PSI - Issue 47

Robert Eriksson et al. / Procedia Structural Integrity 47 (2023) 227–237 R. Eriksson, A. Azeez / Structural Integrity Procedia 00 (2023) 000–000

228

2

a)

K

b)

K

Fracture toughness

Fracture toughness

K 1

K 1 = K 2 K f

K f

K 2

T 2

T 1

T 2

T 1

T

T

Fig. 1. Schematic illustrations of a (a) LUCF cycle and (b) LCF cycle.

All mechanisms have in common that they are the result of plastic deformation at the crack tip. In general, most researchers seems to favor residual stresses as the main contributing mechanism since stress relief heat treatments after WPS load have been shown to reduce the WPS e ff ect Reed and Knott (1996); Blumenauer and Krempe (2001); Chen et al. (2002). There are many possible temperature–load histories, but two cases have become particularly common when char acterizing WPS: the load–cool–fracture (LCF) cycle and the load–unload–cool–fracture (LUCF) cycle, both are illus trated in Fig. 1. The LCF cycle is generally considered to give a larger increase in K f than LUCF Reed and Knott (1996). The following terminology is used to describe the temperature–load cycle • The temperature where the WPS load is applied is denoted T 1 and the applied load is denoted either P 1 (force) or K 1 (stress intensity). • For LUCF, the value of the lowest load at T 1 is denoted P 2 (force) or K 2 (stress intensity). For LCF, P 2 = P 1 , K 2 = K 1 . • At low temperature, T 2 , the load at fracture is denoted P f with the corresponding fracture resistance K f (stress intensity). The WPS load, P 2 , needed to give rise to a WPS e ff ect have been variously described as K 1 > K Ic Reed and Knott (1996) or P 1 P GY ≥ 0 . 5where P GY is the load causing general yield Wang et al. (2002). Models describing the WPS e ff ects include “global” approaches such as that of Wallin (2003, 2004) which do not require the stresses in front of the crack to be known and “local” approaches such as the Beremin Beremin (1983) and modified Beremin models Lefevre et al. (2002); Kordisch et al. (2000) which are based on a weakest link assumption and require that stresses are known. The crack tip plastic zone created during WPS load forms, what Chell (1986) refers to as, a “residual zone” which will not deform further when the crack is loaded to fracture at low temperature. As pointed out by several authors Reed and Knott (1996); Smith et al. (2010), plasticity is necessary for cleavage to occur, meaning that failure will occur at the onset of plasticity outside of the residual zone. In the present work, a physically based analytical model was developed based on fracture mechanics. As most suggested mechanisms of the WPS e ff ect could be, at least, indirectly related to plastic zone size, a strip-yield model was taken as a basis for the developed model. WPS tests were performed in an electromechanical tensile test rig, Alwetron TCT 100, and a 3-zone furnace. Tem perature was controlled by three thermocouples; two mounted on each grip and one mounted on the specimen. Load line displacement was measured using a high temperature extensometer, Epsilon Tech. Corp. The studied material was a wrought, creep resistant 9 % Cr steel for turbine applications with a DBTT around ∼ 50 ◦ C. The tests were performed using standard compact tension specimens with thickness 25 mm (CT25); side grooves where used. The specimen was heated during a small preload of 0.5 kN to avoid going into compression during heating. Once the desired temperature was reached, the specimen was left to dwell for 30 min to reach a homogeneous temperature prior to loading. For the LCF cycle, the furnace was turned o ff immediately after reaching maximum load (except for specimen no. 8, see Table 1, which was left to dwell 60 min prior to cooling). For the LUCF cycle, 2. Experiments