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17 pages, 11707 KiB  
Article
Interfacial Characteristics and Mechanical Properties of TiAl4822/Ti6Al4V Metal–Intermetallic Laminate Composite Prepared Through Vacuum Hot Pressing
by Jianwen Qin, Shouyin Zhang, Zhijian Ma and Baiping Lu
Materials 2025, 18(4), 898; https://doi.org/10.3390/ma18040898 - 19 Feb 2025
Abstract
In this work, the TiAl4822/Ti6Al4V metal–intermetallic laminate (MIL) composite was fabricated using vacuum hot pressing (VHP). The interfacial morphologies and mechanical properties of the composites were investigated. No discernible defect was observed in the well-bonded interface region. This interface region comprised two distinct [...] Read more.
In this work, the TiAl4822/Ti6Al4V metal–intermetallic laminate (MIL) composite was fabricated using vacuum hot pressing (VHP). The interfacial morphologies and mechanical properties of the composites were investigated. No discernible defect was observed in the well-bonded interface region. This interface region comprised two distinct areas: the Ti2Al (6 μm) region near the TiAl layer and the Ti3Al (4 μm) region near the Ti6Al4V layer. Electron backscatter diffraction analysis revealed that dynamic recrystallization (DRX) took place at the interface during the hot pressing process. The ductile brittle nature of Ti6Al4V and TiAl4822 layers and the formation of fine grains within the interface are conducive to enhancing toughness and tensile strength. Room temperature tensile testing exhibited that the tensile strength of TiAl4822/Ti6Al4V MIL composite was 636.9 MPa, approximately 225 MPa higher than single TiAl4822 alloy. The Ti6Al4V layer, as well as the formation of fine grain interface, effectively inhibited further propagation of the main crack through crack passivation, crack deflection, and load transformation. The bending strength of the TiAl4822/Ti6Al4V MIL composite was 1114.1 MPa. The fracture toughness of the TiAl4822/Ti6Al4V MIL composite reached 33.15 MPam1/2, which increased by 78.2% compared with single TiAl4822 alloy. Full article
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Figure 1
<p>OM of the starting alloy: (<b>a</b>) TiAl4822; (<b>b</b>) Ti6Al4V.</p>
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<p>Schematic of the preform.</p>
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<p>Schematic of VHP processing parameters.</p>
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<p>Schematic of the specimen: (<b>a</b>) tensile test; (<b>b</b>) three-point bending test.</p>
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<p>Fracture toughness testing specimen.</p>
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<p>MIL micromorphology: (<b>a</b>) macro-organization interface and (<b>b</b>) MIL interface area image; (<b>c</b>,<b>e</b>) detailed drawing of the interface area; (<b>d</b>) microstructure of the TiAl4822.</p>
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<p>Phase composition of the TiAl4822/Ti6Al4V MIL composites: (<b>a</b>) scan line trace, (<b>b</b>) EDS line scan, (<b>c</b>) EDS spot scan, and (<b>d</b>) XRD spectrum.</p>
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<p>EBSD analysis of the TiAl4822/Ti6Al4V MIL cross-section: (<b>a</b>) grain morphology, (<b>b</b>) phase distribution and grain boundary diagram, (<b>c</b>) grain angular orientation difference, (<b>d</b>) KAM map of MIL, (<b>e</b>) image quality mapping of GOS values and their volume fractions, (<b>f</b>) critical GOS value.</p>
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<p>Texture distribution of TiAl4822/Ti6Al4V MIL composite: (<b>a</b>) EBSD inverse pole figure of region A: Ti6Al4V martrix, regionn B: TiAl4822 martrix and region C: interface layer, (<b>b</b>) EBSD pole figure of α/α2, (<b>c</b>) EBSD pole figure of TiAl.</p>
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<p>Microstructural TEM images of the TiAl4822/Ti6Al4V MIL: (<b>a</b>) Ti<sub>3</sub>Al phase, (<b>b</b>) Ti<sub>2</sub>Al phase, (<b>c</b>) high-resolution image and diffraction spots of Ti<sub>3</sub>Al (<b>d</b>) high-resolution image and diffraction spots of Ti<sub>2</sub>Al.</p>
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<p>Tensile stress–strain curves of: (<b>a</b>) TiAl4822/Ti6Al4V MIL composite and TiAl4822 at room temperature; (<b>b</b>) TiAl4822/Ti6Al4V MIL composite at 650 °C.</p>
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<p>Fractographies of the TiAl4822/Ti6Al4VMIL composite: (<b>a</b>) fracture surface of the cross-section, (<b>b</b>) morphology of the Ti6Al4V ductile fracture, (<b>c</b>,<b>e</b>) fracture morphology of the interface region, and (<b>d</b>,<b>f</b>) brittle cleavage and quasi cleavage fracture morphology of TiAl4822.</p>
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<p>Crack propagation characteristics of TiAl4822/Ti6Al4V MIL fracture toughness test specimen in the direction parallel to the hot pressing direction: (<b>a</b>) overall specimen morphology; (<b>b</b>–<b>d</b>) Local crack growth path.</p>
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<p>Load deformation curve and crack propagation characteristics of the TiAl4822/Ti6Al4V MIL composite: (<b>a</b>) load deformation curve, (<b>b</b>) crack growth morphology, (<b>c</b>–<b>f</b>) enlarged view of the local area in <a href="#materials-18-00898-f012" class="html-fig">Figure 12</a>b.</p>
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19 pages, 3698 KiB  
Article
Synthesis and Characterization of Memantine-Loaded Niosomes for Enhanced Alzheimer’s Disease Targeting
by Hasan Turkez, Sena Oner, Ozge Caglar Yıldırım, Mehmet Enes Arslan, Marilisa Pia Dimmito, Çigdem Yuce Kahraman, Lisa Marinelli, Erdal Sonmez, Özlem Kiki, Abdulgani Tatar, Ivana Cacciatore, Antonio Di Stefano and Adil Mardinoglu
Pharmaceutics 2025, 17(2), 267; https://doi.org/10.3390/pharmaceutics17020267 - 17 Feb 2025
Abstract
Background/Objectives: Over the past 25 years, numerous biological molecules, like recombinant lysosomal enzymes, neurotrophins, receptors, and therapeutic antibodies, have been tested in clinical trials for neurological diseases. However, achieving significant success in clinical applications has remained elusive. A primary challenge has been the [...] Read more.
Background/Objectives: Over the past 25 years, numerous biological molecules, like recombinant lysosomal enzymes, neurotrophins, receptors, and therapeutic antibodies, have been tested in clinical trials for neurological diseases. However, achieving significant success in clinical applications has remained elusive. A primary challenge has been the inability of these molecules to traverse the blood–brain barrier (BBB). Recognizing this hurdle, our study aimed to utilize niosomes as delivery vehicles, leveraging the “molecular Trojan horse” technology, to enhance the transport of molecules across the BBB. Methods: Previously synthesized memantine derivatives (MP1–4) were encapsulated into niosomes for improved BBB permeability, hypothesizing that this approach could minimize peripheral drug toxicity while ensuring targeted brain delivery. Using the human neuroblastoma (SH-SY5Y) cell line differentiated into neuron-like structures with retinoic acid and then exposed to amyloid beta 1–42 peptide, we established an in vitro Alzheimer’s disease (AD) model. In this model, the potential usability of MP1–4 was assessed through viability tests (MTT) and toxicological response analysis. The niosomes’ particle size and morphological structures were characterized using scanning electron microscopy (SEM), with their loading and release capacities determined via UV spectroscopy. Crucially, the ability of the niosomes to cross the BBB and their potential anti-Alzheimer efficacy were analyzed in an in vitro transwell system with endothelial cells. Results: The niosomal formulations demonstrated effective drug encapsulation (encapsulation efficiency: 85.3% ± 2.7%), controlled release (72 h release: 38.5% ± 1.2%), and stable morphology (PDI: 0.22 ± 0.03, zeta potential: −31.4 ± 1.5 mV). Among the derivatives, MP1, MP2, and MP4 exhibited significant neuroprotective effects, enhancing cell viability by approximately 40% (p < 0.05) in the presence of Aβ1-42 at a concentration of 47 µg/mL. The niosomal delivery system improved BBB permeability by 2.5-fold compared to free drug derivatives, as confirmed using an in vitro bEnd.3 cell model. Conclusions: Memantine-loaded niosomes provide a promising platform for overcoming BBB limitations and enhancing the therapeutic efficacy of Alzheimer’s disease treatments. This study highlights the potential of nanotechnology-based delivery systems in developing targeted therapies for neurodegenerative diseases. Further in vivo studies are warranted to validate these findings and explore clinical applications. Full article
(This article belongs to the Section Drug Delivery and Controlled Release)
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Graphical abstract

Graphical abstract
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<p>Synthetic strategies to obtain memantine derivatives (<b>MP1–4</b>). Reagent and conditions: (<b>a</b>) valproic acid, TEA, and ethylchloroformate in dry THF/DMF, 3 h, rt; (<b>b</b>) phenylbutyric acid, TEA, and ethylchloroformate in dry THF/DMF, 3 h, rt; (<b>c</b>) butyric acid, TEA, and ethylchloroformate in dry THF/DMF, 3 h, rt; (<b>d</b>) caffeic acid, TEA, HOBt, and DCC in dry DMF, 15 h, rt.</p>
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<p>Schematic diagram summarizing in vitro BBB permeability analyses.</p>
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<p>Determination of cytotoxic properties of synthesized memantine derivatives (<b>MP1–4</b>) in cultured human fibroblast cells (HDFa) using MTT cell viability assay. Group analyses were performed with the one-way ANOVA procedure, and comparisons were made using Dunnett’s test (against the control).</p>
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<p>Analysis of the genotoxic properties of synthesized memantine derivatives (<b>MP1–4</b>) in cultured human fibroblast cells (HDFa) using Hoechst 33258 fluorescent staining of cell nuclei. (<b>A</b>) <b>MP1</b>, (<b>B</b>) <b>MP2</b>, (<b>C</b>), <b>MP3</b> and (<b>D</b>) <b>MP4</b>.</p>
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<p>Scanning electron microscope (SEM) image showing the surface topography and morphology of niosomes prepared by the thin film method.</p>
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<p>Drug release profile showing the release of <b>MP1–4</b> drugs from niosomal carriers prepared by the thin film hydration technique at certain time intervals for 24 h.</p>
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<p>Morphological cell structures of differentiated SH-SY5Y cells. (<b>A</b>) 20× resolution of undifferentiated cell cultures; (<b>B</b>) 20× resolution of differentiated cell cultures (application of 10 µM all-trans RA to cell culture for 11 days).</p>
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<p>Determination of the cytotoxic properties of <b>MP1–4</b> in the differentiated SH-SY5Y cell cultures resembling mature neurons using the MTT cell viability test. Group analyses were performed with the one-way ANOVA procedure, and comparisons were made using Dunnett’s test (against the control).</p>
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<p>Determination of the neuroprotective properties of <b>MP1–4</b> in an experimental AD model. Group analyses were performed with the one-way ANOVA procedure, and comparisons were made using Dunnett’s test (against the amyloid beta application). The level of significance was set at 5% (<span class="html-italic">p</span> &lt; 0.05). Significant differences compared to the control group are indicated by an asterisk (*).</p>
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<p>Neuroprotective properties of the <b>MP1–4</b> (47 µg/mL) on the experimental AD model for 24 h of application. Group analyses were performed with the one-way ANOVA procedure, and comparisons were made using Dunnett’s test (against the amyloid beta application). The level of significance was set at 5% (<span class="html-italic">p</span> &lt; 0.05). Significant differences compared to the control group are indicated by an asterisk (*).</p>
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<p>The neuroprotective effect of MP1–4 and the drug/carrier system on cell viability in the Alzheimer’s disease model using an in vitro BBB permeability test. Group analyses were performed with the one-way ANOVA procedure, and comparisons were made using Dunnett’s test (against the amyloid beta application). The level of significance was set at 5% (<span class="html-italic">p</span> &lt; 0.05). Significant differences compared to the control group are indicated by an asterisk (*).</p>
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9 pages, 2186 KiB  
Communication
Flexible Hybrid Integration Hall Angle Sensor Compatible with the CMOS Process
by Ye Luo, Youtong Fang, Yang Lv, Huaxiong Zheng and Ke Guan
Sensors 2025, 25(3), 927; https://doi.org/10.3390/s25030927 - 4 Feb 2025
Abstract
Silicon-based Hall application-specific integrated circuit (ASIC) chips have become very successful, making them ideal for flexible electronic and sensor devices. In this study, we designed, simulated, and tested flexible hybrid integration angle sensors that can be made using complementary metal-oxide-semiconductor (CMOS) technology. These [...] Read more.
Silicon-based Hall application-specific integrated circuit (ASIC) chips have become very successful, making them ideal for flexible electronic and sensor devices. In this study, we designed, simulated, and tested flexible hybrid integration angle sensors that can be made using complementary metal-oxide-semiconductor (CMOS) technology. These sensors are manufactured on a 100 µm-thick flexible polyimide (PI) membrane, which is suitable for large-scale production and has strong potential for industrial use. The Hall sensors have a sensitivity of 0.205 V/mT. Importantly, their sensitivity remains stable even after being bent to a minimum radius of 10 mm and after undergoing 100 bending cycles. The experiment shows that these flexible hybrid integration devices are promising as angle sensors. Full article
(This article belongs to the Section Physical Sensors)
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<p>The preparation and mechanical analysis of the flexible hybrid integration Hall sensor. (<b>a</b>) The process for creating the flexible hybrid integration Hall sensor. (<b>b</b>) An experimental image of the finished sensor. (<b>c</b>) An image of the sensing unit. (<b>d</b>) A diagram of the electrical connections for the sensing unit. (<b>e</b>) The mechanical distribution in the xx direction across different thicknesses (50 µm, 100 µm, 200 µm, and 300 µm) at a bending radius of 10 mm. (<b>f</b>) The strain distribution in the silicon wafer along the dashed red line.</p>
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<p>Analysis of the silicon-based Hall plate inside the chip die. (<b>a</b>) Schematic diagram of the 3D structure of the Hall plate; Nepi (N type silicon epitaxial layer); Psub (P type silicon substrate); PW (P type well). (<b>b</b>) Simulation of the electric potential distribution of the Hall plate under a magnetic field in the <span class="html-italic">z</span>-axis direction. (<b>c</b>) Simulation of Hall voltage values under different strengths of the <span class="html-italic">z</span>-axis magnetic field. (<b>d</b>) Schematic diagram of the electrical relationship of the rotating current method used to eliminate offset voltage. (<b>e</b>) SEM image of the Hall plate after removing the upper metal structures. (<b>f</b>) Electrical connection diagram of the rotating current method; LDO (Low dropout regulator).</p>
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<p>Circuit analysis of the chip die. (<b>a</b>) Schematic block diagram of the chip die circuit. (<b>b</b>,<b>c</b>) Simulation of the first-order chopper amplifier waveform and the demodulated filtered waveform. (<b>d</b>) Simulation of output values under different z-direction magnetic field strengths. (<b>e</b>) Layout of the chip die. (<b>f</b>) Image of the chip.</p>
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<p>Experimental testing of the flexible hybrid integration Hall sensor. (<b>a</b>) Schematic diagram of the experimental setup for measuring magnetic field sensitivity. (<b>b</b>) Experimental testing image, which shows different bending radii of the flexible device achieved through planar buckling. (<b>c</b>) Measurement results of magnetic sensing sensitivity before bending, during bending (with a 10 mm bending radius), and after 100 cycles of bending. (<b>d</b>) Schematic diagram of the experimental setup for measuring magnetic field response. (<b>e</b>) Voltage test results with varying distances between the magnet and the sensitive element, the right <span class="html-italic">y</span>-axis represents the distance between the magnet surface and the chip die. (<b>f</b>) Voltage test results when the magnet moves laterally across the sensitive element.</p>
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<p>Flexible hybrid integration angle sensor. (<b>a</b>) Photo of the flexible hybrid integration angle sensor. (<b>b</b>) Schematic diagram of the flexible hybrid integration angle sensor device, which is directly attached to the outer edge of the measurement location, with the magnet mounted on an internal rotating shaft to collect voltage data through the circuit. (<b>c</b>) Photo of the device. (<b>d</b>) Voltage output at different rotation positions of the magnet.</p>
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20 pages, 7285 KiB  
Article
Study on Bending Performance of High-Ductility Composite Slab Floor with Composite Ribs
by Yuchen Jiang, Libo Liu, Xiaolei Wang, Run Liu and Haibo Yang
Materials 2025, 18(3), 658; https://doi.org/10.3390/ma18030658 - 2 Feb 2025
Abstract
In order to solve the problems of high production cost and complex control of the inverted arch of an unsupported prestressed concrete composite slab, a flange truss high-ductility concrete composite slab floor is proposed to change the structure and pouring material to meet [...] Read more.
In order to solve the problems of high production cost and complex control of the inverted arch of an unsupported prestressed concrete composite slab, a flange truss high-ductility concrete composite slab floor is proposed to change the structure and pouring material to meet the requirements of no support during construction. The crack distribution and bending performance of the flange truss high-ductile concrete composite slab floor (CRHDCS) under different structures are clarified through the test and numerical analysis of four different rib plate structure floors. According to the analysis results, the calculation formulas of the cracking moment and short-term stiffness before cracking are modified, and the equivalent short-term stiffness formula of a single web member of the “V” truss to this kind of bottom plate is established. The results show that, unlike the short-term stiffness-change law of typical concrete composite slabs after cracking, the short-term stiffness of the designed bottom plate in this paper includes a short-term increase stage. The numerical simulation results are the same as the experimental ones; the maximum error is 10%. The maximum errors between the modified cracking moment and the short-term stiffness calculation formula are 6% and 8%, respectively. The influence rates of removing flange plate, truss-inverted binding, and adding rib plate on the cracking bending moment of foundation structure are −81.5%, 11.0%, and 22.2% respectively, and the influence rates on short-term stiffness are −87.6%, −1.5%, and 37.5% respectively. Full article
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<p>Composite slab foundation structure diagram.</p>
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<p>Concrete rib composite slab floor.</p>
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<p>Foundation structure of flange truss composite ribbed slab high-ductility concrete composite slab floor.</p>
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<p>Photograph of steel fiber.</p>
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<p>Specimen structure diagram. * The unit of size in the figure is mm.</p>
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<p>Web structure and size diagram.</p>
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<p>Schematic diagram of the test site.</p>
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<p>The actual drawing of the test.</p>
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<p>Layout diagram of steel strain gauge.</p>
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<p>Layout of strain gauges on the lower surface of concrete.</p>
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<p>Layout of measuring points on the side surface of concrete.</p>
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<p>Fracture diagram of the test bottom plate.</p>
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<p>Load-mid-span deflection curve diagram. (<b>a</b>) The curves of DB-1 and DB-2; (<b>b</b>) The curves of DB-1 and DB-3; (<b>c</b>) The curves of DB-1 and DB-4.</p>
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<p>Load-strain curve at mid-span. (<b>a</b>) Load-web mid-span steel strain curve; (<b>b</b>) External load-web mid-span concrete strain curve.</p>
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<p>Numerical model diagram (DB-1).</p>
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<p>Plastic strain program of concrete (DB-1).</p>
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<p>Steel strain cloud diagram (DB-1).</p>
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<p>Flange rib high-ductile concrete composite slab floor stiffness calculation diagram.</p>
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<p>Truss equivalent calculation diagram. (<b>a</b>) Simplified truss model; (<b>b</b>) Equivalent beam model diagram.</p>
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29 pages, 5792 KiB  
Article
Probabilistic Modelling of Fatigue Behaviour of 51CrV4 Steel for Railway Parabolic Leaf Springs
by Vítor M. G. Gomes, Felipe K. Fiorentin, Rita Dantas, Filipe G. A. Silva, José A. F. O. Correia and Abílio M. P. de Jesus
Metals 2025, 15(2), 152; https://doi.org/10.3390/met15020152 - 1 Feb 2025
Abstract
The longevity of railway vehicles is an important factor in their mechanical and structural design. Fatigue is a major issue that affects the durability of railway components, and therefore, knowledge of the fatigue resistance characteristics of critical components, such as the leaf springs, [...] Read more.
The longevity of railway vehicles is an important factor in their mechanical and structural design. Fatigue is a major issue that affects the durability of railway components, and therefore, knowledge of the fatigue resistance characteristics of critical components, such as the leaf springs, must be extensively investigated. This research covers the fatigue resistance of chromium–vanadium alloy steel, usually designated as 51CrV4, from the high-cycle regime (HCF) (103104) up to very high-cycle fatigue (VHCF) (109) under the bending loading conditions typical of leaf springs and under uniaxial tension/compression loading, respectively, for a stress ratio, Rσ, of −1. Different test frequencies were considered (23, 150, and 20,000 Hz) despite the material not exhibiting a relatively significant frequency effect. In order to create a new fatigue prediction model, two prediction models, the Basquin SN linear regression model and the Castillo–Fernandez–Cantelli (CFC) model, were evaluated. According to the analysis carried out, the CFC model provided a better prediction of the fatigue failures than the SN model, especially when outlier failure data were considered. The investigation also examined the failure modes, observing multiple cracks for higher loads and single cracks that initiated on the surface or from internal inclusions at lower loading. The present investigation aims to provide a fatigue resistance prediction model encompassing the HCF and VHCF regions for the fatigue design of railway wagon leaf springs, or even for other components made of 51CrV4 with a tempered martensitic microstructure. Full article
(This article belongs to the Special Issue Fracture Mechanics of Metals (2nd Edition))
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<p>Two-axle wagon suspension with a parabolic leaf spring.</p>
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<p>Typical microstructure of the chromium–vanadium alloyed steel found for all tested specimens using optical microscopy (picture adapted from [<a href="#B30-metals-15-00152" class="html-bibr">30</a>]).</p>
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<p>Rotating bending fatigue testing machine (simple bending).</p>
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<p>Geometry of the smooth fatigue specimen for rotating bending loading. <b>Left:</b> Sample of the actual specimen showing the details (Dt.A) of the finish in the analysis zone: Dt. A1—polished; Dt. A2—unpolished. <b>Right:</b> Rendered image of the CAD model showing the dimensions for the definition of the specimen’s geometry.</p>
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<p>Representation of the Shimadzu machine, its structure, and the testing specimen with the cooling system.</p>
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<p>Geometry of the fatigue specimen for ultrasonic uniaxial tension/compression testing in the Shimadzu machine. <b>Left:</b> Sample of the actual specimen showing the details (Dt. A) of the finish in the analysis zone: Dt. A—polished. <b>Right:</b> Rendered image of the CAD model showing the dimensions for the definition of the specimen’s geometry.</p>
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<p>The Rumul machine and the testing specimen installed.</p>
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<p>Geometry of a smooth fatigue specimen for uniaxial tension/compression testing in the Rumul machine. <b>Left:</b> Rendered image of the CAD model showing the dimensions for the definition of the specimen’s geometry. <b>Right:</b> Sample of an actual specimen.</p>
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<p>SN curve for smooth fatigue specimens under rotating bending loading conditions with turned surface finishing (<span class="html-italic">R</span>—stress ratio; <math display="inline"><semantics> <msub> <mi>P</mi> <mi>f</mi> </msub> </semantics></math>—probability of failure; RB—rotating bending; Avg.—Average curve <math display="inline"><semantics> <msub> <mi>σ</mi> <mi>a</mi> </msub> </semantics></math>—Stress amplitude; <math display="inline"><semantics> <msub> <mi>N</mi> <mi>f</mi> </msub> </semantics></math>—Number of cycles to failure; Dist—Distribution).</p>
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<p>SN curve for smooth fatigue specimens under rotating bending loading conditions with polished plus turned surface finishing (<span class="html-italic">R</span>—stress ratio; <math display="inline"><semantics> <msub> <mi>P</mi> <mi>f</mi> </msub> </semantics></math>—probability of failure; RB—rotating bending; Poli.—polished specimen; Turn—turned specimens; Avg.—average curve; <math display="inline"><semantics> <msub> <mi>σ</mi> <mi>a</mi> </msub> </semantics></math>—stress amplitude; <math display="inline"><semantics> <msub> <mi>N</mi> <mi>f</mi> </msub> </semantics></math>—number of cycles to failure; Dist—Distribution).</p>
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<p>SN curve for smooth fatigue specimens with polished surface finishing under subsonic tension/compression fatigue loading conditions (<span class="html-italic">R</span>—stress ratio; <math display="inline"><semantics> <msub> <mi>P</mi> <mi>f</mi> </msub> </semantics></math>—probability of failure; AT(150Hz)—specimen under axial tension at 150 Hz; Avg.—average curve; <math display="inline"><semantics> <msub> <mi>σ</mi> <mi>a</mi> </msub> </semantics></math>—stress amplitude; <math display="inline"><semantics> <msub> <mi>N</mi> <mi>f</mi> </msub> </semantics></math>—number of cycles to failure; Dist—Distribution).</p>
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<p>SN curve for smooth fatigue specimens with polished surface finishing under ultrasonic tension/compression fatigue loading conditions (<span class="html-italic">R</span>—stress ratio; <math display="inline"><semantics> <msub> <mi>P</mi> <mi>f</mi> </msub> </semantics></math>—probability of failure; AT(20kHz)—specimen under axial tension at 20 kHz; Avg.—average curve; <math display="inline"><semantics> <msub> <mi>σ</mi> <mi>a</mi> </msub> </semantics></math>—stress amplitude; <math display="inline"><semantics> <msub> <mi>N</mi> <mi>f</mi> </msub> </semantics></math>—number of cycles to failure; Dist—Distribution).</p>
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<p>Comparison of fatigue data obtained for the fatigue life of spring steel under uniaxial tension/compression tests (<span class="html-italic">R</span>—stress ratio; <math display="inline"><semantics> <msub> <mi>P</mi> <mi>f</mi> </msub> </semantics></math>—probability of failure; AT(150Hz)—specimen under axial tension at 150 Hz; AT(20kHz)—specimen under axial tension at 20 kHz; Avg.—average curve; <math display="inline"><semantics> <msub> <mi>σ</mi> <mi>a</mi> </msub> </semantics></math>—stress amplitude; <math display="inline"><semantics> <msub> <mi>N</mi> <mi>f</mi> </msub> </semantics></math>—number of cycles to failure; Dist—Distribution).</p>
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<p>Comparison of fatigue data obtained from fatigue tests of spring steel in uniaxial tensile tests at 150 Hz and 20 kHz considering corrected nominal stresses for the frequency effect (<span class="html-italic">R</span>—stress ratio; <math display="inline"><semantics> <msub> <mi>P</mi> <mi>f</mi> </msub> </semantics></math>—probability of failure; AT(150Hz)—specimen under axial tension at 150 Hz; AT(20kHz)—specimen under axial tension at 20 kHz; Avg.—average curve; <math display="inline"><semantics> <msub> <mi>σ</mi> <mi>a</mi> </msub> </semantics></math>—stress amplitude; <math display="inline"><semantics> <msub> <mi>N</mi> <mi>f</mi> </msub> </semantics></math>—number of cycles to failure; Dist—Distribution).</p>
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<p>PSN field for smooth specimens under fatigue rotating bending loading (<span class="html-italic">R</span>—stress ratio; <math display="inline"><semantics> <msub> <mi>P</mi> <mi>f</mi> </msub> </semantics></math>—probability of failure; Est—estimation for run-out data; RB(25Hz)—smooth specimens under rotating bending at 25 Hz; Avg.—average curve; <math display="inline"><semantics> <msub> <mi>σ</mi> <mi>a</mi> </msub> </semantics></math>—stress amplitude; <math display="inline"><semantics> <msub> <mi>N</mi> <mi>f</mi> </msub> </semantics></math>—number of cycles to failure; Dist—Distribution; 3p—3 parameters).</p>
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<p>PSN field curve for smooth specimens under subsonic and ultrasonic fatigue tensile loading (<span class="html-italic">R</span>—stress ratio; <math display="inline"><semantics> <msub> <mi>P</mi> <mi>f</mi> </msub> </semantics></math>—probability of failure; Est—estimation for run-out data; AT(150Hz)—smooth specimens under axial tension at 150 Hz; AT(20kHz)—smooth specimens under a axial load at 20 kHz; Avg.—average curve; <math display="inline"><semantics> <msub> <mi>σ</mi> <mi>a</mi> </msub> </semantics></math>—stress amplitude; <math display="inline"><semantics> <msub> <mi>N</mi> <mi>f</mi> </msub> </semantics></math>—number of cycles to failure; Dist—Distribution; 3p—3 parameters).</p>
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<p>Regression model considering the data sets of rotating bending and axial tensile fatigue loading (<span class="html-italic">R</span>—stress ratio; <math display="inline"><semantics> <msub> <mi>P</mi> <mi>f</mi> </msub> </semantics></math>—probability of failure; RB(25Hz)—specimen under rotating bending at 25 Hz; AT(150 Hz)—specimen under axial tension at 150 Hz; AT(20kHz)—specimen under axial tension at 20 kHz Avg.—average curve; <math display="inline"><semantics> <msub> <mi>σ</mi> <mi>a</mi> </msub> </semantics></math>—stress amplitude; <math display="inline"><semantics> <msub> <mi>N</mi> <mi>f</mi> </msub> </semantics></math>—number of cycles to failure; Dist—Distribution; 3p—3 parameters).</p>
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<p>PSN hyperbolic field considering only the rotating bending and tensile loading (<span class="html-italic">R</span>—stress ratio; <math display="inline"><semantics> <msub> <mi>P</mi> <mi>f</mi> </msub> </semantics></math>—probability of failure; Est—estimation for run-out data; RB— specimens under rotating bending; AT— specimens under axial tension Avg.—average curve; <math display="inline"><semantics> <msub> <mi>σ</mi> <mi>a</mi> </msub> </semantics></math>—stress amplitude; <math display="inline"><semantics> <msub> <mi>N</mi> <mi>f</mi> </msub> </semantics></math>—number of cycles to failure; Dist—Distribution; 3p—3 parameters).</p>
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<p>Comparison between the PSN power and hyperbolic fields from the rotating bending and tensile fatigue data (<math display="inline"><semantics> <msub> <mi>P</mi> <mi>f</mi> </msub> </semantics></math>—probability of failure; Est—estimation for run-out data; RB—rotating bending; AT—axial tension/compression Avg.—average curve; Dist—Distribution; 3p—3 parameters).</p>
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<p>Fracture surfaces for the three stress amplitude regions of the rotating bending specimens: (<b>A</b>) <math display="inline"><semantics> <msub> <mi>σ</mi> <mi>a</mi> </msub> </semantics></math> = 1057.96 MPa, (<b>B</b>) <math display="inline"><semantics> <msub> <mi>σ</mi> <mi>a</mi> </msub> </semantics></math> = 1020.62 MPa, (<b>C</b>) <math display="inline"><semantics> <msub> <mi>σ</mi> <mi>a</mi> </msub> </semantics></math> = 816.56 MPa, and (<b>D</b>) <math display="inline"><semantics> <msub> <mi>σ</mi> <mi>a</mi> </msub> </semantics></math> = 595.67 MPa.</p>
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<p>Fracture surfaces obtained for different stress amplitudes and testing frequencies of specimens under uniaxial testing conditions (<math display="inline"><semantics> <msub> <mi>R</mi> <mi>σ</mi> </msub> </semantics></math> = −1.0): (<b>A</b>) <math display="inline"><semantics> <msub> <mi>σ</mi> <mi>a</mi> </msub> </semantics></math> = 770 MPa (20 kHz), (<b>B</b>) <math display="inline"><semantics> <msub> <mi>σ</mi> <mi>a</mi> </msub> </semantics></math> = 705 MPa (20 kHz), (<b>C</b>) <math display="inline"><semantics> <msub> <mi>σ</mi> <mi>a</mi> </msub> </semantics></math> = 670 MPa (20 kHz), (<b>D</b>) <math display="inline"><semantics> <msub> <mi>σ</mi> <mi>a</mi> </msub> </semantics></math> = 680 MPa (20 kHz), and (<b>E</b>) <math display="inline"><semantics> <msub> <mi>σ</mi> <mi>a</mi> </msub> </semantics></math> = 650 MPa (150 Hz).</p>
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<p>Comparison of different crack propagation zones of a fish-eye fracture surface. (<b>A</b>,<b>B</b>) Close to the non-metallic inclusion and (<b>C</b>) away from the initiation zone.</p>
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<p>Fatigue fracture surfaces depending on the amplitude stress level from the high-cycle up to giga-cycle and for rotating bending and uniaxial tension/compression loads (<span class="html-italic">R</span>—stress ratio; <math display="inline"><semantics> <msub> <mi>P</mi> <mi>f</mi> </msub> </semantics></math>—probability of failure; Est—estimation for run-out data; RB—specimens under rotating bending; AT—specimens under axial tension/compression Avg.—average curve; <math display="inline"><semantics> <msub> <mi>σ</mi> <mi>a</mi> </msub> </semantics></math>—stress amplitude; <math display="inline"><semantics> <msub> <mi>N</mi> <mi>f</mi> </msub> </semantics></math>—number of cycles to failure; Dist—Distribution; 3p—3 parameters).</p>
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<p>Analysis of the chemical composition of the non-metallic inclusion and the metallic matrix using EDS: Z1—Zone 1; Z2—Zone 2.</p>
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<p>Chemical composition analysis of the non-metallic slender inclusions and the matrix using EDS: Z1—Zone 1; Z2—Zone 2; Z3—Zone 3; Z4—Zone 4.</p>
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16 pages, 4554 KiB  
Article
Design of Tool Shape and Evaluation of Deformation Behavior by Digital Image Correlation Method in V-Bending of Sheet Metal Using Plastic Tools Manufactured by 3D Printer
by Naotaka Nakamura, Yuri Hata, Witthaya Daodon, Daiki Ikeda, Nozomu Adachi, Yoshikazu Todaka and Yohei Abe
Materials 2025, 18(3), 608; https://doi.org/10.3390/ma18030608 - 29 Jan 2025
Abstract
In the V-bending of sheet metals using a pair of plastic punch and die manufactured by a 3D printer, the effects of two different dimensions designed with the same tool geometry on the deformation behaviors of the punch, die, and sheet were evaluated. [...] Read more.
In the V-bending of sheet metals using a pair of plastic punch and die manufactured by a 3D printer, the effects of two different dimensions designed with the same tool geometry on the deformation behaviors of the punch, die, and sheet were evaluated. The deformation behavior and strain distribution of the punch, die, and sheet were analyzed using a digital image correlation method. Sheets from pure aluminum to ultra-high-strength steel were bent using the two tools with different spans; one was designed on the assumption of tool steel material, and the other was designed on the assumption of plastic material. In both tools, the large compressive strain appeared around the center of the punch tip and on the corners of the die. The tools with a long span for the plastic material gave a lower bending force and small deformation of the plastic tools. The angle difference between a bent sheet at the bottom dead center and a tool was smaller for the tools with the long span, although the springback in the bent sheet appeared. It was found that the design method on the assumption of the plastic material is effective for the V-bending plastic tools. Full article
(This article belongs to the Special Issue State of the Art in Materials for Additive Manufacturing)
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<p>(<b>a</b>) Specimen for compression test, (<b>b</b>) compression test, and (<b>c</b>) nominal stress-strain curve measured from simple compression test of PLA (unit: mm).</p>
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<p>Relationships between compressive plastic strain in height direction and number of compressions in repeated compression test of PLA for <span class="html-italic">σ</span><sub>c</sub> = 20 MPa and 40 MPa.</p>
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<p>Dimensions of plastic tools and sheet for V-bending (unit: mm).</p>
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<p>Image sensing conditions for observing deformation of tools and bend angles of sheet in V-bending using digital camera. (<b>a</b>) Top view and (<b>b</b>) front view (unit: mm).</p>
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<p>Strain distributions of punch and die in height direction in bending of (<b>a</b>) 440 MPa and (<b>b</b>) 980 MPa steel sheets for <span class="html-italic">W</span> = 12 mm.</p>
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<p>Strain distributions of punch and die in height direction in bending of 980 MPa steel sheet for <span class="html-italic">W</span> = 24 mm.</p>
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<p>Bending force-punch stroke curves for <span class="html-italic">W</span> = (<b>a</b>) 12 mm and (<b>b</b>) 24 mm.</p>
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<p>Variations in bend angle with punch stroke for <span class="html-italic">W</span> = (<b>a</b>) 12 mm and (<b>b</b>) 24 mm.</p>
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<p>(<b>a</b>) Angle from 90° and (<b>b</b>) springback angle of bent sheets for <span class="html-italic">W</span> = 12 mm and 24 mm.</p>
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<p>Strain distributions of punch and die in height direction at bottom dead center for <span class="html-italic">W</span> = (<b>a</b>) 12 mm and (<b>b</b>) 24 mm.</p>
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<p>Relationships between mean strain of punch in height direction and punch stroke for <span class="html-italic">W</span> = (<b>a</b>) 12 mm and (<b>b</b>) 24 mm.</p>
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<p>Relationship between maximum bending force and maximum free bending force.</p>
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<p>Relationship between bend angle of sheet at bottom dead center and maximum bending force.</p>
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<p>Relationships between springback angle of bent sheet and tensile strength of sheet for <span class="html-italic">W</span> = 12 mm and 24 mm.</p>
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21 pages, 7403 KiB  
Article
Low-Temperature, Highly Sensitive Ammonia Sensors Based on Nanostructured Copper Iodide Layers
by Sergey I. Petrushenko, Mateusz Fijalkowski, Kinga Adach, Denis Fedonenko, Yevhenii M. Shepotko, Sergei V. Dukarov, Volodymyr M. Sukhov, Alina L. Khrypunova and Natalja P. Klochko
Chemosensors 2025, 13(2), 29; https://doi.org/10.3390/chemosensors13020029 - 22 Jan 2025
Viewed by 336
Abstract
Chemiresistive ammonia gas sensors with a low limit of detection of 0.15 ppm and moisture-independent characteristics based on p-type copper iodide (CuI) semiconductor films have been developed. CuI films were deposited on glass and polyethylene terephthalate (PET) substrates using a Successive Ionic [...] Read more.
Chemiresistive ammonia gas sensors with a low limit of detection of 0.15 ppm and moisture-independent characteristics based on p-type copper iodide (CuI) semiconductor films have been developed. CuI films were deposited on glass and polyethylene terephthalate (PET) substrates using a Successive Ionic Layer Adsorption and Reaction method to fabricate CuI/glass and CuI/PET gas sensors, respectively. They have a nanoscale morphology, an excess iodine and sulfur impurity content, a zinc blende γ-CuI crystal structure with a grain size of ~34 nm and an optical band gap of about 2.95 eV. The high selective sensitivity of both sensors to NH3 is explained by the formation of the [Cu(NH3)2]+ complex. At 5 °C, the responses to 3 ppm ammonia in air in terms of the relative resistance change were 24.5 for the CuI/glass gas sensor and 28 for the CuI/PET gas sensor, with short response times of 50 s to 210 s and recovery times of 10–70 s. The sensors have a fast response–recovery and their performance was well maintained after long-term stability testing for 45 days. After 1000 repeated bends of the flexible CuI/PET gas sensor in different directions, with bending angles up to 180° and curvature radii up to 0.25 cm, the response changes were only 3%. Full article
(This article belongs to the Special Issue Functional Nanomaterial-Based Gas Sensors and Humidity Sensors)
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<p>(<b>a</b>)—Schematic diagram of the automatic SILAR method used in this work. (<b>b</b>)—Photographs of the high-precision LCR meter GW INSTEK LCR 6002 (GW Instek, Taipei, Taiwan) (top left) and RMS digital multimeter UNI-T UT171C (top right) used in gas sensing measurements. Schematic structure of NH<sub>3</sub> gas sensors (bottom middle) and photograph of CuI/glass sensor (bottom left) and CuI/PET sensor in four bent states (bottom right).</p>
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<p>SEM images of the CuI/glass sample with low (<b>a</b>) and high (<b>b</b>) magnification. (<b>c</b>)—SEM image of a thin-film contact obtained by RF magnetron sputtering of the Au80Pd20 alloy on the copper iodide surface of the CuI/glass gas sensor. (<b>d</b>)—EDS spectrum of the CuI/glass sample. (<b>e</b>)—Overall energy-dispersive X-ray spectroscopy (EDS) map of CuI/glass sample. Elemental EDS mapping of individual elements in the CuI/glass sample: (<b>f</b>)—Cu; (<b>g</b>)—I; (<b>h</b>)—S.</p>
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<p>SEM images of the CuI/PET sample with low (<b>a</b>) and high (<b>b</b>) magnification. (<b>c</b>)—EDS spectrum of the CuI/PET sample. (<b>d</b>)—Overall EDS map of CuI/PET sample. Elemental EDS mapping of individual elements in the CuI/PET sample: (<b>e</b>)—Cu; (<b>f</b>)—I; (<b>g</b>)—S.</p>
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<p>X-ray diffraction pattern of CuI/PET sample.</p>
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<p>Optical properties of copper iodide films in CuI/glass and CuI/PET samples obtained by the automatic SILAR method: (<b>a</b>)—transmission spectra; (<b>b</b>)—diffuse reflectance spectra; (<b>c</b>)—Tauc plots; (<b>d</b>)—graphs for determining the band gap of nanostructured CuI films using the Kubelka–Munk function. Inserts in (<b>a</b>,<b>b</b>) contain data on uncoated glass and PET substrates.</p>
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<p>Diagrams of responses <span class="html-italic">S*</span> = Δ<span class="html-italic">R/R<sub>0</sub></span> to 3 ppm of NH<sub>3</sub>, 12 ppm of H<sub>2</sub>O, 120 ppm of acetone, 120 ppm of ethanol and to <span class="html-italic">RH</span> 100% for CuI/glass and CuI/PET sensors operating at different temperatures: (<b>a</b>)—5 °C; (<b>b</b>)—22 °C; (<b>c</b>)—35 °C; (<b>d</b>)—45 °C; (<b>e</b>)—55 °C. (<b>f</b>)—Plots of responses <span class="html-italic">S*</span> = Δ<span class="html-italic">R/R<sub>0</sub></span> of CuI/glass and CuI/PET sensors to water, acetone and ethanol at different temperatures.</p>
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<p>Response curves <span class="html-italic">S</span> = <span class="html-italic">R<sub>g</sub>/R<sub>0</sub></span> after sensor exposure to 0.75 ppm NH<sub>3</sub> and recovery when the atmosphere was changed back to dry air for the developed gas sensors operating at different temperatures: (<b>a</b>)—CuI/glass sensor at 5 °C; (<b>b</b>)—CuI/glass sensor at 22 °C; (<b>c</b>)—CuI/glass sensor at 55 °C; (<b>d</b>)—CuI/PET sensor at 5 °C; (<b>e</b>)—CuI/ PET sensor at 22 °C; (<b>f</b>)—CuI/ PET sensor at 55 °C.</p>
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<p>Sensitivity of CuI/glass (<b>a</b>) and CuI/PET (<b>b</b>) sensors to different ammonia concentrations depending on operating temperatures in the range of 5–55 °C.</p>
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<p>Dependencies of the responses of CuI/glass (<b>a</b>) and CuI/PET (<b>b</b>) sensors operating at different temperatures on the concentration of ammonia in the air in the range of 0.15–3 ppm.</p>
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<p>Relative resistance change (<span class="html-italic">R<sub>bent</sub></span> − <span class="html-italic">R<sub>0</sub>)</span>/<span class="html-italic">R<sub>0</sub></span> at different bending angles φ of the CuI/PET sensor in dry air at 22 °C (the sensor was bent in the direction along the thin-film strip contacts of Au80Pd20). The photo and diagram at the top show the sensor bending outward with tensile strains in CuI, which corresponds to the red line in the graph. The photo and diagram at the bottom show the sensor bending inward with compressive strains in CuI, which corresponds to the blue line in the graph.</p>
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<p>Relative change in resistance (<span class="html-italic">R<sub>bent</sub></span> − <span class="html-italic">R<sub>0</sub></span>)/<span class="html-italic">R<sub>0</sub></span> at bending angles φ of 170–180° and different radii of curvature <span class="html-italic">r</span> of the CuI/PET sensor in dry air at 22 °C (the sensor was bent in the direction along the thin-film strip contacts of Au80Pd20). (<b>a</b>)—The diagram on the top shows the outward bending of the sensor with tensile strains in CuI, which corresponds to the red line in the graph. The diagram on the bottom shows the inward bending of the sensor with compressive strains in CuI, which corresponds to the blue line in the graph. (<b>b</b>)—Photo of the unbent CuI/PET sensor. (<b>c</b>)—Photo of the CuI/PET sensor bent outward at a bending angle of φ ≈ 180°.</p>
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<p>Experiments on bending of the CuI/PET sensor in the direction perpendicular to the Au80Pd20 thin-film contacts in dry air at 22 °C. (<b>a</b>)—Relative change in resistance (<span class="html-italic">R<sub>bent</sub></span> − <span class="html-italic">R<sub>0</sub></span>)/<span class="html-italic">R<sub>0</sub></span> at bending angles φ of 170–180° and different radii of curvature <span class="html-italic">r</span>. The diagram on the top shows the outward bending of the sensor with tensile strains in CuI, which corresponds to the red line in the graph. The diagram on the bottom shows the inward bending of the sensor with compressive strains in CuI, which corresponds to the blue line in the graph. (<b>b</b>)—Photo of the inward bending of the CuI/PET sensor to <span class="html-italic">r</span> of 5 cm and φ of ~25°. (<b>c</b>)—Photo of the inward bending of the CuI/PET sensor to <span class="html-italic">r</span> of 3.5 cm and φ of ~40°. (<b>d</b>)—Photo of the inward bending of the CuI/PET sensor to <span class="html-italic">r</span> of 1.5 cm and φ of ~170°. (<b>e</b>)—Photo of the inward bending of the CuI/PET sensor to <span class="html-italic">r</span> of 0.25 cm and φ of ~180°.</p>
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<p>Responses of CuI/glass (<b>a</b>) and CuI/PET (<b>b</b>) sensors to 3 ppm NH<sub>3</sub> at 5 °C and 90% relative humidity during a long-term stability test. The CuI/PET sensor was additionally bent in different directions with bending angles up to 180° and curvature radii up to 0.25 cm.</p>
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9 pages, 15780 KiB  
Article
Influence of Printing Orientation on the Mechanical Properties of Provisional Polymeric Materials Produced by 3D Printing
by Fábio Hideo Kaiahara, Eliane Cristina Gava Pizi, Fabiana Gouveia Straioto, Lucas David Galvani, Milton Carlos Kuga, Thalita Ayres Arrué, Ageu Raupp Junior, Marcus Vinícius Reis Só, Jefferson Ricardo Pereira and Hugo Vidotti
Polymers 2025, 17(3), 265; https://doi.org/10.3390/polym17030265 - 21 Jan 2025
Viewed by 391
Abstract
This study investigates the impact of printing layer orientation on the mechanical properties of 3D-printed temporary prosthetic materials. Traditionally, temporary prostheses are fabricated using acrylic resin (polymethyl methacrylate), but advancements have introduced bis-acrylic resins, CAD/CAM-based acrylic resin (milled), and 3D printing technologies. In [...] Read more.
This study investigates the impact of printing layer orientation on the mechanical properties of 3D-printed temporary prosthetic materials. Traditionally, temporary prostheses are fabricated using acrylic resin (polymethyl methacrylate), but advancements have introduced bis-acrylic resins, CAD/CAM-based acrylic resin (milled), and 3D printing technologies. In 3D printing, material is manufactured in overlapping layers, which can be oriented in different directions, directly affecting the material’s resistance. Specimens were designed as bars (2 mm × 2 mm × 25 mm) and grouped according to their printing orientation: BP0 (0 degrees), BP45 (45 degrees), and BP90 (90 degrees). The models were created using Fusion 360 software (version 2.0.12600) and printed on a 3D DLP printer with DLP Slicer software (Chitu DLP Slicer, CBD Tech, version v1.9.0). The bars were then subjected to 3-point bending tests using an Instron Universal Testing Machine to measure Flexural Strength (FS) and Flexural Modulus (FM). Results demonstrated that the BP90 group exhibited the highest Flexural Strength (114.71 ± 7.61 MPa), followed by BP45 (90.10 ± 8.45 MPa) and BP0 (80.90 ± 4.01 MPa). Flexural Modulus was also highest in the BP90 group (3.74 ± 3.64 GPa), followed by BP45 (2.85 ± 2.70 GPa) and BP0 (2.52 ± 2.44 GPa). Significant statistical differences (p < 0.05) were observed, indicating changes in the mechanical properties of the 3D-printed material. The study concludes that printing orientation significantly influences the mechanical properties of temporary prosthetic materials, making the selection of an optimal orientation essential to enhance material performance for its intended application. Full article
(This article belongs to the Section Polymer Analysis and Characterization)
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<p>Specimens removed and identified (red P0, blue P45, and black P90).</p>
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<p>Specimens under polymerization process (series of 8 exposures of 30 s polymerization).</p>
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<p>2-parts metallic matrix (<b>A</b>), joined with type C clamp (<b>B</b>,<b>C</b>).</p>
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<p>PMMA plates with possible bubbles marks and perforations.</p>
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<p>Acrylic resin plates base-fixed and sectioned in a precision cutting machine.</p>
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<p>Emic/Instron 23-2S.</p>
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<p>3-point bending test.</p>
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17 pages, 12970 KiB  
Article
Design of Dielectric Elastomer Actuator and Its Application in Flexible Gripper
by Xiaoyu Meng, Jiaqing Xie, Haoran Pang, Wenchao Wei, Jiping Niu, Mingqiang Zhu, Fang Gu, Xiaohuan Fan and Haiyan Fan
Micromachines 2025, 16(1), 107; https://doi.org/10.3390/mi16010107 - 19 Jan 2025
Viewed by 446
Abstract
Dielectric elastomer actuators (DEAs) are difficult to apply to flexible grippers due to their small deformation range and low output force. Hence, a DEA with a large bending deformation range and output force was designed, and a corresponding flexible gripper was developed to [...] Read more.
Dielectric elastomer actuators (DEAs) are difficult to apply to flexible grippers due to their small deformation range and low output force. Hence, a DEA with a large bending deformation range and output force was designed, and a corresponding flexible gripper was developed to realize the function of grasping objects of different shapes. The relationship between the pre-stretch ratio and DEA deformation degree was tested by experiments. Based on the performance test results of the dielectric elastomer (DE), the bending deformation process of DEAs with different shapes was simulated by Finite Element Method (FEM) simulation. DEAs with different shapes were prepared through laser cutting and the relationship between the voltage and the bending angle, and the output force of the DEAs was measured. The result shows that under uniaxial stretching, the deformation of the DEA in the stretching direction gradually increases and decreases in the unstretched direction with the increase in the pre-stretch ratio. Under biaxial stretching, DEA deformation increases with the increase in the pre-stretch ratio. The shape of the DEA has a certain influence on the bending deformation range under the same conditions, and the elliptical DEA has a larger bending deformation range and higher output force compared with the rectangular DEA and the trapezium DEA. The elliptical DEA can produce a bending deformation of 40° and an output force of 37.2 mN at a voltage of 24 kV. The three-finger flexible gripper composed of an elliptical DEA can realize the grasping of a paper cup. Full article
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<p>The driving principle of DEAs.</p>
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<p>The structure and deformation process of DEAs.</p>
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<p>The preparation schematic diagram of the DEA.</p>
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<p>Hyperelastic tensile test.</p>
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<p>Relative dielectric constant test.</p>
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<p>The test results under uniaxial stretching, (<b>a</b>) the deformation process of square DEA, (<b>b</b>) the relationship between transverse elongation and voltage, and (<b>c</b>) the relationship between vertical elongation and voltage.</p>
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<p>The test results under biaxial stretching, (<b>a</b>) the deformation process of square DEA, (<b>b</b>) the relationship between transverse elongation and voltage, and (<b>c</b>) the relationship between vertical elongation and voltage.</p>
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<p>Material test results, (<b>a</b>) hyperelastic test results, and (<b>b</b>) relative dielectric constant test results.</p>
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<p>The initial bending angle of DEA when the pre-stretch ratio is 1.25, 1.50, and 1.75.</p>
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<p>The displacement nephograms of DEAs with different shapes at voltages of 0 kV, 10 kV, and 20 kV; (<b>a</b>–<b>c</b>) the deformation process of rectangular DEA, (<b>d</b>–<b>f</b>) the deformation process of trapezium DEA, and (<b>g</b>–<b>i</b>) the deformation process of elliptical DEA.</p>
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<p>The bending state of DEAs with different shapes under voltages of 0 kV, 8 kV, 16 kV, and 24 kV; (<b>a</b>–<b>d</b>) the deformation process of rectangular DEA, (<b>e</b>–<b>h</b>) the deformation process of trapezium DEA, and (<b>i</b>–<b>l</b>) the deformation process of elliptical DEA.</p>
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<p>The performance test of DEA; (<b>a</b>) the change in bending angle with voltage, and (<b>b</b>) the change in output force with voltage.</p>
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<p>The difference between simulation data and experimental data.</p>
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<p>The grabbing test of three-fingered gripper, (<b>a</b>) the frame size of the gripper, (<b>b</b>) the three-dimensional model of the gripper, (<b>c</b>) three-finger gripper physical map, (<b>d</b>–<b>f</b>) test of grasping balloons, and (<b>g</b>–<b>i</b>) test of grasping paper cups.</p>
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14 pages, 4800 KiB  
Article
Design and Analysis of Compact High–Performance Lithium–Niobate Electro–Optic Modulator Based on a Racetrack Resonator
by Zixin Chen, Jianping Li, Weiqin Zheng, Hongkang Liu, Quandong Huang, Ya Han and Yuwen Qin
Photonics 2025, 12(1), 85; https://doi.org/10.3390/photonics12010085 - 17 Jan 2025
Viewed by 470
Abstract
With the ever-growing demand for high-speed optical communications, microwave photonics, and quantum key distribution systems, compact electro-optic (EO) modulators with high extinction ratios, large bandwidth, and high tuning efficiency are urgently pursued. However, most integrated lithium–niobate (LN) modulators cannot achieve these high performances [...] Read more.
With the ever-growing demand for high-speed optical communications, microwave photonics, and quantum key distribution systems, compact electro-optic (EO) modulators with high extinction ratios, large bandwidth, and high tuning efficiency are urgently pursued. However, most integrated lithium–niobate (LN) modulators cannot achieve these high performances simultaneously. In this paper, we propose an improved theoretical model of a chip-scale electro-optic (EO) microring modulator (EO-MRM) based on X-cut lithium–niobate-on-insulator (LNOI) with a hybrid architecture consisting of a 180-degree Euler bend in the coupling region, double-layer metal electrode structure, and ground–signal–signal–ground (G-S-S-G) electrode configuration, which can realize highly comprehensive performance and a compact footprint. After parameter optimization, the designed EO-MRM exhibited an extinction ratio of 38 dB. Compared to the structure without Euler bends, the increase was 35 dB. It also had a modulation bandwidth of 29 GHz and a tunability of 8.24 pm/V when the straight waveguide length was 100 μm. At the same time, the proposed device footprint was 1.92 × 104 μm2. The proposed MRM model provides an efficient solution to high-speed optical communication systems and microwave photonics, which is helpful for the fabrication of high-performance and multifunctional photonic integrated devices. Full article
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<p>(<b>a</b>) A schematic diagram of the proposed racetrack resonator with a double-layer electrode. Inset: the cross-section of coupling area. (<b>b</b>) A top view of the racetrack microring resonator. (<b>c</b>) The optical mode field and intensity distribution of the Euler bend with a waveguide width of 0.8 µm, simulated by FDTD.</p>
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<p>(<b>a</b>) Lumerical MODE simulation of the fundamental TE<sub>0</sub> optical mode of the waveguide. (<b>b</b>) The calculated optical effective index of the waveguide.</p>
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<p>(<b>a</b>) The coupling coefficient <span class="html-italic">κ</span><sup>2</sup> and (<b>b</b>) the transmission coefficient <span class="html-italic">t</span><sup>2</sup> vary with w<sub>gap</sub> in the coupling region at the wavelength of 1550 nm.</p>
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<p>(<b>a</b>) The coupling coefficient <span class="html-italic">κ</span><sup>2</sup> and (<b>b</b>) the transmission coefficient <span class="html-italic">t</span><sup>2</sup> vary with w<sub>1</sub> in the coupling region at the wavelength of 1550 nm.</p>
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<p>The BW and <span class="html-italic">Q</span> factor performances with the variation in <span class="html-italic">Lc</span> of the resonator.</p>
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<p>(<b>a</b>) The coupling and transmission coefficients with a variation in wavelength, when w<sub>gap</sub> = 0.7 μm and w<sub>1</sub> = 0.6 μm. (<b>b</b>) Transmission spectrum of the resonator with different bends used in the coupling region at the wavelength of 1550 nm.</p>
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<p>(<b>a</b>) A top view of the proposed tunable racetrack resonator with double-layer electrodes. (<b>b</b>) The simulated TE optical mode field profile at 1550 nm and the electric field between the double-layer electrodes. Here, the TFLN waveguide was formed by a 300 nm × 0.8 µm LN loading ridge. (<b>c</b>) A schematic of a unit cell of the electrode structure. (<b>d</b>) The simulation result of the influence of h and d on metal loss.</p>
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<p>Metal loss analysis for different electrode designs. (<b>a</b>) Metal electrodes were placed directly on the waveguide. (<b>b</b>) A 2.8 μm-wide layer of SiO<sub>2</sub> was added between the double metal electrode and the waveguide.</p>
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<p>(<b>a</b>) The simulated transmission spectrum of the TE mode of the passive racetrack resonator. (<b>b</b>) The detailed spectrum at 1550.118 nm. (<b>c</b>) The spectrum under different voltages of the TE mode at 1550.118 nm. (<b>d</b>) Resonant wavelength shifts as a function of the applied voltage.</p>
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21 pages, 30504 KiB  
Article
Bending Performance of Diamond Lattice Cylindrical Shells
by Sheng Li, Laiyu Liang, Ping Yang, Shaoan Li and Yaozhong Wu
Materials 2025, 18(2), 272; https://doi.org/10.3390/ma18020272 - 9 Jan 2025
Viewed by 281
Abstract
The Diamond lattice cylindrical shell (Diamond LCS) was proposed by a mapping approach based on the triply periodic minimal surfaces (TPMS). The finite element models were built and their accuracy was verified by experimental results. Parameter studies were carried out to investigate the [...] Read more.
The Diamond lattice cylindrical shell (Diamond LCS) was proposed by a mapping approach based on the triply periodic minimal surfaces (TPMS). The finite element models were built and their accuracy was verified by experimental results. Parameter studies were carried out to investigate the effect of geometric and loading parameters on the bending properties of the Diamond LCSs by the finite element model. The results show that Diamond LCS has a stable “V” deformation pattern under a three-point bending load. In the range of relative density (RD) = 15–30%, the higher the RD, the better the lateral bending performance of the Diamond LCS structure. The larger the variation radial coefficient, the higher the lateral load-carrying capacity of the structure. The smaller the loading angle of the punch, the better the lateral bending performance of the Diamond LCS structure. However, if the loading angle is too small, the structure is prone to large torsional deformation, and the deformation tends to destabilize. The increase in punch diameter effectively improves the deformation pattern and bending energy absorption characteristics of the structure. The smaller the span of the cylindrical support, the better the bending energy absorption characteristics of the structure. Full article
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<p>(<b>a</b>) The construction process of the LCS structure using TPMS surfaces. (<b>b</b>) The proposed Diamond LCS structure.</p>
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<p>(<b>a</b>) FE model of Diamond LCS under three-point bending; (<b>b</b>) illustration of the loading angle.</p>
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<p>Effective stress-plastic strain curves of the 316L stainless steel [<a href="#B22-materials-18-00272" class="html-bibr">22</a>].</p>
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<p>Illustration of the energy absorption properties of the energy absorber.</p>
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<p>Verification of the numerical model: (<b>a</b>) Geometric details of three-point bending of thin-walled square tube in reference [<a href="#B28-materials-18-00272" class="html-bibr">28</a>]; (<b>b</b>) Experimental and simulation force-displacement curves of square tube bending; and (<b>c</b>) Comparison of the experimental and simulation flexural deformation modes of thin-walled square tubes with wall thickness t = 0. 95 mm.</p>
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<p>Verification of the numerical model [<a href="#B22-materials-18-00272" class="html-bibr">22</a>]. Comparison of (<b>a</b>) the experimental and simulation deformation modes and (<b>b</b>) Experimental and simulation force-displacement curves.</p>
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<p>Force-displacement curves of Diamond LCS with different <math display="inline"><semantics> <mrow> <mi>R</mi> <mi>D</mi> </mrow> </semantics></math>.</p>
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<p>Bending deformation modes of Diamond LCS with different <math display="inline"><semantics> <mrow> <mi>R</mi> <mi>D</mi> </mrow> </semantics></math>.</p>
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<p>Crashworthiness of the Diamond LCS with different <math display="inline"><semantics> <mrow> <mi>R</mi> <mi>D</mi> </mrow> </semantics></math>: (<b>a</b>) <math display="inline"><semantics> <mrow> <mi>E</mi> <mi>A</mi> </mrow> </semantics></math>, (<b>b</b>) <math display="inline"><semantics> <mrow> <mi>S</mi> <mi>E</mi> <mi>A</mi> </mrow> </semantics></math>, (<b>c</b>) <math display="inline"><semantics> <mrow> <mi>P</mi> <mi>C</mi> <mi>F</mi> </mrow> </semantics></math>, and (<b>d</b>) <math display="inline"><semantics> <mrow> <mi>C</mi> <mi>F</mi> <mi>E</mi> </mrow> </semantics></math>.</p>
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<p>Force-displacement curves of Diamond LCS with different <math display="inline"><semantics> <mrow> <mi>γ</mi> </mrow> </semantics></math>.</p>
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<p>Bending deformation modes of Diamond LCS with different <math display="inline"><semantics> <mrow> <mi>γ</mi> </mrow> </semantics></math>.</p>
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<p>Crashworthiness of the Diamond LCS with different <math display="inline"><semantics> <mrow> <mi>γ</mi> </mrow> </semantics></math>: (<b>a</b>) <math display="inline"><semantics> <mrow> <mi>E</mi> <mi>A</mi> </mrow> </semantics></math>, (<b>b</b>) <math display="inline"><semantics> <mrow> <mi>S</mi> <mi>E</mi> <mi>A</mi> </mrow> </semantics></math>, (<b>c</b>) <math display="inline"><semantics> <mrow> <mi>P</mi> <mi>C</mi> <mi>F</mi> </mrow> </semantics></math>, and (<b>d</b>) <math display="inline"><semantics> <mrow> <mi>C</mi> <mi>F</mi> <mi>E</mi> </mrow> </semantics></math>.</p>
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<p>Force-displacement curves of Diamond LCS with different <math display="inline"><semantics> <mrow> <mi>α</mi> </mrow> </semantics></math>.</p>
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<p>Bending deformation modes of Diamond LCS with different <math display="inline"><semantics> <mrow> <mi>α</mi> </mrow> </semantics></math>.</p>
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<p>Crashworthiness of the Diamond LCS with different <math display="inline"><semantics> <mrow> <mi>α</mi> </mrow> </semantics></math>: (<b>a</b>) <math display="inline"><semantics> <mrow> <mi>E</mi> <mi>A</mi> </mrow> </semantics></math>, (<b>b</b>) <math display="inline"><semantics> <mrow> <mi>S</mi> <mi>E</mi> <mi>A</mi> </mrow> </semantics></math>, (<b>c</b>) <math display="inline"><semantics> <mrow> <mi>P</mi> <mi>C</mi> <mi>F</mi> </mrow> </semantics></math>, and (<b>d</b>) <math display="inline"><semantics> <mrow> <mi>C</mi> <mi>F</mi> <mi>E</mi> </mrow> </semantics></math>.</p>
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<p>Force-displacement curves of Diamond LCS with different <math display="inline"><semantics> <mrow> <mi>θ</mi> </mrow> </semantics></math>.</p>
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<p>Bending deformation modes of Diamond LCS with different <math display="inline"><semantics> <mrow> <mi>θ</mi> </mrow> </semantics></math>.</p>
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<p>Crashworthiness of the Diamond LCS with different <math display="inline"><semantics> <mrow> <mi>θ</mi> </mrow> </semantics></math>: (<b>a</b>) <math display="inline"><semantics> <mrow> <mi>E</mi> <mi>A</mi> </mrow> </semantics></math>, (<b>b</b>) <math display="inline"><semantics> <mrow> <mi>S</mi> <mi>E</mi> <mi>A</mi> </mrow> </semantics></math>, (<b>c</b>) <math display="inline"><semantics> <mrow> <mi>P</mi> <mi>C</mi> <mi>F</mi> </mrow> </semantics></math>, and (<b>d</b>) <math display="inline"><semantics> <mrow> <mi>C</mi> <mi>F</mi> <mi>E</mi> </mrow> </semantics></math>.</p>
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<p>Force-displacement curves of Diamond LCS with different <math display="inline"><semantics> <mrow> <mi>D</mi> </mrow> </semantics></math>.</p>
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<p>Bending deformation modes of Diamond LCS with different <math display="inline"><semantics> <mrow> <mi>D</mi> </mrow> </semantics></math>.</p>
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<p>Crashworthiness of the Diamond LCS with different <math display="inline"><semantics> <mrow> <mi>D</mi> </mrow> </semantics></math>: (<b>a</b>) <math display="inline"><semantics> <mrow> <mi>E</mi> <mi>A</mi> </mrow> </semantics></math>, (<b>b</b>) <math display="inline"><semantics> <mrow> <mi>S</mi> <mi>E</mi> <mi>A</mi> </mrow> </semantics></math>, (<b>c</b>) <math display="inline"><semantics> <mrow> <mi>P</mi> <mi>C</mi> <mi>F</mi> </mrow> </semantics></math>, and (<b>d</b>) <math display="inline"><semantics> <mrow> <mi>C</mi> <mi>F</mi> <mi>E</mi> </mrow> </semantics></math>.</p>
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<p>Force-displacement curves of Diamond LCS with different <math display="inline"><semantics> <mrow> <mi>S</mi> </mrow> </semantics></math>.</p>
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<p>Bending deformation modes of Diamond LCS with different <math display="inline"><semantics> <mrow> <mi>S</mi> </mrow> </semantics></math>.</p>
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<p>Crashworthiness of the Diamond LCS with different <math display="inline"><semantics> <mrow> <mi>S</mi> </mrow> </semantics></math>: (<b>a</b>) <math display="inline"><semantics> <mrow> <mi>E</mi> <mi>A</mi> </mrow> </semantics></math>, (<b>b</b>) <math display="inline"><semantics> <mrow> <mi>S</mi> <mi>E</mi> <mi>A</mi> </mrow> </semantics></math>, (<b>c</b>) <math display="inline"><semantics> <mrow> <mi>P</mi> <mi>C</mi> <mi>F</mi> </mrow> </semantics></math>, and (<b>d</b>) <math display="inline"><semantics> <mrow> <mi>C</mi> <mi>F</mi> <mi>E</mi> </mrow> </semantics></math>.</p>
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12 pages, 3122 KiB  
Article
Effect of p-InGaN Cap Layer on Low-Resistance Contact to p-GaN: Carrier Transport Mechanism and Barrier Height Characteristics
by Mohit Kumar, Laurent Xu, Timothée Labau, Jérôme Biscarrat, Simona Torrengo, Matthew Charles, Christophe Lecouvey, Aurélien Olivier, Joelle Zgheib, René Escoffier and Julien Buckley
Crystals 2025, 15(1), 56; https://doi.org/10.3390/cryst15010056 - 8 Jan 2025
Viewed by 642
Abstract
This study investigated the low contact resistivity and Schottky barrier characteristics in p-GaN by modifying the thickness and doping levels of a p-InGaN cap layer. A comparative analysis with highly doped p-InGaN revealed the key mechanisms contributing to low-resistance contacts. Atomic force microscopy [...] Read more.
This study investigated the low contact resistivity and Schottky barrier characteristics in p-GaN by modifying the thickness and doping levels of a p-InGaN cap layer. A comparative analysis with highly doped p-InGaN revealed the key mechanisms contributing to low-resistance contacts. Atomic force microscopy inspections showed that the surface roughness depends on the doping levels and cap layer thickness, with higher doping improving the surface quality. Notably, increasing the doping concentration in the p++-InGaN cap layer significantly reduced the specific contact resistivity to 6.4 ± 0.8 × 10−6 Ω·cm2, primarily through enhanced tunneling. Current–voltage (I–V) characteristics indicated that the cap layer’s surface properties and strain-induced polarization effects influenced the Schottky barrier height and reverse current. The reduction in barrier height by approximately 0.42 eV in the p++-InGaN layer enhanced hole tunneling, further lowering the contact resistivity. Additionally, polarization-induced free charges at the metal–semiconductor interface reduced band bending, thereby enhancing carrier transport. A transition in current conduction mechanisms was also observed, shifting from recombination tunneling to space-charge-limited conduction across different voltage ranges. This research underscores the importance of doping, cap layer thickness, and polarization effects in achieving ultra-low contact resistivity, offering valuable insights for improving the performance of p-GaN-based power devices. Full article
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<p>Schematic representation of the epitaxial structure of the device, consisting of the following layers: (<b>a</b>) p<sup>+</sup>-GaN top layer, (<b>b</b>) p<sup>++</sup>-GaN heavily doped cap layer, (<b>c</b>) p<sup>++</sup>-In<sub>0.15</sub>Ga<sub>0.85</sub>N heavily doped cap layer, (<b>d</b>) etched circular pad stack, and (<b>e</b>) circular transmission line model (CTLM) configuration with 1, 2, 4, 8, 19, and 49 μm spacing.</p>
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<p>AFM image of samples L.1, L.2, M., N.1, N.2, N.3, and O.</p>
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<p>Room temperature current−voltage (I−V) characteristics of (<b>a</b>) samples L.1, L.2, and M., (<b>b</b>) samples N.1, N.2, N.3, and O., and (<b>c</b>) the analysis of the current conduction mechanism under forward bias for there difference voltage range named region 1, 2 and 3 for the p<sup>++</sup>-InGaN/p<sup>++</sup>-GaN heterojunction in sample O.</p>
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<p>Plot of the specific contact resistance for various samples. The contact transfer length method (CTLM) was employed, where a logarithmic fit to the experimental data was used to determine the specific contact resistance values.</p>
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<p>Current−voltage (I−V) curves and corresponding specific contact resistance measurements at the linear regime close to 15 V before annealing (<b>a</b>,<b>b</b>) and after annealing at 760 °C in an N<sub>2</sub> atmosphere (<b>c</b>,<b>d</b>).</p>
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<p>(<b>a</b>) Plot of the Schottky barrier height (<math display="inline"><semantics> <mrow> <msub> <mrow> <mi>Φ</mi> </mrow> <mrow> <mi>B</mi> </mrow> </msub> </mrow> </semantics></math>) for various samples, highlighting variations in barrier properties. (<b>b</b>) Energy band diagrams depicting (<b>i</b>) the metal/p<sup>+</sup>-GaN and (<b>ii</b>) the metal/p<sup>++</sup>-InGaN/p<sup>+</sup>-GaN superlattice structure.</p>
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10 pages, 2472 KiB  
Article
High-Mobility All-Transparent TFTs with Dual-Functional Amorphous IZTO for Channel and Transparent Conductive Electrodes
by Min-Woo Park, Sohyeon Kim, Su-Yeon Son, Si-Won Kim, Tae-Kyun Moon, Pei-Chen Su and Kyoung-Kook Kim
Materials 2025, 18(2), 216; https://doi.org/10.3390/ma18020216 - 7 Jan 2025
Viewed by 392
Abstract
The increasing demand for advanced transparent and flexible display technologies has led to significant research in thin-film transistors (TFTs) with high mobility, transparency, and mechanical robustness. In this study, we fabricated all-transparent TFTs (AT-TFTs) utilizing amorphous indium-zinc-tin-oxide (a-IZTO) as a dual-functional material for [...] Read more.
The increasing demand for advanced transparent and flexible display technologies has led to significant research in thin-film transistors (TFTs) with high mobility, transparency, and mechanical robustness. In this study, we fabricated all-transparent TFTs (AT-TFTs) utilizing amorphous indium-zinc-tin-oxide (a-IZTO) as a dual-functional material for both the channel layer and transparent conductive electrodes (TCEs). The a-IZTO was deposited using radio-frequency magnetron sputtering, with its composition adjusted for both channel and electrode functionality. XRD analysis confirmed the amorphous nature of the a-IZTO layers, ensuring structural stability post-thermal annealing. The a-IZTO TCEs demonstrated high optical transparency (89.57% in the visible range) and excellent flexibility, maintaining a low sheet resistance with minimal degradation even after 100,000 bending cycles. The fabricated AT-TFTs exhibit superior field-effect mobility (30.12 cm2/V·s), an on/off current ratio exceeding 108, and a subthreshold swing of 0.36 V/dec. The AT-TFT device demonstrated a minimum transmittance of 75.46% in the visible light range, confirming its suitability for next-generation flexible and transparent displays. Full article
(This article belongs to the Section Advanced Nanomaterials and Nanotechnology)
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<p>(<b>a</b>) The three-dimensional schematic image of the AT-TFT device using amorphous SiO<sub>2</sub> and dual-functional a-IZTO for channel and TCEs. (<b>a’</b>) The cross-sectional schematic image of the AT-TFT device. (<b>b</b>) The top view image of AT-TFT which has 60 μm of channel width and 20 μm of channel length. (<b>c</b>) The image of the fabricated AT-TFT wafer which has high transparency.</p>
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<p>(<b>a</b>) The resistivity, mobility, carrier concentration, and metal atom proportion of a-IZTO for TCEs deposited by RF magnetron sputtering with RF power ranging from 30 to 90 W. The black line, blue line, and red line represent the resistivity, carrier mobility, and carrier concentration of a-IZTO for TCEs, respectively. (<b>b</b>) Bending test results of a-IZTO for TCE deposited by sputtering at RF power of 50 W (bending cycles = 0, 20,000, 40,000, 60,000, 80,000, 100,000). The inset image within the graph shows the flexible bending of the a-IZTO deposited on the PI film. The FE-SEM images of (<b>b’</b>) the as-deposited a-IZTO surface and (<b>b”</b>) the a-IZTO surface after 100,000 bending cycles. The red arrow in (<b>b”</b>) indicates the cracks on the a-IZTO surface after 100,000 bending cycles.</p>
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<p>The XRD results of dual-functional a-IZTO films for 30 nm-thick channels and 200 nm-thick TCEs before and after a thermal annealing process at 300 °C for 15 min. The broad peaks in the black box indicate that both IZTO for the channel and TCEs are in an amorphous state.</p>
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<p>(<b>a</b>) Schematic image of the AT-TFT, showing the film thickness and layer structure when measuring the transmittance of a-IZTO for the channel, AT-TFT, and a-IZTO for TCEs. (<b>b</b>) Transmittance results for a-IZTO in the channel, AT-TFT, and a-IZTO for TCEs. The inset table shows the average transmittance in the 400–700 nm range.</p>
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<p>(<b>a</b>) The output and (<b>b</b>) transfer characteristics of AT-TFTs using all-amorphous oxide materials, including dual-functional a-IZTO films for channel and TCEs. The red line represents the drain current on a logarithmic scale, plotted on the left y-axis, while the blue line shows the linear measurement of the drain current, plotted on the right y-axis.</p>
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14 pages, 5690 KiB  
Article
Investigation of Flow Characteristics in Valveless Piezoelectric Pumps with Airfoil Baffles at Varying Angles of Attack
by Jun Huang, Hiba Affane, Bo Zhang, Ming Kuang, Jian Xiong and Siyao Zhang
Appl. Sci. 2025, 15(1), 445; https://doi.org/10.3390/app15010445 - 6 Jan 2025
Viewed by 335
Abstract
To investigate the impact of airfoil angle of attack on the output performance of a valveless piezoelectric pump with airfoil baffles, this study conducted comprehensive performance tests and full-flow field simulations of piezoelectric pumps across a range of angles. At a driving voltage [...] Read more.
To investigate the impact of airfoil angle of attack on the output performance of a valveless piezoelectric pump with airfoil baffles, this study conducted comprehensive performance tests and full-flow field simulations of piezoelectric pumps across a range of angles. At a driving voltage of 100 V and with a Clark Y airfoil set at an angle of 0°, the piezoelectric pump reached a peak output flow rate of 200.7 mL/min. An increase in the angle of attack corresponded to a decline in both the maximum output flow rate and the maximum back pressure of the pump. Flow field simulation results demonstrated that an increased airfoil angle of attack led to a gradual increase in entropy production within the piezoelectric pump. Turbulent dissipation and wall entropy production were found to be more pronounced compared to viscous entropy production. High turbulent dissipation was primarily observed at the pump chamber inlet, the trailing edges of the airfoils in both the inlet and outlet pipes, and the outlet bend. As the angle of attack increased, the complexity of the vortex core structures within the flow field escalated as well. Regions with significant wall entropy production were notably concentrated at the outlet bend. Full article
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<p>Assembly diagram of the valveless piezoelectric pump with airfoil baffles.</p>
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<p>Schematic diagram of the working principle.</p>
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<p>Experimental schematic diagram.</p>
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<p>Flow field simulation model.</p>
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<p>Performance curves of the valveless piezoelectric pump with airfoil baffles: (<b>a</b>) flow rate–frequency curve; (<b>b</b>) back pressure–frequency curve.</p>
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<p>Displacement of vibrator and pump flow rate.</p>
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<p>Entropy production curves of the pumps.</p>
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<p>Turbulent entropy production (left) and vortex distribution (right) of the pump at different angles of attack: (<b>a</b>) 0°; (<b>b</b>) 5°; (<b>c</b>) 10°; and (<b>d</b>) 15°.</p>
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<p>Turbulent entropy production (left) and vortex distribution (right) of the pump at different angles of attack: (<b>a</b>) 0°; (<b>b</b>) 5°; (<b>c</b>) 10°; and (<b>d</b>) 15°.</p>
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<p>Proportion of wall entropy production of different components.</p>
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19 pages, 12807 KiB  
Article
Modification of Mechanical Properties of Ti–6Al–4V Using L-PBF for Anatomical Plates
by Soumyabrata Basak, Sang-Hun Lee, Jeong-Rim Lee, Dong-Hyun Kim, Jeong Hun Lee, Myunghwan Byun and Dong-Hyun Kim
Metals 2025, 15(1), 32; https://doi.org/10.3390/met15010032 - 2 Jan 2025
Viewed by 667
Abstract
In this research, as-built Ti–6Al–4V anatomical plates were successfully fabricated using laser powder bed fusion (LPBF). This study thoroughly examines the microstructural evolution and its role in enhancing the mechanical properties of clavicle bone plates under sub-β-transus heat treatment for medical application. Scanning [...] Read more.
In this research, as-built Ti–6Al–4V anatomical plates were successfully fabricated using laser powder bed fusion (LPBF). This study thoroughly examines the microstructural evolution and its role in enhancing the mechanical properties of clavicle bone plates under sub-β-transus heat treatment for medical application. Scanning electron microscope (SEM) images of the as-built specimens reveal a dense formation of a hard α’ hcp martensite structure, which decomposes during annealing at 650 °C and ultimately transforms into an α + β lamellar structure at 950 °C. Additionally, coarse grains resulting from recrystallization and reduced dislocation density were observed through electron backscatter diffraction (EBSD) following heat treatment. Due to these microstructural evolutions, the desired mechanical properties of as-built Ti64 parts for surgical applications were achieved. Heat treatment of the anatomical plates at 950 °C demonstrated an excellent strength–ductility synergy under tensile deformation and the highest energy absorption capability under bending deformation, indicating sufficient durability for medical implantation applications. Full article
(This article belongs to the Section Additive Manufacturing)
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<p>Schematic representation of (<b>a</b>) LPBF process for Ti–6Al–4V (SEM image of raw powder particles at inset); (<b>b</b>) Z-built clavicle bone plate, cuboid, and tensile specimen; (<b>c</b>) a 4-point bending test for clavicle bone plate.</p>
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<p>(<b>a</b>) 3D X-ray computed tomography analysis of pore characteristics within the as-built Ti-–6Al–4V cube specimen; (<b>b</b>) a bar chart showing the volume distribution and count of the interior pores.</p>
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<p>(<b>a</b>) XRD analyses of as-built and heat-treated Ti64 alloy fabricated by LPBF. (<b>b</b>,<b>c</b>) Magnified images from the marked by purple and green dashed line reveal details of peak profiling and the presence of phase constituents. β-phase peak intensity increased at 950 HT and 800 HT thermal annealing conditions.</p>
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<p>Dislocation density estimated by XRD peak analysis for as-built, 650 HT, 800 HT, and 950 HT.</p>
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<p>SEM images through build direction (<span class="html-italic">Z</span> axis) of (<b>a</b>) as-built showing α’ martensitic structures; (<b>b</b>) 650 HT showing decomposition of martensitic α’ and stable α; (<b>c</b>) 800 HT showing α + β phase; and (<b>d</b>) 950 HT showing primary α, α + β lamella, and rod-shaped β. All the conditions present the nanosized β particles.</p>
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<p>Inverse pole Figure (IPF) maps of the (<b>a</b>) as-built, (<b>b</b>) 650 HT, (<b>c</b>) 800 HT, and (<b>d</b>) 950 HT. The prior β grain boundaries are marked with black dashed lines.</p>
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<p>Kernel average misorientation (KAM) maps of the (<b>a</b>) as-built, (<b>b</b>) 650 HT, (<b>c</b>) 800 HT, and (<b>d</b>) 950 HT. The prior β grain boundaries are indicated by white dashed lines.</p>
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<p>Surface hardness maps sequentially for as-built, 650 HT, 800 HT, and 950 HT, respectively. Hardness decreased as the thermal annealing temperature increased to 950 HT condition.</p>
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<p>Engineering stress–strain curves for as-built, 650 HT, 800 HT, and 950 HT specimens.</p>
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<p>SEM fractography and magnified tensile fracture surface images for (<b>a,e</b>) as-built; (<b>b</b>,<b>f</b>) 650 HT; (<b>c</b>,<b>g</b>) 800 HT; (<b>d</b>,<b>h</b>) 950 HT. The fracture mode change from quasi-cleavage failure of as-built to ductile failure of 950 HT specimen.</p>
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<p>Plots of 4-point bending behavior (force–displacement curve) for various annealing treatment conditions: as-built, 650 HT, 800 HT, and 950 HT.</p>
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<p>4-point bending properties of clavicle bone plate under different conditions: (<b>a</b>) maximum load and rigidity; (<b>b</b>) energy absorption capacity in the elastic and plastic regions for as-built and heat-treated specimens.</p>
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<p>(<b>a</b>) Pictographs of the broken specimens after bending tests and corresponding SEM images reveal morphologies of fracture zone for (<b>b</b>–<b>b2</b>) as-built, (<b>c</b>–<b>c2</b>) 800 HT, (<b>d</b>–<b>d2</b>) 950 HT. The yellow and red arrows indicate the tearing edges and shear dimples, relatively.</p>
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