The influence of antimony-additions on the creep behaviour of a 25wt% Cr-20wt% Ni austenitic stainless steel

The effect of antimony on the creep behaviour (dislocation creep) of a 25 wt% Cr-20 wt% Ni stainless steel with ~ 0.005 wt% C was studied with a view to assessing the segregation effect. The antimony content of the steel was varied up to 4000 ppm. The test temperature range was 1153 to 1193 K, the stress range, 9.8 to 49.0 MPa, and the grain-size range, 40 to 600m. The steady state creep rate, , decreases with increasing antimony content, especially in the range of intermediate grain sizes (100 to 300m). Stress drop tests were performed in the secondary creep stages and the results indicate that antimony causes dislocations in the substructure to be immobile, probably by segregating to them, reducing the driving stress for creep.Nomenclature _a Creep stress in a constant load creep test without stress-drop - _A Initial applied stress in stress-drop tests - Stress decrement - ( _A-) Applied stress after a stress decrement, - t _i Incubation time after stress drop (by the positive creep) - _C Strain-arrest stress - _i Internal stress - _{s s}-component (= _i- _c) - Steady state creep rate (average value) in a constant load creep test - Strain rate at time,t, in a constant load creep test - New steady state creep rate (average value) after stress drop from _A to ( _A-) - Strain rate at time,t, after stress drop.     

Methodology for the determination of the interfacial properties of brittle matrix composites

The interfacial properties of a glass-ceramic matrix composite (SiC/CAS) were determined from single-fibre push-out tests using the interfacial test system. The coefficient of friction, , the residual clamping stress, _c, and fibre axial residual stress, _z , were extracted by fitting the experimental stress versus fibre-end displacement curves using the models of Hsueh, and Kerans and Parthasarathy. Using Hsueh's model, the intrinsic interfacial frictional stress (=_c) was found to be 11.1±3.2 MPa, whereas by using Kerans-Parthasarathy's model it was found to be 8.2±1.5 MPa. Comparisons between these models are included, together with a discussion of data analysis techniques.Nomenclature _z Axial fibre residual stress (Pa) - ^* Effective clamping stress (Pa) - _c Residual clamping stress (Pa) - _p Poisson's effect-induced clamping stress (Pa) - _d ⁰ Debond stress in the absence of residual stresses (Pa) - _d Experimental debond stress (Pa) - Compressive applied stress (Pa) - Interfacial shear stress (Pa) - u Fibre-end displacement (m) - h Debond length (m) - r Fibre radius (m) - E _f Fibre Young's modulus (Pa) - E _m Matrix Young's modulus (Pa) - v _f Fibre Poisson's ratio (dimensionless) - v _m Matrix Poisson's ratio (dimensionless) - f Fibre volume fraction (dimensionless) - k Parameter (dimensionless) - D Parameter (dimensionless) - Interfacial coefficient of friction (dimensionless) - G _i Interface toughness (J m^–2) - C _m Load-train compliance (m N^–1)     

The dielectric behaviour of single-crystal MgO,Fe/MgO and Cr/MgO

The dielectric constants and loss factors,, for pure single-crystal MgO and for Fe-and Cr-doped crystals have been measured at frequencies, , from 500 Hz to 500 kHz at room temperature. For pure MgO at 1 kHz the values of and the loss tangent, tan , (9.62 and 2.16×10^–3, respectively) agree well with the data of Von Hippel; the conductivity, , varies as ⁿ withn=0.98±0.02. In Fe-doped crystals increases with Fe-concentration (at any given frequency); for a crystal doped with 12800 ppm Fe, was about four times the value for pure MgO. At all concentrations the variation of log with log was linear andn=0.98±0.02. A decrease in with increasing Fe-concentration was also observed. A similar, although less pronounced, behaviour was found in Cr-doped crystals. The effects are discussed in terms of hopping mechanisms.     

Application of a new iterative method for elastic-plastic stress analysis

A new iterative method for elastic-plastic stress analysis based on a new approximation of the constitutive equations is proposed and compared with standard methods on the accuracy and the computational time in a test problem. The proposed method appears to be better than the conventional methods on the accuracy and comparable with others on the computational time. Also the present method is applied to a crack problem and the results are compared with experimental ones. The agreement of both results are satisfactory.List of symbols u = (u ₁, u ₂) displacements u ^(H) = u ⁽ⁿ⁺¹⁾ - u ⁽ⁿ⁾ u _k ⁽ⁿ⁾ = u _(k ^{(n + 1)} - u ⁽ⁿ⁾ (n, k = 0, 1, 2, ...) - = ₁₁, ₂₂, ₁₂) stresses - = (₁₁, ₂₂, ₁₂) strains - = (₁₁, ₂₂, ₁₂) center of yield surface - D elastic coeffficient matrix, C = D ^–1 - von Mises yield function. The initial yielding is given by f() = _Y - f {f/} - ^* transposed f - H hardening parameter (assumed to be a positive constant for kinematic hardening problems) - time derivative of - [K] total elastic stiffness matrix - T traction vector - = [B] relation between nodal displacements and strains     

Distribution characteristics of mechanical properties and correlation between the respective properties on S35C carbon steel

Mechanical properties of tensile strength, , upper yield stress, _SU, lower yield stress, _SL elongation, , area reduction, , Vickers hardness, H _v, and impact absorbed energy, E, were examined using 50 specimens of S35C carbon steel, which were machined from two bars supplied from the same charged and heat-treated material. Distribution characteristics of these properties are discussed, and the correlation between each pair of them is investigated from a statistical viewpoint. The main conclusions obtained are summarized as follows; distribution characteristics of _B, _SL, , , H _v and E are well approximated by a normal distribution, but those of asu are not approximated as well by this type of distribution. In the latter case, a Weibull distribution is preferable to represent the distribution pattern. No significant correlation was observed between each pair of the above mechanical properties. Consequently, individual properties have the inherent distribution characteristics independent of the other properties.     

Residual thermal stresses in the pull-out specimen: A finite element calculation

The residual thermal stress field in the pull-out specimen is calculated in the case of a high properties thermoset system (carbon-bismaleimide). The calculation is performed within the framework of the linear theory of elasticity by means of a finite element method. The specimen is modelled as a three-phase composite (holder-fibre-matrix). The meniscus which forms at the fibre entry is taken into account in order to provide a realistic stress concentration. The latter is far higher than the matrix strength. Evidence that fibre debonding propagates from the fibre end during cooling is then produced.Nomenclature T thermal load - L _e embedded length - r _f fibre radius - _c curvature radius of the meniscus (fibre entry) - r _c radial dimension of the finite element mesh - E _m,E _h matrix and holder moduli - E _A,E _T fibre axial and transverse moduli - _m, _h matrix and holder thermal expansion coefficients - _A, _T fibre axial and transverse thermal expansion coefficients - _rr, , _zz, _rz non-zero components of the residual stress field - _rr ⁱ , ^im , _zz ^im , _rz ⁱ stresses at the interface in the matrix (r=r _f + ) - _rr ⁱ , ^if , _zz ^if , _rz ⁱ stresses at the interface in the fibre (r=r _f–) - _p1 maximum principal stress - _zz ^f mean axial stress over the fibre section - _rupt ^m matrix strength - u _r ,u _z non-zero components of the displacement field     

The tensile strength and ductility of continuous fibre composites

The plastic instability approach has been applied to the tensile behaviour of a continuous fibre composite. It is shown that the combination of two components with different strengths and degrees of work-hardening produces a new material with a new degree of work-hardening, which may be determined by the present analysis. Expressions for the elongation at rupture and the strength of a composite have been obtained and the results of the calculation are compared with some experimental data.List of symbols V _f volume fraction of fibres in composite - , , true strain of fibre, matrix and composite - s true stress - , , nominal stress on fibre, matrix and composite - _*, _*, _* critical stress of fibre, matrix and composite (ultimate tensile strength) - _*, _* critical strain of separate fibre and matrix - _* critical strain of composite - Q external load - A cross-sectional area - A ₀ initial value of area     

Fatigue of metallic glasses

The fatigue behaviour of Ni₄₉Fe₂₉P₁₄B₆Si₂, Ni₄₈Fe₂₉P₁₄B₆Al₃ and Pd_77.5Cu₆Si_16.5 metallic glasses is examined. In the finite lifetime regime the relationship between stress amplitude ( _a), fracture stress ( _f), mean stress ( _m) and cycles to failure (N _f) is _a=A( _f–_m) (2N _f) ^b , whereA andb are 16.9 and –0.40 respectively for reduced gauge section Ni₄₉ strips (for _m 140 kg mm^–2) and 27.0 and –0.44 for Pd base wires. These results are unusual in thatA 1. Consequently, a sharp discontinuity exists near _a( _f– _m) –1. In a simple tensile test failure occurs at _f(=_y) and 2N_f=1; for peak stresses only a percent or so less than _f the sample will withstand hundreds of cycles of stress. For uniform cross-section glassy metal filaments, a fatigue limit is observed at stress ratios ( _a/ _f) in the vicinity of 0.07 to 0.15. The fatigue limit for reduced section specimens is a factor of 2 higher. Fatigue failure of the Ni-Fe strips may occur under partially or fully plane stress or plane strain conditions, depending on sample thickness and stress. Final failure of the Pd_77.5Cu₆Si_16.5 wires always occurs by general yielding of the remaining section.     

A modified Weibull treatment for the analysis of strength-test data from non-identical brittle specimens

Powder compacts (e.g., pharmaceutical tablets) manufactured on commerically available machines are not strictly identical but show inevitable variability in their weights, thicknesses and compaction pressures. Consequently, the variability in fracture-stress data obtained from such brittle specimens is greater than that due to the inherent strength variability of the material itself. A modified Weibull analysis has been developed so that a more accurate estimate of the inherent variability of the mechanical strength of the material can be derived from test data obtained from commercially produced compacts; its application is illustrated.Nomenclature D diameter - f() relative frequency of occurrence of specimens with density and volume - F minimization function - i ascending rank number of a fracture stress - m Weibull modulus - N _tot number of specimens in a batch - N() number of specimens with densities in the range to + d and volumes in the range to + d - P _f failure probability - p _u upper punch compaction pressure - t thickness - volume - w weight - W _f fracture load - density - _f fracture stress - ¯ _f mean fracture stress of a batch - ¯ _f() mean fracture stress of specimens with density and volume - ₀ scale parameter or normalizing factor - _u location parameter or threshold stress     

The longitudinal electrical conductivity of metal matrix composites at cryogenic temperature in the presence of a longitudinal magnetic field

In this paper we use the results of Part I to derive two integral expressions for the electrical conductivity of metal matrix composite materials when a magnetic field, B, is added to a small electric field also parallel to fibres. One expression applies to strong magnetic fields meaningR ₀/a < 1, whereR ₀ =m ^* v _F/eB when the Fermi velocity is perpendicular to the magnetic field. WhenB , the integral expression reduces to the well known conductivity value = ₀(1 -V _f), where ₀ is the bulk matrix conductivity andV _f is the fibre volume fraction. For weak magnetic fields,R ₀/a > 1, then the electrical conductivity is expressed by the sum of two integrals. When B 0, the electrical conductivity reduces to the integral expression obtained in our earlier results when there is only a longitudinal electric field. In this paper we correct an incorrect derivation of the composite conductivity in the absence of a magnetic field published earlier [J. Mater. Sci.21 (1986) 2409].     

Experimental study of the boundary between single-phase convection and boiling in underheated cryogenic liquids

The effect of pressure and underheating on the position of the boundary between heat-transfer regimes in liquid helium and hydrogen is investigated.Notation q heat flux - p pressure - =T_s–T underheating - T_s saturation temperature - T temperature of liquid - T=T_wa – T T_s=T_wa – T_s - T_wa temperature of heat-emitting surface - A,a, B, b, C constants - m, n indices - Nu Nusselt number - Ra Rayleigh number - thermal conductivity - coefficient of cubical expansion - kinematic viscosity - g acceleration - standard deviation Indices 01 conditions of convection-boiling transition - 02 conditions of boiling-convection transition Translated from Inzhenerno-Fizicheskii Zhurnal, Vol. 42, No. 1, pp. 5–11, January, 1982.     

Effect of magnesium content on the stress exponent and effective stress in the steady state creep of Al-Mg alloys

The stress exponent of steady state creep,n, and the internal ( _i) and effective stresses ( _e) have been determined using the strain transient dip test for a series of polycrystalline Al-Mg alloys creep tested at 300° C and compared with previously published data. The internal or dislocation back stress, _i, varied with applied stress,, but was insensitive to magnesium content of the alloy, being represented by the empirical equation _i=1.084 ^1.802. Such an applied stress dependence of _i can be explained by using an equation for _i of the form _i (dislocation density)^1/2 and published values for the stress dependence of dislocation density. Values of the friction stress, _f, derived using the equation _e/=(1–c) (1– _f/), indicate that _f is not dependent on the magnesium content. A constant value of _f can best be rationalized by postulating that the creep dislocation structure is relatively insensitive to the magnesium content of the alloy.On leave from Engineering Materials Department, University of Windsor, Windsor, Ontario N9B 3P4, Canada.     

Pyramidal yield criteria for epoxides

The tensile, compressive and shear yield strengths of two epoxides were measured under superposed hydrostatic pressure extending to 300 MN m^–2. For both materials, the ratio of the moduli of the tensile, _T, to compressive, _C, yield stress at atmospheric pressure was approximately 34, as has been reported previously for a number of thermoplastics. The ₂= ₃ envelope in stress space was plotted according to these two-parameter ( _C and _T) yield criteria: conical, paraboloidal and pyramidal; the best correlation was with the last. The experimental tensile and compressive data for tests under pressure, however, fit slightly better two straight lines which are consistent with a three-parameter single hexagonal pyramidal yield surface. For plane stress and shear under pressure yield envelopes of these surfaces, the correlation with experimental data is again best for the pyramidal criteria, except for biaxial or triaxial tension when these resins are brittle. The third independent parameter employed in the pyramidal criterion was the equi-biaxial compressive yield stress, determined by tensile experiments under appropriate superposed hydrostatic pressure; alternatively plane strain compressive yield stress, _PC, may be used.     

Charged defects-controlled conductivity in Ge-ln-Se glasses

The variation of the d.c. electrical conductivity, , with composition and temperature was investigated for glasses of the Ge-In-Se system. The results indicate a decrease in the activation energy for electrical conductivity, E, and an increase in on introduction of indium into Ge-Se glasses. The changes in E and with composition (selenium content in the glasses) are identical for the Ge_x In₅ Se_95–x and Ge_x In₈Se_92–x families. The results have been traced to the conduction controlled by charged defects in these chalcogenide glasses. The changes in E and have been explained by a shift in the Fermi level, being brought by the introduction of indium.     

Deformation of a carbon-epoxy composite under hydrostatic pressure

This paper describes the behaviour of a carbon-fibre reinforced epoxy composite when deformed in compression under high hydrostatic confining pressures. The composite consisted of 36% by volume of continuous fibres of Modmur Type II embedded in Epikote 828 epoxy resin. When deformed under pressures of less than 100 MPa the composite failed by longitudinal splitting, but splitting was suppressed at higher pressures (up to 500 MPa) and failure was by kinking. The failure strength of the composite increased rapidly with increasing confining pressure, though the elastic modulus remained constant. This suggests that the pressure effects were introduced by fracture processes. Microscopical examination of the kinked structures showed that the carbon fibres in the kink bands were broken into many fairly uniform short lengths. A model for kinking in the composite is suggested which involves the buckling and fracture of the carbon fibres.List of symbols d diameter of fibre - E _f elastic modulus of fibre - E _m elastic modulus of epoxy - G _m shear modulus of epoxy - k radius of gyration of fibre section - l length of buckle in fibre - P confining pressure (= ₂ = ₃) - R radius of bent fibre - V _f volume fraction of fibres in composite - _t, _c bending strains in fibres - angle between the plane of fracture and ₁ - ₁ principal stress - ₃ confining pressure - _c strength of composite - _f strength of fibre in buckling mode - _n normal stress on a fracture plane - _m strength of epoxy matrix - shear stress - tangent slope of Mohr envelope - slope of pressure versus strength curves in Figs. 3 and 4.     

Surface tension of liquid4He nearT λ

The surface tension of liquid⁴He in the neighborhood of the transition temperature was measured with very high precision by the surface-wave resonance method. The derivative of the surface tension with respect to temperature decreases by 21.1±0.8 mdyn/cm K as the temperature is increased from T – 20 mK to T + 20 mK. The regular part of the surface tension near T is considered to consist of two parts, a ripplon contribution _R and a bulk contribution _B. In the present analysis, the singular part exists only below T, and is proportional to (T – T)^1.3±0.1. This behavior seems to be related to the Bose-Einstein condensate or the superfluid density.     

Theory of the Surface Tension Maximum of Liquid 3He

Recently Matsumoto et al. performed very precise measurements of the surface tension of liquid ³He, (T), at low temperatures and found that (T) exhibits a small maximum at about 100 mK. Existing theories are unable to explain this anomaly. On the basis of a local approximation for the entropy in which the Fermi liquid effect is included, we can evaluate the variation of (T), (T)=(T)–(0), as a function of T and of the number density (or the interaction strength). It is found that (T) consists of two terms; a T² term and T⁴ ln T term. We predict that, for the density of real liquid ³He, exhibits a tiny minimum and a small but relatively larger maximum. This prediction explains qualitatively and quantitatively all salient features of the observed (T).     

Quantitative representation of stress-corrosion cracking of metals

A method is presented of quantitatively estimating the resistance of metals to stress corrosion on the basis of the reduction in the long-time strength _SC in a given medium. The calculation of ^sc when the reduction in the long-time strength is associated not only with stress corrosion but also with general corrosion and hydrogen embrittlement is discussed.     

Electronic and ionic conduction in some simple lithium salts

The electrical conductivity () and thermoelectric power (S) of solidified melt samples of Li₂MoO₄, Li₂WO₄ and Li₂SO₄ are presented in the temperature range 415 K to melting point of each compound. The ratio of ionic to electronic contribution to has been obtained with the help of a time-dependence study of d.c. electrical conductivity. It has been shown that in Li₂MoO₄ electronic contribution to remains high up to its melting point (about 8% just below the melting point) and it shows no superionic phase. However, in Li₂WO₄ and Li₂SO₄ a superionic phase is obtained in which the ionic contribution to is more than 99.99%. However, in normal ionic (or) phase it is small and decreases with decreasing temperature. Separate temperature variations of ionic ( _i) and electronic ( _e) conductivities are presented and the conduction mechanisms are discussed. It is shown that ionic conduction in the-phase is dominated by Schottky-type defects.     

Stress-strain relations for composites with different stiffnesses in tension and compression

Some fibrous composites exhibit different mechanical properties in tension as compared to those in compression. Proper material modelling is required to enable correct behavioural prediction of components made of such composites. Though the models due to Ambartsumyan, Bert and Jones are often used, there are still many unanswered questions pertaining to material modelling. In Bert's model the strain transverse to the fibres and the shear strain are discontinuous when the fibre strain is zero. In Jones' model, cross compliances in tension-compression zones are assumed to depend on the magnitudes of principal stresses.Here we have proposed a new model for bimodulus orthotropic materials with zonewise symmetric linear constitutive laws valid for biaxial fields and dependent on the signs of both normal stresses and strains, referred to material axes. This model maintains strain continuity in the entire biaxial field and has ten independent elastic constants compared to eight each in Bert's and Jones' model. Applicability of the present model is illustrated by considering limited experimental data available on Aramid cord-rubber and ATJ graphite.List of symbols L, T Axes parallel and transverse to fibres - S _SS ^± Shear compliance for ± _LT - S _Lt ,S _Lc Compliances inL-direction for uniaxial tension and compression applied inL-direction - S _Tt ,S _Tc Compliances inT-direction for uniaxial tension and compression applied inT-direction - S _TL ^(t) Cross compliance corresponding to uniaxial load inL- orT-direction maintaining _L > 0 - S _TL ^(c) Cross compliance corresponding to uniaxial load inL- orT-direction maintaining _L < 0 - S _LL ⁺,S _TT ⁺,S _TL ⁺ Compliances when _L > 0 and _T > 0 - S _LL ^–,S _TT ^–,S _TL ^– Compliances when _L < 0 and _T < 0 - S _Lt ^– Compliance inL-direction when _L < 0 under the combined action of tensile _L and _T - S _Tt ^– Compliance inT-direction when _T < 0 under the combined action of tensile _L and _T - S _Lc ⁺ Compliance inL-direction when _L > 0 under the combined action of compressive _L and _T - S _Tc Compliance inT-direction when _T > 0 under the combined action of compressive _L and _T - ₁ , ₂ ; ₃, ₄ Parameters defined in Eqs. (12); (12) - _L , _T , _LT Strains with respect toL-,T-axes - _L , _T , _LT Stresses with respect toL-,T-axes - Orientation of fibres with respect toL-axis - S ₁₁,S ₂₁ Compliances referred to axes oriented at an angle toL-,T-axes