on surfaces with four different Ra values for applied loads of 5 N and 100 N: (<b>a</b>) Ra 0.01 μm; (<b>b</b>) Ra 0.1 μm; (<b>c</b>) Ra 1 μm; and (<b>d</b>) Ra 5 μm.</p>
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<p>Plastic deformation ratio for surfaces with four different Ra values as a function of increasing loads. The error bars show standard deviations of the simulation results.</p>
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<p>A-weighted SPL values of surfaces with various Ra values for different normal loads under rolling contact values of (<b>a</b>) Ra 0.01 μm, (<b>b</b>) Ra 0.1 μm, (<b>c</b>) Ra 1 μm, and (<b>d</b>) Ra 5 μm.</p>
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<p>A-weighted SPL values of surfaces with various Ra values for different normal loads under rolling contact values of (<b>a</b>) Ra 0.01 μm, (<b>b</b>) Ra 0.1 μm, (<b>c</b>) Ra 1 μm, and (<b>d</b>) Ra 5 μm.</p>
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<p>Maximum A-weighted SPL values with four different Ra values as a function of increasing load. The error bars show standard deviations of the simulation results.</p>
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16 pages, 10367 KiB  
Article
Influence of the Deformation Degree of Combined Loadings on the Structural and Mechanical Properties of Stainless Steels
by Magdalena Gabriela Huțanu, Liviu Andrușcă, Marcelin Benchea, Mihai-Adrian Bernevig, Dragoș Cristian Achiței, Ștefan-Constantin Lupescu, Gheorghe Bădărău and Nicanor Cimpoeșu
J. Manuf. Mater. Process. 2025, 9(2), 45; https://doi.org/10.3390/jmmp9020045 - 1 Feb 2025
Viewed by 506
Abstract
Stainless steels have many practical applications requiring various mechanical or chemical demands in the working environment. By optimizing a device used in mechanical experiments for torsional loading, several cylindrical samples were tested (both ends twisted with the same torque value in opposite directions) [...] Read more.
Stainless steels have many practical applications requiring various mechanical or chemical demands in the working environment. By optimizing a device used in mechanical experiments for torsional loading, several cylindrical samples were tested (both ends twisted with the same torque value in opposite directions) of 316L stainless steel (SS) to evaluate changes in the structural, chemical, and mechanical characteristics. Initially, the experimental samples were pre-loaded by tension in the elastic range (6%) and then subjected to torsion (772°) at different rates: 5, 10, and 20 mm/min. The experimental sequence consisted of a combined loading protocol with an initial tensile test followed by a subsequent torsional test. Two reference tests were performed by fracturing the samples in both torsion and tension to determine the mechanical strength parameters. The macro- and microstructural evolution of the samples as a function of the torsional degree was followed by scanning electron microscopy. The microhardness modification of the material was observed because of the strain (the microhardness variation from the center of the disk sample to the edge was also monitored). Structurally, all samples showed grain size changes because of torsional/compressive deformation zones and an increase in the degree of grain boundary misorientation. From the tensile and torsional behaviors of 316L SS and the structural results obtained, it was concluded that these materials are suitable for complex stress states in the elasto-plastic range through tensile and torsion. A reduction in Young’s modulus of up to four times the initial value at medium and high stress rates was observed when complex stresses were applied. Full article
(This article belongs to the Special Issue Advances in Metal Forming and Additive Manufacturing)
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<p>Experimental set-up used for the torsion test in (<b>a</b>); 3D model of the torsion device in (<b>b</b>); schematic presentation of the combined stress in (<b>c</b>); main dimensions of the specimen in (<b>d</b>).</p>
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<p>Tensile/torsion to failure curves of austenitic 316L steel.</p>
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<p>SEM micrographs of the tensile fracture (<b>a</b>) 2D: 100×, 250× and 1000× and (<b>b</b>) 3D.</p>
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<p>SEM micrographs of the torsion fracture (<b>a</b>) 2D: 100×, 250×, and 1000× and (<b>b</b>) 3D.</p>
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<p>Combined mechanical test stages applied to experimental materials (<b>a</b>) tensile and (<b>b</b>) torsion.</p>
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<p>SEM microstructures (<b>a</b>) middle area and (<b>b</b>) edge for the initial and twisted samples with 5, 10, and 20 mm/min from left to right.</p>
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<p>Mechanical properties of the experimental samples (<b>a</b>) Load vs. depth variation, (<b>b</b>) dwell time vs. twist rate variation, (<b>c</b>) hardness vs. twist rate variation, (<b>d</b>) Indentation modulus vs. twist rate variation and (<b>e</b>) contact stiffness vs. twist rate variation.</p>
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<p>The friction coefficient variation with distance (<b>a</b>) center against the initial and (<b>b</b>) edge against the initial.</p>
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<p>3D images of the scratch (<b>a</b>) initial, (<b>b</b>) tensile + torsion rate of 5mm/min, (<b>c</b>) tensile + torsion rate of 10 mm/min, and (<b>d</b>) tensile + torsion rate of 20 mm/min.</p>
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40 pages, 11072 KiB  
Article
An Approach to Evaluate the Fatigue Life of the Material of Liquefied Gases’ Vessels Based on the Time Dependence of Acoustic Emission Parameters: Part 1
by Oleg G. Perveitalov and Viktor V. Nosov
Metals 2025, 15(2), 148; https://doi.org/10.3390/met15020148 - 31 Jan 2025
Viewed by 536
Abstract
In the first part of this article devoted to the assessment of the fatigue life of structural steels at low temperatures, a study was conducted on the effect of pre-cycling in a low-cycle fatigue mode on the time dependences of acoustic emission parameters. [...] Read more.
In the first part of this article devoted to the assessment of the fatigue life of structural steels at low temperatures, a study was conducted on the effect of pre-cycling in a low-cycle fatigue mode on the time dependences of acoustic emission parameters. Commonly used St-3 steel was tested at −60 °C with varying durabilities, after which it was fractured once during static tests. The multilevel acoustic model used made it possible to estimate the structural parameter γ at the stage of elastoplastic deformation. The stage of active development of microcracks and their coalescence corresponds to a homogeneous fracture with stable acoustic emission characteristics (signal duration, amplitude variation coefficient, etc.). It was shown that regardless of the maximum voltage (460, 480, and 500 MPa) in the cycle and the operating times of up to 0.3, 0.5, and 0.7, the structural parameter remains within the known limits. The parameters of the Weibull law distribution and the logarithmically normal distribution for the coefficient γ were obtained, theoretical and calculated fatigue curves were plotted, and a method was proposed for evaluating the number of cycles before fracture under irregular loading conditions in the real operation of pressure vessels based on the “rainflow” cycles counting method. Full article
(This article belongs to the Special Issue Fatigue Assessment of Metals)
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<p>The dependence of durability on stress for various loading modes for St-3 steel (according to [<a href="#B31-metals-15-00148" class="html-bibr">31</a>]).</p>
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<p>Geometric dimensions of the low-cycle fatigue test specimen.</p>
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<p>Test equipment: (<b>a</b>) Instron 8802 250 kN servo-hydraulic machine and Zwick/Roell BW91250 thermal camera; (<b>b</b>) arrangement of waveguides on the specimen.</p>
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<p>Specimen location in a thermal chamber with a fixed Instron DIN 2620-604 dynamic deformation sensor (extensometer).</p>
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<p>Cycle scheme during fatigue loading.</p>
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<p>A-Line work screen for processing AE signals during Specimen VII tests. Red color—1 channel; green color—2 channel.</p>
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<p>Time dependences of the cumulative AE count (black points) and stress changes (red points) during single tensile tests after various degrees of operation at 213 K: (<b>a</b>) without operation (Specimen I); (<b>b</b>) <math display="inline"><semantics> <mrow> <mrow> <mrow> <msub> <mrow> <mi>N</mi> </mrow> <mrow> <mi>i</mi> </mrow> </msub> </mrow> <mo>/</mo> <mrow> <msub> <mrow> <mi>N</mi> </mrow> <mrow> <mi>f</mi> </mrow> </msub> </mrow> </mrow> <mo>=</mo> <mn>0.3</mn> </mrow> </semantics></math> at <math display="inline"><semantics> <mrow> <msub> <mrow> <mi>σ</mi> </mrow> <mrow> <mi>m</mi> <mi>a</mi> <mi>x</mi> </mrow> </msub> <mo>=</mo> <mn>460</mn> <mtext> </mtext> <mi mathvariant="normal">M</mi> <mi mathvariant="normal">P</mi> <mi mathvariant="normal">a</mi> </mrow> </semantics></math> (Specimen VI); (<b>c</b>) <math display="inline"><semantics> <mrow> <mrow> <mrow> <msub> <mrow> <mi>N</mi> </mrow> <mrow> <mi>i</mi> </mrow> </msub> </mrow> <mo>/</mo> <mrow> <msub> <mrow> <mi>N</mi> </mrow> <mrow> <mi>f</mi> </mrow> </msub> </mrow> </mrow> <mo>=</mo> <mn>0.5</mn> </mrow> </semantics></math> at <math display="inline"><semantics> <mrow> <msub> <mrow> <mi>σ</mi> </mrow> <mrow> <mi>m</mi> <mi>a</mi> <mi>x</mi> </mrow> </msub> <mo>=</mo> <mn>480</mn> <mtext> </mtext> <mi mathvariant="normal">M</mi> <mi mathvariant="normal">P</mi> <mi mathvariant="normal">a</mi> </mrow> </semantics></math> (Specimen VII); and (<b>d</b>) <math display="inline"><semantics> <mrow> <mrow> <mrow> <msub> <mrow> <mi>N</mi> </mrow> <mrow> <mi>i</mi> </mrow> </msub> </mrow> <mo>/</mo> <mrow> <msub> <mrow> <mi>N</mi> </mrow> <mrow> <mi>f</mi> </mrow> </msub> </mrow> </mrow> <mo>=</mo> <mn>0.7</mn> </mrow> </semantics></math> at <math display="inline"><semantics> <mrow> <msub> <mrow> <mi>σ</mi> </mrow> <mrow> <mi>m</mi> <mi>a</mi> <mi>x</mi> </mrow> </msub> <mo>=</mo> <mn>500</mn> <mtext> </mtext> <mi mathvariant="normal">M</mi> <mi mathvariant="normal">P</mi> <mi mathvariant="normal">a</mi> </mrow> </semantics></math> (Specimen VIII).</p>
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<p>Time dependence of the logarithm of the cumulative AE count: (<b>a</b>) without pre-fatigue (Specimen I); (<b>b</b>) <math display="inline"><semantics> <mrow> <mrow> <mrow> <msub> <mrow> <mi>N</mi> </mrow> <mrow> <mi>i</mi> </mrow> </msub> </mrow> <mo>/</mo> <mrow> <msub> <mrow> <mi>N</mi> </mrow> <mrow> <mi>f</mi> </mrow> </msub> </mrow> </mrow> <mo>=</mo> <mn>0.3</mn> </mrow> </semantics></math> at <math display="inline"><semantics> <mrow> <msub> <mrow> <mi>σ</mi> </mrow> <mrow> <mi>m</mi> <mi>a</mi> <mi>x</mi> </mrow> </msub> <mo>=</mo> <mn>460</mn> <mtext> </mtext> <mi mathvariant="normal">M</mi> <mi mathvariant="normal">P</mi> <mi mathvariant="normal">a</mi> </mrow> </semantics></math> (Specimen VI); (<b>c</b>) <math display="inline"><semantics> <mrow> <mrow> <mrow> <msub> <mrow> <mi>N</mi> </mrow> <mrow> <mi>i</mi> </mrow> </msub> </mrow> <mo>/</mo> <mrow> <msub> <mrow> <mi>N</mi> </mrow> <mrow> <mi>f</mi> </mrow> </msub> </mrow> </mrow> <mo>=</mo> <mn>0.5</mn> </mrow> </semantics></math> at <math display="inline"><semantics> <mrow> <msub> <mrow> <mi>σ</mi> </mrow> <mrow> <mi>m</mi> <mi>a</mi> <mi>x</mi> </mrow> </msub> <mo>=</mo> <mn>480</mn> <mtext> </mtext> <mi mathvariant="normal">M</mi> <mi mathvariant="normal">P</mi> <mi mathvariant="normal">a</mi> </mrow> </semantics></math> (Specimen VII); and (<b>d</b>) <math display="inline"><semantics> <mrow> <mrow> <mrow> <msub> <mrow> <mi>N</mi> </mrow> <mrow> <mi>i</mi> </mrow> </msub> </mrow> <mo>/</mo> <mrow> <msub> <mrow> <mi>N</mi> </mrow> <mrow> <mi>f</mi> </mrow> </msub> </mrow> </mrow> <mo>=</mo> <mn>0.7</mn> </mrow> </semantics></math> at <math display="inline"><semantics> <mrow> <msub> <mrow> <mi>σ</mi> </mrow> <mrow> <mi>m</mi> <mi>a</mi> <mi>x</mi> </mrow> </msub> <mo>=</mo> <mn>500</mn> <mtext> </mtext> <mi mathvariant="normal">M</mi> <mi mathvariant="normal">P</mi> <mi mathvariant="normal">a</mi> </mrow> </semantics></math> (Specimen VIII).</p>
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<p>Time dependence of the logarithm of the cumulative AE count: (<b>a</b>) without pre-fatigue (Specimen I); (<b>b</b>) <math display="inline"><semantics> <mrow> <mrow> <mrow> <msub> <mrow> <mi>N</mi> </mrow> <mrow> <mi>i</mi> </mrow> </msub> </mrow> <mo>/</mo> <mrow> <msub> <mrow> <mi>N</mi> </mrow> <mrow> <mi>f</mi> </mrow> </msub> </mrow> </mrow> <mo>=</mo> <mn>0.3</mn> </mrow> </semantics></math> at <math display="inline"><semantics> <mrow> <msub> <mrow> <mi>σ</mi> </mrow> <mrow> <mi>m</mi> <mi>a</mi> <mi>x</mi> </mrow> </msub> <mo>=</mo> <mn>460</mn> <mtext> </mtext> <mi mathvariant="normal">M</mi> <mi mathvariant="normal">P</mi> <mi mathvariant="normal">a</mi> </mrow> </semantics></math> (Specimen VI); (<b>c</b>) <math display="inline"><semantics> <mrow> <mrow> <mrow> <msub> <mrow> <mi>N</mi> </mrow> <mrow> <mi>i</mi> </mrow> </msub> </mrow> <mo>/</mo> <mrow> <msub> <mrow> <mi>N</mi> </mrow> <mrow> <mi>f</mi> </mrow> </msub> </mrow> </mrow> <mo>=</mo> <mn>0.5</mn> </mrow> </semantics></math> at <math display="inline"><semantics> <mrow> <msub> <mrow> <mi>σ</mi> </mrow> <mrow> <mi>m</mi> <mi>a</mi> <mi>x</mi> </mrow> </msub> <mo>=</mo> <mn>480</mn> <mtext> </mtext> <mi mathvariant="normal">M</mi> <mi mathvariant="normal">P</mi> <mi mathvariant="normal">a</mi> </mrow> </semantics></math> (Specimen VII); and (<b>d</b>) <math display="inline"><semantics> <mrow> <mrow> <mrow> <msub> <mrow> <mi>N</mi> </mrow> <mrow> <mi>i</mi> </mrow> </msub> </mrow> <mo>/</mo> <mrow> <msub> <mrow> <mi>N</mi> </mrow> <mrow> <mi>f</mi> </mrow> </msub> </mrow> </mrow> <mo>=</mo> <mn>0.7</mn> </mrow> </semantics></math> at <math display="inline"><semantics> <mrow> <msub> <mrow> <mi>σ</mi> </mrow> <mrow> <mi>m</mi> <mi>a</mi> <mi>x</mi> </mrow> </msub> <mo>=</mo> <mn>500</mn> <mtext> </mtext> <mi mathvariant="normal">M</mi> <mi mathvariant="normal">P</mi> <mi mathvariant="normal">a</mi> </mrow> </semantics></math> (Specimen VIII).</p>
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<p>AE features for determining the start and end times of the uniform fracture stage for specimen VII (<math display="inline"><semantics> <mrow> <mrow> <mrow> <msub> <mrow> <mi>N</mi> </mrow> <mrow> <mi>i</mi> </mrow> </msub> </mrow> <mo>/</mo> <mrow> <msub> <mrow> <mi>N</mi> </mrow> <mrow> <mi>f</mi> </mrow> </msub> </mrow> </mrow> <mo>=</mo> <mn>0.7</mn> </mrow> </semantics></math> at <math display="inline"><semantics> <mrow> <msub> <mrow> <mi>σ</mi> </mrow> <mrow> <mi>m</mi> <mi>a</mi> <mi>x</mi> </mrow> </msub> <mo>=</mo> <mn>500</mn> <mtext> </mtext> <mi mathvariant="normal">M</mi> <mi mathvariant="normal">P</mi> <mi mathvariant="normal">a</mi> </mrow> </semantics></math>): (<b>a</b>) time dependence of the signal overlap coefficient; (<b>b</b>) time dependence of the amplitude variation coefficient.</p>
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<p>Photo of specimens that underwent static tensile and fatigue tests: specimens I, II, and specimens IV–XI.</p>
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<p>Results of fatigue life calculations based on low-temperature AE testing: (<b>a</b>) calculated and experimental fatigue life values; (<b>b</b>) fatigue curve of St-3 steel at low temperatures for similar specimens according to Strizhalo [<a href="#B51-metals-15-00148" class="html-bibr">51</a>].</p>
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<p>Results of acoustic emission tests for single static fracture of standard specimens after various levels of fatigue accumulation: (<b>a</b>) time dependencies of the cumulative AE count for specimens after an accumulation of 0.3, 0.5, and 0.7 of the life and in the initial state during static tests and (<b>b</b>) time dependencies of the logarithm of the cumulative AE count and stress–strain curves (steel 20) [<a href="#B58-metals-15-00148" class="html-bibr">58</a>]; (<b>c</b>) time dependencies of the cumulative AE count for specimens after an accumulation of 0.3, 0.5, and 0.7 of the life and in the initial state during static tests and (<b>d</b>) time dependencies of the logarithm of the cumulative AE count and stress–strain curves (low-carbon steel 20) [<a href="#B55-metals-15-00148" class="html-bibr">55</a>]; (<b>e</b>) time dependencies of the cumulative AE count for specimens after an accumulation of 0.3, 0.5, and 0.7 of the life, in the initial state and a tensile specimen during static tests and (<b>f</b>) time dependencies of the logarithm of the cumulative AE count and stress–strain curves (steel 15Kh2GMF) [<a href="#B53-metals-15-00148" class="html-bibr">53</a>].</p>
Full article ">Figure 12 Cont.
<p>Results of acoustic emission tests for single static fracture of standard specimens after various levels of fatigue accumulation: (<b>a</b>) time dependencies of the cumulative AE count for specimens after an accumulation of 0.3, 0.5, and 0.7 of the life and in the initial state during static tests and (<b>b</b>) time dependencies of the logarithm of the cumulative AE count and stress–strain curves (steel 20) [<a href="#B58-metals-15-00148" class="html-bibr">58</a>]; (<b>c</b>) time dependencies of the cumulative AE count for specimens after an accumulation of 0.3, 0.5, and 0.7 of the life and in the initial state during static tests and (<b>d</b>) time dependencies of the logarithm of the cumulative AE count and stress–strain curves (low-carbon steel 20) [<a href="#B55-metals-15-00148" class="html-bibr">55</a>]; (<b>e</b>) time dependencies of the cumulative AE count for specimens after an accumulation of 0.3, 0.5, and 0.7 of the life, in the initial state and a tensile specimen during static tests and (<b>f</b>) time dependencies of the logarithm of the cumulative AE count and stress–strain curves (steel 15Kh2GMF) [<a href="#B53-metals-15-00148" class="html-bibr">53</a>].</p>
Full article ">Figure 12 Cont.
<p>Results of acoustic emission tests for single static fracture of standard specimens after various levels of fatigue accumulation: (<b>a</b>) time dependencies of the cumulative AE count for specimens after an accumulation of 0.3, 0.5, and 0.7 of the life and in the initial state during static tests and (<b>b</b>) time dependencies of the logarithm of the cumulative AE count and stress–strain curves (steel 20) [<a href="#B58-metals-15-00148" class="html-bibr">58</a>]; (<b>c</b>) time dependencies of the cumulative AE count for specimens after an accumulation of 0.3, 0.5, and 0.7 of the life and in the initial state during static tests and (<b>d</b>) time dependencies of the logarithm of the cumulative AE count and stress–strain curves (low-carbon steel 20) [<a href="#B55-metals-15-00148" class="html-bibr">55</a>]; (<b>e</b>) time dependencies of the cumulative AE count for specimens after an accumulation of 0.3, 0.5, and 0.7 of the life, in the initial state and a tensile specimen during static tests and (<b>f</b>) time dependencies of the logarithm of the cumulative AE count and stress–strain curves (steel 15Kh2GMF) [<a href="#B53-metals-15-00148" class="html-bibr">53</a>].</p>
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<p>Fatigue curves for the experimental materials with marked points of the calculated fatigue life: (<b>a</b>) steel 20 with preliminary cycling at 390 MPa; (<b>b</b>) steel 20 with preliminary loading at 330 MPa; and (<b>c</b>) steel 15Kh2GMF with preliminary loading at 800 MPa.</p>
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<p>Fatigue curves for the experimental materials with marked points of the calculated fatigue life: (<b>a</b>) steel 20 with preliminary cycling at 390 MPa; (<b>b</b>) steel 20 with preliminary loading at 330 MPa; and (<b>c</b>) steel 15Kh2GMF with preliminary loading at 800 MPa.</p>
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<p>Distribution of the structural parameter <math display="inline"><semantics> <mrow> <mi>γ</mi> </mrow> </semantics></math> during uniform fracture of specimens VI, VII, and VIII. (<b>a</b>) Weibull distribution; (<b>b</b>) log-normal distribution.</p>
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<p>Comparison of experimental and calculated values of cumulative AE count based on numerical simulation data.</p>
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<p>Results of modeling the stress state of the gas receiver under operating conditions: (<b>a</b>) stress distribution over the object’s surface and maximum stresses in the area near the flange (as a natural stress concentrator); (<b>b</b>) results of the residual fatigue life calculation based on the low-cycle fatigue criterion using the Manson–Coffin model, taking into account the accumulation of plastic deformation.</p>
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<p>Example diagram for determining the range of acting stresses in the concentrator—the source of AE signals: (<b>a</b>) logarithm of the cumulative AE count and the pressure increase graph during hydraulic testing of the vessel [<a href="#B70-metals-15-00148" class="html-bibr">70</a>]; (<b>b</b>) logarithm of the cumulative count and stress during the testing of laboratory specimens from the pressure vessel material (<math display="inline"><semantics> <mrow> <msub> <mrow> <mi>t</mi> </mrow> <mrow> <mn>1</mn> </mrow> </msub> </mrow> </semantics></math>—time moment of the beginning of the uniform microcracking stage).</p>
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<p>Operational data for converting pressure drop cycles in a liquid ethylene storage tank 47D01 into a fatigue curve relationship – pressure drop oscillogram over the course of a year.</p>
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<p>Simulation of the operational pressure drop process using laboratory AE testing of standard specimens and Comsol Multiphysics 5.6 software: (<b>a</b>) a graph of pressure changes in the tank during the year, modeled using the software; (<b>b</b>) stress–strain state of a standard specimen for low-cycle fatigue.</p>
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<p>Matrix histograms of the distribution of loading blocks of a real vessel (<b>a</b>) by stress amplitude <math display="inline"><semantics> <mrow> <mover accent="true"> <mrow> <mi>σ</mi> </mrow> <mo stretchy="false">^</mo> </mover> </mrow> </semantics></math> and mean cycle stress <math display="inline"><semantics> <mrow> <mover accent="true"> <mrow> <mi>σ</mi> </mrow> <mo stretchy="false">¯</mo> </mover> </mrow> </semantics></math> and (<b>b</b>) by relative contribution to damage accumulation.</p>
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<p>Flowchart of the method implementation.</p>
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18 pages, 13259 KiB  
Article
Impact of Ni Doping on the Microstructure and Mechanical Properties of TiB2 Films
by Ying Wang, Xu Wang, Hailong Shang, Xiaotong Liu, Yu Qi, Xiaoben Qi and Ning Zhong
Nanomaterials 2025, 15(3), 229; https://doi.org/10.3390/nano15030229 - 31 Jan 2025
Viewed by 542
Abstract
The TiB2 film exhibits exceptional hardness and chemical stability due to its unique crystal structure and robust covalent bonds, but it also demonstrates high brittleness and poor toughness, which restricts its practical applications in engineering. By appropriately incorporating metal dopants, the toughness [...] Read more.
The TiB2 film exhibits exceptional hardness and chemical stability due to its unique crystal structure and robust covalent bonds, but it also demonstrates high brittleness and poor toughness, which restricts its practical applications in engineering. By appropriately incorporating metal dopants, the toughness of the ceramic matrix can be enhanced without compromising its inherent hardness. In this study, TiB2 films with different nickel contents (0–32.22 at.%) were fabricated through radio frequency magnetron sputtering. The microstructure, chemical composition, phase structure, and mechanical properties were analyzed using scanning electron microscopy, transmission electron microscopy, X-ray diffraction, X-ray photoelectron spectroscopy and nanoindentation tester. The pure TiB2 film exhibited (0001) and (0002) peaks; however, the addition of nickel resulted in broadening of the (0001) peak and disappearance of the (0002) peak, and no crystalline nickel or other nickel-containing phases could be identified. It was found that the incorporation of nickel refines the grain structure of titanium diboride, with nickel present in an amorphous form at the boundaries of titanium diboride, thereby forming a wrapped structure. The enrichment of nickel at the grain boundary becomes more pronounced as the nickel content is further increased, which hinders the growth of TiB2 grains, resulting in the thinning of columnar crystals and formation of nanocrystalline in the film, and the coating hardness remains above 20 GPa, when the nickel content is less than 10.83 at.%. With the increase in nickel content, titanium diboride exhibited a tendency to form an amorphous structure, while nickel became increasingly enriched at the boundaries, and the coating hardness and elastic modulus decreased. The wrapped microstructure could absorb the energy generated by compressive shear stress through plastic deformation, which should be beneficial to improve the toughness of the coatings. The addition of nickel enhanced the adhesion between the film and substrate while reducing the friction coefficient of the film. Specifically, when the nickel content reached 4.26 at.%, a notable enhancement in both nanohardness and toughness was observed for nanocomposite films. Full article
(This article belongs to the Special Issue Design and Applications of Heterogeneous Nanostructured Materials)
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Figure 1
<p>Schematic diagram of crack length (c) and half of the diagonal length of the indentation (a).</p>
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<p>X-ray diffraction patterns of TiB<sub>2</sub> films with different Ni contents.</p>
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<p>XPS spectra of B, Ni Ti, obtained from the surface of TiB<sub>2</sub> film with 4.26 at.% Ni (<b>a</b>) -B 1s; (<b>b</b>) -Ni 2p; (<b>c</b>) -Ti 2p.</p>
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<p>XPS spectra of TiB<sub>2</sub>-Ni films with different nickel content after etching 120 s (<b>a</b>) -B 1s; (<b>b</b>) -Ni 2p; (<b>c</b>) -Ti 2p.</p>
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<p>Cross-sectional SEM micrographs of TiB<sub>2</sub> film with different Ni contents: (<b>a</b>) 4.26 at.%; (<b>b</b>) 10.83 at.%; (<b>c</b>) 23.45 at.%; (<b>d</b>) 32.22 at.%.</p>
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<p>TEM results for TiB<sub>2</sub> films with different Ni contents: (<b>a</b>) bright-field cross-sectional image (4.26 at.%); (<b>b</b>) SAED pattern from the area marked as circle in (<b>a</b>); (<b>c</b>) HRTEM image; (<b>d</b>) bright-field cross-sectional image (32.22 at.%); (<b>e</b>) SAED pattern from the area marked as circle in (<b>d</b>); (<b>f</b>) HRTEM image.</p>
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<p>HAADF images and the corresponding EDX map scan of B, Ti and Ni elements of TiB<sub>2</sub>-Ni films. (<b>a</b>) Image of TiB<sub>2</sub> with 4.26 at.% Ni film; (<b>b</b>) map scan of B; (<b>c</b>) map scan of Ti; (<b>d</b>) map scan of Ni; (<b>e</b>) Image of TiB<sub>2</sub> with 32.22 at.% Ni film; (<b>f</b>) map scan of B; (<b>g</b>) map scan of Ti; (<b>h</b>) map scan of Ni.</p>
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<p>Plan-view TEM micrograph of TiB<sub>2</sub> with 10.83 at.% Ni coating. (<b>a</b>) low-magnification; (<b>b</b>) high-magnification.</p>
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<p>The indentation curves of TiB<sub>2</sub> film with different Ni contents.</p>
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<p>The nanohardness and elastic modulus of TiB<sub>2</sub> film with different Ni contents.</p>
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<p>The values of H/E, H<sup>3</sup>/E<sup>2</sup>, normalized plastic depth value (δ<sub>H</sub>) of TiB<sub>2</sub> film with different Ni contents.</p>
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<p>The calculated fracture toughness with different Ni contents.</p>
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<p>Indentation morphology of TiB<sub>2</sub> film with different Ni contents (<b>a</b>) 4.26 at.%;(<b>b</b>) 32.22 at.%.</p>
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<p>The nanoscratch morphology of TiB<sub>2</sub>-Ni coatings.</p>
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<p>The friction coefficient of TiB<sub>2</sub>-Ni composite coatings.</p>
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<p>Wear track morphology of TiB<sub>2</sub>-Ni coatings. (<b>a</b>) TiB<sub>2</sub> coating, (<b>b</b>)TiB<sub>2</sub>-4.26 at.% Ni coating, (<b>c</b>) TiB<sub>2</sub>-32.22 at.% Ni coating.</p>
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14 pages, 29079 KiB  
Article
Molecular Dynamics Investigation on Grain Size-Dependent Superelastic Behavior of CuZr Shape Memory Alloys
by Mixun Zhu, Kai Wang, Hongtao Zhong, Huahuai Shen, Yong Zhang, Xiaoling Fu and Yuanzheng Yang
Metals 2025, 15(2), 142; https://doi.org/10.3390/met15020142 - 29 Jan 2025
Viewed by 641
Abstract
The superelasticity of CuZr shape memory alloys (SMAs) originates from stress-induced transformations between the B2 (austenite) and B19’ (martensite) phases. Grain size is a key parameter affecting the superelasticity of shape memory alloys. Previous studies on NiTi, Fe-based, and Cu-based SMAs confirm that [...] Read more.
The superelasticity of CuZr shape memory alloys (SMAs) originates from stress-induced transformations between the B2 (austenite) and B19’ (martensite) phases. Grain size is a key parameter affecting the superelasticity of shape memory alloys. Previous studies on NiTi, Fe-based, and Cu-based SMAs confirm that altering grain size effectively regulates superelasticity. Current research on the influence of grain size on the superelasticity of CuZr shape memory alloys (SMAs) is relatively sparse. This study employs molecular dynamics simulations to evaluate the effect of grain size on the superelasticity of CuZr SMAs through uniaxial loading–unloading tests. Polycrystalline samples with grain sizes of 6.59 nm, 5 nm, and 4 nm were analyzed. The results indicate that reducing grain size can decrease the irrecoverable strain, thereby enhancing superelasticity. The improvement in superelasticity is attributed to a higher recovery rate of the martensite-to-austenite transformation, allowing more plastic deformation within the grain interior to recover during unloading, and thereby reducing the irrecoverable strain. The recovery rate of the martensite-to-austenite transformation is closely related to the elastic strain energy accumulated within the grain interior during loading. Full article
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<p>Polycrystalline B2-CuZr samples with different average grain sizes of (<b>a</b>) <span class="html-italic">d</span><sub>ave</sub> = 6.59 nm, (<b>b</b>) <span class="html-italic">d</span><sub>ave</sub> = 5 nm, and (<b>c</b>) <span class="html-italic">d</span><sub>ave</sub> = 4 nm. The B2-CuZr crystals within the grain interior are colored in blue and the grain boundaries are colored in white.</p>
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<p>Engineering stress–strain curves of polycrystalline B2-CuZr samples with grain sizes of 6.59 nm, 5 nm, and 4 nm during tensile loading and unloading.</p>
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<p>The evolution of local shear strain distribution and the corresponding phase configuration in polycrystalline B2-CuZr samples with grain sizes of (<b>a</b>) 6.59 nm, (<b>b</b>) 5 nm, and (<b>c</b>) 4 nm. The initial state without applied strain (<span class="html-italic">ε</span> = 0), the critical strain for martensitic transformation, the maximum strain of loading (<span class="html-italic">ε</span> = 0.08), and the end point of unloading were selected for display.</p>
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<p>Phase content evolution of polycrystalline B2-CuZr samples with grain sizes of (<b>a</b>) 6.59 nm, (<b>b</b>) 5 nm, and (<b>c</b>) 4 nm during loading and unloading.</p>
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<p>(<b>a</b>) The volumetric strain <math display="inline"><semantics> <mrow> <msub> <mi>ε</mi> <mi mathvariant="normal">V</mi> </msub> </mrow> </semantics></math> and irrecoverable strain of polycrystalline B2-CuZr samples with different grain sizes. (<b>b</b>) The volumetric strain <math display="inline"><semantics> <mrow> <msub> <mi>ε</mi> <mi mathvariant="normal">V</mi> </msub> </mrow> </semantics></math> of polycrystal, grain interior, and grain boundaries with different grain sizes.</p>
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<p>The engineering stress–strain curves of polycrystalline B2-CuZr samples with grain sizes of (<b>a</b>) 6.59 nm, (<b>b</b>) 5 nm, and (<b>c</b>) 4 nm and the corresponding potential energy changes with applied strain.</p>
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<p>The correlation between the recovery rate of the R phase and the elastic energy accumulation within the grain interior during loading.</p>
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24 pages, 8403 KiB  
Article
Experimental Study on the Seismic Performance of Confined High Walls of Autoclaved Aerated Concrete Panels Used in Subway Stations
by Xiaowei Li, Qidi Zhang, Han Bao, Yong Yao and Zhaoqiang Zhang
Buildings 2025, 15(3), 416; https://doi.org/10.3390/buildings15030416 - 28 Jan 2025
Viewed by 508
Abstract
This study addresses the unique challenge of the partition walls in subway stations, featuring high height, fire prevention, and located outside the main frames, by introducing a confined autoclaved aerated concrete (AAC) panel wall system. Different from studies on a main frame with [...] Read more.
This study addresses the unique challenge of the partition walls in subway stations, featuring high height, fire prevention, and located outside the main frames, by introducing a confined autoclaved aerated concrete (AAC) panel wall system. Different from studies on a main frame with infill walls, this study aimed to explore the seismic performance of partition walls, which were fabricated with confined high AAC panel walls and located outside the main frames. A custom-designed partition wall, measuring 6600 mm in height, 3400 mm in width, and 200 mm in thickness, underwent cyclic testing. A detailed analysis of specimen’s failure modes was conducted, focusing on seismic behavior such as hysteresis curves, envelope curves, ductility, stiffness degradation, and energy-dissipation capacity. Additionally, the study delved into shear deformation, relative slippage between AAC panels, and reinforcement strains within the specimen. Finally, the D-value method for calculating the initial stiffness of the confined high AAC panel walls and the weak sub-structural approach for determining the load-bearing capacity of confined high AAC panel walls were proposed and validated. The results indicate that the strength degradation factor of the confined high AAC panel walls ranges from 0.971 to 0.716. The drift of its upper portion accounts for 76.94–83.63% of the total drift, while the energy dissipation factor of its upper portion is 0.8–4.8% higher than that of the entire specimen. The yield and ultimate drift rotations of the entire confined high AAC panel wall and its upper portions satisfy the elastic and elastic-plastic inter-story drift rotation limits specified in the Chinese code. The calculated initial stiffness of the confining frame, obtained using the D-value method, closely aligns with experimental results, with a deviation of only 2.48%. Additionally, the load-bearing capacity calculated using the weak sub-structural approach deviates from the experimental average by just 4.30%. Full article
(This article belongs to the Section Building Structures)
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<p>Research framework.</p>
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<p>Details of partition walls used in subway stations. (<b>a</b>) Specimen; (<b>b</b>) cast-in-place concrete joint; (<b>c</b>) L-shaped connector. (In mm).</p>
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<p>Details of panel P1. (in mm).</p>
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<p>Details of prefabricated beams and columns. (<b>a</b>) B1, B2, C1, and C2; (<b>b</b>) B3; (<b>c</b>) C3. (In mm).</p>
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<p>Setup and instruments. (In mm).</p>
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<p>Loading procedure.</p>
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<p>Failure of the specimen: (<b>a</b>) Deformation; (<b>b</b>) cracks. (In mm).</p>
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<p>Cracks on panels. (<b>a</b>) Crack on the lower left side of panel ‘P1d’; (<b>b</b>) crack on lower left corner of the panel ‘P2d’.</p>
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<p>Cracks on upper joints. (<b>a</b>) J6; (<b>b</b>) J5; (<b>c</b>) J4; (<b>d</b>) J3.</p>
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<p>Diagonal strut effect of infilled AAC panels.</p>
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<p>Hysteresis curves of (<b>a</b>) specimen and (<b>b</b>) upper portion.</p>
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<p>Geometric plotting method.</p>
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<p>Envelope curves.</p>
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<p>Strength degradation curves.</p>
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<p>Stiffness degradation curves.</p>
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<p>Hysteretic loop.</p>
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<p>Energy dissipation factors.</p>
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<p>Shear deformation.</p>
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<p>Shear angle.</p>
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<p>Slippage vs. drift curves.</p>
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<p>Strains in the reinforcement of columns (<b>a</b>) C2a and (<b>b</b>) C2b.</p>
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<p>Calculation diagram for <span class="html-italic">D</span>-Value method.</p>
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<p>Flowchart for calculating the lateral stiffness of confining frame.</p>
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<p>Calculation process of the load-bearing capacity of the sub-structure.</p>
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16 pages, 12609 KiB  
Article
Microstructure and Micro-Mechanical Properties of Thermally Sprayed HA-TiO2 Coating on Beta-Titanium Substrate
by Abdulaziz Kurdi, Doaa Almalki, Ahmed Degnah and Animesh Kumar Basak
Materials 2025, 18(3), 540; https://doi.org/10.3390/ma18030540 - 24 Jan 2025
Viewed by 279
Abstract
Metallic biomaterials in a solid form cause stress-shielding in orthopedic applications. Such implants also suffer from limited tissue attachment to become a part of the living system. In view of that, hydroxyapatite (HA) coating reinforced with titanium oxide (TiO2) was deposited [...] Read more.
Metallic biomaterials in a solid form cause stress-shielding in orthopedic applications. Such implants also suffer from limited tissue attachment to become a part of the living system. In view of that, hydroxyapatite (HA) coating reinforced with titanium oxide (TiO2) was deposited in a beta (β)-Titanium (Ti-35Nb-7Ta-5Zr) substrate by plasma spray. This allows us to exploit the best of the two materials, namely the relatively low modulus of β-Ti, together with the porous and bone-like structure/composition of the HA to facilitate cell growth. This is foreseen to be used as an implant, particularly for musculoskeletal-related disability. Detailed scanning electron microscopy (SEM) investigation shows the lamellar structure of the coating that is composed of different phases and some porosities. Transmission electron microscopy (TEM) confirms the co-existence of both the amorphous and crystalline phases that build up the coating structure. In situ micro-mechanical tests revealed that the HA-TiO2 coating was low in strength and modules compared to that of the substrate material, together with lower ductility. The yield stress and modulus of elasticity of the coating were about 877 ± 174 MPa and 447 ± 24 MPa, respectively. In contrast, the beta (β)-Ti substrate possesses about 990 ± 85 MPa of yield stress and 259 ± 19 MPa modulus of elasticity. The deformation mechanism was also quite different, where the coating crumbled under compressive loading, featuring limited ductility with cleavage (brittle)-type fracture, and the substrate showed plastic flow of materials in the form of slip/shear planes with extended ductility. Full article
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<p>Back-scattered SEM micrographs on HA-TiO<sub>2</sub> coating: (<b>a</b>,<b>b</b>) planner view of as-received coating (without metallographic polishing) and (<b>c</b>,<b>d</b>) cross-sectional view (after metallographic polishing).</p>
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<p>Elemental mapping on the cross-section of HA-TiO<sub>2</sub> coating: (<b>a</b>) secondary electron image, (<b>b</b>) O map, (<b>c</b>) Ca map, (<b>d</b>) P map, (<b>e</b>) Ti map, (<b>f</b>) EDS layered image, and (<b>g</b>) map sum spectrum.</p>
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<p>TEM micrographs on HA-TiO<sub>2</sub> coating: (<b>a</b>) overall view of the coating representing different areas, (<b>b</b>–<b>d</b>) high resolution (HR)-TEM images of areas 1–3, respectively, as marked in (<b>a</b>) with corresponding electron diffraction pattern (<b>e</b>–<b>g</b>).</p>
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<p>FIB-prepared micro-pillars for compression (<b>a</b>) on the substrate (β-Ti) and (<b>b</b>) on the HA-TiO<sub>2</sub> coating. Enlarged views of the micro-pillars are shown as inserts in respective images.</p>
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<p>Stress–strain curves on the cross-section of the HA-TiO<sub>2</sub> coating and substrate β-Ti metallic biomaterial subjected to micro-pillar compression.</p>
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<p>Deformed micro-pillars on the coating cross-section after compression: (<b>a</b>) micro-pilalr 1, (<b>b</b>) micro-pillar 2 and (<b>c</b>) micro-pillar 3.</p>
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<p>Deformed micro-pillars on the β-Ti substrate after compression: (<b>a</b>) micro-pilalr 1, (<b>b</b>) micro-pillar 2 and (<b>c</b>) micro-pillar 3.</p>
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<p>Deformed micro-pillars on coating (<b>a</b>,<b>b</b>) and substrate (<b>c</b>,<b>d</b>) after compression.</p>
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26 pages, 26651 KiB  
Article
Deformation and Failure Mechanism and Control of Water-Rich Sandstone Roadways in the Huaibei Mining Area
by Zhisen Zhang, Yukuan Fan, Qiang Xu, Kai Li, Minkang Han and Lixiang Fei
Appl. Sci. 2025, 15(3), 1177; https://doi.org/10.3390/app15031177 - 24 Jan 2025
Viewed by 254
Abstract
The sandstone roof rock in the Huaibei mining area contains abundant water at depths of 2–3 m. Water–rock interactions in the rock-surrounding roadway can cause significant deformation, seriously threatening the safety of mine operations. Investigating the deformation and failure mechanisms of water-rich sandstone [...] Read more.
The sandstone roof rock in the Huaibei mining area contains abundant water at depths of 2–3 m. Water–rock interactions in the rock-surrounding roadway can cause significant deformation, seriously threatening the safety of mine operations. Investigating the deformation and failure mechanisms of water-rich sandstone is therefore of critical importance. In this study, X-ray diffraction and scanning electron microscopy were used to analyze the composition and microstructure of water-rich sandstone. Based on the stress state during the roadway excavation, a true triaxial loading scheme with four different stress paths was designed to study the effects of different moisture contents and loading methods on the mechanical properties of the sandstone. The results show that the deviatoric stress decreased for all stress paths. Acoustic emission (AE) characteristics during the deformation and failure processes were also studied, which indicated that the AE b-value decreased, increased, and then decreased again corresponding to the primary compaction, elastic deformation, and plastic deformation evolutionary processes in the internal microstructure of the rock. The variation in the b-value reflected the development and expansion of internal fractures. These findings provide useful insights for controlling the stability of the surrounding rock in water-rich roadways in coal mines. Full article
(This article belongs to the Special Issue Novel Research on Rock Mechanics and Geotechnical Engineering)
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<p>Water inflow into the roadway roof.</p>
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<p>Schematic diagram of roadway roof degradation under water-rich conditions.</p>
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<p>Thin section of sandstone and a physical sample.</p>
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<p>Identification of the mineral composition of thin rock sections.</p>
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<p>Design of experimental scheme.</p>
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<p>True triaxial stress loading and unloading paths for sandstone samples with different water contents: (<b>a</b>) Path I; (<b>b</b>) Path II; (<b>c</b>) Path III; (<b>d</b>) Path IV.</p>
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<p>Sampling location.</p>
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<p>Pre-treatment of samples: (<b>a</b>) parent rock sample—part; (<b>b</b>) wave velocity test; (<b>c</b>) wave velocity results.</p>
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<p>Water content vs. time curve.</p>
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<p>X-ray diffraction (XRD) pattern.</p>
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<p>Scanning electron microscopy (SEM) results for sandstone.</p>
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<p>Three-dimensional deviatoric stress–strain relationship diagrams: (<b>a</b>) dry; (<b>b</b>) 30% saturated; (<b>c</b>) saturated.</p>
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<p>Failure modes of sandstone with different water contents: (<b>a</b>) dry; (<b>b</b>) 30% saturated; (<b>c</b>) saturated.</p>
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<p>Three-dimensional deviatoric stress–strain relationship diagrams: (<b>a</b>) dry; (<b>b</b>) 30% saturated; (<b>c</b>) saturated.</p>
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<p>Failure modes of sandstone with different water contents: (<b>a</b>) dry; (<b>b</b>) 30% saturated; (<b>c</b>) saturated.</p>
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<p>Three-dimensional deviatoric stress–strain relationship diagrams: (<b>a</b>) dry; (<b>b</b>) 30% saturated; (<b>c</b>) saturated.</p>
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<p>Failure modes of sandstone with different water contents: (<b>a</b>) dry; (<b>b</b>) 30% saturated; (<b>c</b>) saturated.</p>
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<p>Three-dimensional deviatoric stress–strain relationship diagrams: (<b>a</b>) dry; (<b>b</b>) 30% saturated; (<b>c</b>) saturated.</p>
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<p>Failure modes of sandstone with different water contents: (<b>a</b>) dry; (<b>b</b>) 30% saturated; (<b>c</b>) saturated.</p>
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<p>Influence of water content on mechanical behavior of sandstone.</p>
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<p>Damage variable at characteristic points with different water contents.</p>
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<p>Relationship between acoustic emission (AE) count, cumulative count, and strain.</p>
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<p>Relationship between AE <span class="html-italic">b</span>-value and offsetting force.</p>
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<p>Sensor arrangement.</p>
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<p>Industrial-scale test results: (<b>a</b>) comparison of rock mass deformations; (<b>b</b>) comparison of bolt (cable) loading conditions; (<b>c</b>) comparison of roof separation conditions.</p>
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