Failure in notched tension bars due to high-temperature creep: Interaction between nucleation controlled cavity growth and continuum cavity growth

Abstract

For axi-symmetrically notched tension bars [Dyson, B.F., Loveday, M.S., 1981, Creep Fracture in Nimonic 80A under Tri-axial Tensile Stressing, In: Ponter A.R.S., Hayhurst, D.R. (Eds.), Creep in Structures, Berlin, show two types of damage propagation are shown: for low stress, failure propagates from the outside notch surface to the centre-line; and for high stress, failure propagates from the centre-line to the outside notch surface. The objectives of the paper are to: identify the physics of the processes controlling global failure modes; and, describe the global behaviour using physics-based constitutive equations.Two sets of constitutive equations are used to model the softening which takes place in tertiary creep of Nimonic 80A at 750°C. Softening by multiplication of mobile dislocations is firstly combined, for low stress, with softening due to nucleation controlled creep constrained cavity growth; and secondly combined, for high stress, with softening due to continuum void growth. The Continuum Damage Mechanics, CDM, Finite Element Solver DAMAGE XX has been used to study notch creep fracture. Low stress notch behaviour is accurately predicted provided that the constitutive equations take account of the effect of stress level on creep ductility. High stress notch behaviour is accurately predicted from a normalized inverse cavity spacing d/2' = 6, and an initial normalized cavity radius r hi /' = 3.16 · 10 À3 , where 2' is the cavity spacing, and d is the grain size; however, the constants in the strain rate equation required recalibration against high stress notch data. A void nucleation mechanism is postulated for high stress behaviour which involves decohesion where slip bands intersect second phase grain boundary particles. Both equation sets accurately predict experimentally observed global failure modes. Nomenclature Stress r 1Uni-axial stress r ij Stress tensor r e (=3r ij r ij /2) 1/2 Effective stress r 0Normalizing stress S ij Stress deviation tensor J 1 Normalized first stress invariant R ij ð¼ r ij =r 0 Þ Normalized stress tensor S ij ð¼S ij =r 0 Þ Normalized stress deviation tensor R e (=r e /r 0 ) Normalized effective stress Strain _ e 1 Uni-axial creep strain rate _ e ij Creep strain rate tensor _ e o ð¼ A ¼ P Þ Steady-state uni-axial strain rate e o (= r 0 /E) Normalizing strain k ij (= e ij /e o ) Normalized strain tensor _ e e ð¼ 2_ e ij _ e ij =3Þ 1=2 Effective creep strain rate tensor

Application of small punch creep testing to a thermally sprayed CoNiCrAlY bond coat

“…5(a) supports this view. It is important to recall that at 750 C the -NiAl phase has little ductility [28] and brittle second phase particles have previously been noted as preferential sites of void nucleation during creep of, for example, duplex stainless steel [29], Nimonic 80A [30] and SiC reinforced aluminium [31].…”

Thermal-mechanical fatigue simulation of a P91 steel in a temperature range of 400–600°C

“…Similar studies have been carried out, for a P91 steel, in order to develop a creep constitutive model with a damage capability [3]. The development of these creep constitutive models have contributed to the gaining of a better understanding of the material behaviour in such applications as welding process modelling [4] and failure prediction in multiaxial components [5].…”

Characterization of viscoplasticity behaviour of P91 and P92 power plant steels

“…Similar studies have been carried out, for a P91 steel, in order to develop a creep constitutive model with a damage capability [6]. The development of these creep constitutive models have contributed to the gaining of better understanding of the material behaviour in such applications as welding process modelling [7] and failure prediction in multiaxial components [8]. The accuracy of the FE modelling works depends on the material constitutive equations.…”