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23 pages, 3342 KiB  
Article
Tuning Electromagnetically Induced Transparency in a Double GaAs/AlGaAs Quantum Well with Modulated Doping
by C. A. Dagua-Conda, J. A. Gil-Corrales, R. V. H. Hahn, R. L. Restrepo, M. E. Mora-Ramos, A. L. Morales and C. A. Duque
Crystals 2025, 15(3), 248; https://doi.org/10.3390/cryst15030248 - 6 Mar 2025
Viewed by 267
Abstract
Including an n-doped layer in asymmetric double quantum wells restricts confined carriers into V-shaped potential profiles, forming discrete conduction subbands and enabling intersubband transitions. Most studies on doped semiconductor heterostructures focus on how external fields and structural parameters dictate optical absorption. However, [...] Read more.
Including an n-doped layer in asymmetric double quantum wells restricts confined carriers into V-shaped potential profiles, forming discrete conduction subbands and enabling intersubband transitions. Most studies on doped semiconductor heterostructures focus on how external fields and structural parameters dictate optical absorption. However, electromagnetically induced transparency remains largely unexplored. Here, we show that the effect of an n-doped layer GaAs/AlxGa1−xAs in an asymmetric double quantum well system is quite sensitive to the width and position of the doped layer. By self-consistently solving the Poisson and Schrödinger’s equations, we determine the electronic structure using the finite element method within the effective mass approximation. We found that the characteristics of the n-doped layer can modulate the resonance frequencies involved in the electromagnetically induced transparency phenomenon. Our results demonstrate that an n-doped layer can control the electromagnetically induced transparency effect, potentially enhancing its applications in optoelectronic devices. Full article
(This article belongs to the Section Materials for Energy Applications)
Show Figures

Figure 1

Figure 1
<p>(<b>a</b>) Schematic diagram of a GaAs/Al<sub>0.3</sub>Ga<sub>0.7</sub> As double quantum well system centered at the origin of the x-axis subject to an external electric field <span class="html-italic">F</span> polarized along the heterostructure growth direction, with <math display="inline"><semantics> <msub> <mi>L</mi> <mi>L</mi> </msub> </semantics></math> (<math display="inline"><semantics> <msub> <mi>L</mi> <mi>R</mi> </msub> </semantics></math>) denoting the left (right) well and <math display="inline"><semantics> <msub> <mi>L</mi> <mi>b</mi> </msub> </semantics></math> the barrier width, respectively. The doping layer (dashed red line) of <math display="inline"><semantics> <mi>δ</mi> </semantics></math>-width is inside the well (GaAs) at a <math display="inline"><semantics> <mi>ξ</mi> </semantics></math> separation from the origin. (<b>b</b>) Three-level ladder system configuration to study the electromagnetically induced transparency process.</p>
Full article ">Figure 2
<p>(<b>a</b>,<b>c</b>) Self-consistent potential <math display="inline"><semantics> <mrow> <msub> <mi>V</mi> <mi>sc</mi> </msub> <mrow> <mo>(</mo> <mi>x</mi> <mo>)</mo> </mrow> </mrow> </semantics></math> (solid black line) and electron confining potential <math display="inline"><semantics> <msub> <mi>V</mi> <mi>in</mi> </msub> </semantics></math> (dashed black line) for GaAs/Al<sub>0.3</sub>Ga<sub>0.7</sub> As, demonstrating two scenarios for the variations in the self-consistent potential seen by the electrons and probability densities <math display="inline"><semantics> <mrow> <mrow> <mo>|</mo> </mrow> <msub> <mi>ψ</mi> <mi>ν</mi> </msub> <msup> <mrow> <mo>|</mo> </mrow> <mn>2</mn> </msup> </mrow> </semantics></math> (<math display="inline"><semantics> <mrow> <mi>ν</mi> <mo>=</mo> <mn>1</mn> <mo>,</mo> <mn>2</mn> <mo>,</mo> <mn>3</mn> </mrow> </semantics></math>), with right-hand well width <math display="inline"><semantics> <mrow> <msub> <mi>L</mi> <mi>R</mi> </msub> <mo>=</mo> <mn>2</mn> </mrow> </semantics></math> nm (<b>a</b>) and <math display="inline"><semantics> <mrow> <msub> <mi>L</mi> <mi>R</mi> </msub> <mo>=</mo> <mn>20</mn> </mrow> </semantics></math> nm (<b>c</b>). In (<b>b</b>), the energy levels of the first three subbands as a function of the right-hand well width <math display="inline"><semantics> <msub> <mi>L</mi> <mi>R</mi> </msub> </semantics></math>. Note that the energies associated with the states presented in (<b>a</b>,<b>c</b>) correspond to the minimum and maximum values of the parameter sweep, respectively.</p>
Full article ">Figure 3
<p>(<b>a</b>,<b>c</b>) Modifications in the self-consistent potential of GaAs/Al<sub>0.3</sub>Ga<sub>0.7</sub> As induced by the external electric field with <math display="inline"><semantics> <mrow> <mi>F</mi> <mo>=</mo> <mn>0</mn> </mrow> </semantics></math> kV/cm (<b>a</b>) and <math display="inline"><semantics> <mrow> <mi>F</mi> <mo>=</mo> <mn>33</mn> </mrow> </semantics></math> kV/cm (<b>c</b>). In (<b>b</b>), the energy levels of the first three subbands are shown as a function of the external electric field.</p>
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<p>(<b>a</b>,<b>c</b>) Modifications in the GaAs/Al<sub>0.3</sub>Ga<sub>0.7</sub> As self-consistent potential induced by the location of the doped layer of <math display="inline"><semantics> <mi>δ</mi> </semantics></math>-width with <math display="inline"><semantics> <mrow> <mi>ξ</mi> <mo>=</mo> <mo>−</mo> <mn>13</mn> </mrow> </semantics></math> nm (<b>a</b>) and <math display="inline"><semantics> <mrow> <mi>ξ</mi> <mo>=</mo> <mn>15</mn> </mrow> </semantics></math> nm (<b>c</b>). In (<b>b</b>), the energy levels of the first three subbands are shown as a function of the doped delta layer location.</p>
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<p>(<b>a</b>,<b>c</b>) Modifications in the GaAs/Al<sub>0.3</sub>Ga<sub>0.7</sub> As self-consistent potential induced by the doped layer width with <math display="inline"><semantics> <mrow> <mi>δ</mi> <mo>=</mo> <mn>1</mn> </mrow> </semantics></math> nm (<b>a</b>) and <math display="inline"><semantics> <mrow> <mi>δ</mi> <mo>=</mo> <mn>4</mn> </mrow> </semantics></math> nm (<b>c</b>). In (<b>b</b>), the energy levels of the first three subbands are shown as a function of the doped delta layer width.</p>
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<p>(<b>a</b>–<b>d</b>) Electron density in terms of the <span class="html-italic">x</span>-coordinate for different configurations of the double quantum well. Two different values are shown for the width of the right-hand well (<b>a</b>), the applied electric field (<b>b</b>), the doped layer position (<b>c</b>), and the width of the doped layer (<b>d</b>).</p>
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<p>The product <math display="inline"><semantics> <mrow> <mrow> <mo>|</mo> </mrow> <msub> <mi>E</mi> <mn>10</mn> </msub> <msup> <mrow> <mo>|</mo> </mrow> <mn>2</mn> </msup> <msup> <mrow> <mo>|</mo> <msub> <mi>M</mi> <mn>10</mn> </msub> <mo>|</mo> </mrow> <mn>2</mn> </msup> </mrow> </semantics></math> with respect to the right-hand well width (<b>a</b>), of the externally applied electric field intensity (<b>b</b>), of the doped layer position (<b>c</b>), and of the doped layer width (<b>d</b>). For each figure, the parameters are the same as those in <a href="#crystals-15-00248-f001" class="html-fig">Figure 1</a>, <a href="#crystals-15-00248-f002" class="html-fig">Figure 2</a>, <a href="#crystals-15-00248-f003" class="html-fig">Figure 3</a>, <a href="#crystals-15-00248-f004" class="html-fig">Figure 4</a> and <a href="#crystals-15-00248-f005" class="html-fig">Figure 5</a>.</p>
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<p>(<b>a</b>–<b>d</b>) 2D contour plots for the electromagnetically induced transparency as a function of the photon energy. In (<b>a</b>), the effect of variations in the width of the right-hand well, the applied electric field (<b>b</b>), the location of the doped layer (<b>c</b>), and the width of the doped layer (<b>d</b>). For each figure, the parameters are the same as those in <a href="#crystals-15-00248-f001" class="html-fig">Figure 1</a>, <a href="#crystals-15-00248-f002" class="html-fig">Figure 2</a>, <a href="#crystals-15-00248-f003" class="html-fig">Figure 3</a>, <a href="#crystals-15-00248-f004" class="html-fig">Figure 4</a> and <a href="#crystals-15-00248-f005" class="html-fig">Figure 5</a>.</p>
Full article ">Figure 9
<p>Linear optical absorption coefficient (shaded curves), and electromagnetically induced transparency (unshaded curves) with respect to the incident photon energy. In (<b>a</b>) the effect of the position of the doped layer with <math display="inline"><semantics> <mi>δ</mi> </semantics></math>-width, in (<b>b</b>) the effect of the doped layer position. For each panel, the parameters are the same as those in <a href="#crystals-15-00248-f001" class="html-fig">Figure 1</a>, <a href="#crystals-15-00248-f002" class="html-fig">Figure 2</a>, <a href="#crystals-15-00248-f003" class="html-fig">Figure 3</a>, <a href="#crystals-15-00248-f004" class="html-fig">Figure 4</a> and <a href="#crystals-15-00248-f005" class="html-fig">Figure 5</a>.</p>
Full article ">Figure A1
<p>(<b>a</b>,<b>c</b>) Self-consistent potential <math display="inline"><semantics> <mrow> <msub> <mi>V</mi> <mi>sc</mi> </msub> <mrow> <mo>(</mo> <mi>x</mi> <mo>)</mo> </mrow> </mrow> </semantics></math> (solid black line) and electron confining potential <math display="inline"><semantics> <msub> <mi>V</mi> <mi>in</mi> </msub> </semantics></math> (dashed black line) for GaAs/Al<sub>0.3</sub>Ga<sub>0.7</sub> As, demonstrating two scenarios for the variations in the self-consistent potential seen by the electrons and probability densities <math display="inline"><semantics> <mrow> <mrow> <mo>|</mo> </mrow> <msub> <mi>ψ</mi> <mi>ν</mi> </msub> <msup> <mrow> <mo>|</mo> </mrow> <mn>2</mn> </msup> </mrow> </semantics></math> (<math display="inline"><semantics> <mrow> <mi>ν</mi> <mo>=</mo> <mn>1</mn> <mo>,</mo> <mn>2</mn> <mo>,</mo> <mn>3</mn> </mrow> </semantics></math>), with right-hand well width <math display="inline"><semantics> <mrow> <msub> <mi>L</mi> <mi>R</mi> </msub> <mo>=</mo> <mn>2</mn> </mrow> </semantics></math> nm (a) and <math display="inline"><semantics> <mrow> <msub> <mi>L</mi> <mi>R</mi> </msub> <mo>=</mo> <mn>20</mn> </mrow> </semantics></math> nm (<b>c</b>). In (<b>b</b>), the energy levels of the first three subbands as a function of the right-hand well width <math display="inline"><semantics> <msub> <mi>L</mi> <mi>R</mi> </msub> </semantics></math>. Coloured solid lines show the results obtained assuming the same effective mass for well and barrier, whereas dots represent the results obtained distinguishing the effective masses of well (<math display="inline"><semantics> <mrow> <msup> <mi>m</mi> <mo>*</mo> </msup> <mo>=</mo> <mn>0.067</mn> <msub> <mi>m</mi> <mn>0</mn> </msub> </mrow> </semantics></math>) and barrier (<math display="inline"><semantics> <mrow> <msup> <mi>m</mi> <mo>*</mo> </msup> <mo>=</mo> <mn>0.09</mn> <msub> <mi>m</mi> <mn>0</mn> </msub> </mrow> </semantics></math>).</p>
Full article ">
18 pages, 4478 KiB  
Article
Numerical Simulation of Fine Particle Migration in Loose Soil Under Groundwater Seepage Based on Computational Fluid Dynamics–Discrete Element Method
by Hongkun Yang, Yinger Deng, Hu Su, Pengjie Li, Lin Chen and Ning Wang
Water 2025, 17(5), 740; https://doi.org/10.3390/w17050740 - 3 Mar 2025
Viewed by 234
Abstract
The seepage of groundwater in loose soil causes the migration of fine particles within the soil, which can significantly contribute to slope instability and trigger a series of geological issues, such as soil erosion, landslides, and debris flow. This study employed a coupled [...] Read more.
The seepage of groundwater in loose soil causes the migration of fine particles within the soil, which can significantly contribute to slope instability and trigger a series of geological issues, such as soil erosion, landslides, and debris flow. This study employed a coupled computational fluid dynamics and discrete element method (CFD-DEM) to investigate the migration process of soil particles under groundwater seepage. It elucidated the effects of key factors, including particle size ratio, particle quantity, and weight, on the migration behavior of fine particles within porous media. The results indicated that when the particle size ratio was less than or equal to 5, over 90% of fine particles accumulated on the surface of the medium. Additionally, an increase in the weight or quantity of fine particles intensified their accumulation. However, when the particle size ratio exceeded five, it became the dominant factor affecting displacement. Under the same weight conditions, the larger the particle size ratio, the longer the particle migration distance. Compared to a particle size ratio of 3, the accumulation percentages of fine particles with a particle size ratio of 20 increased by 26.88% and 31.46% in the middle and tail sections, respectively. Full article
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Figure 1

Figure 1
<p>Schematic diagram of the initial setting of the model: (<b>a</b>) Particles; (<b>b</b>) fluid element.</p>
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<p>The relationship between drag coefficient and Reynolds number.</p>
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<p>Fine particle migration and flow characteristics of fluid at different times: (<b>a</b>) Time = 0.1 s; (<b>b</b>) Time = 0.7 s; (<b>c</b>) Time = 2.0 s. (<b>d</b>) Time = 0.1 s; (<b>e</b>) Time = 0.7 s; (<b>f</b>) Time = 2.0 s.</p>
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<p>Velocity variation in monitored particles in the Z direction: (<b>a</b>) Series A; (<b>b</b>) Series B; (<b>c</b>) Series C.</p>
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<p>Displacement variation in monitored particles in the Z direction: (<b>a</b>) Series A; (<b>b</b>) Series B; (<b>c</b>) Series C.</p>
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<p>The clogging location of monitored particles in AG-2: (<b>a</b>) Time = 0.1 s; (<b>b</b>) Time = 0.7 s; (<b>c</b>) Time = 2.0 s. The clogging location of monitored particles in BG-4: (<b>d</b>) Time = 0.1 s; (<b>e</b>) Time = 0.7 s; (<b>f</b>) Time = 2.0 s. The clogging location of monitored particles in CG-5: (<b>g</b>) Time = 0.1 s; (<b>h</b>) Time = 0.7 s; (<b>i</b>) Time = 2.0 s.</p>
Full article ">Figure 6 Cont.
<p>The clogging location of monitored particles in AG-2: (<b>a</b>) Time = 0.1 s; (<b>b</b>) Time = 0.7 s; (<b>c</b>) Time = 2.0 s. The clogging location of monitored particles in BG-4: (<b>d</b>) Time = 0.1 s; (<b>e</b>) Time = 0.7 s; (<b>f</b>) Time = 2.0 s. The clogging location of monitored particles in CG-5: (<b>g</b>) Time = 0.1 s; (<b>h</b>) Time = 0.7 s; (<b>i</b>) Time = 2.0 s.</p>
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<p>Percentage of fine particles in section Q at different times: (<b>a</b>) Series A; (<b>b</b>) Series B; (<b>c</b>) Series C.</p>
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<p>Accumulation of fine particles at different depths.</p>
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12 pages, 6163 KiB  
Article
Study on the Wellbore Instability Mechanism in the Longtan Formation with Soft/Hard Thin Interlayers in the South Sichuan Basin
by Jianhua Guo, Yu Sang, Beiqiao Meng, Lianbin Xia, Yangsong Wang, Chengyu Ma, Tianyi Tan and Bin Yang
Processes 2025, 13(3), 727; https://doi.org/10.3390/pr13030727 - 3 Mar 2025
Viewed by 245
Abstract
The lithology of the transitional facies of the Longtan Formation in the southern Sichuan Basin is complex, with soft/hard thin interlayers of mud shale, sandstone, and limestone. Drilling this layer often results in wellbore instability, including frequent blockages, tripping resistance, and sticking. This [...] Read more.
The lithology of the transitional facies of the Longtan Formation in the southern Sichuan Basin is complex, with soft/hard thin interlayers of mud shale, sandstone, and limestone. Drilling this layer often results in wellbore instability, including frequent blockages, tripping resistance, and sticking. This study focuses on a shale gas block in the Longtan Formation in Zigong, where a geomechanical profile was established by integrating ground stress, rock parameter tests, and logging data. The critical collapse pressure was calculated, and wellbore instability was simulated using the Mohr–Coulomb failure criterion and the discrete element method. Results indicate significant variability in the mechanical strength of the rocks, with notable longitudinal heterogeneity and a high risk of wellbore instability. The critical collapse pressure equivalent density ranges from 1.05–1.69 g/cm3. Under low-density conditions, wellbore expansion and reduction coexist due to local shear and dropping. Even when the drilling fluid density exceeds the collapse pressure equivalent, stress imbalance can still cause localized dropping at lithologic interfaces. These findings offer valuable insights into the mechanical mechanisms behind wellbore instability in formations with soft/hard thin interlayers and provide guidance for the prevention and control of wellbore instability and associated risks. Full article
(This article belongs to the Section Energy Systems)
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Figure 1

Figure 1
<p>Lithology and logging curve diagram of the Longtan Formation in a shale gas well in southern Sichuan.</p>
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<p>Uniaxial compressive strength test results of Longtan Formation.</p>
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<p>Geomechanical parameters and collapse pressure equivalent density interpretation results of the Longtan Formation.</p>
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<p>Discrete element model of wellbore instability in the Longtan Formation.</p>
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<p>Simulation results of wellbore instability in the thin interbedded formations of the Longtan Formation: (<b>a</b>) instability characteristics of different lithological sections; (<b>b</b>) distribution of shear failure bands in particle cementation.</p>
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<p>Variation trends of wellbore instability characteristics under different drilling fluid densities.</p>
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27 pages, 25794 KiB  
Article
Numerical Investigation of the Influence of Temperature on Fluidization Behavior: Importance of Particle Collision Parameters and Inter-Particle Forces
by Milan Mihajlović, Juan G. Ramírez, Ildefonso Campos Velarde, Martin Van Sint Annaland and Ivo Roghair
Fluids 2025, 10(3), 60; https://doi.org/10.3390/fluids10030060 - 27 Feb 2025
Viewed by 213
Abstract
Fluidized bed reactors (FBRs) are integral to various industries due to their exceptional capability in facilitating efficient gas–solid interactions, resulting in superior mixing and heat and mass transfer. This research delves into the impact of temperature on fluidization dynamics, particularly focusing on the [...] Read more.
Fluidized bed reactors (FBRs) are integral to various industries due to their exceptional capability in facilitating efficient gas–solid interactions, resulting in superior mixing and heat and mass transfer. This research delves into the impact of temperature on fluidization dynamics, particularly focusing on the collisional properties of particles within the bed. The investigation builds upon foundational research, notably Geldart’s classification of fluidization regimes and recent advancements in high-temperature experimental techniques, such as High-Temperature Endoscopic-Laser particle image velocimetry/digital image analysis. To explore these temperature effects, a coupled Discrete Element Method and Computational Fluid Dynamics (cfd–dem) model was employed. This approach enables a detailed examination of gas–particle and particle–particle interactions under varying temperature conditions. The simulations in this study explore the friction coefficient, as well as changes in both tangential and normal restitution coefficients, which affect the fluidization behavior. These changes were systematically analyzed to determine their influence on minimum fluidization velocity and bubble formation. The numerical results are compared with experimental data from high-temperature fluidization studies, highlighting the necessity of incorporating inter-particle forces to fully capture the observed phenomena. The findings underscore the critical role of particle collisional properties in high-temperature fluidization and suggest the potential increasing role of inter-particle forces. Overall, this paper provides new insights into the complex dynamics of fluidized beds at elevated temperatures, emphasizing the need for further experimental–numerical research to enhance the reliability and understanding of these systems in industrial applications. Full article
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Figure 1

Figure 1
<p>Minimum fluidization conditions as a function of selected particle collisional parameters.</p>
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<p>Energy states of the system affected by changes in friction coefficient. (<b>a</b>) Average energy dissipation due to friction contact—sliding. (<b>b</b>) Average potential energy of the system.</p>
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<p>Screenshots of single-bubble injection for different normal restitution coefficients.</p>
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<p>Probability density function of the porosity for a single-bubble injection at different values of normal restitution coefficient and friction coefficient.</p>
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<p>Screenshots of single-bubble injection for different friction coefficients.</p>
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<p>Time-averaged solid circulation patterns for different normal restitution coefficients.</p>
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<p>Probability density function of the porosity for the fluidization conditions presented for different values of normal restitution coefficient and friction coefficient.</p>
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<p>Bubble diameter over height for the fluidization conditions presented for different values of normal restitution coefficient and friction coefficient.</p>
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<p>Number of bubbles depending on their size for the fluidization conditions presented for different values of normal restitution coefficient and friction coefficient. (<b>a</b>) Number of bubbles depending on their size <math display="inline"><semantics> <mrow> <mo>(</mo> <msub> <mi>e</mi> <mi>n</mi> </msub> <mo>)</mo> </mrow> </semantics></math>. (<b>b</b>) Number of bubbles depending on their size <math display="inline"><semantics> <mrow> <mo>(</mo> <msub> <mi>μ</mi> <mi>f</mi> </msub> <mo>)</mo> </mrow> </semantics></math>.</p>
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<p>Time-averaged solid circulation patterns for different friction coefficients.</p>
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<p>Time-averaged solid circulation patterns for different tangential restitution coefficients.</p>
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<p>Particle velocity at different heights of the bed for different tangential restitution coefficients.</p>
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<p>Bubble size and porosity distributions for the cases with different tangential restitution coefficients. (<b>a</b>) Bubble diameter over height for cases with different <math display="inline"><semantics> <msub> <mi>e</mi> <mi>t</mi> </msub> </semantics></math>. (<b>b</b>) PDF of porosity for cases with different <math display="inline"><semantics> <msub> <mi>e</mi> <mi>t</mi> </msub> </semantics></math>.</p>
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<p>Minimum fluidization velocity as a function of temperature of 528 <math display="inline"><semantics> <mi mathvariant="sans-serif">μ</mi> </semantics></math><math display="inline"><semantics> <mi mathvariant="normal">m</mi> </semantics></math> glass beads fluidized with different gases compared to literature correlations from Ergun, Wen and Yu, and Carman-Kozeny [<a href="#B37-fluids-10-00060" class="html-bibr">37</a>,<a href="#B38-fluids-10-00060" class="html-bibr">38</a>,<a href="#B44-fluids-10-00060" class="html-bibr">44</a>,<a href="#B45-fluids-10-00060" class="html-bibr">45</a>,<a href="#B46-fluids-10-00060" class="html-bibr">46</a>]. Experimental data as reported by Campos Velarde [<a href="#B8-fluids-10-00060" class="html-bibr">8</a>].</p>
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<p>Minimum fluidization velocity simulated at different temperatures, changing the gas and particle properties, in comparison with experimental results.</p>
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<p>(<b>a</b>) Experimental behavior of 400–600 <math display="inline"><semantics> <mi mathvariant="sans-serif">μ</mi> </semantics></math><math display="inline"><semantics> <mi mathvariant="normal">m</mi> </semantics></math> glass beads fluidized with nitrogen at 473 <math display="inline"><semantics> <mi mathvariant="normal">K</mi> </semantics></math> and a He:Ar mixture at 323 <math display="inline"><semantics> <mi mathvariant="normal">K</mi> </semantics></math> with the same excess velocity. (<b>b</b>) Time-averaged solid velocity profiles at H = 0.3 <math display="inline"><semantics> <mi mathvariant="normal">m</mi> </semantics></math>; adapted from Campos Velarde [<a href="#B8-fluids-10-00060" class="html-bibr">8</a>].</p>
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<p>(<b>a</b>) Time-averaged solid circulation patterns for two cases of same gas properties. (<b>b</b>) Particle velocity at the middle of the bed for the same gas properties at different temperatures.</p>
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<p>Bubble size and porosity distributions for the cases with same gas properties. (<b>a</b>) Bubble diameter over height for cases with same gas properties. (<b>b</b>) PDF of porosity for cases with same gas properties.</p>
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<p>Solid circulation patterns of 400–600 <math display="inline"><semantics> <mi mathvariant="sans-serif">μ</mi> </semantics></math><math display="inline"><semantics> <mi mathvariant="normal">m</mi> </semantics></math> glass beads fluidized with nitrogen at the same excess velocity at different operating temperatures; experiments by Campos Velarde [<a href="#B8-fluids-10-00060" class="html-bibr">8</a>].</p>
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<p>Time-averaged lateral solid velocity profiles at two positions above the porous plate fluidized with nitrogen at different temperatures; experiments by Campos Velarde [<a href="#B8-fluids-10-00060" class="html-bibr">8</a>].</p>
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<p>Time-averaged glass particle phase velocity for different temperatures.</p>
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<p>Particle velocity profiles at different heights of the bed for different temperatures.</p>
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<p>Bubble size distributions for glass particle fluidization for different temperatures with(out) <span class="html-small-caps">vdw</span> force. (<b>a</b>) Bubble diameter over height for glass particle fluidization. (<b>b</b>) Bubble diameter over height for glass particle fluidization for 573 <math display="inline"><semantics> <mi mathvariant="normal">K</mi> </semantics></math>.</p>
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<p>PDF of porosity for cases with different temperatures and variable gas properties.</p>
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20 pages, 5623 KiB  
Article
A Study of the Scale Dependency and Anisotropy of the Permeability of Fractured Rock Masses
by Honglue Qian and Yanyan Li
Water 2025, 17(5), 697; https://doi.org/10.3390/w17050697 - 27 Feb 2025
Viewed by 200
Abstract
Affected by discontinuities, the hydraulic properties of rock masses are characterized by significant scale dependency and anisotropy. Sampling a rock mass at any scale smaller than the representative elementary volume (REV) size may result in incorrect characterization and property upscaling. Here, a three-dimensional [...] Read more.
Affected by discontinuities, the hydraulic properties of rock masses are characterized by significant scale dependency and anisotropy. Sampling a rock mass at any scale smaller than the representative elementary volume (REV) size may result in incorrect characterization and property upscaling. Here, a three-dimensional discrete fracture network (DFN) model was built using the joint data obtained from a dam site in southwest China. A total of 504 two-dimensional sub-models with sizes ranging from 1 m × 1 m to 42 m × 42 m were extracted from the DFN model and then used as geometric models for equivalent permeability tensor calculations. A series of steady-state seepage numerical simulations were conducted for these models using the finite element method. We propose a new method for estimating the REV size of fractured rock masses based on permeability. This method provides a reliable estimate of the REV size by analyzing the tensor characteristic of the directional permeability, as well as its constant characteristic beyond the REV size. We find that the hydraulic REV sizes in different directions vary from 6 to 36 m, with the maximum size aligning with the average orientation of joint sets and the minimum along the angle bisector of intersecting joints. Additionally, the REV size is negatively correlated with the average trace length of the two intersecting joint sets. We find that the geometric REV size, determined by the joint connectivity and density, falls into the range of the hydraulic REV size. The findings could provide guidance for determining the threshold values of numerical rock mass models. Full article
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<p>Development of joints in the rock mass around the tunnel. A total of 128 joints were obtained from a relatively homogeneous and undisturbed zone with a length of 80 m. Set 1 refers to steeply dipping joints, set 2 consists of shallow-dipping joints, and set 3 consists of moderately dipping joints. (<b>a</b>) Structural plane measurements; (<b>b</b>) upper-hemisphere and equal area projection of the joint orientations.</p>
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<p>The 3D DFN model and distribution of joint traces on the <span class="html-italic">y</span>–<span class="html-italic">z</span> plane with <span class="html-italic">x</span> = 50 m. The <span class="html-italic">x</span>-direction refers to the west, and the <span class="html-italic">y</span>-direction refers to the south. (<b>a</b>) The 3D analysis area; (<b>b</b>) the 2D analysis area. (yellow lines, red lines, and blue lines represent the joints in set 1, set 2, and set 3, respectively).</p>
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<p>Permeability ellipse and principal direction.</p>
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<p>Fracture network models used for validation and their boundary conditions. (<b>a</b>) Single-fracture model; (<b>b</b>) pressure distribution in the single-fracture model; (<b>c</b>) double-fracture model; and (<b>d</b>) pressure distribution in the double-fracture model. (The arrow lines represent the streamline).</p>
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<p>Extraction of 2D sub-models used for fluid flow simulations. (<b>a</b>) DFN original network; (<b>b</b>) sub-models.</p>
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<p>Schematic diagram of the rotation of the sub-models.</p>
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<p>Boundary conditions for the numerical models (yellow lines, red lines, and blue lines represent the joints in set 1, set 2, and set 3, respectively).</p>
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<p>Variation in the directional permeability <span class="html-italic">k</span> with the model size.</p>
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<p>Fluid pressure distribution in different sized models when <span class="html-italic">θ</span> = 0°.</p>
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<p>Schematic diagram of the REV size estimation.</p>
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<p>Values of <span class="html-italic">k</span><sub>AVG</sub> and <span class="html-italic">k</span> in different flow directions and the fitted permeability ellipses. (<b>a</b>) <span class="html-italic">CV</span> = 0.20; (<b>b</b>) <span class="html-italic">CV</span> = 0.19; (<b>c</b>) <span class="html-italic">CV</span> = 0.18; (<b>d</b>) <span class="html-italic">CV</span> = 0.17; (<b>e</b>) <span class="html-italic">CV</span> = 0.16; (<b>f</b>) <span class="html-italic">CV</span> = 0.15; (<b>g</b>) <span class="html-italic">CV</span> = 0.14; (<b>h</b>) <span class="html-italic">CV</span> = 0.13; (<b>i</b>) <span class="html-italic">CV</span> = 0.12; (<b>j</b>) <span class="html-italic">CV</span> = 0.11; (<b>k</b>) <span class="html-italic">CV</span> = 0.10.</p>
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<p>Values of <span class="html-italic">RMS</span><sub>Norm</sub> for different <span class="html-italic">CV</span> values.</p>
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<p>(<b>a</b>) <span class="html-italic">L</span><sub>REV</sub> size, (<b>b</b>) permeability, and (<b>c</b>) fitted ellipse in each seepage direction for <span class="html-italic">CV</span> = 0.18.</p>
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<p>Variations in the <span class="html-italic">AI</span><sub>p</sub> and <span class="html-italic">RMS</span><sub>Norm</sub> with model sizes.</p>
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<p>Variation in the permeability <span class="html-italic">k</span> in each direction with model sizes.</p>
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<p>Distributions of the average trace length and orientation for the three joint sets in the 2D DFN model.</p>
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<p>Variations in joint connectivity, <span class="html-italic">P</span><sub>20</sub>, and <span class="html-italic">P</span><sub>21</sub> with model sizes.</p>
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17 pages, 10499 KiB  
Article
Numerical Investigation into the Runout Dynamics of Reservoir Landslides: Insights from the Yanguan Landslide
by Hao Fang, Bing Li, Kai Liu and Yaobin Meng
Water 2025, 17(5), 695; https://doi.org/10.3390/w17050695 - 27 Feb 2025
Viewed by 219
Abstract
Understanding the dynamic behavior of landslides is essential for effective risk assessment. This study examines the Yanguan landslide, which occurred on 29 October 2017, in the Three Gorges Reservoir (TGR) region of China. Due to its unique capability in modeling discontinuum behaviors during [...] Read more.
Understanding the dynamic behavior of landslides is essential for effective risk assessment. This study examines the Yanguan landslide, which occurred on 29 October 2017, in the Three Gorges Reservoir (TGR) region of China. Due to its unique capability in modeling discontinuum behaviors during landslide fragmentation, the discrete element method was utilized to analyze the movement characteristics of this landslide. The investigation began with a field survey to assess the geological features and failure mechanism of the landslide, which indicates that the landslide was likely triggered by prolonged variations in reservoir water levels and heavy rainfall preceding the event. Following this, a three-dimensional numerical model of the landslide was constructed using pre- and post-event terrain data. The accuracy of the numerical model was validated by comparing its simulation results with field survey data. Finally, the landslide’s movement behavior and energy transformation were analyzed based on the validated model. This work can enhance landslide risk assessment by quantifying dynamic parameters critical for impact prediction, further provide a scientific basis for the study of the landslides in the TGR area, and contribute to disaster prevention. Full article
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<p>Geological background of the research area: (<b>a</b>) Geomorgraphy of the Zigui County; (<b>b</b>) Lithologic map of the study area.</p>
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<p>Images taken after the Yanguan landslide occurrence: (<b>a</b>) Post-failure form of the landslide; (<b>b</b>) Destroyed house; (<b>c</b>) Ruined road; (<b>d</b>) Cracks.</p>
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<p>Geological profile of the Yanguan landslide.</p>
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<p>Water level variation and rainfall distribution in YGR from 2003 to 2019.</p>
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<p>Morphology and movement deposits characteristics of the Yanguan landslide.</p>
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<p>Numerical modeling flowchart of the PFC model of the Yanguan landslide.</p>
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<p>Numerical model of the landslide: (<b>a</b>) PFC model; (<b>b</b>) Terrain and stratum distribution.</p>
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<p>A 3-D uniaxial compression test of the sliding mass using PFC.</p>
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<p>Distribution of the Yanguan landslide velocity: (<b>a</b>) t = 0 s; (<b>b</b>) t = 10 s; (<b>c</b>) t = 20 s; (<b>d</b>) t = 30 s; (<b>e</b>) t = 40 s; (<b>f</b>) t = 50 s; (<b>g</b>) t = 60 s; (<b>h</b>) t = 70 s; (<b>i</b>) t = 80 s.</p>
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<p>Average velocities of the Yanguan landslide: (<b>a</b>) average velocities of the road and sliding mass; (<b>b</b>) average velocities monitored at different locations of landslide.</p>
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<p>Average displacements of the Yanguan landslide: (<b>a</b>) average displacements of the road and sliding mass; (<b>b</b>) average displacements monitored at different locations of landslide.</p>
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<p>Energy conversions during the landslide runout process.</p>
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20 pages, 10891 KiB  
Article
Calibration of DEM Polyhedron Model for Wheat Seed Based on Angle of Repose Test and Semi-Resolved CFD-DEM Coupling Simulation
by Longbao Wang, Hanyu Yang, Zhinan Wang, Qingjie Wang, Caiyun Lu, Chao Wang and Jin He
Agriculture 2025, 15(5), 506; https://doi.org/10.3390/agriculture15050506 - 26 Feb 2025
Viewed by 110
Abstract
The shape of particles is a critical determinant that significantly influences the accuracy of discrete element simulations. To reduce the discrepancies between the discrete element model of wheat seeds and the actual particle shapes, and to enhance the accuracy of Computational Fluid Dynamics-Discrete [...] Read more.
The shape of particles is a critical determinant that significantly influences the accuracy of discrete element simulations. To reduce the discrepancies between the discrete element model of wheat seeds and the actual particle shapes, and to enhance the accuracy of Computational Fluid Dynamics-Discrete Element Method (CFD-DEM) coupling simulations in gas–solid two-phase flow studies, We employed laser scanning and inverse modeling techniques to develop a three-dimensional (3D) reconstruction of the wheat seed. Subsequently, we employed Rocky DEM simulation software to develop a polyhedron model and an Angle of Repose (AOR) test model. The interval range of material parameters was determined through a series of physical experiments and subsequently employed to delineate the high and low levels of parameters for the simulation tests. The simulation parameters were calibrated using data from AOR simulation tests. The Plackett–Burman test, Steepest-Ascent test, and Box–Behnken test were conducted sequentially to determine the optimal parameter configuration. A test bench for wheat gas-assisted seeding was constructed, and a semi-resolved CFD-DEM coupling simulation model was developed to perform comparative analysis. The results demonstrated that the optimal parameters were as follows: the static friction coefficient of wheat seed was 0.15, the dynamic friction coefficient of wheat seed was 0.11694, and the dynamic friction coefficient between wheat seed and resin was 0.0797. In this scenario, the relative error of AOR was 2.3% and the maximum relative error of ejection velocity observed was 4.1%. The reliability of the polyhedron model and its calibration parameters was rigorously validated, thereby providing a robust reference for studies on gas–solid two-phase flows. Full article
(This article belongs to the Section Agricultural Technology)
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<p>Triaxial size measurement position.</p>
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<p>Triaxial size distribution of LY502 wheat seed particles.</p>
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<p>Construction process of wheat seed discrete element model.</p>
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<p>TMS-Touch texture analyzer: 1. control interface; 2. loadcell slider; 3. mechanical transducer; 4. metallic probe; 5. wheat seed.</p>
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<p>Bulk density measurement: 1. the hopper; 2. wheat seeds; 3. conventional cylindrical container; 4. the tray; 5. control interface; 6. electronic weighing scale.</p>
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<p>Determination of the friction coefficient: 1. mechanical transducer; 2. sliding block mechanism; 3. seed plate; 4. liquid crystal display panel; 5. operational control panel.</p>
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<p>Determination of the collision restitution coefficient: 1. supplemental lighting; 2. high-speed camera; 3. portable computer; 4. graph paper; 5. resin plate.</p>
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<p>The procedure for determining the angle of repose: 1. the hopper; 2. ambient illumination source; 3. camera device; 4. wheat seeds; 5. resin plate; 6. the tray.</p>
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<p>AOR test model: 1. particle inlet; 2. seeds in free fall; 3. stock pile; 4. resin plate.</p>
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<p>The procedure for determining the AOR.</p>
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<p>Air-assisted seeding test platform: 1. pressure control valve; 2. air compressor; 3. background panel; 4. seeding device; 5. portable computer; 6. high-speed camera; 7. illumination unit; 8. digital barometer; 9. direct current power supply.</p>
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<p>Semi-resolved CFD-DEM coupling simulation model.</p>
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<p>Pareto chart illustrating the normalization effect.</p>
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<p>Response surface of AOR relative error.</p>
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21 pages, 4966 KiB  
Article
Influence of Particle Shape and Size on Gyratory Crusher Simulations Using the Discrete Element Method
by Manuel Moncada, Christian Rojas, Patricio Toledo, Cristian G. Rodríguez and Fernando Betancourt
Minerals 2025, 15(3), 232; https://doi.org/10.3390/min15030232 - 26 Feb 2025
Viewed by 125
Abstract
Gyratory crushers are fundamental machines in aggregate production and mineral processing. Discrete Element Method (DEM) simulations offer detailed insights into the performance of these machines and serve as a powerful tool for their design and analysis. However, these simulations are computationally intensive due [...] Read more.
Gyratory crushers are fundamental machines in aggregate production and mineral processing. Discrete Element Method (DEM) simulations offer detailed insights into the performance of these machines and serve as a powerful tool for their design and analysis. However, these simulations are computationally intensive due to the large number of particles involved and the need to account for particle breakage. This study aims to investigate the effect of particle shape and size distribution on the performance of a DEM model of a gyratory crusher. The selected study case corresponds to a primary gyratory crusher operating in a copper processing industry. As particle shapes, spheres and polyhedrons are used with a particle replacement scheme. This study utilizes two different size distributions, with variations also applied to the minimum particle size. The results are analyzed in terms of the impact of these factors on the power draw, mass flow, and product size distribution for each of the combinations explained. The findings demonstrate that particle shape primarily influences the product size distribution, whereas variations in particle size distribution have a pronounced effect on power draw, mass flow rate, and product size distribution. Based on the results, recommendations are provided regarding the selection of the minimum particle size. It is concluded that the minimum particle size should not exceed a third of the closed-side setting to ensure accurate and reliable simulation outcomes. Full article
(This article belongs to the Special Issue Process Modelling and Applications for Aggregate Production)
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<p>Particle size distribution of the feed of the DEM simulations.</p>
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<p>Setup of the DEM model of the gyratory crusher.</p>
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<p>Images of the DEM simulation with different feed cut size. Sample 1: (<b>a</b>) 50 mm, (<b>b</b>) 70 mm, (<b>c</b>) 122 mm, (<b>d</b>) 300 mm, (<b>e</b>) 500 mm, (<b>f</b>) 650 mm, (<b>g</b>) 750 mm, (<b>h</b>) 1000 mm. Sample 2: (<b>i</b>) 100 mm, (<b>j</b>) 500 mm, (<b>k</b>) 850 mm, (<b>l</b>) 1000 mm.</p>
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<p>Particle size distribution of the product with S1 showing the effect of the feed cut size.</p>
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<p>Particle size distribution of the product showing the effect of the feed cut size with S2 and 4P as particle shape.</p>
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<p>Effect of the feed cut size: (<b>a</b>) crushing power, (<b>b</b>) mass flow rate of the product, and (<b>c</b>) particle mass inside the crushing chamber.</p>
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<p>Particle size distribution of the product showing the effect of the minimum particle size with S2 and four polyhedron as particle shape.</p>
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<p>Particle size distribution of the product for cases A, B, and C.</p>
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<p>Particle size distribution of the product for cases A, D, and D’.</p>
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<p>Effect of the minimum particle size: (<b>a</b>) Crushing power, and (<b>b</b>) mass flow rate of the product.</p>
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<p>(<b>a</b>) Images of the DEM simulation with different particle size distribution with S2: (<b>a</b>) 4P fine, (<b>b</b>) 4P base, (<b>c</b>) 4P coarse, (<b>d</b>) S fine, (<b>e</b>) S base, (<b>f</b>) S coarse.</p>
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<p>Particle size distribution of the product showing the effect of changing the feed size distribution.</p>
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<p>Effect of the particle size distribution: (<b>a</b>) Crushing power, and (<b>b</b>) mass flow rate of the product.</p>
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<p>Images of the DEM simulation with different particle shapes and size. (<b>a</b>) S, (<b>b</b>) 1P, (<b>c</b>) 4P, (<b>d</b>) CP, (<b>e</b>) SP. Particle shapes are presented at the top and a close-up of the particles at the bottom.</p>
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<p>Particle size distribution of the product showing the effect of the particle shape.</p>
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<p>Effect of the particle shape: (<b>a</b>) Crushing power, and (<b>b</b>) mass flow rate of the product.</p>
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<p>Elapsed simulation times: (<b>a</b>) particle size (each <span class="html-italic">x</span>-axis for the respective case), (<b>b</b>) particle shape. S* was simulated with a different GPU card.</p>
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16 pages, 5727 KiB  
Article
Numerical Analysis of Influence Mechanism of Orifice Eccentricity on Silo Discharge Rate
by Yinglong Wang, Yanlong Han, Anqi Li, Hao Li, Haonan Gao, Ze Sun, Shouyu Ji, Zhuozhuang Li and Fuguo Jia
Agriculture 2025, 15(5), 490; https://doi.org/10.3390/agriculture15050490 - 25 Feb 2025
Viewed by 211
Abstract
Eccentric silo is an extremely common type of silo, but it is still unclear how to accurately control the discharge by adjusting eccentric orifices, limiting the application and development of eccentric silo. In this study, the rice particle discharging process on silos with [...] Read more.
Eccentric silo is an extremely common type of silo, but it is still unclear how to accurately control the discharge by adjusting eccentric orifices, limiting the application and development of eccentric silo. In this study, the rice particle discharging process on silos with different eccentricities was simulated by the discrete element method (DEM), and the influence mechanism of orifice eccentricity on silo discharge rate was analyzed. The results show that eccentricity has a direct influence on the particle volume fraction and vertical velocity that determine the discharge rate of the silo. In fully eccentric silo, it is not easy for particle flow to achieve balance, particles will pass through outlet with more kinetic energy. Moreover, continuous force network cannot be formed between particles with shear resistance, resulting in weak interlocking action between particles. The orientation of particle in fully eccentric silo is more vertical, especially near the silo wall, which will produce larger local particle volume fraction above the orifice. When the eccentricity exceeds the critical eccentricity, the sparse flow area on the discharge orifice becomes larger, and the particle acceleration area increases accordingly. Research findings may offer valuable insights for the accurate control of discharge rate of eccentric silo, as well as for optimizing silo design. Full article
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<p>The rice particle modeling process: (<b>a</b>) photo of rice grains, (<b>b</b>) ellipsoid model of rice, and (<b>c</b>) 3D rice particle model in simulation.</p>
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<p>Silo modeling process: (<b>a</b>) the three-dimensional schematic diagram of eccentric silo and (<b>b</b>) the division of silo-monitoring area.</p>
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<p>Schematic diagram of self-built silo-unloading platform.</p>
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<p>Schematic diagram of rice particle flow pattern in silo during discharging process in simulation and experiment.</p>
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<p>Experimental and simulation comparison of particle discharge rate change under different orifice eccentricity.</p>
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<p>The volume fraction of particles at the discharge orifice of each eccentric silo.</p>
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<p>Vertical velocity distribution of particles at discharge orifices of different eccentric silos.</p>
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<p>The movement trajectories of particle in the steady state of flow: (<b>a</b>) concentric silo and (<b>b</b>) fully eccentric silo.</p>
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<p>The normal contact force network in concentric silo and fully eccentric silo at the same moment.</p>
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<p>Schematic diagram for determining orientation angle of long axis of rice grains.</p>
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<p>Changes trend of rice orientation angles at different heights above the orifice: (<b>a</b>) concentric silo and (<b>b</b>) fully eccentric silo.</p>
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<p>Shear rate nephogram in silos with different eccentricities (<span class="html-italic">e</span>): (<b>a</b>) 0; (<b>b</b>) 0.1; (<b>c</b>) 0.2; (<b>d</b>) 0.5; (<b>e</b>) 0.7; (<b>f</b>) 0.8; (<b>g</b>) 0.84; (<b>h</b>) 0.85; and (<b>i</b>) 0.857.</p>
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<p>Particle coordination number nephogram in silos with different eccentricities (<span class="html-italic">e</span>): (<b>a</b>) 0; (<b>b</b>) 0.1; (<b>c</b>) 0.2; (<b>d</b>) 0.5; (<b>e</b>) 0.7; (<b>f</b>) 0.8; (<b>g</b>) 0.84; (<b>h</b>) 0.85; and (<b>i</b>) 0.857.</p>
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18 pages, 7859 KiB  
Article
Study on Sand-Steel Interface Shear Test Method and Mechanism
by Xiaofei Hu, Long Yu, Yunrui Han and Qing Yang
J. Mar. Sci. Eng. 2025, 13(3), 407; https://doi.org/10.3390/jmse13030407 - 22 Feb 2025
Viewed by 228
Abstract
Soil-structure interface properties play an essential role in geotechnical engineering. The interface shear test is widely used to measure the interface properties. However, in the traditional interface shear test (TIST), distribution of shear stresses along contact surface is not uniform due to boundary [...] Read more.
Soil-structure interface properties play an essential role in geotechnical engineering. The interface shear test is widely used to measure the interface properties. However, in the traditional interface shear test (TIST), distribution of shear stresses along contact surface is not uniform due to boundary effects. Thus, average mechanical response of the whole interface measured by TIST cannot be used to evaluate interface friction properties. In this paper, a novel interface shear apparatus (MIDST) is presented to investigate the shear behaviours of the soil-structure interface. A series of shear tests were conducted on Fujian standard sand-steel interface. Two shear force sensors simultaneously monitor the shear force along the interface: a pre-embedded sensor inside the interface/steel plate (responding to MIDST), while the other outside the interface (responding to TIST). Laboratory test results show that the pre-embedded internal sensor successfully detects the weakening characteristics of the interface, while the external sensor monitors the hardening law. The interface shear strength measured by internal sensor is significantly higher than that monitored by external sensor. A commercial DEM software Version 5.0, Particle Flow Code in Two Dimensions (PFC2D), is employed to study the soil-structure interaction mechanism, and numerical test results show that the main reasons for the internal and external differences are the uneven shear stress distribution at the soil-structure interface and the boundary effect. In addition, numerical test results agree with the laboratory test results, indicating that the shear behaviours monitored by MIDST are relatively accurate and can provide a reference for engineering design. Full article
(This article belongs to the Section Ocean Engineering)
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<p>Limitations of traditional interface direct shear test and the modification of the new method.</p>
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<p>Schematic of the Modified Interface Direct Shear Test (MIDST) apparatus.</p>
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<p>Schematic of the interface-measurement block.</p>
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<p>Friction measurement of the linear slide rail.</p>
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<p>Specimen preparation.</p>
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<p>Shear behaviour of the interface shear verification tests under different normal stress levels: (<b>a</b>) shear force and (<b>b</b>) shear stress curves.</p>
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<p>Load analysis for two test methods, both for those using internal and external sensors.</p>
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<p>Calibration of <span class="html-italic">f</span><sub>r</sub>.</p>
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<p>Peak shear stress vs. normal stress.</p>
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<p>Numerical modelling of soil-structure interface direct shear test with the DEM method.</p>
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<p>Relationship between the shear stress and displacement at different positions on the soil-structure interface: (<b>a</b>) shear stress of different blocks and (<b>b</b>) particle movement mechanism.</p>
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<p>A comparison of different test methods between DEM and experimental results: (<b>a</b>) DEM simulation results and (<b>b</b>) laboratory test results.</p>
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<p>Displacement nephogram of the soil particle obtained by PFC<sup>2D</sup>: (<b>a</b>) <span class="html-italic">x</span>-direction; (<b>b</b>) <span class="html-italic">y</span>-direction; (<b>c</b>) particles move direction.</p>
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<p>Displacement nephogram of the soil particle obtained by PFC<sup>2D</sup>: (<b>a</b>) <span class="html-italic">x</span>-direction; (<b>b</b>) <span class="html-italic">y</span>-direction; (<b>c</b>) particles move direction.</p>
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15 pages, 5907 KiB  
Article
Markov-Chain-Based Statistic Model for Predicting Particle Movement in Circulating Fluidized Bed Risers
by Yaming Zhuang
Processes 2025, 13(3), 614; https://doi.org/10.3390/pr13030614 - 21 Feb 2025
Viewed by 269
Abstract
To increase the calculation speed of the computational fluid dynamics (CFD)-based simulation for the gas–solid flow in fluidized beds, a Markov chain model (MCM) was developed to simulate the particle movement in a two-dimensional (2D) circulating fluidized bed (CFB) riser. As a statistic [...] Read more.
To increase the calculation speed of the computational fluid dynamics (CFD)-based simulation for the gas–solid flow in fluidized beds, a Markov chain model (MCM) was developed to simulate the particle movement in a two-dimensional (2D) circulating fluidized bed (CFB) riser. As a statistic model, the MCM takes the results obtained from a CFD–discrete element method (DEM) as samples for calculating transition probability matrixes of particle movement. The transition probability matrixes can be directly used to describe the macroscopic regularities of particle movement and further used to simulate the particle motion combined with the Monte Carlo method. Particle distribution snapshots, residence time distribution (RTD), and mixing obtained from both MCM and CFD-DEM are compared. The results indicate that the MCM offers a computational speed that is approximately 100 times faster than that of the CFD-DEM. The discrepancy in the mean particle residence time, as computed by the two models, is under 2%. Furthermore, the MCM provides an accurate depiction of time-averaged particle motion. In sum, the MCM can well describe the time-averaged particle mixing compared to the CFD-DEM. Full article
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<p>Discretized CFB riser with <span class="html-italic">n·m</span> cells (space states).</p>
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<p>Diagram of the transitions of Markov chain states.</p>
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<p>Bar graphs of the transition probability matrixes. Fluidized air velocity: (<b>a</b>) 3 m/s; (<b>b</b>) 4 m/s; (<b>c</b>) 6 m/s; (<b>d</b>) 8 m/s.</p>
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<p>Probability distribution for the positions of a batch of particles after being put into the inlet of the CFB riser. Fluidization velocity of 3 m/s at time (<b>a</b>) 0.25 s; (<b>b</b>) 0.5 s; and (<b>c</b>) 0.75 s.</p>
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<p>Probability distribution for the positions of a batch of particles after being put into the inlet of the CFB riser. Fluidization velocity of 4 m/s at time (<b>a</b>) 0.05 s; (<b>b</b>) 0.1 s; and (<b>c</b>) 0.15 s.</p>
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<p>Probability distribution for the positions of a batch of particles after being put into the inlet of the CFB riser. Fluidization velocity of 6 m/s at time (<b>a</b>) 0.05 s; (<b>b</b>) 0.1 s; and (<b>c</b>) 0.15 s.</p>
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<p>Probability distribution for the positions of a batch of particles after being put into the inlet of the CFB riser. Fluidization velocity of 8 m/s at time (<b>a</b>) 0.05 s; (<b>b</b>) 0.1 s; and (<b>c</b>) 0.15 s.</p>
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<p>Particle RTD calculated by two models. Fluidization velocity: (<b>a</b>) 3 m/s; (<b>b</b>) 4 m/s; (<b>c</b>) 6 m/s; (<b>d</b>) 8 m/s.</p>
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<p>Particle distribution comparison of the two models. Fluidization velocity: 4 m/s. (<b>a</b>) CFD-DEM; (<b>b</b>) MCM.</p>
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<p>Axial particle mixing of two models. Fluidization velocity: (<b>a</b>) 3 m/s; (<b>b</b>) 4 m/s; (<b>c</b>) 6 m/s; (<b>d</b>) 8 m/s.</p>
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19 pages, 18180 KiB  
Article
Analysis of Crushing Performance of Toothed Double-Roll Crusher for Coal Particle Based on Discrete Element Method
by Zeren Chen, Guoqiang Wang, Zhengjie Lei, Duomei Xue and Zhengbin Liu
Processes 2025, 13(3), 613; https://doi.org/10.3390/pr13030613 - 21 Feb 2025
Viewed by 307
Abstract
The large toothed double-roll crusher, as key crushing equipment for open pit mines, is very necessary to analyse its crushing performance at different feed particle sizes, compressive strengths of coal, and rotation speeds of toothed rollers. Firstly, a toothed double-roll crusher is taken [...] Read more.
The large toothed double-roll crusher, as key crushing equipment for open pit mines, is very necessary to analyse its crushing performance at different feed particle sizes, compressive strengths of coal, and rotation speeds of toothed rollers. Firstly, a toothed double-roll crusher is taken as the research object in this paper; the coupling simulation model of the toothed double-roll crusher based on the DEM-MBD and Ab-T10 breakage model is constructed. The validity of the coupling simulation model is verified through the actual measurement data. On this basis, the crushing performance under variable factors is analysed by integrating comprehensive tests. The results show that the rotation speed of the toothed roller is the main influence factor on the crushing performance of the toothed double-roll crusher when it works at 25~33.3 r/min. With the increase in compressive strength of coal, the productivity decreases, and this phenomenon disappears gradually at 33.3~42 r/min. Further, a 5–20% increase in the large size of the coal particles can improve 10% crushing quality with a discharge size lower than 300 mm and approximately 25% productivity of the toothed double-roll crusher. Finally, the power density is reduced as the mass percentage of large-sized coal particles increases, and this phenomenon is weakened with the increase in the compressive strength of coal. Full article
(This article belongs to the Section Manufacturing Processes and Systems)
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<p>Toothed double-roll crusher.</p>
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<p>Coal particles and DEM model.</p>
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<p>Schematic diagram of a typical loading and unloading cycle.</p>
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<p>Schematic diagram of uniaxial compression test process.</p>
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<p>Calibration process of the crushing parameters. (<b>a</b>) Calibration process of 5.5 MPa coal, (<b>b</b>) Calibration process of 16.8 MPa coal.</p>
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<p>DEM-MBD coupling simulation model of TDRC.</p>
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<p>Validation simulation analysis of TDRC. The light blue area in the figure shows the discharge particle size within 0.3 m.</p>
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<p>Normal stress distribution of crushing teeth.</p>
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<p>Discharge particle size at 33.3 r/min crushing condition. (<b>a</b>) Full size discharge particle size distribution, (<b>b</b>) Discharge particle size distribution within 0.35 m.</p>
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27 pages, 21307 KiB  
Article
A POD-Based Reduced-Dimension Method for Solution Coefficient Vectors in the Crank–Nicolson Mixed Finite Element Method for the Fourth-Order Parabolic Equation
by Xiaohui Chang and Hong Li
Fractal Fract. 2025, 9(3), 137; https://doi.org/10.3390/fractalfract9030137 - 21 Feb 2025
Viewed by 286
Abstract
This research proposes a method for reducing the dimension of the coefficient vector for Crank–Nicolson mixed finite element (CNMFE) solutions to solve the fourth-order variable coefficient parabolic equation. Initially, the CNMFE schemes and corresponding matrix schemes for the equation are established, followed by [...] Read more.
This research proposes a method for reducing the dimension of the coefficient vector for Crank–Nicolson mixed finite element (CNMFE) solutions to solve the fourth-order variable coefficient parabolic equation. Initially, the CNMFE schemes and corresponding matrix schemes for the equation are established, followed by a thorough discussion of the uniqueness, stability, and error estimates for the CNMFE solutions. Next, a matrix-form reduced-dimension CNMFE (RDCNMFE) method is developed utilizing proper orthogonal decomposition (POD) technology, with an in-depth discussion of the uniqueness, stability, and error estimates of the RDCNMFE solutions. The reduced-dimension method employs identical basis functions, unlike standard CNMFE methods. It significantly reduces the number of unknowns in the computations, thereby effectively decreasing computational time, while there is no loss of accuracy. Finally, numerical experiments are performed for both fourth-order and time-fractional fourth-order parabolic equations. The proposed method demonstrates its effectiveness not only for the fourth-order parabolic equations but also for time-fractional fourth-order parabolic equations, which further validate the universal applicability of the POD-based RDCNMFE method. Under a spatial discretization grid 40×40, the traditional CNMFE method requires 2×412 degrees of freedom at each time step, while the RDCNMFE method reduces the degrees of freedom to 2×6 through POD technology. The numerical results show that the RDCNMFE method is nearly 10 times faster than the traditional method. This clearly demonstrates the significant advantage of the RDCNMFE method in saving computational resources. Full article
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<p>(<b>a</b>) The exact solution <math display="inline"><semantics> <msup> <mi>u</mi> <mi>n</mi> </msup> </semantics></math>. (<b>b</b>) The CNMFE solution <math display="inline"><semantics> <msubsup> <mi>u</mi> <mi>h</mi> <mi>n</mi> </msubsup> </semantics></math>. (<b>c</b>) The RDCNMFE solution <math display="inline"><semantics> <msubsup> <mi>u</mi> <mi>d</mi> <mi>n</mi> </msubsup> </semantics></math>.</p>
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<p>(<b>a</b>) The exact solution <math display="inline"><semantics> <msup> <mi>q</mi> <mi>n</mi> </msup> </semantics></math>. (<b>b</b>) The CNMFE solution <math display="inline"><semantics> <msubsup> <mi>q</mi> <mi>h</mi> <mi>n</mi> </msubsup> </semantics></math>. (<b>c</b>) The RDCNMFE solution <math display="inline"><semantics> <msubsup> <mi>q</mi> <mi>d</mi> <mi>n</mi> </msubsup> </semantics></math>.</p>
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<p>(<b>a</b>) The exact solution <math display="inline"><semantics> <msup> <mi>u</mi> <mi>n</mi> </msup> </semantics></math>. (<b>b</b>) The CNMFE solution <math display="inline"><semantics> <msubsup> <mi>u</mi> <mi>h</mi> <mi>n</mi> </msubsup> </semantics></math>. (<b>c</b>) The RDCNMFE solution <math display="inline"><semantics> <msubsup> <mi>u</mi> <mi>d</mi> <mi>n</mi> </msubsup> </semantics></math>.</p>
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<p>(<b>a</b>) The exact solution <math display="inline"><semantics> <msup> <mi>q</mi> <mi>n</mi> </msup> </semantics></math>. (<b>b</b>) The CNMFE solution <math display="inline"><semantics> <msubsup> <mi>q</mi> <mi>h</mi> <mi>n</mi> </msubsup> </semantics></math>. (<b>c</b>) The RDCNMFE solution <math display="inline"><semantics> <msubsup> <mi>q</mi> <mi>d</mi> <mi>n</mi> </msubsup> </semantics></math>.</p>
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<p>Comparison of error results of u when <math display="inline"><semantics> <mrow> <mi>ε</mi> <mo>=</mo> <mn>1</mn> <mo>,</mo> <mn>0.01</mn> <mo>,</mo> <mn>0.0001</mn> </mrow> </semantics></math>.</p>
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<p>Comparison of error results of q when <math display="inline"><semantics> <mrow> <mi>ε</mi> <mo>=</mo> <mn>1</mn> <mo>,</mo> <mn>0.01</mn> <mo>,</mo> <mn>0.0001</mn> </mrow> </semantics></math>.</p>
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<p>(<b>a</b>) The exact solution <math display="inline"><semantics> <msup> <mi>u</mi> <mi>n</mi> </msup> </semantics></math>. (<b>b</b>) The CNMFE solution <math display="inline"><semantics> <msubsup> <mi>u</mi> <mi>h</mi> <mi>n</mi> </msubsup> </semantics></math>. (<b>c</b>) The RDCNMFE solution <math display="inline"><semantics> <msubsup> <mi>u</mi> <mi>d</mi> <mi>n</mi> </msubsup> </semantics></math>.</p>
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<p>(<b>a</b>) The exact solution <math display="inline"><semantics> <msup> <mi>q</mi> <mi>n</mi> </msup> </semantics></math>. (<b>b</b>) The CNMFE solution <math display="inline"><semantics> <msubsup> <mi>q</mi> <mi>h</mi> <mi>n</mi> </msubsup> </semantics></math>. (<b>c</b>) The RDCNMFE solution <math display="inline"><semantics> <msubsup> <mi>q</mi> <mi>d</mi> <mi>n</mi> </msubsup> </semantics></math>.</p>
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18 pages, 2790 KiB  
Article
Particle Size-and Structure-Dependent Breakage Behaviors of EnAM-Containing Slags
by Simon Bahnmüller, Paul Hirschberger, Thu Trang Võ, Cindytami Rachmawati, Arno Kwade, Urs Peuker, Harald Kruggel-Emden and Carsten Schilde
Minerals 2025, 15(2), 195; https://doi.org/10.3390/min15020195 - 19 Feb 2025
Viewed by 223
Abstract
Slags containing critical minerals concentrated in artificial phases, so-called engineered artificial minerals (EnAMs), are a novel source of critical raw materials. To liberate the EnAMs, the slags need to be comminuted, reducing the size of the particles. This work investigated the dependence of [...] Read more.
Slags containing critical minerals concentrated in artificial phases, so-called engineered artificial minerals (EnAMs), are a novel source of critical raw materials. To liberate the EnAMs, the slags need to be comminuted, reducing the size of the particles. This work investigated the dependence of the breakage behavior on particle size and mineral structure during the comminution of an EnAM-containing slag. Piston-die experiments were performed for particles in the 3 mm to 5 mm size range. Nanoindentation and two-roller breakage tester experiments were performed for those in the 50 µm to 200 µm size range. The investigations were accompanied by X-ray computed tomography (XCT) and scanning electron microscope/energy dispersive X-ray spectroscopy (SEM/EDX) measurements as well as a micro X-ray fluorescence analysis to examine the mineral microstructure. It was found that the commonly assumed exponential connection between particle size and strength differed in the two size ranges. This behavior can be linked to different grain and cluster sizes, which were found in the investigation of the mineral microstructure. In addition to particle size, it was found that mineral structure plays an important role when characterizing the breakage behavior. Full article
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<p>(<b>a</b>) A part of the homogeneous structure shown in element map overlay from micro X-ray fluorescence analysis and (<b>b</b>) the binary image of the observed clusters.</p>
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<p>Representative schematic representation of the structures: mineral grains and clusters in the slag sample are shown for three exemplary clusters, where a single hexagonal shape represents a single mineral grain and the accumulation of the connected single grains forms a cluster.</p>
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<p>Three-dimensional visualization of XCT data of three clusters consisting of smaller grains. Every grain is labelled with a unique color for easier differentiation.</p>
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<p>Size distributions of the determined grain size determined based on XCT image data (<b>a</b>) and the cluster size based on µXRF imaging (<b>b</b>) of the analyzed slag.</p>
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<p>SEM (<b>a</b>,<b>b</b>) and SEM/EDX (<b>c</b>,<b>d</b>) images of two samples (<b>a</b>–<b>d</b>) in the micro size range to investigate the structural composition on the surface.</p>
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<p>Side view of the load cell (Shimadzu AG-50kNX) during the piston tests (<b>a</b>) and a schematic representation of the load cell’s setup (<b>b</b>).</p>
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<p>Schematic representation of the two-roller breakage tester.</p>
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<p>Distribution of the breakage strength of particles in different size classes in the micro and macro size ranges.</p>
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<p>The specific breakage strength as a function of particle size, fitted based on the model by Tavares and King [<a href="#B33-minerals-15-00195" class="html-bibr">33</a>]. Grain and cluster sizes are highlighted as vertical lines.</p>
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<p>Normalized fragment size distribution in the micro and macro size ranges and normalized cluster and grain size distributions.</p>
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<p>Normalized fragment size distributions in the micro and macro ranges and fits for the breakage function by Zhu et al. [<a href="#B42-minerals-15-00195" class="html-bibr">42</a>].</p>
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18 pages, 3063 KiB  
Article
Numerical Investigation of the Wave Equation for the Convergence and Stability Analysis of Vibrating Strings
by Md Joni Alam, Ahmed Ramady, M. S. Abbas, K. El-Rashidy, Md Tauhedul Azam and M. Mamun Miah
AppliedMath 2025, 5(1), 18; https://doi.org/10.3390/appliedmath5010018 - 19 Feb 2025
Viewed by 191
Abstract
The modeling of the one-dimensional wave equation is the fundamental model for characterizing the behavior of vibrating strings in different physical systems. In this work, we investigate numerical solutions for the one-dimensional wave equation employing both explicit and implicit finite difference schemes. To [...] Read more.
The modeling of the one-dimensional wave equation is the fundamental model for characterizing the behavior of vibrating strings in different physical systems. In this work, we investigate numerical solutions for the one-dimensional wave equation employing both explicit and implicit finite difference schemes. To evaluate the correctness of our numerical schemes, we perform extensive error analysis looking at the L1 norm of error and relative error. We conduct thorough convergence tests as we refine the discretization resolutions to ensure that the solutions converge in the correct order of accuracy to the exact analytical solution. Using the von Neumann approach, the stability of the numerical schemes are carefully investigated so that both explicit and implicit schemes maintain the stability criteria over simulations. We test the accuracy of our numerical schemes and present a few examples. We compare the solution with the well-known spectral and finite element method. We also show theoretical proof of the stability and convergence of our numerical scheme. Full article
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<p>Comparison of numerical and exact solutions.</p>
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<p>Evolution of maximum <math display="inline"><semantics> <mrow> <msup> <mrow> <mi>L</mi> </mrow> <mrow> <mn>1</mn> </mrow> </msup> </mrow> </semantics></math> norm of error over timesteps.</p>
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<p>Log-log plot of the maximum error versus the number of intervals <math display="inline"><semantics> <mrow> <mi>n</mi> <mo>.</mo> </mrow> </semantics></math></p>
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<p>Maximum relative error at each timestep.</p>
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<p>Stable solution for explicit scheme satisfying CFL conditions <math display="inline"><semantics> <mrow> <mi>γ</mi> <mo>=</mo> <mn>0.9</mn> <mo>≤</mo> <mn>1</mn> </mrow> </semantics></math>.</p>
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<p>Unstable solution for explicit scheme violating CFL conditions <math display="inline"><semantics> <mrow> <mi>γ</mi> <mo>=</mo> <mn>1.1</mn> <mo>&gt;</mo> <mn>1</mn> </mrow> </semantics></math>.</p>
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<p>Stable solution for implicit scheme for CFL number <math display="inline"><semantics> <mrow> <mi>γ</mi> <mo>=</mo> <mn>2.5</mn> <mo>&gt;</mo> <mn>1</mn> </mrow> </semantics></math>.</p>
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