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Search Results (356)

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11 pages, 948 KiB  
Article
Contact Interaction of a Rigid Stamp and a Porous Elastic Cylinder of Finite Dimensions
by Mikhail I. Chebakov, Elena M. Kolosova and Maria D. Datcheva
Mathematics 2025, 13(1), 104; https://doi.org/10.3390/math13010104 - 30 Dec 2024
Viewed by 235
Abstract
This article investigates an axisymmetric contact problem involving the interaction between a rigid cylindrical stamp and a poroelastic cylinder of finite dimensions, based on the Cowin–Nunziato theory of media with voids. The stamp is assumed to have a flat base and to be [...] Read more.
This article investigates an axisymmetric contact problem involving the interaction between a rigid cylindrical stamp and a poroelastic cylinder of finite dimensions, based on the Cowin–Nunziato theory of media with voids. The stamp is assumed to have a flat base and to be in frictionless contact with the cylinder. The cylinder, in turn, rests on a rigid base without friction, with no normal displacements or tangential stresses on its lateral surface. Under an applied vertical force, the stamp undergoes displacement, compressing the poroelastic cylinder. The mathematical formulation of this problem involves expressing the unknown displacements within the cylinder and the variation in pore volume fraction as a series of Bessel functions. This representation reduces the problem to an integral equation of the first kind, describing the distribution of contact stresses beneath the stamp. The kernel of the integral equation is explicitly provided in its transformed form. The collocation method is employed to solve the integral equation, enabling the determination of contact stresses and the relationship between the indenter’s displacement and the applied force. A comparative model parameter analysis is performed to examine the effects of different material porosity parameters and model geometrical characteristics on the results. Full article
(This article belongs to the Section Engineering Mathematics)
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Figure 1
<p>Sketch of the problem geometry, including the notations for the dimensions of the rigid cylindrical stamp and the finite-sized deformable cylinder with voids, along with the applied load configuration (axis of symmetry—<span class="html-italic">Oz</span>).</p>
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<p>Contact stresses for different values of <span class="html-italic">N</span> and two different values of the substrate radius: (<b>a</b>) <span class="html-italic">R</span> = 1.1, (<b>b</b>) <span class="html-italic">R</span> = 1.5, with a cylindrical substrate height of <span class="html-italic">h</span> = 0.5.</p>
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17 pages, 11160 KiB  
Article
Influence of the Spray Swirl Flow on the Gas–Liquid Interfacial Area Morphology: Multiparametric Qualitative Analysis
by Grzegorz Ligus, Barbara Wasilewska, Marek Krok and Laura Pałys-Żyta
Energies 2025, 18(1), 91; https://doi.org/10.3390/en18010091 - 29 Dec 2024
Viewed by 352
Abstract
In this study, the authors carried out a multiparametric assessment of the influence of swirl patterns during aerosol flow on the shape of the interfacial area that forms the cone based on data obtained from experimental measurements using the PIV and LLS methods. [...] Read more.
In this study, the authors carried out a multiparametric assessment of the influence of swirl patterns during aerosol flow on the shape of the interfacial area that forms the cone based on data obtained from experimental measurements using the PIV and LLS methods. The results were correlated with the disinfection process occurring in the near and far fields of the aerosol (direct surface disinfection and volume fogging). In this study, parameters such as turbulent kinetic energy (TKE), swirl strength (SS), pressure fields, and Sauter mean diameter (d32) are used to investigate the relationship between aerosol spray morphology and flow dynamics under different operating conditions. Three different geometrical settings of the aerosol-generating system and two different pressures corresponding to the air supply to the spray nozzle have been adopted. By evaluating the results obtained, the influence of each parameter on the formation of the aerosol displacement trajectory, the stabilization of the spray cone, and its degradation was identified. The shape of the boundary between the dynamically moving aerosol and the surrounding air was also evaluated. The conditions for swirling and straight-line flows within the aerosol cone, and, thus, the conditions for the volumetric development of swirling phenomena, were further clarified. Full article
(This article belongs to the Special Issue Heat Transfer and Multiphase Flow)
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Figure 1
<p>Methodology of multiparametric qualitative analysis of the swirl spray pattern.</p>
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<p>Liquid output Q<sub>L</sub> of the adopted studies.</p>
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22 pages, 3247 KiB  
Article
Experimental Identification of the Void Fraction in a Large Hydrodynamic Offset Halves Bearing
by Alexander Engels, Sören Wettmarshausen, Michael Stottrop, Thomas Hagemann, Christoph Weißbacher, Hubert Schwarze and Beate Bender
Lubricants 2025, 13(1), 7; https://doi.org/10.3390/lubricants13010007 - 29 Dec 2024
Viewed by 411
Abstract
A common approach to optimising hydrodynamic journal bearings for power loss is to reduce the lubricant supply and direct the oil to specific bearing areas where it is needed to guarantee safe operation. This requires information on the processes in the gap and [...] Read more.
A common approach to optimising hydrodynamic journal bearings for power loss is to reduce the lubricant supply and direct the oil to specific bearing areas where it is needed to guarantee safe operation. This requires information on the processes in the gap and the surrounding pocket areas for both pre-design and simulation. In this paper, a system consisting of a total of eight cameras is used to determine the void fraction in deep grooves outside the lubricant film. The void fraction in the lubrication gap is determined using a novel method for the evaluation of two proximity measurements. While the variation of the deep groove void fraction is realised by a special oil supply and radially adjustable deep groove elements, the gap void fraction is adjusted by the oil supply in the lube oil pockets at the pad leading edges. On the one hand, the experimental investigations show that the void fraction of the deep groove areas has hardly any influence on the general operating behaviour. On the other hand, the void fraction in the lubrication gap can be measured quantitatively for the first time, and the operating point-dependent gas fractions can be visualised. It is also shown that gaseous cavitation is the main mechanism in partially filled regions of the lubrication gap. Full article
(This article belongs to the Special Issue Advances in Lubricated Bearings, 2nd Edition)
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Figure 1
<p>Test bearing design of OHB; (<b>a</b>): OHB in isometric view; (<b>b</b>): OHB in axial section from drive direction at <math display="inline"><semantics> <mrow> <msub> <mi>b</mi> <mi mathvariant="normal">b</mi> </msub> <mo>=</mo> <mn>97.5</mn> </mrow> </semantics></math> mm; (<b>c</b>): OHB in radial section.</p>
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<p>Test rig for large hydrodynamic journal bearings (side view); (<b>a</b>): test rig in radial half section; (<b>b</b>): enlarged view of the test bearing with measurement technology (radial section in the measuring plane of the distance sensors).</p>
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<p>Schematic view: radial section of the shaft in the sensor plane.</p>
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<p>Radially adjustable deep groove elements, which can be fed with lubricant; (<b>a</b>): initial design; (<b>b</b>): recessed deep groove; (<b>c</b>): recessed deep groove with deep groove element.</p>
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<p>Camera system for identifying the lower deep groove and ring groove void fraction; <span class="html-italic">1,4</span>: radial cameras in deep groove element; <span class="html-italic">2,3</span>: axial cameras for the ring groove void fraction.</p>
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<p>Comparison of the models for determining the void fraction.</p>
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<p>Comparison and dependencies of the distance sensors.</p>
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<p>Calibration functions of the capacitive distance sensor and the eddy current distance sensor.</p>
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<p>Relative permittivity as a function of temperature of different VG 32 lubricants.</p>
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<p>Deep groove void fraction @ <span class="html-italic">n</span> <math display="inline"><semantics> <mrow> <mo>=</mo> <mn>1500</mn> <mspace width="3.33333pt"/> <msup> <mi>min</mi> <mrow> <mo>−</mo> <mn>1</mn> </mrow> </msup> <mo>,</mo> <mo> </mo> <mover> <mi>p</mi> <mo>¯</mo> </mover> <mo>=</mo> <mn>1.00</mn> <mspace width="3.33333pt"/> <mi>MPa</mi> <mo>,</mo> <mo> </mo> <msub> <mi>h</mi> <mi mathvariant="normal">P</mi> </msub> <mo>=</mo> <mn>45</mn> <mspace width="3.33333pt"/> <mi>mm</mi> <mo>,</mo> <mo> </mo> <msub> <mover accent="true"> <mi>V</mi> <mo>˙</mo> </mover> <mi mathvariant="normal">L</mi> </msub> <mo>=</mo> <mn>7.5</mn> <mspace width="3.33333pt"/> <mi mathvariant="normal">L</mi> <mspace width="0.166667em"/> <mi>·</mi> <mspace width="0.166667em"/> <msup> <mi mathvariant="normal">s</mi> <mrow> <mo>−</mo> <mn>1</mn> </mrow> </msup> </mrow> </semantics></math>.</p>
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<p>Deep groove depth influence on <math display="inline"><semantics> <msub> <mi>h</mi> <mi>min</mi> </msub> </semantics></math> and <math display="inline"><semantics> <msub> <mi>P</mi> <mi>fric</mi> </msub> </semantics></math> at different operating points.</p>
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<p>Deep groove depth influence on operational safety and power loss at different operating points.</p>
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<p>Lubrication gap void fraction at low rotational speed and load (<b>a</b>) and high rotational speed and load (<b>b</b>) under standard oil supply and fully opened deep grooves.</p>
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<p>Lubrication gap pressure distribution (<b>a</b>) and void fraction (<b>b</b>) under standard oil supply and fully opened deep grooves.</p>
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<p>Lubrication gap void fraction @ <span class="html-italic">n</span> <math display="inline"><semantics> <mrow> <mo>=</mo> <mn>3000</mn> <mspace width="3.33333pt"/> <msup> <mi>min</mi> <mrow> <mo>−</mo> <mn>1</mn> </mrow> </msup> </mrow> </semantics></math>, <math display="inline"><semantics> <mrow> <mover> <mi>p</mi> <mo>¯</mo> </mover> <mo>=</mo> <mn>1.00</mn> <mo>.</mo> <mo>.</mo> <mo>.</mo> <mn>4.50</mn> <mspace width="3.33333pt"/> <mi>MPa</mi> </mrow> </semantics></math> with standard oil supply and fully opened deep grooves.</p>
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<p>Lubrication gap void fraction @ <span class="html-italic">n</span> <math display="inline"><semantics> <mrow> <mo>=</mo> <mn>1000</mn> <mo>.</mo> <mo>.</mo> <mn>3600</mn> <mspace width="3.33333pt"/> <msup> <mi>min</mi> <mrow> <mo>−</mo> <mn>1</mn> </mrow> </msup> </mrow> </semantics></math>, <math display="inline"><semantics> <mrow> <mover> <mi>p</mi> <mo>¯</mo> </mover> <mo>=</mo> <mn>1.00</mn> <mspace width="3.33333pt"/> <mi>MPa</mi> </mrow> </semantics></math> with standard oil supply at maximum deep groove depth.</p>
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<p>Lubrication gap void fraction @ <span class="html-italic">n</span> <math display="inline"><semantics> <mrow> <mo>=</mo> <mn>1500</mn> <mspace width="3.33333pt"/> <msup> <mi>min</mi> <mrow> <mo>−</mo> <mn>1</mn> </mrow> </msup> </mrow> </semantics></math>, <math display="inline"><semantics> <mrow> <mover> <mi>p</mi> <mo>¯</mo> </mover> <mo>=</mo> <mn>1.00</mn> <mspace width="3.33333pt"/> <mi>MPa</mi> </mrow> </semantics></math> with varying <math display="inline"><semantics> <msub> <mover accent="true"> <mi>V</mi> <mo>˙</mo> </mover> <mi mathvariant="normal">P</mi> </msub> </semantics></math>.</p>
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<p>Lubrication gap void fraction @ <span class="html-italic">n</span> <math display="inline"><semantics> <mrow> <mo>=</mo> <mn>3000</mn> <mspace width="3.33333pt"/> <msup> <mi>min</mi> <mrow> <mo>−</mo> <mn>1</mn> </mrow> </msup> </mrow> </semantics></math>, <math display="inline"><semantics> <mrow> <mover> <mi>p</mi> <mo>¯</mo> </mover> <mo>=</mo> <mn>2.0</mn> <mspace width="3.33333pt"/> <mi>MPa</mi> </mrow> </semantics></math> with varying <math display="inline"><semantics> <msub> <mover accent="true"> <mi>V</mi> <mo>˙</mo> </mover> <mi mathvariant="normal">P</mi> </msub> </semantics></math>.</p>
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<p>Deep groove volume flow influence on <math display="inline"><semantics> <msub> <mi>h</mi> <mi>min</mi> </msub> </semantics></math> and <math display="inline"><semantics> <msub> <mi>P</mi> <mi>fric</mi> </msub> </semantics></math> at different operating points.</p>
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<p>Deep groove flow rate influence on maximum film pressure and temperature at different operating points.</p>
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20 pages, 12886 KiB  
Article
Microstructural Analysis, Compressive Strength, and Wear Properties of Spark-Plasma-Sintered Al–Mg–PPA Composites
by Osarue Osaruene Edosa, Francis Kunzi Tekweme, Peter A. Olubambi and Kapil Gupta
Quantum Beam Sci. 2024, 8(4), 32; https://doi.org/10.3390/qubs8040032 - 17 Dec 2024
Viewed by 340
Abstract
One technique for sintering green compacts and imparting the required qualities to meet the specific application requirements is spark plasma sintering (SPS). This study examines the effects of SPS parameters (sintering temperature and pressure, holding time, and heating rate) and plantain peel ash [...] Read more.
One technique for sintering green compacts and imparting the required qualities to meet the specific application requirements is spark plasma sintering (SPS). This study examines the effects of SPS parameters (sintering temperature and pressure, holding time, and heating rate) and plantain peel ash (PPA) reinforcement concentrations (0, 5 wt%, 10 wt%, 15 wt%, and 20 wt%) on the microstructure, compressive strength, and wear characteristics of the fabricated Al–Mg–PPA composites. As a result of the ball milling machine’s high efficiency, the PPA reinforcement was evenly dispersed throughout the aluminum matrix after 90 min of milling. At lower sintering temperatures and pressures, microstructural flaws such as weak grain boundaries, micro-pores, and micro-cracks were more noticeable than at higher ones. The PPA reinforcement and magnesium powder (wetting agent) increased the composites’ compressive strength by improving the wettability between the PPA reinforcement and the Al matrix. At a weight fraction of 5 wt% PPA, the maximum compressive strength of 432 MPa was attained for the sintered composites, which is a 222% improvement over the sintered aluminum matrix. Additionally, the PPA reinforcement enhanced the wear properties of the sintered Al–Mg–PPA composites by reducing the wear loss. Increasing the wear load resulted in a higher wear rate. The COF for the sintered composites ranges from 0.049 to 0.727. The most consistent correlation between the wear rate and the COF is that as the wear rate decreases, the COF decreases, and vice versa. Abrasive wear was the dominant wear mechanism observed. Tear ridges, shear steps, micro-voids, and cleavages were seen on the composites’ fracture surfaces, an indication of a ductile-brittle fracture. Full article
(This article belongs to the Special Issue Quantum Beam Science: Feature Papers 2024)
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Figure 1
<p>Processing sequence and equipment for fabricating, characterizing, and testing of (L<sub>25</sub>5<sup>5</sup>) Al/Mg/PPA composites.</p>
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<p>Processing sequence and equipment for fabricating, characterizing, and testing of (L<sub>25</sub>5<sup>5</sup>) Al/Mg/PPA composites.</p>
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<p>Optical micrographs of AMCs (<b>a</b>) Al–2Mg, (<b>b</b>) Al–2Mg–5PPA, (<b>c</b>) Al–2Mg–15PPA, and (<b>d</b>) Al–2Mg–20PPA.</p>
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<p>Optical micrographs of AMCs (<b>a</b>) Al–2Mg, (<b>b</b>) Al–2Mg–5PPA, (<b>c</b>) Al–2Mg–15PPA, and (<b>d</b>) Al–2Mg–20PPA.</p>
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<p>SEM micrographs of AMCs (<b>a</b>) Al–2Mg, (<b>b</b>) Al–2Mg–5PPA, (<b>c</b>) Al–2Mg–20PPA@430 °C, (<b>d</b>) Al–2Mg–20PPA@550 °C; (<b>e</b>) Al–2Mg–5PPA—EDS and (<b>f</b>) Al–2Mg–20PPA—EDS.</p>
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<p>SEM micrographs of AMCs (<b>a</b>) Al–2Mg, (<b>b</b>) Al–2Mg–5PPA, (<b>c</b>) Al–2Mg–20PPA@430 °C, (<b>d</b>) Al–2Mg–20PPA@550 °C; (<b>e</b>) Al–2Mg–5PPA—EDS and (<b>f</b>) Al–2Mg–20PPA—EDS.</p>
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<p>XRD patterns for Al–Mg–PPA samples (<b>a</b>) Al–2Mg–15PPA, (<b>b</b>) Al–2Mg–20PPA@520 °C, (<b>c</b>) Al–2Mg, and (<b>d</b>) Al–2Mg–20PPA@550 °C.</p>
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<p>Stress–strain curves of Al–Mg–PPA composites under compressive load.</p>
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<p>Main effect plots for (<b>a</b>) means and (<b>b</b>) SNRs of compressive strength.</p>
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<p>Pareto chart of standardized effects for wear rate.</p>
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<p>SEM fracture surfaces of Al–Mg–PPA composites (<b>a</b>) Al<b>–</b>2Mg–5PPA, (<b>b</b>) Al–2Mg–20PPA, (<b>c</b>) Al–2Mg, (<b>d</b>) Al–2Mg–15PPA, (<b>e</b>) Al–2Mg–10PPA, and (<b>f</b>) Al–2Mg.</p>
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<p>Effects of input parameters on the wear rates of Al–Mg–PPA composites (<b>a</b>) samples sintered at 430° (<b>b</b>) samples sintered at 550°.</p>
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<p>Main effects plots for SNRs of wear loss.</p>
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<p>Pareto chart of standardized effects for wear loss.</p>
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<p>Relationship between wear rate and COF of Al–Mg–PPA samples sintered at different temperatures (<b>a</b>) sintered at 430° (<b>b</b>) sintered at 550°.</p>
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<p>Effect of applied load on the wear rate of Al–Mg–PPA samples sintered at different temperatures (<b>a</b>) sintered at 430° (<b>b</b>) sintered at 550°.</p>
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<p>Effects of wear load on the COFs of Al–Mg–PPA composites.</p>
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<p>Worn out surfaces of Al–Mg–PPA composites (<b>a</b>) Al–2Mg–20PPA, (<b>b</b>) Al–2Mg–15PPA, (<b>c</b>) Al–2Mg–5PPA, and (<b>d</b>) Al–2Mg.</p>
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19 pages, 4071 KiB  
Article
Aged Lignocellulose Fibers of Cedar Wood (9th and 12th Century): Structural Investigation Using FTIR-Deconvolution Spectroscopy, X-Ray Diffraction (XRD), Crystallinity Indices, and Morphological SEM Analyses
by Yousra Bouramdane, Mustapha Haddad, Adil Mazar, Saadia Aît Lyazidi, Hicham Oudghiri Hassani and Abdellatif Boukir
Polymers 2024, 16(23), 3334; https://doi.org/10.3390/polym16233334 - 28 Nov 2024
Viewed by 739
Abstract
The characterization of lignocellulosic biomass present in archaeological wood is crucial for understanding the degradation processes affecting wooden artifacts. The lignocellulosic fractions in both the external and internal parts of Moroccan archaeological cedar wood (9th, 12th, and 21st centuries) were characterized using infrared [...] Read more.
The characterization of lignocellulosic biomass present in archaeological wood is crucial for understanding the degradation processes affecting wooden artifacts. The lignocellulosic fractions in both the external and internal parts of Moroccan archaeological cedar wood (9th, 12th, and 21st centuries) were characterized using infrared spectroscopy (FTIR-ATR deconvolution mode), X-ray diffraction (XRD), and SEM analysis. The XRD demonstrates a significant reduction in the crystallinity index of cellulose from recent to aging samples. This finding is corroborated by the FTIR analysis, which shows a significant reduction in the area profiles of the C-H crystalline cellulosic bands (1374, 1315, and 1265 cm−1) and C-O-C (1150–1000 cm−1). The alterations in the lignin fraction of aging samples (from the 9th and 12th centuries) were demonstrated by a reduction in the intensity of the bands at 1271 and 1232 cm−1 (Car-O) and the formation of new compounds, such as quinones and/or diaryl carbonyl structures, within the 1700–1550 cm−1 range. The SEM images of cedar wood samples from the 9th and 12th centuries reveal voids, indicating that the entire cell wall component has been removed, a characteristic feature of simultaneous white rot fungi. In addition, horizontal “scratches” were noted, indicating possible bacterial activity. Full article
(This article belongs to the Special Issue Aging Behavior and Durability of Polymer Materials, 2nd Edition)
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<p>Hydrolysis reaction of acetyl groups in hemicelluloses.</p>
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<p>(<b>a</b>) Oxidation reaction of primary alcohol in cellulose. (<b>b</b>) Degradation reaction of vicinal diol and its oxidation to diacedic fraction in cellulose.</p>
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<p>(<b>a</b>) Chemical structure of lignin. (<b>b</b>) Oxidation and degradation of lateral chain in lignin. (<b>c</b>) Oxidation of aromatic ring in lignin.</p>
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<p>The IR spectra from bottom to top in black color represent the normal spectra with overlapping bands of three internal samples (C9: sample dating to 9th century, C12: sample dating to 12th century, and C21: sample dating to 21st century), while the other remaining colored spectra are recorded in deconvolution mode representing nonoverlapped and well-resolved bands.</p>
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<p>The IR spectra from bottom to top in black color represent the normal spectra with overlapping bands of three external samples (C’9: sample dating to 9th century, C’12: sample dating to 12th century, and C’21: sample dating to 21st century), while the other remaining colored spectra are recorded in deconvolution mode representing nonoverlapped and well-resolved bands.</p>
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<p>Superposition of FTIR spectra of both cedar wood samples: internal (C21: 21st century, C12: 12th century, C9: 9th century) and external (C’21: 21st century, C’12: 12th century, C’9: 9th century).</p>
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<p>XRD diffractogram of three superposed internal cedar wood samples (<b>C9</b>: sample dating to 9th century, <b>C12</b>: sample dating to 12th century, <b>C21</b>: sample dating to 21st century).</p>
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<p>XRD diffractogram of three superposed external cedar wood samples (<b>C’9</b>: sample dating to 9th century, <b>C’12</b>: sample dating to 12th century, <b>C’21</b>: sample dating to 21st century).</p>
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<p>SEM micrographs of three internal cedar wood samples (<b>C9</b>: sample dating to 9th century, <b>C12</b>: sample dating to 12th century, and <b>C21</b>: sample dating to 21st century).</p>
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<p>SEM micrographs of three external cedar wood samples (<b>C’9</b>: sample dating to 9th century, <b>C’12</b>: sample dating to 12th century, and <b>C’21</b>: sample dating to 21st century).</p>
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21 pages, 2929 KiB  
Article
Porosity Effects on the Composite Girder by Rheological Dynamics and FEM
by Nataša Mrđa Bošnjak, Dragan D. Milašinović, Danica Goleš, Jelena Gučević and Arpad Čeh
Materials 2024, 17(23), 5779; https://doi.org/10.3390/ma17235779 - 25 Nov 2024
Viewed by 504
Abstract
A theoretical model for porous viscoelastoplastic (VEP) materials in the dry state is investigated in this research study. The model is based on the principles of conservation of mass and energy using the rheological dynamic theory (RDT). The model provides expressions for the [...] Read more.
A theoretical model for porous viscoelastoplastic (VEP) materials in the dry state is investigated in this research study. The model is based on the principles of conservation of mass and energy using the rheological dynamic theory (RDT). The model provides expressions for the creep coefficient, Poisson’s ratio, modulus of elasticity, damage variable, and strength as a function of porosity and/or void volume fraction (VVF). The reliability of the proposed model was analyzed by comparing numerical results with experimental ones on hardened concrete. A numerical model was created and analyzed in the commercial software Abaqus and validated by comparison with experimental data obtained by geodetic measurements on a composite wood–lightweight concrete girder. The deflections and stresses of the beam resulting from the influence of concrete creep and porosity were analyzed at the initial moment of time and after 6 years. The results showed that the RDT provided a reliable model for estimating parameters after exposure to long-term loads. Full article
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<p>RDT method implementation in FEM analysis.</p>
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<p>Tested porosities versus concrete densities.</p>
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<p>Quasi-static stress–strain curves, determined by rheological dynamical analysis (RDA) modeling, for the SG concretes.</p>
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<p>Dynamic stress–strain curves for the SG concretes determined by RDT modeling.</p>
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<p>Porosity as a function of compressive strength for the SG concrete.</p>
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<p>Tested dimensionless Young’s modulus as a function of porosity in comparison with the straight line for the SG2 concrete.</p>
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<p>Tested compressive strengths compared with the curves obtained by RDT modeling for the SG2 and SG3 concretes.</p>
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<p>Cross-section of a concrete–wood beam (cm).</p>
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<p>Longitudinal section of a concrete–wood beam (cm).</p>
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<p>Concrete–wood beam after 6 years of exposure to atmospheric influences.</p>
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<p>Beam deflections at the initial moment of time <math display="inline"><semantics> <mrow> <msub> <mrow> <mi>t</mi> </mrow> <mrow> <mn>0</mn> </mrow> </msub> </mrow> </semantics></math> for the model of girder with screws.</p>
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<p>Beam deflections at t = 6 years for the model of a girder with screws.</p>
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<p>Longitudinal and transverse cross-section through the middle of the girder with the distribution of maximum stress at the moment t = 6 years.</p>
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<p>Beam deflections at the initial moment of time t<sub>0</sub> for the model of the girder with only frictional connection.</p>
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<p>Beam deflections at t = 6 years for the model of girder with only frictional connection.</p>
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<p>Beam deflections at t = 6 years for the model of the girder with only frictional connection and with porosity influence (half of the girder).</p>
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<p>Measured geodetic profiles.</p>
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15 pages, 10127 KiB  
Article
Electrical Tortuosities of Porous Structures Based on Triply Periodic Minimal Surfaces and Honeycombs for Power-to-Heat Systems
by Thorsten Ott and Volker Dreißigacker
Energies 2024, 17(22), 5781; https://doi.org/10.3390/en17225781 - 19 Nov 2024
Viewed by 407
Abstract
Power-to-heat (P2H) systems offer an efficient solution for decarbonization by facilitating the integration of renewable energy into the industrial, heating, and transport sectors. Its key requirements include high thermal efficiency and an appropriate electrical resistivity to meet application-specific electrical needs. When designing P2H [...] Read more.
Power-to-heat (P2H) systems offer an efficient solution for decarbonization by facilitating the integration of renewable energy into the industrial, heating, and transport sectors. Its key requirements include high thermal efficiency and an appropriate electrical resistivity to meet application-specific electrical needs. When designing P2H systems, materials and electrical boundary conditions are often limited by application-specific requirements, whereas geometric structures offer high degrees of freedom. While thermal design calculations are often straightforward due to a variety of available Nusselt and pressure loss correlations, simplified design pathways, particularly for porous structures, are lacking in electrical design. Given the wide range of geometric degrees of freedom for porous structures and the fact that detailed modeling involves substantial computational effort, this work employed electrical tortuosity to capture and correlate the geometry-dependent impacts on the effective electrical resistance in a compact way. Honeycomb and triply periodic minimal surface (TPMS)-based structures were selected for this purpose, as they are characterized by high specific surfaces, allowing for high total heat transfer coefficients. The results show that the effective electrical resistance of both TPMS and honeycomb structures can be adjusted by the geometric structure. It was found that the electrical tortuosities of the investigated TPMS structures are nearly identical, while honeycomb structures show slightly higher values. Furthermore, the electrical tortuosity is mainly a function of the void fraction and does not change with the specific surface when the void fraction is kept constant. Finally, correlations for electrical tortuosity depending on geometric parameters with a mean error below 5% are derived for the first time, thereby providing a basis for simplified and computationally efficient electrical design calculations for P2H systems. Full article
(This article belongs to the Section J: Thermal Management)
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<p>Different TPMS structures in a unit cell, from left to right: Fischer-Koch-S, Gyroid, Schoen IWP, Schwarz-D, and Schwarz-P.</p>
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<p>(<b>a</b>) Different cubic Gyroid structures in a unit cell for different level-set parameters (from left to right: <math display="inline"><semantics> <mrow> <mi>t</mi> </mrow> </semantics></math> = 0, <math display="inline"><semantics> <mrow> <mi>t</mi> </mrow> </semantics></math> = 0.5 and <math display="inline"><semantics> <mrow> <mi>t</mi> </mrow> </semantics></math> = −0.5); (<b>b</b>) noncubic unit cell of a Gyroid structure for a level-set parameter of <math display="inline"><semantics> <mrow> <mi>t</mi> </mrow> </semantics></math> = 0 with an exemplary aspect ratio <math display="inline"><semantics> <mrow> <mi>B</mi> </mrow> </semantics></math> = 1.5.</p>
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<p>Unit cells of thickened cubic TPMS structures (<math display="inline"><semantics> <mrow> <msub> <mrow> <mi>t</mi> </mrow> <mrow> <mi mathvariant="normal">r</mi> <mi mathvariant="normal">e</mi> <mi mathvariant="normal">f</mi> </mrow> </msub> <mo>=</mo> <mn>0</mn> <mo>)</mo> </mrow> </semantics></math> with a void fraction of 80% and a specific surface of 100 m<sup>2</sup>/m<sup>3</sup>.</p>
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<p>(<b>a</b>) Unit cells with homogeneous wall thickness and quadratic and rectangular channel; (<b>b</b>) honeycomb structures with an outer rectangular shape, consisting of 10 × 10 unit cells having a quadratic and rectangular channel geometry.</p>
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<p>Methodology for deriving parameterized electrical tortuosity.</p>
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<p>Example of a meshed gyroid structure (<b>left</b>) and inlet and outlet surfaces for setting the boundary conditions (<b>right</b>).</p>
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<p>Electrical tortuosity as a function of void fraction and specific surface for honeycomb structures.</p>
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<p>Electrical tortuosity as a function of void fraction for different TPMS structures with reference level-set parameter <math display="inline"><semantics> <mrow> <msub> <mrow> <mi>t</mi> </mrow> <mrow> <mi mathvariant="normal">r</mi> <mi mathvariant="normal">e</mi> <mi mathvariant="normal">f</mi> </mrow> </msub> <mo>=</mo> <mn>0</mn> </mrow> </semantics></math> in comparison to a honeycomb with quadratic channels.</p>
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<p>Electric current density distribution for a honeycomb with quadratic channels and a Schwarz-P structure for a void fraction <math display="inline"><semantics> <mrow> <mi>ε</mi> </mrow> </semantics></math> = 0.8; red is the highest and blue is the lowest current density. The arrow indicates the direction of the current flow.</p>
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<p>(<b>a</b>) Electrical tortuosity as a function of the level-set parameter <math display="inline"><semantics> <mrow> <msub> <mrow> <mi>t</mi> </mrow> <mrow> <mi mathvariant="normal">r</mi> <mi mathvariant="normal">e</mi> <mi mathvariant="normal">f</mi> </mrow> </msub> </mrow> </semantics></math> and void fraction <math display="inline"><semantics> <mrow> <mi>ε</mi> </mrow> </semantics></math> illustrated for a Gyroid structure; (<b>b</b>) increase in electrical tortuosity for a level-set parameter of <math display="inline"><semantics> <mrow> <msub> <mrow> <mi>t</mi> </mrow> <mrow> <mi mathvariant="normal">r</mi> <mi mathvariant="normal">e</mi> <mi mathvariant="normal">f</mi> </mrow> </msub> <mo>=</mo> <mo>±</mo> <mtext> </mtext> <mn>0.7</mn> <msub> <mrow> <mi>t</mi> </mrow> <mrow> <mi mathvariant="normal">m</mi> <mi mathvariant="normal">a</mi> <mi mathvariant="normal">x</mi> </mrow> </msub> </mrow> </semantics></math>, related to the base surface (<math display="inline"><semantics> <mrow> <msub> <mrow> <mi>t</mi> </mrow> <mrow> <mi mathvariant="normal">r</mi> <mi mathvariant="normal">e</mi> <mi mathvariant="normal">f</mi> </mrow> </msub> </mrow> </semantics></math> = 0) for all investigated TPMS structures.</p>
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<p>Cuboidal unit cells with a quadratic base for a Schwarz-P structure (<math display="inline"><semantics> <mrow> <msub> <mrow> <mi>t</mi> </mrow> <mrow> <mi mathvariant="normal">r</mi> <mi mathvariant="normal">e</mi> <mi mathvariant="normal">f</mi> </mrow> </msub> <mo>=</mo> <mn>0</mn> <mo>)</mo> </mrow> </semantics></math> with a void fraction of 80% and different aspect ratios <math display="inline"><semantics> <mrow> <mi>B</mi> </mrow> </semantics></math>; the primarily cubic unit cell (<math display="inline"><semantics> <mrow> <mi>B</mi> </mrow> </semantics></math> = 1) is elongated in the ‘∥-direction’ and compressed in the two spatial directions perpendicular to it (the ‘⊥-direction’).</p>
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<p>Electrical tortuosity as a function of void fraction <math display="inline"><semantics> <mrow> <mi>ε</mi> </mrow> </semantics></math> and the aspect ratio <math display="inline"><semantics> <mrow> <mi>B</mi> </mrow> </semantics></math> for a Schwarz-P structure: (<b>a)</b> orthogonal to the direction of stretching; (<b>b</b>) parallel to the direction of stretching.</p>
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<p>Electrical tortuosity orthogonal to the direction of stretching as a function of void fraction <math display="inline"><semantics> <mrow> <mi>ε</mi> </mrow> </semantics></math> for an aspect ratio of <math display="inline"><semantics> <mrow> <mi>B</mi> <mo>=</mo> <mn>2</mn> </mrow> </semantics></math> for TPMS and honeycomb structures.</p>
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<p>Comparison of the electrical tortuosity as a function of the void fraction with the results of Abueidda et al. [<a href="#B20-energies-17-05781" class="html-bibr">20</a>] and Sauermoser-Yri et al. [<a href="#B21-energies-17-05781" class="html-bibr">21</a>].</p>
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17 pages, 23471 KiB  
Article
An Analysis of Dynamic Recrystallization During the Reduction Pretreatment Process Using a Multiscale Model
by Die Wu, Zhen Ning, Yanlin Zhu and Wei Yu
Metals 2024, 14(11), 1290; https://doi.org/10.3390/met14111290 - 14 Nov 2024
Viewed by 574
Abstract
In this study, a multiscale model is developed through secondary development (UMAT and UEXTERNALDB) in Abaqus with the objective of simulating the thermal deformation process with dynamic recrystallization behavior. The model couples the finite element method (FEM) with the multiphase field model (MPFM), [...] Read more.
In this study, a multiscale model is developed through secondary development (UMAT and UEXTERNALDB) in Abaqus with the objective of simulating the thermal deformation process with dynamic recrystallization behavior. The model couples the finite element method (FEM) with the multiphase field model (MPFM), thereby establishing bidirectional coupling between macroscopic mechanical behavior and microstructural evolution. A comparison between the single-element hot compression simulation and experimental results demonstrates that the model accurately simulates both the macroscopic mechanical behavior and microstructural evolution during the thermal deformation process, thereby exhibiting high precision. Simulations of the reduction pretreatment (RP) process under different reduction amounts and billet surface temperatures demonstrate that increasing the reduction amount and billet surface temperature significantly enhances both plastic deformation and the volume fraction of dynamic recrystallization in the billet core. This results in the closure of core voids and the refinement of the core microstructure, thereby providing valuable guidance for the development of optimal reduction pretreatment (RP) processes. Full article
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<p>Flowchart of UMAT subroutine.</p>
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<p>A schematic diagram of the coupling between the FEM and the MPFM: (<b>a</b>) the macroscopic bulk body; (<b>b</b>) discretization with finite elements; (<b>c</b>) the MPFM.</p>
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<p>Flowchart of multiscale model.</p>
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<p>Multiscale model of single-element hot compression simulation.</p>
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<p>Multiscale model of RP process.</p>
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<p>Temperature distribution in the billet at different billet surface temperatures: (<b>a</b>) 900 °C; (<b>b</b>) 950 °C; (<b>c</b>) 1000 °C.</p>
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<p>A schematic diagram of the conditions for stopping the MPFM calculation.</p>
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<p>Comparison of experimental and simulated results of stress–strain curves (—: experimental results; ○: simulated results): (<b>a</b>) 1200 °C; (<b>b</b>) 1100 °C; (<b>c</b>) 1050 °C; (<b>d</b>) 1000 °C; (<b>e</b>) 950 °C.</p>
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<p>Variation in dynamic recrystallization volume fraction with strain: (<b>a</b>) 0.001 s<sup>−1</sup>; (<b>b</b>) 0.01 s<sup>−1</sup>; (<b>c</b>) 0.1 s<sup>−1</sup>.</p>
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<p>Distribution of plastic strain in billet under different reduction ratios: (<b>a</b>) 6.67%; (<b>b</b>) 13.3%; (<b>c</b>) 20.0%.</p>
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<p>Variation in plastic strain with position under different reduction ratios.</p>
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<p>Distribution of dynamic recrystallization volume fraction in billet under different reduction ratios: (<b>a</b>) 6.67%; (<b>b</b>) 13.3%; (<b>c</b>) 20.0%.</p>
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<p>Variation in dynamic recrystallization volume fraction with position under different reduction ratios.</p>
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<p>Distribution of grain size in billet under different reduction ratios: (<b>a</b>) 6.67%; (<b>b</b>) 13.3%; (<b>c</b>) 20.0%.</p>
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<p>Variation in grain size with position under different reduction ratios.</p>
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<p>Distribution of plastic strain in billet under different billet surface temperatures: (<b>a</b>) 900 °C; (<b>b</b>) 950 °C; (<b>c</b>) 1000 °C.</p>
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<p>Variation in plastic strain with position under different billet surface temperatures.</p>
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<p>Distribution of dynamic recrystallization volume fraction in billet under different billet surface temperatures: (<b>a</b>) 900 °C; (<b>b</b>) 950 °C; (<b>c</b>) 1000 °C.</p>
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<p>Variation in dynamic recrystallization volume fraction with position under different billet surface temperatures.</p>
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<p>Distribution of dynamic recrystallization grain size in billet under different billet surface temperatures: (<b>a</b>) 900 °C; (<b>b</b>) 950 °C; (<b>c</b>) 1000 °C.</p>
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<p>Variation in grain size with position under different billet surface temperatures.</p>
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19 pages, 8216 KiB  
Article
Damage Evolution Mechanism of Railway Wagon Bogie Adapter 1035 Steel and Damage Parameter Calibration Based on Gursone–Tvergaarde–Needleman Model
by Jiachuan Yin, Xiaomin Huang, Guangzhi Ma, Changzhe Song, Xuefeng Tang and Hongchao Ji
Materials 2024, 17(20), 5070; https://doi.org/10.3390/ma17205070 - 17 Oct 2024
Viewed by 753
Abstract
As a critical component of a train, the railway wagon bogie adapter has higher quality requirements. During the forging process, external loads can induce voids, cracks, and other defects in the forging, thereby reducing its service life. Hence, studying the damage behavior of [...] Read more.
As a critical component of a train, the railway wagon bogie adapter has higher quality requirements. During the forging process, external loads can induce voids, cracks, and other defects in the forging, thereby reducing its service life. Hence, studying the damage behavior of the forging material, specifically AISI 1035 steel, becomes crucial. This study involved obtaining stress–strain curves for AISI 1035 steel through uniaxial tensile tests at temperatures of 900 °C, 1000 °C, and 1100 °C, with strain rates of 0.1 s−1, 1 s−1, and 10 s−1. Subsequently, SEM was used to observe samples at various deformation stages. The damage parameters, q1,  q2 and q3 in the GTN model “a computational model used to analyze and simulate material damage which can effectively capture the damage behavior of materials under different loading conditions” were then calibrated using the Ramberg–Osgood model and stress–strain curve fitting. Image Pro Plus software v11.1 quantified the sample porosity as f0, fn, fc and fF. A finite element model was established to simulate the tensile behavior of the AISI 1035 steel samples. By comparing the damage parameters of f0, fn, fc and fF obtained by the finite element method and experimental method, the validity of the damage parameters obtained by the finite element inverse method could be verified. Full article
(This article belongs to the Special Issue Research on Metal Cutting, Casting, Forming, and Heat Treatment)
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<p>Geometry and size of tensile sample.</p>
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<p>Uniaxial thermal tensile test flow.</p>
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<p>Tensile sample geometric model and mesh.</p>
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<p>Stress–strain curve of AISI 1035 steel: (<b>a</b>) 0.1 s<sup>−1</sup>, (<b>b</b>) 1 s<sup>−1</sup>, and (<b>c</b>) 10 s<sup>−1</sup>.</p>
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<p>Fitting results of the Ramberg–Osgood hardening model.</p>
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<p>Calibration of void nucleation strain <math display="inline"><semantics> <mrow> <msub> <mi>ε</mi> <mi>N</mi> </msub> </mrow> </semantics></math> of AISI 1035 steel: (<b>a</b>) 0.1 s<sup>−1</sup>; (<b>b</b>) 1s<sup>−1</sup>; (<b>c</b>) 10 s<sup>−1</sup>.</p>
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<p>Initial voids in the volume fraction measurement area and the undeformed area of the neck of the stretched sample.</p>
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<p>Digital image processing of SEM micrographs was used to calculate the porosity. (<b>a</b>) <math display="inline"><semantics> <mrow> <msub> <mi>f</mi> <mn>0</mn> </msub> </mrow> </semantics></math>; (<b>b</b>) <math display="inline"><semantics> <mrow> <msub> <mi>f</mi> <mi>n</mi> </msub> </mrow> </semantics></math>; (<b>c</b>) <math display="inline"><semantics> <mrow> <msub> <mi>f</mi> <mi>c</mi> </msub> </mrow> </semantics></math>; and (<b>d</b>) <math display="inline"><semantics> <mrow> <msub> <mi>f</mi> <mi>F</mi> </msub> </mrow> </semantics></math>.</p>
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<p>Matrix with partially grown void and nucleation void.</p>
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<p>Fracture morphology of A1S1 1035 steel at 1100 °C, 10 s<sup>−1</sup>: (<b>a</b>) 1000×; and (<b>b</b>) 5000×.</p>
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<p>Fracture morphology of AISI 1035 steel tensile parts at different temperatures: (<b>a</b>) 900 °C; (<b>b</b>) 1000 °C; and (<b>c</b>) 1100 °C.</p>
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<p>Fracture morphology of AISI 1035 steel tensile parts at different strain rates: (<b>a</b>) 0.1 s<sup>−1</sup>; (<b>b</b>) 1 s<sup>−1</sup>; and (<b>c</b>) 10 s<sup>−1</sup>.</p>
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<p>Response value of stress–strain curve.</p>
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<p>Response surface model of the effect of different parameter combinations on material strain at breaking point (R3). (<b>a</b>) Combination of <math display="inline"><semantics> <mrow> <msub> <mi>f</mi> <mn>0</mn> </msub> </mrow> </semantics></math> and <math display="inline"><semantics> <mrow> <msub> <mi>f</mi> <mi>n</mi> </msub> </mrow> </semantics></math>, (<b>b</b>) Combination of <math display="inline"><semantics> <mrow> <msub> <mi>f</mi> <mi>n</mi> </msub> </mrow> </semantics></math> and <math display="inline"><semantics> <mrow> <msub> <mi>f</mi> <mi>c</mi> </msub> </mrow> </semantics></math>, (<b>c</b>) Combination of <math display="inline"><semantics> <mrow> <msub> <mi>f</mi> <mi>c</mi> </msub> </mrow> </semantics></math> and <math display="inline"><semantics> <mrow> <msub> <mi>f</mi> <mi>F</mi> </msub> </mrow> </semantics></math>, (<b>d</b>) Combination of <math display="inline"><semantics> <mrow> <msub> <mi>f</mi> <mi>F</mi> </msub> </mrow> </semantics></math> and <math display="inline"><semantics> <mrow> <msub> <mi>f</mi> <mn>0</mn> </msub> </mrow> </semantics></math>.</p>
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<p>The 35 steel’s final fracture zone was obtained by experiment and finite element analysis.</p>
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<p>The 1100 °C thermal tensile simulation void volume fraction (VVF) distribution cloud image.</p>
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<p>Impact of each damage parameter on the tensile curve. (<b>a</b>) <span class="html-italic">f</span><sub>0</sub>, (<b>b</b>) <math display="inline"><semantics> <mrow> <msub> <mi>f</mi> <mi>n</mi> </msub> </mrow> </semantics></math>, (<b>c</b>) <math display="inline"><semantics> <mrow> <msub> <mi>f</mi> <mi>c</mi> </msub> </mrow> </semantics></math>, (<b>d</b>) <math display="inline"><semantics> <mrow> <msub> <mi>f</mi> <mi>F</mi> </msub> </mrow> </semantics></math>.</p>
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25 pages, 7613 KiB  
Article
The Influence and Mechanism of Polyvinyl Alcohol Fiber on the Mechanical Properties and Durability of High-Performance Shotcrete
by Ge Zhang, Like Li, Huawei Shi, Chen Chen and Kunpeng Li
Buildings 2024, 14(10), 3200; https://doi.org/10.3390/buildings14103200 - 8 Oct 2024
Viewed by 882
Abstract
This study investigates the impact of polyvinyl alcohol (PVA) fibers on the mechanical properties and durability of high-performance shotcrete (HPS). Results demonstrate that PVA fibers have a dual impact on the performance of HPS. Positively, PVA fibers enhance the tensile strength and toughness [...] Read more.
This study investigates the impact of polyvinyl alcohol (PVA) fibers on the mechanical properties and durability of high-performance shotcrete (HPS). Results demonstrate that PVA fibers have a dual impact on the performance of HPS. Positively, PVA fibers enhance the tensile strength and toughness of shotcrete due to their intrinsic high tensile strength and fiber-bridging effect, which significantly improves the material’s splitting tensile strength, deformation resistance, and toughness, and the splitting tensile strength and peak strain have been found to be increased by up to 30.77% and 31.51%, respectively. On the other hand, the random distribution and potential agglomeration of PVA fibers within the HPS matrix can lead to increased air-void formations. This phenomenon raises the volume content of large bubbles and increases the average bubble area and diameter, thereby elevating the pore volume fraction within the 500–1200 μm and >1200 μm ranges. Therefore, these microstructural changes reduce the compactness of the HPS matrix, resulting in a decrease in compressive strength and elastic modulus. The compressive strength exhibited a reduction ranging from 10.44% to 15.11%, while the elastic modulus showed a decrease of between 8.09% and 12.67%. Overall, the PVA-HPS mixtures with different mix proportions demonstrated excellent frost resistance, chloride ion penetration resistance, and carbonation resistance. The electrical charge passed ranged from 133 to 370 C, and the carbonation depth varied between 2.04 and 6.12 mm. Although the incorporation of PVA fibers reduced the permeability and carbonation resistance of shotcrete, it significantly mitigated the loss of tensile strength during freeze–thaw cycles. The findings offer insights into optimizing the use of PVA fibers in HPS applications, balancing enhancements in tensile properties with potential impacts on compressive performance. Full article
(This article belongs to the Section Building Materials, and Repair & Renovation)
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<p>Complete research methodology and experimental procedure.</p>
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<p>Shotcrete constitutive relation experimental under uniaxial compression: (<b>a</b>) operation and data acquisition system; (<b>b</b>) experimental process; (<b>c</b>) load-destroying.</p>
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<p>Bubble characterization parameter testing for (<b>a</b>) hardened concrete pore structure testing; (<b>b</b>) image of the air bubbles in hardened concrete.</p>
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<p>Effect of PVA fiber content on shotcrete strength for (<b>a</b>) compressive strength, and (<b>b</b>) splitting tensile strength.</p>
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<p>Effect of PVA fiber content on HPS tensile-to-compression ratio.</p>
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<p>The effect of fiber types and dosage on the strength of shotcrete [<a href="#B14-buildings-14-03200" class="html-bibr">14</a>,<a href="#B26-buildings-14-03200" class="html-bibr">26</a>,<a href="#B29-buildings-14-03200" class="html-bibr">29</a>,<a href="#B30-buildings-14-03200" class="html-bibr">30</a>,<a href="#B33-buildings-14-03200" class="html-bibr">33</a>,<a href="#B34-buildings-14-03200" class="html-bibr">34</a>]. (<b>a</b>) compressive strength, (<b>b</b>) splitting tensile strength.</p>
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<p>The relationship curve between HPS strength development and age, with different PVA fiber contents for (<b>a</b>) f<sub>cc</sub>(<span class="html-italic">t</span>)-ln(<span class="html-italic">t</span>), and (<b>b</b>) f<sub>ts</sub>(<span class="html-italic">t</span>)-ln(<span class="html-italic">t</span>).</p>
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<p>Effect of PVA fiber content on k<sub>cc</sub> and k<sub>ts</sub> value for (<b>a</b>) PVA-K<sub>cc</sub>, and (<b>b</b>) PVA-K<sub>ts</sub>.</p>
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<p>Compressive stress–strain curve of shotcrete with different mix proportions.</p>
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<p>Compressive stress–strain and equation fitting curve of shotcrete (dimensionless).</p>
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<p>Comparative study of pore characteristics of PVA-HPS under different fiber contents. (<b>a</b>) The relationship between dV/dlogD and pore size; (<b>b</b>) Proportion of pore volume.</p>
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<p>Microstructure of shotcrete with different mix proportions: (<b>a</b>) HPS; (<b>b</b>) PVA-0.50%; (<b>c</b>) PVA-1.0%.</p>
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<p>The variation curve of shotcrete bubble characteristic parameters under different PVA fiber contents for (<b>a</b>) bubble number and bubble volume content and (<b>b</b>) the average bubble area and average bubble diameter.</p>
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<p>Bubble pore volume distribution of HPS under different PVA fiber contents.</p>
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<p>The impact of PVA fibers on the mass loss rate and relative dynamic elastic modulus of HPS. (<b>a</b>) Mass loss rate and (<b>b</b>) relative dynamic elastic modulus.</p>
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<p>The variation compressive strength of PVA-HPS under freeze–thaw cycles. (<b>a</b>) Compressive strength, (<b>b</b>) compressive strength loss rate, and (<b>c</b>) relative compressive strength.</p>
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<p>The variation splitting tensile strength of PVA-HPS under freeze–thaw cycles. (<b>a</b>) Splitting tensile strength; (<b>b</b>) splitting tensile strength loss rate; (<b>c</b>) relative splitting tensile strength.</p>
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<p>Five-dimensional evaluation diagram of PVA-HPS.</p>
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<p>Electric flux of concrete with different PVA fiber contents.</p>
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<p>The impact of PVA fibers on the carbonation performance of HPS. (<b>a</b>) Carbonation depth and (<b>b</b>) carbonization coefficient.</p>
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17 pages, 13381 KiB  
Article
Vacuum Chamber Infusion for Fiber-Reinforced Composites
by Benjamin Grisin, Stefan Carosella and Peter Middendorf
Polymers 2024, 16(19), 2763; https://doi.org/10.3390/polym16192763 - 30 Sep 2024
Viewed by 1041
Abstract
A new approach to an automatable fiber impregnation and consolidation process for the manufacturing of fiber-reinforced composite parts is presented in this article. Therefore, a vacuum chamber sealing machine classically used in food packaging is modified for this approach—Vacuum Chamber Infusion (VCI). Dry [...] Read more.
A new approach to an automatable fiber impregnation and consolidation process for the manufacturing of fiber-reinforced composite parts is presented in this article. Therefore, a vacuum chamber sealing machine classically used in food packaging is modified for this approach—Vacuum Chamber Infusion (VCI). Dry fiber placement (DFP) preforms, made from 30 k carbon fiber tape, with different layer amounts and fiber orientations, are infused with the VCI and with the state-of-the-art process—Vacuum Assisted Process (VAP)—as the reference. VCI uses a closed system that is evacuated once, while VAP uses a permanently evacuated open system. Since process management greatly influences material properties, the mechanical properties, void content, and fiber volume fraction (FVF) are analyzed. In addition, the study aims to identify how the complexity of a resin infusion process can be reduced, the automation potential can be increased, and the number of consumables can be reduced. Comparable material characteristics and a reduction in consumables, setup complexity, and manufacturing time by a factor of four could be approved for VCI. A void content of less than 2% is measured for both processes and an FVF of 39% for VCI and 45% for VAP is achieved. Full article
(This article belongs to the Special Issue Manufacturing of Polymer-Matrix Composites)
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<p>VAP infusion setup and principle.</p>
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<p>Modified Vacuum chamber machine for resin infusion, modifications are highlighted.</p>
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<p>Scheme of the vacuum chamber infusion process. 1–4: Infusion with the inner bag. 5 and 6: Consolidation of the impregnated preform with an outer bag and 2 consolidation plates.</p>
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<p>The geometry of sample plates. A, B, and C: areas for void content and fiber volume fraction. Grey area: area for tensile and bending specimens.</p>
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<p>Representative cross-sections of the unidirectional configurations. Magnification: 250×, (<b>a</b>) VCI_UD_1 sample area B, (<b>b</b>) VCI_UD_2 sample area A, (<b>c</b>) VCI_UD_3 sample area A, (<b>d</b>) VAP_UD_1 sample area C, (<b>e</b>) VAP_UD_2 sample area B, (<b>f</b>) VAP_UD_2 sample area B.</p>
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<p>Representative cross-sections of the unidirectional configurations. Magnification: 250×, (<b>a</b>) VCI_UD_1 sample area B, (<b>b</b>) VCI_UD_2 sample area A, (<b>c</b>) VCI_UD_3 sample area A, (<b>d</b>) VAP_UD_1 sample area C, (<b>e</b>) VAP_UD_2 sample area B, (<b>f</b>) VAP_UD_2 sample area B.</p>
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<p>Representative cross-section with voids of the unidirectional configuration VCI. Magnification: 500×.</p>
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<p>Representative cross-sections of the 0/90° configuration magnification 250×, (<b>a</b>) VCI_0_90_1 sample area A, (<b>b</b>) VCI_0_90_2 sample area B, (<b>c</b>) VCI_0_90_3 sample area A, (<b>d</b>) VAP_0_90_1 sample area A, (<b>e</b>) VAP_0_90_2 sample area C, (<b>f</b>) VAP_0_90_3 sample area B.</p>
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<p>Representative cross-sections of the 0/90° configuration magnification 250×, (<b>a</b>) VCI_0_90_1 sample area A, (<b>b</b>) VCI_0_90_2 sample area B, (<b>c</b>) VCI_0_90_3 sample area A, (<b>d</b>) VAP_0_90_1 sample area A, (<b>e</b>) VAP_0_90_2 sample area C, (<b>f</b>) VAP_0_90_3 sample area B.</p>
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15 pages, 9688 KiB  
Article
Effect of Vibration Pretreatment–Microwave Curing Process Parameters on the Mechanical Performance of Resin-Based Composites
by Dechao Zhang, Lihua Zhan, Bolin Ma, Jinzhan Guo, Wentao Jin, Xin Hu, Shunming Yao and Guangming Dai
Polymers 2024, 16(17), 2518; https://doi.org/10.3390/polym16172518 - 4 Sep 2024
Viewed by 979
Abstract
The vibration pretreatment–microwave curing process can achieve high-quality molding under low-pressure conditions and is widely used in the curing of resin-based composites. This study investigated the effects of the vibration pretreatment process parameters on the void content and the fiber weight fraction of [...] Read more.
The vibration pretreatment–microwave curing process can achieve high-quality molding under low-pressure conditions and is widely used in the curing of resin-based composites. This study investigated the effects of the vibration pretreatment process parameters on the void content and the fiber weight fraction of T700/TRE231; specifically, their influence on the interlaminar shear strength and impact strength of the composite. Initially, an orthogonal experimental design was employed with interlaminar shear strength as the optimization target, where vibration acceleration was determined as the primary factor and dwell time as the secondary factor. Concurrently, thermogravimetric analysis (TGA) was performed based on process parameters that corresponded to the extremum of interlaminar shear strength, revealing a 2.17% difference in fiber weight fraction among specimens with varying parameters, indicating a minimal effect of fiber weight fraction on mechanical properties. Optical digital microscope (ODM) analysis identified interlaminar large-size voids in specimens treated with vibration energy of 5 g and 15 g, while specimens subjected to a vibration energy of 10 g exhibited numerous small-sized voids within layers, suggesting that vibration acceleration influences void escape pathways. Finally, impact testing revealed the effect of the vibration pretreatment process parameters on the impact strength, implying a positive correlation between interlaminar shear strength and impact strength. Full article
(This article belongs to the Special Issue Advances in Functional Polymers and Composites)
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<p>Vibration pretreatment–microwave curing equipment; (<b>a</b>) composite laminates in the vibration pretreatment equipment; (<b>b</b>) composite laminates in the microwave furnace.</p>
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<p>Vibration pretreatment–microwave curing process curve for T700/TRE231.</p>
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<p>The testing process schematic and specimen dimensions of three-point bending.</p>
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<p>Sampling area for three-point bending and ODM specimens.</p>
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<p>Variation in the weight of TRE231 resin with temperature.</p>
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<p>Drop hammer impact test equipment and impact schematic; (<b>a</b>) drop hammer impact equipment; (<b>b</b>) specimen dimensions and impact clamping; (<b>c</b>) simply supported beam impact process.</p>
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<p>The temperature–viscosity curve of TRE231 resin.</p>
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<p>Weight retention rate of samples under vibration pretreatment parameters.</p>
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<p>The fiber weight fraction of samples under vibiration pretreatment parameters.</p>
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<p>Microscopic morphology of samples under vibration pretreatment parameters.</p>
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<p>The impact time–force curves and impact strength of samples under pretreatment parameters.</p>
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11 pages, 4967 KiB  
Article
Small Punch Testing of a Ti6Al4V Titanium Alloy and Simulations under Different Stress Triaxialities
by Kun Wang, Xilong Zhao and Zeyu Cao
Materials 2024, 17(17), 4203; https://doi.org/10.3390/ma17174203 - 25 Aug 2024
Viewed by 658
Abstract
The mechanical properties of local materials subjected to various stress triaxialities were investigated via self-designed small punch tests and corresponding simulations, which were tailored to the geometry and notch forms of the samples. The finite element model was developed on the basis of [...] Read more.
The mechanical properties of local materials subjected to various stress triaxialities were investigated via self-designed small punch tests and corresponding simulations, which were tailored to the geometry and notch forms of the samples. The finite element model was developed on the basis of the actual test method. After verifying the accuracy of the simulation, the stress, strain, and void volume fraction distributions of the Ti6Al4V titanium alloy under different stress states were compared and analyzed. The results indicate that the mechanical properties of the local material significantly differ during downward pressing depending on the geometric shape. A three-dimensional tensile stress state was observed in the center area, where the void volume fraction was greater than the fracture void volume fraction. The fracture morphology of the samples further confirmed the presence of different stress states. Specifically, the fracture morphology of the globular head samples (with or without U-shaped notches) predominantly featured dimples. Modifying the specimen’s geometry effectively increased stress triaxiality, facilitating the determination of the material’s constitutive relationship under varying stress states. Full article
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<p>Diagram of the specimen for the small punch test and small punch test with a U-shaped notch.</p>
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<p>Three-dimensional model and profile display of small punch testing device.</p>
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<p>Finite element model of small punch test specimen and small punch test specimen with a U-shaped notch.</p>
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<p>Damage evolution of laser welded joints of Ti6Al4V titanium alloy [<a href="#B14-materials-17-04203" class="html-bibr">14</a>].</p>
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<p>Microstructure of Ti6Al4V titanium alloy.</p>
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<p>Macroscopic fracture of small punch test and small punch test specimens with a U-shaped notch.</p>
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<p>Fracture morphology of small punch test and small punch test with a U-shaped notch.</p>
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<p>Comparison of margin of error between small punch test specimen and the small punch test specimen with U-shaped notch obtained by experiment and simulation.</p>
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<p>Mises stress distribution at different iteration steps. (<b>a</b>) Small punch test without a U-shaped notch. (<b>b</b>) Small punch test with a U-shaped notch.</p>
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<p>Evolution distribution of Mises stress under different iteration steps. (<b>a</b>) Small punch test without a U-shaped notch. (<b>b</b>) Small punch test with a U-shaped notch.</p>
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<p>Stress triaxiality and void volume fraction curve of small punch test specimen with or without U-shaped notch.</p>
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12 pages, 2826 KiB  
Article
Water Diffusion in Additively Manufactured Polymers: Effect of Voids
by Boyu Li, Konstantinos P. Baxevanakis and Vadim V. Silberschmidt
J. Compos. Sci. 2024, 8(8), 319; https://doi.org/10.3390/jcs8080319 - 12 Aug 2024
Cited by 1 | Viewed by 737
Abstract
This study investigates the effect of void features in additively manufactured polymers on water diffusion, focusing on polyethylene terephthalate glycol (PETG) composites. The additive manufacturing (AM) of polymers, specifically, material extrusion AM (MEAM), results in manufacturing-induced voids, therefore affecting the water resistance of [...] Read more.
This study investigates the effect of void features in additively manufactured polymers on water diffusion, focusing on polyethylene terephthalate glycol (PETG) composites. The additive manufacturing (AM) of polymers, specifically, material extrusion AM (MEAM), results in manufacturing-induced voids, therefore affecting the water resistance of the printed parts. The research analyses the effects of size, shape, orientation and the hydrophilicity of voids on moisture diffusion in PETG composites employing numerical (finite-element) simulations. Two void types were examined: voids of Type I that retard the moisture propagation and voids of Type II that enhance it. Simulations demonstrate that a higher volume fraction of voids and their orientation with regard to the diffusion direction significantly hinder the moisture transport for Type I voids. Conversely, due to their high diffusivity, Type II voids serve as channels for rapid moisture transmission. Consequently, for such materials, the global diffusion rates mainly depend on the volume fraction of voids rather than their shape. These findings indicate the critical role of voids in the design of AM parts for environments exposed to moisture, such as marine and offshore applications. Understanding the void effects is critical for optimising the durability and performance of MEAM components underwater exposure. Full article
(This article belongs to the Special Issue Progress in Polymer Composites, Volume III)
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<p>Schematic presentation of 3D-printed fibre-reinforced composite manufactured with MEAM [<a href="#B18-jcs-08-00319" class="html-bibr">18</a>].</p>
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<p>Transversal cross-section of 3D-printed PETG specimen reinforced with short carbon fibres.</p>
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<p>Schematics of 0.2 mm × 0.2 mm FEA models: (PM) no void; (C010), (C015) and (C005) circular voids; (T) triangular void; (D) diamond-shape void.</p>
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<p>Evolution of water uptake for different void shapes: (<b>a</b>) Type I; (<b>b</b>) Type II. (<b>c</b>) Effect of solubility on water uptake.</p>
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<p>Spatial evolution of water uptake process for different voids of Type I.</p>
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<p>Mass flow rate at 5 ms calculated with FEM models.</p>
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<p>Spatial evolution of water uptake process for different voids of Type II.</p>
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20 pages, 1595 KiB  
Article
Thermomechanical Analysis of PBF-LB/M AlSi7Mg0.6 with Respect to Rate-Dependent Material Behaviour and Damage Effects
by Lukas Richter, Irina Smolina, Andrzej Pawlak, Daniela Schob, Robert Roszak, Philipp Maasch and Matthias Ziegenhorn
Appl. Mech. 2024, 5(3), 533-552; https://doi.org/10.3390/applmech5030030 - 9 Aug 2024
Viewed by 1039
Abstract
This paper describes the self-heating effects resulting from mechanical deformation in the additively manufactured aluminium alloy AlSi7Mg0.6. The material’s self-heating effect results from irreversible changes in the material’s microstructure that are directly coupled with the inelastic deformations. These processes are highly dissipative, which [...] Read more.
This paper describes the self-heating effects resulting from mechanical deformation in the additively manufactured aluminium alloy AlSi7Mg0.6. The material’s self-heating effect results from irreversible changes in the material’s microstructure that are directly coupled with the inelastic deformations. These processes are highly dissipative, which is reflected in the heat generation of the material. To describe such effects, a numerical framework that combines an elasto-viscoplastic Chaboche model with the Gurson Tvergaard Needleman damage approach is analysed and thermomechanically extended. This paper characterises the sample preparation, the experimental set-up, the development of the thermomechanical approach, and the material model. A user material subroutine applies the complete material model for the finite element software Abaqus 2022. To validate the material model and the parameters, a complex tensile test is performed. In order to check the finite element model, the energy transformation ratio is included in the evaluation. The numerical analyses of the mechanical stress evolution and the self-heating behaviour demonstrate good agreement with the experimental test. In addition, the calculation shows the expected behaviour of the void volume fraction that rises from the initial value of 0.0373% to a higher value under a complex mechanical load. Full article
(This article belongs to the Special Issue Applied Thermodynamics: Modern Developments (2nd Volume))
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<p>(<b>A</b>) Sample orientation during the manufacturing process (<b>B</b>) Sample geometry.</p>
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<p>Concept of a thermomechanical experimental set-up.</p>
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<p>User interface of the post−processing tool.</p>
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<p>Input strain path. <math display="inline"><semantics> <msub> <mi mathvariant="normal">v</mi> <mn>1</mn> </msub> </semantics></math>: complex load path with multiple holding times and an unloading step. <math display="inline"><semantics> <msub> <mi mathvariant="normal">v</mi> <mn>2</mn> </msub> </semantics></math>: tension test with one holing time and an unloading step.</p>
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<p>Results of the thermomechanical experiments: (<b>A</b>) the material’s stress response for the strain rate <math display="inline"><semantics> <msub> <mi mathvariant="normal">v</mi> <mn>1</mn> </msub> </semantics></math>; (<b>B</b>) the material’s stress response for the strain rate <math display="inline"><semantics> <msub> <mi mathvariant="normal">v</mi> <mn>2</mn> </msub> </semantics></math>; (<b>C</b>) the material’s temperature response for the strain rate <math display="inline"><semantics> <msub> <mi mathvariant="normal">v</mi> <mn>1</mn> </msub> </semantics></math>; (<b>D</b>) the material’s temperature response for the strain rate <math display="inline"><semantics> <msub> <mi mathvariant="normal">v</mi> <mn>2</mn> </msub> </semantics></math>.</p>
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<p>Boundary conditions: (<b>A</b>) displacement boundary conditions and radiator, (<b>B</b>) surface film condition on clamped area, and (<b>C</b>) surface film condition for air convection.</p>
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<p><span class="html-small-caps">FE</span>−Results at t = 26 s: (<b>A</b>) region of interest, (<b>B</b>) stress distribution in the y-direction, and (<b>C</b>) temperature distribution.</p>
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<p>Thermomechanical material behaviour for loading path <math display="inline"><semantics> <msub> <mi mathvariant="normal">v</mi> <mn>1</mn> </msub> </semantics></math> (<b>A</b>) Mechanical behaviour (<b>B</b>) Temperature behaviour.</p>
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<p>Energy behaviour for loading path <math display="inline"><semantics> <msub> <mi mathvariant="normal">v</mi> <mn>1</mn> </msub> </semantics></math> (<b>A</b>) Energy transformation ratio (<b>B</b>) Stored energy of cold work.</p>
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<p>Numerical void volume fraction <math display="inline"><semantics> <msup> <mover accent="true"> <mi>f</mi> <mo stretchy="false">^</mo> </mover> <mo>*</mo> </msup> </semantics></math> for loading path <math display="inline"><semantics> <msub> <mi mathvariant="normal">v</mi> <mn>1</mn> </msub> </semantics></math>.</p>
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<p>Thermomechanical material behaviour for the loading path <math display="inline"><semantics> <msub> <mi mathvariant="normal">v</mi> <mn>2</mn> </msub> </semantics></math> (<b>A</b>) Mechanical behaviour (<b>B</b>) Temperature behaviour.</p>
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