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Welding of Steels
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Welding of Steels

A special issue of Applied Sciences (ISSN 2076-3417). This special issue belongs to the section "Mechanical Engineering".

Deadline for manuscript submissions: closed (15 January 2019) | Viewed by 75711

Special Issue Editor

Prof. Dr. Filippo Berto

E-Mail Website
Guest Editor
Department of Chemical Engineering, Materials and Environment, Sapienza University of Rome, 00184 Rome, Italy
Interests: fatigue and fracture behavior of materials; mechanical characterization; structural integrity of conventional and innovative materials
Special Issues, Collections and Topics in MDPI journals

Special Issue Information

Dear Colleagues,

Steels are still an undisputable material for many of constructions made today. Both stainless and ordinary C-Mn steels have developed significantly over the last few years, as a response to new demands for strength, corrosion resistance, etc., for a more sustainable future. However, to make effective use of these steels, they need to be weldable. This Special Issue will deal with welding of modern steels, both newer steels and application of new welding techniques for the lower strength levels as well as new very high strength steels. It will also cover welding of stainless steels, with its multitude of alloys.

Significant research is going on around the world about welding of these new steels, in terms of steel design, consumable design and welding techniques, which need to be highlighted. Papers on experimental investigations as well as numerical analyses are welcome to illustrate the progress made within this field.

Prof. Dr. Filippo Berto


Guest Editors

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Keywords

  • Welding
  • High strength steels
  • Stainless steels
  • New alloys
  • Metallurgy
  • Welding techniques

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Published Papers (15 papers)

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Research

18 pages, 3181 KiB  
Article
Numerical Investigation on Ultimate Compressive Strength of Welded Stiffened Plates Built by Steel Grades of S235–S390
by Chenfeng Li, Sen Dong, Tingce Wang, Weijun Xu and Xueqian Zhou
Appl. Sci. 2019, 9(10), 2088; https://doi.org/10.3390/app9102088 - 21 May 2019
Cited by 13 | Viewed by 4770
Abstract
The welded stiffened plate is widely used in naval architecture and offshore engineering as a basic structural member. The aim of this study is to investigate the effect of welding residual stress and steel grade on the ultimate strength of stiffened plates under [...] Read more.
The welded stiffened plate is widely used in naval architecture and offshore engineering as a basic structural member. The aim of this study is to investigate the effect of welding residual stress and steel grade on the ultimate strength of stiffened plates under uniaxial compressive load by non-linear finite element analysis. Nineteen stiffened plates built with three types of stiffeners with various column slenderness ratios provided in the ISSC’2000 VI.2 benchmark calculations are employed in the present study. The commercial finite element code ABAQUS is applied to simulate the collapse behavior of the stiffened plates and verified against the benchmark calculations. Fabrication-related imperfections, such as initial deflections and residual stresses, are accounted for in the simulations. The ultimate strength of stiffened plates built in common shipbuilding steels, namely S235, S315, S355 and S390, are investigated by varying the yield strength of materials in the simulation. Analysis of the numerical results shows that the welding residual stress reduces the ultimate strength of stiffened plates, and increase in yield strength of the material can effectively improve the ultimate strength of common ship stiffened plates; and quantitative analyses of their influences have also been performed. Full article
(This article belongs to the Special Issue Welding of Steels)
►▼ Show Figures

Figure 1

Figure 1
<p>Definition of cross-sectional dimension of stiffened plate.</p> Full article ">Figure 2
<p>1/2 + 1/2 span stiffened plate model.</p> Full article ">Figure 3
<p>Assumed initial deformation of stiffened plate.</p> Full article ">Figure 4
<p>Assumed residual stress distribution in the fillet weld.</p> Full article ">Figure 5
<p>Results for the ultimate strength of flat-bat stiffened plate versus column slenderness ratio compared with the ISSC’2000 Report [<a href="#B11-applsci-09-02088" class="html-bibr">11</a>].</p> Full article ">Figure 6
<p>Results for the ultimate strength of angle-bar stiffened plate versus column slenderness ratio compared with the ISSC’2000 Report [<a href="#B11-applsci-09-02088" class="html-bibr">11</a>].</p> Full article ">Figure 7
<p>Results for ultimate strength of tee-bar stiffened plate versus column slenderness ratio compared with the ISSC’2000 Report [<a href="#B11-applsci-09-02088" class="html-bibr">11</a>].</p> Full article ">Figure 8
<p>Ultimate strength of flat-bar stiffened plates versus column slenderness ratio.</p> Full article ">Figure 9
<p>Ultimate strength of angle-bar stiffened plates versus column slenderness ratio.</p> Full article ">Figure 10
<p>Ultimate strength of tee-bar stiffened plates versus column slenderness ratio.</p> Full article ">Figure 11
<p>The trend of ultimate strength of the stiffened plates with various built-up steels.</p> Full article ">Figure 11 Cont.
<p>The trend of ultimate strength of the stiffened plates with various built-up steels.</p> Full article ">Figure 12
<p>The increase ratio of ultimate strength of the stiffened plates plotted against column slenderness ratio.</p> Full article ">
12 pages, 7841 KiB  
Article
A New Plasma Surface Alloying to Improve the Wear Resistance of the Metallic Card Clothing
by Dongbo Wei, Fengkun Li, Shuqin Li, Xiaohu Chen, Feng Ding, Pingze Zhang and Zhangzhong Wang
Appl. Sci. 2019, 9(9), 1849; https://doi.org/10.3390/app9091849 - 6 May 2019
Cited by 4 | Viewed by 2591
Abstract
A new surface strengthening process: Plasma surface chromizing was implemented on the metallic card clothing to improve its wear resistance based on double glow plasma surface metallurgy technology. A chromizing coating was prepared in the process, which consisted of a deposited layer and [...] Read more.
A new surface strengthening process: Plasma surface chromizing was implemented on the metallic card clothing to improve its wear resistance based on double glow plasma surface metallurgy technology. A chromizing coating was prepared in the process, which consisted of a deposited layer and diffusion layer. The surface morphologies, microstructure, phase composition, and hardness were analyzed in detail. The friction behaviors of the metallic card clothing before and after plasma surface alloying were comparatively analyzed under various sliding speeds at room temperature. The results showed that: 1. The chromizing coating on the surface of metallic card clothing was dense and homogeneous without defects, and the metallic card clothing still maintained its integrity and sharpness. 2. The chromizing coating consist of [Fe,Cr], Cr, Cr23C6, and Cr7C3, which contribute to the high hardness. 3. The average microhardness of metallic card clothing increased from 365.4 HV0.05 to 564.9 HV0.05 after plasma surface chromizing. Nano hardness of the chromizing coating was approximately 1.87 times than the metallic card clothing. 4. At various sliding velocities of 2 m/min, 4 m/min, and 6 m/min, the specific wear rates of metallic card clothing were 16.38, 9.06 and 6.26 × 10−4·mm3·N−1·m−1, and the specific wear rates of metallic card clothing after plasma surface chromizing were 2.91, 3.30, and 2.95 × 10−4·mm3·N−1·m−1. Furthermore, the wear mechanism of the chromizing coating gradually changed from adhesive wear to abrasive wear as the sliding velocity increased. The results indicate that the wear resistance of metallic card clothing was improved obviously after plasma surface chromizing. Full article
(This article belongs to the Special Issue Welding of Steels)
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Figure 1

Figure 1
<p>Schematic illustration (<b>a</b>) and equipment drawing (<b>b</b>) of double glow plasma technology.</p> Full article ">Figure 2
<p>The schematic diagram of the test point. (<b>a</b>) The tooth point position of the metallic card clothing; (<b>b</b>) The position of the tooth base of the metal needle cloth; (<b>c</b>) The bottom of the metallic card clothing.</p> Full article ">Figure 3
<p>The schematic diagram of reciprocating friction (<b>a</b>), the equipment drawing of CET-I friction-wear tester (<b>b</b>) and the schematic diagram of profilometer equipped by the friction-wear tester (<b>c</b>).</p> Full article ">Figure 4
<p>Surface morphologies (<b>a</b>,<b>b</b>) and XRD (X-ray diffraction method) patterns (<b>c</b>) of chromizing coating.</p> Full article ">Figure 5
<p>Cross-sectional microstructure (SEM) and composition distribution (EDS) of chromizing coating. (<b>a</b>) cross-sectional microstructure; (<b>b</b>) composition distribution</p> Full article ">Figure 6
<p>Nano indentation curve, nanohardness and elastic modulus bar graph of the metallic card clothing before and after plasma surface alloying. (<b>a</b>) Force-depth curves of uncoated and coated samples, and (<b>b</b>) hardness and elastic modulus of uncoated and coated samples.</p> Full article ">Figure 7
<p>Friction coefficient vs. sliding time for untreated and treated samples under various sliding speeds: (<b>a</b>) 2 m/min, 420 g; (<b>b</b>) 4 m/min, 420 g; (<b>c</b>) 6 m/min, 420 g; and (<b>d</b>) variation tendency of friction coefficient.</p> Full article ">Figure 8
<p>Wear profiles for uncoated and coated samples under various speeds: (<b>a</b>) 2 m/min; (<b>b</b>) 4 m/min; (<b>c</b>) 6 m/min; and (<b>d</b>) variation tendency of wear profile.</p> Full article ">Figure 9
<p>Wear track morphology of chromizing coating under various sliding speeds (<b>a</b>) 2 m/min; (<b>b</b>) 4 m/min; and (<b>c</b>) 6 m/min.</p> Full article ">
17 pages, 8370 KiB  
Article
Metallurgical Effects of Niobium and Molybdenum on Heat-Affected Zone Toughness in Low-Carbon Steel
by Hardy Mohrbacher
Appl. Sci. 2019, 9(9), 1847; https://doi.org/10.3390/app9091847 - 5 May 2019
Cited by 13 | Viewed by 6276
Abstract
Modern weldable high strength steel grades are typically based on low-carbon alloy concepts using microalloying for obtaining a good strength-toughness balance. Such steel grades having a yield strength in the range of 420 to 690 MPa are very commonly used in pipelines, heavy [...] Read more.
Modern weldable high strength steel grades are typically based on low-carbon alloy concepts using microalloying for obtaining a good strength-toughness balance. Such steel grades having a yield strength in the range of 420 to 690 MPa are very commonly used in pipelines, heavy vehicles, shipbuilding and general structural applications. Thermomechanical processing during hot rolling combined with accelerated cooling is an established means of producing such steel grades. Considering the alloying concepts, the use of niobium and molybdenum, and in selected cases boron, is very efficient to achieve high strength and good toughness. However, all targeted applications of such high strength steels involve extensive welding. Thus, heat affected zone properties are of particular importance. The present paper investigates the effects of Nb, Mo and Ti on the heat affected zone properties. Variations of the Mn and Si contents are considered as well. Additionally, the influence of post-weld heat treatment in the coarse-grained heat-affected zone (HAZ) is considered. In this approach, HAZ subzones were generated using laboratory weld cycle simulations in combination with systematic variation of alloying elements to scrutinize and interpret their specific effects. The results indicate that Mo and Nb, when alloyed in the typical range, provide excellent HAZ toughness and guarantee sufficiently low ductile-to-brittle transition temperature. An alloy combination of Nb, Mo and Ti improves performance under hot deformation conditions and toughness after post-weld heat treatment. Full article
(This article belongs to the Special Issue Welding of Steels)
►▼ Show Figures

Figure 1

Figure 1
<p>Design criteria for weldable high strength steels: (<b>a</b>) The Graville diagram classifying weldability of steels and historical evolution of chemical composition represented by the carbon equivalent (CE) in high strength steels; (<b>b</b>) ductile-to-brittle transition (DBTT) behavior and criteria influencing the toughness behavior in steel.</p> Full article ">Figure 2
<p>Properties of the as-rolled materials: (<b>a</b>) Relationship of yield strength (R<sub>p0.2</sub>) and tensile strength; (<b>b</b>) ductile-to-brittle transition temperature (50%) determined by Charpy testing (PF: polygonal ferrite matrix, all other acicular ferrite matrices).</p> Full article ">Figure 3
<p>Light optical micrographs (Le Pera etching) of the as-rolled materials: (<b>a</b>) V2-a (polygonal ferrite); (<b>b</b>) V2-d (mostly acicular ferrite); (<b>c</b>) V4-a (coarse acicular ferrite); (<b>d</b>) V4-d (fine acicular ferrite).</p> Full article ">Figure 4
<p>(<b>a</b>) Setup for HAZ simulation and (<b>b</b>) sample geometries for mechanical testing.</p> Full article ">Figure 5
<p>Heat-affected zone simulation of alloy V3: (<b>a</b>) Evolution of hardness and yield-to-tensile ratio (YTR) and measured transformation temperatures during up-heating; (<b>b</b>) microstructure in the inter-critical zone (ICZ) of the Nb-free alloy V3-a; (<b>c</b>) microstructure of alloy V3-d containing 0.10%Nb (etching with Le Pera agent).</p> Full article ">Figure 6
<p>Evolution of hardness and yield-to-tensile ratio (YTR) in the heat-affected zone of (<b>a</b>) alloys V4; (<b>b</b>) alloys V6. Measured transformation temperatures during up-heating are indicated.</p> Full article ">Figure 7
<p>Ductile-to-brittle transition temperature (DBTT T50%) determined by Charpy testing: (<b>a</b>) Simulated intercritcal HAZ (ICZ) with a peak temperature of 900 °C; (<b>b</b>) simulated coarse-grained HAZ (CGZ) with a peak temperature of 1300 °C.</p> Full article ">Figure 8
<p>Effect of Ti and Nb micro-alloy addition on the microstructure in the CGZ (etching with Le Pera agent); abnormal prior austenite grain size in alloy V4-d (0.10%Nb) indicated by arrows.</p> Full article ">Figure 9
<p>Reduction of area vs. hot-tensile strength of coarse-grained zone (CGZ) treated material at a test temperature of (<b>a</b>) 560 °C and (<b>b</b>) 620 °C.</p> Full article ">Figure 10
<p>Influence of post-weld heat treatment (PWHT) on strength and toughness: (<b>a</b>) DBTT (T50%) vs. yield strength after a relaxation treatment at 560 °C for 30 min; (<b>b</b>) DBTT (T50%) vs. yield strength before relaxation treatment (same data as <a href="#applsci-09-01847-f007" class="html-fig">Figure 7</a>b, same trendline as <a href="#applsci-09-01847-f009" class="html-fig">Figure 9</a>a).</p> Full article ">Figure 11
<p>Precipitation strengthening effect in HAZ sub-zones for alloy variants containing 0.10%Nb: (<b>a</b>) Sub-critical zone at peak temperatures of 600 and 800 °C; (<b>b</b>) coarse-grained zone (peak temperature 1300 °C), precipitation effect of PWHT (560 °C/30 min.) and sum of precipitation strengthening (CGZ + PWHT).</p> Full article ">
20 pages, 16705 KiB  
Article
Filler Metal Mixing Behaviour of 10 mm Thick Stainless Steel Butt-Joint Welds Produced with Laser-Arc Hybrid and Laser Cold-Wire Processes
by Miikka Karhu, Veli Kujanpää, Harri Eskelinen and Antti Salminen
Appl. Sci. 2019, 9(8), 1685; https://doi.org/10.3390/app9081685 - 24 Apr 2019
Cited by 9 | Viewed by 4831
Abstract
In thick section laser welding, filler metal addition is usually required to improve joint fit-up tolerances or to control the chemical composition of the weld metal. With deep and narrow welds produced using an over-alloyed filler metal, it may be challenging to ensure [...] Read more.
In thick section laser welding, filler metal addition is usually required to improve joint fit-up tolerances or to control the chemical composition of the weld metal. With deep and narrow welds produced using an over-alloyed filler metal, it may be challenging to ensure that the filler metal and its elements are homogeneously mixed and evenly distributed throughout the fusion zone. Inhomogeneous filler metal mixing can cause unfavourable changes to weld metal chemistry and microstructure. Filler metal mixing behaviour in laser-arc hybrid and laser cold-wire welding is studied in this work. Welding tests were conducted on 10 mm thick butt-welded joints of AISI 316L austenitic stainless steel. An overmatching type 2205 duplex stainless steel filler wire was used to obtain a composition contrast between the base metal and filler metal. Energy dispersive spectroscopy (EDS) with chromium as the trace element was used for element mapping and stepwise characterization of the weld cross-section samples. Optical metallography was used to observe possible inhomogeneous filler metal mixing behaviour like local acute changes in macro- and microstructural features. The results showed a clear difference in filler metal mixing between the weld surface part (upper half) of the weld and the weld root part (lower half) in 10 mm thick welded cross-sections for closed root gap of I-groove welds or when the gap was only 0.4 mm. In narrow I-groove preparations, inhomogeneous mixing phenomena were more pronounced in laser cold-wire welds than in laser-arc hybrid welds. In both welding processes, a combination of trailing wire feeding and the use of a wider groove enabled filler metal to be introduced deeper into the bottom of the groove and improved mixing in the root portion of the welds. Full article
(This article belongs to the Special Issue Welding of Steels)
►▼ Show Figures

Figure 1

Figure 1
<p>Butt-joint configurations used in the experiments: (<b>a</b>) Closed square (root gap = 0 mm); (<b>b</b>) open square (root gap = 0.4 mm and 0.8 mm); and (<b>c</b>) V-groove with 10-degree groove angle, closed gap and without root face.</p> Full article ">Figure 2
<p>Welding head set-up used in the laser-arc hybrid and laser cold-wire experiments. (<b>A</b>) process fibre with connector; (<b>B</b>) laser welding head-unit; (<b>C</b>) adjustable gas metal arc welding (GMAW) torch unit.</p> Full article ">Figure 3
<p>Set-up of laser beam and filler wire feeding arrangements used in laser-arc hybrid welding (upper row; <b>a</b>,<b>b</b>) and laser cold-wire welding (lower row; <b>c</b>,<b>d</b>). A1 = leading wire feeding. A2 = trailing wire feeding.</p> Full article ">Figure 4
<p>Relationship between chromium content in the weld metal and corresponding filler metal mixing-%. On the horizontal axis, 17.6 represents chromium weight-% content equal to the base metal and 22.9 is the chromium weight-% content of the filler metal.</p> Full article ">Figure 5
<p>Example showing transversal EDS-measurement line locations (in red) ML1, ML2, ML3 and ML4 in the studied weld cross-sections.</p> Full article ">Figure 6
<p>Example showing the cross-sectional area defined from the macrograph of test weld LAH1.</p> Full article ">Figure 7
<p>Weld cross-sections (<b>a</b>–<b>f</b>) from laser-arc hybrid welding tests.</p> Full article ">Figure 8
<p>Weld cross-sections (<b>a</b>–<b>h</b>) from laser cold-wire welding tests.</p> Full article ">Figure 9
<p>Energy dispersive spectroscopy (EDS) mapping images (<b>a</b>–<b>f</b>) showing the distribution of chromium in laser arc-hybrid test welds.</p> Full article ">Figure 10
<p>Average values of filler metal mixing-% from the data of measurement lines ML1–ML4 in laser-arc hybrid test welds.</p> Full article ">Figure 11
<p>Filler metal mixing results from EDS stepwise characterization. (<b>a</b>–<b>d</b>) Filler metal mixing-% profiles of measurement lines 1–4 (ML1–ML4) from test weld LAH3 (<b>e</b>). In figures (<b>a</b>–<b>d</b>) one step between adjacent measurement points in the horizontal axis equals 100 micrometres.</p> Full article ">Figure 12
<p>Filler metal mixing results from EDS stepwise characterization. (<b>a</b>–<b>d</b>) Filler metal mixing-% profiles of measurement lines 1–4 (ML1–ML4) from test weld LAH4 (<b>e</b>). In figures (<b>a</b>–<b>d</b>) one step between adjacent measurement points in the horizontal axis equals 100 micrometres.</p> Full article ">Figure 13
<p>Filler metal mixing results from EDS stepwise characterization. (<b>a</b>–<b>d</b>) Filler metal mixing-% profiles of measurement lines 1–4 (ML1–ML4) from test weld LAH5 (<b>e</b>). In figures (<b>a</b>–<b>d</b>) one step between adjacent measurement points in the horizontal axis equals 100 micrometres.</p> Full article ">Figure 14
<p>Filler metal mixing results from EDS stepwise characterization. (<b>a</b>–<b>d</b>) Filler metal mixing-% profiles of measurement lines 1–4 (ML1–ML4) from test weld LAH6 (<b>e</b>). In figures (<b>a</b>–<b>d</b>) one step between adjacent measurement points in the horizontal axis equals 100 micrometres.</p> Full article ">Figure 15
<p>EDS mapping images (<b>a</b>–<b>h</b>) showing the distribution of chromium in laser cold-wire test welds.</p> Full article ">Figure 16
<p>Tabular data presenting average values of filler metal mixing-% from the data of measurement lines ML1–ML4 in laser cold-wire test welds.</p> Full article ">Figure 17
<p>Filler metal mixing results from EDS stepwise characterization. (<b>a</b>–<b>d</b>) Filler metal mixing-% profiles of measurement lines 1–4 (ML1–ML4) from test weld LCW6 (<b>e</b>). In figures (<b>a</b>–<b>d</b>) one step between adjacent measurement points in the horizontal axis equals 100 micrometres.</p> Full article ">Figure 18
<p>Example of mixing behaviour in a laser cold-wire weld produced in I-groove configuration with closed gap. Micrographs (<b>b</b>,<b>c</b>) and filler metal mixing profiles (<b>d</b>,<b>e</b>) at corresponding locations (<b>a</b>) showing inhomogeneous mixing occurred in laser cold-wire test weld LCW1.</p> Full article ">Figure 19
<p>Filler metal mixing results from EDS stepwise characterization. (<b>a</b>–<b>d</b>) Filler metal mixing-% profiles of measurement lines 1–4 (ML1–ML4) from test weld LCW7 (<b>e</b>). In figures (<b>a</b>–<b>d</b>) one step between adjacent measurement points in the horizontal axis equals 100 micrometres.</p> Full article ">Figure 20
<p>Filler metal mixing results from EDS stepwise characterization. (<b>a</b>–<b>d</b>) Filler metal mixing-% profiles of measurement lines 1–4 (ML1–ML4) from test weld LCW8 (<b>e</b>). In figures (<b>a</b>–<b>d</b>) one step between adjacent measurement points in the horizontal axis equals 100 micrometres.</p> Full article ">
18 pages, 8609 KiB  
Article
The Resistance of Welded Joints of Galvanized RHS Trusses with Different Vent Hole Geometries
by Miguel A. Serrano, Carlos López-Colina, Fernando L. Gayarre, Tim Wilkinson and Jesús Suárez
Appl. Sci. 2019, 9(8), 1553; https://doi.org/10.3390/app9081553 - 15 Apr 2019
Cited by 2 | Viewed by 4368
Abstract
A worldwide-accepted technique to protect steel lattice girders with welded hollow sections against corrosion is the hot-dip galvanizing process. In this process, vent holes are required in braces to fill the inner part protecting them from corrosion, to allow the immersion of the [...] Read more.
A worldwide-accepted technique to protect steel lattice girders with welded hollow sections against corrosion is the hot-dip galvanizing process. In this process, vent holes are required in braces to fill the inner part protecting them from corrosion, to allow the immersion of the structure in the zinc bath and to recover the excess fluid after the bath. The cross-section reduction due to the vent hole could lead to a decrease in the effective brace resistance; this is not easily quantified, because there are neither prescriptions nor recommendations in the design codes to assess this effect. Therefore, the hollow structural sections could be underutilized due to doubts regarding the safety of this type of joint. This research was conducted in order to categorize different geometries and positions of vent holes in order to determine the best in terms of joint efficiency. A validated finite element model considering welds on lattice girders joints was extended to take into account different vent hole shapes. This research concludes that the presence of ventilation holes such as the ones considered in this study does not significantly affect the joint resistance, and that all the analyzed hole shapes could be proposed as a valid solution for machining vent holes. The conclusions drawn up from this work could be useful for structural steel designers, providing them with valuable design recommendations. Full article
(This article belongs to the Special Issue Welding of Steels)
►▼ Show Figures

Figure 1

Figure 1
<p>Weld geometry modeling with shell elements (A–B and C–D).</p> Full article ">Figure 2
<p>Configuration of different vent holes.</p> Full article ">Figure 3
<p>Mesh models of holes and welds.</p> Full article ">Figure 4
<p>Global boundary conditions in FE Models.</p> Full article ">Figure 5
<p>Tested K-joint configuration.</p> Full article ">Figure 6
<p>Curve compression force vs. indentation at point 1.</p> Full article ">Figure 7
<p>Curve compression force vs. indentation at point 2.</p> Full article ">Figure 8
<p>Von Mises stresses in the simulation of the K-joint.</p> Full article ">Figure 9
<p>Model vs. tests for 2nd validation.</p> Full article ">Figure 10
<p>Comparison of maximum load in the joints with and without vent holes.</p> Full article ">Figure 11
<p>Comparison of load for a 3%b<sub>0</sub> indentation in joints with and without vent holes.</p> Full article ">Figure 12
<p>Comparison of maximum load for a 3% b<sub>0</sub> indentation.</p> Full article ">Figure 13
<p>Failure in joint KB1 with a semi-circular hole and brace angles of 45°.</p> Full article ">Figure 14
<p>Failure in joint KB4 with semi-circular hole and brace angles of 55°.</p> Full article ">Figure 15
<p>Failure in joint KC8 with V-notched hole and brace angles of 60°.</p> Full article ">Figure 16
<p>Failure in joint NP10 with V-notched hole and braces angles of 45–90°.</p> Full article ">
10 pages, 2213 KiB  
Article
Fracture Assessment of Weld Joints of High-Strength Steel in Pre-Strained Condition
by Gyubaek An, Jeongung Park, Mituru Ohata and Fumiyoshi Minami
Appl. Sci. 2019, 9(7), 1306; https://doi.org/10.3390/app9071306 - 28 Mar 2019
Cited by 4 | Viewed by 3627
Abstract
Unstable fractures tend to occur after ductile crack initiation or propagation. In most collapsed steel structures, a maximum 15% pre-strain was recorded, at the steel structural connections, during the great earthquake of 1995, in Japan. Almost-unstable fractures were observed in the beam-to-column connections, [...] Read more.
Unstable fractures tend to occur after ductile crack initiation or propagation. In most collapsed steel structures, a maximum 15% pre-strain was recorded, at the steel structural connections, during the great earthquake of 1995, in Japan. Almost-unstable fractures were observed in the beam-to-column connections, where geometrical discontinuities existed. Structural collapse and unstable failure occurred after large-scale plastic deformations. Ship structures can also suffer from unstable fractures in the welded joints. The fracture resistance of butt-welded joints subjected to tension in the pre-strained condition was estimated by considering the toughness deterioration, due to pre-strain and toughness correction for constraint loss in a tension specimen. The target specimen for this fracture assessment was a double-edged, through-thickness crack panel, with a crack in the weld joint (heat-affected zone (HAZ)). The critical fracture toughness value (crack tip opening displacement (CTOD)) of a large structure with pre-strain, which was applied to the HAZ region, was estimated from a small-scale, pre-stained, three-point bend specimen. Fracture toughness values, evaluated by a CTOD test, were recently mandated for shipbuilding steel plates. The critical fracture toughness value is a very useful parameter to evaluate the safety of huge ship structures. Full article
(This article belongs to the Special Issue Welding of Steels)
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Figure 1

Figure 1
<p>Groove geometry and macro section of the weld joint.</p> Full article ">Figure 2
<p>Double-edged through-thickness crack panel with a crack in the heat-affected zone (HAZ).</p> Full article ">Figure 3
<p>Effect of pre-strain on ductile-to-brittle transition temperature [<a href="#B19-applsci-09-01306" class="html-bibr">19</a>].</p> Full article ">Figure 4
<p>Monographs of equivalent CTOD ratio <span class="html-italic">β</span><sub>0</sub> for double-edged, through-thickness crack panel (ETCP) with a reference crack length of 2a = 11 mm.</p> Full article ">Figure 5
<p>Stress-strain curves for HSB600 steel.</p> Full article ">Figure 6
<p>Critical CTODs for the 3.0% pre-strained ETCP estimated from 3.0% pre-strained (three-point bend) 3PB test results.</p> Full article ">Figure 7
<p>Temperature shift of fracture toughness Δ<span class="html-italic">T</span><sub>PD</sub> in pre-strained and dynamic conditions.</p> Full article ">Figure 8
<p>Temperature shift ∆<span class="html-italic">T<sub>PD</sub></span> for the pre-strained HAZ predicted with the procedure specified in WES 2808.</p> Full article ">Figure 9
<p>Critical CTOD of the 3.0% pre-strained ETCP estimated from the predicted value.</p> Full article ">
18 pages, 7814 KiB  
Article
A New Approach to the Study of Multi-Pass Welds–Microstructure and Properties of Welded 20-mm-Thick Superduplex Stainless Steel
by Maria Asuncion Valiente Bermejo, Kjell Hurtig, Daniel Eyzop and Leif Karlsson
Appl. Sci. 2019, 9(6), 1050; https://doi.org/10.3390/app9061050 - 13 Mar 2019
Cited by 18 | Viewed by 3534
Abstract
Type 2507 superduplex stainless steel 20 mm in thickness was multi-pass-welded with Gas Metal Arc Welding (GMAW) and Flux-Cored Arc Welding (FCAW) processes. Recommended and higher arc energies and inter-pass temperatures were used. Thermal cycles were monitored using a recently developed procedure involving [...] Read more.
Type 2507 superduplex stainless steel 20 mm in thickness was multi-pass-welded with Gas Metal Arc Welding (GMAW) and Flux-Cored Arc Welding (FCAW) processes. Recommended and higher arc energies and inter-pass temperatures were used. Thermal cycles were monitored using a recently developed procedure involving the successive instrumentation of the multi-pass welds, pass by pass, by addition of thermocouples in each weld pass. The repeatability of temperature measurements and survival rate of more than 90% of thermocouples confirmed the reliability of the procedure. Reheating by subsequent passes caused a progressive increase in the austenite content of the weld metal. The as-deposited GMAW passes with higher-than-recommended arc energy showed the lowest presence of nitrides. Therefore, the cooling rate—and not the time exposed at the critical temperature range—seems to be the key factor for nitride formation. The welding sequence layout also plays an important role in the distribution of secondary phases. A larger amount and concentration of secondary austenite and σ-phase was found for a larger number of subsequent passes in the immediate vicinity of a specific weld pass. The impact toughness exceeded requirements for all welds. Differences in absorbed energies were related to the amount of micro-inclusions found with the FCAW weld showing the lowest absorbed energies and highest amount of micro-inclusions. Pitting corrosion preferentially initiated in locations with secondary austenite and σ-phase. However, in the absence of these secondary phases, the HAZ containing nitrides was the weakest location where pitting initiated. The results of this work have implications on practical welding for superduplex stainless steels: the current recommendations on maximum arc energy should be revised for large thickness weldments, and the importance of the welding sequence layout on the formation of secondary phases should be considered. Full article
(This article belongs to the Special Issue Welding of Steels)
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<p>Welding sequence layout. (<b>Left</b>) FR weld; (<b>Center</b>) GR weld; (<b>Right</b>) GH weld.</p> Full article ">Figure 2
<p>Welding configuration used to extract coupons to study the evolution of the microstructure caused by the subsequent weld passes. The numbers from 1 to 7 correspond to the locations where coupons with passes from 1 to 7 were extracted and contained that number of weld passes, i.e., 1 means one weld pass, 2 means two weld passes, and so forth until 7, including the seven weld passes in the GH weldment.</p> Full article ">Figure 3
<p>Example of two inserted thermocouples to register temperatures in Weld Pass 2. Holes of 2 mm in diameter were used to insert the thermocouples from the backside by drilling through the root pass. After the insertion, the thermocouples were spot welded on the surface of the joint—in this case, the surface of the root pass.</p> Full article ">Figure 4
<p>(<b>Left</b>) Back side of the joint showing the layout of the thermocouples inserted through drilled holes. (<b>Right</b>) Front side of the joint showing the layout of the harpooned thermocouples in the last passes. Note that the image shows the harpooned thermocouples only in the last passes because most of them have to be removed prior to welding of the following weld pass.</p> Full article ">Figure 5
<p>Temperature vs. time plots in Weld Pass 7 in the GR weld registered by using two thermocouples harpooned in the weld metal and two thermocouples inserted into drilled holes. The table embedded in the plot shows the time at which the deposited weld metal was exposed at the three ranges of temperatures studied.</p> Full article ">Figure 6
<p>Temperature vs. time plot given by the thermocouples used in the FR weld inserted in a furnace. All the thermocouples yielded signal and reacted to the heat pulses in the furnace. Most importantly, the measurements proved to be accurate, within a range of dispersion of 2 °C.</p> Full article ">Figure 7
<p>Average cooling rate between 1200 and 800 °C vs. weld pass number and welding procedure. Data for GR weld in Weld Pass 4 is not available because of a failure in the software that monitored the temperatures. The GH weld had only seven weld passes, which is why there are no data for Weld Passes 8, 9, and 10. Weld Pass 7 for FR and GR welds had exactly the same cooling rate (75 °C/s), which is why only one point is shown.</p> Full article ">Figure 8
<p>Ferrite content (Y-axis) measured by magnetic permeability pass by pass to evaluate the influence of the successive reheating on the ferrite content. A progressive reduction in ferrite content was observed in the weld metal. For better clarity of the reduction in ferrite, dotted lines show a linear regression of ferrite content in specific weld metal passes (Weld Pass 1 in the GH and GR welds and Weld Pass 4 in the FR weld). “WM pX” means the weld metal deposited in Weld Pass X.</p> Full article ">Figure 9
<p>Micrographs of Weld Pass 1 in the GH weld. (<b>Left</b>) As-deposited. It shows Widmanstätten austenite laths within the ferritic matrix. (<b>Right</b>) After six reheating cycles. Widmanstätten austenite laths grew (an average of 4 μm thicker) and secondary austenite (small needle-shaped) was formed in the intragranular areas of the ferritic matrix.</p> Full article ">Figure 10
<p>Influence of the grain size on the nitrides precipitation. First and second passes in GR weld cooled at the same cooling rate (38 °C/s). (<b>Left</b>) Root pass (GTAW) showing significant amounts of nitrides after only 6 s in the critical temperature range. Average ferrite grain size was 147 μm. (<b>Right</b>) Second pass (GMAW) not presenting nitrides after 6 s in the critical temperature range. Average ferrite grain size was 87 μm.</p> Full article ">Figure 11
<p>Preferential locations for precipitation of nitrides in the HAZ: primary ferrite grain boundaries, δ/ɣ boundaries and in the ferrite grains.</p> Full article ">Figure 12
<p>(<b>Left</b>) HTHAZ of Weld Pass 1 of the GH weld showing nitrides. (<b>Right</b>) HTHAZ of Weld Pass 1 of the GH weld after depositing two passes. Secondary austenite is found in locations of previous nitrides.</p> Full article ">Figure 13
<p>The σ-phase in Weld Pass 4, close to Weld Passes 5 and 6 in the FR weld.</p> Full article ">Figure 14
<p>Micrographs from the GR welds showing the microstructure of Weld Pass 1 (GTAW Root). (<b>Left</b>) Massive amounts of nitrides in the ferrite grains in the as-deposited condition. (<b>Center</b>) Secondary austenite formation in the ferrite grains after reheated caused by Weld Pass 2. (<b>Right</b>) The σ-phase after six reheats.</p> Full article ">Figure 15
<p>Charpy-V impact toughness results (CW: coupons extracted from the center of the weld section and with notch in the weld metal. CF: coupons extracted from the center of the weld section and with notch in the fusion line +2 mm). Tests performed at RT (room temperature) and at −46 °C.</p> Full article ">Figure 16
<p>Oxygen content in the root and the center of the welds.</p> Full article ">Figure 17
<p>Micro-inclusions in the weld deposited metal (shown as black dots in the micrographs). (<b>Left</b>) FCAW experiment with recommended arc energy. (<b>Center</b>) GMAW experiment with higher-than-recommended arc energy. (<b>Right</b>) GMAW experiment with recommended arc energy.</p> Full article ">Figure 18
<p>A pore (in black) surrounded in the majority of its perimeter by austenite phase (in white).</p> Full article ">
18 pages, 13806 KiB  
Article
MAG Welding Tests of Modern High Strength Steels with Minimum Yield Strength of 700 MPa
by Teemu Lahtinen, Pedro Vilaça, Pasi Peura and Saara Mehtonen
Appl. Sci. 2019, 9(5), 1031; https://doi.org/10.3390/app9051031 - 12 Mar 2019
Cited by 35 | Viewed by 6025
Abstract
The modern high strength steel plates have an excellent combination of strength and toughness based on micro-alloying and complex microstructure. Retaining this combination of properties in the weld zone is a major challenge for applications in high-demanding structural construction. This work investigates the [...] Read more.
The modern high strength steel plates have an excellent combination of strength and toughness based on micro-alloying and complex microstructure. Retaining this combination of properties in the weld zone is a major challenge for applications in high-demanding structural construction. This work investigates the weldability of three different modern high strength steel plates, with a thickness of 8 mm. Two of the test materials were produced by a thermo-mechanically controlled process (TMCP) and one by a quenching and tempering method (Q&T). Two-passes MAG (metal active gas) welding was used with four different heat inputs. The tests implemented on all the materials included tensile, hardness profiles (HV5), Charpy-V impact toughness tests, and microstructure analysis using scanning electron microscope (SEM). For one of the TMCP steels, some extended tests were conducted to define how the tensile properties change along the weld line. These tests included tensile tests with digital image correlation (DIC), and 3-point bending tests. The most notable differences in mechanical properties of the welds between the materials were observed in Charpy-V impact toughness tests, mostly at the vicinity of the fusion line, with the Q&T steel more prone to embrittlement of the heat affected zone (HAZ) than the TMCP steels. Microstructural analysis revealed carbide concentration combined with coarse bainitic structures in HAZ of Q&T steel, explaining the more severe embrittlement. During the tensile tests, the DIC measurements have shown a strain localization in the softest region of the HAZ. Increasing the heat input resulted in earlier localization of the strain and less maximum strength. The tensile properties along the weld line were investigated in all welding conditions, and the results emphasize relevant and systematic differences of the yield strength at the transient zones near the start and end of the weld compared with the intermediate stationary domain. Full article
(This article belongs to the Special Issue Welding of Steels)
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<p>Optical (above) and SEM (below) micrographs of microstructures of the three base materials. The microstructures of TMCP (MC) steels are fairly similar between each other, consisting of fine ferritic grains with small amount of second phase constituents at the grain boundaries. Deformation from the rolling can be seen as elongated grains. Base material of Q&amp;T steel consists mainly of tempered martensite.</p> Full article ">Figure 2
<p>Schematic presentation of the weld joint: (<b>a</b>) joint design; (<b>b</b>) welding sequence. All dimensions in [mm].</p> Full article ">Figure 3
<p>Locations of the specimens for tensile tests testing the difference in strength along the weld axis.</p> Full article ">Figure 4
<p>Charpy-V impact test specimens: (<b>a</b>) dimension; (<b>b</b>) notch locations with red vertical lines representing the locations: WM (weld metal), FL (fusion line) + 1 mm, FL + 3 mm and FL + 5 mm.</p> Full article ">Figure 5
<p>Elongations of the welds: (<b>a</b>) machined specimens; (<b>b</b>) specimens with reinforcements.</p> Full article ">Figure 6
<p>Yield and ultimate tensile strength of the welds: (<b>a</b>) machined specimens; (<b>b</b>) specimens with reinforcements.</p> Full article ">Figure 7
<p>The effect of the distance from the start of the weld on tensile test properties in the S700MC-1. The distance is measured to the center of the tensile specimen.</p> Full article ">Figure 8
<p>Results of DIC-measurements in specimens extracted from the steady-state domain from the welds of S700MC-1.</p> Full article ">Figure 9
<p>Average displacement-load/width curves for each heat input: (<b>a</b>) face side bending; (<b>b</b>) root side bending; (<b>c</b>) formation of the gap during testing induces the change in slope of the curve.</p> Full article ">Figure 10
<p>Charpy-V impact toughness results for the base materials. The error bars illustrate the standard deviation of the three individual tests.</p> Full article ">Figure 11
<p>Charpy-V impact toughness results for all the welds. Specimens having the notch at FL + 1 mm were tested at −20, −40 and −60 °C. Specimens having the notch at WM, FL + 3 mm and FL + 5 mm were tested only at −40 and −60 °C.</p> Full article ">Figure 12
<p>Hardness profiles of the welds of different materials with four different heat inputs.</p> Full article ">Figure 13
<p>Locations of the SEM images in the cross section of weld specimens: (<b>a</b>) low heat input, t<sub>8/5</sub> = 5 s; (<b>b</b>) high heat input, t<sub>8/5</sub> = 20 s.</p> Full article ">Figure 14
<p>SEM images from low heat input (t<sub>8/5</sub> = 5 s) weld HAZ for steel S700MC-1. (<b>A</b>–<b>D</b>) are locations established in <a href="#applsci-09-01031-f013" class="html-fig">Figure 13</a>.</p> Full article ">Figure 15
<p>SEM images from high heat input (t<sub>8/5</sub> = 20 s) weld HAZ for steel S700MC-1. (<b>A</b>–<b>D</b>) are locations established in <a href="#applsci-09-01031-f013" class="html-fig">Figure 13</a>.</p> Full article ">Figure 16
<p>SEM images from low heat input weld (t<sub>8/5</sub> = 5 s) HAZ for steel S700MC-2. (<b>A</b>–<b>D</b>) are locations established in <a href="#applsci-09-01031-f013" class="html-fig">Figure 13</a>.</p> Full article ">Figure 17
<p>SEM images from high heat input weld (t<sub>8/5</sub> = 20 s) HAZ for steel S700MC-2. (<b>A</b>–<b>D</b>) are locations established in <a href="#applsci-09-01031-f013" class="html-fig">Figure 13</a>.</p> Full article ">Figure 18
<p>SEM images from low heat input weld (t<sub>8/5</sub> = 5 s) HAZ for steel S690QL. (<b>A</b>–<b>D</b>) are locations established in <a href="#applsci-09-01031-f013" class="html-fig">Figure 13</a>.</p> Full article ">Figure 19
<p>SEM images from high heat input weld (t<sub>8/5</sub> = 20 s) HAZ for steel S690QL. (<b>A</b>–<b>D</b>) are locations established in <a href="#applsci-09-01031-f013" class="html-fig">Figure 13</a>.</p> Full article ">Figure 20
<p>Carbide concentrations in detail from location D of the HAZ of S690QL produced with the high heat input weld (t<sub>8/5</sub> = 20 s).</p> Full article ">
15 pages, 6148 KiB  
Article
Correlation of Microstructure and Mechanical Properties of Metal Big Area Additive Manufacturing
by Benjamin Shassere, Andrzej Nycz, Mark W. Noakes, Christopher Masuo and Niyanth Sridharan
Appl. Sci. 2019, 9(4), 787; https://doi.org/10.3390/app9040787 - 23 Feb 2019
Cited by 51 | Viewed by 8947
Abstract
Metal Big Area Additive Manufacturing (MBAAM) is a novel wire-arc additive manufacturing method that uses a correction-based approach developed at the Oak Ridge National Laboratory (ORNL). This approach is an integrated software method that minimizes the dynamic nature of welding and compensates for [...] Read more.
Metal Big Area Additive Manufacturing (MBAAM) is a novel wire-arc additive manufacturing method that uses a correction-based approach developed at the Oak Ridge National Laboratory (ORNL). This approach is an integrated software method that minimizes the dynamic nature of welding and compensates for build height. The MBAAM process is used to fabricate simple geometry thin walled specimens, using a C-Mn steel weld wire, to investigate the scatter in mechanical properties and correlate them to the underlying microstructure. The uni-axial tensile tests show isotropic tensile and yield properties with respect to building directions, although some scatter in elongation is observed. Large scatter is observed in the Charpy Impact tests. The microstructure characterization reveals mostly homogenous ferrite grains with some pearlite, except for some changes in morphology and grain size at the interface between the build and the base plate. The measured properties and microstructure are compared with the toughness and strength values reported in the literature, and a hypothesis is developed to rationalize the differences. Overall, the MBAAM process creates stable, isotropic, and weld-like mechanical properties in the deposit, while achieving a precise geometry obtained through a real-time feedback sensing, closed loop control system. Full article
(This article belongs to the Special Issue Welding of Steels)
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<p>Compensation effect on bead/layer roughness.</p> Full article ">Figure 2
<p>(<b>a</b>) Metal Big Area Additive Manufacturing (MBAAM) system with a printed excavator arm. (<b>b</b>) Build without the compensation-based approach. (<b>c</b>) Build with the compensation-based approach.</p> Full article ">Figure 3
<p>One of the MBAAM manufactured test walls.</p> Full article ">Figure 4
<p>Pictorial representation of the (<b>a</b>) metallographic, tensile (572 mm tall by 470 mm long) and (<b>b</b>) Charpy v-notch specimens (305 mm tall by 362 mm long).</p> Full article ">Figure 5
<p>Stress versus strain curves (<b>a</b>) perpendicular to build direction, (<b>b</b>) 45° to build direction, (<b>c</b>) parallel to build direction, and (<b>d</b>) graphical representation of yield and ultimate tensile stress and elongation averages with the standard deviation of results (the elongation average does not account for the red/orange curves from the interface).</p> Full article ">Figure 6
<p>Charpy v-notch impact toughness curves. The figure shows a Ductile to Brittle Transition Temperature (DBTT) increase from −36 °C to −17 °C with build height.</p> Full article ">Figure 7
<p>Fractography of samples fractured at −20 °C from the top of the build (<b>a</b>) and the bottom of the build (<b>b</b>). High magnification images (<b>c</b>) and (<b>d</b>) from the top and bottom of the build respectively. White arrows indicate regions of cleavage, and the black arrows indicate regions of ductile failure.</p> Full article ">Figure 8
<p>Hardness maps from different locations in the build height. (<b>a</b>) The last layer, top. (<b>b</b>) The middle of the build. (<b>c</b>)Bottom of the build. (<b>d</b>) interface between base plate and beginning of the build. Images show the Z–Y plane (thickness) of the build.</p> Full article ">Figure 9
<p>Microstructure from different locations in the build height. (<b>a</b>) The last layer, top. (<b>b</b>) The middle of the build. (<b>c</b>) Bottom of the build. (<b>d</b>) Interface between base plate and beginning of the build.</p> Full article ">Figure 10
<p>High magnification micrographs from different locations in the build height. (<b>a</b>) The last layer, top. (<b>b</b>) The middle of the build. (<b>c</b>) Bottom of the build. (<b>d</b>) Interface region between base plate and beginning of the build.</p> Full article ">Figure 11
<p>Continuous Cooling Transformation (CCT) diagrams for 0.1 wt% C. The CCT diagram is used to understand the evolution of the microstructure on cooling from face centered cubic (FCC) austenite. Drawn using the thermodynamic software JMatPro<sup>®</sup>.</p> Full article ">
20 pages, 3575 KiB  
Article
Determination of Residual Welding Stresses in a Steel Bridge Component by Finite Element Modeling of the Incremental Hole-Drilling Method
by Evy Van Puymbroeck, Wim Nagy, Ken Schotte, Zain Ul-Abdin and Hans De Backer
Appl. Sci. 2019, 9(3), 536; https://doi.org/10.3390/app9030536 - 5 Feb 2019
Cited by 21 | Viewed by 3798
Abstract
For welded bridge components, the knowledge of residual stresses induced by welding is essential to determine their effect on the fatigue life behavior resulting in optimal fatigue design and a better knowledge about the fatigue strength of these welded connections. The residual stresses [...] Read more.
For welded bridge components, the knowledge of residual stresses induced by welding is essential to determine their effect on the fatigue life behavior resulting in optimal fatigue design and a better knowledge about the fatigue strength of these welded connections. The residual stresses of a welded component in an orthotropic steel bridge deck are determined with the incremental hole-drilling method. This method is specified by the American Society for Testing and Materials ASTM E837-13a and it can be used only when the material behavior is linear-elastic. However in the region of the bored hole, there are relaxed plastic strains present that can lead to significant error of the measured stresses. The hole-drilling procedure is simulated with three-dimensional finite element modeling including a simplistic model of plasticity. The effect of plasticity on uniform in-depth residual stresses is determined and it is concluded that residual stresses obtained under the assumption of linear-elastic material behavior are an overestimation. Including plasticity for non-uniform in-depth residual stress fields results in larger tensile and smaller compressive residual stresses. Larger tensile residual stresses cause premature fatigue failure. Therefore, it is important to take these larger tensile residual stresses into account for the fatigue design of a welded component. Full article
(This article belongs to the Special Issue Welding of Steels)
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Graphical abstract
Full article ">Figure 1
<p>Plastic zone introduced by hole drilling.</p> Full article ">Figure 2
<p>RS-200 Milling Guide (<b>left</b>) and hole-drilled strain gauge rosette (<b>right</b>).</p> Full article ">Figure 3
<p>Schematic geometry of a strain gauge rosette (<b>a</b>) and detail of one strain gauge (<b>b</b>).</p> Full article ">Figure 4
<p>Linear-elastic stress–strain behavior.</p> Full article ">Figure 5
<p>Elastic-plastic stress–strain behavior.</p> Full article ">Figure 6
<p>(<b>a</b>) Uniform in-depth and (<b>b</b>) non-uniform in-depth residual stress field.</p> Full article ">Figure 7
<p>Finite element model with drilled hole.</p> Full article ">Figure 8
<p>Mesh sensitivity study.</p> Full article ">Figure 9
<p>Comparison between linear-elastic and elastic-plastic material behavior from 3D models.</p> Full article ">Figure 10
<p>Full scale test specimen (<b>a</b>) top view; (<b>b</b>) cross section; (<b>c</b>) bottom view between two longitudinal stiffeners.</p> Full article ">Figure 11
<p>Locations of measured strain gauge rosettes (dimensions in mm).</p> Full article ">Figure 12
<p>General residual stress distribution of the orthotropic steel deck from experimental measurements (f<sub>y</sub> = yield strength) for the longitudinal direction (<b>a</b>) and the transverse direction (<b>b</b>).</p> Full article ">Figure 13
<p>Comparison of residual stress distribution on the top of the deck plate between ASTM and the 3D model in the longitudinal (<b>a</b>) and transverse (<b>b</b>) directions.</p> Full article ">Figure 14
<p>Comparison of residual stress distribution on the bottom of the deck plate between ASTM and the 3D model in the longitudinal (<b>a</b>) and transverse (<b>b</b>) directions.</p> Full article ">Figure 15
<p>Comparison of residual stress distribution on the longitudinal stiffener between ASTM and the 3D model in the longitudinal (<b>a</b>) and transverse (<b>b</b>) directions.</p> Full article ">Figure 16
<p>Residual stress distribution of the orthotropic steel deck according FEM for the (<b>a</b>) longitudinal and (<b>b</b>) transverse directions.</p> Full article ">
18 pages, 14852 KiB  
Article
Stability and Heat Input Controllability of Two Different Modulations for Double-Pulse MIG Welding
by Jiaxiang Xue, Min Xu, Wenjin Huang, Zhanhui Zhang, Wei Wu and Li Jin
Appl. Sci. 2019, 9(1), 127; https://doi.org/10.3390/app9010127 - 1 Jan 2019
Cited by 16 | Viewed by 4502
Abstract
Aluminum alloy welding frequently experiences difficulties such as heat input control, poor weld formation, and susceptibility to pore generation. We compared the use of two different modulations for double-pulse metal inert gas (MIG) welding to reduce the heat input required to generate oscillations [...] Read more.
Aluminum alloy welding frequently experiences difficulties such as heat input control, poor weld formation, and susceptibility to pore generation. We compared the use of two different modulations for double-pulse metal inert gas (MIG) welding to reduce the heat input required to generate oscillations in the weld pool. The stabilities of rectangular wave-modulated and trapezoidal wave-modulated double-pulse MIG welding (DP-MIG and TP-MIG) were analyzed by examining their welding processes and weld profiles. We found that the transitional pulse in TP-MIG welding results in smoother current transitions, softer welding arc sounds, and a highly uniform fish-scale pattern. Therefore, TP-MIG welding is more stable than DP-MIG welding. The effects of these double-pulse modulation schemes on welding input energy are presented. We propose methods for reducing welding input energy by varying the number of pulses or the pulse base time of low-energy pulse train while keeping the welding current and welding arc stable and unchanged. Compared to DP-MIG welding, TP-MIG welding reduces the input energy by 12% and produces finer grain sizes, which increases weld hardness. Therefore, TP-MIG welding offers a new approach for heat input control in DP-MIG welding of aluminum alloys. The results of this work are significant for aluminum alloy welding. Full article
(This article belongs to the Special Issue Welding of Steels)
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<p>Relationship between transfer mode and current.</p> Full article ">Figure 2
<p>Sketch of the P-MIG scheme.</p> Full article ">Figure 3
<p>DP-MIG current waveform diagram.</p> Full article ">Figure 4
<p>TP-MIG low-frequency modulation waveform.</p> Full article ">Figure 5
<p>High-frequency modulation signal energy pulse group graph: (<b>a</b>) High-energy pulse train of the high-frequency modulation signal; (<b>b</b>) Low-energy pulse train of the high-frequency modulation signal.</p> Full article ">Figure 6
<p>Transition pulse group of high-frequency modulation signal: (<b>a</b>) High- and low-energy transition pulse group; (<b>b</b>) Low-energy transition pulse group.</p> Full article ">Figure 7
<p>1.2-mm calibration curve step distance of aluminum alloy welding wire.</p> Full article ">Figure 8
<p>Waveform and weld formation at 80 A and 160 A: (<b>a</b>) 80 A welding waveform; (<b>b</b>) 160 A welding waveform; (<b>c</b>) 80 A weld formation; (<b>d</b>) 160 A weld formation.</p> Full article ">Figure 8 Cont.
<p>Waveform and weld formation at 80 A and 160 A: (<b>a</b>) 80 A welding waveform; (<b>b</b>) 160 A welding waveform; (<b>c</b>) 80 A weld formation; (<b>d</b>) 160 A weld formation.</p> Full article ">Figure 9
<p>Weld seam obtained by parameter self-adjusting welding test; (<b>a</b>) 83 A weld formation; (<b>b</b>) 92 A weld formation; (<b>c</b>) 135 A weld formation.</p> Full article ">Figure 10
<p>120 A waveform and weld formation at different frequencies: (<b>a</b>) 2 Hz; (<b>b</b>) 5 Hz; (<b>c</b>) 15 Hz.</p> Full article ">Figure 11
<p>DP-MIG weld under different average currents: (<b>a</b>) Average current of 88 A; (<b>b</b>) Average current of 80 A.</p> Full article ">Figure 12
<p>TP-MIG weld under different average currents: (<b>a</b>) Average current of 88 A; (<b>b</b>) Average current of 80 A.</p> Full article ">Figure 13
<p>DP-MIG current waveform diagram of different number of pulses in the low-energy pulse train: (<b>a</b>) Average current of 88 A; (<b>b</b>) Average current of 80 A.</p> Full article ">Figure 14
<p>TP-MIG current waveform diagram of different number of pulses in the low-energy pulse train: (<b>a</b>) Average current of 88 A; (<b>b</b>) Average current of 80 A.</p> Full article ">Figure 15
<p>Micrograph of the welds of DP-MIG under different average currents: (<b>a</b>) Average current of 88 A; (<b>b</b>) Average current of 80 A.</p> Full article ">Figure 16
<p>Micrograph of the welds of TP-MIG under different average currents: (<b>a</b>) Average current of 88 A; (<b>b</b>) Average current of 80 A.</p> Full article ">Figure 17
<p>DP-MIGW at different frequencies: (<b>a</b>) Average current of 90 A, frequency of 3.8 Hz; (<b>b</b>) Average current of 80 A, frequency of 3 Hz.</p> Full article ">Figure 18
<p>TP-MIGW at different frequencies; (<b>a</b>) Average current of 90 A, frequency of 3 Hz; (<b>b</b>) Average current of 80 A, frequency of 2.4 Hz.</p> Full article ">Figure 18 Cont.
<p>TP-MIGW at different frequencies; (<b>a</b>) Average current of 90 A, frequency of 3 Hz; (<b>b</b>) Average current of 80 A, frequency of 2.4 Hz.</p> Full article ">Figure 19
<p>Sample extraction of welded joint.</p> Full article ">Figure 20
<p>DP-MIGW welding joint at 1.0 m/min.</p> Full article ">Figure 21
<p>TP-MIGW welding joint at 1.0 m/min.</p> Full article ">Figure 22
<p>Hardness distribution of welded joints at two speeds: (<b>a</b>) Comparison of DP-MIGW hardness of welded joints; (<b>b</b>) Comparison of TP-MIGW hardness of welded joints.</p> Full article ">Figure 23
<p>Difference in the hardness of welded joints at two speeds: (<b>a</b>) DP-MIGW hardness of welded joint; (<b>b</b>) DP-MIGW hardness of welded joint.</p> Full article ">Figure 24
<p>Microstructure photograph of welded joints: (<b>a</b>) DP-MIGW, 0.6 m/min, (<b>b</b>) DP-MIGW, 1.0 m/min, (<b>c</b>) TP-MIGW, 0.6 m/min, (<b>d</b>) TP-MIGW, 1.0 m/min.</p> Full article ">Figure 25
<p>Tensile properties of welded joints: (<b>a</b>) Tensile curves; (<b>b</b>) Ultimate tensile strength.</p> Full article ">Figure 26
<p>Photographs of fractured tensile samples: (<b>a</b>) DP-MIGW, 0.6 m/min, (<b>b</b>) DP-MIGW, 1.0 m/min, (<b>c</b>) TP-MIGW, 0.6 m/min, (<b>d</b>) TP-MIGW, 1.0 m/min.</p> Full article ">
24 pages, 12231 KiB  
Article
Enhancement of Exit Flow Uniformity by Modifying the Shape of a Gas Torch to Obtain a Uniform Temperature Distribution on a Steel Plate during Preheating
by Thien Tu Ngo, Junho Go, Tianjun Zhou, Hap Van Nguyen and Geun Sik Lee
Appl. Sci. 2018, 8(11), 2197; https://doi.org/10.3390/app8112197 - 9 Nov 2018
Cited by 10 | Viewed by 5934
Abstract
The objective of this study is to improve the exit flow uniformity of a gas torch with multiple exit holes for effective heating of a steel plate. The torch was simulated, and combustion experiments were performed for validation. Based on a basic model, [...] Read more.
The objective of this study is to improve the exit flow uniformity of a gas torch with multiple exit holes for effective heating of a steel plate. The torch was simulated, and combustion experiments were performed for validation. Based on a basic model, three different revised models were designed and analyzed with the software ANSYS FLUENT 18.2. The flow uniformity (γ) of the velocity distribution at the multiple exit holes was investigated with the pressure drop ranging from 100 to 500 Pa. The basic model had flow uniformity ranging from 0.849 to 0.852, but the three new models had γ1 = 0.901–0.912, γ2 = 0.902–0.911, and γ3 = 0.901–0.914, respectively. The maximum percentage difference of the flow uniformity index between the three new models and the basic model was 7.3%. The basic model with nonuniform flow distribution made a temperature difference of the back side of the steel plate from the center to the edge of around 229 °C, while the modified model with uniform flow distribution had a smaller temperature difference of 90 °C. The simulation results showed good agreement with our experimental results for both the basic model and the modified model. The modified gas torch made a wider and more uniform temperature distribution on a preheated steel plate than the basic one. The results revealed that a trade-off between cost and flow uniformity, as well as the new gas torch, could be applied to a steel-plate preheating process before welding. Full article
(This article belongs to the Special Issue Welding of Steels)
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<p>Comparison of typical flow field and temperature distribution of a steel plate: (<b>a</b>) basic gas torch; (<b>b</b>) modified gas torch.</p> Full article ">Figure 2
<p>Diagram of a basic gas torch.</p> Full article ">Figure 3
<p>The inner and outer shapes of three new gas torch models: (<b>a</b>) model 1; (<b>b</b>) model 2; (<b>c</b>) model 3.</p> Full article ">Figure 3 Cont.
<p>The inner and outer shapes of three new gas torch models: (<b>a</b>) model 1; (<b>b</b>) model 2; (<b>c</b>) model 3.</p> Full article ">Figure 4
<p>Meshes and boundary conditions of a three-dimensional gas torch.</p> Full article ">Figure 5
<p>Grid independence check for velocity at the outlets along the lengthwise (x) direction.</p> Full article ">Figure 6
<p>(<b>a</b>) Basic gas torch pressure contours; (<b>b</b>) inner streamlines and velocity vectors at the outlets.</p> Full article ">Figure 6 Cont.
<p>(<b>a</b>) Basic gas torch pressure contours; (<b>b</b>) inner streamlines and velocity vectors at the outlets.</p> Full article ">Figure 7
<p>Pressure contours at the middle plane of the three new gas torches: (<b>a</b>) model 1; (<b>b</b>) model 2; (<b>c</b>) model 3.</p> Full article ">Figure 7 Cont.
<p>Pressure contours at the middle plane of the three new gas torches: (<b>a</b>) model 1; (<b>b</b>) model 2; (<b>c</b>) model 3.</p> Full article ">Figure 8
<p><b>Figure</b><b>8.</b> Inner streamlines and velocity vectors at the outlet of the three new gas torches: (<b>a</b>) model 1; (<b>b</b>) model 2; (<b>c</b>) model 3.</p> Full article ">Figure 8 Cont.
<p><b>Figure</b><b>8.</b> Inner streamlines and velocity vectors at the outlet of the three new gas torches: (<b>a</b>) model 1; (<b>b</b>) model 2; (<b>c</b>) model 3.</p> Full article ">Figure 9
<p>Exit velocity distribution along the lengthwise direction of the basic and three new gas torches.</p> Full article ">Figure 10
<p>Comparison of velocity uniformity index at the outlet between the basic torch and the three new gas torches as a function of pressure drop.</p> Full article ">Figure 11
<p>Schematic diagram of steel preheating process with gas torch, fluid domain, and steel plate. (<b>a</b>) A half model (x-y plane cut view); (<b>b</b>) three-dimensional magnified view of the quarter model; (<b>c</b>) mesh model of the quarter model.</p> Full article ">Figure 11 Cont.
<p>Schematic diagram of steel preheating process with gas torch, fluid domain, and steel plate. (<b>a</b>) A half model (x-y plane cut view); (<b>b</b>) three-dimensional magnified view of the quarter model; (<b>c</b>) mesh model of the quarter model.</p> Full article ">Figure 12
<p>The temperature contours of a combustion flame zone at the center-line view during the steel-plate preheating process: (<b>a</b>) basic model; (<b>b</b>) modified model.</p> Full article ">Figure 12 Cont.
<p>The temperature contours of a combustion flame zone at the center-line view during the steel-plate preheating process: (<b>a</b>) basic model; (<b>b</b>) modified model.</p> Full article ">Figure 13
<p>The species mole fraction and temperature variations of the stoichiometric LPG/air mixture for the basic model, modified model, and Akram’s results [<a href="#B25-applsci-08-02197" class="html-bibr">25</a>].</p> Full article ">Figure 14
<p>Experimental apparatus setup: (<b>a</b>) basic model; (<b>b</b>) modified model; (<b>c</b>) back side (1: thermocouples, 2: TVS-200EX infrared camera, 3: recorder for thermocouples).</p> Full article ">Figure 14 Cont.
<p>Experimental apparatus setup: (<b>a</b>) basic model; (<b>b</b>) modified model; (<b>c</b>) back side (1: thermocouples, 2: TVS-200EX infrared camera, 3: recorder for thermocouples).</p> Full article ">Figure 15
<p>Steel-plate temperature distribution at the back side with incomplete combustion (Φ = 1.4): (<b>a</b>) basic model experiment; (<b>b</b>) basic model simulation; (<b>c</b>) modified model experiment; (<b>d</b>) modified model simulation.</p> Full article ">Figure 15 Cont.
<p>Steel-plate temperature distribution at the back side with incomplete combustion (Φ = 1.4): (<b>a</b>) basic model experiment; (<b>b</b>) basic model simulation; (<b>c</b>) modified model experiment; (<b>d</b>) modified model simulation.</p> Full article ">Figure 16
<p>Steel-plate temperature distribution at the back side with complete combustion (Φ = 1.0): (<b>a</b>) basic model experiment; (<b>b</b>) basic model simulation; (<b>c</b>) modified model experiment; (<b>d</b>) modified model simulation.</p> Full article ">Figure 17
<p>Comparison of the temperature distribution at the back side of the steel plate from the center to the edge for complete combustion.</p> Full article ">
13 pages, 7010 KiB  
Article
High-Speed Welding of Stainless Steel with Additional Compensatory Gas Jet Blow Molten Pool
by Changwen Dong, Jiaxiang Xue, Zhanhui Zhang, Li Jin, Yu Hu and Wei Wu
Appl. Sci. 2018, 8(11), 2170; https://doi.org/10.3390/app8112170 - 6 Nov 2018
Cited by 7 | Viewed by 5357
Abstract
To avoid humping bead defects in high-speed welding, this paper proposes the method of an additional and compensatory gas jet blow molten pool. A pulsed metal inert gas high-speed welding test platform was constructed for compensatory gas jet blow molten pool. A total [...] Read more.
To avoid humping bead defects in high-speed welding, this paper proposes the method of an additional and compensatory gas jet blow molten pool. A pulsed metal inert gas high-speed welding test platform was constructed for compensatory gas jet blow molten pool. A total of 304 stainless steel sheets were used as the welding workpieces under equal heat inputs. Two high-speed butt welding processes were conducted and compared, in which the workpieces were welded with and without compensatory gas jets at 154 cm/min and 167 cm/min, respectively. After high-speed welding with compensatory gas jet blow, the weld appearance was straight, uniform, and high-quality, with no humping bead or undercut defects. The macroscopic morphologies and microstructures of cross-sections of the weld at the toe, near the surface, the middle, and the bottom portion all showed the stirring effect of the gas jet on the molten pool and improved grain refinement degrees. Hardness was enhanced in the weld center and the heat-affected zone. At welding speeds of 154 cm/min and 167 cm/min, the fracture load capacities of the welds were increased by 24.9 and 10.4%, respectively. Full article
(This article belongs to the Special Issue Welding of Steels)
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<p>Schematic of the relative speeds of molten steel flows and parent material: (<b>a</b>) Low relative speed (high-speed welding); (<b>b</b>) high relative speed (normal welding).</p> Full article ">Figure 2
<p>Schematic of the molten pool blown with compensatory gas jet during welding.</p> Full article ">Figure 3
<p>Self-made compensatory gas jet test platform: (<b>a</b>) Diagram of the compensatory gas jet system; (<b>b</b>) diagram of compensatory gas jet nozzle structure.</p> Full article ">Figure 4
<p>Current and voltage waveforms during the butt welding of 304 stainless steel: (<b>a</b>) Sample 1#; (<b>b</b>) sample 2#.</p> Full article ">Figure 5
<p>Weld bead shape of sample 1#: (<b>a</b>) Front view; (<b>b</b>) back view.</p> Full article ">Figure 6
<p>Weld bead shape of sample 2#: (<b>a</b>) Front view; (<b>b</b>) back view.</p> Full article ">Figure 7
<p>Microstructures of butt welds formed at high-speed welding.</p> Full article ">Figure 8
<p>Weld hardness test point.</p> Full article ">Figure 9
<p>Comparison of hardness curves of welds formed with and without compensatory gas jets.</p> Full article ">Figure 10
<p>Dimensions of tensile strength test specimens.</p> Full article ">Figure 11
<p>Tensile properties of welded joints: (<b>A</b>) Tensile samples; (<b>B</b>) tensile curves; (<b>C</b>) result of tensile tests.</p> Full article ">Figure 12
<p>Photographs of fractured tensile samples: (<b>a</b>) Weld joint, (<b>b</b>) base material.</p> Full article ">
13 pages, 8222 KiB  
Article
Effect of Additional Shielding Gas on Welding Seam Formation during Twin Wire DP-MIG High-Speed Welding
by Yu Hu, Jiaxiang Xue, Changwen Dong, Li Jin and Zhanhui Zhang
Appl. Sci. 2018, 8(9), 1658; https://doi.org/10.3390/app8091658 - 14 Sep 2018
Cited by 8 | Viewed by 4582
Abstract
For diminishing welding defects such as incomplete penetration, which may easily occur during the twin wire Double Pulsed Metal Inert Gas (DP-MIG) high-speed welding, a novel method using additional shielding gas is introduced in this paper. A branch for the additional shielding gas [...] Read more.
For diminishing welding defects such as incomplete penetration, which may easily occur during the twin wire Double Pulsed Metal Inert Gas (DP-MIG) high-speed welding, a novel method using additional shielding gas is introduced in this paper. A branch for the additional shielding gas was specially set near the back end of the protection hood for the DP-MIG nozzle. The constructed gas branch was used for enabling manual intervention in the formation of a high-temperature solid–liquid weld seam just emerging from the nozzle and also for secondary gas protection on the surface of the weld seam. The butt welding test was carried out in the 2205 duplex stainless steel plate and the weld seam was then characterized by a tensile test, metallographic analysis, X-ray non-destructive testing (NDT), hardness analysis, and impact test. The results showed that the introduction of an appropriate amount of additional shielding gas can effectively improve and diminish the unfused weld seam and also improve the mechanical properties such as the tensile properties of the weld joint, the hardness and toughness of the weld joints. Therefore, the introduction of additional shielding gas has further research potential in theory and process practice. Full article
(This article belongs to the Special Issue Welding of Steels)
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<p>Schematic diagram of the twin wire pulsed MIG welding under additional shielding gas.</p> Full article ">Figure 2
<p>Dimensions of test specimens (unit: mm): (<b>a</b>) Draft of plate and all dimensions, (<b>b</b>) Specimen for tensile test, (<b>c</b>) Specimen for Charpy impact test.</p> Full article ">Figure 3
<p>Morphology of the weld seam: (<b>a</b>) Penetration depth and width/depth, (<b>b</b>) The cross-section of joint of Sample 1, (<b>c</b>) The cross-section of joint of Sample 2, (<b>d</b>) The cross-section of joint of Sample 3, (<b>e</b>) The cross-section of joint of Sample 4, (<b>f</b>) The cross-section of joint of Sample 5, (<b>g</b>) The cross-section of joint of Sample 6, (<b>h</b>) the comparison before and after the introduction of additional shielding gas.</p> Full article ">Figure 3 Cont.
<p>Morphology of the weld seam: (<b>a</b>) Penetration depth and width/depth, (<b>b</b>) The cross-section of joint of Sample 1, (<b>c</b>) The cross-section of joint of Sample 2, (<b>d</b>) The cross-section of joint of Sample 3, (<b>e</b>) The cross-section of joint of Sample 4, (<b>f</b>) The cross-section of joint of Sample 5, (<b>g</b>) The cross-section of joint of Sample 6, (<b>h</b>) the comparison before and after the introduction of additional shielding gas.</p> Full article ">Figure 4
<p>X-ray NDT photograph: (<b>a</b>) X-ray NDT for weld seam in Sample 1, (<b>b</b>) X-ray NDT for weld seam in Sample 2.</p> Full article ">Figure 5
<p>Metallographic test results: (<b>a</b>) Metallographic image of HAZ in Sample 1, (<b>b</b>) Metallographic image of HAZ in Sample 2, (<b>c</b>) Metallographic image of HAZ in Sample 3, (<b>d</b>) Metallographic image of HAZ in Sample 4, (<b>e</b>) Metallographic image of HAZ in Sample 5, (<b>f</b>) Metallographic image of HAZ in Sample 6, (<b>g</b>) Metallographic image of fusion zone in Sample 1, (<b>h</b>) Metallographic image of fusion zone of in Sample 2, (<b>i</b>) Metallographic image of fusion zone in Sample 3, (<b>j</b>) Metallographic image of fusion zone in Sample 4, (<b>k</b>) Metallographic image of fusion zone in Sample 5, (<b>l</b>) Metallographic image of fusion zone in Sample 6.</p> Full article ">Figure 5 Cont.
<p>Metallographic test results: (<b>a</b>) Metallographic image of HAZ in Sample 1, (<b>b</b>) Metallographic image of HAZ in Sample 2, (<b>c</b>) Metallographic image of HAZ in Sample 3, (<b>d</b>) Metallographic image of HAZ in Sample 4, (<b>e</b>) Metallographic image of HAZ in Sample 5, (<b>f</b>) Metallographic image of HAZ in Sample 6, (<b>g</b>) Metallographic image of fusion zone in Sample 1, (<b>h</b>) Metallographic image of fusion zone of in Sample 2, (<b>i</b>) Metallographic image of fusion zone in Sample 3, (<b>j</b>) Metallographic image of fusion zone in Sample 4, (<b>k</b>) Metallographic image of fusion zone in Sample 5, (<b>l</b>) Metallographic image of fusion zone in Sample 6.</p> Full article ">Figure 5 Cont.
<p>Metallographic test results: (<b>a</b>) Metallographic image of HAZ in Sample 1, (<b>b</b>) Metallographic image of HAZ in Sample 2, (<b>c</b>) Metallographic image of HAZ in Sample 3, (<b>d</b>) Metallographic image of HAZ in Sample 4, (<b>e</b>) Metallographic image of HAZ in Sample 5, (<b>f</b>) Metallographic image of HAZ in Sample 6, (<b>g</b>) Metallographic image of fusion zone in Sample 1, (<b>h</b>) Metallographic image of fusion zone of in Sample 2, (<b>i</b>) Metallographic image of fusion zone in Sample 3, (<b>j</b>) Metallographic image of fusion zone in Sample 4, (<b>k</b>) Metallographic image of fusion zone in Sample 5, (<b>l</b>) Metallographic image of fusion zone in Sample 6.</p> Full article ">Figure 6
<p>Rockwell hardness distribution in welded joints.</p> Full article ">Figure 7
<p>Tensile properties of the base metal and welded joint: (<b>a</b>) Tensile curves, (<b>b</b>) Result of tensile tests.</p> Full article ">Figure 8
<p>Photographs of fractured tensile samples: (<b>a</b>) Sample 1, (<b>b</b>) Sample 2.</p> Full article ">
11 pages, 1964 KiB  
Article
Fatigue Strength Assessment of Steel Rollers: On the Reliability of the Strain Energy Density Approach on Real Components
by Francesco Chebat, Mirco Peron, Luigi Mario Viespoli, Torgeir Welo and Filippo Berto
Appl. Sci. 2018, 8(7), 1015; https://doi.org/10.3390/app8071015 - 21 Jun 2018
Cited by 14 | Viewed by 4227
Abstract
Welded joints are one of the most widely applied methods to join different steel components. However, they introduce stress concentrators that are commonly known to reduce the fatigue strength of structures. Several methods have been developed to assess the fatigue behavior of welded [...] Read more.
Welded joints are one of the most widely applied methods to join different steel components. However, they introduce stress concentrators that are commonly known to reduce the fatigue strength of structures. Several methods have been developed to assess the fatigue behavior of welded components, such as the Notch Stress Intensity Factors (NSIFs) approach. However, this approach has been reported to be geometry dependent, and does not allow for a direct comparison of failures occurring at the weld toes with those occurring at the weld roots. This drawback has been overcame by considering the value of the strain energy density (SED) range averaged in a control volume ahead of the notch tip. More than 900 fatigue data of welded joints have been summarized within a single scatter band ΔW-N (strain energy range–umber of cycles to failure) using this approach. The reliability of the just mentioned scatter band in summarizing the fatigue data of real components such as steel welded rollers produced by Rulmeca is herein evaluated. The results prove the reliability of the SED approach to assess the fatigue behavior of welded rollers, paving the way to its diffusion in assessing real components. Full article
(This article belongs to the Special Issue Welding of Steels)
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<p>Fatigue behavior of welded joints as a function of the averaged local strain energy density range Δ<span class="html-italic">W</span>; <span class="html-italic">R</span> is the nominal load ratio.</p> Full article ">Figure 2
<p>Typical roller assembly.</p> Full article ">Figure 3
<p>Schematic representation of the weld root, characterized by a weld root length “<span class="html-italic">c</span>”, and the control volume of the strain energy density (SED) approach (blue circle).</p> Full article ">Figure 4
<p>Lack of penetration at the weld root (<b>a</b>), a detail of the crack through the welded zone (<b>b</b>) and typical roller failure (<b>c</b>).</p> Full article ">Figure 4 Cont.
<p>Lack of penetration at the weld root (<b>a</b>), a detail of the crack through the welded zone (<b>b</b>) and typical roller failure (<b>c</b>).</p> Full article ">Figure 5
<p>Geometry of the rollers.</p> Full article ">Figure 6
<p>Synthesis of fatigue data for different roller geometries in terms of local SED and a comparison with the scatter band shown in <a href="#applsci-08-01015-f001" class="html-fig">Figure 1</a>.</p> Full article ">
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