Issue 41

S. E. Ferreira et alii, Frattura ed Integrità Strutturale, 41 (2017) 129-138; DOI: 10.3221/IGF-ESIS.41.18 131 Since constant amplitude SIF range induce constant FCG rates, the VE widths in such cases can also be assumed fixed and equal to the crack increment per cycle. Any given VE suffers damage in each load cycle, caused by the strain loop range induced by that cycle, which depends on the distance x i between the i -th VE and the fatigue crack tip (Fig. 1). The strain ranges acting in any given VE increase at every load cycle, as the crack tip approaches it. The fracture of the VE adjacent to the crack tip occurs when its accumulated damage reaches a critical value, estimated by the linear damage accumulation rule (or by any other suitable damage accumulation rule) as: (1) where N i ( x i ) is the number of cycles that the i th VE located at a distance x i from the crack tip would last if only that cycle loading would act during its whole life, and n i is the number of cycles that acted at that load, in this case just one. The CDM uses the elastoplastic strain distribution ahead of the crack tip to calculate damage in every VE. However, like in the LE case, EP models for the stress and strain fields inside the plastic zones pz ahead of a crack tip assumed to have a zero tip radius  = 0 , like the HRR field [21-22], are singular at x = 0 as well. To eliminate this undesired and physically inadmissible feature (since no loaded cracks can sustain infinite strains at their tips), the necessarily finite EP stress and strain fields can be estimated by shifting the HRR field origin into the crack by a distance X . This procedure is inspired by Creager and Paris’ idea originally used to estimate stress concentration factors K t from the SIF of geometrically similar cracks [23]. Under constant SIF range conditions the crack advances a distance equal to the VE width in each cycle ( n i  1 ), so the sum in Eq. (1) can be approximated by an integral along, say, the reverse or cyclic plastic zone ( pz r ), neglecting in a first approximation fatigue damage induced outside it: (2) The HRR field origin shift can be estimated e.g. assuming X   /2 , as Creager and Paris did, where  is the (finite) crack tip radius under the peak load K max . To calculate the cyclic plastic strain range  p ahead of the crack tip, the modification proposed by Schwalbe [24] can be used as in [19]: (3) where S Yc is the cyclic yield strength of the material, E is its Young's modulus, and h c is its Ramberg-Osgood strain- hardening exponent. Since the elastic strain amplitude inside the cyclic plastic zone is neglected in Eq. (3), its associated fatigue life N(x + X) can be estimated from the plastic part of Coffin-Manson’s equation by N , x+X . = , 1 - 2 .,,, ∆ , ε - p . (x+X) - 2 , ε - c ...-, 1 - c .. (4) where c and  c are Coffin-Manson's plastic exponent and coefficient, respectively. Hence, estimating the X displacement of the modified HRR field in the same way as Creager and Paris did with the LE Williams field, i.e. assuming it as   CTOD/2 , where CTOD is Schwalbe’s estimate for the Crack Tip Opening Displacement induced by K max , then (5) The FCG rate induced by constant SIF ranges, i.e. by fixed {  K, R } conditions, can then be estimated by Eq. (2)-(5). Finally, these resulting da/dN values can be used to calculate the constant C in a modified McEvily’s rule that simulates all 3 phases of typical da/dN  K curves, considering the FCG threshold  K th and the toughness K c limits in FCG rates: (6)   ( ) 1 i i i i n N x     0 pz r da dx dN N x X            1 1 2 h c p Yc r x X S E pz x X                2 1 2 1 2 4 2 1 max Yc c K CTOD X E S h                 2 th c c max da dN C K K K K K           