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Metals
Volume 9
Issue 5
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Metals, Volume 9, Issue 5 (May 2019) – 135 articles

Cover Story (view full-size image): Computational fluid dynamics is used to study postcombustion in an electric arc furnace. The furnace is equipped with three virtual lance burners having inlets for the main and secondary oxygen and fuel. The model includes the gas phase inside the burners and the furnace above the molten bath surface. The off-gas analysis is used to estimate the flow rate of CO arising from the bath. The results demonstrate the regions inside the furnace having high concentration of CO capable of being combusted. The model is used to adjust the flow rate of secondary oxygen in order to improve the postcombustion. View this paper.
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13 pages, 6457 KiB  
Article
Stress Corrosion Behavior of AM50Gd Magnesium Alloy in Different Environments
by Miao Yang, Xiaobo Liu, Zhiyi Zhang and Yulai Song
Metals 2019, 9(5), 616; https://doi.org/10.3390/met9050616 - 27 May 2019
Cited by 12 | Viewed by 3566
Abstract
A new type of high strength corrosion-resistant magnesium alloy was prepared by adding 1% rare earth Gd to AM50 and then treated with hot extrusion method. The stress corrosion properties of the new materials in air, pure water, 0.5 mol/L NaCl, and 0.5 [...] Read more.
A new type of high strength corrosion-resistant magnesium alloy was prepared by adding 1% rare earth Gd to AM50 and then treated with hot extrusion method. The stress corrosion properties of the new materials in air, pure water, 0.5 mol/L NaCl, and 0.5 mol/L Na2SO4 solution were studied by the slow strain rate tensile (SSRT) test, in situ open circuit potential test, Tafel curve test, stereomicroscope, SEM, and EDS. The results showed the following. The stress corrosion sensitivity of the material in different environments was Na2SO4 > NaCl > distilled water > air. According to the Tafel curves measured at 0 and 100 MPa, the corrosion voltage decreased little and the corrosion current density increased rapidly under 100 Pa. This was because the film of the corrosion product ruptured to form a large cathode and a small anode, which resulted in a large instantaneous corrosion current. The mechanism of hydrogen embrittlement and anodic dissolution together affected the stress corrosion behavior of the alloy. In distilled water, hydrogen embrittlement played a major role, while in NaCl and Na2SO4 solution, hydrogen embrittlement and anodic dissolution were both affected. The direct reason of the stress corrosion crack (SCC) samples’ failure was the cracks expanding rapidly at the bottom of pit, which was caused by corrosion. Full article
(This article belongs to the Special Issue Fatigue and Fracture of Mg Alloys)
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Figure 1

Figure 1
<p>Diagrammatic sketch of the slow strain rate tensile (SSRT) test sample with its dimension.</p> Full article ">Figure 2
<p>Schematic diagram of electrochemical test method (<b>a</b>) and photo of the device (<b>b</b>).</p> Full article ">Figure 3
<p>SEM morphology of the AM50Gd magnesium alloy. Radial (<b>a</b>) and axial (<b>b</b>) sections.</p> Full article ">Figure 4
<p>The stress vs. elongation plots of AM50Gd magnesium alloy at a strain rate 1.0 × 10<sup>−6</sup> under different environmental conditions.</p> Full article ">Figure 5
<p>Comparison of the SSRT data (<b>a</b>) and SCC susceptibility index (<b>b</b>).</p> Full article ">Figure 6
<p>The corrosion open circuit potential curves during SSRT in different aqueous solutions (<b>a</b>), the corrosion open circuit potential curve with SSRT curve in distilled water (<b>b</b>), the NaCl aqueous solution (<b>c</b>), and the Na<sub>2</sub>SO<sub>4</sub> aqueous solution (<b>d</b>).</p> Full article ">Figure 6 Cont.
<p>The corrosion open circuit potential curves during SSRT in different aqueous solutions (<b>a</b>), the corrosion open circuit potential curve with SSRT curve in distilled water (<b>b</b>), the NaCl aqueous solution (<b>c</b>), and the Na<sub>2</sub>SO<sub>4</sub> aqueous solution (<b>d</b>).</p> Full article ">Figure 7
<p>Tafel curves under 0 and 100 MPa in NaCl (<b>a</b>) and Na<sub>2</sub>SO<sub>4</sub> (<b>b</b>) aqueous solution.</p> Full article ">Figure 8
<p>The surface morphologies of the alloy in air (<b>a</b>), distilled water (<b>b</b>), NaCl (<b>c</b>), and Na<sub>2</sub>SO<sub>4</sub> (<b>d</b>) aqueous solutions.</p> Full article ">Figure 8 Cont.
<p>The surface morphologies of the alloy in air (<b>a</b>), distilled water (<b>b</b>), NaCl (<b>c</b>), and Na<sub>2</sub>SO<sub>4</sub> (<b>d</b>) aqueous solutions.</p> Full article ">Figure 9
<p>Fractographies of the SSRT test sample in air: (<b>a</b>) ×23, (<b>b</b>) ×500, (<b>c</b>) ×3000, and (<b>d</b>) EDS results of Point A.</p> Full article ">Figure 10
<p>Fractographies of the SSRT test sample in distilled water. (<b>a</b>) Overall, (<b>b</b>) edge.</p> Full article ">Figure 11
<p>Fractographies of the SSRT test sample in 0.5 mol/L NaCl aqueous solution (<b>a</b>). (<b>b</b>) Overall, (<b>c</b>) Corrosion field, and (<b>d</b>) Field without corrosion.</p> Full article ">Figure 12
<p>Fractographies of the SSRT test sample in 0.5 mol/L Na<sub>2</sub>SO<sub>4</sub> aqueous solution. (<b>a</b>) Overall, (<b>b</b>) Field in the middle.</p> Full article ">
17 pages, 4068 KiB  
Article
A Fast Metals Recovery Method for the Synthesis of Lithium Nickel Cobalt Aluminum Oxide Material from Cathode Waste
by Soraya Ulfa Muzayanha, Cornelius Satria Yudha, Adrian Nur, Hendri Widiyandari, Hery Haerudin, Hanida Nilasary, Ferry Fathoni and Agus Purwanto
Metals 2019, 9(5), 615; https://doi.org/10.3390/met9050615 - 27 May 2019
Cited by 32 | Viewed by 6788
Abstract
An approach for a fast recycling process for Lithium Nickel Cobalt Aluminum Oxide (NCA) cathode scrap material without the presence of a reducing agent was proposed. The combination of metal leaching using strong acids (HCl, H2SO4, HNO3) [...] Read more.
An approach for a fast recycling process for Lithium Nickel Cobalt Aluminum Oxide (NCA) cathode scrap material without the presence of a reducing agent was proposed. The combination of metal leaching using strong acids (HCl, H2SO4, HNO3) and mixed metal hydroxide co-precipitation followed by heat treatment was investigated to resynthesize NCA. The most efficient leaching with a high solid loading rate (100 g/L) was obtained using HCl, resulting in Ni, Co, and Al leaching efficiencies of 99.8%, 95.6%, and 99.5%, respectively. The recycled NCA (RNCA) was successfully synthesized and in good agreement with JCPDS Card #87-1562. The highly crystalline RNCA presents the highest specific discharge capacity of a full cell (RNCA vs. Graphite) of 124.2 mAh/g with capacity retention of 96% after 40 cycles. This result is comparable with commercial NCA. Overall, this approach is faster than that in the previous study, resulting in more efficient and facile treatment of the recycling process for NCA waste and providing 35 times faster processing. Full article
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Graphical abstract

Graphical abstract
Full article ">Figure 1
<p>Flow chart of the recycled Lithium Nickel Cobalt Aluminum Oxide (RNCA) recycling process.</p> Full article ">Figure 2
<p>XRD pattern of NCA powder before heat treatment (B-HT) and after heat treatment at 800 °C for 2 or 4 h (P-01 and P-02 samples).</p> Full article ">Figure 3
<p>Specific Discharge Capacity (mAh/g) of the B-HT and P-02 Samples.</p> Full article ">Figure 4
<p>The effect of various acids on the leaching efficiency of NCA scrap (4M, 80 °C, 1 h).</p> Full article ">Figure 5
<p>Comparison of leaching between P-02, B-HT, and commercial NCA (4M HCl, 80 °C, 1 h).</p> Full article ">Figure 6
<p>Leaching behavior using various acids during 60 min of processing (4M, 80 °C): (<b>a</b>) Nickel; (<b>b</b>) Cobalt; (<b>c</b>) Aluminum.</p> Full article ">Figure 7
<p>FT-IR analysis of the –OH precursor from the leachates of H<sub>2</sub>SO<sub>4</sub>, HNO<sub>3</sub>, and HCl; denoted as P-SA, P-NA, P-HA (M: metal).</p> Full article ">Figure 8
<p>XRD patterns of RNCA using different leaching agents.</p> Full article ">Figure 9
<p>Crystallite sizes of RNCA Samples.</p> Full article ">Figure 10
<p>SEM images of RNCA material: (<b>a</b>,<b>b</b>) RNCA-HA; (<b>c</b>,<b>d</b>) RNCA-NA; (<b>e</b>,<b>f</b>) RNCA-SA.</p> Full article ">Figure 11
<p>Specific Discharge Capacity of all RNCA samples at 0.02 C.</p> Full article ">Figure 12
<p>Performance test of RNCA-HA: (<b>a</b>) Ratability; (<b>b</b>) Cyclability at 100 mA/g.</p> Full article ">
10 pages, 2256 KiB  
Article
Modelling the Sintering Neck Growth Process of Metal Fibers under the Surface Diffusion Mechanism Using the Lattice Boltzmann Method
by Houping Dai, Dongdong Chen and Zhoushun Zheng
Metals 2019, 9(5), 614; https://doi.org/10.3390/met9050614 - 27 May 2019
Cited by 8 | Viewed by 3540
Abstract
In this paper, the sintering neck growth process of metal fibers under the surface diffusion mechanism is simulated by using the Lattice Boltzmann method (LBM). The surface diffusion model is developed considering the geometrical characteristic of metal fibers. Then, the LBM scheme is [...] Read more.
In this paper, the sintering neck growth process of metal fibers under the surface diffusion mechanism is simulated by using the Lattice Boltzmann method (LBM). The surface diffusion model is developed considering the geometrical characteristic of metal fibers. Then, the LBM scheme is constructed for solving the developed surface diffusion model. The sintering neck growth process of two metal fibers with different fiber angles is simulated by LBM. The simulated morphologies of sintering metal fibers well agree with ones obtained by experiments. Moreover, the numerical simulation results show that the sintering neck radius of two metal fibers is increased with the increase of fiber angle, which implies that the initial geometrical characteristic plays an important role in the sintering neck formation and growth of metal fibers. Full article
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Figure 1

Figure 1
<p>Scanning electron microscopy image of 316L stainless steel fiber felts with fiber diameters of 8 μm sintered at 1200 °C for 2 h.</p> Full article ">Figure 2
<p>Diagram of two metal fibers in polar coordinates (α and β represent the fiber angle and polar angle, respectively).</p> Full article ">Figure 3
<p>Section of two metal fibers in rectangular coordinates (<span class="html-italic">Y</span> is the length of the sintering neck; <span class="html-italic">O</span><sub>1</sub> and <span class="html-italic">O</span><sub>2</sub> represent the centers of two ovals; <span class="html-italic">R</span> is the radius of oval).</p> Full article ">Figure 4
<p>One-dimensional three-velocity (D1Q3) model (<span class="html-italic">C</span><sub>1</sub>, <span class="html-italic">C</span><sub>2</sub> and <span class="html-italic">C</span><sub>3</sub> represent three adjacent points on the axis).</p> Full article ">Figure 5
<p>Sintering neck growth of two metal fibers with different fiber angles: (<b>a</b>) α = 0; (<b>b</b>) α = π/6; (<b>c</b>) α = π/4; (<b>d</b>) α = π/3; (<b>e</b>) α = π/2. <span class="html-italic">T</span> = time.</p> Full article ">Figure 5 Cont.
<p>Sintering neck growth of two metal fibers with different fiber angles: (<b>a</b>) α = 0; (<b>b</b>) α = π/6; (<b>c</b>) α = π/4; (<b>d</b>) α = π/3; (<b>e</b>) α = π/2. <span class="html-italic">T</span> = time.</p> Full article ">Figure 6
<p>Experimental results of two 8 μm 316L stainless steel fibers with angles of 0~π/2 sintered at 1200 °C for 2 h: (<b>a</b>) section morphology of sintering metal fibers; (<b>b</b>–<b>f</b>) high magnification morphologies.</p> Full article ">Figure 6 Cont.
<p>Experimental results of two 8 μm 316L stainless steel fibers with angles of 0~π/2 sintered at 1200 °C for 2 h: (<b>a</b>) section morphology of sintering metal fibers; (<b>b</b>–<b>f</b>) high magnification morphologies.</p> Full article ">Figure 7
<p>Sintering neck radii of two metal fibers with different fiber angles.</p> Full article ">
11 pages, 1777 KiB  
Article
Nanoindentation Investigation on the Size-Dependent Creep Behavior in a Zr-Cu-Ag-Al Bulk Metallic Glass
by Z. Y. Ding, Y. X. Song, Y. Ma, X. W. Huang and T. H. Zhang
Metals 2019, 9(5), 613; https://doi.org/10.3390/met9050613 - 27 May 2019
Cited by 24 | Viewed by 3269
Abstract
Nanoindentation technology has been widely adopted to study creep behavior in small regions. However, nanoindentation creep behavior of metallic glass is still not well understood. In the present work, we investigated nanoindentation size effects on creep deformation in a Zr-based bulk metallic glass [...] Read more.
Nanoindentation technology has been widely adopted to study creep behavior in small regions. However, nanoindentation creep behavior of metallic glass is still not well understood. In the present work, we investigated nanoindentation size effects on creep deformation in a Zr-based bulk metallic glass at room temperature. The total creep strain and strain rate of steady-state creep were gradually decreased with increasing holding depth under a Berkovich indenter, indicating a length-scale-dependent creep resistance. For a spherical indenter, creep deformations were insignificant in elastic regions and then greatly enhanced by increasing holding strain in plastic regions. Strain rate sensitivities (SRS) decreased with increasing holding depth and holding strain at first, and then stabilized as holding depth was beyond about 500 nm for both indenters. SRS values were 0.4–0.5 in elastic regions, in which atomic diffusion and free volume migration could be the creep mechanism. On the other hand, evolution of the shear transformation zone was suggested as a creep mechanism in plastic regions, and the corresponding SRS values were in the range of 0.05 to 0.3. Full article
(This article belongs to the Special Issue Creep and High Temperature Deformation of Metals and Alloys)
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Figure 1

Figure 1
<p>(<b>a</b>) The typical creep <span class="html-italic">P-h</span> curves at various holding depths under Berkovich indenter. <span class="html-italic">P-h</span> curves at shallow depths were enlarged in the inset. (<b>b</b>) Creep displacements at various holding loads were plotted with holding time.</p> Full article ">Figure 2
<p>The total creep displacements in the end of holding stage were plotted with holding depth for both indenters.</p> Full article ">Figure 3
<p>(<b>a</b>) The typical <span class="html-italic">P-h</span> curve under 200 mN spherical indenter with 5 s holding. The distribution range of the critical load at the first pop-in event is shown in the inset, which was plotted with measurement. (<b>b</b>) The corresponding holding strain for each holding load was calculated for spherical nanoindentation and was plotted with holding depth. The creep tests could be divided as elastic holding and plastic holding.</p> Full article ">Figure 4
<p>The total creep strain in the end of holding stage and strain rate of steady-state creep were estimated for (<b>a</b>) Berkovich and (<b>b</b>) spherical indenters, which were plotted with holding depth and holding strain.</p> Full article ">Figure 5
<p>(<b>a</b>) The creep displacements versus holding time, which is perfectly fitted by an empirical law; (<b>b</b>) the creep strain rate versus holding time; (<b>c</b>) the hardness versus holding time; (<b>d</b>) the log–log correlation between hardness and strain rate obtained from the creep, strain rate sensitivity was estimated by linear fitting of the steady-state part.</p> Full article ">Figure 6
<p>Strain rate sensitivities were plotted with holding depth for (<b>a</b>) a Berkovich indenter and plotted with holding strain for (<b>b</b>) a spherical indenter.</p> Full article ">
11 pages, 4745 KiB  
Article
Numerical Analysis of the Effects of Pulsed Laser Spot Heating Parameters on Brazing of Diamond Tools
by Yangguang Wang, Guoqin Huang, Yanfang Su, Meiqin Zhang, Zhen Tong and Changcai Cui
Metals 2019, 9(5), 612; https://doi.org/10.3390/met9050612 - 27 May 2019
Cited by 10 | Viewed by 3286
Abstract
A 3D finite element (FE) model is built to numerically analyze heating parameters on temperature during brazing diamond grains by the pulsed laser spot heating. A pulsed Nd:YAG laser is used for experimental validation. The results show that during laser heating, the temperature [...] Read more.
A 3D finite element (FE) model is built to numerically analyze heating parameters on temperature during brazing diamond grains by the pulsed laser spot heating. A pulsed Nd:YAG laser is used for experimental validation. The results show that during laser heating, the temperature varies periodically because of the pulsed heat flux. Four key thermal indices, the maximum temperature Tmax, the minimum temperature Tmin, the average temperature Tav and the temperature fluctuation amplitude ΔT are addressed. The primary factor affecting Tmax, ΔT and Tav is the pulse power and on Tmin is the pulse frequency. The secondary effect factor on Tmax, Tav and ΔT is the pulse width and on Tmin is the pulse power. For engineering practice, the order of designing heating parameters is recommended as: pulse power, second frequency and last width. Full article
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Figure 1

Figure 1
<p>Experimental setup: (<b>a</b>) laser heating device, (<b>b</b>) temperature monitor, and (<b>c</b>) laser heating zone in the chamber.</p> Full article ">Figure 2
<p>Diamond brazing by spot heating with pulsed laser.</p> Full article ">Figure 3
<p>Waveform of the output power of a pulsed laser beam.</p> Full article ">Figure 4
<p>Finite element model of laser brazing. (<b>a</b>) The cross-sectional illustration of the diamond/filler alloy/substrate assembly; (<b>b</b>) the meshed model.</p> Full article ">Figure 5
<p>Comparison of experimental and simulation temperature curves: (<b>a</b>) calibration and (<b>b</b>) validation.</p> Full article ">Figure 6
<p>Simulated temperature curve and distributions during pulsed laser spot heating: (<b>a</b>) temperature curve, (<b>b</b>) <span class="html-italic">t</span> = 0.966667 s, (<b>c</b>) <span class="html-italic">t</span> = 0.969667 s, and (<b>d</b>) <span class="html-italic">t</span> = 3 s.</p> Full article ">Figure 7
<p>Stable heating stage in <a href="#metals-09-00612-f006" class="html-fig">Figure 6</a>: (<b>a</b>) temperature vs. heating time and (<b>b</b>) temperature evolution within a heat pulse duration.</p> Full article ">Figure 8
<p>Temperature curve induced within a heat pulse duration.</p> Full article ">Figure 9
<p>Diamond grains brazed by: (<b>a</b>) Case D, (<b>b</b>) Case G, and (<b>c</b>) Case E.</p> Full article ">Figure 10
<p>Comparisons between the simulated and the calculated results: (<b>a</b>) <span class="html-italic">T</span><sub>max</sub>, (<b>b</b>) <span class="html-italic">T</span><sub>min</sub>, (<b>c</b>) Δ<span class="html-italic">T</span> and (<b>d</b>) <span class="html-italic">T</span><sub>av</sub>.</p> Full article ">
30 pages, 5984 KiB  
Article
An Evolutionary Yield Function Model Based on Plastic Work and Non-Associated Flow Rule
by Taejoon Park, Fadi Abu-Farha and Farhang Pourboghrat
Metals 2019, 9(5), 611; https://doi.org/10.3390/met9050611 - 25 May 2019
Cited by 11 | Viewed by 4511
Abstract
A constitutive law was developed based on the evolutionary yield function to account for the evolution of anisotropy induced by the plastic deformation. For the effective description of anisotropy, the yield stress function and plastic potential were separately defined based on the non-associated [...] Read more.
A constitutive law was developed based on the evolutionary yield function to account for the evolution of anisotropy induced by the plastic deformation. For the effective description of anisotropy, the yield stress function and plastic potential were separately defined based on the non-associated flow rule. In particular, for the description of the equivalent status, the accumulated plastic work was employed as an alternative to the accumulated plastic strain. Numerical formulations based on the plastic work were also derived in case the hardening rule, as well as the evolution of the plastic potential and yield stress function, were defined in terms of the plastic work. The developed constitutive law was characterized using the mechanical properties of the multi-phase BAO QP980 steel and niobium sheets at room temperature. From the uniaxial tension tests and the balanced biaxial tension test, separate sets of anisotropic coefficients for each of the plastic potential and yield stress functions were obtained as a function of the plastic work. By comparing with non-evolving yield functions, the importance of the developed constitutive law to properly describe the evolution of the plastic potential and yield function were validated. Full article
(This article belongs to the Special Issue Constitutive Modelling for Metals)
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Figure 1

Figure 1
<p>Comparison of the calculated yield surfaces based on crystal plasticity simulations and isotropic version of the Yld2000-2D function.</p> Full article ">Figure 2
<p>Comparison of the calculated strain-rate potentials based on crystal plasticity simulations and isotropic version of the Yld2000-2D function.</p> Full article ">Figure 3
<p>Measured <span class="html-italic">r</span>-values for the QP980 steel averaged from 5 experiments for each tensile direction: (<b>a</b>) evolutions with respect to plastic work; (<b>b</b>) averaged <span class="html-italic">r</span>-value distributions.</p> Full article ">Figure 3 Cont.
<p>Measured <span class="html-italic">r</span>-values for the QP980 steel averaged from 5 experiments for each tensile direction: (<b>a</b>) evolutions with respect to plastic work; (<b>b</b>) averaged <span class="html-italic">r</span>-value distributions.</p> Full article ">Figure 4
<p>Evolution of the <span class="html-italic">r</span>-values for the niobium sheet averaged from 3 experiments for each tensile direction.</p> Full article ">Figure 5
<p>Measured (<b>a</b>) yield stress and (<b>b</b>) normalized yield stress evolutions for the Q&amp;P 980 steel with respect to the plastic work.</p> Full article ">Figure 6
<p>Measured (<b>a</b>) yield stress and (<b>b</b>) normalized yield stress evolutions for the niobium sheet with respect to the plastic work.</p> Full article ">Figure 6 Cont.
<p>Measured (<b>a</b>) yield stress and (<b>b</b>) normalized yield stress evolutions for the niobium sheet with respect to the plastic work.</p> Full article ">Figure 7
<p>Comparison of the averaged <span class="html-italic">r</span>-values for the Q&amp;P 980 steel and the calculated distribution based on the plastic potential function.</p> Full article ">Figure 8
<p>Evolutions of the parameters of the yield function (Yld2000-2d function) for the Q&amp;P 980 steel with respect to the plastic work.</p> Full article ">Figure 9
<p>Evolutions of the parameters of the plastic potential function (Yld2000-2d function) for the niobium sheet with respect to the plastic work.</p> Full article ">Figure 10
<p>Evolutions of the parameters of yield function (Yld2000-2d function) for the niobium sheet with respect to the plastic work.</p> Full article ">Figure 11
<p>Comparison of the measured and calculated hardening behavior of (<b>a</b>) the QP980 steel and (<b>b</b>) niobium sheet with respect to the plastic work.</p> Full article ">Figure 12
<p>Comparison of the normalized contours of the yield functions (at 0 MPa and 80 MPa of the plastic work) and plastic potential function for the Q&amp;P 980 steel.</p> Full article ">Figure 13
<p>Comparison of the normalized contours of (<b>a</b>) the yield functions and (<b>b</b>) plastic potential function for the niobium sheet.</p> Full article ">Figure 14
<p>The tensor product of the gradients of the yield function and the plastic potential for the elastic stiffness tensor for the niobium sheet (<b>a</b>) at 0 MPa plastic work and (<b>b</b>) at 40 MPa plastic work.</p> Full article ">Figure 15
<p>Comparison of the measured and simulated true stress–strain curves for (<b>a</b>) 0, (<b>b</b>) 45 and (<b>c</b>) 90 degrees off tensile directions for the Q&amp;P 980 steel.</p> Full article ">Figure 15 Cont.
<p>Comparison of the measured and simulated true stress–strain curves for (<b>a</b>) 0, (<b>b</b>) 45 and (<b>c</b>) 90 degrees off tensile directions for the Q&amp;P 980 steel.</p> Full article ">Figure 16
<p>Comparison of the measured and simulated true stress–strain curves for (<b>a</b>) 0, (<b>b</b>) 45 and (<b>c</b>) 90 degrees off tensile directions for the niobium sheet.</p> Full article ">Figure 16 Cont.
<p>Comparison of the measured and simulated true stress–strain curves for (<b>a</b>) 0, (<b>b</b>) 45 and (<b>c</b>) 90 degrees off tensile directions for the niobium sheet.</p> Full article ">Figure 17
<p>Comparison of the measured and simulated ultimate tensile strength (UTS) distributions for (<b>a</b>) the Q&amp;P 980 steel and (<b>b</b>) niobium sheet.</p> Full article ">Figure 18
<p>A schematic view of the initial circular blank and drawn cup: (<b>a</b>) top view of the initial blank and material element in the flange area; (<b>b</b>) cross-section view of the initial blank and the final cup.</p> Full article ">Figure 18 Cont.
<p>A schematic view of the initial circular blank and drawn cup: (<b>a</b>) top view of the initial blank and material element in the flange area; (<b>b</b>) cross-section view of the initial blank and the final cup.</p> Full article ">Figure 19
<p>Comparison of the measured and calculated (<b>a</b>) <span class="html-italic">r</span>-values and (<b>b</b>) normalized yield stresses based on associated flow rule (Yld2000-2D and CPB06ex3 yield functions) for the niobium sheet.</p> Full article ">Figure 20
<p>Comparison of the predicted cup height profiles: (<b>a</b>) based on the non-associated flow rule with Yld2000-2D functions; (<b>b</b>) based on the associated flow rule with Yld2000-2D and CPB06ex3 functions at 0 MPa of the accumulated plastic work; (<b>c</b>) based on the associated flow rule with Yld2000-2D and CPB06ex3 functions at 40 MPa of the accumulated plastic work.</p> Full article ">Figure 20 Cont.
<p>Comparison of the predicted cup height profiles: (<b>a</b>) based on the non-associated flow rule with Yld2000-2D functions; (<b>b</b>) based on the associated flow rule with Yld2000-2D and CPB06ex3 functions at 0 MPa of the accumulated plastic work; (<b>c</b>) based on the associated flow rule with Yld2000-2D and CPB06ex3 functions at 40 MPa of the accumulated plastic work.</p> Full article ">
22 pages, 8703 KiB  
Article
Main Issues in Quality of Friction Stir Welding Joints of Aluminum Alloy and Steel Sheets
by Mian Wasif Safeen and Pasquale Russo Spena
Metals 2019, 9(5), 610; https://doi.org/10.3390/met9050610 - 25 May 2019
Cited by 51 | Viewed by 7211
Abstract
Joining of aluminum alloys through friction stir welding (FSW) is effectively employed in several industries (e.g., aeronautics and aerospace) since it guarantees proper weld strength as compared to other joining technologies. Contrarily, dissimilar FSW of aluminum alloys and steels often poses important issues [...] Read more.
Joining of aluminum alloys through friction stir welding (FSW) is effectively employed in several industries (e.g., aeronautics and aerospace) since it guarantees proper weld strength as compared to other joining technologies. Contrarily, dissimilar FSW of aluminum alloys and steels often poses important issues in the selection of welding parameters due to the difficulty to join different materials. Improper welding parameters give rise to the formation of intermetallic compounds, and internal and external defects (e.g., tunnel formation, voids, surface grooves, and flash). Intermetallic compounds are brittle precipitates of Al/Fe, which chiefly initiate crack nucleation, whereas internal and external defects mainly act as stress concentration factors. All these features significantly reduce joint strength under static and dynamic loading conditions. With reference to the literature, the influence of main welding parameters (rotational speed, welding speed, tool geometry, tilt angle, offset distance, and plunge depth) on the formation of intermetallic compounds and defects in FSW of aluminum alloys and steels is discussed here. Possible countermeasures to avoid or limit the above-mentioned issues are also summarily reported. Full article
(This article belongs to the Special Issue Recent Achievements in Rotary, Linear and Friction Stir Welding of Metals Alloys)
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Figure 1

Figure 1
<p>Most common friction stir welding (FSW) material configurations: (<b>a</b>) butt welding (<b>b</b>) lap welding (Reproduced from [<a href="#B3-metals-09-00610" class="html-bibr">3</a>], with permission from Taylor &amp; Francis, 2017).</p> Full article ">Figure 2
<p>Example of dissimilar FSW with the placement of the softer material (Al 5083) at the (<b>a</b>) retreating and (<b>b</b>) advancing side (Reproduced from [<a href="#B4-metals-09-00610" class="html-bibr">4</a>], with permission from Elsevier, 2016).</p> Full article ">Figure 3
<p>Iron-Aluminum phase diagram (Reproduced from [<a href="#B39-metals-09-00610" class="html-bibr">39</a>], with permission from Springer Nature, 2003).</p> Full article ">Figure 4
<p>Detection of intermetallic compounds (IMCs) by X-ray diffraction at different temperatures (Reproduced from [<a href="#B37-metals-09-00610" class="html-bibr">37</a>], with permission from Elsevier, 2009).</p> Full article ">Figure 5
<p>Effect of welding speed on joint temperature in FSW of AA6061 and TRIP steel (Reproduced from [<a href="#B41-metals-09-00610" class="html-bibr">41</a>], with permission from Elsevier, 2018). Arrows in the graph identify the temperature profiles as a function of welding speed.</p> Full article ">Figure 6
<p>Effect of (<b>a</b>) welding speed and (<b>b</b>) rotational speed on IMC thickness in FSW between AA6082-T6 and Q235A steel (Reproduced from [<a href="#B13-metals-09-00610" class="html-bibr">13</a>], with permission from the authors).</p> Full article ">Figure 7
<p>Effect of rotational speed on tensile strength in FSW of commercially pure aluminum and AISI304 stainless steel (Reproduced from [<a href="#B31-metals-09-00610" class="html-bibr">31</a>], with permission from Springer Nature, 2018).</p> Full article ">Figure 8
<p>Thickness of IMCs at the top and bottom of AA5005/St-52 steel FSW welded joints at different welding and rotational speeds (Reproduced from [<a href="#B20-metals-09-00610" class="html-bibr">20</a>], with permission from Springer Nature, 2018).</p> Full article ">Figure 9
<p>Relationship between joint strength and heat input factor in FSW of AA5186 and mild steel (Reproduced from [<a href="#B42-metals-09-00610" class="html-bibr">42</a>], with permission from Elsevier, 2013).</p> Full article ">Figure 10
<p>Dimensions of tapered and cylindrical pin profiles used in FSW of AA5052 and HSLA steel (Reproduced from [<a href="#B22-metals-09-00610" class="html-bibr">22</a>], with permission from Elsevier, 2015).</p> Full article ">Figure 11
<p>IMCs found in FSW of AA5052 and HSLA steel in the upper and bottom region of joints (Image adapted from results of [<a href="#B22-metals-09-00610" class="html-bibr">22</a>]).</p> Full article ">Figure 12
<p>Effect of plunge depth on temperature in FSW of AA5083 and SS400 steel (Image adapted from results of [<a href="#B21-metals-09-00610" class="html-bibr">21</a>]).</p> Full article ">Figure 13
<p>Effect of pin depth on shear load and IMC thickness in FSW of AA5083 and SS400 steel (Reproduced from [<a href="#B21-metals-09-00610" class="html-bibr">21</a>], with permission from The Japan Institute of Metals and Materials, 2019).</p> Full article ">Figure 14
<p>Weld temperature for different tool tilt angles during FSW of AA1100 aluminum alloy and A441 AISI steel (Reproduced from [<a href="#B52-metals-09-00610" class="html-bibr">52</a>], with permission from SAGE UK, 2019).</p> Full article ">Figure 15
<p>Effect of tool tilt angle on shear strength and IMC thickness during FSW of AA5083 and SS400 steel (Reproduced from [<a href="#B24-metals-09-00610" class="html-bibr">24</a>], with permission from The Japan Institute of Metals and Materials, 2019).</p> Full article ">Figure 16
<p>Tunnel defect due to excessive heat input (700 rpm, 12 mm/min) during FSW of AA 3003-H18 and mild steel (st-52) (Reproduced from [<a href="#B14-metals-09-00610" class="html-bibr">14</a>], with permission from Elsevier, 2013).</p> Full article ">Figure 17
<p>Tunnel defect due to insufficient heat input (355 rpm, 40 mm/min) in FSW of AA5186 and mild steel st-52 (Reproduced from [<a href="#B42-metals-09-00610" class="html-bibr">42</a>], with permission from Elsevier, 2013).</p> Full article ">Figure 18
<p>Voids in FSW weld joint of AA5083 and st-12 steel (Reproduced from [<a href="#B15-metals-09-00610" class="html-bibr">15</a>], with permission from Taylor &amp; Francis, 2012).</p> Full article ">Figure 19
<p>Tensile strength of defective and sound FSW joints of AA 6082-T6 and DP-600 steel (Reproduced from [<a href="#B54-metals-09-00610" class="html-bibr">54</a>], with permission from Springer Nature, 2018).</p> Full article ">Figure 20
<p>Flash and large surface and in FSW of AA6061 and 430 stainless steel (Reproduced from [<a href="#B46-metals-09-00610" class="html-bibr">46</a>], with permission from SAGE UK, 2019).</p> Full article ">Figure 21
<p>Microcracks at the Al/steel interface in FSW joints of AA5083-H321 and 316L stainless steel (Reproduced from [<a href="#B16-metals-09-00610" class="html-bibr">16</a>] with permission from Springer Nature, 2016).</p> Full article ">Figure 22
<p>Effect of rotational speed on tensile strength in FSW of AA5083-H321 and 316L stainless steel (Reproduced from [<a href="#B16-metals-09-00610" class="html-bibr">16</a>], with permission from Springer Nature, 2016).</p> Full article ">Figure 23
<p>Process window for FSW (Reproduced from [<a href="#B56-metals-09-00610" class="html-bibr">56</a>], with permission from the authors).</p> Full article ">Figure 24
<p>Zinc foil sandwiched between aluminum and steel (Reproduced from [<a href="#B57-metals-09-00610" class="html-bibr">57</a>], with permission from Elsevier, 2016).</p> Full article ">Figure 25
<p>Comparison of mechanical strength of FSW joints obtained during welding of AA5754 with uncoated and zinc-coated DP1000 steel (Reproduced from [<a href="#B58-metals-09-00610" class="html-bibr">58</a>], with permission from Taylor &amp; Francis, 2017).</p> Full article ">
23 pages, 637 KiB  
Article
Radar Detection-Based Modeling in a Blast Furnace: A Prediction Model of Burden Surface Descent Speed
by Jiuzhou Tian, Akira Tanaka, Qingwen Hou and Xianzhong Chen
Metals 2019, 9(5), 609; https://doi.org/10.3390/met9050609 - 25 May 2019
Cited by 6 | Viewed by 3113
Abstract
The distribution of burden layers is a vital factor that affects the production of a blast furnace. Radars are advanced instruments that can provide the detection results of the burden surface shape inside a blast furnace in real time. To better estimate the [...] Read more.
The distribution of burden layers is a vital factor that affects the production of a blast furnace. Radars are advanced instruments that can provide the detection results of the burden surface shape inside a blast furnace in real time. To better estimate the burden layer thicknesses through improving the prediction accuracy of the burden descent during charging periods, an innovative data-driven model for predicting the distribution of the burden surface descent speed is proposed. The data adopted were from the detection results of an operating blast furnace, collected using a mechanical swing radar system. Under a kinematic continuum modeling mechanism, the proposed model adopts a linear combination of Gaussian radial basis functions to approximate the equivalent field of burden descent speed along the burden surface radius. A proof of the existence and uniqueness of the prediction solution is given to guarantee that the predicted radial profile of the burden surface can always be calculated numerically. Compared with the plain data-driven descriptive model, the proposed model has the ability to better characterize the variability in the radial distribution of burden descent speed. In addition, the proposed model provides prediction results of higher accuracy for both the future surface shape and descent speed distribution. Full article
(This article belongs to the Special Issue Mathematical Modeling and Simulation in Ironmaking and Steelmaking)
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<p>Radar placement on top of the BF.</p> Full article ">Figure 2
<p>Mechanical servo system of radar.</p> Full article ">Figure 3
<p>The variation in BSRDs within half an hour: (<b>a</b>–<b>f</b>) detection results obtained at different time.</p> Full article ">Figure 4
<p>Kinematic modeling mechanism diagram.</p> Full article ">Figure 5
<p>Flow of the proposed prediction model.</p> Full article ">Figure 6
<p>Solution domain of the descent speed along the burden radius.</p> Full article ">Figure 7
<p>EBDFs from two data segments. (<b>a</b>) EBDF corresponds to Case 1; (<b>b</b>) EBDF corresponds to Case 2.</p> Full article ">Figure 8
<p>The relationship between <span class="html-italic">m</span> and the approximation RMSE.</p> Full article ">Figure 9
<p>Comparison of prediction accuracy of BSRP. (<b>a</b>) Results for Case 1; and (<b>b</b>) results for Case 2.</p> Full article ">Figure 10
<p>Comparison of BSRD prediction accuracy. (<b>a</b>) Results for Case 1; (<b>b</b>) results for Case 2; and (<b>c</b>) results presented on equal scaled axes. The upper subgraph corresponds to Case 1, and the lower one corresponds to Case 2.</p> Full article ">Figure 11
<p>Model comparison of the variation in prediction performance with time.</p> Full article ">
9 pages, 2926 KiB  
Article
Wire and Arc Additive Manufacturing of Aluminum Components
by Markus Köhler, Sierk Fiebig, Jonas Hensel and Klaus Dilger
Metals 2019, 9(5), 608; https://doi.org/10.3390/met9050608 - 24 May 2019
Cited by 112 | Viewed by 11664
Abstract
An increasing demand for flexibility and product integration, combined with reduced product development cycles, leads to continuous development of new manufacturing technologies such as additive manufacturing. Wire and arc additive manufacturing (WAAM) provides promising technology for the near net-shape production of large structures [...] Read more.
An increasing demand for flexibility and product integration, combined with reduced product development cycles, leads to continuous development of new manufacturing technologies such as additive manufacturing. Wire and arc additive manufacturing (WAAM) provides promising technology for the near net-shape production of large structures with complex geometry, using cost efficient production resources such as arc welding technology and wire materials. Compared to powder-based additive manufacturing processes, WAAM offers high deposition rates as well as enhanced material utilization. Because of the layer-by-layer built up approach, process conditions such as energy input, arc characteristics, and material composition result in a different processability during the additive manufacturing process. This experimental study aims to describe the effects of the welding process on buildup accuracy and material properties during wire arc additive manufacturing of aluminum structures. Following a process development using pulse cold metal transfer (CMT-P), linear wall samples were manufactured with variations of the filler metal. The samples were analyzed in terms of surface finishing, hardness, and residual stress. Furthermore, mechanical properties were determined in different building directions. Full article
(This article belongs to the Special Issue Arc-based Additive Manufacturing)
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<p>Experimental setup for robot-guided wire and arc additive manufacturing (WAAM) (<b>a</b>) and a schematic process sequence for sample manufacturing (<b>b</b>).</p> Full article ">Figure 2
<p>Dimensions of tensile sample in mm, <span class="html-italic">R</span><sub>z</sub> in µm (<b>a</b>) and schematic representation of sample extraction (<b>b</b>).</p> Full article ">Figure 3
<p>Manufactured samples: (<b>a</b>) Al-4047 and (<b>b</b>) Al-5356.</p> Full article ">Figure 4
<p>Highspeed images during buildup of Al-4047 (<b>a</b>) and Al-5356 (<b>b</b>).</p> Full article ">Figure 5
<p>Cross-section micrographs of WAAM samples (<b>a</b>) Al-4047, (<b>b</b>) Al-5356, and (<b>c</b>) hardness distribution depending on buildup height.</p> Full article ">Figure 6
<p>Effect of process adjustments on droplet transition and waviness of the structure: (<b>a</b>) negative arc and pulse corrections, (<b>b</b>) neutral settings, and (<b>c</b>) positive arc and pulse corrections.</p> Full article ">Figure 7
<p>Tensile properties of processed Al-5356 in (<b>a</b>) horizontal and (<b>b</b>) vertical directions.</p> Full article ">Figure 8
<p>Scanning electron microscope microstructure of the fracture sample after the tensile test.</p> Full article ">
18 pages, 4494 KiB  
Article
Effects of Zr Addition on Thermodynamic and Kinetic Properties of Liquid Mg-6Zn-xZr Alloys
by Ye Yuan, Yuan Huang and Qiang Wei
Metals 2019, 9(5), 607; https://doi.org/10.3390/met9050607 - 24 May 2019
Cited by 8 | Viewed by 3543
Abstract
Mg-6Zn-xZr (ZK60) alloys are precipitation strengthened by Mg-Zn intermetallics. Therefore, it is important to investigate the thermodynamic and kinetic effects of Zr addition on formations of these reinforcement phases (Mg7Zn3, MgZn2, and MgZn) in Mg-6Zn- [...] Read more.
Mg-6Zn-xZr (ZK60) alloys are precipitation strengthened by Mg-Zn intermetallics. Therefore, it is important to investigate the thermodynamic and kinetic effects of Zr addition on formations of these reinforcement phases (Mg7Zn3, MgZn2, and MgZn) in Mg-6Zn-xZr melts. Because it is difficult to gain thermodynamic and kinetic data in melts by experiment, obtaining these data points still depends on a theoretical calculation. Based on the Miedema formation enthalpy model and the Chou model, the thermodynamic properties (the mixing enthalpies, the Gibbs free energies, and the component activities) of Mg-6Zn-xZr ternary alloys and their constitutive binary alloys are calculated. The thermodynamic effects of Zr addition on formations of Mg7Zn3, MgZn2, and MgZn are predicted. Using a ternary model for predicting diffusion coefficients, the diffusion coefficients of Zn and Zr in liquid Mg-6Zn-xZr alloys are calculated and the kinetic effects of Zr addition on the diffusion coefficient of Zn is discussed. The results show that the Zr addition can hinder the formations of Mg7Zn3, MgZn2, and MgZn inter-metallics in thermodynamics and kinetics. Full article
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<p>Calculated mixing enthalpy <math display="inline"><semantics> <mrow> <mo>Δ</mo> <mi>H</mi> </mrow> </semantics></math> of liquid Mg-Zn (<b>a</b>), Zn-Zr (<b>b</b>), and Mg-Zr (<b>c</b>) alloys with experimental data and calculated Calphad-data from the literature [<a href="#B29-metals-09-00607" class="html-bibr">29</a>,<a href="#B30-metals-09-00607" class="html-bibr">30</a>,<a href="#B31-metals-09-00607" class="html-bibr">31</a>].</p> Full article ">Figure 2
<p>The Mg-Zn phase diagram (<b>a</b>), Zn-Zr phase diagram (<b>b</b>) adapted from [<a href="#B32-metals-09-00607" class="html-bibr">32</a>], and the Mg-Zr phase diagram (<b>c</b>) adapted from [<a href="#B35-metals-09-00607" class="html-bibr">35</a>], with permission from Springer, 2002.</p> Full article ">Figure 3
<p>Calculated activities of the component of liquid Mg-Zn (<b>a</b>), Zn-Zr (<b>b</b>), and Mg-Zr (<b>c</b>) binary alloys with experimental data from the literature [<a href="#B36-metals-09-00607" class="html-bibr">36</a>,<a href="#B37-metals-09-00607" class="html-bibr">37</a>] at different temperatures.</p> Full article ">Figure 4
<p>Relationship between enthalpies of formation and melting temperature [<a href="#B32-metals-09-00607" class="html-bibr">32</a>] of some binary inter-metallics.</p> Full article ">Figure 5
<p>Three-dimensional diagram (<b>a</b>) and isogram (<b>b</b>) of the mixing enthalpy of liquid Mg-Zn-Zr alloys.</p> Full article ">Figure 6
<p>Effect of Zr addition on the activity coefficients of Mg and Zn in Mg-6Zn-<span class="html-italic">x</span>Zr melts at different temperatures, (<b>a</b>) Mg; (<b>b</b>) Zn.</p> Full article ">Figure 7
<p>Effect of Zr addition on the change of <math display="inline"><semantics> <mrow> <mo>Δ</mo> <mi>G</mi> </mrow> </semantics></math> associated with formation of Mg-Zn compounds in Mg-6Zn-<span class="html-italic">x</span>Zr (<span class="html-italic">x</span> = 0–0.6 wt%) melts at different temperatures, (<b>a</b>) MgZn; (<b>b</b>) MgZn<sub>2</sub>; (<b>c</b>) Mg<sub>7</sub>Zn<sub>3</sub>.</p> Full article ">Figure 7 Cont.
<p>Effect of Zr addition on the change of <math display="inline"><semantics> <mrow> <mo>Δ</mo> <mi>G</mi> </mrow> </semantics></math> associated with formation of Mg-Zn compounds in Mg-6Zn-<span class="html-italic">x</span>Zr (<span class="html-italic">x</span> = 0–0.6 wt%) melts at different temperatures, (<b>a</b>) MgZn; (<b>b</b>) MgZn<sub>2</sub>; (<b>c</b>) Mg<sub>7</sub>Zn<sub>3</sub>.</p> Full article ">Figure 8
<p>Dependence of the main interdiffusion coefficient <math display="inline"><semantics> <mrow> <msub> <mi>D</mi> <mrow> <mi>ZrZr</mi> </mrow> </msub> </mrow> </semantics></math> (<b>a</b>) of Zr atoms and the cross interdiffusion coefficient <math display="inline"><semantics> <mrow> <msub> <mi>D</mi> <mrow> <mi>ZnZr</mi> </mrow> </msub> </mrow> </semantics></math> (<b>b</b>) of Zn atoms on the composition of Zr in the liquid Mg-6 wt% Zn-<span class="html-italic">x</span>Zr alloys at 950 K, 1050 K, and 1150 K.</p> Full article ">Figure 9
<p>Dependence of the main interdiffusion coefficient <math display="inline"><semantics> <mrow> <msub> <mi>D</mi> <mrow> <mi>ZnZn</mi> </mrow> </msub> </mrow> </semantics></math> (<b>a</b>) of Zn atoms and the cross interdiffusion coefficient <math display="inline"><semantics> <mrow> <msub> <mi>D</mi> <mrow> <mi>ZrZn</mi> </mrow> </msub> </mrow> </semantics></math> (<b>b</b>) of Zr atoms on the composition of Zn in the liquid Mg-<span class="html-italic">x</span>Zn-0.6 wt% Zr alloys at 950 K, 1050 K, and 1150 K.</p> Full article ">
19 pages, 9384 KiB  
Article
Numerical Simulation Analysis of New Steel Sets Used for Roadway Support in Coal Mines
by Qinghai Li, Jingkai Li, Jinpeng Zhang, Changxiang Wang, Kai Fang, Limin Liu and Wenjing Wang
Metals 2019, 9(5), 606; https://doi.org/10.3390/met9050606 - 24 May 2019
Cited by 6 | Viewed by 2660
Abstract
The surrounding rock control is a tough issue in the roadway with the swelling soft rock. The steel set is an important material for the control of swelling soft rock roadways. However, traditional steel sets failed to prevent the expansive pressure of the [...] Read more.
The surrounding rock control is a tough issue in the roadway with the swelling soft rock. The steel set is an important material for the control of swelling soft rock roadways. However, traditional steel sets failed to prevent the expansive pressure of the soft rock. Based on traditional steel sets, this paper developed a new steel set through both theoretical analysis and numerical simulation. The results showed that the new steel set was the set with the roof beam 1000 mm from the top of the set and the floor beam 400 mm from the bottom end of the set. The maximum deformations of the roof-floor and two sides of the ventilation roadway controlled by the best-improved set at the observation point were 147 mm and 108 mm, respectively. So, the best-improved set can effectively control the surrounding rock of the ventilation roadway. This provides an effective method for the surrounding rock control in extremely soft rock roadways. Full article
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<p>Relative position and section of the roadway.</p> Full article ">Figure 2
<p>Destroyed roadway surrounding rock.</p> Full article ">Figure 3
<p>Cross section of the 12H I-beam.</p> Full article ">Figure 4
<p>Set model size.</p> Full article ">Figure 5
<p>Numerical model of set.</p> Full article ">Figure 6
<p>Application method of the displacement load.</p> Full article ">Figure 7
<p>Models of the roof beam (the RB sets).</p> Full article ">Figure 8
<p>Models of the floor beam (the FB sets).</p> Full article ">Figure 9
<p>The stress nephogram and the deformation nephogram of the traditional set.</p> Full article ">Figure 10
<p>The stress nephogram and the deformation nephogram of RB set I.</p> Full article ">Figure 11
<p>The stress nephogram and the deformation nephogram of RB set II.</p> Full article ">Figure 12
<p>The stress nephogram and the deformation nephogram of RB set III.</p> Full article ">Figure 13
<p>The stress nephogram and the deformation nephogram of RB set IV.</p> Full article ">Figure 14
<p>The stress nephogram and the deformation nephogram of FB set I.</p> Full article ">Figure 15
<p>The stress nephogram and the deformation nephogram of FB set II.</p> Full article ">Figure 16
<p>The stress nephogram and the deformation nephogram of FB set III.</p> Full article ">Figure 17
<p>The stress nephogram and the deformation nephogram of FB set IV.</p> Full article ">Figure 18
<p>The layout of the monitoring points.</p> Full article ">Figure 19
<p>The stress-strain curve of the RB sets.</p> Full article ">Figure 19 Cont.
<p>The stress-strain curve of the RB sets.</p> Full article ">Figure 20
<p>The stress-strain of the FB sets.</p> Full article ">Figure 20 Cont.
<p>The stress-strain of the FB sets.</p> Full article ">Figure 21
<p>The structure of the improved steel set.</p> Full article ">Figure 22
<p>The stress nephogram and deformation nephogram of the RFB set.</p> Full article ">Figure 23
<p>The displacement-load relationship of the four models.</p> Full article ">Figure 24
<p>The parameters of the RFB set.</p> Full article ">Figure 25
<p>The deformation of the surrounding rock with time in the ventilation roadway.</p> Full article ">
21 pages, 1016 KiB  
Article
Metallic Glasses: A New Approach to the Understanding of the Defect Structure and Physical Properties
by Vitaly Khonik and Nikolai Kobelev
Metals 2019, 9(5), 605; https://doi.org/10.3390/met9050605 - 24 May 2019
Cited by 42 | Viewed by 3350
Abstract
The work is devoted to a brief overview of the Interstitialcy Theory (IT) as applied to different relaxation phenomena occurring in metallic glasses upon structural relaxation and crystallization. The basic hypotheses of the IT and their experimental verification are shortly considered. The main [...] Read more.
The work is devoted to a brief overview of the Interstitialcy Theory (IT) as applied to different relaxation phenomena occurring in metallic glasses upon structural relaxation and crystallization. The basic hypotheses of the IT and their experimental verification are shortly considered. The main focus is given on the interpretation of recent experiments on the heat effects, volume changes and their link with the shear modulus relaxation. The issues related to the development of the IT and its relationship with other models on defects in metallic glasses are discussed. Full article
(This article belongs to the Special Issue Recent Advancements in Metallic Glasses)
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<p>Octahedral (<b>a</b>) and dumbbell (split) (<b>b</b>) interstitial defects in a computer model of a face-centered cubic lattice [<a href="#B27-metals-09-00605" class="html-bibr">27</a>]. All of the dumbbell atoms (marked by red circles) are characterized by <math display="inline"><semantics> <mrow> <mo>&lt;</mo> <mn>0</mn> <mo>,</mo> <mn>2</mn> <mo>,</mo> <mn>8</mn> <mo>,</mo> <mn>0</mn> <mo>&gt;</mo> </mrow> </semantics></math> Voronoi indexes. With permission from JETP Letters, 2019.</p> Full article ">Figure 2
<p>Histogram illustrating the distribution of the ratio of dilatation <math display="inline"><semantics> <msub> <mi>U</mi> <mrow> <mi>b</mi> <mi>u</mi> <mi>l</mi> <mi>k</mi> </mrow> </msub> </semantics></math> and shear <math display="inline"><semantics> <msub> <mi>U</mi> <mrow> <mi>s</mi> <mi>h</mi> <mi>e</mi> <mi>a</mi> <mi>r</mi> </mrow> </msub> </semantics></math> components of the elastic energy for dumbbell interstitials in 63 polycrystalline elemental metals. The same data for interstitial-type defects in 189 metallic glasses are also shown [<a href="#B27-metals-09-00605" class="html-bibr">27</a>]. With permission from JETP Letters, 2019.</p> Full article ">Figure 3
<p>Estimates of interstitial and vacancy concentrations in crystalline aluminum (<b>a</b>) and indium (<b>b</b>) derived from the diaelastic effect measurements. The melting temperatures are indicated [<a href="#B35-metals-09-00605" class="html-bibr">35</a>,<a href="#B36-metals-09-00605" class="html-bibr">36</a>]. With permission from Pleiades Publishing, LTD, 2019.</p> Full article ">Figure 4
<p>Temperature dependences of the quantities <math display="inline"><semantics> <mrow> <mo>Δ</mo> <mi>W</mi> </mrow> </semantics></math> and <math display="inline"><semantics> <mrow> <mo>∂</mo> <mo>Δ</mo> <mi>G</mi> <mo>/</mo> <mo>∂</mo> <mi>T</mi> </mrow> </semantics></math> entering Equation (<a href="#FD6-metals-09-00605" class="html-disp-formula">6</a>) derived from calorimetric and shear modulus measurements [<a href="#B49-metals-09-00605" class="html-bibr">49</a>]. The data correspond to structural relaxation below the glass transition. With permission from Elsevier, 2019.</p> Full article ">Figure 5
<p>Experimental and calculated using Equation (<a href="#FD7-metals-09-00605" class="html-disp-formula">7</a>) temperature dependences of the shear modulus <span class="html-italic">G</span> of glassy Zr<sub>65</sub>Cu<sub>15</sub>Al<sub>10</sub>Ni<sub>10</sub> in the initial and relaxed states. Temperature dependence of the shear modulus <math display="inline"><semantics> <mi>μ</mi> </semantics></math> after full crystallization is also shown. Calorimetric <math display="inline"><semantics> <msub> <mi>T</mi> <mi>g</mi> </msub> </semantics></math> is indicated by the arrow [<a href="#B51-metals-09-00605" class="html-bibr">51</a>]. With permission from Elsevier, 2019.</p> Full article ">Figure 6
<p>Dependence of the integral heat <span class="html-italic">Q</span> absorbed upon heating from 330 K to 610 K (supercooled liquid region) as a function of the shear modulus <math display="inline"><semantics> <msub> <mi>G</mi> <mn>330</mn> </msub> </semantics></math> measured at 330 K just after heating onset. The points correspond to different preannealing treatment applied for shear modulus measurements and DSC tests as indicated. The solid line gives the lest square fit [<a href="#B52-metals-09-00605" class="html-bibr">52</a>]. With permission from Elsevier, 2019.</p> Full article ">Figure 7
<p>Dependence of the relative density change on the quantity <math display="inline"><semantics> <mrow> <mi>l</mi> <mi>n</mi> <mo>(</mo> <mi>G</mi> <mo>/</mo> <msub> <mi>G</mi> <mn>0</mn> </msub> <mo>)</mo> </mrow> </semantics></math>, where <math display="inline"><semantics> <msub> <mi>G</mi> <mn>0</mn> </msub> </semantics></math> is the initial room-temperature shear modulus and <span class="html-italic">G</span> is the room-temperature shear modulus after annealing of bulk glassy <math display="inline"><semantics> <mrow> <mi>P</mi> <msub> <mi>d</mi> <mn>40</mn> </msub> <mi>C</mi> <msub> <mi>u</mi> <mn>30</mn> </msub> <mi>N</mi> <msub> <mi>i</mi> <mn>10</mn> </msub> <msub> <mi>P</mi> <mn>20</mn> </msub> </mrow> </semantics></math> at <math display="inline"><semantics> <mrow> <mi>T</mi> <mo>=</mo> <mn>533</mn> </mrow> </semantics></math> K [<a href="#B57-metals-09-00605" class="html-bibr">57</a>]. With permission from Elsevier, 2019.</p> Full article ">Figure 8
<p>The melting enthalpy vs. the absolute value of sum of the enthalpies of structural relaxation and crystallization. The numbers correspond to different Zr-, Pd- and La-based MGs [<a href="#B64-metals-09-00605" class="html-bibr">64</a>]. With permission from Elsevier, 2019.</p> Full article ">Figure 9
<p>The height of the boson peak as a function of the defect concentration <span class="html-italic">c</span> calculated using Equation (<a href="#FD2-metals-09-00605" class="html-disp-formula">2</a>). The line gives the least square fit [<a href="#B73-metals-09-00605" class="html-bibr">73</a>]. With permission from John Wiley and Sons, 2019.</p> Full article ">Figure 10
<p>Experimental and calculated height of the boson peak <math display="inline"><semantics> <msub> <mi>h</mi> <mi>B</mi> </msub> </semantics></math> as a function of the excess enthalpy <math display="inline"><semantics> <mrow> <mo>Δ</mo> <mi>H</mi> </mrow> </semantics></math> plotted according to Equation (<a href="#FD15-metals-09-00605" class="html-disp-formula">15</a>). The numbers near the data points indicate the corresponding preannealing temperatures in Kelvins [<a href="#B74-metals-09-00605" class="html-bibr">74</a>]. With permission from John Wiley and Sons, 2019.</p> Full article ">Figure 11
<p>Formation of a perfect icosahedron by the creation of dumbbell interstitials on the opposite faces of the FCC cell: (<b>Left</b>) elementary FCC cell where the arrows show how two atoms are inserted instead of one atom; and (<b>Right</b>) perfect icosahedron with <math display="inline"><semantics> <mrow> <mo>&lt;</mo> <mn>0</mn> <mo>,</mo> <mn>0</mn> <mo>,</mo> <mn>12</mn> <mo>,</mo> <mn>0</mn> <mo>&gt;</mo> </mrow> </semantics></math> Voronoi indexes formed by six dumbbell interstitials on the faces of the cell and one interstitial in the octahedral position (i.e., in center of the cell, see <a href="#metals-09-00605-f001" class="html-fig">Figure 1</a>a) [<a href="#B81-metals-09-00605" class="html-bibr">81</a>].</p> Full article ">Figure 12
<p>Dumbbell [001]-oriented interstitial (two red circles) and the relative changes of the volume of the Voronoi polyhedra <math display="inline"><semantics> <mrow> <mo>Δ</mo> <mrow> <mi>V</mi> <mo>/</mo> <mi>V</mi> </mrow> </mrow> </semantics></math> (indicated by the numbers) as compared with the ideal lattice for the atoms (shaded circles) in a copper crystal. <math display="inline"><semantics> <mrow> <mo>Δ</mo> <mrow> <mi>V</mi> <mo>/</mo> <mi>V</mi> </mrow> </mrow> </semantics></math>-changes are shown along the dumbbell axis, perpendicular to it and in the [111] direction. It is seen that the defect creates both positive and negative changes of the Voronoi polyhedra volume (i.e., changes of the local density) [<a href="#B27-metals-09-00605" class="html-bibr">27</a>]. With permission from JETP Letters, 2019.</p> Full article ">
11 pages, 5294 KiB  
Article
Interfacial Reaction and Microstructure Evolution of Sn-9Zn/Ni(Cu) Solder Joints
by Xuewei Zhu, Jian Peng, Xiaofeng Wei, Pengpeng Yan and Fu Wang
Metals 2019, 9(5), 604; https://doi.org/10.3390/met9050604 - 24 May 2019
Cited by 4 | Viewed by 3070
Abstract
Sn-9Zn solder is a promising Pb-free solder, but it tends to form bulky intermetallic compounds (IMC) grains at the interface when soldered with common simple metal Cu or Ni substrates. Interfacial reaction between Sn-9Zn solder and Ni(Cu) solid solution substrates at 250 °C [...] Read more.
Sn-9Zn solder is a promising Pb-free solder, but it tends to form bulky intermetallic compounds (IMC) grains at the interface when soldered with common simple metal Cu or Ni substrates. Interfacial reaction between Sn-9Zn solder and Ni(Cu) solid solution substrates at 250 °C and 350 °C were systematically probed in this study. Results showed that when soldered at 250 °C, a Ni5Zn21 layer is formed at Sn-Zn/Ni-20Cu and Sn-Zn/Ni-40Cu joints; and Ni2Sn2Zn + Cu5Zn8 and Cu5Zn8 phases are formed in Sn-Zn/Ni-60Cu and Sn-Zn/Ni-80Cu joints, respectively. Fine-grained IMCs formed at the interface are formed even when the soldered time is prolonged to 16 h. This result indicates that Ni(Cu) solid solution substrates inhibit the rapid growth of IMC at the Sn-Zn/Ni-Cu interface. Ni(Cu) solid solution substrate can also provide various combinations of reaction products at the Sn-Zn/Ni-Cu joints. The Ni5Zn21 transfers to Ni2Sn2Zn + Cu5Zn8 phases when the Cu content increased to 60%, and a bi-layered structure Ni2Sn2Zn + Cu5Zn8 IMCs was formed in Sn-Zn/Ni(Cu) joints at 350 °C regardless of the Cu content in Ni(Cu) substrate (20–80%). These results would provide an effective support in designing Sn-Zn soldering system with optimized IMC layer to improve mechanical performance. Full article
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Figure 1
<p>Schematic diagram of Sn-Zn/Ni(Cu) joints.</p> Full article ">Figure 2
<p>The Sn-Zn/Ni-20Cu joint of 16 h soldering at 250 °C: (<b>a</b>) cross-section micrograph; (<b>b</b>) Etched interface; (<b>c</b>) morphology of Ni<sub>5</sub>Zn<sub>21</sub> grains; (<b>d</b>) morphology of Ni<sub>2</sub>Sn<sub>2</sub>Zn grains.</p> Full article ">Figure 3
<p>The X-ray diffraction (XRD) pattern of etched interface of Sn-Zn/Ni-20Cu joint after soldered at 250 °C for 16 h.</p> Full article ">Figure 4
<p>Cross-section micrographs (<b>a</b>,<b>b</b>) and X-ray diffraction (XRD) patterns (<b>c</b>,<b>d</b>) of Sn-Zn/Ni-20Cu joints after soldered at 250 °C for 1 h (<b>a</b>,<b>c</b>) and 9 h (<b>b</b>,<b>d</b>).</p> Full article ">Figure 5
<p>Cross-section micrographs and X-ray diffraction (XRD) patterns of etched interface of joints after soldered 9 h at 250 °C: (<b>a</b>,<b>d</b>) Sn-Zn/Ni-40Cu; (<b>b</b>,<b>e</b>) Sn-Zn/Ni-60Cu; (<b>c</b>,<b>f</b>) Sn-Zn/Ni-80Cu.</p> Full article ">Figure 6
<p>Etched surface micrograph of joints after soldered 9 h at 250 °C: (<b>a</b>) Sn-Zn/Ni-20Cu; (<b>b</b>) Sn-Zn/Ni-40Cu; (<b>c</b>) Sn-Zn/Ni-60Cu; (<b>d</b>) Sn-Zn/Ni-80Cu.</p> Full article ">Figure 7
<p>Cross-section micrograph of joints after soldered 2 min at 350 °C: (<b>a</b>) Sn-Zn/Ni-20Cu; (<b>b</b>) Sn-Zn/Ni-40Cu; (<b>c</b>) Sn-Zn/Ni-60Cu; (<b>d</b>) Sn-Zn/Ni-80Cu.</p> Full article ">Figure 8
<p>Cross-section micrograph of joints after soldered 60 min at 350 °C: (<b>a</b>) Sn-Zn/Ni-20Cu; (<b>b</b>) Sn-Zn/Ni-40Cu; (<b>c</b>) Sn-Zn/Ni-60Cu; (<b>d</b>) Sn-Zn/Ni-80Cu.</p> Full article ">Figure 9
<p>Two reaction stages of the Sn-Zn/Ni(Cu) joints superimpose on the Zn-Sn-Ni-Cu isothermal tetrahedron at 250 °C during (<b>a</b>) early stage; and (<b>b</b>) later stage.</p> Full article ">Figure 10
<p>Two reaction stages of the Sn-Zn/Ni(Cu) joints superimpose on the Zn-Sn-Ni-Cu isothermal tetrahedron at 350 °C during (<b>a</b>) early stage; and (<b>b</b>) later stage.</p> Full article ">
12 pages, 12000 KiB  
Article
Springback Calibration of a U-Shaped Electromagnetic Impulse Forming Process
by Xiaohui Cui, Zhiwu Zhang, Hailiang Yu, Xiaoting Xiao and Yongqi Cheng
Metals 2019, 9(5), 603; https://doi.org/10.3390/met9050603 - 24 May 2019
Cited by 21 | Viewed by 3434
Abstract
A three-dimensional (3D) finite-element model (FEM), including quasi-static stamping, sequential coupling for electromagnetic forming (EMF) and springback, was established to analyze the springback calibration by electromagnetic force. Results show that the tangential stress at the sheet bending region is reduced, and even the [...] Read more.
A three-dimensional (3D) finite-element model (FEM), including quasi-static stamping, sequential coupling for electromagnetic forming (EMF) and springback, was established to analyze the springback calibration by electromagnetic force. Results show that the tangential stress at the sheet bending region is reduced, and even the direction of tangential stress at the bending region is changed after EMF. The springback can be significantly reduced with a higher discharge voltage. The simulation results are in good agreement with the experiment results, and the simulation method has a high accuracy in predicting the springback of quasi-static stamping and electromagnetic forming. Full article
(This article belongs to the Special Issue Forming Processes of Modern Metallic Materials)
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<p>Simulation strategy for springback control by EMF.</p> Full article ">Figure 2
<p>Structure-field model. (<b>a</b>) Geometric dimension; (<b>b</b>) 2D view of stamping; (<b>c</b>) 3D view of stamping; (<b>d</b>) angle in 2D view after springback.</p> Full article ">Figure 3
<p>Field analysis of the (<b>a</b>) electromagnetic field model; (<b>b</b>) updated model of the sheet and the coil; and (<b>c</b>) induced current on the sheet metal.</p> Full article ">Figure 4
<p>Current through the coil as a function of time.</p> Full article ">Figure 5
<p>Distribution of displacement in the sheet for 1.5 kV at different times: (<b>a</b>) 0 μs, (<b>b</b>) 75 μs, (<b>c</b>) 300 μs, (<b>d</b>) 600 μs, (<b>e</b>) 1500 μs, and (<b>f</b>) 3000 μs.</p> Full article ">Figure 6
<p>Distribution of tangential stresses in the sheet for 1.5 kV at different times: (<b>a</b>) 0 μs, (<b>b</b>) 75 μs, (<b>c</b>) 300 μs, (<b>d</b>) 600 μs, (<b>e</b>) 1500 μs, and (<b>f</b>) 3000 μs.</p> Full article ">Figure 7
<p>Springback after coil discharge at 1.5 kV: (<b>a</b>) 3D view; (<b>b</b>) 2D view profile.</p> Full article ">Figure 8
<p>Stress changes at special nodes: (<b>a</b>) node 3694; (<b>b</b>) node 7893.</p> Full article ">Figure 9
<p>Distribution of tangential stresses in the sheet for 2 kV at different times: (<b>a</b>) 0 μs, (<b>b</b>) 75 μs, (<b>c</b>) 300 μs, (<b>d</b>) 600 μs, (<b>e</b>) 1500 μs, and (<b>f</b>) 3000 μs.</p> Full article ">Figure 10
<p>Springback after the coil discharge at 2 kV: (<b>a</b>) 3D view; (<b>b</b>) 2D view.</p> Full article ">Figure 11
<p>Distribution of tangential stresses in the sheet for 3 kV at different times: (<b>a</b>) 75 μs, (<b>b</b>) 150 μs, (<b>c</b>) 300 μs, (<b>d</b>) 600 μs, (<b>e</b>) 1500 μs, and (<b>f</b>) 3000 μs.</p> Full article ">Figure 12
<p>Springback after coil discharge at 3 kV: (<b>a</b>) 3D view; (<b>b</b>) 2D view.</p> Full article ">Figure 13
<p>Equivalent plastic strain. (<b>a</b>) Quasi-static stamping; (<b>b</b>) 1.5 kV; (<b>c</b>) 2 kV; (<b>d</b>) 3 kV.</p> Full article ">Figure 14
<p>Strain rate versus time for (<b>a</b>) quasi-static stamping, and (<b>b</b>) coil discharge with different voltages.</p> Full article ">
10 pages, 1670 KiB  
Article
Finite Fracture Mechanics Assessment in Moderate and Large Scale Yielding Regimes
by Ali Reza Torabi, Filippo Berto and Alberto Sapora
Metals 2019, 9(5), 602; https://doi.org/10.3390/met9050602 - 24 May 2019
Cited by 12 | Viewed by 3210
Abstract
The coupled Finite Fracture Mechanics (FFM) criteria are applied to investigate the ductile failure initiation at blunt U-notched and V-notched plates under mode I loading conditions. The FFM approaches are based on the simultaneous fulfillment of the energy balance and a stress requirement, [...] Read more.
The coupled Finite Fracture Mechanics (FFM) criteria are applied to investigate the ductile failure initiation at blunt U-notched and V-notched plates under mode I loading conditions. The FFM approaches are based on the simultaneous fulfillment of the energy balance and a stress requirement, and they involve two material properties, namely the fracture toughness and the tensile strength. Whereas the former property is obtained directly from experiments, the latter is estimated through the Equivalent Material Concept (EMC). FFM results are presented in terms of the apparent generalized fracture toughness and compared with experimental data already published in the literature related to two different aluminium alloys, Al 7075-T6 and Al 6061-T6, respectively. It is shown that FFM predictions can be accurate even under moderate or large scale yielding regimes. Full article
(This article belongs to the Special Issue Fracture, Fatigue and Structural Integrity of Metallic Materials)
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<p>Blunt V-notch geometry. The presence of a “virtual” crack of length <span class="html-italic">c</span> stemming from the notch tip is also depicted.</p> Full article ">Figure 2
<p>U-notched structures (<span class="html-italic">ω</span> = 0°), apparent fracture toughness: Predictions by punctual FFM (dashed line) and average FFM (continuous line) related to experimental data on Al 7075-T6 plates (circles) and on Al 6061-T6 plates (triangles).</p> Full article ">Figure 3
<p>Blunt V-notched structures (<span class="html-italic">ω</span> = 30°), apparent generalized fracture toughness: Predictions by punctual FFM (dashed line) and average FFM (continuous line) related to experimental data on Al 7075-T6 plates (circles) and on Al 6061-T6 plates (triangles).</p> Full article ">Figure 4
<p>Blunt V-notched structures (<span class="html-italic">ω</span> = 60°), apparent generalized fracture toughness: Predictions by punctual FFM (dashed line) and average FFM (continuous line) related to experimental data on Al 7075-T6 plates (circles) and on Al 6061-T6 plates (triangles).</p> Full article ">Figure 5
<p>Blunt V-notched structures (<span class="html-italic">ω</span> = 90°), apparent generalized fracture toughness: Predictions by punctual FFM (dashed line) and average FFM (continuous line) related to experimental data on Al 7075-T6 plates (circles) and on Al 6061-T6 plates (triangles).</p> Full article ">Figure 6
<p>Average FFM predictions: Percentage discrepancy with respect to experimental data referring to blunt V-notched structures made of Al 7075-T6 (circles), and made of Al 6061-T6 (triangles).</p> Full article ">Figure 7
<p>Average FFM: Dimensionless apparent generalized fracture toughness (<b>a</b>) and dimensionless critical crack extension (<b>b</b>) compared to dimensionless notch root radius.</p> Full article ">
17 pages, 8402 KiB  
Article
Roasting Pretreatment Combined with Ultrasonic Enhanced Leaching Lead from Electrolytic Manganese Anode Mud
by Huimin Xie, Shiwei Li, Libo Zhang, Yongmi Wang and Hailin Long
Metals 2019, 9(5), 601; https://doi.org/10.3390/met9050601 - 24 May 2019
Cited by 10 | Viewed by 3428
Abstract
A method of conventional roasting pretreatment combined with ultrasonic enhanced leaching with ammonium acetate was proposed to solve the difficult problem of lead in electrolytic manganese anode mud. The effects of concentration, liquid–solid ratio, temperature, leaching time and rotating speed on the leaching [...] Read more.
A method of conventional roasting pretreatment combined with ultrasonic enhanced leaching with ammonium acetate was proposed to solve the difficult problem of lead in electrolytic manganese anode mud. The effects of concentration, liquid–solid ratio, temperature, leaching time and rotating speed on the leaching process under conventional and ultrasonic conditions were studied, and the lead leaching rate can be as high as 93.09% under optimized process parameters. A leaching kinetic model under conventional and ultrasonic conditions was established to explore the restrictive links of the leaching process. The results show that the leaching process under both conventional and ultrasonic conditions is controlled by diffusion, and the activation energies are 29.40 kJ/mol and 26.95 kJ/mol for the conventional and ultrasound enhance leaching processes, respectively. Full article
(This article belongs to the Special Issue Metal Removal and Recycling)
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<p>X-ray diffraction (XRD) pattern of the electrolytic manganese anode mud.</p> Full article ">Figure 2
<p>Scanning electron microscopy (SEM) and element distribution map of the electrolytic manganese anode mud. (<b>a</b>): SEM at low magnification under conventional conditions; (<b>b</b>): a partial enlarged view of (<b>a</b>).</p> Full article ">Figure 3
<p>Conventional roasting device (1 = intake valve; 2 = flow meter; 3 = pressure gauge; 4 = crucible; 5 = furnace; 6 = sample; 7 = quartz tube; 8 = water; 9 = ammonia solution; 10 = dilute NaOH solution; 11 = water; 12 = small vacuum pump).</p> Full article ">Figure 4
<p>Connection diagram of experimental device (1 = ultrasonic generator console; 2 = ultrasonic radiation rod; 3 = thermocouple; 4 = beaker; 5 = magnetic rotor; 6 = water bath control panel; 7 = support frame).</p> Full article ">Figure 5
<p>Effects of roasting pretreatment on leaching of lead.</p> Full article ">Figure 6
<p>XRD of 1123 K roasted electrolytic manganese anode mud.</p> Full article ">Figure 7
<p>SEM and element distribution map of 1123 K roasted electrolytic manganese anode mud. (<b>a</b>): SEM at low magnification under conventional conditions; (<b>b</b>): a partial enlarged view of (<b>a</b>).</p> Full article ">Figure 8
<p>Effects of ammonium acetate concentration on leaching of Pb.</p> Full article ">Figure 9
<p>Effects of liquid to solid ratio on leaching of Pb.</p> Full article ">Figure 10
<p>Effects of temperature on leaching of Pb.</p> Full article ">Figure 11
<p>Effects of leaching time on leaching of Pb.</p> Full article ">Figure 12
<p>Effects of rotating speed on leaching of Pb.</p> Full article ">Figure 13
<p>Effects of ultrasonic power on leaching of Pb.</p> Full article ">Figure 14
<p>XRD of leaching residues under conventional and ultrasonic conditions. (<b>a</b>): XRD at conventional conditions; (<b>b</b>): XRD at ultrasonic conditions.</p> Full article ">Figure 15
<p>SEM and element distribution map of leaching residues under conventional and ultrasonic conditions. (<b>a</b>): SEM at low magnification under conventional conditions; (<b>b</b>): a partial enlarged view of (<b>a</b>); (<b>c</b>): SEM at low magnification under ultrasonic conditions; (<b>d</b>): a partial enlarged view of (<b>c</b>).</p> Full article ">Figure 16
<p>Size distribution of lead residues under conventional and ultrasonic conditions. (<b>a</b>): under conventional conditions, (<b>b</b>): under ultrasonic conditions. (cumu is the cumulative distribution of granularity; diff is difference distribution of granularity).</p> Full article ">Figure 17
<p>Sketch of the unreacted shrinking core model. (t<sub>1</sub> is reaction time).</p> Full article ">Figure 18
<p>Plot of 1 − (1 − x)1/3 versus time for different temperatures under conventional and ultrasonic conditions. (<b>a</b>): under conventional conditions; (<b>b</b>): under ultrasonic conditions.</p> Full article ">Figure 19
<p>Plot of 1 − 2/3 x − (1 − x) <sup>2/3</sup> versus time for different temperatures under conventional and ultrasonic conditions. (<b>a</b>): under conventional conditions; (<b>b</b>): under ultrasonic conditions.</p> Full article ">Figure 20
<p>Plot of 1/3ln(1 − x) − 1 + (1 − x)<sup>−1/3</sup> versus time for different temperatures under conventional and ultrasonic conditions. (<b>a</b>): under conventional conditions; (<b>b</b>): under ultrasonic conditions.</p> Full article ">Figure 21
<p>Arrhenius curve obtained for 1/3ln(1 − x) − 1 − (1 − x)<sup>−1/3</sup> model under conventional and ultrasonic conditions. (<b>a</b>): under conventional conditions; (<b>b</b>): under ultrasonic conditions.</p> Full article ">
15 pages, 18518 KiB  
Article
Comparative Study of Jet Slurry Erosion of Martensitic Stainless Steel with Tungsten Carbide HVOF Coating
by Galileo Santacruz, Antonio Shigueaki Takimi, Felipe Vannucchi de Camargo, Carlos Pérez Bergmann and Cristiano Fragassa
Metals 2019, 9(5), 600; https://doi.org/10.3390/met9050600 - 24 May 2019
Cited by 18 | Viewed by 4709
Abstract
This work evaluates the behavior of a martensitic stainless steel (AISI 410) thermally treated by quenching and tempering with a tungsten carbide (86WC-10Co-4Cr) coating obtained by high-velocity oxygen fuel (HVOF) thermal spray deposition, analyzing the volume loss under erosive attacks at 30 [...] Read more.
This work evaluates the behavior of a martensitic stainless steel (AISI 410) thermally treated by quenching and tempering with a tungsten carbide (86WC-10Co-4Cr) coating obtained by high-velocity oxygen fuel (HVOF) thermal spray deposition, analyzing the volume loss under erosive attacks at 30 ∘ and 90 ∘ incidence angles by using jet slurry erosion equipment with electrofused alumina erodent particles. Firstly, the characterization of the samples was carried out in terms of the microstructure (SEM), thickness, roughness, porosity, and microhardness. Then, samples were structurally characterized in the identification of the phases (XRD and EDS) present in the coating, as well as the particle size distribution (LG) and morphology of the erodent. It was determined that the tungsten carbide coating presented better resistance to jet slurry erosion wear when compared to the martensitic stainless steel analyzed, which is approximately two times higher for the 30 ∘ angle. The more ductile and brittle natures of the substrate and the coating, respectively, were evidenced by their higher volumetric erosion at 30 ∘ for the first and 90 ∘ for the latter, as well as their particular material removal mechanisms. The enhanced resistance of the coating is mainly attributed to its low porosity and high WC-Co content, resulting in elevated mechanical resistance. Full article
(This article belongs to the Special Issue Advances in Design by Metallic Materials: Synthesis, Characterization, Simulation and Applications)
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<p>Resulting microstructure of AISI 410 stainless steel after thermal treatments by (<b>a</b>) LM 1000×, and (<b>b</b>) SEM 5000×, after Vilella acid etching.</p> Full article ">Figure 2
<p>Particle size distribution of aluminium oxide and cumulative distribution.</p> Full article ">Figure 3
<p>SEM image showing the abrasive particles’ shape and size (56×).</p> Full article ">Figure 4
<p>Schematic diagram of the jet slurry erosion tester.</p> Full article ">Figure 5
<p>XRD pattern of the HVOF 86WC-10Co-4Cr coating.</p> Full article ">Figure 6
<p>SEM image of the elemental microanalysis region in the cross-section of the HVOF 86WC-10Co4Cr coating using the EDS system (400×).</p> Full article ">Figure 7
<p>Quantitative pattern of the elemental microanalysis region in the cross-section of the HVOF 86WC-10Co-4Cr coating using the EDS system.</p> Full article ">Figure 8
<p>SEM image of the microstructure of the HVOF 86WC-10Co-4Cr coating in the cross-section (500×).</p> Full article ">Figure 9
<p>Variation of the accumulated volumetric erosion rate as a function of the impact angle of 30<math display="inline"><semantics> <msup> <mrow/> <mo>∘</mo> </msup> </semantics></math> for the martensitic stainless steel (AISI 410) and tungsten carbide (86WC-10Co-4Cr) coating.</p> Full article ">Figure 10
<p>Variation of the accumulated volumetric erosion rate as a function of the impact angle of 90<math display="inline"><semantics> <msup> <mrow/> <mo>∘</mo> </msup> </semantics></math> for the martensitic stainless steel (AISI 410) and tungsten carbide (86WC-10Co-4Cr) coating.</p> Full article ">Figure 11
<p>Eroded regions of the materials submitted to the jet slurry test obtained by Image J software (LM, 50×).</p> Full article ">Figure 12
<p>Three-dimensional scanning of the samples submitted to the jet slurry erosion test obtained the eroded depth by using the Geomagic Studio software (Version 2012.1.1, 3D Systems Inc, Rock Hill, SC, USA).</p> Full article ">
14 pages, 7325 KiB  
Article
Tooth Root Bending Fatigue Strength of High-Density Sintered Small-Module Spur Gears: The Effect of Porosity and Microstructure
by Vigilio Fontanari, Alberto Molinari, Michelangelo Marini, Wolfgang Pahl and Matteo Benedetti
Metals 2019, 9(5), 599; https://doi.org/10.3390/met9050599 - 24 May 2019
Cited by 11 | Viewed by 3854
Abstract
The present paper is aimed at investigating the effect of porosity and microstructure on tooth root bending fatigue of small-module spur gears produced by powder metallurgy (P/M). Specifically, three steel variants differing in powder composition and alloying route were subjected either to case-hardening [...] Read more.
The present paper is aimed at investigating the effect of porosity and microstructure on tooth root bending fatigue of small-module spur gears produced by powder metallurgy (P/M). Specifically, three steel variants differing in powder composition and alloying route were subjected either to case-hardening or sinter-hardening. The obtained results were interpreted in light of microstructural and fractographic inspections. On the basis of the Murakami a r e a method, it was found that fatigue strength is mainly dictated by the largest near-surface defect and by the hardness of the softest microstructural constituent. Owing to the very complicated shape of the critical pore, it was found that its maximum Feret diameter is the geometrical parameter that best captures the detrimental effect on fatigue. Full article
(This article belongs to the Special Issue Fatigue Design and Defects in Metals and Alloys)
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Graphical abstract

Graphical abstract
Full article ">Figure 1
<p>Geometry of the powder metallurgy (P/M) small-module spur gears. Image taken from [<a href="#B26-metals-09-00599" class="html-bibr">26</a>].</p> Full article ">Figure 2
<p>Tooth bending fatigue testing configuration. Detail of the anvils and the gear under testing. Image taken from [<a href="#B26-metals-09-00599" class="html-bibr">26</a>].</p> Full article ">Figure 3
<p>Sub-model of the finite element (FE) analyses carried out to estimate the tooth root bending stress. The applied loads are represented by the red arrows. Image taken from [<a href="#B26-metals-09-00599" class="html-bibr">26</a>].</p> Full article ">Figure 4
<p>Microstructure of the investigated material variants. (<b>a</b>) case-hardened A85Mo05, (<b>b</b>) sinter-hardened DHP, (<b>c</b>) sinter-hardened DDH.</p> Full article ">Figure 5
<p>Microhardness profiles measured at the gear tooth root. Each point represents the average of five measurements.</p> Full article ">Figure 6
<p>Results of the fatigue tests. (<b>a</b>) Case-hardened A85Mo05, (<b>b</b>) sinter-hardened DHP, (<b>c</b>) sinter-hardened DDH. The run-out samples are indicated by arrows.</p> Full article ">Figure 7
<p>Fractographic inspections of A85Mo05: (<b>a</b>) crack path (LOM), (<b>b</b>) crack initiation site (SEM).</p> Full article ">Figure 8
<p>Fractographic inspections of DHP: (<b>a</b>) crack path (SEM), (<b>b</b>) crack initiation site (SEM).</p> Full article ">Figure 9
<p>Fractographic inspections of DDH: (<b>a</b>) crack path (SEM), (<b>b</b>) crack initiation site (SEM).</p> Full article ">Figure 10
<p>Non-propagating cracks found by SEM inspection in the DDH variant. (<b>a</b>) Cluster of pores interconnected by the crack, (<b>b</b>) single pore.</p> Full article ">Figure 11
<p>Sketch illustrating the procedure for evaluating the characteristic dimensions of the pore, defined by the maximum Feret diameter. Image taken from [<a href="#B26-metals-09-00599" class="html-bibr">26</a>].</p> Full article ">
32 pages, 9799 KiB  
Review
A Review on Heterogeneous Nanostructures: A Strategy for Superior Mechanical Properties in Metals
by Yan Ma, Muxin Yang, Fuping Yuan and Xiaolei Wu
Metals 2019, 9(5), 598; https://doi.org/10.3390/met9050598 - 24 May 2019
Cited by 51 | Viewed by 6760
Abstract
Generally, strength and ductility are mutually exclusive in homogeneous metals. Nanostructured metals can have much higher strength when compared to their coarse-grained counterparts, while simple microstructure refinement to nanoscale generally results in poor strain hardening and limited ductility. In recent years, heterogeneous nanostructures [...] Read more.
Generally, strength and ductility are mutually exclusive in homogeneous metals. Nanostructured metals can have much higher strength when compared to their coarse-grained counterparts, while simple microstructure refinement to nanoscale generally results in poor strain hardening and limited ductility. In recent years, heterogeneous nanostructures in metals have been proven to be a new strategy to achieve unprecedented mechanical properties that are not accessible to their homogeneous counterparts. Here, we review recent advances in overcoming this strength–ductility trade-off by the designs of several heterogeneous nanostructures in metals: heterogeneous grain/lamellar/phase structures, gradient structure, nanotwinned structure and structure with nanoprecipitates. These structural heterogeneities can induce stress/strain partitioning between domains with dramatically different strengths, strain gradients and geometrically necessary dislocations near domain interfaces, and back-stress strengthening/hardening for high strength and large ductility. This review also provides the guideline for optimizing the mechanical properties in heterogeneous nanostructures by highlighting future challenges and opportunities. Full article
(This article belongs to the Special Issue Strengthening Mechanisms in Metallic Materials)
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<p>Microstructural characteristics and enhanced strength–ductility synergy of the MEA with HGS: (<b>a</b>) Architected HGS; (<b>b</b>) Engineering tensile stress–strain curves for the MEA with HGS under various testing temperatures. Adapted from [<a href="#B17-metals-09-00598" class="html-bibr">17</a>].</p> Full article ">Figure 2
<p>Unprecedented strengthening and extra strain-hardening in Ti with heterogeneous lamella structure (HLS): (<b>a</b>) Microstructural characteristics of HLS; (<b>b</b>) Outstanding tensile properties of Ti with HLS; (<b>c</b>) Engineering tensile stress–strain curves for various samples; (<b>d</b>) True strain hardening rate vs. true stress for various samples; (<b>e</b>) Back stress hardening behaviors for Ti with HLS; (<b>f</b>) Dislocation pile-ups at hetero-interfaces. Adapted from [<a href="#B16-metals-09-00598" class="html-bibr">16</a>].</p> Full article ">Figure 3
<p>In situ x-ray diffraction measurements for the HSSS with HPS during tensile deformation. (<b>a</b>) Lattice strains in both hard B2 phase and soft austenitic matrix vs. applied tensile strain; (<b>b</b>) Close-up view for the transient region. Adapted from [<a href="#B28-metals-09-00598" class="html-bibr">28</a>].</p> Full article ">Figure 4
<p>Schematic illustrations of SMAT (<b>a</b>) and SMGT (<b>b</b>) techniques. Adapted from [<a href="#B38-metals-09-00598" class="html-bibr">38</a>] and [<a href="#B41-metals-09-00598" class="html-bibr">41</a>].</p> Full article ">Figure 5
<p>Tensile properties and corresponding deformation mechanisms of gradient grained copper: (<b>a</b>) Engineering stress–strain curves; (<b>b</b>) TEM image of the top layer after a true tensile strain of 33% showing “in-situ” coarsened dislocation-free grains; (<b>c</b>) Variation of average grain sizes with applied tensile strain for the gradient layer. Adapted from [<a href="#B42-metals-09-00598" class="html-bibr">42</a>].</p> Full article ">Figure 6
<p>Tensile properties and extra strain hardening mechanisms of gradient grained IF steel: (<b>a</b>) Strain hardening rate vs. true strain curves; (<b>b</b>) Extra strain hardening by gradient structures; (<b>c</b>) Distributions of lateral strain and strain gradient along the depth; (<b>d</b>) Superior tensile properties of gradient structures over homogeneous counterparts. Adapted from [<a href="#B14-metals-09-00598" class="html-bibr">14</a>].</p> Full article ">Figure 7
<p>Tensile properties and corresponding deformation mechanisms of a TWIP steel with gradient twinned structure: (<b>a</b>) Twin volume fraction and twin thickness distributions along the radial direction; (<b>b</b>) True stress–strain curves; (<b>c</b>) Hierarchical twin structures formed during the pre-torsion deformation and the subsequent tensile loading; (<b>d</b>) Activated different twinning systems in the sequential torsion and tension deformation. Adapted from [<a href="#B47-metals-09-00598" class="html-bibr">47</a>].</p> Full article ">Figure 8
<p>Tensile properties and corresponding deformation mechanisms of gradient nanotwinned copper with highly tunable structural gradients: (<b>a</b>) Grain size and twin thickness distributions along the thickness; (<b>b</b>) Tensile properties; (<b>c</b>) Deformation microstructure of gradient nanotwinned structure at 1% strain showing bundles of concentrated dislocations; (<b>d</b>) High density of dislocations in the bundles of concentrated dislocations revealed by MD. Adapted from [<a href="#B13-metals-09-00598" class="html-bibr">13</a>].</p> Full article ">Figure 9
<p>Dynamic shear properties and corresponding deformation mechanisms of gradient structure: (<b>a</b>) Set-up of dynamic shear experiments in Hopkinson bar; (<b>b</b>) Shear stress-shear displacement curves; (<b>c</b>) Impact shear toughness vs. dynamic shear yield strength; (<b>d</b>–<b>f</b>) Propagation process of ASB from the surface to the center. Adapted from [<a href="#B58-metals-09-00598" class="html-bibr">58</a>].</p> Full article ">Figure 10
<p>Fatigue behaviors of an AISI 316L stainless steel with gradient structure: (<b>a</b>) Distributions of microhardness along the depth for different samples; (<b>b</b>) Cyclic deformation curves of the CG and the gradient structured samples at different stress amplitudes as indicated; (<b>c</b>) S-N curves of different samples; (<b>d</b>) Correlation between the tensile strength and fatigue ratio (adapted from [<a href="#B49-metals-09-00598" class="html-bibr">49</a>]).</p> Full article ">Figure 11
<p>Dislocation density-based continuum plasticity modeling for gradient structure: (<b>a</b>) The predicted height profile on the lateral surface of a gradient IF steel and comparison with experimental data; (<b>b</b>) The predicted strain hardening rate and comparison with experimental data; (<b>c</b>) Strength–ductility maps of the gradient IF steel considering GNDs and back stress; (<b>d</b>) The predicted GNDs density distributions for one case. Adapted from [<a href="#B54-metals-09-00598" class="html-bibr">54</a>].</p> Full article ">Figure 12
<p>Typical tensile strain-stress curves. (<b>a</b>) as deposited Cu (adapted from Refs. [<a href="#B75-metals-09-00598" class="html-bibr">75</a>]); (<b>b</b>) DPD-Cu (adapted from Refs. [<a href="#B90-metals-09-00598" class="html-bibr">90</a>]); (<b>c</b>) DPD-Cu-Zn (adapted from Refs. [<a href="#B80-metals-09-00598" class="html-bibr">80</a>]); (<b>d</b>) Cu-Al (adapted from [<a href="#B91-metals-09-00598" class="html-bibr">91</a>]).</p> Full article ">Figure 13
<p>TEM observations of the typical microstructure in an as-deposited NT-Cu sample. Bright-field TEM image (<b>a</b>) and the electron diffraction pattern (inset) show roughly equiaxed submicrometer-sized grains with random orientations separated by high-angle GBs. The statistical distributions for grain size (<b>b</b>) and for thickness of the twin/matrix lamellae (<b>c</b>) (obtained from the many TEM images of the same sample). Electron diffraction patterns (inset in (<b>d</b>)) indicate that the twins in each grain are parallel to each other in {111} planes (<b>d</b>), and high-resolution TEM images (<b>e</b>) show that the twins follow a sequence of ATATA, with twinning elements, for example, A: (<math display="inline"><semantics> <mrow> <mover accent="true"> <mn>1</mn> <mo>¯</mo> </mover> <mn>1</mn> <mover accent="true"> <mn>1</mn> <mo>¯</mo> </mover> </mrow> </semantics></math>)/[<math display="inline"><semantics> <mrow> <mover accent="true"> <mn>1</mn> <mo>¯</mo> </mover> <mn>12</mn> </mrow> </semantics></math>] and T: (<math display="inline"><semantics> <mrow> <mover accent="true"> <mn>1</mn> <mo>¯</mo> </mover> <mn>11</mn> </mrow> </semantics></math>)/[<math display="inline"><semantics> <mrow> <mn>1</mn> <mover accent="true"> <mn>1</mn> <mo>¯</mo> </mover> <mn>2</mn> </mrow> </semantics></math>]. Adapted from [<a href="#B74-metals-09-00598" class="html-bibr">74</a>].</p> Full article ">Figure 14
<p>In situ TEM images of crack growth process in a thin foil of NT Cu. (<b>a</b>) A zigzag crack formed during tensile loading. (<b>b</b>) Magnified images of region 1 in (<b>a</b>), circles indicate the short crack edges on TBs. Adapted from [<a href="#B107-metals-09-00598" class="html-bibr">107</a>].</p> Full article ">Figure 15
<p>(<b>a</b>) Engineering tensile stress–strain curves at a strain rate of 4 × 10<sup>−4</sup> s<sup>−1</sup>: Curve A: as-annealed CG Ni with an average grain size of 27 µm; Curve B: electrodeposited (ED) Ni (<span class="html-italic">d</span> = 1μm); Curve C: electrodeposited UFG Ni (<span class="html-italic">d</span> = 200 nm); Curve D: UFG Ni obtained via equal channel angular pressing (ECAP) for one pass; Curve E: ED nanocrystalline Ni (<span class="html-italic">d</span> = 18 nm); Curve F: electroplated nanodomained Ni (<span class="html-italic">d</span> = 150 nm, <span class="html-italic">d<sub>domain</sub></span> = 7 nm). (<b>b</b>) Normalized yield strength versus normalized tensile uniform elongation for metals. (<b>c</b>) TEM images of nanodomained Ni, showing low-angle domain boundaries (LADBs). (<b>d</b>) High-resolution electron microscope (HREM) images showing the dislocation pile-ups. (adapted from [<a href="#B119-metals-09-00598" class="html-bibr">119</a>]).</p> Full article ">Figure 16
<p>MD simulations of slip-nanodomain interactions. (<b>a</b>) Configuration for simulation cell with a straight edge dislocation and two nanodomains. (<b>b</b>) Simulated shear stress-shear strain curves as the straight dislocation is blocked by the nanodomains. (<b>c</b>) A sequence of snapshots at varying shear strains showing the pinning of the dislocation and its subsequent bowing around the nanodomains with LADB. (<b>d</b>) A sequence of snapshots at varying shear strains showing the pinning of the dislocation and its subsequent bowing around the nanodomains with high-angle domain boundaries (HADB). Adapted from [<a href="#B119-metals-09-00598" class="html-bibr">119</a>].</p> Full article ">
12 pages, 2440 KiB  
Article
Biomechanical Assessment of Design Parameters on a Self-Developed 3D-Printed Titanium-Alloy Reconstruction/Prosthetic Implant for Mandibular Segmental Osteotomy Defect
by Sheng-Ni Huang, Ming-You Shie, Yen-Wen Shen, Jui-Ting Hsu, Heng-Li Huang and Lih-Jyh Fuh
Metals 2019, 9(5), 597; https://doi.org/10.3390/met9050597 - 24 May 2019
Cited by 7 | Viewed by 3287
Abstract
Patients with oral cancer often have to undergo the surgery for mandibular excision. Once the bone in the cancerous area is removed, not only the facial area but also chewing function of the patient is needed to be repaired by clinicians. In recent [...] Read more.
Patients with oral cancer often have to undergo the surgery for mandibular excision. Once the bone in the cancerous area is removed, not only the facial area but also chewing function of the patient is needed to be repaired by clinicians. In recent years, the rapid growth of three-dimensional (3D) metal printing technology has meant that higher-quality facial reconstructions are now possible, which could even restore chewing function. This study developed 3D-printed titanium (Ti)-alloy reconstruction implant for a prosthesis designed for mandibular segmental osteotomy defects, and 3D finite element (FE) analysis was conducted to evaluate its biomechanical performance. The analyzed parameters in the FE models were as follows: (1) two prosthesis designs, namely a prosthesis retaining the residual mandibular bone (for patients with mild oral cancer) and a prosthesis with complete mandibular resection (for patients with severe oral cancer); (2) two lengths of prosthesis, namely 20 and 25 mm; and (3) three thicknesses of prosthesis, namely 0.8, 1, and 1.5 mm. A 45° lateral bite force (100 N) was applied to the top of the prosthesis as the loading condition. The results revealed that for the two prosthesis designs, the prosthesis retaining the residual mandibular bone showed higher stress on the prosthesis and cortical bone compared with the prosthesis with complete mandibular resection. Regarding the two prosthesis lengths, no fixed trend of prosthesis stress was found, but stress in the cortical bone was relatively high for a prosthesis length of 20 mm compared with that of 25 mm. For the three prosthesis thicknesses, as the thickness of the prosthesis decreased, the stress in the prosthesis decreased but the stress in the cortical bone increased. These findings require confirmation in future clinical investigations. Full article
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<p>Synbone manufactured artificial edentulous mandibular bone model (model number: 8571).</p> Full article ">Figure 2
<p>Resection conditions of the two mandibular defects modelled in this study—(<b>a</b>) mild oral cancer and (<b>b</b>) severe oral cancer. Based on these conditions, two designs of 3D-printed Ti-alloy mandibular implants were conducted—(<b>c</b>) Model 1 and (<b>d</b>) Model 2. Computer-aided design (CAD) models of (<b>e</b>) bone screws were used to (<b>f</b>) secure the 3D-printed Ti-alloy mandibular implant on the defective area of the mandible (Model 1 was shown as example).</p> Full article ">Figure 3
<p>Using Model 2 as an example: (<b>a</b>) implant body length of 20 mm; under this length, two abutments are allocated; (<b>b</b>) implant body length of 30 mm; under this length, three abutments are allocated; (<b>c</b>) implant body thickness of 0.8 mm; (<b>d</b>) implant body thickness of 1.0 mm; and (<b>e</b>) implant body thickness of 1.5 mm.</p> Full article ">Figure 4
<p>(<b>a</b>) FE model of Model 1. (<b>b</b>) The temporomandibular joint is the region specified by the boundary conditions (blue area). (<b>c</b>) Oblique occlusal force (100 N) was applied above the abutment (the red arrow).</p> Full article ">Figure 5
<p>Von Mises stress distribution of the implant: (<b>a</b>) Model 1 (length: 20 mm; thickness: 0.8 mm) and (<b>b</b>) Model 2 (length: 20 mm; thickness: 0.8 mm) (Unit: MPa).</p> Full article ">Figure 6
<p>Von Mises stress distribution of the cortical bones near the 3D-printed implant in (<b>a</b>) Model 1 (length: 20 mm; thickness: 0.8 mm) and (<b>b</b>) Model 2 (length: 20 mm; thickness: 0.8 mm) (Unit: MPa).</p> Full article ">Figure 7
<p>Geometry and loading conditions of the three-points-bending model (upper level) and the stress distributions of the three-points-bending model (lower left) and Model 1 (lower right).</p> Full article ">Figure 8
<p>(<b>a</b>) A 3D-printed Ti-alloy reconstruction sample was constructed by referring to CAD model of Model 1 (length: 20 mm; thickness: 0.8 mm) and (<b>b</b>) was well placed in the defect of the artificial mandible.</p> Full article ">
17 pages, 5990 KiB  
Article
Analysis of the Depth of Immersion of the Submerged Entry Nozzle on the Oscillations of the Meniscus in a Continuous Casting Mold
by F. Saldaña-Salas, E. Torres-Alonso, J.A. Ramos-Banderas, G. Solorio-Díaz and C.A. Hernández-Bocanegra
Metals 2019, 9(5), 596; https://doi.org/10.3390/met9050596 - 24 May 2019
Cited by 12 | Viewed by 3652
Abstract
In this study the effects of the depth of immersion of the Submerged Entry Nozzles (SEN) on the fluid-dynamic structure, oscillations of the free surface and opening of the slag layer, in a continuous casting mold for conventional slab of steel were analyzed. [...] Read more.
In this study the effects of the depth of immersion of the Submerged Entry Nozzles (SEN) on the fluid-dynamic structure, oscillations of the free surface and opening of the slag layer, in a continuous casting mold for conventional slab of steel were analyzed. For this work, a water/oil/air system was used in a 1:1 scale model, using the techniques of Particle Image Velocimetry (PIV), colorimetry and mathematical multiphase simulation. The results of the fluid dynamics by PIV agree with those obtained in the mathematical simulation, as well as with the dispersion of dye. It was observed that working with immersion depths of 100 mm or less could be detrimental to steel quality because they promote surface oscillations of a higher degree of Stokes with high elevations and asymmetry in their three dimensions. In addition, this generates an excessive opening of the oil layer which was corroborated through the quantification of the F index. On the other hand, with depths of immersion in the range of 150–200 mm, lower oscillations were obtained as well as zones of low speed near the wall of the SEN and a smaller opening of the oil layer. Full article
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<p>Geometrical mold and Submerged Entry Nozzle (SEN) dimensions and planes of analysis; (<b>a</b>) frontal view and (<b>b</b>) lateral view.</p> Full article ">Figure 2
<p>Initial and boundary conditions.</p> Full article ">Figure 3
<p>Experimental setup for Particle Image Velocimetry (PIV).</p> Full article ">Figure 4
<p>Velocity fields in the P<sub>III</sub> plane (<b>a</b>) PIV and (<b>b</b>) mathematical modelling. (<b>c</b>) Colorant dispersion technique.</p> Full article ">Figure 5
<p>Velocity fields in the P<sub>III</sub> plane. (<b>a</b>) Case I, (<b>b</b>) Case II and (<b>c</b>) Case III.</p> Full article ">Figure 6
<p>Velocity fields in the P<sub>V</sub> plane. (<b>a</b>) Case I, (<b>b</b>) Case II and (<b>c</b>) Case III.</p> Full article ">Figure 7
<p>Meniscus level oscillations and close up. (<b>a</b>) Case I, (<b>b</b>) Case II, (<b>c</b>) Case III.</p> Full article ">Figure 8
<p>Meniscus level oscillations at different times and planes. (<b>a</b>) Case I, (<b>b</b>) Case II, (<b>c</b>) Case III.</p> Full article ">Figure 9
<p>Scheme of <span class="html-italic">F</span> index.</p> Full article ">Figure 10
<p><span class="html-italic">F</span> index for all cases at different times in both sides of the mold.</p> Full article ">Figure 11
<p>Oil layer opening. (<b>a</b>) Case I, (<b>b</b>) Case II, (<b>c</b>) Case III.</p> Full article ">
19 pages, 9612 KiB  
Article
A Novel Damage Model to Predict Ductile Fracture Behavior for Anisotropic Sheet Metal
by Hua Zhang, Hong Zhang, Fuguo Li and Jun Cao
Metals 2019, 9(5), 595; https://doi.org/10.3390/met9050595 - 23 May 2019
Cited by 8 | Viewed by 4032
Abstract
The purpose of the present work is to investigate the fracture behavior of anisotropic sheet metal under various stress states. Notched tension and flat-grooved tension tests at 0°, 45°, and 90° directions with respect to rolling direction were carried out by a hybrid [...] Read more.
The purpose of the present work is to investigate the fracture behavior of anisotropic sheet metal under various stress states. Notched tension and flat-grooved tension tests at 0°, 45°, and 90° directions with respect to rolling direction were carried out by a hybrid experimental–numerical approach, and then a novel damage model was proposed by coupling Hill48’s criterion. Based on this, finite element method (FEM) analysis models were established. The force–displacement responses of experiments and simulations are in good agreement, which verify the FEM models. The predictability of the damage model established for the fracture behavior of anisotropic materials was studied by comparing the fracture displacements between experiments and simulations. It is found that the predictability of novel damage model is basically consistent with predictive results. The difference of damage locations and local strain evolutions at a 45° direction is greater than the other directions. In addition, stress triaxiality does not play a predominant role in the fracture process for notched tension specimens, while it does play a predominant role for flat-grooved tension specimens. Full article
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Graphical abstract

Graphical abstract
Full article ">Figure 1
<p>Shapes of specimens for (<b>a</b>) notched tension and (<b>b</b>) flat-grooved tension (all units are in mm; black dots indicate the start and end points of the digital image correlation (DIC) extensometer).</p> Full article ">Figure 2
<p>Seven stress states in the space of (<math display="inline"><semantics> <mrow> <msub> <mi>R</mi> <mi>H</mi> </msub> <mo>,</mo> <msub> <mi>L</mi> <mi>e</mi> </msub> <mo>,</mo> <mi>ω</mi> </mrow> </semantics></math>) with plane stress conditions.</p> Full article ">Figure 3
<p>The relationship between anisotropy and principal stress coordinate systems under uniaxial tension with a notch state.</p> Full article ">Figure 4
<p>The difference between the definitions of (<b>a</b>) <math display="inline"><semantics> <mi>θ</mi> </semantics></math> and (<b>b</b>) <math display="inline"><semantics> <mi>ω</mi> </semantics></math>.</p> Full article ">Figure 5
<p>The stress strain curves of loading and unloading at (<b>a</b>) 0°, (<b>b</b>) 45°, and (<b>c</b>) 90° directions, as well as (<b>d</b>) the linear fitting relation between ln(1-<span class="html-italic">D</span>) and plastic strain <math display="inline"><semantics> <mrow> <msub> <mi>ε</mi> <mi>p</mi> </msub> </mrow> </semantics></math>.</p> Full article ">Figure 6
<p>The relation between the critical damage variable and <math display="inline"><semantics> <mi>θ</mi> </semantics></math>.</p> Full article ">Figure 7
<p>(<b>a</b>) Optimizing strain hardening extrapolations over larger strains; (<b>b</b>) predictions of force–displacement responses using different strain hardening laws.</p> Full article ">Figure 8
<p>Mesh design of notched tension specimens (<b>a</b>) R1.5, (<b>b</b>) R4.5, and (<b>c</b>) R 25, as well as flat-grooved tension specimens (<b>d</b>) PLT-R4, (<b>e</b>) PLT-R8, and (<b>f</b>) PLT-R12.</p> Full article ">Figure 9
<p>Comparisons of force–displacement responses between experiments and simulations, as well as the strain evolutions at the failure material point for notched tension <span class="html-italic">R</span> = 1.5 mm, at the (<b>a</b>) 0°, (<b>b</b>) 45°, and (<b>c</b>) 90° directions (PEEQ represents equivalent plastic strain).</p> Full article ">Figure 10
<p>Comparisons of force–displacement responses between experiments and simulations, as well as the strain evolutions at the failure material point for notched tension <span class="html-italic">R</span> = 4.5 mm, at the (<b>a</b>) 0°, (<b>b</b>) 45°, and (<b>c</b>) 90° directions (PEEQ represents equivalent plastic strain).</p> Full article ">Figure 11
<p>Comparisons of force–displacement responses between experiments and simulations, as well as the strain evolutions at the failure material point for notched tension <span class="html-italic">R</span> = 25 mm, at the (<b>a</b>) 0°, (<b>b</b>) 45°, and (<b>c</b>) 90° directions (PEEQ represents equivalent plastic strain).</p> Full article ">Figure 12
<p>Comparisons of force–displacement responses between experiments and simulations, as well as the strain evolutions at the failure material point for flat-grooved tension <span class="html-italic">R</span> = 4 mm at the (<b>a</b>) 0°, (<b>b</b>) 45°, and (<b>c</b>) 90° directions (PEEQ represents equivalent plastic strain).</p> Full article ">Figure 13
<p>Comparisons of force–displacement responses between experiments and simulations, as well as the strain evolutions at the failure material point for flat-grooved tension <span class="html-italic">R</span> = 8 mm at the (<b>a</b>) 0°, (<b>b</b>) 45°, and (<b>c</b>) 90° directions (PEEQ represents equivalent plastic strain).</p> Full article ">Figure 14
<p>Comparisons of force–displacement responses between experiments and simulations, as well as the strain evolutions at the failure material point for flat-grooved tension <span class="html-italic">R</span> = 12 mm at the (<b>a</b>) 0°, (<b>b</b>) 45°, and (<b>c</b>) 90° directions (PEEQ represents equivalent plastic strain).</p> Full article ">Figure 15
<p>The relative difference of predicting fracture displacement between experiments and simulations using an anisotropic damage model.</p> Full article ">Figure 16
<p>Damage distributions at the fracture for notched tension specimens at the 0°, 45°, and 90° directions.</p> Full article ">Figure 17
<p>Damage distributions at the fracture for flat-grooved tension specimens at 0°, 45°, and 90° directions.</p> Full article ">Figure 18
<p>Comparisons between stress triaxiality and damage distribution at fracture along the half cross-section for notched specimens (<b>a</b>) R1.5, (<b>b</b>) R4.5, and (<b>c</b>) R25.</p> Full article ">Figure 19
<p>Comparisons between stress triaxiality and damage distribution at fracture along the half cross-section for flat-grooved tension specimens (<b>a</b>) PLT-R4, (<b>b</b>) PLT-R8, and (<b>c</b>) PLT-R12.</p> Full article ">Figure 20
<p>Strain evolutions at fracture onset point for (<b>a</b>) notched tension specimens and (<b>b</b>) flat-grooved tension specimens.</p> Full article ">Figure 21
<p>Damage evolution at the (<b>a</b>) 0°, (<b>b</b>) 45°, and (<b>c</b>) 90° directions for notched and flat-grooved tension specimens.</p> Full article ">Figure 22
<p>Schematic diagrams of diffuse and localized necking.</p> Full article ">Figure 23
<p>Scanning electron microscope (SEM) images of fracture surfaces for (<b>a</b>–<b>c</b>) notched tension specimens and (<b>d</b>–<b>f</b>) flat-grooved tension specimens.</p> Full article ">
10 pages, 5367 KiB  
Article
Improvement of Filler Wire Dilution Using External Oscillating Magnetic Field at Full Penetration Hybrid Laser-Arc Welding of Thick Materials
by Ömer Üstündağ, Vjaceslav Avilov, Andrey Gumenyuk and Michael Rethmeier
Metals 2019, 9(5), 594; https://doi.org/10.3390/met9050594 - 23 May 2019
Cited by 18 | Viewed by 3858
Abstract
Hybrid laser-arc welding offers many advantages, such as deep penetration, good gap bridge-ability, and low distortion due to reduced heat input. The filler wire which is supplied to the process is used to influence the microstructure and mechanical properties of the weld seam. [...] Read more.
Hybrid laser-arc welding offers many advantages, such as deep penetration, good gap bridge-ability, and low distortion due to reduced heat input. The filler wire which is supplied to the process is used to influence the microstructure and mechanical properties of the weld seam. A typical problem in deep penetration high-power laser beam welding with filler wire and hybrid laser-arc welding is an insufficient mixing of filler material in the weld pool, leading to a non-uniform element distribution in the seam. In this study, oscillating magnetic fields were used to form a non-conservative component of the Lorentz force in the weld pool to improve the element distribution over the entire thickness of the material. Full penetration hybrid laser-arc welds were performed on 20-mm-thick S355J2 steel plates with a nickel-based wire for different arrangements of the oscillating magnetic field. The Energy-dispersive X-ray spectroscopy (EDS) data for the distribution of two tracing elements (Ni and Cr) were used to analyze the homogeneity of dilution of the filler wire. With a 30° turn of the magnetic field to the welding direction, a radical improvement in the filler material distribution was demonstrated. This would lead to an improvement of the mechanical properties with the use of a suitable filler wire. Full article
(This article belongs to the Special Issue Electromagnetic Stirring Technique in Metallurgy and Material Processing)
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<p>(<b>a</b>) Geometrical sizes of the weld pool, which influence the formation of droplets; (<b>b</b>) stability threshold for liquid steel, see Equation (1).</p> Full article ">Figure 2
<p>Nickel concentration of a hybrid laser-arc welding (HLAW) on a 10-mm-thick steel plate [<a href="#B20-metals-09-00594" class="html-bibr">20</a>].</p> Full article ">Figure 3
<p>(<b>a</b>) Scheme of electromagnetic weld pool support system; (<b>b</b>) root side view of electromagnetic weld pool support system (the gap does not affect the current density). GMAW: gas metal arc welding.</p> Full article ">Figure 4
<p>(<b>a</b>) Magnetic field parallel to the welding direction and influence of the electric current by the gap, non-conservative component of the Lorentz force; (<b>b</b>) optimal balance of conservative and non-conservative components of the Lorentz force at a turning through 30°.</p> Full article ">Figure 5
<p>(<b>a</b>) Experimental setup; (<b>b</b>) AC magnet used for the experiments.</p> Full article ">Figure 6
<p>Cross sections of HLAW on 20-mm-thick plates of S355J2 using 12.2 kW laser power and wire feeding rate 11 m min<sup>−1</sup>, at a welding speed of 0.5 m min<sup>−1</sup>: (<b>a</b>) without electromagnetic weld pool support; (<b>b</b>) with electromagnetic support only (magnetic field perpendicular to the welding direction); (<b>c</b>) with electromagnetic support and stirring (magnetic field turned through 30°).</p> Full article ">Figure 7
<p>Schematic drawing of the different zones for results of the EDS analysis.</p> Full article ">Figure 8
<p>Results of the EDS analysis for the distribution of two tracing elements nickel and chromium of a HLAW 20 mm thick S355J2 with electromagnetic weld pool support (magnetic field perpendicular to the welding direction): (<b>a</b>) upper part; (<b>b</b>) root part.</p> Full article ">Figure 9
<p>Results of the EDS analysis for the distribution of the two tracing elements nickel and chromium of a HLAW 20 mm thick S355J2 with electromagnetic weld pool support and stirring (magnetic field turned through 30° to the welding direction): (<b>a</b>) upper part; (<b>b</b>) root part.</p> Full article ">
12 pages, 6004 KiB  
Article
Study on Microstructure and Properties of Tailored Hot-Stamped U-shaped Parts Based on Temperature Field Control
by Xiangji Li, Limei Xiao, Qifeng Zheng, Huan Zhang and Yanjiao Gu
Metals 2019, 9(5), 593; https://doi.org/10.3390/met9050593 - 23 May 2019
Cited by 2 | Viewed by 2871
Abstract
In order to meet the needs of the automotive industry, it is necessary to produce “tailored” parts. The U-shaped die equipped with a high-speed airflow device was designed to conduct the hot stamping experiments. The microstructure, micro-hardness, tensile properties, and fracture behavior of [...] Read more.
In order to meet the needs of the automotive industry, it is necessary to produce “tailored” parts. The U-shaped die equipped with a high-speed airflow device was designed to conduct the hot stamping experiments. The microstructure, micro-hardness, tensile properties, and fracture behavior of the parts were analyzed. The experimental results showed that the quenched phase of the hardened section was mainly martensite, and the micro-hardness and tensile strength could reach 445 HV and 1454 MPa, respectively. The fracture mechanism was brittle fracture. For the toughness section, as the tool temperature increased from 300 to 600 °C, both micro-hardness and tensile strength decreased. Meanwhile, the area fractions of bainite and ferrite increased, and the area fraction of martensite reduced. The fracture behavior was plastic fracture. Full article
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<p>Phase composition: (<b>a</b>) prediction based on the continuous cooling transformation (CCT) diagram for B1500HS steel [<a href="#B6-metals-09-00593" class="html-bibr">6</a>], (<b>b</b>) in examined U-shaped parts.</p> Full article ">Figure 2
<p>Schematic of the die: (<b>a</b>) overall structure, (<b>b</b>) slow cooling zone, and (<b>c</b>) quenching zone.</p> Full article ">Figure 3
<p>Predicted temperature distribution with a different holding time of (<b>a</b>) 5 s, (<b>b</b>) 8 s, (<b>c</b>) 10 s, and (<b>d</b>) 12 s.</p> Full article ">Figure 4
<p>The relationship between the predicted maximum temperature of the as-quenched blank and the holding time at different die temperatures.</p> Full article ">Figure 5
<p>Predicted temperature distribution of the blank at different air flow rates of (<b>a</b>) 1 m/s, (<b>b</b>) 3 m/s, (<b>c</b>) 7 m/s, and (<b>d</b>) 10 m/s.</p> Full article ">Figure 6
<p>The relationships between predicted maximum temperature of the as-quenched blanks and the number of hot stampings under various air velocities.</p> Full article ">Figure 7
<p>The schematic illustration of heat treatment of the blank.</p> Full article ">Figure 8
<p>The two regions of the U-shaped part and the sampling position.</p> Full article ">Figure 9
<p>The sample size of the tensile test (unit: mm).</p> Full article ">Figure 10
<p>The scanning electron microscope (SEM) image at: (<b>a</b>) location C1 with a 300 °C die temperature, (<b>b</b>) location C2 with a 500 °C die temperature, (<b>c</b>) location H1 with a 400 °C die temperature, (<b>d</b>) location H2 with a 400 °C die temperature, (<b>e</b>) location H1 with a 500 °C die temperature, and (<b>f</b>) location H1 with a 600 °C die temperature.</p> Full article ">Figure 11
<p>Various micrographs showing (from top to bottom) two-stage color tint etched optical micrographs, and the manually generated microstructure images at: (<b>a</b>) location C1 with a 300 °C die temperature, (<b>b</b>) location H1 with a 300 °C die temperature, and (<b>c</b>) location H1 with a 600 °C die temperature.</p> Full article ">Figure 12
<p>The area fraction under different tool temperatures of: (<b>a</b>) martensite, (<b>b</b>) bainite, and (<b>c</b>) ferrite.</p> Full article ">Figure 13
<p>Hardness of different locations at different die temperatures.</p> Full article ">Figure 14
<p>The mechanical properties for two regions at different tool temperatures: (<b>a</b>) the average ultimate tensile strength (UTS), and (<b>b</b>) the average elongation.</p> Full article ">Figure 15
<p>Fracture morphology at (<b>a</b>) location C1 at 300 °C, and (<b>b</b>) 500 °C; and (<b>c</b>) location H1 at 300 °C, (<b>d</b>) 400 °C, (<b>e</b>) 500 °C, and (<b>f</b>) 600 °C.</p> Full article ">
21 pages, 24828 KiB  
Article
Development of High-Fidelity Imaging Procedures to Establish the Local Material Behavior in Friction Stir Welded Stainless Steel Joints
by S. Ramachandran, A. K. Lakshminarayanan, P. A. S. Reed and J. M. Dulieu-Barton
Metals 2019, 9(5), 592; https://doi.org/10.3390/met9050592 - 23 May 2019
Cited by 4 | Viewed by 3910
Abstract
Friction stir welded (FSW) 304 austenitic stainless steel (SS) joints are studied using a range of microstructural characterization techniques to identify various sub-regions across the weld. A high-resolution (HR) 2D-digital image correlation (DIC) methodology is developed to assess the local strain response across [...] Read more.
Friction stir welded (FSW) 304 austenitic stainless steel (SS) joints are studied using a range of microstructural characterization techniques to identify various sub-regions across the weld. A high-resolution (HR) 2D-digital image correlation (DIC) methodology is developed to assess the local strain response across the weld surface and cross-section in the elastic regime. The HR-DIC methodology includes the stitching of multiple images, as it is only possible to partially cover the FSW region using a single camera with the high-resolution optical set-up. An image processing procedure is described to stitch the strain maps as well as strain data sets that allow full-field strain to be visualized and interrogated over the entire FSW region. It is demonstrated that the strains derived from the DIC can be associated with the local weld geometry and the material microstructure in the region of the FSW. The procedure is validated in the material elastic range and provides an important first step in enabling detailed mechanical assessments of the local effects in the FSW process. Full article
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<p>(<b>a</b>) Friction stir welded (FSW) (stainless steel-stainless steel (SS-SS)) joint; (<b>b</b>) schematic for extracting the specimens for various material characterizations (AS—Advancing side; RS—Retreating side).</p> Full article ">Figure 2
<p>(<b>a</b>) Water jet machined transverse tensile specimen of the FSW (SS-SS) joint; (<b>b</b>) fine speckle pattern produced through an airbrush.</p> Full article ">Figure 3
<p>(<b>a</b>) Region of interest (ROI) for the high-resolution (HR) 2D-digital image correlation (DIC) strain measurements; (<b>b</b>) the HR 2D-DIC experimental setup.</p> Full article ">Figure 4
<p>Image stitching procedures involved in Image-J to stitch grey scale images.</p> Full article ">Figure 5
<p>Sequentially stitched optical micrographs of the FSW (SS-SS) weld: (<b>a</b>) weld surface (AS); (<b>b</b>) weld surface (RS); (<b>c</b>) weld cross-section.</p> Full article ">Figure 5 Cont.
<p>Sequentially stitched optical micrographs of the FSW (SS-SS) weld: (<b>a</b>) weld surface (AS); (<b>b</b>) weld surface (RS); (<b>c</b>) weld cross-section.</p> Full article ">Figure 6
<p>(<b>a</b>) SEM-BSE of Weld nugget; (<b>b</b>) W-SEM-EDS map; (<b>c</b>) W-EDS line scan.</p> Full article ">Figure 7
<p>SEM-EBSD Micrographs along the FSW (SS-SS) weld cross-section: (<b>a</b>) BM; (<b>b</b>) HAZ/TMAZ/Weld nugget interfaces.</p> Full article ">Figure 8
<p>(<b>a</b>) ROI for hardness mapping; (<b>b</b>) optical micrograph showing the microhardness indents; (<b>c</b>) microhardness contour map along the cross-section of the weld; (<b>d</b>) microhardness line profile (L1) along the line as shown in (<b>b</b>).</p> Full article ">Figure 8 Cont.
<p>(<b>a</b>) ROI for hardness mapping; (<b>b</b>) optical micrograph showing the microhardness indents; (<b>c</b>) microhardness contour map along the cross-section of the weld; (<b>d</b>) microhardness line profile (L1) along the line as shown in (<b>b</b>).</p> Full article ">Figure 9
<p>(<b>a</b>) ROIs (1, 2) for the Alicona profile measurements along the FSW (SS-SS) weld cross-section; (<b>b</b>) Alicona 3D surface topography of the ROI-1 located in (<b>a</b>); (<b>c</b>) Alicona 3D surface topography of the ROI-2 located in (<b>a</b>).</p> Full article ">Figure 10
<p>(<b>a</b>) Alicona ROI on the FSW (SS-SS) weld surface; (<b>b</b>) Alicona surface profile of the FSW (SS-SS) weld surface. (<b>c</b>) line profiles showing the thickness variations along the horizontal lines (1, 2) as shown in (b).</p> Full article ">Figure 10 Cont.
<p>(<b>a</b>) Alicona ROI on the FSW (SS-SS) weld surface; (<b>b</b>) Alicona surface profile of the FSW (SS-SS) weld surface. (<b>c</b>) line profiles showing the thickness variations along the horizontal lines (1, 2) as shown in (b).</p> Full article ">Figure 11
<p>(<b>a</b>) DIC ROIs on the FSW (SS-SS) weld; (<b>b</b>) sequentially stitched speckle map using Image-J.</p> Full article ">Figure 12
<p>ε<sub><span class="html-italic">yy</span></sub> strain map derived from the DIC correlation on the stitched speckle maps of the FSW (SS-SS) weld at 2.5 kN. O—Overlap between the images.</p> Full article ">Figure 13
<p>Speckle map and its corresponding DIC strain map at 2.5 kN.</p> Full article ">Figure 14
<p>Image stitching procedures in Image-J to stitch strain maps (RGB) using reduced pixel coordinates determined from the stitched grey scale DIC speckle images.</p> Full article ">Figure 15
<p>(<b>a</b>) MATLAB and Image-J based strain data stitching procedures; (<b>b</b>) Correlation between the stitched strain data and manually stitched data (right) along a vertical line as shown in the strain map (left) at 2.5 kN.</p> Full article ">Figure 16
<p>(<b>a</b>) ε<sub><span class="html-italic">yy</span></sub> strain map of the FSW (SS-SS) weld along the weld face (WF), weld root (WR), and weld cross-section (CS) at 2.5 kN; (<b>b</b>) defect locations on the WR.</p> Full article ">Figure 17
<p>Strain distribution through the weld region: (<b>a</b>) DIC ROI; (<b>b</b>) strain distribution on the WF and WR along the vertical line shown in (<b>a</b>) at 2.5 kN load; (<b>c</b>) average strain (ε<sub><span class="html-italic">xx</span></sub><sub>,</sub> ε<sub><span class="html-italic">yy</span></sub><sub>,</sub> and ε<sub><span class="html-italic">xy</span></sub>) distributions.</p> Full article ">Figure 18
<p>(<b>a</b>) ROIs on WF and WR for local stress-strain curves; (<b>b</b>) local stress-strain curves of WF and WR.</p> Full article ">Figure 19
<p>(<b>a</b>) Out-of-plane displacement measurements along the vertical line of weld cross-section at 2.5 kN; (<b>b</b>) correction factor applied stress-strain curves of the weld nugget.</p> Full article ">
17 pages, 21336 KiB  
Article
Effect of Temperature on the Corrosion Behavior of Biodegradable AZ31B Magnesium Alloy in Ringer’s Physiological Solution
by Sebastian Feliu, Jr., Lucien Veleva and Federico García-Galvan
Metals 2019, 9(5), 591; https://doi.org/10.3390/met9050591 - 22 May 2019
Cited by 11 | Viewed by 4481
Abstract
In this work, the corrosion behaviors of the AZ31B alloy in Ringer’s solution at 20 °C and 37 °C were compared over four days to better understand the influence of temperature and immersion time on corrosion rate. The corrosion products on the surfaces [...] Read more.
In this work, the corrosion behaviors of the AZ31B alloy in Ringer’s solution at 20 °C and 37 °C were compared over four days to better understand the influence of temperature and immersion time on corrosion rate. The corrosion products on the surfaces of the AZ31B alloys were examined by scanning electron microscopy (SEM), energy dispersive X-ray spectroscopy (EDS) and X-ray diffraction (XRD). Electrochemical impedance spectroscopy (EIS) provided information about the protective properties of the corrosion layers. A significant acceleration in corrosion rate with increasing temperature was measured using mass loss and evolved hydrogen methods. This temperature effect was directly related to the changes in chemical composition and thickness of the Al-rich corrosion layer formed on the surface of the AZ31B alloy. At 20 °C, the presence of a thick (micrometer scale) Al-rich corrosion layer on the surface reduced the corrosion rate in Ringer’s solution over time. At 37 °C, the incorporation of additional Mg and Al compounds containing Cl into the Al-rich corrosion layer was observed in the early stages of exposure to Ringer’s solution. At 37 °C, a significant decrease in the thickness of this corrosion layer was noted after four days. Full article
(This article belongs to the Special Issue Surface Chemistry and Corrosion of Light Alloys)
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<p>BSE (Back scattered electron) images of AZ31B specimen surfaces non-exposed (<b>a</b>) and immersed in Ringer’s solution for 2 days (<b>b</b>,<b>c</b>) and 4 days (<b>d</b>,<b>e</b>) at 20 °C (<b>b</b>,<b>d</b>) and 37 °C (<b>c</b>,<b>e</b>).</p> Full article ">Figure 2
<p>(<b>a</b>,<b>b</b>) BSE images and (<b>c</b>,<b>d</b>) variations in the Al/(Al + Mg) × 100 atomic ratio with the distance, according to quantitative energy dispersive X-ray spectroscopy (EDS) analysis of cross-sections of corrosion layers formed on AZ31B specimens after immersion in Ringer’s solution at 20 °C for (<b>a</b>,<b>c</b>) 2 days and (<b>b</b>,<b>d</b>) 4 days. Scatter bands shown in (<b>c</b>,<b>d</b>) represent the standard deviation of 3 measurements.</p> Full article ">Figure 3
<p>(<b>a</b>,<b>b</b>) BSE images and (<b>c</b>,<b>d</b>) variations in the Al/(Al + Mg) × 100 atomic ratio with the distance, according to quantitative EDS analysis of cross-sections of corrosion layers formed on AZ31B specimens after immersion in Ringer’s solution at 37 °C for (<b>a</b>,<b>c</b>) 2 days and (<b>b</b>,<b>d</b>) 4 days. Scatter bands shown in (<b>c</b>,<b>d</b>) represent the standard deviation of 3 measurements.</p> Full article ">Figure 4
<p>(<b>a</b>) BSE image of the cross-section of a pit formed on the AZ31B specimen after immersion in Ringer’s solution at 37 °C for 4 d. (<b>b</b>) Variations in (<b>b</b>) the Al/(Al + Mg) × 100 atomic ratio and (<b>c</b>) Cl atomic percentage with the distance according to quantitative EDS analysis.</p> Full article ">Figure 5
<p>Optical microscopy images of cross-sections of AZ31B specimens after immersion in Ringer’s solution for 4 days: at 20 °C (<b>a</b>) and (<b>b</b>) at 37 °C.</p> Full article ">Figure 6
<p>3D analysis of pits in AZ31B specimen surfaces immersed in Ringer’s solution for 4 days, after removal of the corrosion products: at (<b>a</b>) 20 °C and (<b>b</b>) 37 °C.</p> Full article ">Figure 7
<p>Low-angle XRD patterns of AZ31B specimens after immersion in Ringer’s solution at 20 °C and 37 °C, for 2 days and 4 days.</p> Full article ">Figure 8
<p>Variations in corrosion of AZ31B specimens over 4 days of immersion in Ringer’s solution, as a function of temperature: (<b>a</b>) Hydrogen evolution volume; (<b>b</b>) corresponding corrosion rates, <span class="html-italic">P</span><sub>H</sub>.</p> Full article ">Figure 9
<p>Variations in corrosion rates (mm·y<sup>−1</sup>) of AZ31B specimens as a function of temperature over 4 days of immersion in Ringer’s solution, obtained from mass loss measurements.</p> Full article ">Figure 10
<p>Nyquist plots of AZ31B specimens immersed for 1 h, 1 day, and 4 days in Ringer’s solution at 20 °C and 37 °C.</p> Full article ">Figure 11
<p>The equivalent circuit used for fitting experimental electrochemical impedance spectroscopy (EIS) plots of AZ31B specimens immersed in Ringer’s solution at 20 °C and 37 °C.</p> Full article ">Figure 12
<p>Variations in charge transfer resistance (<span class="html-italic">R</span><sub>t</sub>) values as a function of immersion time, obtained from fitting of the EIS plots of AZ31B specimens immersed in Ringer’s solution at 20 °C and 37 °C.</p> Full article ">
18 pages, 5877 KiB  
Review
Effects of Different Parameters on Initiation and Propagation of Stress Corrosion Cracks in Pipeline Steels: A Review
by M.A. Mohtadi-Bonab
Metals 2019, 9(5), 590; https://doi.org/10.3390/met9050590 - 22 May 2019
Cited by 71 | Viewed by 8568
Abstract
The demand for pipeline steels has increased in the last several decades since they were able to provide an immune and economical way to carry oil and natural gas over long distances. There are two important damage modes in pipeline steels including stress [...] Read more.
The demand for pipeline steels has increased in the last several decades since they were able to provide an immune and economical way to carry oil and natural gas over long distances. There are two important damage modes in pipeline steels including stress corrosion cracking (SCC) and hydrogen induced cracking (HIC). The SCC cracks are those cracks which are induced due to the combined effects of a corrosive environment and sustained tensile stress. The present review article is an attempt to highlight important factors affecting the SCC in pipeline steels. Based on a literature survey, it is concluded that many factors, such as microstructure of steel, residual stresses, chemical composition of steel, applied load, alternating current (AC) current and texture, and grain boundary character affect the SCC crack initiation and propagation in pipeline steels. It is also found that crystallographic texture plays a key role in crack propagation. Grain boundaries associated with {111}∥rolling plane, {110}∥rolling plane, coincidence site lattice boundaries and low angle grain boundaries are recognized as crack resistant paths while grains with high angle grain boundaries provide easy path for the SCC intergranular crack propagation. Finally, the SCC resistance in pipeline steels is improved by modifying the microstructure of steel or controlling the texture and grain boundary character. Full article
(This article belongs to the Special Issue Corrosion and Protection of Metals)
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<p>Formation of rust by the oxidation of iron to ferrous ions [<a href="#B13-metals-09-00590" class="html-bibr">13</a>]. Reproduced with permission from [<a href="#B13-metals-09-00590" class="html-bibr">13</a>], Noria Corporation and Machinery Lubrication, 2018.</p> Full article ">Figure 2
<p>Effective factors for the stress corrosion cracking (SCC) crack initiation in pipeline steels.</p> Full article ">Figure 3
<p>(<b>a</b>,<b>b</b>) SCC cracks in the cross section of X80 pipeline steel after SSRT test in high pH solution [<a href="#B22-metals-09-00590" class="html-bibr">22</a>]. Reproduced with permission from [<a href="#B22-metals-09-00590" class="html-bibr">22</a>], Springer Nature, 2014.</p> Full article ">Figure 4
<p>Transgranular type of cracks from corrosion pits on (<b>a</b>) internal surface, and (<b>b</b>) external surface [<a href="#B26-metals-09-00590" class="html-bibr">26</a>]. Reproduced with permission from [<a href="#B26-metals-09-00590" class="html-bibr">26</a>], Elsevier, 2014.</p> Full article ">Figure 5
<p>(<b>a</b>) Pit formation from an inclusion in X65 pipeline steel, (<b>b</b>) SCC crack initiation from a pit and (<b>c</b>) EDS analysis on the inclusion showing the existence of Na, Mg, Al, Si, P, S, Cl, K, and Ca elements [<a href="#B27-metals-09-00590" class="html-bibr">27</a>]. Reproduced with permission from [<a href="#B27-metals-09-00590" class="html-bibr">27</a>], Springer Nature, 2009.</p> Full article ">Figure 6
<p>Effect of Δ<span class="html-italic">K</span> upon SCC velocity of pipeline steel exposed to carbonate, bicarbonate solution [<a href="#B28-metals-09-00590" class="html-bibr">28</a>]. Reproduced with permission from [<a href="#B28-metals-09-00590" class="html-bibr">28</a>], Elsevier, 2017.</p> Full article ">Figure 7
<p>Initiation of transgranular SCC cracks from the surface oxide: (<b>a</b>) region 1, and (<b>b</b>) region 2 [<a href="#B32-metals-09-00590" class="html-bibr">32</a>]. Reproduced with permission from [<a href="#B32-metals-09-00590" class="html-bibr">32</a>], Elsevier, 2003.</p> Full article ">Figure 8
<p>(<b>a</b>,<b>b</b>) Crack tip morphology after the pipeline specimens subjected to constant load for 7 days in different soil solutions [<a href="#B33-metals-09-00590" class="html-bibr">33</a>]. Reproduced with permission from [<a href="#B33-metals-09-00590" class="html-bibr">33</a>], Elsevier, 2007.</p> Full article ">Figure 9
<p>SEM image of fracture surfaces of X80 pipeline steel in air and NS4 solution in (<b>a</b>) in air, (<b>b</b>) 0 A/m<sup>2</sup>, (<b>c</b>) 5 A/m<sup>2</sup>, (<b>d</b>) 10 A/m<sup>2</sup>, (<b>e</b>) 30 A/m<sup>2</sup> and (<b>f</b>) 50 A/m<sup>2</sup> [<a href="#B42-metals-09-00590" class="html-bibr">42</a>]. Reproduced with permission from [<a href="#B42-metals-09-00590" class="html-bibr">42</a>], Elsevier, 2017.</p> Full article ">Figure 10
<p>The SCC model for L360NS pipeline steel in sulfur melting cladding condition [<a href="#B9-metals-09-00590" class="html-bibr">9</a>]. Reproduced with permission from [<a href="#B9-metals-09-00590" class="html-bibr">9</a>], Elsevier, 2017.</p> Full article ">Figure 11
<p>Electron backscatter diffraction (EBSD) map of the SCC crack propagation in X65 pipeline steel [<a href="#B12-metals-09-00590" class="html-bibr">12</a>]. Reproduced with permission from [<a href="#B12-metals-09-00590" class="html-bibr">12</a>], Elsevier, 2009.</p> Full article ">
11 pages, 3951 KiB  
Article
Influence of Niobium or Molybdenum Addition on Microstructure and Tensile Properties of Nickel-Chromium Alloys
by Marisa Aparecida de Souza, Bárbara de Oliveira Fiorin, Tomaz Manabu Hashimoto, Ana Paula Rosifini, Carlos Angelo Nunes, Carlos Antônio Reis Pereira Baptista and Alfeu Saraiva Ramos
Metals 2019, 9(5), 589; https://doi.org/10.3390/met9050589 - 22 May 2019
Cited by 1 | Viewed by 3937
Abstract
This work discusses on influence of niobium or molybdenum addition on microstructure and tensile properties of NiCr-based dental alloys. In this regard, the Ni-24Cr-8Nb, Ni-22Cr-10Nb and Ni-20Cr-12Nb (wt. %) alloys produced by arc melting process. To compare the typical Ni-22Cr-10Mo dental alloy was [...] Read more.
This work discusses on influence of niobium or molybdenum addition on microstructure and tensile properties of NiCr-based dental alloys. In this regard, the Ni-24Cr-8Nb, Ni-22Cr-10Nb and Ni-20Cr-12Nb (wt. %) alloys produced by arc melting process. To compare the typical Ni-22Cr-10Mo dental alloy was also produced. These ternary alloys were analyzed by chemical analyses, X-ray diffraction (XRD), scanning electron microscopy (SEM), electron dispersive spectrometry (EDS), thermogravimetric analysis (TG), Vickers hardness and tensile tests. Although the mass losses of the samples during arc melting, the optical emission spectrometry showed that the initial compositions were kept. The Ni-22Cr-10Mo alloy produced a matrix of Niss (ss—solid solution), whereas Ni3Nb disperse in a Niss matrix was also identified in Ni-Cr-Nb alloys. Excepting for the Ni-22Cr-10Nb alloy with mass gain of 0.23%, the as-cast Ni-Cr alloys presented mass gains close to 0.4% after heating up to 1000 °C under synthetic airflow. The hardness values, the modulus of elasticity, yield strength and ultimate tensile strength have enhanced, whereas the ductility was reduced with increasing niobium addition of up to 12 wt.-%.The Ni-22Cr-10Mo alloy presented an intergranular fracture mechanism containing deep dimples and quasi-cleavage planes, whereas the shallow dimples were identified on fracture surface of the as-cast Nb-richer Ni-Cr alloys due to the presence of higher Ni3Nb amounts. Full article
(This article belongs to the Special Issue Numerical Modelling and Simulation of Metal Processing)
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Graphical abstract
Full article ">Figure 1
<p>XRD patterns of Ni-Cr alloys evaluated in this work: (<b>a</b>) Commercial Ni-Cr alloy, (<b>b</b>) Ni-22Cr-10Mo alloy and (<b>c</b>) Ni-20Cr-12Nb alloy.</p> Full article ">Figure 2
<p>SEM images of the (<b>a</b>) commercial Ni-Cr alloy and (<b>b</b>) as-cast Ni-22Cr-10Mo alloy without a chemical attack.</p> Full article ">Figure 3
<p>SEM images of the as-cast (<b>a</b>,<b>d</b>) Ni-24Cr-8Nb, (<b>b</b>,<b>e</b>) Ni-22Cr-10Nb and (<b>c</b>,<b>f</b>) Ni-20Cr-12Nb alloys: (<b>a</b>,<b>c</b>) Without chemical attack and (<b>b</b>,<b>d</b>,<b>f</b>) after chemical attack.</p> Full article ">Figure 4
<p>Representative thermogravimetric curves of the commercial Ni-Cr alloy, and as-cast Ni-Cr-Nb and Ni-Cr-Mo alloys evaluated in this work.</p> Full article ">Figure 5
<p>Representative tensile σ-ε curves of the as-cast Ni-Cr-Mo and Ni-Cr-Nb alloys investigated in this work.</p> Full article ">Figure 6
<p>SEM fractographies of the as-cast (<b>a</b>,<b>b</b>) Ni-22Cr-10Mo and (<b>c</b>,<b>d</b>) Ni-20Cr-12Nb alloys after tensile tests.</p> Full article ">
17 pages, 5504 KiB  
Article
Optimization and Validation of Sound Absorption Performance of 10-Layer Gradient Compressed Porous Metal
by Fei Yang, Xinmin Shen, Panfeng Bai, Xiaonan Zhang, Zhizhong Li and Qin Yin
Metals 2019, 9(5), 588; https://doi.org/10.3390/met9050588 - 21 May 2019
Cited by 17 | Viewed by 3570
Abstract
Sound absorption performance of a porous metal can be improved by compression and optimal permutation, which is favorable to promote its application in noise reduction. The 10-layer gradient compressed porous metal was proposed to obtain optimal sound absorption performance. A theoretical model of [...] Read more.
Sound absorption performance of a porous metal can be improved by compression and optimal permutation, which is favorable to promote its application in noise reduction. The 10-layer gradient compressed porous metal was proposed to obtain optimal sound absorption performance. A theoretical model of the sound absorption coefficient of the multilayer gradient compressed porous metal was constructed according to the Johnson-Champoux-Allard model. Optimal parameters for the best sound absorption performance of the 10-layer gradient compressed porous metal were achieved by a cuckoo search algorithm with the varied constraint conditions. Preliminary verification of the optimal sound absorber was conducted by the finite element simulation, and further experimental validation was obtained through the standing wave tube measurement. Consistencies among the theoretical data, the simulation data, and the experimental data proved accuracies of the theoretical sound absorption model, the cuckoo search optimization algorithm, and the finite element simulation method. For the investigated frequency ranges of 100–1000 Hz, 100–2000 Hz, 100–4000 Hz, and 100–6000 Hz, actual average sound absorption coefficients of optimal 10-layer gradient compressed porous metal were 0.3325, 0.5412, 0.7461, and 0.7617, respectively, which exhibited the larger sound absorption coefficients relative to those of the original porous metals and uniform 10-layer compressed porous metal with the same thickness of 20 mm. Full article
(This article belongs to the Special Issue Cellular Metals: Fabrication, Properties and Applications)
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Graphical abstract
Full article ">Figure 1
<p>Comparisons of sound absorption coefficients of the optimal 10-layer gradient compressed porous metals and those of the original porous metal and uniform compressed porous metal.</p> Full article ">Figure 2
<p>The constructed finite element simulation model for 10-layer gradient compressed porous metal. (<b>a</b>) Frame diagram of the system of the standing wave tube measurement; (<b>b</b>) schematic drawing compositions of the 10-layer gradient compressed porous metal.</p> Full article ">Figure 3
<p>The constructed finite element simulation model for the original porous metal and the uniform compressed porous metal.</p> Full article ">Figure 4
<p>Comparisons of the sound absorption coefficients in theory and those in simulation. (<b>a</b>) The optimal 10-layer gradient compressed porous metals for the varied frequency ranges; (<b>b</b>) the original porous metal and uniform compressed porous metal with same thicknesses of 20 mm.</p> Full article ">Figure 5
<p>Schematic diagram of the used CTM2050 universal testing machine for fabrication of the optimal 10-layer gradient compressed porous metals.</p> Full article ">Figure 6
<p>The prepared 10-layer gradient compressed porous metals for the varied frequency ranges. (<b>a</b>) 100–1000 Hz; (<b>b</b>) 100–2000 Hz; (<b>c</b>) 100–4000 Hz; (<b>d</b>) 100–6000 Hz.</p> Full article ">Figure 7
<p>The AWA6128A detector for the standing wave tube measurement. (<b>a</b>) Schematic diagram; (<b>b</b>) actual picture.</p> Full article ">Figure 7 Cont.
<p>The AWA6128A detector for the standing wave tube measurement. (<b>a</b>) Schematic diagram; (<b>b</b>) actual picture.</p> Full article ">Figure 8
<p>Comparisons of the theoretical data, the simulation data, and the experimental data of the sound absorption coefficient of the investigated sound absorbers. (<b>a</b>) Optimal 10-layer gradient compressed porous metals for 100–1000 Hz; (<b>b</b>) optimal 10-layer gradient compressed porous metals for 100–2000 Hz; (<b>c</b>) optimal 10-layer gradient compressed porous metals for 100–4000 Hz; (<b>d</b>) optimal 10-layer gradient compressed porous metals for 100–6000 Hz; (<b>e</b>) uniform compressed porous metal with the thickness of 20 mm; (<b>f</b>) original porous metal with the thickness of 20 mm.</p> Full article ">Figure 8 Cont.
<p>Comparisons of the theoretical data, the simulation data, and the experimental data of the sound absorption coefficient of the investigated sound absorbers. (<b>a</b>) Optimal 10-layer gradient compressed porous metals for 100–1000 Hz; (<b>b</b>) optimal 10-layer gradient compressed porous metals for 100–2000 Hz; (<b>c</b>) optimal 10-layer gradient compressed porous metals for 100–4000 Hz; (<b>d</b>) optimal 10-layer gradient compressed porous metals for 100–6000 Hz; (<b>e</b>) uniform compressed porous metal with the thickness of 20 mm; (<b>f</b>) original porous metal with the thickness of 20 mm.</p> Full article ">Figure 9
<p>Distributions of compression ratio of optimal 10-layer gradient compressed porous metals.</p> Full article ">Figure 10
<p>Distribution of structural parameters of the optimal 10-layer gradient compressed porous metals along the thickness direction. (<b>a</b>) Porosity; (<b>b</b>) static flow resistivity.</p> Full article ">
18 pages, 987 KiB  
Article
A Prediction Model for Internal Cracks during Slab Continuous Casting
by Yiwen Kong, Dengfu Chen, Qiang Liu and Mujun Long
Metals 2019, 9(5), 587; https://doi.org/10.3390/met9050587 - 21 May 2019
Cited by 21 | Viewed by 5669
Abstract
Slab continuous casting internal cracking is a common quality defect in the production process. The ability to predict the quality of each continuous casting product and assess whether it is suitable for hot delivery or needs to be cleaned down will greatly increase [...] Read more.
Slab continuous casting internal cracking is a common quality defect in the production process. The ability to predict the quality of each continuous casting product and assess whether it is suitable for hot delivery or needs to be cleaned down will greatly increase the rolled product rate and reduce the scrap rate and production management cost. According to the quality defects of internal cracks during slab continuous casting and based on the solidification and heat transfer simulations, stress and strain calculations and theoretical analysis of metallurgical processes related to continuous casting combined with an abnormal casting event expert system, the internal crack generation index of the slice unit is used to predict the crack occurrence rating of each sized slab. Moreover, the internal crack prediction model for the slab is successfully developed and applied in a domestic steel mill. The accuracy of the model prediction reached 86.85%. This method achieved the organic combination of theoretical analysis and an expert system and provides an important theoretical tool for the prediction of crack quality defects in slab continuous casting; the method can be applied in slab continuous casting production. Full article
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<p>Schematic of space-time dispersion of the internal crack generation index of the slice unit in the slab casting process.</p> Full article ">Figure 2
<p>The improved BP algorithm flow.</p> Full article ">Figure 3
<p>Model software function.</p> Full article ">Figure 4
<p>The internal crack generation index distribution at the beginning stage of casting.</p> Full article ">
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