numero25

J.T.P de Castro et alii, Frattura ed Integrità Strutturale, 25 (2013) 79-86; DOI: 10.3221/IGF-ESIS.25.12 84 cannot be answered by SN or  N procedures alone. Such problem can be avoided by adding short crack concepts to their “infinite” life design criteria which, in its simplest version, may be given by [5]:     1 2 ( ) 1 R F R K a g a w a a               (17) where   1 /2 ( ) 1 R R R K a K a a            , and     2 1 R R R a K S      Since the fatigue limit  S R   S L (R) already reflects the effect of the microstructural defects inherent to the material, equation (17) complements it by describing the tolerances to small cracks that may pass unnoticed in actual service conditions. The practical usefulness of this sensible criterion is well illustrated by a practical example, as follows. Due to a rare manufacturing problem, a batch of an important component left the factory with tiny surface cracks (only detectable by a microscope), causing some unexpected and embarrassing failures. Hence, it became necessary to properly quantify the actual effect of such tiny cracks in that component fatigue strength. Assume its rectangular cross-section has 2 by 3.4mm ; its (uncracked) fatigue limit is S L = 246MPa ; and it is made from steel with S U = 990MPa . This measured fatigue limit is about ¼ of the steel S U /4 , whereas it would be traditionally estimated by S L  S U /2  495MPa . This difference may be due to a surface finish factor k sf  0.5 , a value between those proposed by Juvinall for S U  1GPa steels with cold-drawn ( k sf  0.45 ) and machined surfaces ( k sf  0.7 ) [17]. The surface finish should not affect cracks, but as this difference could be due to other factors too, like tensile residual stresses, the only safe option is to use S L = 246MPa to evaluate the tiny crack effects. Therefore, by Goodman S L (R) = S R = S L S U (1 – R)/[S U (1 – R) + S L (1 + R)] for R  1 (or  m  0 ), for example. The FCG threshold  K R is also needed to model short crack effects, but if data is not available, as in this case, it must be estimated e.g. by  K R (R  0.17)   K 0  6MPa  m and  K R (R > 0.17)  7  (1 - 0.85R) [18]. This practice increases the predictions uncertainty, but it is the only option available. Besides, it tends to be conservative. Moreover, it assumes that  K R (R < 0)  K 0 , a safe estimate too (except if the load history contained severe compressive underloads which might accelerate the crack, not the case here.) Using the SIF of an edge cracked strip of width w loaded in mode I, then the tolerable stress ranges under pulsating loads shown in Fig. 4 are estimated within a fatigue safety factor  F as:   0 0 1 2 3 0 2 [ 0.752 2.02 0.37(1 sin ) ]sec tan 1 2 2 2 F K a a a a a w a w a w w w a                            (18) Figure 4 : Larger stress ranges tolerable under several R –ratios by the analyzed component considering it contains an edge crack with size a , for w  3 .4mm ,   1 .12 ,  K 0  6MPa  m , a 0  59  m ,   6 , and  F  1.6 . N OTCH S ENSITIVITY IN E NVIRONMENTALLY A SSISTED C RACKING P ROBLEMS he behavior of materials on aggressive environments is an important problem for many industries, because the costs and specially the delivery times for special EAC-resistant alloys keep increasing. Major problems occur e.g. in the oil industry, where oil and gas fields can contain considerably amounts of H 2 S, and in the aeronautical T