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Structural Integrity

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Procedía Structural Integrity 2 (2016) 2367-2374

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21st European Conference on Fracture, ECF21, 20-24 June 2016, Catania, Italy

Thermal load-induced notch stress intensity factors derived from

averaged strain energy density

P. Ferro*, F. Berto, T. Borsato

University ofPadova, Department of Engineering and Management, Stradella S. Nicola, 31-36100 Vicenza, Italy

Abstract

Under the hypothesis of steady-state heat transfer and plane-strain conditions, the intensity of the stress distributions ahead of sharp V-notch tips can be expressed in terms of thermal notch stress intensity factors (thermal NSIFs) which can be used for fatigue strength assessments of notched components. The calculation of thermal NSIFs requires both an uncoupled thermal-mechanical numerical analysis and a very refined mesh. For these reasons, the numerical simulation becomes considerably expensive and time-consuming above all if large 2D or 3D models have to be solved. Refined meshes are not necessary when the aim of the finite element analysis is to determine the mean value of the local strain energy density on a control volume surrounding the points of stress singularity. On the other hand, the NSIFs value can be directly determined by the strain energy density. In this work, the method for rapid calculations of NSIFs based on averaged strain energy density, recently published in literature, is extended to thermal problems.

Copyright © 2016) The Authors. Published by Elsevier B.V. This is an openaccess article under the CC BY-NC-ND license

(http://creativecommons.org/licenses/by-nc-nd/4.0/).

Peer-review under responsibility of the Scientific Committee of ECF21.

Keywords: Thermal Stress; Strain Energy Density; Thermal Notch Stress Intensity Factors; Finite Element Method; Thermal Fatigue

1. Intro duction

In presence of components with geometric discontinuities like sharp V-notches, the stress distribution near such regions is singular. Under linear-elastic hypothesis and plain strain or plain stress conditions, the solution of this problem was first obtained by Williams (1952) and later extended to the elastic-plastic regime by Hutchinson (1968), Rice and Rosengreen (19)685) (HRR). It was found that the singularity grade of the stress asymptotic distribution

* Corresponding author. Tel.: +39 0444 998769; fax: +39 0444 998888. E-mail address: ferro@gest.unipd.it

CrossMark

Copyright © 2016 The Authors. Published by Elsevier B.V. This is an open access article under the CC BY-NC-ND license

(http://creativecommons.org/licenses/by-nc-nd/4.0/).

Peer review under responsibility of the Scientific Committee of ECF21.

10.1016/j.prostr.2016.06.296

depends only on the V-notch opening angle, while its intensity is given by the notch stress intensity factor (NSIF). The definition of this last parameter allowed to formulate a local approach useful to predict the static and fatigue resistance of V-notched components. For example, when mode II contribution is not singular, the mode I NSIF was used to summarize the high-cycle fatigue strength of welded joints having very different geometries (Lazzarin and Tovo (1998), Atzori et al. (1999), Lazzarin and Livieri (2001), Atzori and Meneghetti (2001)). In those works, the local parameter Ki was correlated only with the geometry and the external loads without including the influence of residual stresses.

In order to take into account the effect of thermal loads on fatigue strength of notched components, the local approach based on NSIFs was extended to thermal problems. First works about this topic date back to 1979 and extend until 1998 (Stern (1979), Babuska and Miller (1984), Szabo ' and Yosibash (1996), Yosibash (1997), Yosibash (1998)). Many years later, this problem was analytically solved by P. Ferro at al. (2006) by using the 'stress function approach'. The thermal load-induced NSIFs were found to be natural extension of the conventional force-induced NSIFs. In that work, the authors introduced also the concept of Residual-NSIF which is the parameter used to quantify the intensity of residual singular stress field arising near the weld toe and induced by the solidification of the fusion zone (FZ). The residual stress field near the weld toe was then extensively studied in further works that highlighted how the intensity and the sign of the induced residual stress depend on boundaries conditions and phase transformations (Ferro and Petrone (2009), Ferro (2012)). Finally, a local model was developed to predict the fatigue strength of welded joints which takes into account the effects of the residual stress (Ferro (2014)). Unfortunately, the local approach based on NSIFs suffers from two drawbacks. Fatigue strength of notched components with different V-opening angles can not be compared each other because of the different singularities degree which appears in the NSIF units; furthermore, a very fine mesh is required near the notch tip to capture the real NSIF value. This last drawback is particularly heavy in 3D numerical modelling and even more in the 3D simulation of welding process since a non-linear-transient thermo-metallurgical and mechanical analysis is required. In order to overcome these problems, a local energy based approach was developed by Lazzarin and Zambardi (2001). In this formulation it is hypothesized that the static and fatigue behaviour of components with sharp V-shaped notches depends on the strain energy density (SED) averaged over a circular sector (structural volume) of radius Rc near the notch tip. The local strain energy density is directly linked to the relevant NSIF for modes I and II while Rc is treated as a material property. This criterion was verified over a great number of experimental tests collected in different papers by Lazzarin and co-workers (Livieri and Lazzarin (2005), Berto and Lazzarin (2009)). The great advantage of the SED approach is that the units of the strain energy density do not change with the opening angle and its value is related to the nodal displacements. This last characteristic makes the strain energy density almost mesh independent. In other words, a coarse mesh is sufficient to capture the SED value averaged over the control volume, which makes this approach easy to use also when 3D and large components have to be modelled (Lazzarin et al. (2010)). Another interesting consequence of the above mentioned SED characteristic is that the NSIF value can be indirectly derived from the numerically calculated value of the strain energy density averaged over a control volume of radius R by means of a coarse mesh (Lazzarin et al. (2010)). Starting from previous results, the possibility to derive the thermal load-induced NSIFs through the SED approach and a coarse mesh is in this work demonstrated.

2. Preliminary considerations

In a previous work, Lazzarin et al. (2010) observed how the total elastic stain energy (Et) stored in a finite element depends only on the nodal displacements and finite element stiffness matrix according to the following relation:

Et = /V WdV = i{d}t[K]{d} (1)

with V and [K] being the volume and the stiffness matrix of the finite element, respectively, while {d} being the vector of nodal displacements. Thus, the elastic strain energy determination does not require any strain and stress calculation; it can be derived directly from nodal displacements. On the other hand, the calculation of stresses involves the

displacements derivative. For this reason, the mesh refinement necessary to capture the stress values must be much higher than that required for the accurate determination of the strain energy. Now, in a thermal-mechanical problem, it can be observed that the displacements are related to temperature through the thermal expansion coefficient, a. It is thus evident that the elastic strain energy induced by thermal loads depends only on nodal temperatures and shape functions of the finite element. No derivation or integration process are involved in the calculation. In such conditions, the degree of mesh refinement required for the determination of the strain energy induced by thermal loads is expected to be low as in the previous case.

3. Steady-state thermal problem

Under steady-state conditions, the temperature distribution near a sharp V-notch (Fig. 1) of an isotropic and homogeneous material is given by the following equation (Ferro et al. (2006)):

where T is the temperature, r and 0are the polar coordinates in the cylindrical reference system of Fig. 1, Ci are the generalized flux intensity factors (GFIFs) and Si and f are the eigenvalues and the eigenfunctions of the problem, respectively. These last parameters depend only on the V-notch angle and boundary conditions on the free edges of the notches (ri and T2 in Fig. 1).

It was found (Ferro et al. (2006)) that Eq. (2) can be simplified as follows:

where subscripts 5 and a stand for symmetric and anti-symmetric component, respectively. Furthermore, the following relations hold true:

T(r,0)=Srcirs'fi(0)

T(r, 0) = CsrSsfs(0) + CarSafa(0)

fs(0) = COS(0Ss), fa(0) = Sin(0Sa),

Fig. 1. Domain Q for the sharp V-notch problem

The thermal load was applied in terms of symmetric (Eq. 6) and anti-symmetric (Eq. 7) heat fluxes on the inner line (r) shown in Fig. 1. The V-notch surfaces (Ti and T2) were considered adiabatic while a reference value for the temperature was imposed in the notch apex (T(0,0) = 0).

<7(r = W0) = -Ag) =*(l"Tïhfl) f°r 0°<ö< 112-5°

'—'max

'd(r = rmax,e) = -A(£) ^(l+^ö) for -112.5°<0<O°

112.5° <9 < 112.5°

The parameters of the thermal problem are summarized in Table 1. The material properties refer to AISI 1008 steel.

Table 1. Parameters of the analysed problem

2ßO(°) rmax X X (thermal p (mass density) c (specific heat) aO □ linear expansion E (Young modulus)

(mm) (W/m2) conductivity) (kg/m3) (J/(kg°C)) coefficient) (m/(m°C)) (GPa)

(W/(mK))

135 10 5 65.2 7872 481 13.1 x 10-6 200

The thermal problem was solved by using three different mesh densities in order to evaluate the mesh sensitivity to the temperature distribution (Tab. 2). In particular, the analyses were carried out with ANSYS numerical code by using the finite element PLANE 77.

Fig. 2 shows the temperature distribution along the notch bisector induced by the symmetric heat flux and the temperature distribution along one adiabatic notch surface induced by the anti-symmetric heat flux. Values obtained with the highest mesh density are plotted with continuous lines, while those obtained with the lowest mesh density are marked with circles. Even if the results of such thermal problem were just published in a previous work (Ferro et al. (2006)), it is shown here that they are almost insensitive to the mesh density like nodal displacement values derived from mechanical analysis.

Table 2. Mesh densities used for the thermal numerical solution

Number FE

1000 100 10

0.1 0.01

Anti-symmetric heat flux 0 = 112.5°

Symmetric heat tlux 0 = 0° q/^

Ss= 1.6

sa = 0.8

0.1 r(nim)

Fig. 2. Comparison between thermal solutions obtained with the highest (continuous lines) and lowest (circles) mesh density

Fig. 3. Plots of the stress components along the notch bisector (9 = 0)

4. Strain energy density induced by thermal loads

Under linear thermo-elastic conditions, the stress distribution induced by a thermal load near the V-notch was shown to follow the William's solution (Ferro et al. (2006)):

ffij(r-e) = -¡^g[j°(0) + g[j2)(0) (i,j = r,0) (8)

where h and h are the first eigenvalues for mode I and II, respectively, while K% and Kth2 are the thermo-elastic stress intensity factors (TSIFs) which can be determined according to the definition proposed by Gross and Mendelson (1972):

K? = V2E lim r1"^iCT00(r, 0 = 0) (9)

K2h = V2S lim r1"^2xr0(r,0 = 0) (10)

The relevant eigenfunctions gij have closed form expressions (Lazzarin and Tovo (1998)). Now, under plane stress or plane strain conditions, the strain energy density averaged over a control volume of radius R is given by:

W = —

where E is the Young's modulus, ei and e2 are parameters which vary under plane stress or plane strain conditions and depend also on the notch opening angle 2p and the Poisson's ratio v (Lazzarin and Zambardi (2001)). R is a material parameter which in this work was set equal to 0.28 mm according to the value obtained for steel welded joints (Berto and Lazzarin (2009)). By rearranging Eq. (11) and by considering separately pure mode I and pure mode II loading, the TSIFs can by derived through the strain energy density as follows:

The mechanical analyses were carried out by using the same mashes of the previous thermal simulations (Tab 2) and by switching the thermal to mechanical element (PLANE 183). The temperature distribution obtained in the previous step was used as load for the stress and strain calculation. Material properties are summarized in Table 1. Node displacements belonging to the r surface were set equal to 0. In particular, Table 3 shows the mesh densities used to model the circular sector of radius R = 0.28 mm and the most relevant results obtained. Fig. 3 shows the thermal stress distribution along the notch bisector obtained with the finest mesh. It is noted the very good correlation between the numerical and theoretical solution by Williams (1952). In this case, by using Eqs. (9) and (10), the theoretical values of Kth1 and Kth2 were found equal to 2388.99 MPa mm0326 and 289.48 MPa mm-0302, respectively. It is evident from Table 3 that the strain energy density averaged over a control volume of radius R is almost mesh insensitive. This phenomenon was expected in force of the direct correlation between nodal temperatures and displacements and it is now demonstrated. It is interesting to note that the K% and Kth2 values obtained with the coarsest mesh and Eq. (12) differ from the theoretical values of 1.71% and 3.49%, respectively.

Table 3. Mesh densities used for the thermo-mechanical problem numerical solution

Number R = 0.28 mm K1th W2 K%h

FE (MJ/m3) (MPa mm0326) (MJ/m3) (MPa mm-0302)

5. Conclusions

The possibility to calculate a posteriori the thermal notch intensity factor through the strain energy density averaged over a control volume is demonstrated. The advantage of this approach is due to the possibility to use a very coarse mesh. As a matter of fact, it was shown that the strain energy value is almost mesh insensitive since it can be determined via the nodal displacements, without involving their derivatives. In thermo-mechanical analysis nodal temperatures are correlated to displacements. Thus, the strain energy density can be directly derived from nodal temperatures. This result is important in view of the possibility to calculate the thermal notch intensity factor values of 3D large structures without the necessity of high density meshes.

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