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19 pages, 3364 KiB  
Article
Determination of Constrained Modulus of Granular Soil from In Situ Tests—Part 2 Application
by K. Rainer Massarsch
Geotechnics 2024, 4(2), 636-654; https://doi.org/10.3390/geotechnics4020034 - 14 Jun 2024
Viewed by 800
Abstract
The paper demonstrates how the concepts presented in the companion paper: “Determination of Constrained Modulus of Granular Soil from In Situ Tests—Part 1 Analyses” can be applied in practice. A settlement design based on the tangent modulus method is described. Extensive in situ [...] Read more.
The paper demonstrates how the concepts presented in the companion paper: “Determination of Constrained Modulus of Granular Soil from In Situ Tests—Part 1 Analyses” can be applied in practice. A settlement design based on the tangent modulus method is described. Extensive in situ tests were performed on a well-documented test site consisting of sand with silt and clay layers. The field tests comprised different types of penetration tests, such as the cone penetration test, the flat dilatometer, and the seismic down-hole test. The modulus number and the constrained tangent modulus were derived from the cone penetration test with pore water pressure measurement and the flat dilatometer test. In addition, the shear wave speed was determined from two seismic down-hole tests, from which the small-strain shear modulus could be evaluated. The constrained modulus obtained from the cone penetration test with pore water pressure measurement (CPTU) and the flat dilatometer (DMT) was compared with that from the seismic down-hole tests. The importance of the stress history on the constrained modulus was demonstrated. The range of modulus numbers, derived from different in situ tests, compares favorably with empirical values reported in the literature. Full article
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Figure 1
<p>Soil sample from borehole A3 and SPT <span class="html-italic">N</span>-values and cone resistance, <span class="html-italic">q</span><sub>c</sub>, adapted from [<a href="#B31-geotechnics-04-00034" class="html-bibr">31</a>].</p>
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<p>Results of CPTU: (<b>a</b>) cone resistance, (<b>b</b>) sleeve resistance, (<b>c</b>) hydrostatic and pore water pressure, and (<b>d</b>) soil behavior index.</p>
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<p>Overconsolidation ratio determined from CPTU based on Equations (3) and (4).</p>
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<p>Modulus number derived from CPTU for assuming normally consolidated (<span class="html-italic">NC</span>). The effect of preloading (<span class="html-italic">OC</span>) was considered according to Equation (5).</p>
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<p>Tangent constrained modulus as a function of depth, determined from modulus number, <span class="html-italic">m</span>, according to Equation (9) given in the companion paper.</p>
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<p>Results of DMT: (<b>a</b>) pressure readings, (<b>b</b>) material index, (<b>c</b>) horizontal stress index, and (<b>d</b>) dilatometer modulus.</p>
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<p>Variation in constrained modulus from DMT according to Equation (8).</p>
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<p>The variation in modulus number with depth from DMT derived from the tangent modulus according to Equation (9).</p>
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<p>(<b>a</b>) Shear wave speed as determined from seismic down-hole tests (SCPT and SDMT) and (<b>b</b>) Small-strain shear modulus, <span class="html-italic">G</span><sub>0</sub>, from SCPT and SDMT.</p>
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<p>Relationship between maximum shear modulus, <span class="html-italic">G</span><sub>0</sub>, and modulus number, <span class="html-italic">m</span>, cf. Equation (10) at 0.25% shear strain. Dashed lines <span class="html-italic">σ</span>′<sub>v</sub> = 50 kPa; full line: <span class="html-italic">σ</span>′<sub>v</sub> = 100 kPa. The blue line indicates the modulus number for medium-dense, normally consolidated sand.</p>
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<p>Modulus number derived from seismic tests (SCPTU and SDMT) for normally consolidated (NC) and overconsolidated (OC) conditions.</p>
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<p>Tangent constrained modulus from seismic tests (SCPTU and SDMT) for normally consolidated (NC) and overconsolidated (OC) conditions.</p>
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13 pages, 58262 KiB  
Article
Modelling the Evolution of Phases during Laser Beam Welding of Stainless Steel with Low Transformation Temperature Combining Dilatometry Study and FEM
by Karthik Ravi Krishna Murthy, Fatma Akyel, Uwe Reisgen, Simon Olschok and Dhamini Mahendran
J. Manuf. Mater. Process. 2024, 8(2), 50; https://doi.org/10.3390/jmmp8020050 - 1 Mar 2024
Cited by 1 | Viewed by 1743
Abstract
In this study, the evolution of volume fractions during laser beam welding (LBW) of stainless steel, with a specific focus on incorporating the low transformation temperature (LTT) effect using the dilatometer, has been proposed. The LTT effect refers to the phase transformations that [...] Read more.
In this study, the evolution of volume fractions during laser beam welding (LBW) of stainless steel, with a specific focus on incorporating the low transformation temperature (LTT) effect using the dilatometer, has been proposed. The LTT effect refers to the phase transformations that occur at lower temperatures and lead to the formation of a martensitic microstructure, which will significantly influence the residual stresses and distortion of the welded joints. In this research, the LTT conditions are achieved by varying the Cr and Ni content in the weld seam by varying the weld parameter, including laser power, welding speed and filler wire speed. The dilatometer analysis technique is employed to simulate the thermal conditions encountered during LBW. By subjecting the stainless steel samples to controlled heating and cooling cycles, the kinetics of the volume fractions can be measured using the lever rule and empirical method (KOP and Lee). The phase transformation simulation model is computed by integrating the thermal and metallurgical effects to predict the volume fractions in LBW joints and has been validated using dilatometer results. This provides valuable insight into the relationship between welding parameters and phase transformations in stainless steel with the LTT effect during laser beam welding. Using this relationship, the weld quality can be improved by reducing the residual stresses and distortion. Full article
(This article belongs to the Special Issue Advanced Joining Processes and Techniques 2023)
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<p>(<b>a</b>) Welding setup at Welding and Joining Institute, RWTH Aachen; (<b>b</b>) thermocouple setup for temperature measurement; (<b>c</b>) weld cross section for Cr 13%; (<b>d</b>) weld cross section for Cr 14%; (<b>e</b>) weld cross section for Cr 15% [<a href="#B25-jmmp-08-00050" class="html-bibr">25</a>].</p>
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<p>(<b>a</b>) Dilatometer sample preparation and (<b>b</b>) dilation curves for Cr 13%, Cr14% and Cr15%.</p>
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<p>Extrapolated dilation curves during cooling for Cr 13%, Cr 14% and Cr 15%.</p>
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<p>Combined spherical and conical heat source model.</p>
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<p>Transient thermal model with temperature distribution along the plate.</p>
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<p>Phase transformation model (martensite).</p>
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<p>Volume fraction for Cr 13%, Cr14% and Cr15% using lever rule.</p>
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<p>Volume fraction for Cr 13%, Cr14% and Cr15% using lattice parameter, thermal expansion and carbon content.</p>
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<p>Austenitic, martensitic and average atomic volume for Cr 13%, Cr14% and Cr15% using lattice parameter, thermal expansion and carbon content.</p>
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<p>Experimental (TE) and simulation (TS) temperature distribution comparison.</p>
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<p>Volume fractions comparison of simulation, empirical and graphical, for Cr13%, Cr14% and Cr15%.</p>
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23 pages, 27417 KiB  
Article
Comparison of the Piezocone Penetrometer (CPTU) and Flat Dilatometer (DMT) Methods for Landslide Characterisation
by Kristijan Grabar, Jasmin Jug, Anja Bek and Stjepan Strelec
Geosciences 2024, 14(3), 64; https://doi.org/10.3390/geosciences14030064 - 26 Feb 2024
Cited by 1 | Viewed by 1593
Abstract
The increasing occurrence of landslides worldwide causes many human casualties and huge socio-economic losses. Therefore, the fastest and most accurate characterisation of landslides is important. The objective of this study is to compare how well the flat dilatometer (DMT) test and the piezocone [...] Read more.
The increasing occurrence of landslides worldwide causes many human casualties and huge socio-economic losses. Therefore, the fastest and most accurate characterisation of landslides is important. The objective of this study is to compare how well the flat dilatometer (DMT) test and the piezocone penetration (CPTU) test can find the depth of a sliding zone. Inclinometers were used to measure horizontal changes in the soil to ensure the depth of the sliding zone was correct. The coincidence of the results of in situ static probes, and the displacements of the inclinometers is a sure confirmation of the depth of the sliding zone. In the example of Bedekovčina and Kravarsko landslides, in situ static probes were used to obtain values of input parameters on the sliding zone for parametric sensitivity analysis of parameters. Sensitivity analysis was performed by plotting the relationship between the above parameters and the vertical effective stress σ′vo on the sliding zone. The sensitivity analysis of the parameters of 11 tested samples shows that for the parameters of the obtained DMT probe, a higher sensitivity of the parameters is obtained, closer to the values concerning the expected range, and a minor standard deviation. The parameter Kd obtained by dilatometer probing is the best indicator of the depth of the sliding zone. The literature value Kd = 1.8–2.0 on the sliding zone in this paper is extended to the range Kd = 1.8–2.5, and its detection sensitivity is influenced by over-consolidation in shallow soil layers. In general, the research results show that the dilatometer probe has an advantage over the piezocone penetrometer test for the needs of landslide characterisation. Full article
(This article belongs to the Topic Geotechnics for Hazard Mitigation)
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<p>Locations of investigated landslides [<a href="#B22-geosciences-14-00064" class="html-bibr">22</a>,<a href="#B23-geosciences-14-00064" class="html-bibr">23</a>].</p>
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<p>Investigation flow chart.</p>
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<p>Geological profile of the wider area around the investigation location in Bedekovčina [<a href="#B20-geosciences-14-00064" class="html-bibr">20</a>].</p>
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<p>Different clay samples from the area of the Bedekovčina landslide.</p>
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<p>Plan view of in-situ investigations on the Bedekovčina landslide.</p>
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<p>Situation plan of in-situ investigations on landslides in Kravarsko.</p>
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<p>Groups of <span class="html-italic">c<sub>u</sub></span> profiles measured at the geotechnical test site of soft clays, Bothkennar in Great Britain. The tags are: FV—downhole wing probe; TC—triaxle compression (CKoUC); DSS—triaxle simple shear; TE—triaxle extension [<a href="#B35-geosciences-14-00064" class="html-bibr">35</a>].</p>
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<p>Self-propelled machine for hydraulic pressing of CPTU probe, Pagani TG 63-150.</p>
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<p>Detection of the sliding zone based on the index <span class="html-italic">K<sub>d</sub></span> [<a href="#B41-geosciences-14-00064" class="html-bibr">41</a>].</p>
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<p>Results of DMT-1 investigation—Bedekovčina landslide.</p>
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<p>Results of for DMT-2 investigation—Bedekovčina landslide.</p>
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<p>Overlay of research results obtained by DMT-1 and DMT-2 sounding—Bedekovčina landslide.</p>
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<p>Results of DMT investigation for <span class="html-italic">p</span><sub>0</sub>, <span class="html-italic">p</span><sub>1</sub>, and interpreted data for <span class="html-italic">E<sub>D</sub>:</span> (<b>a</b>) KL8; (<b>b</b>) KL9; (<b>c</b>) KL10.</p>
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<p>Results of DMT investigation on the Kravarsko landslides.</p>
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<p>Prognostic profiles and the analysis of the obtained data by depth in the first group of the investigation area in Bedekovčina: (<b>a</b>) undrained shear strength <span class="html-italic">c<sub>u</sub></span>; (<b>b</b>) over-consolidation ratio <span class="html-italic">OCR</span>; (<b>c</b>) compressibility modulus <span class="html-italic">M</span>; (<b>d</b>) horizontal stress index <span class="html-italic">K<sub>d</sub></span>.</p>
Full article ">Figure 15 Cont.
<p>Prognostic profiles and the analysis of the obtained data by depth in the first group of the investigation area in Bedekovčina: (<b>a</b>) undrained shear strength <span class="html-italic">c<sub>u</sub></span>; (<b>b</b>) over-consolidation ratio <span class="html-italic">OCR</span>; (<b>c</b>) compressibility modulus <span class="html-italic">M</span>; (<b>d</b>) horizontal stress index <span class="html-italic">K<sub>d</sub></span>.</p>
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<p>Comparison diagrams of in situ probing results with the obtained values of horizontal displacements on inclinometers: (<b>a</b>) Group 1 data <span class="html-italic">K<sub>d</sub></span>—horizontal displacements; (<b>b</b>) Group 3 data <span class="html-italic">K<sub>d</sub>—</span>horizontal displacements; (<b>c</b>) Group 3 data <span class="html-italic">c<sub>u</sub></span>—horizontal displacements.</p>
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<p>Geotechnical soil profile for KL10 landslide on road DC31 in Kravarsko.</p>
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<p>The relationship between the horizontal stress index <span class="html-italic">K<sub>d</sub></span> and the vertical effective stress <span class="html-italic">σ’<sub>vo</sub></span> on the detected sliding zones.</p>
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<p>The relationship between the undrained shear strength <span class="html-italic">c<sub>u</sub></span> and vertical effective stresses <span class="html-italic">σ’<sub>vo</sub></span> on the detected sliding zones.</p>
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<p>The relationship between the over-consolidation ratio <span class="html-italic">OCR</span> and vertical effective stresses <span class="html-italic">σ′<sub>vo</sub></span> on the detected sliding zones.</p>
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<p>The relationship between the compressibility modulus <span class="html-italic">M</span> and vertical effective stresses <span class="html-italic">σ’<sub>vo</sub></span> on the detected sliding zones.</p>
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17 pages, 9363 KiB  
Article
Microscopic Investigation for Experimental Study on Transverse Cracking of Ti-Nb Containing Micro-Alloyed Steels
by Serkan Turan, Hossam Shafy and Heinz Palkowski
Materials 2024, 17(4), 900; https://doi.org/10.3390/ma17040900 - 15 Feb 2024
Cited by 1 | Viewed by 902
Abstract
The influence of Ti on the behavior of hot ductility was examined in four different Ti-containing micro-alloyed steels with a constant content of Nb. Thermomechanical investigations using a dilatometer were carried out to simulate the conditions during casting and cooling in the strand [...] Read more.
The influence of Ti on the behavior of hot ductility was examined in four different Ti-containing micro-alloyed steels with a constant content of Nb. Thermomechanical investigations using a dilatometer were carried out to simulate the conditions during casting and cooling in the strand of a continuous caster with temperatures in the range of 650–1100 °C, strain rates of 0.01 s−1 and 0.001 s−1, and reheating rates between 60 and 180 Kmin−1. To understand the fracture mechanism, optical (LOM) and scanning electron microscopy (SEM), elemental analysis via energy dispersive X-ray spectroscopy (EDX), MatCalc “Scheil–Gulliver” calculations, and precipitation kinetics calculations were carried out for the critical conditions, showing low hot ductility between Ar3 and Ae3 temperatures and a brittle to ductile transition temperature at 900 °C. The existence of TiNb(CN), thin ferrite formation, and grain boundary sliding (GBs) due to limited dynamic recrystallization (DRX) has been documented and discussed. As a result, the reheating rate has no sufficient effect on the ductility. The existence of Nb-rich TiNb(CN) of sizes below ~1 μm triggers brittle fracture by increasing the frequency of micro-voids around grain boundaries. It can be stated that if the conditions in the hot ductility trough are avoided, the addition of Ti and high strain support minimize the risk of crack formation. Full article
(This article belongs to the Section Metals and Alloys)
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Figure 1
<p>DIL 805 T/A hot tensile testing system: (<b>a</b>) installation of tensile test; (<b>b</b>) thermomechanical treatment; (<b>c</b>) sample dimensions in mm (1.8 mm thickness).</p>
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<p>Thermomechanical cycle used.</p>
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<p>Scheil–Gulliver calculations for Ti-Nb-containing steels: (<b>a</b>) A1; (<b>b</b>) A2; (<b>c</b>) A3.</p>
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<p>Hot ductility by reduction of area vs. temperature: (<b>a</b>) 0.001 s<sup>−1</sup>, 60 Kmin<sup>−1</sup>; (<b>b</b>) 0.01 s<sup>−1</sup>, 60 Kmin<sup>−1</sup>; (<b>c</b>) 0.001 s<sup>−1</sup>, 180 Kmin<sup>−1</sup>; (<b>d</b>) 0.01 s<sup>−1</sup>, 180 Kmin<sup>−1</sup>.</p>
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<p>SEM micrographs of the fracture surface (parallel to the tensile force): (<b>a</b>) A0 at 700 °C, 0.001 s<sup>−1</sup>, 60 Kmin<sup>−1</sup> (RA = 19%); (<b>b</b>) A0 at 700 °C, 0.01 s<sup>−1</sup> 60 Kmin<sup>−1</sup> (RA = 46%); (<b>c</b>) A3 at 700 °C, 0.001 s<sup>−1</sup>, 60 Kmin<sup>−1</sup> (RA = 19%); (<b>d</b>) A3 at 700 °C, 0.01 s<sup>−1</sup>, 60 Kmin<sup>−1</sup> (RA = 36%).</p>
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<p>SEM micrographs near the fracture surface (perpendicular to the tensile force)—ferrite formation at GBs (AF and PF represent acicular and polygonal ferrite morphologies, respectively): (<b>a</b>) A0 (RA = 19%); (<b>b</b>) A1 (RA = 19%); (<b>c</b>) A2 (RA = 15%); (<b>d</b>) A3 (RA = 23%).</p>
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<p>SEM micrographs near the fracture surface (perpendicular to the tensile force) of A3 at 700 °C, 0.01 s<sup>−1</sup>, 180 Kmin<sup>−1</sup> (RA = 33%): (<b>a</b>) SEM overview image; (<b>b</b>) phase fraction of secondary precipitates (via MatCalc); (<b>c</b>) SEM image of micro-crack on GB induced by TiNb(CN); (<b>d</b>) EDX analysis.</p>
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<p>LOM micrographs near the fracture surface (perpendicular to the tensile force): DIF on GBs in (<b>a</b>) A0 (RA = 17%); (<b>b</b>) A3 (RA = 18%).</p>
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<p>SEM micrographs of the fracture surface (perpendicular to the tensile force) for A2 at 750 °C, 0.001 s<sup>−1</sup>, 180 Kmin<sup>−1</sup> (RA = 16%): (<b>a</b>) secondary precipitates and micro-voids; (<b>b</b>) EDX analysis.</p>
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<p>SEM micrographs near the fracture surface (perpendicular to the tensile force) for A3 at 750 °C, 0.001 s<sup>−1</sup>, 180 Kmin<sup>−1</sup> (RA = 19%): (<b>a</b>) primary and secondary TiNb(CN) clusters; (<b>b</b>) secondary TiNb(CN) cluster and micro-void on GB; (<b>c</b>) EDX for Spec. 1; (<b>d</b>) EDX for Spec. 2; (<b>e</b>) EDX for Spec. 3.</p>
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<p>Calculations for phase fraction and mean radius via MatCalc at 850 °C, 0.001 s<sup>−1</sup>, 180 Kmin<sup>−1</sup>: primary and secondary precipitates for (<b>a</b>) A1; (<b>b</b>) A2; (<b>c</b>) A3.</p>
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<p>SEM micrographs near the fracture surface (perpendicular to the tensile force) for A1 at 850 °C, 0.001 s<sup>−1</sup>, 180 Kmin<sup>−1</sup> (RA = 30%): (<b>a</b>) secondary TiNb(CN) on GB, micro-voids, GB sliding, and V-shape pile-up of secondary TiNb(CN)s; (<b>b</b>) EDX analysis.</p>
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<p>SEM micrograph near the fracture surface (perpendicular to the tensile force) for A2 at 850 °C, 0.001 s<sup>−1</sup>, 180 Kmin<sup>−1</sup> (RA = 34%): (<b>a</b>) secondary TiNb(CN) on GB, GB sliding and micro-defect, zig zag-formed secondary TiNb(CN) on secondarily formed GB; (<b>b</b>) EDX analysis.</p>
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<p>SEM micrograph near the fracture surface (perpendicular to the tensile force) of A3 at 850 °C, 0.001 s<sup>−1</sup>, 180 Kmin<sup>−1</sup> (RA = 31%): (<b>a</b>) secondary TiNb(CN) on GB, GB sliding with lateral slip and micro-voids; (<b>b</b>) EDX analysis.</p>
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<p>SEM micrographs of the fracture surface (parallel to the tensile force) at 900 °C, 0.001 s<sup>−1</sup>, and 0.01 s<sup>−1</sup>: (<b>a</b>) A0 at 900 °C, 0.001 s<sup>−1</sup>, 180 Kmin<sup>−1</sup> (RA = 38%); (<b>b</b>) A0 at 900 °C, 0.01 s<sup>−1</sup> 180 Kmin<sup>−1</sup> (RA = 56%); (<b>c</b>) A3 at 900 °C, 0.001 s<sup>−1</sup>, 180 Kmin<sup>−1</sup> (RA = 49%); (<b>d</b>) A3 at 900 °C, 0.01 s<sup>−1</sup>, 180 Kmin<sup>−1</sup> (RA = 68%).</p>
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<p>SEM micrograph near the fracture surface (perpendicular to the tensile force) for A0 at 900 °C, 0.001 s<sup>−1</sup>, 180 Kmin<sup>−1</sup> (RA = 37%): (<b>a</b>) secondary precipitates of NbC &lt; 500 nm dimension accumulation near micro-void; (<b>b</b>) MatCalc calculation.</p>
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12 pages, 7920 KiB  
Article
Secondary Phase Precipitation in Fe-22Mn-9Al-0.6C Low-Density Steel during Continuous Cooling Process
by Yihao Zhou, Tinghui Man, Jun Wang, Hongshan Zhao and Han Dong
Materials 2024, 17(3), 631; https://doi.org/10.3390/ma17030631 - 28 Jan 2024
Cited by 2 | Viewed by 977
Abstract
Secondary phase precipitation in Fe-22Mn-9Al-0.6C low-density steel was investigated during a continuous cooling process with different cooling rates through a DIL805A thermal expansion dilatometer, and the changes in microstructures and hardness by different cooling rates were discussed. The results showed that the matrix [...] Read more.
Secondary phase precipitation in Fe-22Mn-9Al-0.6C low-density steel was investigated during a continuous cooling process with different cooling rates through a DIL805A thermal expansion dilatometer, and the changes in microstructures and hardness by different cooling rates were discussed. The results showed that the matrix of the Fe-22Mn-9Al-0.6C was composed of austenite and δ-ferrite; moreover, the secondary phases included κ-carbide, β-Mn and DO3 at room temperature. The precipitation temperatures of 858 °C, 709 °C and 495 °C corresponded to the secondary phases B2, κ-carbide and β-Mn, respectively, which were obtained from the thermal expansion curve by the tangent method. When the cooling rate was slow, it had enough time to accommodate C-poor and Al-rich regions in the austenite due to amplitude modulation decomposition. Furthermore, the Al enrichment promoted δ-ferrite formation. Meanwhile, the subsequent formation of κ-carbide and β-Mn occurred through the continuous diffusion of C and Mn into austenite. In addition, the hardness of austenite was high at 0.03 °C/s due to the κ-carbide and β-Mn production and C enrichment, and it was inversely proportional to the cooling rate. It can be concluded that the presence of κ-carbide, DO3 and β-Mn produced at the austenitic/ferrite interface when the cooling rate was below 0.1 °C/s resulted in κ-carbide and β-Mn precipitating hardly at cooling rates exceeding 0.1 °C/s, which provides a guideline for the industrial production of Fe-Mn-Al-C low-density steel in the design of the hot working process. Full article
(This article belongs to the Section Metals and Alloys)
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<p>Lattice diagram of Fe<sub>3-x</sub>Mn<sub>x</sub>AlC (0 ≤ x ≤ 3)carbide in Fe-Mn-Al-C low-density steel.</p>
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<p>Lattice diagram of B2 ordered phase in Fe-Mn-Al-C low-density steel.</p>
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<p>Lattice diagram of DO<sub>3</sub> ordered phase in Fe-Mn-Al-C low-density steel.</p>
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<p>Schematic illustration of the continuous cooling process: (<b>a</b>) thermal expansion test; (<b>b</b>) continuous cooling curve.</p>
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<p>Hardness of Fe-22Mn-9Al-0.6C low-density steel with different cooling rates.</p>
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<p>(<b>a</b>) Thermo-Calc equilibrium state property diagram; (<b>b</b>) Expansion curve.</p>
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<p>OM images of the microstructure under different cooling rates.</p>
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<p>XRD diagram of Fe-22Mn-9Al-0.6C low-density steel at (<b>a</b>) 0.03 °C/s, (<b>b</b>) 0.1 °C/s, (<b>c</b>) 50 °C/s.</p>
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<p>SEM images and corresponding energy spectrum of different precipitates of the steel at cooling rate of 0.03 °C/s: (<b>a</b>) DO<sub>3,</sub> (<b>b</b>) κ-carbide.</p>
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<p>The microstructure of the steel at varying cooling rates.</p>
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13 pages, 40013 KiB  
Article
A Study of Pigment, Adhesive, and Firing Temperature in Pottery Figurines Excavated from the Tomb of Qibi Ming, China
by Yanli Li, Haiqiang Guo, Ke Xiao, Panpan Liu, Xiaolian Chao, Peng Fu, Huiping Xing and Yuhu Li
Molecules 2023, 28(23), 7739; https://doi.org/10.3390/molecules28237739 - 24 Nov 2023
Cited by 2 | Viewed by 1435
Abstract
Some painted pottery figurines were excavated from the tomb of Qibi Ming of the Tang Dynasty. A series of analytical techniques were employed to understand the craftsmanship of these painted pottery figurines. The pigment, cross-section, adhesive, and firing temperature were analyzed using microscopy [...] Read more.
Some painted pottery figurines were excavated from the tomb of Qibi Ming of the Tang Dynasty. A series of analytical techniques were employed to understand the craftsmanship of these painted pottery figurines. The pigment, cross-section, adhesive, and firing temperature were analyzed using microscopy (OM), energy X-ray fluorescence spectrometry (EDX), micro-Raman spectroscopy, pyrolysis–gas chromatography–mass spectrometry (Py-GC/MS), and a dilatometer (DIL). The results demonstrated that the surface of the pigment layers had degraded to different degrees. The pigment particles were litharge, gypsum, malachite, cinnabar, hematite, minium, white lead, and carbon black. The cross-sectional images show that the painted layer of figurines 10-0966 and 10-0678 included a pigment layer and a preparation layer. The preparation layer of both pigments was lead white. Animal glue was used as an adhesive. The firing temperature of the pottery figurines was likely 1080 °C. This study can provide more accurate information with regard to the composition of the raw materials utilized in the making of these artifacts and support the selection of appropriate substances for the purposes of conservation and restoration of the painted pottery figurines. Full article
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<p>Locality map of the tomb of Qibi Ming.</p>
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<p>Photograph of the pottery figurines.</p>
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<p>Optical microscopic images of samples (<b>a</b>) 10-0852, (<b>b</b>) 10-0873, (<b>c</b>) 10-0904, (<b>d</b>) 10-0675, (<b>e</b>) 10-0857, (<b>f</b>) 10-0966, (<b>g</b>) 10-0678, and (<b>h</b>) 10-0888.</p>
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<p>Raman spectra of samples (<b>a</b>) 10-0852, (<b>b</b>) 10-0873, (<b>c</b>) 10-0904, (<b>d</b>) 10-0675, (<b>e</b>) 10-0857, (<b>f</b>) 10-0966, (<b>g</b>) 10-0678, and (<b>h</b>) 10-0888.</p>
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<p>Images of the cross-sections of samples (<b>a</b>) 10-0852, (<b>b</b>) 10-0873, (<b>c</b>) 10-0904, (<b>d</b>) 10-0675, (<b>e</b>) 10-0857, (<b>f</b>) 10-0966, (<b>g</b>) 10-0678, and (<b>h</b>) 10-0888.</p>
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<p>Raman spectra of figurines (<b>a</b>) 10-0966 and (<b>b</b>) 10-0678.</p>
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<p>Chromatogram of sample 10-0904.</p>
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<p>Thermal expansion curve of sample 10-1108.</p>
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<p>Images of the sampling position of samples (<b>a</b>) 10-0852, (<b>b</b>) 10-0873, (<b>c</b>) 10-0904, (<b>d</b>) 10-0675, (<b>e</b>) 10-0857, (<b>f</b>) 10-0966, (<b>g</b>) 10-0678, (<b>h</b>) 10-0888, and (<b>i</b>) 10-1108.</p>
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14 pages, 1751 KiB  
Article
Characterizing and Modeling Transformation-Induced Plasticity in 13Cr-4Ni Welds upon Cooling
by Jean-Benoit Lévesque, Carlo Baillargeon, Daniel Paquet, Jacques Lanteigne and Henri Champliaud
Materials 2023, 16(22), 7166; https://doi.org/10.3390/ma16227166 - 15 Nov 2023
Viewed by 856
Abstract
Dilatometric experiments were conducted with the main purpose of measuring the transformation-induced coefficients of 13% chromium and 4% nickel, which are martensitic stainless steel base and filler materials used for hydraulic turbine manufacturing. To this end, a set of experiments was conducted in [...] Read more.
Dilatometric experiments were conducted with the main purpose of measuring the transformation-induced coefficients of 13% chromium and 4% nickel, which are martensitic stainless steel base and filler materials used for hydraulic turbine manufacturing. To this end, a set of experiments was conducted in a quenching dilatometer equipped with loading capabilities. The measurement system was further improved by means of modified pushrods to allow for the use of specimens with geometries that are compliant with tensile test standards. This improvement allowed for the measurement of the materials’ phases and respective yield strengths. The dataset was further used to determine the relationship between the applied external stress and the martensitic start temperature (Ms) upon cooling. The TRIP coefficient’s K values for both the S41500 steel and E410NiMo filler material were measured at 8.12×105 and 7.11×105, respectively. Additionally, the solid phase transformation model parameters for both the austenitic and martensitic transformation of the filler material were measured. These parameters were then used to model austenitic-phase-transformation kinetics and martensite transformation, including transformation-induced plasticity effects. Good agreement was achieved between the calculation and the experiments. Full article
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<p>Micrographs of the materials under study: (<b>a</b>) UNS S41500 base metal; (<b>b</b>) E410NiMo base metal.</p>
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<p>Dilatation as a function of temperature during the heating and cooling of S41500 and E410NiMo soft martensitic stainless steels.</p>
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<p>Schematic of the test chamber of the quenching dilatometer. The force, F, compressive or tensile, is applied to the sample through the Load cell. The other end of the sample is fixed to the wall.</p>
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<p>TRIP experiments: thermal and loading history.</p>
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<p>Total strain measured during the cooling of the S41500 specimens for various magnitudes of uniaxial stress (compressive and tensile).</p>
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<p>Total strain measured during the cooling of the E410NiMo specimens for various magnitudes of uniaxial stress (compressive and tensile).</p>
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<p>Final transformation-induced strains as a function of the applied stress magnitude for S41500 and E410NiMo.</p>
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<p>Martensitic start transformation as a function of applied stress.</p>
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<p>Comparison of modeled total strains (solid) and experimental total strains data (dashed) for S41500 stainless steel.</p>
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<p>Comparison of modeled total strains (solid) and experimental total strains data (dashed) for E410NiMo stainless steel.</p>
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13 pages, 5549 KiB  
Article
The Study of Hardness Evolution during the Tempering Process of 38MnB5Nb Ultra-High-Strength Hot Stamping Steel: Experimental Analysis and Constitutive Models
by Ping Luo, Xianjun Li, Wenliang Zhang, Zhunli Tan, Minghao Zhang, Kaize Wang, Pengdi Hou, Junjie Liu and Xiao Liang
Metals 2023, 13(10), 1642; https://doi.org/10.3390/met13101642 - 25 Sep 2023
Cited by 1 | Viewed by 1270
Abstract
To elucidate the hardness evolution behaviors for 38MnB5Nb ultra-high-strength hot stamping steel, a series of tempering processes with varying tempering temperatures and times were carried out with a dilatometer. Meanwhile, the hardness of each sample was measured after dilatometer experiments. The results indicated [...] Read more.
To elucidate the hardness evolution behaviors for 38MnB5Nb ultra-high-strength hot stamping steel, a series of tempering processes with varying tempering temperatures and times were carried out with a dilatometer. Meanwhile, the hardness of each sample was measured after dilatometer experiments. The results indicated that the tempering process parameters (including the tempering temperature and time) play an important role in the hardness of the studied steel. The hardness of 38MnB5Nb ultra-high-strength hot stamping steel at the quenched state is about 580 Hv, while it is 240 Hv for the quasi-annealed state. As the tempering time extends, the hardness is decreased sharply at the initial stage; then, the hardness is decreased in a quasi-linear trend with a slight slope; finally, the hardness almost keeps a constant value, which depends on the tempering temperature. In addition, the tempering process has a big effect on the mechanical properties of 38MnB5Nb ultra-high-strength hot stamping steel by increasing the product of the strength and elongation by about 40%. Full article
(This article belongs to the Section Structural Integrity of Metals)
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<p>Schematic graphs of preliminary heat treatments: (<b>a</b>) homogenization processes; (<b>b</b>) quenching processes.</p>
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<p>Microstructure after quenching of 38MnB5Nb steel: (<b>a</b>) SEM; (<b>b</b>) TEM for displaying phase; (<b>c</b>) TEM for displaying dislocation.</p>
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<p>Schematic graph of the tempering processes for the samples with sizes of 10 mm in length, 4 mm in width, and 2 mm in thickness.</p>
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<p>Schematic graph of tempering experiments for samples with sizes of 150 mm in length, 50 mm in width, and 2 mm in thickness.</p>
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<p>Hardness evolution with time during tempering at different temperatures ranging from 300 to 700 °C (<b>a</b>), and at a temperature of 800 °C (<b>b</b>).</p>
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<p>Evolution of hardness ratio with time during tempering for different temperatures between 300 and 700 °C.</p>
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<p>(<b>a</b>) Relationship between <math display="inline"><semantics> <mrow> <mrow> <mrow> <mi mathvariant="normal">ln</mi> <mo stretchy="false">(</mo> <mi mathvariant="normal">ln</mi> </mrow> <mo>⁡</mo> <mrow> <mfrac> <mrow> <mn>1</mn> </mrow> <mrow> <mn>1</mn> <mo>−</mo> <mi>τ</mi> </mrow> </mfrac> <mo stretchy="false">)</mo> </mrow> </mrow> </mrow> </semantics></math> and <math display="inline"><semantics> <mrow> <mrow> <mrow> <mi mathvariant="normal">ln</mi> </mrow> <mo>⁡</mo> <mrow> <mi>t</mi> </mrow> </mrow> </mrow> </semantics></math> with different tempering temperatures; (<b>b</b>) relationship between <math display="inline"><semantics> <mrow> <mrow> <mrow> <mi mathvariant="normal">ln</mi> </mrow> <mo>⁡</mo> <mrow> <mi>D</mi> </mrow> </mrow> </mrow> </semantics></math> and <math display="inline"><semantics> <mrow> <mfrac> <mrow> <mn>1</mn> </mrow> <mrow> <mi mathvariant="normal">R</mi> <mi>T</mi> </mrow> </mfrac> </mrow> </semantics></math> with different tempering temperatures.</p>
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<p>Comparison between the experimental and predicted hardness of the 38MnB5Nb steels after undergoing various tempering processes: (<b>a</b>) tempering at 300 °C; (<b>b</b>) tempering at 400 °C; (<b>c</b>) tempering at 500 °C; (<b>d</b>) tempering at 600 °C; (<b>e</b>) tempering at 700 °C.</p>
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<p>The comparison between measured hardness and predicted hardness.</p>
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<p>SEM microstructure of the samples with the tempering time of 40 s and different temperatures: (<b>a</b>) 500 °C; (<b>b</b>) 600 °C; (<b>c</b>) 700 °C; (<b>d</b>) TEM microstructure of the samples with tempering time of 40 s and temperature of 700 °C; (<b>e</b>) TEM dark field microstructure of the samples with tempering time of 40 s and temperature of 700 °C.</p>
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<p>SEM microstructure of the samples with the tempering time of 40 s and different temperatures: (<b>a</b>) 500 °C; (<b>b</b>) 600 °C; (<b>c</b>) 700 °C; (<b>d</b>) TEM microstructure of the samples with tempering time of 40 s and temperature of 700 °C; (<b>e</b>) TEM dark field microstructure of the samples with tempering time of 40 s and temperature of 700 °C.</p>
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26 pages, 9462 KiB  
Article
New Correlations for the Determination of Undrained Shear, Elastic Modulus, and Bulk Density Based on Dilatometer Tests (DMT) for Organic Soils in the South of Quito, Ecuador
by Jorge Mayanquer, Mariela Anaguano-Marcillo, Nicolás Játiva and Jorge Albuja-Sánchez
Appl. Sci. 2023, 13(15), 8570; https://doi.org/10.3390/app13158570 - 25 Jul 2023
Cited by 1 | Viewed by 1922
Abstract
The Marchetti Dilatometer test is a non-destructive in situ test that can be used to determine the geotechnical properties of soils. This paper presents the results of a study that investigated the correlations between the parameters obtained from the Marchetti Dilatometer test and [...] Read more.
The Marchetti Dilatometer test is a non-destructive in situ test that can be used to determine the geotechnical properties of soils. This paper presents the results of a study that investigated the correlations between the parameters obtained from the Marchetti Dilatometer test and geomechanical parameters for soft soils, mainly organic soils, obtained in the laboratory. The study was conducted in the El Garrochal sector in Southern Quito, Ecuador. The results of the study showed that there are significant correlations between the Marchetti Dilatometer test and the undrained shear strength, modulus of elasticity, and density of soil. The equations that were developed in this study can be used to estimate these geomechanical parameters from the results of the Marchetti Dilatometer test for the South Quito sector, which are valuable for geotechnical engineers to design structures in these types of soils. The equations that were developed in this study can be used to improve the accuracy of the design of these structures. Full article
(This article belongs to the Section Civil Engineering)
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<p>The location where samples were collected.</p>
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<p>(<b>a</b>) Long Year drilling platform, (<b>b</b>) DMT platform, (<b>c</b>) wooden body found in a Shelby tube, and (<b>d</b>) mud found after Shelby’s extraction.</p>
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<p>(<b>a</b>) Extracted Shelby tubes, (<b>b</b>) sample being extracted from a Shelby tube, (<b>c</b>) photographed samples, and (<b>d</b>) stored samples.</p>
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<p>Basic principles of Marchetti Dilatometer: pushing, contact stress A, expansion stress B, pressure C, adapted from [<a href="#B4-applsci-13-08570" class="html-bibr">4</a>].</p>
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<p>Marchetti’s nomogram for estimating soil type and unit weight γ, [<a href="#B26-applsci-13-08570" class="html-bibr">26</a>].</p>
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<p>Plasticity chart with results of each borehole.</p>
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<p>(<b>a</b>) Variation of the LL, with respect to depth; (<b>b</b>) variation of PI, with respect to depth.</p>
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<p>Particle size distribution of samples of all boreholes.</p>
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<p>(<b>a</b>) Organic material retained above the N°200 sieve; (<b>b</b>) fine sand retained above the N°200 sieve.</p>
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<p>Variation of (<b>a</b>) ash content and (<b>b</b>) organic content of each borehole about depth.</p>
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<p>Variation of density values of each borehole about depth.</p>
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<p>Relationship between the specific gravity of the soil and its ash content, using equation 1.</p>
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<p>Relationship between the specific gravity of the soil and its ash content, using equation 2.</p>
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<p>Samples that were not suitable to be molded.</p>
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<p>(<b>a</b>) Variation of <span class="html-italic">C<sub>u</sub></span> and (<b>b</b>) the elasticity modulus of each borehole about depth.</p>
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<p>Variation of (<b>a</b>) material index (<span class="html-italic">I<sub>D</sub></span>), (<b>b</b>) horizontal stress index (<span class="html-italic">K<sub>D</sub></span>), and (<b>c</b>) dilatometer modulus (<span class="html-italic">E<sub>D</sub></span>) of each borehole about depth.</p>
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<p>Results of the DMT plotted on Marchetti’s nomogram.</p>
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<p>Relationship between the <span class="html-italic">I<sub>D</sub></span> and <span class="html-italic">E<sub>D</sub></span> of all collected data.</p>
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<p>Correlation and regression equation to obtain the modulus of elasticity.</p>
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<p>Correlation equation based on test results and depth for modulus of elasticity.</p>
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<p>Comparison between different elasticity modulus values.</p>
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<p>Undrained shear distribution, <span class="html-italic">K<sub>D</sub></span>, and vertical effective stress about depth.</p>
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<p>Relationship between variables (<span class="html-italic">C<sub>u</sub></span>, <span class="html-italic">K<sub>D</sub></span>, and <span class="html-italic">σ′<sub>v</sub></span>).</p>
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<p>Correlation equation based on test results and depth for undrained shear strength.</p>
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<p>Comparison between different undrained shear strengths.</p>
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<p>Correlation and regression between density and I<sub>D</sub> value.</p>
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<p>Correlation equation based on test results and depth for density.</p>
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<p>Comparison between density found by the Marchetti nomogram, density found by the equation in the previous graph, and density found in the laboratory.</p>
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12 pages, 4574 KiB  
Article
An Experimental Investigation of the Solid State Sintering of Cemented Carbides Aiming for Mechanical Constitutive Modelling
by Louise Rosenblad, Hjalmar Staf, Henrik Larsson and Per-Lennart Larsson
Crystals 2023, 13(6), 978; https://doi.org/10.3390/cryst13060978 - 20 Jun 2023
Viewed by 1291
Abstract
The densification of cemented carbides during sintering was studied using an existing constitutive model based on powder particle size and material composition. In the present analysis, we study how well the constitutive model can capture the experimental results of a dilatometer test. Three [...] Read more.
The densification of cemented carbides during sintering was studied using an existing constitutive model based on powder particle size and material composition. In the present analysis, we study how well the constitutive model can capture the experimental results of a dilatometer test. Three experiments were performed, where the only difference was the transition between the debinding and sintering process. From magnetic measurements, it is concluded that the carbon level in the specimen is affected by changes to the experimental setup. It is shown, using parameter adjustments, that the constitutive model is more suited for a certain experimental setup and carbon level, which is a limitation of the model. In order to capture the mechanical behaviour under different experimental conditions, further constitutive modelling relevant to the carbon level is recommended. Full article
(This article belongs to the Section Hybrid and Composite Crystalline Materials)
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<p>The dilatometer used to measure strain during the sintering process.</p>
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<p>Schematics of the three different experimental setups. Observe that the time is not to scale.</p>
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<p>The three experimental strain curves as a function of time after 450 °C was reached.</p>
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<p>Strain rate curves corresponding to the strain curves in <a href="#crystals-13-00978-f003" class="html-fig">Figure 3</a>.</p>
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<p>Experimental results and optimisation of the reference experiment.</p>
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<p>Experimental results and optimisation of the separated debinding step—aired experiment.</p>
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17 pages, 7334 KiB  
Article
Effects of Pressure and Cooling Rates on Crystallization Behavior and Morphology of Isotactic Polypropylene
by Vito Speranza, Rita Salomone and Roberto Pantani
Crystals 2023, 13(6), 922; https://doi.org/10.3390/cryst13060922 - 8 Jun 2023
Cited by 5 | Viewed by 3347
Abstract
Isotactic Polypropylene (iPP) is a widely used polymer due to its excellent mechanical and thermal properties, as well as its chemical resistance. The crystallization behavior of polypropylene is influenced by several factors, such as temperature, cooling rate, and pressure. The effect of pressure [...] Read more.
Isotactic Polypropylene (iPP) is a widely used polymer due to its excellent mechanical and thermal properties, as well as its chemical resistance. The crystallization behavior of polypropylene is influenced by several factors, such as temperature, cooling rate, and pressure. The effect of pressure is significant for both scientific and technological points of view, since in important industrial processing techniques the polymer solidifies under high pressures. In this paper, the study of the effect of pressure on the crystallization kinetics of iPP was conducted using a dilatometer in the pressure range from 100 to 600 bar and under two cooling rates: 0.1 and 1 °C/s. The morphology of the samples was characterized using DSC, optical microscopy, and X-ray diffraction. The results showed that pressure had a larger effect on specific volume changes at higher temperatures (in the melt state) than at lower temperatures (in the solid state). The polymer crystallization, which determined the transition between the melt and solid state, occurred at higher temperatures with increasing pressure. The cooling rate affected the crystallization process, with higher cooling rates leading to crystallization at lower temperatures. The size of the spherulites decreased with increasing cooling rates. The crystallinity evolution curves showed a linear relationship between the crystallization temperature and pressure. The study used a Kolmogoroff–Avrami–Evans model to describe the evolution into isotropic structures, and the predictions of the model accurately described the effect of pressure and cooling rates on the final spherulite radii. Full article
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<p>Thermal and pressure protocol adopted during the specific volume measurements.</p>
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<p>Sketch showing the position of slices adopted for morphology characterization.</p>
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<p>Thermal histories recorded during isobaric cooling test adopting two different cooling conditions.</p>
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<p>Specific volume of iPP (T30g, Montell) measurements at 100, 200, 400, and 600 bar with piston-die dilatometer with 0.1 °C/s. Measurement with Gnomix dilatometer at 100 bar with a constant cooling rate of 0.02 °C/s are reported for comparison.</p>
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<p>Specific volume of iPP (T30g, Montell) measured in isobaric mode at 100, 200, 400, and 600 bar with piston-die dilatometer with 1 °C/s.</p>
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<p>Specific volume of iPP measurements at 600 bar with 0.1 and 1 °C/s.</p>
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<p>Thermograms collected during heating of the samples solidified under different pressure and temperature histories. In the topmost of the plot are represented thermograms of the samples obtained by cooling the dilatometer passively (namely, characterized by a cooling rate of 0.1 °C/s) whereas in the bottommost of the plot are represented thermograms of the samples obtained by cooling the dilatometer actively (namely, characterized by a cooling rate of 1 °C/s). The vertical dotted lines identify the range of temperatures from the peak to the end of the melting process.</p>
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<p>WAXS patterns for samples obtained at 100 and 600 bar with cooling rate of 0.1 °C/s (<b>a</b>,<b>b</b>), and with cooling rate of 1 °C/s (<b>c</b>,<b>d</b>).</p>
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<p>Polarized optical micrograph of the sample obtained from isobaric tests.</p>
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<p>Specific volume of iPP measurement at 100 bar with 0.1 °C/s. Specific volumes of polymer in melt and solid state are represented as solid and dashed lines, respectively.</p>
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<p>Crystallinity evolution as evaluated from the dilatometer tests adopting Equation (1): (<b>a</b>) cooling rate of 0.1 °C/s, (<b>b</b>) cooling rate of 1 °C/s.</p>
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<p>Dependence of crystallization temperature upon pressure: (<b>a</b>) cooling rate of 0.1 °C/s, (<b>b</b>) cooling rate of 1 °C/s.</p>
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<p>X-ray diffractograms of samples obtained at 100 and 600 bar pressure: (<b>a</b>) cooling rate of 0.1 °C/s, (<b>b</b>) cooling rate of 1 °C/s.</p>
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<p>Phase content of samples obtained at 100 and 600 bar pressure with cooling rate of 0.1 °C/s and 1 °C/s.</p>
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<p>Comparison of experimental relative crystallinity evolutions (solid lines) with model predictions (dotted lines): (<b>a</b>) cooling rate of 0.1 °C/s, (<b>b</b>) cooling rate of 1 °C/s.</p>
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<p>Comparison of half crystallization temperatures evaluated from PvT measurements and the predicted values: (<b>a</b>) cooling rate of 0.1 °C/s, (<b>b</b>) cooling rate of 1 °C/s.</p>
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<p>Comparison of the final spherulites radii evaluated from the analysis of the polarized optical micrographs of the samples (shown in <a href="#crystals-13-00922-f009" class="html-fig">Figure 9</a>) and model predictions obtained using Equations (7) and (8).</p>
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12 pages, 6525 KiB  
Communication
Aspects of Austenitization for the Bearing Steel Induction Quenching Design
by Daniela Nachazelova, Jaromir Dlouhy, Petr Motycka and Jakub Kotous
Materials 2023, 16(9), 3523; https://doi.org/10.3390/ma16093523 - 4 May 2023
Cited by 2 | Viewed by 1334
Abstract
The dissolution of carbides during the heating to the quenching temperature has a significant effect on the martensite oversaturation and the resulting mechanical properties. The kinetics of dissolution can be influenced by various external factors. This work deals with monitoring the carbide dissolution [...] Read more.
The dissolution of carbides during the heating to the quenching temperature has a significant effect on the martensite oversaturation and the resulting mechanical properties. The kinetics of dissolution can be influenced by various external factors. This work deals with monitoring the carbide dissolution utilizing dilatometer analysis. The austenitization of 100CrMnSi6-4 bearing steel in two initial states was compared—after accelerated spheroidization annealing and conventional soft annealing. The main objective was to determine the amount of undissolved cementite during austenitization in the temperature range where only austenite and cementite are present in the structure. The austenitization temperature determines the degree of cementite dissolution and, consequently, the carbon content in austenite and thus the final properties after quenching. The cementite dissolution was quantified from dilatometric curves and image analysis. Full article
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<p>Microstructure after ASR (accelerated spheroidization and refinement) process.</p>
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<p>Microstructure after conventional soft annealing (SA).</p>
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<p>Dilatometric curves from linear heating (ASR lin-13) and the simulation of real induction heating in a medium-frequency converter (ASR sim-13).</p>
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<p>The elongation of the sample ASR-lin-6 and its temperature derivative. A<sub>1</sub> is the starting point of transformation, T<sub>1</sub> is the point at which the elongation derivative equals to the derivative for homogeneous austenite, T<sub>i</sub> is the point of the maximum elongation derivative. T<sub>Q</sub> is the quenching temperature for the metallographic sample for the evaluation of the absolute amount of carbides.</p>
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<p>The illustration of the use of lever rule to determine M<sub>s</sub> temperature.</p>
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<p>Specimen SA—initial state. Globular carbides in the ferrite matrix. Polished specimen.</p>
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<p>Specimen SA-sim-13, quenched from T<sub>Q</sub> temperature. Undissolved carbides in a martensitic matrix. Polished specimen.</p>
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<p>Dilatometric curve of ASR-lin-6 (green line); amount of undissolved carbides (brown line). T<sub>Q</sub> for this sample was 923 °C, undissolved carbide content at T<sub>Q</sub> was 2.5 vol.% (see <a href="#materials-16-03523-t003" class="html-table">Table 3</a>).</p>
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<p>Dilatometric curve of SA-lin-6 (green line); amount of undissolved carbides (brown line). T<sub>Q</sub> for this sample was 954 °C; undissolved carbide content at T<sub>Q</sub> was 4.1 vol.% (see <a href="#materials-16-03523-t003" class="html-table">Table 3</a>).</p>
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<p>The plot of “Hardness-C<sub>γ</sub>” relationship. Encircled groups of points represent T<sub>Q</sub> specimens, heated by both linear and simulated regimes. The points out of the circles represent specimens quenched at temperatures T<sub>Q</sub>—30°C and T<sub>Q</sub>—60 °C.</p>
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<p>The plot of “Hardness-M<sub>s</sub>” relationship.</p>
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23 pages, 19196 KiB  
Article
Influence of Vanadium Micro-Alloying on the Microstructure of Structural High Strength Steels Welded Joints
by Giulia Stornelli, Anastasiya Tselikova, Daniele Mirabile Gattia, Michelangelo Mortello, Rolf Schmidt, Mirko Sgambetterra, Claudio Testani, Guido Zucca and Andrea Di Schino
Materials 2023, 16(7), 2897; https://doi.org/10.3390/ma16072897 - 5 Apr 2023
Cited by 17 | Viewed by 2762
Abstract
The inter-critically reheated grain coarsened heat affected zone (IC GC HAZ) has been reported as one of the most brittle section of high-strength low-alloy (HSLA) steels welds. The presence of micro-alloying elements in HSLA steels induces the formation of microstructural constituents, capable to [...] Read more.
The inter-critically reheated grain coarsened heat affected zone (IC GC HAZ) has been reported as one of the most brittle section of high-strength low-alloy (HSLA) steels welds. The presence of micro-alloying elements in HSLA steels induces the formation of microstructural constituents, capable to improve the mechanical performance of welded joints. Following double welding thermal cycle, with second peak temperature in the range between Ac1 and Ac3, the IC GC HAZ undergoes a strong loss of toughness and fatigue resistance, mainly caused by the formation of residual austenite (RA). The present study aims to investigate the behavior of IC GC HAZ of a S355 steel grade, with the addition of different vanadium contents. The influence of vanadium micro-alloying on the microstructural variation, RA fraction formation and precipitation state of samples subjected to thermal cycles experienced during double-pass welding was reported. Double-pass welding thermal cycles were reproduced by heat treatment using a dilatometer at five different maximum temperatures of the secondary peak in the inter-critical area, from 720 °C to 790 °C. Although after the heat treatment it appears that the addition of V favors the formation of residual austenite, the amount of residual austenite formed is not significant for inducing detrimental effects (from the EBSD analysis the values are always less than 0.6%). Moreover, the precipitation state for the variant with 0.1 wt.% of V (high content) showed the presence of vanadium rich precipitates with size smaller than 60 nm of which, more than 50% are smaller than 15 nm. Full article
(This article belongs to the Topic Microstructure and Properties in Metals and Alloys)
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<p>Hot rolled material (2% Nital etching) ((<b>a</b>) Reference material, (<b>b</b>) Variant I, (<b>c</b>) Variant II, (<b>d</b>) Variant III).</p>
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<p>Experimental thermal profiles as acquired by thermocouples as obtained by dilatometry.</p>
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<p>Microstructures as obtained by dilatometric cycles (reference material).</p>
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<p>Microstructures as obtained by dilatometric cycles (variant I).</p>
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<p>Microstructures as obtained by dilatometric cycles (variant II).</p>
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<p>Microstructures as obtained by dilatometric cycles (variant III).</p>
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<p>EBSD polar figure maps of specimens subjected to dilatometric cycles (reference material).</p>
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<p>EBSD polar figure maps of specimens subjected to dilatometric cycles (variant I).</p>
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<p>EBSD polar figure maps of specimens subjected to dilatometric cycles (variant II).</p>
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<p>EBSD polar figure maps of specimens subjected to dilatometric cycles (variant III).</p>
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<p>High-angle grain boundaries quantification (HAGBs %) (ϕ &gt; 10°) for each condition.</p>
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<p>Hardness dependence on inter-critical temperature for the different considered materials.</p>
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<p>XRD spectra for reference material (<b>a</b>) and variant III (<b>b</b>) as a function of second temperature peak.</p>
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<p>EBSD phase maps. Red zones: RA phase (reference material).</p>
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<p>EBSD phase maps. Red zones: RA phase (variant I).</p>
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<p>EBSD phase maps. Red zones: RA phase (variant II).</p>
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<p>EBSD phase maps. Red zones: RA phase (variant III).</p>
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<p>Microstructures of the considered specimens after LePerà etching. Red zones: RA phase (Variant I).</p>
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<p>Microstructures of the considered specimens after LePerà etching. Red zones: RA phase (Variant II).</p>
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<p>Quantified RA % with EBSD phase maps (<b>a</b>) and selective etching (<b>b</b>) as a function of the second peak temperature for the different considered steels.</p>
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<p>TEM micrograph of Variant II steel after inter-critical treatment with second peak temperature at 735 °C. Highlighted are the areas of cementite.</p>
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<p>TEM micrograph of Variant II steel after inter-critical treatment with second peak temperature at 735 °C. Highlighted are the fine V-rich precipitates in the matrix.</p>
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<p>TEM micrograph detail of Variant II steel after inter-critical treatment with second peak temperature at 735 °C. Highlighted are the fine V-rich precipitates in the matrix.</p>
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<p>Precipitates size distribution (Variant II, 735 °C).</p>
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<p>TEM micrograph of Variant III steel after inter-critical treatment with second peak temperature at 735 °C. Highlighted are the areas of cementite.</p>
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<p>TEM micrograph of Variant III steel after inter-critical treatment with second peak temperature at 735 °C. Highlighted are the Nb-V rich precipitates in the matrix.</p>
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<p>TEM micrograph of Variant III steel after inter-critical treatment with second peak temperature at 735 °C. Highlighted are the Nb rich precipitates in the matrix.</p>
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<p>Precipitates size distribution (Variant III, 735 °C).</p>
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21 pages, 5843 KiB  
Article
Experimental Studies of Thermophysical Properties and Microstructure of X37CrMoV5-1 Hot-Work Tool Steel and Maraging 350 Steel
by Piotr Koniorczyk, Mateusz Zieliński, Judyta Sienkiewicz, Janusz Zmywaczyk and Andrzej Dębski
Materials 2023, 16(3), 1206; https://doi.org/10.3390/ma16031206 - 31 Jan 2023
Cited by 9 | Viewed by 2737
Abstract
Measurements of thermal diffusivity, heat capacity and thermal expansion of X37CrMoV5-1 (1.2343) hot-work tool steel and Maraging 350 (1.6355) steel in the temperature range from −50 °C to 1000 °C were carried out in this paper. Both X37CrMoV5-1 and Maraging 350 are tested [...] Read more.
Measurements of thermal diffusivity, heat capacity and thermal expansion of X37CrMoV5-1 (1.2343) hot-work tool steel and Maraging 350 (1.6355) steel in the temperature range from −50 °C to 1000 °C were carried out in this paper. Both X37CrMoV5-1 and Maraging 350 are tested for military use as barrel steels. Thermophysical properties were tested using specialised test stands from NETZSCH. Thermal diffusivity was studied using both the LFA 427 laser flash apparatus in the temperature range of RT–1000 °C and the LFA 467 laser flash apparatus in the temperature range of −50 °C–500 °C. Specific heat capacity was investigated using a DSC 404 F1 Pegasus differential scanning calorimeter in the range RT–1000 °C, and thermal expansion was investigated using both a DIL 402 Expedis pushrod dilatometer in the range −50 °C–500 °C and a DIL 402 C in the range RT–1000 °C. Inconel 600 was selected as the reference material during the thermal diffusivity test using LFA467. Tests under the light microscope (LM), scanning electron microscopy (SEM) and Vickers microhardness measurements were carried out to detect changes in the microstructure before and after thermophysical measurements. This paper briefly characterises the research procedures used. In conclusion, the results of testing the thermophysical properties of X37CrMoV5-1 hot-work tool steel and Maraging 350 steel are compared with our results on 38HMJ (1.8509), 30HN2MFA and Duplex (1.4462) barrel steels. The thermophysical properties of X37CrMoV5-1 (1.2343) hot-work tool steel and Maraging 350 (1.6355) steel are incomplete in the literature. The paper presents the thermophysical properties of these steels over a wide range of temperatures so that they can be used as input data for numerical simulations of heat transfer in cannon barrels. Full article
(This article belongs to the Section Construction and Building Materials)
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<p>TTT curves (together with Continuous Cooling Transformation (CCT) curves) corresponding to the start of the precipitation of various phases in X37CrMoV5-1 hot-work tool steel; the letters stand for austenite (A), cementite (C), martensite (M), bainite (B), martensite start temperature (M<sub>S</sub>), ferrite (F) and austenite transformation temperature (Ac<sub>1e</sub>—start and Ac<sub>1b</sub>—end temperature of austenite transformation) [<a href="#B24-materials-16-01206" class="html-bibr">24</a>].</p>
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<p>TTT curves correspond to the start of the precipitation of various phases in Maraging 350 steel (cooling rate was counted from 1000 °C) [<a href="#B14-materials-16-01206" class="html-bibr">14</a>].</p>
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<p>Image of X37CrMoV5-1 in (<b>a</b>) as-delivered state, (<b>b</b>) after DSC testing made by using a digital microscope.</p>
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<p>SEM image of X37CrMoV5-1 in (<b>a</b>) as-delivered state, (<b>b</b>) after DSC testing.</p>
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<p>Qualitative EDS of X37CrMoV5-1 in the as-delivered state.</p>
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<p>Image of Maraging 350 steel in (<b>a</b>) as-delivered state, (<b>b</b>) after DSC testing made by using a digital microscope (together with a schematic representation of lath martensite structures, e.g., laths, blocks, packets, within a prior austenite grain).</p>
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<p>Thermal diffusivity as a function of temperature for the X37CrMoV5 −1 hot-work tool steel and Maraging 350 steel obtained from the first heating runs on LFA 467 and LFA 427.</p>
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<p>Thermal conductivity as a function of temperature for the X37CrMoV5 −1 hot-work tool steel and Maraging 350 steel obtained from the first heating runs on LFA 467 and LFA 427.</p>
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<p>Specific heat as a function of temperature for the X37CrMoV5 −1 hot-work tool steel and Maraging 350 steel obtained from the first heating run on LFA 467 in the range from −50 °C to 500 °C and from the approximation of the experimental data on DSC in the range from RT to 1000 °C.</p>
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<p>Temperature characteristics of apparent specific heat for the X37CrMoV5−1 hot-work tool steel obtained from the first and second heating run: dashed black line—results obtained from LFA 467; dashed green line—approximation of measurement on DSC.</p>
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<p>Density as a function of temperature for the X37CrMoV5−1 hot-work tool steel obtained from the first and second heating runs on DIL 402 C.</p>
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<p>Thermal expansion dL/L<sub>o</sub> (T) and CLTE as a function of temperature for the X37CrMoV5 −1 hot-work tool steel obtained from the first and second heating runs on DIL 402 Expedis and DIL 402 C.</p>
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<p>Thermal diffusivity as a function of temperature for the X37CrMoV5−1 steel obtained from the first heating runs on LFA 427 vs. literature data [<a href="#B15-materials-16-01206" class="html-bibr">15</a>].</p>
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<p>Temperature characteristics of apparent specific heat for the X37CrMoV5−1 steel obtained from the first heating run vs. literature data: dashed red line—approximation of data from [<a href="#B15-materials-16-01206" class="html-bibr">15</a>].</p>
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<p>Temperature characteristics of apparent specific heat for the Maraging 350 steel obtained from the first and second heating run: dashed black line—results obtained from LFA 467; dashed green line—approximation of measurement on DSC.</p>
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<p>Density as a function of temperature for the Maraging 350 steel obtained from the first and second heating runs on DIL 402 C.</p>
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<p>Thermal expansion and CLTE as a function of temperature for the Maraging steel obtained from the first and second heating runs on DIL 402 Expedis and DIL 402C.</p>
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<p>Comparison of thermal diffusivity measurements as a function of temperature for X37CrMoV5−1, Maraging 350, 38HMJ, 30HN2MFA and Duplex 2205 sheets of steel: (<b>a</b>) obtained by the authors from LFA 467; (<b>b</b>) obtained by the authors from LFA 467 [<a href="#B1-materials-16-01206" class="html-bibr">1</a>,<a href="#B13-materials-16-01206" class="html-bibr">13</a>].</p>
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<p>Comparison of thermal conductivity characteristics as a function of temperature for X37CrMoV5−1, Maraging 350, 38HMJ, 30HN2MFA and Duplex 2205 sheets of steel: (<b>a</b>) in the range −50–500 °C; (<b>b</b>) in the range RT–1000 °C [<a href="#B1-materials-16-01206" class="html-bibr">1</a>,<a href="#B13-materials-16-01206" class="html-bibr">13</a>].</p>
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<p>Comparison of thermal expansion characteristics as a function of temperature for X37CrMoV5−1, Maraging 350, 38HMJ, 30HN2MFA and Duplex 2205 sheets of steel [<a href="#B1-materials-16-01206" class="html-bibr">1</a>,<a href="#B13-materials-16-01206" class="html-bibr">13</a>].</p>
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<p>Comparison of density characteristics as a function of temperature for X37CrMoV5−1, Maraging 350, 38HMJ, 30HN2MFA and Duplex 2205 sheets of steel [<a href="#B1-materials-16-01206" class="html-bibr">1</a>,<a href="#B13-materials-16-01206" class="html-bibr">13</a>].</p>
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<p>Comparison of CLTE as a function of temperature for X37CrMoV5−1, Maraging 350, 38HMJ, 30HN2MFA and Duplex 2205 sheets of steel [<a href="#B1-materials-16-01206" class="html-bibr">1</a>,<a href="#B13-materials-16-01206" class="html-bibr">13</a>].</p>
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<p>Comparison of apparent specific heat as a function of temperature for X37CrMoV5−1, Maraging 350, 38HMJ, 30HN2MFA and Duplex 2205 sheets of steel [<a href="#B1-materials-16-01206" class="html-bibr">1</a>,<a href="#B13-materials-16-01206" class="html-bibr">13</a>].</p>
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16 pages, 8678 KiB  
Article
Suction Influence on Load–Settlement Curves Predicted by DMT in a Collapsible Sandy Soil
by Alfredo Lopes Saab, André Luís de Carvalho Rodrigues, Breno Padovezi Rocha, Roger Augusto Rodrigues and Heraldo Luiz Giacheti
Sensors 2023, 23(3), 1429; https://doi.org/10.3390/s23031429 - 27 Jan 2023
Cited by 1 | Viewed by 2154
Abstract
The plate load test (PLT) is the most reliable in situ testing for studying the load–settlement behaviour of footings on unsaturated collapsible soils. In these soils, the suction profile is not constant along the depth, and the scale effect between the prototype and [...] Read more.
The plate load test (PLT) is the most reliable in situ testing for studying the load–settlement behaviour of footings on unsaturated collapsible soils. In these soils, the suction profile is not constant along the depth, and the scale effect between the prototype and footing leads to different suction averages and, consequently, different data. One method to eliminate the effect of soil suction on the test data is to fully saturate the soil prior to the test, which is also recommended at the design process for footing on collapsible soils. However, the inundation process on PLTs is expensive and time-consuming, which makes this procedure difficult to incorporate into engineering practice. This study presents a device that can be attached to flat dilatometer (DMT) to allow local inundation of the soil as part of the in situ test campaign and obtain the DMT-constrained modulus (MDMT) for both natural and inundated conditions. The MDMT presented an average reduction of 56% from natural to inundated condition. This parameter can be used in a model to predict load–settlement curves by DMT data considering the suction influence on this behaviour. The curves obtained from the prediction model were compared to curves determined by PLT conducted under the same in situ conditions. Good agreement was found between the curves predicted by DMT and those measured by PLT for both conditions. The proposed procedure, which uses a device attached to the DMT blade, provides an investigation method to obtain the load–settlement curve under different suction conditions, which can help in the selection and performance prediction of shallow foundations, taking into account suction and collapse phenomenon-related problems. Full article
(This article belongs to the Section Physical Sensors)
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<p>Parts of inundation device and system attached to the DMT blade. Adapted from [<a href="#B32-sensors-23-01429" class="html-bibr">32</a>].</p>
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<p>Schematic representation of the DMT test with the inundation device. Adapted from [<a href="#B32-sensors-23-01429" class="html-bibr">32</a>].</p>
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<p>Schematic representation of DMT inundation process, (<b>a</b>) test under natural condition, (<b>b</b>) depth of interest; (<b>c</b>) local inundation and (<b>d</b>) readings under inundated condition. Adapted from [<a href="#B32-sensors-23-01429" class="html-bibr">32</a>].</p>
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<p>Equipment and setup for the plate load tests carried out at the site.</p>
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<p>(<b>a</b>) Base of the pit prior to the inundation, (<b>b</b>) test assembled prior to the inundation, (<b>c</b>) process of inundation and (<b>d</b>) PLT carried out under inundated condition.</p>
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<p>In situ and laboratory tests previously carried out in the experimental site. Adapted from [<a href="#B38-sensors-23-01429" class="html-bibr">38</a>,<a href="#B39-sensors-23-01429" class="html-bibr">39</a>].</p>
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<p>Suction monitoring by granular matrix sensors and precipitation data.</p>
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<p>Average profile as well as plus and minus one standard deviation (<span class="html-italic">SD</span>) of <span class="html-italic">I<sub>D</sub></span>, <span class="html-italic">K<sub>D</sub></span>, and <span class="html-italic">E<sub>D</sub></span> determined in natural conditions.</p>
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<p><span class="html-italic">I<sub>D</sub></span>, <span class="html-italic">K<sub>D</sub></span>, and <span class="html-italic">E<sub>D</sub></span> profiles obtained after local inundation and average profile for the natural condition.</p>
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<p>Constrained modulus values determined from the DMT, oedometer test and PLT, under natural and inundated condition for the study site.</p>
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<p>Load–settlement curves measured by PLT and predicted by DMT results, under natural and inundated conditions, for 1.0, 2.0, 3.0, and 4.0 m depth.</p>
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<p>Settlements estimated from DMT vs. those obtained by PLT for the natural (<b>a</b>) and inundated (<b>b</b>) conditions of the study site.</p>
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