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23 pages, 11763 KiB  
Article
Strength and Stiffness of Corrugated Plates Subjected to Bending
by Yuki Yoshino and Yoshihiro Kimura
Buildings 2025, 15(3), 469; https://doi.org/10.3390/buildings15030469 (registering DOI) - 2 Feb 2025
Abstract
When a beam twists, localized flexural deformation in the plate occurs at the connection between the corrugated plate and the beam, resulting in a reduction in the bracing stiffness of the corrugated plate. To accurately assess the bracing stiffness of a beam with [...] Read more.
When a beam twists, localized flexural deformation in the plate occurs at the connection between the corrugated plate and the beam, resulting in a reduction in the bracing stiffness of the corrugated plate. To accurately assess the bracing stiffness of a beam with a connected corrugated plate, it is essential to independently determine the bending stiffness of the corrugated plate and the rotational stiffness of the plate–beam joint. This study conducts bending experiments on corrugated plates subjected to in-plane bending to elucidate the stress mechanisms within the plates. The influence of the width-to-thickness ratio on the bending load and bending stiffness of the corrugated plates is examined using the width-to-thickness ratio regulations of various countries. The findings revealed that when the width-to-thickness ratio of the corrugated plates used in this study was lower than the threshold specified in the Eurocode, the bending load at the onset of stiffness degradation was approximately 80% of the maximum bending load, and the initial bending stiffness corresponded closely to the theoretical value. Conversely, when the web width-to-thickness ratio of the corrugated plates exceeded the limit prescribed in the Eurocode, it was demonstrated that the maximum bending load decreased to approximately 50% of the yield load, and the initial bending stiffness was reduced to about 95% of the theoretical value. Full article
(This article belongs to the Section Building Structures)
Show Figures

Figure 1

Figure 1
<p>Corrugated plates for actual structures. (<b>a</b>) Example A and (<b>b</b>) example B.</p>
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<p>Spring model of continuous stiffening for lateral buckling deformation of H beams.</p>
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<p>Connector at the edge of the specimen. (<b>a</b>) Elevation of <span class="html-italic">y</span>–<span class="html-italic">z</span> plane (A–A′ line), (<b>b</b>) elevation of <span class="html-italic">x</span>–<span class="html-italic">y</span> plane (C–C′ line), (<b>c</b>) ground plan, and (<b>d</b>) names of the parts of the corrugated plate.</p>
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<p>Specimen. (<b>a</b>) Type A and (<b>b</b>) Type B.</p>
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<p>Detail of loading beam. (<b>a</b>) Type A: (<b>a-1</b>) <span class="html-italic">x</span>–<span class="html-italic">z</span> plane, (<b>a-2</b>) <span class="html-italic">y</span>–<span class="html-italic">z</span> plane, (<b>a-3</b>) <span class="html-italic">x</span>–<span class="html-italic">y</span> plane of one unit, and (<b>a-4</b>) <span class="html-italic">x</span>–<span class="html-italic">y</span> plane of three units. (<b>b</b>) Type B: (<b>b-1</b>) <span class="html-italic">x</span>–<span class="html-italic">z</span> plane, (<b>b-2</b>) <span class="html-italic">y</span>–<span class="html-italic">z</span> plane, (<b>b-3</b>) <span class="html-italic">x</span>–<span class="html-italic">y</span> plane of one unit, and (<b>b-4</b>) <span class="html-italic">x</span>–<span class="html-italic">y</span> plane of three units.</p>
Full article ">Figure 6
<p>Strain measurement position. (<b>a</b>) Type A and (<b>b</b>) Type B.</p>
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<p>Imperfection deflection before experimentation: (<b>a</b>) At −L/18. (<b>b</b>) At +L/18.</p>
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<p>Load–displacement curve under monotonic loading. (<b>a</b>) Number of Units and (<b>b</b>) differences in the corrugated plate cross-section.</p>
Full article ">Figure 9
<p>Axial strain. (<b>a</b>) A1-0.6-18, (<b>b</b>) A3-0.6-18, and (<b>c</b>) B3-0.6-12.</p>
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<p>Bending strain. (<b>a</b>) A1-0.6-18, (<b>b</b>) A3-0.6-18, and (<b>c</b>) B3-0.6-12.</p>
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<p>Axial strain of the <span class="html-italic">y–z</span> plane. (<b>a</b>) A1-0.6-18, (<b>b</b>) A3-0.6-18, and (<b>c</b>) B3-0.6-12.</p>
Full article ">Figure 12
<p>Ultimate deformation of specimens. (<b>a</b>) Type A (A3-0.6-18) and (<b>b</b>) Type B (B3-0.6-12).</p>
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<p>Analysis model. (<b>a</b>) Type A and (<b>b</b>) Type B.</p>
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<p>A bilinear isotropic hardening law.</p>
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<p>The static Riks method in the ABAQUS software.</p>
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<p>Comparison of experimentally obtained and analytically obtained results. (<b>a</b>) Load–displacement curve under monotonic loading and (<b>b</b>) axial strain of the <span class="html-italic">y</span>–<span class="html-italic">z</span> plane: (<b>b-1</b>) A1-0.6-18, (<b>b-2</b>) B1-0.6-12, (<b>b-3</b>) A3-1.0-18, and (<b>b-4</b>) B3-0.6-12.</p>
Full article ">Figure 17
<p>Ultimate deformation of experimentally obtained and analytically obtained results. (<b>a</b>) A1-0.6-18, (<b>b</b>) A3-0.6-18, and (<b>c</b>) B3-0.6-12.</p>
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<p>The effect of the number of units on the load ratio.</p>
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<p>Comparison of initial bending stiffness and shear span ratio of corrugated plates. (<b>a</b>) Image of bending deformation for corrugated plates and (<b>b</b>) shear span ratio.</p>
Full article ">Figure 20
<p>Comparison of the road ratio and width-to-thickness ratio. (<b>a</b>) Width-to-thickness ratio of flange and (<b>b</b>) width-to-thickness ratio of web [<a href="#B20-buildings-15-00469" class="html-bibr">20</a>,<a href="#B21-buildings-15-00469" class="html-bibr">21</a>,<a href="#B22-buildings-15-00469" class="html-bibr">22</a>].</p>
Full article ">Figure 21
<p>Comparison of the bending stiffness and width-to-thickness ratio of corrugated plates. (<b>a</b>) Definition of stiffness, (<b>b</b>) width-to-thickness ratio of flange, and (<b>c</b>) width-to-thickness ratio of web [<a href="#B20-buildings-15-00469" class="html-bibr">20</a>,<a href="#B21-buildings-15-00469" class="html-bibr">21</a>,<a href="#B22-buildings-15-00469" class="html-bibr">22</a>].</p>
Full article ">
16 pages, 10367 KiB  
Article
Influence of the Deformation Degree of Combined Loadings on the Structural and Mechanical Properties of Stainless Steels
by Magdalena Gabriela Huțanu, Liviu Andrușcă, Marcelin Benchea, Mihai-Adrian Bernevig, Dragoș Cristian Achiței, Ștefan-Constantin Lupescu, Gheorghe Bădărău and Nicanor Cimpoeșu
J. Manuf. Mater. Process. 2025, 9(2), 45; https://doi.org/10.3390/jmmp9020045 (registering DOI) - 1 Feb 2025
Viewed by 239
Abstract
Stainless steels have many practical applications requiring various mechanical or chemical demands in the working environment. By optimizing a device used in mechanical experiments for torsional loading, several cylindrical samples were tested (both ends twisted with the same torque value in opposite directions) [...] Read more.
Stainless steels have many practical applications requiring various mechanical or chemical demands in the working environment. By optimizing a device used in mechanical experiments for torsional loading, several cylindrical samples were tested (both ends twisted with the same torque value in opposite directions) of 316L stainless steel (SS) to evaluate changes in the structural, chemical, and mechanical characteristics. Initially, the experimental samples were pre-loaded by tension in the elastic range (6%) and then subjected to torsion (772°) at different rates: 5, 10, and 20 mm/min. The experimental sequence consisted of a combined loading protocol with an initial tensile test followed by a subsequent torsional test. Two reference tests were performed by fracturing the samples in both torsion and tension to determine the mechanical strength parameters. The macro- and microstructural evolution of the samples as a function of the torsional degree was followed by scanning electron microscopy. The microhardness modification of the material was observed because of the strain (the microhardness variation from the center of the disk sample to the edge was also monitored). Structurally, all samples showed grain size changes because of torsional/compressive deformation zones and an increase in the degree of grain boundary misorientation. From the tensile and torsional behaviors of 316L SS and the structural results obtained, it was concluded that these materials are suitable for complex stress states in the elasto-plastic range through tensile and torsion. A reduction in Young’s modulus of up to four times the initial value at medium and high stress rates was observed when complex stresses were applied. Full article
(This article belongs to the Special Issue Advances in Metal Forming and Additive Manufacturing)
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Figure 1

Figure 1
<p>Experimental set-up used for the torsion test in (<b>a</b>); 3D model of the torsion device in (<b>b</b>); schematic presentation of the combined stress in (<b>c</b>); main dimensions of the specimen in (<b>d</b>).</p>
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<p>Tensile/torsion to failure curves of austenitic 316L steel.</p>
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<p>SEM micrographs of the tensile fracture (<b>a</b>) 2D: 100×, 250× and 1000× and (<b>b</b>) 3D.</p>
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<p>SEM micrographs of the torsion fracture (<b>a</b>) 2D: 100×, 250×, and 1000× and (<b>b</b>) 3D.</p>
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<p>Combined mechanical test stages applied to experimental materials (<b>a</b>) tensile and (<b>b</b>) torsion.</p>
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<p>SEM microstructures (<b>a</b>) middle area and (<b>b</b>) edge for the initial and twisted samples with 5, 10, and 20 mm/min from left to right.</p>
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<p>Mechanical properties of the experimental samples (<b>a</b>) Load vs. depth variation, (<b>b</b>) dwell time vs. twist rate variation, (<b>c</b>) hardness vs. twist rate variation, (<b>d</b>) Indentation modulus vs. twist rate variation and (<b>e</b>) contact stiffness vs. twist rate variation.</p>
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<p>The friction coefficient variation with distance (<b>a</b>) center against the initial and (<b>b</b>) edge against the initial.</p>
Full article ">Figure 9
<p>3D images of the scratch (<b>a</b>) initial, (<b>b</b>) tensile + torsion rate of 5mm/min, (<b>c</b>) tensile + torsion rate of 10 mm/min, and (<b>d</b>) tensile + torsion rate of 20 mm/min.</p>
Full article ">
16 pages, 7739 KiB  
Article
Development of Short Jute Fiber-Reinforced Thermoplastic Pre-Preg Tapes
by Mengyuan Dun, Haitao Fu, Jianxiu Hao and Weihong Wang
Polymers 2025, 17(3), 388; https://doi.org/10.3390/polym17030388 - 31 Jan 2025
Viewed by 418
Abstract
Jute fibers are renewable, light, and strong, allowing them to be considered as attractive materials in composite manufacturing. In the present work, a simple and effective method for preparing continuous pre-preg tapes from short jute fiber bundles (without twist) is developed and its [...] Read more.
Jute fibers are renewable, light, and strong, allowing them to be considered as attractive materials in composite manufacturing. In the present work, a simple and effective method for preparing continuous pre-preg tapes from short jute fiber bundles (without twist) is developed and its application in winding forming is evaluated. Linear low-density polyethylene film (LLDPE) with good flexibility and weather resistance was used as the thermoplastic matrix; jute fiber bundles were first spread parallel to each other on an LLDPE film and then rolled up to form a pre-roll. The pre-roll enclosing fiber bundles was hot-pressed in a designed mold to form a pre-preg tape, where the fiber bundles were more parallel to the tape than the fibers in twine. Although the untwisted structure exhibited a lower tensile strength for the fiber bundle, it could be processed into a continuous pre-preg with higher tensile strength than the jute twine-impregnated pre-preg. This is based on the good impregnation of the short fiber bundle and its unidirectional, uniform strengthening in the continuous pre-preg. The tensile strength and modulus of the fiber bundle-reinforced pre-preg increased by 16.70% and 257.14%, respectively, compared with jute twine-reinforced pre-preg (within the fiber proportion of 40.wt%). When applied to winding, the fiber bundle-reinforced pre-preg showed advantages of interlayer fusion, surface flatness, and ring stiffness. In contrast, the twisted continuous structure did not retain its advantage in pre-preg. The development of pre-preg tapes by discontinuous fibers might be a good way for utilizing natural fibers in the field of green engineering due to its diverse secondary processing. Full article
(This article belongs to the Special Issue Fiber Reinforced Polymers: Manufacture, Properties and Applications)
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Figure 1

Figure 1
<p>Twisted jute twine and carded short fiber bundle.</p>
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<p>Frequency distribution of jute short fiber: (<b>a</b>) length and (<b>b</b>) diameter in bundles.</p>
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<p>Preparation of jute fiber bundle/LLDPE and jute twine/LLDPE pre-preg tape.</p>
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<p>Fixture for longitudinal tensile testing (<b>a</b>), annular tensile testing (<b>b</b>), and ring stiffness testing (<b>c</b>) of the developed pre-preg.</p>
Full article ">Figure 5
<p>Microscopic observation of jute fiber distribution in continuous pre-preg tape: JFB/LLDPE pre-preg with jute mass fractions of 20% (<b>a</b>) and 50% (<b>b</b>); JT/LLDPE pre-preg with jute mass fractions of 20% (<b>c</b>) and 50% (<b>d</b>).</p>
Full article ">Figure 6
<p>Cross-section of JFB/LLDPE pre-pregs under scanning electron microscopy with mass fractions of 20% (<b>a</b>,<b>c</b>,<b>e</b>) and 50% (<b>b</b>,<b>d</b>,<b>f</b>).</p>
Full article ">Figure 7
<p>Fracture surface observation under scanning electron microscopy: (<b>a</b>,<b>c</b>,<b>e</b>) JT/LLDPE pre-pregs with mass fractions of 20%, (<b>b</b>,<b>d</b>,<b>f</b>) JT/LLDPE pre-pregs with mass fractions of 50%.</p>
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<p>(<b>a</b>) Tensile breaking strength–strain curves and (<b>b</b>) force diagram of twisted jute twine and jute fiber bundle under tensile load.</p>
Full article ">Figure 9
<p>Longitudinal tensile properties of jute fiber bundle- and jute twine-reinforced pre-pregs: (<b>A</b>) tensile strength (red line represents the tensile strength of pure LLDPE), (<b>B</b>) tensile modulus. Note: When the same letter appears at the top of the column, there is no significant difference between any two groups. The letters mean the same in the following figures.</p>
Full article ">Figure 10
<p>Filament winding process of continuous JFB/LLDPE and JT/LLDPE pre-preg tapes, as well as images of the obtained winding ring products (including enlarged cross-sectional morphology).</p>
Full article ">Figure 11
<p>(<b>a</b>) Annular tensile stress–strain curves and (<b>b</b>) annular compression load–displacement curves of 40JFB/LLDPE and 40JT/LLDPE winding rings.</p>
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<p>(<b>a</b>) Bamboo short fiber bundle/LLDPE pre-preg tape, (<b>b</b>–<b>d</b>) pictures of small-diameter pipes (outer diameters: 40 mm) wound by 40 wt% jute short fiber bundle/LLDPE pre-preg tape and 40 wt% bamboo short fiber bundle/LLDPE pre-preg tape.</p>
Full article ">
10 pages, 3066 KiB  
Article
Spontaneous Emission Mediated by Moiré Hyperbolic Metasurfaces
by Yuying Liu, Zhanrong Yang, Tongbiao Wang, Jianrong Yang, Tianbao Yu and Qinghua Liao
Nanomaterials 2025, 15(3), 228; https://doi.org/10.3390/nano15030228 - 31 Jan 2025
Viewed by 278
Abstract
We investigate the spontaneous emission of a quantum emitter (QE) placed near the twisted hyperbolic metasurfaces (HMTSs) made of graphene strips. We demonstrate that the spontaneous emission can be enhanced distinctly due to the existence of moiré hyperbolic plasmon polaritons (HPPs) supported by [...] Read more.
We investigate the spontaneous emission of a quantum emitter (QE) placed near the twisted hyperbolic metasurfaces (HMTSs) made of graphene strips. We demonstrate that the spontaneous emission can be enhanced distinctly due to the existence of moiré hyperbolic plasmon polaritons (HPPs) supported by the twisted HMTSs. Moreover, the spontaneous emission decay rate can be efficiently modulated by the chemical potential of graphene, the thickness of the dielectric spacer, and the twist angle between two HMTSs. The maximum spontaneous emission is achieved when topological transition occurs. The spontaneous emission will be enhanced as the thickness of the dielectric spacer increases for most cases. In particular, the twisted HMTSs make it possible to flexibly modify the spontaneous emission through the external field. The findings in this work not only extend past studies of twisted photonic structures but also have important applications in optical sensing and integrated photonics. Full article
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Figure 1

Figure 1
<p>(<b>a</b>) The schematic of a QE placed near the surface of the twisted HMTSs with distance <math display="inline"><semantics> <mrow> <mi>z</mi> </mrow> </semantics></math>. The thickness of spacer between two HMTSs is <math display="inline"><semantics> <mrow> <mi>d</mi> </mrow> </semantics></math>. (<b>b</b>) The top view of the system with a twist angle <math display="inline"><semantics> <mrow> <mi>ϕ</mi> </mrow> </semantics></math> with respect to the <span class="html-italic">x</span>-axis. (<b>c</b>) The top view of the system without twist between HMTSs. <math display="inline"><semantics> <mrow> <mi>W</mi> </mrow> </semantics></math> and <math display="inline"><semantics> <mrow> <mi>P</mi> </mrow> </semantics></math> are the strip width and period, respectively.</p>
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<p>The imaginary parts of reflection coefficients about moiré HMTSs with twist angles of (<b>a</b>) <math display="inline"><semantics> <mrow> <mi>ϕ</mi> <mo>=</mo> <msup> <mrow> <mn>0</mn> </mrow> <mrow> <mo>°</mo> </mrow> </msup> </mrow> </semantics></math>, (<b>b</b>) <math display="inline"><semantics> <mrow> <mi>ϕ</mi> <mo>=</mo> <msup> <mrow> <mn>30</mn> </mrow> <mrow> <mo>°</mo> </mrow> </msup> </mrow> </semantics></math>, (<b>c</b>) <math display="inline"><semantics> <mrow> <mi>ϕ</mi> <mo>=</mo> <msup> <mrow> <mn>60</mn> </mrow> <mrow> <mo>°</mo> </mrow> </msup> </mrow> </semantics></math>, (<b>d</b>) <math display="inline"><semantics> <mrow> <mi>ϕ</mi> <mo>=</mo> <msup> <mrow> <mn>80</mn> </mrow> <mrow> <mo>°</mo> </mrow> </msup> </mrow> </semantics></math>, and (<b>e</b>) <math display="inline"><semantics> <mrow> <mi>ϕ</mi> <mo>=</mo> <msup> <mrow> <mn>90</mn> </mrow> <mrow> <mo>°</mo> </mrow> </msup> </mrow> </semantics></math>. The parameters are set to <math display="inline"><semantics> <mrow> <mi>d</mi> <mo>=</mo> <mn>1.0</mn> </mrow> </semantics></math> nm, <math display="inline"><semantics> <mrow> <mi>μ</mi> <mo>=</mo> <mn>0.4</mn> </mrow> </semantics></math> eV, <math display="inline"><semantics> <mrow> <mi>W</mi> <mo>=</mo> <mn>6</mn> </mrow> </semantics></math> nm, <math display="inline"><semantics> <mrow> <mi>P</mi> <mo>=</mo> <mn>10</mn> </mrow> </semantics></math> nm, and <math display="inline"><semantics> <mrow> <msub> <mrow> <mi>ε</mi> </mrow> <mrow> <mn>2</mn> </mrow> </msub> <mo>=</mo> <mn>1</mn> </mrow> </semantics></math>. The angular frequency is <math display="inline"><semantics> <mrow> <mn>0.15</mn> <mo> </mo> <mi mathvariant="normal">e</mi> <mi mathvariant="normal">V</mi> <mo>/</mo> <mi mathvariant="normal">ℏ</mi> </mrow> </semantics></math>. The green lines denote the dispersion relation. (<b>f</b>) Topological transition regions as a function of frequency and twist angles. The green dashed line represents the topological transition angle.</p>
Full article ">Figure 3
<p>The Purcell factor of the QE as a function of twist angle for different (<b>a</b>) thicknesses of dielectric spacer with a fixed chemical potential of 0.4 eV and (<b>b</b>) the chemical potential of graphene with a fixed thickness of a dielectric spacer of 10 nm. The distance <math display="inline"><semantics> <mrow> <mi>z</mi> </mrow> </semantics></math> is fixed at 10 nm. The period <math display="inline"><semantics> <mrow> <mi>P</mi> </mrow> </semantics></math> is 10 nm and the filling factor is <math display="inline"><semantics> <mrow> <mi>f</mi> <mo>=</mo> <mn>0.6</mn> </mrow> </semantics></math> for the moiré HMTSs. The transition frequencies are set to 0.15 <math display="inline"><semantics> <mrow> <mrow> <mrow> <mi mathvariant="normal">e</mi> <mi mathvariant="normal">V</mi> </mrow> <mo>/</mo> <mrow> <mi mathvariant="normal">ℏ</mi> </mrow> </mrow> </mrow> </semantics></math>.</p>
Full article ">Figure 4
<p>The Purcell factor of the QE as a function of the thickness of the dielectric spacer for different twist angles. The filling fractions are 0.2, 0.4, 0.6, and 0.8 from (<b>a</b>–<b>d</b>), respectively. The twist angles in (<b>b</b>–<b>d</b>) are the same as those in (<b>a</b>). The distance <math display="inline"><semantics> <mrow> <mi>z</mi> </mrow> </semantics></math> is set to 10 nm, the period <math display="inline"><semantics> <mrow> <mi>P</mi> </mrow> </semantics></math> is 10 nm, the chemical potential is 0.4 eV, and the transition frequency is 0.15 <math display="inline"><semantics> <mrow> <mrow> <mrow> <mi mathvariant="normal">e</mi> <mi mathvariant="normal">V</mi> </mrow> <mo>/</mo> <mrow> <mi mathvariant="normal">ℏ</mi> </mrow> </mrow> </mrow> </semantics></math>.</p>
Full article ">Figure 5
<p>The Purcell factor of the QE as a function of chemical potential for different filling factors. The distance <math display="inline"><semantics> <mrow> <mi>z</mi> </mrow> </semantics></math> is set to 10 nm, the period <math display="inline"><semantics> <mrow> <mi>P</mi> </mrow> </semantics></math> is 10 nm, the thickness of the spacer is 10 nm, and the transition frequency is 0.15 <math display="inline"><semantics> <mrow> <mrow> <mrow> <mi mathvariant="normal">e</mi> <mi mathvariant="normal">V</mi> </mrow> <mo>/</mo> <mrow> <mi mathvariant="normal">ℏ</mi> </mrow> </mrow> </mrow> </semantics></math>.</p>
Full article ">Figure 6
<p>The Purcell factor as a function of distance <math display="inline"><semantics> <mrow> <mi>z</mi> </mrow> </semantics></math> for different twist angles. The filling fractions in (<b>a</b>–<b>d</b>) are <math display="inline"><semantics> <mrow> <mi>f</mi> <mo>=</mo> <mn>0.2</mn> </mrow> </semantics></math>, <math display="inline"><semantics> <mrow> <mi>f</mi> <mo>=</mo> <mn>0.4</mn> </mrow> </semantics></math>, <math display="inline"><semantics> <mrow> <mi>f</mi> <mo>=</mo> <mn>0.6</mn> </mrow> </semantics></math>, and <math display="inline"><semantics> <mrow> <mi>f</mi> <mo>=</mo> <mn>0.8</mn> </mrow> </semantics></math>, respectively. The twist angles in (<b>b</b>–<b>d</b>) are the same as those used in (<b>a</b>). The period is fixed at <math display="inline"><semantics> <mrow> <mi>P</mi> <mo>=</mo> <mn>10</mn> </mrow> </semantics></math> nm, and the other parameters are the same with those used in <a href="#nanomaterials-15-00228-f002" class="html-fig">Figure 2</a>.</p>
Full article ">
14 pages, 2614 KiB  
Article
Left Ventricular Twist and Circumferential Strain from MRI Tagging Predict Early Cardiovascular Disease in Duchenne Muscular Dystrophy
by Zhan-Qiu Liu, Patrick Magrath, Nyasha G. Maforo, Michael Loecher, Holden H. Wu, Ashley Prosper, Pierangelo Renella, Nancy Halnon and Daniel B. Ennis
Diagnostics 2025, 15(3), 326; https://doi.org/10.3390/diagnostics15030326 - 30 Jan 2025
Viewed by 238
Abstract
Background: Duchenne Muscular Dystrophy (DMD) is a prevalent fatal genetic disorder, and heart failure is the leading cause of mortality. Peak left ventricular (LV) circumferential strain (Ecc), twist, and circumferential-longitudinal shear angle (θCL) are promising biomarkers for the improved [...] Read more.
Background: Duchenne Muscular Dystrophy (DMD) is a prevalent fatal genetic disorder, and heart failure is the leading cause of mortality. Peak left ventricular (LV) circumferential strain (Ecc), twist, and circumferential-longitudinal shear angle (θCL) are promising biomarkers for the improved and early diagnosis of incipient heart failure. Our goals were as follows: 1) to characterize a spectrum of functional and rotational LV biomarkers in boys with DMD compared with healthy age-matched controls; and 2) to identify LV biomarkers of early cardiomyopathy in the absence of abnormal LVEF or LGE. Methods: Boys with DMD (N = 43) and age-matched healthy volunteers (N = 16) were prospectively enrolled and underwent a 3T CMR exam after obtaining informed consent. Breath-held MRI tagging was used to estimate left ventricular Ecc at the mid-ventricular level as well as the twist, torsion, and θCL between basal and apical LV short-axis slices. A two-tailed t-test with unequal variance was used to test group-wise differences. Multiple comparisons were performed with Holm–Sidak post hoc correction. Multiple-regression analysis was used to test for correlations among biomarkers. A binomial logistic regression model assessed each biomarker’s ability to distinguish the following: (1) healthy volunteers vs. DMD patients, (2) healthy volunteers vs. LGE(−) DMD patients, and (3) LGE(−) DMD patients vs. LGE(+) DMD patients. Results: There was a significant impairment in the peak mid-wall Ecc [−17.0 ± 4.2% vs. −19.5 ± 1.9%, p < 7.8 × 10−3], peak LV twist (10.4 ± 4.3° vs. 15.6 ± 3.1°, p < 8.1 × 10−4), and peak LV torsion (2.03 ± 0.82°/mm vs. 2.8 ± 0.5°/mm, p < 2.6 × 10−3) of LGE(−) DMD patients when compared to healthy volunteers. There was a further significant reduction in the Ecc, twist, torsion, and θCL for LGE(+) DMD patients when compared to LGE(−) DMD patients. In the LGE(+) DMD patients, age significantly correlated with LVEF (r2 = 0.42, p = 9 × 10−3), peak mid-wall Ecc (r2 = 0.27, p = 0.046), peak LV Twist (r2 = 0.24, p = 0.06), peak LV torsion (r2 = 0.28, p = 0.04), and peak LV θCL (r2 = 0.23, p = 0.07). In the LGE(−) DMD patients, only the peak mid-wall Ecc was significantly correlated with age (r2 = 0.25, p = 0.006). The peak LV twist outperformed the peak mid-wall LV Ecc and EF in distinguishing DMD patients from healthy volunteer groups (AUC = 0.88, 0.80, and 0.72), as well as in distinguishing LGE(−) DMD patients from healthy volunteers (AUC = 0.83, 0.74, and 0.62). The peak LV twist and peak mid-wall LV Ecc performed similarly in distinguishing the LGE(−) and LGE(+) DMD cohorts (AUC = 0.74, 0.77, and 0.79). Conclusions: The peak mid-wall LV Ecc, peak LV twist, peak LV torsion, and peak LV θCL were significantly impaired in advance of the decreased LVEF and the development of focal myocardial fibrosis in boys with DMD and therefore were apparent prior to significant irreversible injury. Full article
(This article belongs to the Special Issue New Trends in Cardiovascular Imaging)
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<p>Representative tagged MRI images and measurement of LV rotational biomarkers: (<b>A</b>) Representative tagged MRI images. (<b>B</b>) Derivation of peak LV twist, torsion, and CL-shear angle (θ<sub>CL</sub>).</p>
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<p>Age vs. LV ejection fraction (LVEF) in volunteers (blue circles), LGE(−) patients with DMD (orange squares), and LGE(+) patients with DMD (yellow diamonds). Only LGE(+) patients with DMD have a significant negative correlation. There is no significant relationship between age and LVEF for healthy volunteers and LGE(−) patients with DMD.</p>
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<p>Age vs. the peak mid-wall LV Ecc and peak LV twist in healthy volunteers (blue circle), LGE(−) DMD boys (orange square), and LGE(+) DMD boys (yellow diamonds): (<b>A</b>) peak mid-wall LV Ecc vs. age; and (<b>B</b>) peak LV twist vs. age. A significant correlation is identified with multiple linear regression (solid lines), and non-significant correlations are shown (dashed lines) for completeness. The peak mid-wall LV Ecc and peak LV twist decrease moderately with age in the LGE(+) DMD patients. The peak mid-wall LV Ecc also decreases with age in the LGE(−) DMD patients. There is no correlation with age in the healthy volunteer group for either measure or for the peak LV twist in the LGE(−) DMD patients.</p>
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<p>Peak mid-wall LV E<sub>cc</sub> and peak LV rotational biomarkers as functions of LVEF in healthy volunteers (blue circles), LGE(−) DMD boys (orange squares), and LGE(+) DMD boys (yellow diamonds): (<b>A</b>) peak mid-wall LV E<sub>cc</sub>; (<b>B</b>) peak LV twist; (<b>C</b>) peak LV torsion; and (<b>D</b>) peak LV θ<sub>CL</sub>. Significant correlations are identified with multiple linear regression (solid lines). Non-significant correlations are shown (dashed lines) for completeness. All cardiac MRI biomarkers are significantly correlated with LVEF for the LGE(+) patients with DMD. There is also a significant correlation between the peak LV twist and LVEF in the healthy volunteers. For all other groups, and, particularly, the LGE(−) DMD patients, these cardiac MRI biomarkers are uncorrelated with EF.</p>
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<p>ROCs for the binomial logistic regression analysis differentiating (<b>A</b>) all patients with DMD vs. healthy volunteers, (<b>B</b>) LGE(−) patients with DMD vs. healthy volunteers, and (<b>C</b>) LGE(+) vs. LGE(−) DMD patients. (<b>D</b>) The area under the ROC curve for the peak LV twist, peak mid-wall E<sub>cc</sub>, and EF for each comparison. The peak LV twist (followed by E<sub>cc</sub>) is the best biomarker to differentiate both all patients with DMD from healthy volunteers and LGE(−) patients with DMD from healthy volunteers. All cardiac MRI biomarkers are effective at differentiating patients with DMD and advanced cardiac disease (LGE(+)) from DMD patients without advanced cardiac disease (LGE(−)).</p>
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<p>Correlation analysis and receiver operator characteristic (ROC) curves: (<b>A</b>) The correlations between the peak mid-wall LV E<sub>cc</sub> and peak LV twist. Significant correlations identified with multiple linear regression are shown as solid lines. The peak LV twist and peak mid-wall E<sub>cc</sub> are highly correlated in the LGE(+) DMD patients and moderately correlated in the LGE(−) DMD and healthy volunteer groups. (<b>B</b>) ROC analysis of the binary logistic regression generalized linear model incorporating the combination of the peak mid-wall LV E<sub>cc</sub>, peak LV twist, and LVEF together, and the logistic regression for each shown separately. Only the peak LV twist was a significant co-factor in the combined model, and the AUC of the combined model was marginally less than that of the peak LV twist alone (0.830 vs. 0.837).</p>
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63 pages, 14085 KiB  
Review
Insights from the Last Decade in Computational Fluid Dynamics (CFD) Design and Performance Enhancement of Darrieus Wind Turbines
by Saïf ed-Dîn Fertahi, Shafiqur Rehman, Ernesto Benini, Khadija Lahrech, Abderrahim Samaouali, Asmae Arbaoui, Imad Kadiri and Rachid Agounoun
Processes 2025, 13(2), 370; https://doi.org/10.3390/pr13020370 - 28 Jan 2025
Viewed by 371
Abstract
This review provides an analysis of advancements in the design and performance assessment of Darrieus wind turbines over the past decade, with a focus on the contributions of computational fluid dynamics (CFD) to this field. The primary objective is to present insights from [...] Read more.
This review provides an analysis of advancements in the design and performance assessment of Darrieus wind turbines over the past decade, with a focus on the contributions of computational fluid dynamics (CFD) to this field. The primary objective is to present insights from studies conducted between 2014 and 2024, emphasizing the enhancement of Darrieus wind turbine performance through various technological innovations. The research methodology employed for this review includes a critical analysis of published articles related to Darrieus turbines. The focus on the period from 2014 to 2024 was considered to highlight recent parametric CFD studies on Darrieus turbines, avoiding overlap with previously published reviews and maintaining originality relative to existing review works in the literature. By synthesizing a collection of articles, the review discusses a wide range of recent investigations utilizing CFD modeling techniques, including both 2D and 3D simulations. These studies predominantly utilize the “Ansys-Fluent” V12.0 and “STAR CCM+” V9.02 solvers to evaluate the aerodynamic performance of Darrieus rotors. Technological advancements focus on modifying the geometry of Darrieus, including alterations to blade profiles, chord length, rotor diameter, number of blades, turbine height, rotor solidity, and the integration of multiple rotors in various configurations. Additionally, the incorporation of flow deflectors, the use of advanced blade shapes, such as V-shaped or twisted blades, and the application of an opening ratio on the blades are explored to enhance rotor efficiency. The review highlights the significant impact of these geometric modifications on key performance metrics, particularly the moment and power coefficients. A dedicated section presents CFD-derived visualizations, including vorticity fields, turbulence contours illustrated through the Q-criterion, velocity vectors, and dynamic pressure contours. These visualizations provide a description of the flow structures around the modified Darrieus rotors. Moreover, the review includes an analysis of the dynamic performance curves of Darrieus, which show improvements resulting from the modifications of the baseline design. This analysis covers the evolution of pressure coefficients, moment coefficients, and the increased power output of Darrieus. Full article
(This article belongs to the Special Issue Turbulence Models for Turbomachinery)
18 pages, 1612 KiB  
Article
Liquid–Liquid Flow and Mass Transfer Enhancement in Tube-in-Tube Millireactors with Structured Inserts and Advanced Inlet Designs
by Feng Zhu, Xingxing Pan, Xichun Cao, Yandan Chen, Rijie Wang, Jiande Lin and Hanyang Liu
Fluids 2025, 10(2), 26; https://doi.org/10.3390/fluids10020026 - 24 Jan 2025
Viewed by 321
Abstract
Liquid–liquid mass transfer is crucial in chemical processes like extraction and desulfurization. Traditional tube-in-tube millireactors often overlook internal flow dynamics, focusing instead on entry modifications. This study explores mass transfer enhancement through structured inserts (twisted tapes, multi-blades) and inlet designs (multi-hole injectors, T-mixers). [...] Read more.
Liquid–liquid mass transfer is crucial in chemical processes like extraction and desulfurization. Traditional tube-in-tube millireactors often overlook internal flow dynamics, focusing instead on entry modifications. This study explores mass transfer enhancement through structured inserts (twisted tapes, multi-blades) and inlet designs (multi-hole injectors, T-mixers). Using high-speed imaging and water–succinic acid–butanol experiments, flow patterns and mass transfer rates were analyzed. Results show annular and dispersion flows dominate under tested conditions with structured inserts lowering the threshold for dispersion flow. Multi-hole injectors improved mass transfer by over 40% compared to T-mixers in plain tubes, while C-tape inserts achieved the highest volumetric mass transfer coefficient (2.43 s−1) due to increased interfacial area and droplet breakup from energy dissipation. This approach offers scalable solutions to enhance tube-in-tube millireactor performance for industrial applications. Full article
(This article belongs to the Special Issue Mass Transfer in Multiphase Reactors)
21 pages, 2483 KiB  
Review
Similarities in Mechanisms of Ovarian Cancer Metastasis and Brain Glioblastoma Multiforme Invasion Suggest Common Therapeutic Targets
by Gia A. Jackson and David Cory Adamson
Cells 2025, 14(3), 171; https://doi.org/10.3390/cells14030171 - 23 Jan 2025
Viewed by 451
Abstract
Epithelial-to-mesenchymal transition (EMT) is a critical process in malignant ovarian cancer metastasis. EMT involves the conversion of epithelial cells to mesenchymal cells, conferring enhanced migratory and invasive capabilities. Glioblastoma multiforme (GBM) is the most common malignant primary brain tumor and exhibits an aggressive [...] Read more.
Epithelial-to-mesenchymal transition (EMT) is a critical process in malignant ovarian cancer metastasis. EMT involves the conversion of epithelial cells to mesenchymal cells, conferring enhanced migratory and invasive capabilities. Glioblastoma multiforme (GBM) is the most common malignant primary brain tumor and exhibits an aggressive invasive phenotype that mimics some steps of EMT but does not undergo true metastasis, i.e., the invasion of other organ systems. This study conducts a comparative genomic analysis of EMT in ovarian cancer and invasion in GBM—two malignancies characterized by poor prognosis and limited therapies. Investigating the molecular biology in ovarian cancer and GBM demonstrates shared mechanisms of tumor progression, such as similar genetic and molecular pathways influencing cell plasticity, invasion, and resistance to therapy. The comparative analysis reveals commonalities and differences in the regulatory networks and gene expression profiles associated with EMT and invasion in these cancers. Key findings include the identification of core EMT regulators, such as TWIST1, SNAIL, and ZEB1, which are upregulated in both ovarian cancer and GBM, promoting mesenchymal phenotypes and metastasis. Additionally, the analysis uncovers EMT-related pathways, such as the PI3K/AKT and TGF-β signaling, which are critical in both cancers but exhibit distinct regulatory dynamics. Understanding the intricacies of EMT in ovarian cancer and invasion in GBM provides valuable insights into their aggressive behavior and identifies potential common therapeutic targets. The findings stress the importance of targeting EMT/invasion transitions to develop effective treatments to halt progression and improve patient outcomes in these malignancies. Full article
(This article belongs to the Special Issue Ovary and Brain—Series II)
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<p>Commonly mutated signaling pathways driving both EMT in HGSC and invasion in GBM. The figure highlights key signaling pathways in cancer, particularly during EMT and invasion. Key players include TGFβ, EGFR, and Wnt. These pathways are activated when ligands bind to receptors, triggering a series of phosphorylation events. For example, the RAS/RAF/MEK/ERK pathway drives cell growth and movement, but mutations can cause uncontrolled proliferation and EMT. The PI3K/AKT/mTOR pathway promotes survival and growth; mutations keep AKT signaling active, aiding EMT and invasion. The Wnt pathway stabilizes β-catenin and affects cell proliferation and migration, with mutations increasing EMT and metastasis. TGFβ is crucial for EMT, enhancing cell invasion. These pathways work together to regulate cancer progression, with mutations making EMT and metastasis more likely [<a href="#B6-cells-14-00171" class="html-bibr">6</a>].</p>
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<p>Key genomic, molecular, pathway, and cellular changes involved in EMT. This figure illustrates the key genomic, molecular, pathway, and cellular changes involved in EMT. It highlights genomic alterations, such as amplifications and deletions that drive EMT. The molecular section details the role of transcription factors like SNAIL, TWIST1, and ZEB1. As well as mesenchymal markers like N-cadherin and vimentin. The pathway segment depicts key signaling pathways, including TGF-β, Wnt, RAS-RAF-MEK-ERK, and PI3K-AKT, which regulate EMT. The cellular changes panel shows the morphological transformation from epithelial to mesenchymal phenotype, characterized by loss of cell–cell adhesion and increased migratory capacity [<a href="#B11-cells-14-00171" class="html-bibr">11</a>].</p>
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<p><b>The genomic landscape of ovarian cancer.</b> This figure presents the genomic landscape of ovarian cancer, highlighting key features of the tumor microenvironment, chromosomal rearrangements, mutations, and CNVs. The tumor microenvironment section shows interactions between cancer cells, stromal cells, immune cells, and the extracellular matrix. The chromosomal rearrangements panel displays common structural alterations like translocations and inversions. The mutations section details frequently mutated genes such as TP53, BRCA1, and BRCA2. Lastly, the CNVs panel highlights genomic amplification and deletion regions, affecting critical oncogenes and tumor suppressor genes [<a href="#B24-cells-14-00171" class="html-bibr">24</a>].</p>
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<p><b>Genomic landscape of GBM.</b> This figure illustrates the genomic landscape of GBM, highlighting the tumor microenvironment, chromosomal rearrangements, mutations, and CNVs. The tumor microenvironment section shows interactions among cancer cells, stromal cells, stem cells, immune cells, and the extracellular matrix with AJAP1, uPA, ADAMs, and various MMPs in facilitating tumor invasion and progression. The chromosomal rearrangements panel highlights gains and losses involving key genes such as EGFR and PTEN. The mutations section details altered genes, including TP53, EGFRvIII, and IDH1/2, in GBM pathogenesis. The CNVs panel shows genomic amplification and deletion regions, affecting crucial oncogenes and tumor suppressor genes [<a href="#B30-cells-14-00171" class="html-bibr">30</a>].</p>
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<p><b>Comparative analysis of invasion and EMT-related genomic alterations in ovarian cancer and GBM.</b> This Venn diagram presents a comparative analysis of genomic alterations related to invasion and EMT in ovarian cancer and GBM. The diagram shows unique genomic alterations associated with invasion and EMT specific to each GBM and ovarian cancer in the non-overlapping sections. The overlapping section illustrates shared genomic changes, highlighting common pathways and molecular mechanisms. This comparison underscores the similarities and differences in the genetic underpinnings of these aggressive behaviors in the two cancer types [<a href="#B46-cells-14-00171" class="html-bibr">46</a>].</p>
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16 pages, 5087 KiB  
Article
Cyanobiphenyl- and Cyanoterphenyl-Based Liquid Crystal Dimers (CBnCT): The Enantiotropic Twist-Bend Nematic Phase
by Yamato Shimoura and Yuki Arakawa
Crystals 2025, 15(2), 120; https://doi.org/10.3390/cryst15020120 - 23 Jan 2025
Viewed by 604
Abstract
We report the first homologous series of methylene-linked cyanobiphenyl- and cyanoterphenyl-based liquid crystal (LC) dimers (CBnCT). To induce the heliconical twist-bend nematic (NTB) phase through bent molecular shapes, the CBnCT homologs have an odd-numbered flexible alkylene spacer [...] Read more.
We report the first homologous series of methylene-linked cyanobiphenyl- and cyanoterphenyl-based liquid crystal (LC) dimers (CBnCT). To induce the heliconical twist-bend nematic (NTB) phase through bent molecular shapes, the CBnCT homologs have an odd-numbered flexible alkylene spacer (n) ranging from 1 to 17. Polarized optical microscopy and differential scanning calorimetry are used to identify phases and analyze the phase-transition behavior. Except for n = 1, all the CBnCT homologs exhibit the conventional nematic (N) and NTB phases. The CBnCT dimers with n = 3 and 5 show a monotropic NTB phase, while those with n = 7, 9, 11, 13, 15, and 17 demonstrate an enantiotropic NTB phase below the conventional N phase temperature. The NTB phases of the CBnCT dimers (n = 7, 9, and 11) remain stable down to room temperature and vitrify without crystallization. Compared with cyanobiphenyl-based LC dimer homologs (CBnCB), the CBnCT dimers show significantly broader N and NTB phase temperature ranges with higher isotropic and NTB–N phase-transition temperatures. The NTB phase temperature ranges of CBnCT (n = 7, 9, 11, and 13) are over 100 °C. Additionally, more CBnCT homologs exhibit the enantiotropic NTB phase than the CBnCB ones. These enhancements result from increased π-conjugation and asymmetric molecular structures. Furthermore, CB9CT exhibits higher birefringence than CB9CB owing to its longer π-conjugated terphenyl moiety. Full article
(This article belongs to the Special Issue Advances in Liquid Crystal Dimers and Oligomers)
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<p>Molecular structures of the methylene-linked CB<span class="html-italic">n</span>CB dimers and ether-linked CBO<span class="html-italic">n</span>OCB dimers.</p>
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<p>Molecular structures of the methylene-linked CB<span class="html-italic">n</span>CT homologs synthesized in this study.</p>
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<p>Phase-transition temperatures of CB<span class="html-italic">n</span>CT as a function of <span class="html-italic">n</span> upon (<b>a</b>) heating and (<b>b</b>) cooling.</p>
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<p>POM images of the N<sub>TB</sub> and N phases of CB9CT upon (<b>a</b>,<b>b</b>) heating and (<b>c</b>,<b>d</b>) cooling.</p>
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<p>POM images of the N<sub>TB</sub> and N phases of CB3CT upon cooling in (<b>a</b>,<b>b</b>) a non-treated glass cell and (<b>c</b>–<b>e</b>) a uniaxially rubbed planar alignment cell with a 7 µm thickness. The scale bars in (<b>a</b>,<b>c</b>) show 50 μm. The arrows with A, P, and R in (<b>e</b>) represent the directions of the analyzer, polarizer, and rubbing, respectively, for (<b>c</b>–<b>e</b>). The inset in (<b>c</b>) is a zoomed-in image of the double helical textures.</p>
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<p>DSC curves of CB<span class="html-italic">n</span>CT (<span class="html-italic">n</span> = 3, 5, 7, 9, 11, and 13) upon cooling. Liq. N<sub>2</sub> indicates peaks resulting from the liquid nitrogen added for cooling.</p>
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<p>(<b>a</b>) <span class="html-italic">T</span><sub>m</sub>, (<b>b</b>) Δ<span class="html-italic">T</span><sub>N</sub>, (<b>c</b>) <span class="html-italic">T</span><sub>IN</sub>, and (<b>d</b>) <span class="html-italic">T</span><sub>NNTB</sub> of CB<span class="html-italic">n</span>CT and CB<span class="html-italic">n</span>CB. The data for CB3CB (green triangles) in panels (<b>c</b>,<b>d</b>) represent the Iso–N<sub>TB</sub> phase transition temperature (i.e., <span class="html-italic">T</span><sub>INTB</sub>).</p>
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<p>Molecular structures of CB9CT and CB9CB, optimized using Gaussian 16 with the B3LYP/6-31G(d) basis set.</p>
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<p>Temperature dependence of Δ<span class="html-italic">n</span> of CB9CT and CB9CB as functions of (<b>a</b>) <span class="html-italic">T</span> and (<b>b</b>) Δ<span class="html-italic">T</span>.</p>
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<p>Synthesis routes to CB<span class="html-italic">n</span>CT.</p>
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15 pages, 6812 KiB  
Article
Rotor Position Estimation Algorithm for Surface-Mounted Permanent Magnet Synchronous Motor Based on Improved Super-Twisting Sliding Mode Observer
by Zhuoming Liang, Lanxian Cheng, Li Cheng and Canqing Li
Electronics 2025, 14(3), 436; https://doi.org/10.3390/electronics14030436 - 22 Jan 2025
Viewed by 449
Abstract
In response to the chattering issue inherent in sliding mode observers during rotor position estimation and to enhance the stability and robustness of sensorless control systems for surface-mounted permanent magnet synchronous motors (SPMSM), this study proposes a rotor position estimation algorithm for SPMSM [...] Read more.
In response to the chattering issue inherent in sliding mode observers during rotor position estimation and to enhance the stability and robustness of sensorless control systems for surface-mounted permanent magnet synchronous motors (SPMSM), this study proposes a rotor position estimation algorithm for SPMSM based on an improved super-twisting sliding mode observer (ISTSMO) and a second-order generalized integrator (SOGI) structure. Firstly, the super-twisting algorithm is introduced to design the observer, which effectively attenuates the sliding mode chattering by using continuous control signals. Secondly, SOGI is introduced in the filtering stage, which not only effectively addresses the time delay issues caused by traditional low-pass filters but also enables the observer to extract rotor position information by monitoring only the back electromotive force (back-EMF) signal of the α-phase, thereby simplifying the observer structure. Finally, the proposed scheme is experimentally compared with the traditional sliding mode observer on the YXMBD-TE1000 platform. The experimental results showed that during motor acceleration and deceleration tests, the average speed estimation error was reduced from 141 r/min to 40 r/min, and the maximum position estimation error was reduced from 0.74 rad to 0.29 rad. In load disturbance experiments, the speed variation decreased from 781 r/min to 451 r/min, and the steady-state speed fluctuation was significantly reduced. These results confirm that the proposed observer exhibits superior stability and robustness. Full article
(This article belongs to the Section Power Electronics)
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<p>The physical model of a three-phase PMSM.</p>
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<p>Block diagram of PLL.</p>
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<p>Block diagram of SOGI.</p>
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<p>Bode diagram of SOGI: (<b>a</b>) bode of <math display="inline"><semantics> <mrow> <msub> <mi>G</mi> <mi>v</mi> </msub> <mo stretchy="false">(</mo> <mi>s</mi> <mo stretchy="false">)</mo> </mrow> </semantics></math>; (<b>b</b>) bode of <math display="inline"><semantics> <mrow> <msub> <mi>G</mi> <mrow> <mi>i</mi> <mi>v</mi> </mrow> </msub> <mo stretchy="false">(</mo> <mi>s</mi> <mo stretchy="false">)</mo> </mrow> </semantics></math>.</p>
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<p>Control block diagram of ISTSMO.</p>
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<p>Block diagram of sensorless control of SPMSM.</p>
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<p>Experiment platform: (<b>a</b>) host computer and controller; (<b>b</b>) SPMSM and motor driver.</p>
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<p>Experimental results of the sudden change process of rotational speed: (<b>a</b>) SMO; (<b>b</b>) arcsin SMO; (<b>c</b>) ISTSMO.</p>
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<p>Back-EMF signal during the acceleration phase: (<b>a</b>) three observer; (<b>b</b>) zoomed-in waveform of ISTSMO.</p>
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<p>Experimental results of angle estimation at low speed: (<b>a</b>) SMO; (<b>b</b>) ISTSMO.</p>
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<p>Experimental results of load disturbance: (<b>a</b>) SMO; (<b>b</b>) arcsin SMO; (<b>c</b>) ISTSMO.</p>
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24 pages, 2980 KiB  
Article
Super-Twisting Algorithm Backstepping Adaptive Terminal Sliding-Mode Tracking Control of Quadrotor Drones Subjected to Faults and Disturbances
by Ye Zhang, Yihao Fu, Zhiguo Han and Jingyu Wang
Drones 2025, 9(2), 82; https://doi.org/10.3390/drones9020082 - 22 Jan 2025
Viewed by 380
Abstract
The rapid advancement of quadrotor systems has introduced significant challenges across multiple disciplines. Among these, fault tolerance and trajectory tracking in complex environments have long been recognized as critical challenges in quadrotor control research. To address issues such as rotor performance degradation and [...] Read more.
The rapid advancement of quadrotor systems has introduced significant challenges across multiple disciplines. Among these, fault tolerance and trajectory tracking in complex environments have long been recognized as critical challenges in quadrotor control research. To address issues such as rotor performance degradation and external disturbances, a novel position-attitude control system was developed, aimed to achieve precise position and attitude tracking. Initially, a dynamic model of the quadrotor was formulated, serving as the foundation for the controller design. Super-twisting algorithm terminal sliding-mode control (STATSMC) was then employed within the position loop to suppress chattering by the super-twisting algorithm. Subsequently, a new super-twisting algorithm beckstepping adaptive terminal sliding-mode control (STABATSMC) was proposed to mitigate the controller output and merge enable adherence to the desired Euler angles in case of failure. This approach enables the quadrotor to accurately follow position commands and achieve the desired attitude angles. The introduction of terminal sliding-mode control enhances convergence speed and tracking precision, while the super-twisting algorithm mitigates chattering and smoothens the control output. Finally, a series of simulation experiments were conducted within the Simulink environment to validate the proposed control system. The experimental results are compared with the state-of-art terminal sliding-mode control method, demonstrating the superior performance and effectiveness of the proposed method. Full article
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<p>The model of the quadrotor.</p>
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<p>Control loop of the quadrotor.</p>
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<p>Position control loop of the quadrotor.</p>
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<p>Attitude control loop of the quadrotor.</p>
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<p>The three–dimensional hexagon trajectory tracking of quadrotor without fault.</p>
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<p>Displacement motion in hexagonal trajectory tracking (no faults).</p>
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<p>Rotational motion in hexagonal trajectory tracking (no faults).</p>
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<p>The three–dimensional hexagon trajectory tracking of the quadrotor.</p>
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<p>Displacement motion in hexagonal trajectory tracking.</p>
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<p>Displacement motion error in hexagonal trajectory tracking.</p>
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<p>Rotational motion in hexagonal trajectory tracking.</p>
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<p>Rotational motion error in hexagonal trajectory tracking.</p>
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<p>The three–dimensional spiral trajectory tracking of the quadrotor.</p>
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<p>Displacement motion in spiral trajectory tracking.</p>
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<p>Displacement motion error in spiral trajectory tracking.</p>
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<p>Rotational motion in spiral trajectory tracking.</p>
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<p>Rotational motion error in spiral trajectory tracking.</p>
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<p>Displacement motion in rectangle trajectory tracking.</p>
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<p>Displacement motion error in rectangle trajectory tracking.</p>
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<p>Rotational motion in rectangle trajectory tracking.</p>
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<p>Rotational motion error in rectangle trajectory tracking.</p>
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27 pages, 6011 KiB  
Article
Sensitivity of Dynamic Stall Models to Dynamic Excitation on Large Flexible Wind Turbine Blades in Edgewise Vibrations
by Galih Bangga
Energies 2025, 18(3), 470; https://doi.org/10.3390/en18030470 - 21 Jan 2025
Viewed by 475
Abstract
Present studies are specifically aimed at investigating the sensitivity of different dynamic stall models when exposed to various excitation frequencies. The investigations are targeted at blade edgewise vibrations. This is carried out on a modified version of the IEA 15 MW reference wind [...] Read more.
Present studies are specifically aimed at investigating the sensitivity of different dynamic stall models when exposed to various excitation frequencies. The investigations are targeted at blade edgewise vibrations. This is carried out on a modified version of the IEA 15 MW reference wind turbine employing a wind turbine design tool, DNV Bladed. State-of-the-art dynamic stall models for wind turbine applications, such as the Øye model, Beddoes–Leishman (BL) model and the newly developed IAG model, are evaluated. The beginning of the research work starts by evaluating different dynamic stall models on rigid blade section forces against known airfoil datasets. Then, the blade flexibility is considered to enable systematic evaluations of the blade flexibility influences in comparison to the rigid blade cases. It is observed that the range of the angle of attack grows depending on the excitation frequency and the adopted dynamic stall model. The critical excitation frequency range and the effects of twist distribution are then identified from the studies, which can be useful as a rough guidance when designing wind turbine blades. Full article
(This article belongs to the Topic Advances in Wind Energy Technology)
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<p>Illustration of the wind turbine blade and placement of the target airfoil. Note: Not to scale with the actual blade used in the studies.</p>
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<p>Polar extrapolation to cover the entire operation range and interpolation based on relative thickness.</p>
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<p>Geometric and structural properties of the blade adopted in the present studies.</p>
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<p>Comparison of the aerodynamic force calculations of the S809 airfoil section against measurement data. The results represent dynamic characteristics at <span class="html-italic">k</span> = 0.079.</p>
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<p>Dynamic lift polar calculated at a constant reduced frequency of <span class="html-italic">k</span> = 0.079 at 10 m/s.</p>
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<p>Dynamic lift polar calculated at a constant reduced frequency of <span class="html-italic">k</span> = 0.079 at 30 m/s.</p>
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<p>Dynamic lift polar calculated at a constant reduced frequency of <span class="html-italic">k</span> = 0.079 at 50 m/s.</p>
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<p>Dynamic lift polar calculated at a constant resonance frequency of <span class="html-italic">f</span> = 0.687 Hz at 10 m/s (at <span class="html-italic">k</span> = 0.598).</p>
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<p>Dynamic lift polar calculated at a constant resonance frequency of <span class="html-italic">f</span> = 0.687 Hz at 30 m/s (at <span class="html-italic">k</span> = 0.199).</p>
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<p>Dynamic lift polar calculated at a constant resonance frequency of <span class="html-italic">f</span> = 0.687 Hz at 50 m/s (at <span class="html-italic">k</span> = 0.12).</p>
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<p>Range of angle of attack change due to aeroelastic response exposed to a constant reduced frequency of <span class="html-italic">k</span> = 0.079.</p>
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<p>Range of angle of attack change due to aeroelastic response exposed to various excitation frequencies. The non-dimensional excitation frequency (<math display="inline"><semantics> <mrow> <mi>f</mi> <mo>/</mo> <msub> <mi>f</mi> <mn>0</mn> </msub> </mrow> </semantics></math>) is: (<b>a</b>) 0.25, (<b>b</b>) 0.5, (<b>c</b>) 0.75, (<b>d</b>) 1.0, (<b>e</b>) 1.25 and (<b>f</b>) 1.5. The plots are made over the wind speed.</p>
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<p>Range of angle of attack change due to aeroelastic response exposed to various excitation frequencies. The non-dimensional excitation frequency (<math display="inline"><semantics> <mrow> <mi>f</mi> <mo>/</mo> <msub> <mi>f</mi> <mn>0</mn> </msub> </mrow> </semantics></math>) is: (<b>a</b>) 0.25, (<b>b</b>) 0.5, (<b>c</b>) 0.75, (<b>d</b>) 1.0, (<b>e</b>) 1.25 and (<b>f</b>) 1.5. The plots are made over the wind speed.</p>
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<p>Range of angle of attack change due to aeroelastic response exposed to various excitation frequencies. The wind speed is (<b>a</b>) 10 m/s, (<b>b</b>) 30 m/s and (<b>c</b>) 50 m/s. The plots are made over the non-dimensional excitation frequency (<math display="inline"><semantics> <mrow> <mi>f</mi> <mo>/</mo> <msub> <mi>f</mi> <mn>0</mn> </msub> </mrow> </semantics></math>).</p>
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<p>Range of angle of attack change due to aeroelastic response exposed to various excitation frequencies. The wind speed is (<b>a</b>) 10 m/s, (<b>b</b>) 30 m/s and (<b>c</b>) 50 m/s. The plots are made over the non-dimensional excitation frequency (<math display="inline"><semantics> <mrow> <mi>f</mi> <mo>/</mo> <msub> <mi>f</mi> <mn>0</mn> </msub> </mrow> </semantics></math>).</p>
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<p>Range of angle of attack change due to aeroelastic response exposed to the edgewise resonance frequency for the twisted blade (<span class="html-italic">f</span> = 0.687 Hz) at different wind speeds. (<b>a</b>) No dynamic stall model, (<b>b</b>) Øye model, (<b>c</b>) BL model and (<b>d</b>) IAG model. The percentage of reduction is presented, but the first data instance at 10 m/s is excluded in the plot due to a denominator close to zero when calculating the percentage.</p>
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<p>Illustration of <math display="inline"><semantics> <msub> <mi>x</mi> <mrow> <mi>c</mi> <mi>p</mi> </mrow> </msub> </semantics></math> movement when the angle of attack changes.</p>
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<p>Comparison of <math display="inline"><semantics> <msub> <mi>x</mi> <mrow> <mi>c</mi> <mi>p</mi> </mrow> </msub> </semantics></math> between the extrapolated and reference data of the NACA 63421 airfoil.</p>
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<p>Comparison of the aerodynamic polar between the extrapolated and reference data of the NACA 63421 airfoil.</p>
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<p>Comparison of the aerodynamic polar between the extrapolated and reference data of the NACA 63421 airfoil.</p>
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<p>Dynamic polar calculated at a constant reduced frequency of <span class="html-italic">k</span> = 0.079 at 10 m/s.</p>
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<p>Dynamic polar calculated at a constant reduced frequency of <span class="html-italic">k</span> = 0.079 at 30 m/s.</p>
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<p>Dynamic polar calculated at a constant reduced frequency of <span class="html-italic">k</span> = 0.079 at 50 m/s.</p>
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<p>Dynamic polar calculated at a constant resonance frequency of <span class="html-italic">f</span> = 0.687 Hz at 10 m/s (at <span class="html-italic">k</span> = 0.598).</p>
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<p>Dynamic polar calculated at a constant resonance frequency of <span class="html-italic">f</span> = 0.687 Hz at 30 m/s (at <span class="html-italic">k</span> = 0.199).</p>
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<p>Dynamic polar calculated at a constant resonance frequency of <span class="html-italic">f</span> = 0.687 Hz at 50 m/s (at <span class="html-italic">k</span> = 0.12).</p>
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22 pages, 4661 KiB  
Article
Unraveling Circular Conundrums with a Cheeky Twist: Proposal for a New Way of Measuring Circular Economy Efforts at the Product Level Within Procurement-to-Waste System Boundaries—A Case Study from the Airline Industry
by Christine Grimm
Sustainability 2025, 17(3), 807; https://doi.org/10.3390/su17030807 - 21 Jan 2025
Viewed by 490
Abstract
This study addresses the challenge of evaluating circularity within the procurement-to-waste system boundaries, using the example of single-use in-flight drinking cups provided by SWISS International Air Lines Ltd., the national airline of Switzerland. A comprehensive review of the academic literature, market-based tools, and [...] Read more.
This study addresses the challenge of evaluating circularity within the procurement-to-waste system boundaries, using the example of single-use in-flight drinking cups provided by SWISS International Air Lines Ltd., the national airline of Switzerland. A comprehensive review of the academic literature, market-based tools, and political regulations highlights the absence of adequate methodologies for assessing circularity within these specific system boundaries. Existing approaches, primarily designed at the product level, are often either excessively complex or focused solely on waste management. To address this gap, the research proposes an extension to the Circular Material Use rate (CMU), currently implemented at the European Union level. The traditional CMU rate does not account for circular inflow, thereby neglecting procurement decisions. In response, this study introduces an extended version of the CMU, expressed as CMU Extended = (Circular Inflow + Circular Outflow)/(2 × Total Material). This modification enables a more holistic evaluation of circularity by incorporating both inflows and outflows of materials in relation to total material use. Empirical testing demonstrated the applicability of this extended CMU in the context of SWISS, allowing for an efficient assessment of circularity for single-use in-flight drinking cups. From these initial results, we hypothesize that this ratio is expected to be broadly applicable beyond the airline industry, providing a valuable tool for businesses seeking to measure circularity within similar system boundaries. Full article
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<p>Example from the offline data-capturing file Circulytics.</p>
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<p>Example from the offline data-capturing file MCI Calculator.</p>
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<p>Example CTI tool inflow tab (excerpt).</p>
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<p>Example CTI tool.</p>
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<p>Example CTI report, inflow details (excerpt).</p>
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<p>Example of Circularity Calculator.</p>
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<p>CMU rate displayed and published by the Swiss government (Federal Statistik Office Switzerland, 2024) [<a href="#B29-sustainability-17-00807" class="html-bibr">29</a>].</p>
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<p>Testing of CMU at the company level.</p>
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<p>Summary of advantages and disadvantages of the CMU application at the company level.</p>
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<p>Selected key figures for the business year 2022 (Weleda Group and Weleda AG, 2023) [<a href="#B42-sustainability-17-00807" class="html-bibr">42</a>].</p>
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<p>Application of CMU Extended to the cup case.</p>
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<p>Advantages and disadvantages of CMU Extended.</p>
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12 pages, 2848 KiB  
Article
A 3D-Printed Enclosed Twist Dielectric Resonator Antenna with Circular Polarization
by Andrea Ávila-Saavedra, Marcos Diaz and Francisco Pizarro
Appl. Sci. 2025, 15(2), 992; https://doi.org/10.3390/app15020992 - 20 Jan 2025
Viewed by 576
Abstract
This article presents a circular polarized enclosed dielectric resonator antenna (DRA), operating at 5.8 GHz. The design consists of a twist DRA, which is enclosed in a box to give stability to the structure. The circular polarization of the antenna depends on the [...] Read more.
This article presents a circular polarized enclosed dielectric resonator antenna (DRA), operating at 5.8 GHz. The design consists of a twist DRA, which is enclosed in a box to give stability to the structure. The circular polarization of the antenna depends on the sense of twisting the top with respect to its base to achieve Left Hand Circular Polarization (LHCP) or Right Hand Circular Polarization (RHCP). The antenna was manufactured using 3D printing and low-loss dielectric filament. The measurement results show the two resonance frequencies and an axial ratio below 3 dB at the operational frequency, while exhibiting a bandwidth and gain compatible for unmanned aerial vehicle (UAV) applications. Full article
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<p>Schematic of the proposed antenna. (<b>a</b>) Dimension of the rectangles. (<b>b</b>) Rotation of the rectangles.</p>
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<p>Parametric study of the reflection coefficient <math display="inline"><semantics> <mrow> <mrow> <mo>|</mo> </mrow> <msub> <mi>S</mi> <mn>11</mn> </msub> <mrow> <mo>|</mo> </mrow> </mrow> </semantics></math> in dB as a function of the frequency and the inner box dimensions for the twist.</p>
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<p>Axial ratio in dB as a function of the frequency and the inner box dimensions for the twist.</p>
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<p>Reflection coefficient <math display="inline"><semantics> <mrow> <mrow> <mo>|</mo> </mrow> <msub> <mi>S</mi> <mn>11</mn> </msub> <mrow> <mo>|</mo> </mrow> </mrow> </semantics></math> as a function of the frequency for the twist implementations.</p>
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<p>Axial ratio as a function of the frequency for the twist implementations for different rotation angles.</p>
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<p>Schematic of the proposed antenna twist with box.</p>
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<p>Parametric study of the twist with box using different lateral dimensions. (<b>a</b>) Reflection coefficient <math display="inline"><semantics> <mrow> <mrow> <mo>|</mo> </mrow> <msub> <mi>S</mi> <mn>11</mn> </msub> <mrow> <mo>|</mo> </mrow> </mrow> </semantics></math> as a function of the frequency. (<b>b</b>) Axial ratio as a function of the frequency.</p>
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<p>Schematic of the proposed antenna. (<b>a</b>) Twist DRA with enclosure. (<b>b</b>) Feeding network detail.</p>
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<p>Simulation results for the LHCP implementation. (<b>a</b>) Reflection coefficient <math display="inline"><semantics> <mrow> <mrow> <mo>|</mo> </mrow> <msub> <mi>S</mi> <mn>11</mn> </msub> <mrow> <mo>|</mo> </mrow> </mrow> </semantics></math> as a function of the frequency. (<b>b</b>) Axial ratio as a function of the frequency.</p>
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<p>Simulated electric field distribution in the center of the DRA and a cut-side view. (<b>a</b>) Side view at 5.3 GHz. (<b>b</b>) Side view at 6.0 GHz. (<b>c</b>) Top view at 5.3 GHz. (<b>d</b>) Top view at 6.0 GHz.</p>
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<p>Simulated gain radiation pattern at 5.8 GHz for the LHCP implementation.</p>
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<p>The 3D-printed antenna. (<b>a</b>) CURA implementation with the infill percentages. (<b>b</b>) Implemented antenna. From left to right: twist alone, twist in the enclosure with the exposed half, and full antenna.</p>
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<p>Measurement results for the LHCP implementation. (<b>a</b>) Reflection coefficient <math display="inline"><semantics> <mrow> <mrow> <mo>|</mo> </mrow> <msub> <mi>S</mi> <mn>11</mn> </msub> <mrow> <mo>|</mo> </mrow> </mrow> </semantics></math> as a function of the frequency. (<b>b</b>) Axial ratio as a function of the frequency.</p>
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<p>Results of the measured radiation pattern for the implementation of the LHCP implementation.</p>
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<p>Top view of the antenna misplacement sensitivity study.</p>
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<p>Simulated results for the sensitivity analysis. (<b>a</b>) Displacement error. (<b>b</b>) Permittivity uncertainty.</p>
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21 pages, 1835 KiB  
Article
Optimized Super-Twisting Sliding Mode Control with Parameter Estimation for Autonomous Vehicle Longitudinal Motion on Downhill Road
by Kbrom Lbsu Gdey and Woo Young Choi
Appl. Sci. 2025, 15(2), 981; https://doi.org/10.3390/app15020981 - 20 Jan 2025
Viewed by 478
Abstract
Longitudinal motion control is a critical aspect of autonomous vehicle area, requiring a well-designed controller to ensure optimal system performance, safety, and comfort under varying driving conditions. Previous control methods often face challenges such as chattering effects, and model uncertainties. To address such [...] Read more.
Longitudinal motion control is a critical aspect of autonomous vehicle area, requiring a well-designed controller to ensure optimal system performance, safety, and comfort under varying driving conditions. Previous control methods often face challenges such as chattering effects, and model uncertainties. To address such challenges, in this paper, we propose the application of an optimized super-twisting sliding mode control (OST-SMC) for the longitudinal motion control of autonomous vehicles. The motivation is to enhance the robustness and efficiency of the control system while minimizing the chattering problem. The proposed system’s mathematical modeling and control design are presented in detail with stability analyzed using Lyapunov theory. To enhance the controller’s performance, uncertain parameters are optimized using the gradient descent method, a linear regression-based technique. The OST-SMC algorithm shows enhanced robustness against disturbances and parameter uncertainties compared to conventional sliding mode controllers. Simulations in MATLAB/Simulink and CarMaker validate the proposed method, demonstrating strong performance even on downhill roads. The OST-SMC reduces chattering more effectively than traditional SMCs, achieving smooth tracking and consistent robustness under varying road conditions. Full article
(This article belongs to the Section Transportation and Future Mobility)
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<p>Schematic diagram of the proposed system.</p>
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<p>Conceptual visualization of experimental scenario.</p>
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<p>Visualization of the experimental test with CarMaker: At about 6 (s), the car is driving on a straight road. At about 38.5 (s), the road starts to go downhill. At about 44.5 (s), the downhill ends.</p>
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<p>Process of the proposed method: integration of OST-SMC controller with CarMaker and MATLAB/Simulink framework.</p>
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<p>Tracking performance of the proposed system: vehicle velocity, acceleration, and control input under parameter variations.</p>
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<p>Vehicle pitch and pitch rate comparison under parameter variations and optimal conditions: how the vehicle tilts up and down during the whole scenario can be seen.</p>
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<p>Proposed method performance analysis at different road slopes.</p>
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<p>Proposed method performance analysis at different friction coefficients.</p>
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<p>Proposed method performance analysis at different desired speeds.</p>
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<p>Tracking performance comparison of OST-SMC vs SMC.</p>
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