according to the standard ASTM-E8 . The dog bone type specimens had a gauge length of 55 mm and a diameter of 9.86 mm . The tensile tests were carried out at 5 mm/min at room temperature. A DCR-HC32E handy cam with 30 frames per second was used to capture the deformation of specimens before and after necking. From the recorded images, the dimensions and profile of the neck were measured. The resolutions of the images were enhanced using graphical software and the required data were extracted using point detection software named as Gate Data and Digitizer. Typical deformation of marked specimens is illustrated in Figure 2. The load-displacement and stressstrain curves of the specimen are shown in Figures 3 and 4, respectively. The engineering stress-straincurve, the truestress-strain diagram obtained using Equation (4) and the truestress-straincurve extracted from image processing (strain is calculated from Equation (5) and stress is computed from P/A) are illustrated in Figure 4. As the figure suggests, the two types of the truestress- strain curves are quite different after necking. In order to validate the image processing used in this work, the smallest neck area was measured from the images taken by camera and was calculated using the relation  :
For a U-bending specimen is stressed, outer surface of the material may undergo plastic deformation of the truestress-straincurve. Figure 6(a) is shown. Figure 6(b) to 6(e) are show several stress-strain relationships that may exist in the outer surface of the U-bending process. Method of the stressing is influence the actual relationship.
condition that the two curves could be ﬁtted by similar functions, the cyclic curve could be deduced by the engineering curve theoretically. Meanwhile, although the stress-straincurve of many metals in the region of uniform plastic deformation could be described by the classical Hollomon equation, the equation might not be absolutely appropriate to describe the engineering stress-straincurve. The engineering stress-straincurve does not give a true indication of the deformation characteristics, because it is based entirely on the original dimensions of the specimen, and these dimensions would change continuously during the tensile test. Actually, the truestress-straincurve 13) presents a larger stress value compared with the engineering curve as the cross-sectional area of the specimen is gradually decreasing.
The shape of a load-displacement curve or a nominal stress-straincurve also changes according to deformation method such as tension, compression and bend even if the same ductile metal with an identical truestress-strain relation is considered. 10) In other words, boundary condition deﬁ- nitely aﬀects the shape of the nominal stress-straincurve. Finite element (FE) analysis has been successful in simulat- ing and assessing the plastic deformation behavior for metal forming and also widely accepted as a powerful tool to
Figure 2 shows the truestress–strain curves and variation of nð"Þ with strain. The values in the uniform elongation range, plotted in the ﬁgure, were achieved directly from the measured load–nominal strain data. After the diﬀuse necking starts, at ﬁrst almost no change of nð"Þ was observed. Subsequently, because of the decrease in nð"Þ with increasing the strain, the stress–straincurve, obtained from the power- hardening law, was overestimated. However, the diﬀerence was not so large, and a quite good approximation could be achieved even using the power-hardening law. In addition to HTSS sheets, other annealed metals with large n-values also exhibit good approximations of stress–strain curves when the power-hardening law is used (see Section 4.3). On the other hand, it was observed from the stress–strain curves shown in Fig. 2 that the stress increase due to strain-hardening became gentle in the post-uniform elongation range, although a large
Fig. 2. Typical deformation curve obtained in the current tests. The specimen rupture is described by the strength parameters: Truestrain at the strength limit (TSNS) and truestress at the strength limit (TSSS). The curve is divided into two parts, the initial quasi-elastic part below the yield and the flow part above the yield. The yield is defined by two parameters: Truestrain at the yield limit (TSNY) and truestress at the yield limit (TSSY). The linear part below yield point gives by their slope modulus of elasticity (E), the slope of linear part above yield point gives modulus of flowing (E w )
basic relationship for a wrought magnesium alloy with a wide range of strain rates; the results showed that the flow curve reaches a steady state at low strain values. Mishra et al.  encountered an abnormal change in the behaviour of Mg due to the addition of Ce. Luo et al.  studied high ductility magnesium-zinc-cerium extrusion alloys. Their studies showed that by adding a small amount of zinc, the strength of the Mg-Ce alloy can be improved notably. Luo et al.  also examined the microstructure and mechanical properties of extruded magnesium alloy tubes; the results revealed that the mechanical properties can be improved with the addition of Ce to pure magnesium. Chino et al.  investigated the compressive properties of Mg alloys from room temperature up to 500 °C to understand the effects of Ce on the deformation process; the results showed that the addition of Ce increased the ductility of Mg alloy at room temperature, but decreased it at 300 °C. El-Morsy et al.  investigated the microstructural evolution of AZ61. They achieved fine grain size via a combination of hot extrusion and thermomechanical processing.
In recent decades, mitigation of susceptibility to stress corrosion cracking has been a major challenge for material scientists in the nuclear industry. Sensitization has been prevented mainly by reducing carbon content of the steels. As a side effect, high cycle fatigue strength of the nuclear grade stainless steels have probably decreased. If this is true, it might explain part of the differences between recent experimental data and the Langer curve. On the other hand, our current results provide supporting arguments for use of stabilized stainless steels in components, where high cycle fatigue is a concern.
In this paper, authors describe the derivation of the truestress-true plastic straincurve and other mechanical parameters from the indentation load-depth curve measured by the ABI tests. To consider SA508Gr3, SA516Gr70, and SA533B as the representative for the steels used in nuclear vessels, by comparing the truestress-true plastic straincurve derived from ABI tests with those measured directly by the uniaxial tensile tests, the author verified the feasibility of using the ABI Technique to obtain mechanical properties of the steels used in nuclear vessels.
Twenty six specimens with different compressive strength and porosity ratios were tested. Strain gaug- ing and the platen-to-platen methods were used to find the stress-strain relationship. Before testing, the cylinders were capped with a sulphur compound on both ends to produce a smooth surface to ensure uni- form transfer of load. Two diametrically opposite 60 mm long strain gauges, were attached to each specimen in its middle third position. A prepared specimen ready to be tested is shown in Figure 1. The signals from the two strain gauges were aver- aged for a more accurate result. The results from the strain gauges were used to find the ascending branch of the stress-straincurve until the peak stress; the strain gauge would not give reliable results after that due to the development of vertical cracks in the sur- face of the specimens. The residual parts of the curve were determined using the platen-to-platen method. An Avery 500kN testing machine was em- ployed for this purpose and the experimental set-up is shown in Figure 2.
JIS Z2101 prescribes three specimen systems: the direc- tion of the tensile load and annual rings must form angles of 0° (T-system), 45° (TR-system), and 90° (R-system), where “T” and “R” indicate the tangential and radial direc- tions, respectively. In this study, the strain was measured over a small area. Thus, the variance of mechanical proper- ties such as the moduli of elasticity in the plane vertical to the loading axis was small for the R-system because this plane was located either in earlywood or latewood. This tendency simplifi es the analysis. However, for T- and TR- systems, the plane vertical to the loading axis had stripes of earlywood and latewood; therefore, the variance of mechan- ical properties in this plane was quite large. Therefore, in this study, the R-system was adopted.
where σ and ε are the stress and strain, ρ is the material density, c v is its specific heat capacity, and the Taylor- Quinney coeﬃcient κ is the fraction of energy transferred to heat. Recent estimates for the actual value of this constant are: κ 0.95 − 1.0, as discussed in  and , for example. It is clear from eq. (9) that a significant temperature rise can be expected for low density materials, having high strength, that are loaded to high strains. High strength aluminum alloys, such as the 7075-T6 alloy, are possible candidates for such a significant thermal softening eﬀect.
longitudinal compression, the mean value of the Young's modulus of the palmyra petiole obtained with the specimens without strain gauge is 1415.74 MPa. The average value of its breaking stress obtained with these test pieces is 29.81±3.17 MPa. That obtained with specimens equipped with strain gauges is 3568.2±651.07 MPa. We noted a difference of 2152.46 MPa between the values of the Young's modulus obtained with the two methods. This difference highlights the errors that are made in the compression tests by measuring the overall deformation of the specimen. Indeed, the over all deformation Consists largely of the crushing of the two heads of the specimen. The use of the gauges allows a measuring only the deformation of the specimen by eliminating the crushing of the heads. The value of tensile breaking stress in the direction of the fibers obtained is σ r =37.77±5.69 MPa. The value of longitudinal
Masaharu et al.  discussed the effects of shape and volume fraction of a second phase on stress states and deformation behavior of two-phase materials with the help of empirical relations. They embedded inhomoge- neous spheroidal (second phase) inclusions in a matrix. Analytical expressions to describe the stress states in elastically and plastically deformed two-phase materials are obtained with the Eshelby method and the Mori- Tanaka concept of the “average stress”. Considering that the second phase is also plastically deformable, the over- all deformation behavior of the two-phase materials is discussed with the results obtained by the evaluation of the stress and strain distributions in the materials. Some of the authors predicted the stressstraincurve with the help of empirical relations [25-28] and phase transfor- mations of titanium alloys [29,30].
The loading-unloading-reloading process could affect the tensile deformation of metals with the combined function of stress relaxation and work hardening, which has been reported in multiple experiments. Nevertheless, the effects of different unloading positions and unloading times have not been investigated. In this study, unloading-reloading tests on three materials (AL6061, HSLA and Q195) were conducted. The stress exhibits a rapid rise momentarily upon reloading and stabilizes afterward while the post stress-straincurve deviates up or down from the monotonic tensile curve. The ductility is enhanced by the unloading-reloading process in general. Different unloading positions and unloading times have different degrees of influence on the stretching of these metals. The effect of loading conditions on a medium manganese steel was further studied. The functions to modify the post stress-strain relationship af- ter unloading-reloading were established.
In order to illustrate the models, exact data generated directly from each model will be used. However, in reality experimental data is not always ‘clean’ and therefore cannot be expected to match perfectly any one model. For example, creep experiments are best done at low stress levels. With too high stress (above the stress threshold), the results are compromised and informative features cannot be extracted. Consider Figure 1.7 showing agarose data at a stress of 2 pascals. The compliance has no curvature - it essentially ramps up continuously until the stress is removed at 10 seconds, at which point it relaxes to a high value and remains there, similar to a pure liquid. The stress was high enough to ‘overpower’ the elasticity in the sample, and any stress at this level or higher will be unable to capture the viscoelastic properties of the material.
Ramezanianpour  listed his results as follows. One of the factors affecting the stress- straincurve of concrete is loading rate. Strains caused by the creep under long-term load changed the stress-straincurve. The maximum compressive stress was reduced by slowing down the loading, but the strain, such as maximum stress, increased, and the final strain of concrete failure increased. In other words, by reducing the loading rate, the compressive concrete became softer. Since, the loading rate was obtained through the approximate role of time variations, it could have effects on the compressive strength of concrete, and its probable cause was the creep phenomenon.
of the yield strength on the normal stress acting on the slip plane, described by a Mohr-Coulomb criterion, 3) the numer- ical resolution of free volume equations developed below at- tempts to quantitatively outline the hydrostatic pressure on the emergence of multiple shear band in constrained geome- tries. The concentration of free volume is the order parameter adopted in the present work to explain the process of the ap- pearance of multiple shear bands.
Methods: Five different absorbable (Vicryl, Maxon, Monocryl, PDS II, Vicryl rapide) and one non-absorbable (Ethibond) suture materials were tested. Measurements were made at five time points during the 56 days of incubation under physiological conditions (37.0 ± 0.02 °C; pH 7.4 ± 0.2). The following variables were recorded: load to failure, strain at maximal load as elongation normalized to original length, stiffness as the ratio of load to displacement on the linear proportion of the stressstraincurve, and hysteresis as area under the curve of the stressstraincurve.
parameter ψ . When the value of ψ decreases, the ESCSR of mismatched welded cylinder closes to the stress-creep straincurve of homogeneous cylinder made of weld metal. While with the value of ψ increasing, the ESCSR gradually closes to the stress-creep straincurve of homogeneous cylinder made of base metal, as seen in Fig. 3 (a) and (b). It is expected according to the above results that as long as the welding seam is wide enough, the weld metal will control the creep resistance of mismatched weld and the influence of base metal can be negligible in this instance.