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Crack Modelling with the eXtended Finite Element Method

Francisco Xavier Girão Zenóglio de Oliveira

Thesis to obtain the Master of Science Degree in

Aerospace Engineering

Examination Committee

Chairperson: Prof. Fernando José Parracho Lau

Supervisor: Prof. Virgínia Isabel Monteiro Nabais Infante

Co-supervisor: Prof. Ricardo Miguel Gomes Simões Baptista

Members of the Committee: Prof. Ricardo António Lamberto Duarte Cláudio

Prof. José Miguel Almeida da Silva

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Agradecimentos

Deixo aqui o meu especial agradecimento `a senhora professora Virg´ınia Isabel Monteiro Nabais Infante, e ao senhor professor Ricardo Miguel Gomes Sim ˜oes Baptista, que sempre se revelaram dispon´ıveis na realizac¸ ˜ao desta tese.

Agradec¸o tamb ´em `a minha fam´ılia, que me tem sempre apoiado ao longo da minha vida acad ´emica. Por ultimo, a todos os meus amigos, de secund ´ario, bem como aqueles que conheci nesta casa, Instituto Superior T ´ecnico, que sempre me mostraram que a vida ´e mais do que somente o nosso trabalho.

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Acknowledgements

I leave here my special thanks to professor Virg´ınia Isabel Monteiro Nabais Infante, and professor Ricardo Miguel Gomes Sim ˜oes Baptista, who always proved to be available in the realization of this thesis.

I also thank to my family, who has always supported me through my academic life.

Finally, to all my friends from high school, as well as those I met in this house, Instituto Superior T ´ecnico, who always showed me that life is more than just work.

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Resumo

O mais importante para a ind ´ustria aeroespacial ´e a seguranc¸a do equipamento. Os engenheiros fazem um grande esforc¸o para garantir elevados padr ˜oes de qualidade. O estudo de fendas ´e extrema-mente importante para este prop ´osito.

A modelac¸ ˜ao de fendas, tem sido desde sempre um t ´opico muito importante. As abordagens tradi-cionais com o m ´etodo dos elementos finitos podem fornecer soluc¸ ˜oes precisas, no entanto, a construc¸ ˜ao das malhas ´e demorada e n ˜ao ´e ´obvia.

Um novo conceito emerge, conhecido como o M ´etodo de Elementos Finitos Extendidos, XFEM, em que as descontinuidades geom ´etricas e as sigularidades, s ˜ao introduzidas numericamente, com a adic¸ ˜ao de novos termos `as func¸ ˜oes de forma. Assim, a formulac¸ ˜ao em elementos finitos permanece a mesma, a representac¸ ˜ao da fenda ´e mais f ´acil, com uma soluc¸ ˜ao mais precisa.

Esta tese verifica a validade deste novo conceito para fendas estacion ´arias com ajuda do XFEM, implementado no Abaqus R. O crit ´erio de comparac¸ ˜ao ´e o factor de intensidade de tens ˜oes para

ge-ometrias simples. Os resultados computacionais s ˜ao pr ´oximos dos valores obtidos com as soluc¸ ˜oes dispon´ıveis na literatura e qualitativamente a simplicidade do m ´etodo ´e verificada.

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Abstract

The most important thing for the aerospace industry is the equipment’s safety. Engineers, make a great effort to guarantee high standards of quality. The study of crack phenomena is major for this purpose.

Crack modelling, has ever been an important topic. Traditional approaches of the finite element method can provide accurate solutions, nevertheless the meshing techniques are time consuming and not obvious.

A new concept emerges, known as the eXtended Finite Element Method, XFEM, where the geo-metric discontinuities and singularities, are introduced numerically with the addition of new terms to the classical shape functions. So, the finite element formulation remains the same, the crack representation is easier, with an approximate solution more precise.

This thesis, verifies the validity of this new concept for stationary cracks with Abaqus R’s XFEM aid.

The comparison criterion is the stress intensity factor for simple geometries. The computational results are near to the values obtained from the closed-forms available on the literature and qualitatively the simplicity of this method is checked.

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Table of Contents

Agradecimentos . . . V Acknowledgements . . . VII Resumo . . . IX Abstract . . . XI List of Tables . . . XV List of Figures . . . XVIII Abbreviations . . . XIX Nomenclature . . . XXIII 1 Introduction 1 1.1 Context . . . 1 1.2 Objectives . . . 2 1.3 Method . . . 2 1.4 Structure . . . 3 2 Bibliographic Research 5 2.1 Theory of Linear Elastic Fracture Mechanics . . . 5

2.1.1 Stress Distribution Around a Crack . . . 5

2.1.2 Loading Modes . . . 6

2.1.3 Stress Intensity Factors . . . 7

2.1.4 The Griffith Energy Balance . . . 8

2.1.5 The Energy Release Rate . . . 10

2.1.6 The J-Integral . . . 11

2.2 The Finite Element Method . . . 16

2.2.1 System of Equations . . . 16

2.2.2 Constitutive Relations . . . 16

2.2.3 Element Types . . . 17

2.3 The Classical Approach to the Stress Intensity Factor Calculation . . . 19

2.4 The eXtended Finite Element Method . . . 21

2.4.1 XFEM Enrichment: Jump Function . . . 21

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2.4.3 XFEM Limitations . . . 23

3 Case Studies 25 3.1 The SENT Specimen . . . 27

3.2 The CCT Specimen . . . 28

3.3 The SENB Specimen . . . 29

3.4 Cylindrical Pressure Vessel . . . 30

4 Numerical Study 33 4.1 The SENT Specimen Study . . . 33

4.1.1 Mesh Construction . . . 33

4.1.2 Requested Contours and GRef Influence . . . 36

4.1.3 Influence of the Ratio a/W . . . 41

4.1.4 DRef Influence . . . 42

4.1.5 Interpolation and Integration . . . 43

4.1.6 Standard Element Size Attribution . . . 46

4.1.7 SENT Classical Approach . . . 47

4.1.8 SENT Discussion . . . 49

4.2 The CCT Specimen Study . . . 51

4.2.1 Mesh Construction . . . 51

4.2.2 GRef Influence . . . 52

4.2.3 Standard Element Size Attribution . . . 53

4.2.4 CCT Classical Approach . . . 54

4.3 The SENB Specimen Study . . . 55

4.3.1 Mesh Construction . . . 55

4.3.2 GRef Influence . . . 56

4.3.3 Standard Element Size Attribution . . . 57

4.3.4 SENB Classical Approach . . . 58

4.4 Vessel Under Pressure, Closed-Form Deduction . . . 59

4.4.1 The Geometry, Mesh Construction and Boundaries Conditions . . . 59

4.4.2 Results And Discussion . . . 62

5 Conclusions 65 5.1 Achievements . . . 65

5.2 Future Work . . . 66

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List of Tables

4.1 Impact in the average error of the requested number of contours . . . 37

4.2 SENT analyses characteristics . . . 38

4.3 Numerical study, files appearance . . . 39

4.4 SENT GRef influence results . . . 40

4.5 Ratio a/W influence for GRef =80 divisions . . . 41

4.6 SENT DRef influence results . . . 42

4.7 C3D8R versus C3D4R analyses characteristics . . . 43

4.8 Integration effect . . . 44

4.9 Standard element size attribution SENT results . . . 46

4.10 SENT classical results . . . 48

4.11 CCT analyses characteristics . . . 52

4.12 CCT GRef influence results . . . 52

4.13 Standard element size attribution CTT results . . . 53

4.14 CCT classical results . . . 54

4.15 SENB analyses characteristics . . . 56

4.16 SENB GRef influence results . . . 56

4.17 Standard element size attribution SENB results . . . 57

4.18 SENB classical results . . . 58

4.19 Vessel convergence evaluation, analyses characteristics . . . 60

4.20 Vessel convergence evaluation . . . 61

4.21 Vessel results . . . 62

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List of Figures

2.1 Real and ideal crack tension behaviour . . . 5

2.2 Stresses near the crack tip and polar coordinates . . . 6

2.3 Load modes . . . 7

2.4 A through-thickness crack in an infinitely wide plate subjected to a remote tensile stress . 9 2.5 J-Integral, 2D contours . . . 12

2.6 3D J-Integral . . . 14

2.7 Finite element domain . . . 16

2.8 Abaqus R elements . . . . 17

2.9 Degeneration of a quadrilateral element into a triangle at the crack tip . . . 19

2.10 Degeneration of a brick element into a wedge . . . 19

2.11 Crack-tip elements for elastic and elastic-plastic analyses . . . 20

2.12 Spider-web mesh . . . 20

2.13 XFEM deduction, first mesh . . . 21

2.14 XFEM deduction, final mesh . . . 22

2.15 Abaqus R enrichment scheme . . . . 24

3.1 Edge crack in a semi-infinite plate subject to a remote tensile stress . . . 26

3.2 The SENT specimen . . . 27

3.3 The CCT specimen . . . 28

3.4 The SENB specimen . . . 29

3.5 Broken pressure vessel . . . 30

3.6 Vessel with semicircular crack . . . 31

3.7 Vessel geometry . . . 31

4.1 Perfectly structured mesh . . . 33

4.2 Unstructured mesh . . . 34

4.3 The SENT specimen final mesh . . . 34

4.4 Mesh refinement control . . . 35

4.5 Error evolution for different number of requested contours . . . 37

4.6 Time evolution with the GRef parameter . . . 39

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4.8 Error behaviour versus a/W ratio . . . 41

4.9 DRef influence . . . 42

4.10 Stresses comparison between the two kinds of integration . . . 45

4.11 Standard element size attribution, SENT absolute average error evolution . . . 46

4.12 Classical partition scheme . . . 47

4.13 SENT, Classical mesh . . . 47

4.14 CTT mesh . . . 51

4.15 CCT contour error evolution without GRef = 10 . . . 52

4.16 Standard element size attribution, CTT absolute average error evolution . . . 53

4.17 CCT classical partition scheme . . . 54

4.18 CCT Classical mesh . . . 54

4.19 SENB partition scheme and mesh . . . 55

4.20 SENB mesh detail . . . 55

4.21 SENB contour error evolution . . . 56

4.22 Standard element size attribution, SENB absolute average error evolution . . . 57

4.23 SENB classical partition scheme . . . 58

4.24 SENB Classical mesh . . . 58

4.25 Vessel geometry with boundary conditions and internal pressure . . . 59

4.26 Vessel lateral view with semicircular crack . . . 59

4.27 Vessel radial stress distributions for different ESize values . . . 60

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Abbreviations

C3D4 Abaqus R Linear Tetrahedral Element with 4 nodes

C3D8 Abaqus R Linear Hexahedral Element with 8 nodes

C3D10 Abaqus R Quadratic Tetrahedral Element with 10 nodes

C3D20 Abaqus R Quadratic Hexahedral Element with 20 nodes

CAE Computed Assisted Engineering CCT Centre Crack Tension

DOF Degrees of Freedom DRef Depth Refinement

ESize Element Size for the Standard Element Size Attribution FEM Finite Element Method

GRef Global Refinement

LEFM Linear Elastic Fracture Mechanics SENB Single Edge Notch Bending SENT Single Edge Notched Tension XFEM eXtended Finite Element Method

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Nomenclature

 Strain

Γ Contour containing the crack tip γs Surface energy µ Shear modulus ν Poisson coefficient Ω Finite domain Φ Total energy Π Potential energy

Ψi Crack tip asymptotic solutions

σf Fracture stress

σh Hoop stress

σij Stress tensor

σrr Radial stress

θ Polar coordinate, angle

A Crack area

a Crack length

ai Vector of enriched nodes with the Jump function

Am Amplitude, dimensionless function of θ

B Thickness

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C Closed contour

E Young’s modulus

f Volume forces

fij Dimensionless function of θ in the leading term

G Energy release rate

H Jump function

I Identity matrix

i Number of the Depth Refinement

J J-Integral

j Number of requested contours

K Stress Intensity Factor

k Constant

Kc Fracture toughness

Ki Stress intensity factor, i mode

m Interior normal

n Exterior normal

Ni Shape functions

P Pre-logarithmic, energy factor tensor

pint Vessel internal pressure

q Vector within the virtual displacement

r Polar coordinate, radius

ro Outer radius

rp Plastic radius

ri Inner radius

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t Tension forces

ui Displacement on the i direction

W Elastic strain energy

Ws Work necessary to create new crack faces

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Chapter 1

Introduction

1.1

Context

Nowadays, the study and evaluation of the integrity of the various mechanical components is tremely important for the industry. The two major goals are: increase the components fatigue life ex-pectancy, and their safety.

Previously, the study and analyses were constantly made from laboratory experiments, which were lengthy, expensive [1] and sometimes difficult to implement, due to several rules [2]. Today, there are new materials and mechanical components appearing to a daily rhythm, which must be tested and certified to be available for the consumers.

Inevitably, the engineers presented with systematic solutions, with higher precision, such as the Finite Element Method, making the whole process effective and efficient. However, some mechanical phenomena have always been difficult to model with the Finite Element Method (FEM), amongst which the study of cracks, stationary and especially its propagation. This particular study is necessary in order to predict the mechanical behaviour of the equipment but also in order to increase its life expectancy.

Thus, there has always been a great need of representing correctly cracks, with accurate results and an expeditious manner. There were several attempts, which proved to be difficult to enforce for several reasons, the main one being the mesh generation around the crack tip [1] to obtain good results. It is in these terms that a new concept emerges recently; XFEM, eXtended Finite Element Method, creating a new paradigm in the study of cracks [3, 4].

The method of extended elements permits a representation of cracks by finite elements, which does not require to change the mesh to monitor crack propagation [5], causing a revolution when compared with the classical methods. The discontinuity in the elements is described from enrichments functions overlapping the elements.

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1.2

Objectives

The XFEM, although a relatively recent concept, 1999, is now available in commercial versions of finite elements. In fact, the XFEM has been already implemented in the Finite Element Method software Abaqus R / Standard, owned by Simulia [6].

A need therefore arises of realizing the potential and the Abaqus R’s XFEM validity in order to

un-derstand the possibilities for subsequent analyses. The aim is thus to evaluate the XFEM for crack modelling in Abaqus R.

The basic idea behind the theme is to understand how accurate is the increase brought by the XFEM in the fracture mechanics domain, to predict the integrity of mechanical systems, according to standard methods. The main objective is to conduct a convergence analysis of the FEM discretizations, identifying aspects and important parameters in the cases studied.

The understanding of which are the best modelling techniques with the aid of XFEM, to achieve the best precision and efficiency within the mechanical behaviour of materials, will be the main objectives of this thesis. So, some simulations will be performed, with solutions known by the scientific community, using Abaqus R’s XFEM.

1.3

Method

As first instance, it is made a bibliographic research to understand the principles of the linear elastic fracture mechanics theory, in order to explain the calculation of the stress intensity factor; the main parameter under analysis in this thesis. Then, the fundamentals behind the XFEM are also briefly presented.

Next, it is done an evaluation of the XFEM in Abaqus R, where for a given geometry, with a solution

available in the literature for its stress intensity factor, it is studied the advantages and disadvantages of using the XFEM for different meshes, elements, rules of integration and interpolation.

Furthermore, a better validation of the XFEM is achieved with two others geometries, to verify and confirm the conclusions from the first geometry analysis, to once again investigate the quality and appli-cability of XFEM.

Finally, a more complex geometry is used in order to achieve a better understating of the XFEM limitations, allowing this thesis conclusion.

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1.4

Structure

Chapter 2 discusses initially the basic concepts of the Linear Elastic Fracture Mechanics, where it can be understood how to calculate the stress intensity factor. In the second phase, the traditional finite elements and their constitutive laws are presented, allowing after the introduction of the XFEM, where its formulation is also presented.

Chapter 3 makes a brief presentation of the geometries under analysis. Their geometric character-istics are briefly commented, and the closed-forms for theirs stress intensity factors are introduced.

In chapter 4, all the numerical analyses are carried. First, several analyses are preformed. In the end the more important are identified in order to be better investigated with the other geometries. Finally, a more complex structure, consisting of a vessel under internal pressure, is analysed in order to conclude about the XFEM use for structures analysis.

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Chapter 2

Bibliographic Research

2.1

Theory of Linear Elastic Fracture Mechanics

2.1.1

Stress Distribution Around a Crack

The cracks in mechanical components subject to applied loads behave very close to what is observed when there are notches, which are responsible for stress concentration due to reduction of area against the nominal area. The geometry of the crack creates high stress concentrations in its tip. This behaviour is illustrated in figure 2.1. Due to the high tension observable on the edge of the crack, a plastic zone appears. However, following the LEFM theory, the plastic behaviour is not taken into account, and tension is given by an ideal crack following the linear elastic model. Consequently, the LEFM reveals a large gap, by not taking into account areas that could be in the plastic domain [7].

Figure 2.1: Real and ideal crack tension behaviour [7]

Consulting Anderson [7], for cracked geometries subjected to external forces, it is possible to derive closed-forms or analytical expressions for the stresses in the body, assuming the LEFM. Irwin [8],

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Sneddon [9], Westergaard [10], and Williams [11] were among the first to publish such solutions.

Figure 2.2: Stresses near the crack tip and polar coordinates [7]

Considering a polar coordinate system, (r, θ), with the origin at the crack tip represented on figure 2.2, it can be shown that the stress field in any linear elastic cracked body is given by,

σij = ( k √ r)fij(θ) + ∞ X m=0 Amrm/2g (m) ij (θ) (2.1)

Where σij is the stress tensor, k a constant and fij a dimensionless function of θ in the leading

term. For the higher-order terms, Am is the amplitude and a dimensionless function of θ for the mth

term. It should be noticed, that the solution for any given configuration contains a leading term that is proportional to 1/√r. As r → 0, the leading term goes to infinity, but the other terms remain finite or approach zero. Thus, stress near the crack tip varies with 1/√r, independently of the geometries. It can also be shown that displacement near the crack tip varies with√r. Equation 2.1 describes also a stress singularity, since stress is asymptotic to r = 0.

2.1.2

Loading Modes

In fracture mechanics, there are three types of loading that a crack can experience, presented on figure 2.3.

Mode I loading, where the principal load is applied in the normal direction to the crack plane, opening the crack (a traction mode). Mode II corresponds to an in-plane shear loading and tends to slide one crack face over the other (a shear mode). Mode III refers to the out-of-plane shear (a torsion mode).

A cracked body can be loaded in any one of these modes, or a combination of two or three modes. For each of these modes, it can be deduced a stress intensity factor, which are presented next.

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Figure 2.3: Loading modes I, II and III [7]

2.1.3

Stress Intensity Factors

The stress intensity factors are used as a measure that quantifies the severity of a crack relatively to others cracks [7]. They are so, of extreme importance for the cracks study. They are also related to the mechanisms of crack initialization but also their propagation, and in some cases, the stress in-tensity factor may reach an extreme value: the fracture toughness KC, leading to the fracture of the

components.

Each mode of loading produces the1

r singularity at the crack’s tip, but the proportionality constants

kand fij depend on the mode. For further considerations it is important to substitute k of equation 2.1

by the stress intensity factor K, where K = k√2π. The stress intensity factor is usually given with a subscript to denote the mode of loading, i.e., KI, KII, or KIII.

Considering the LEFM, the stress fields ahead of a crack tip in an isotropic linear elastic material can be written as, limr→0σ (I) ij = KI √ 2πrf (I) ij (θ) (2.2) limr→0σ (II) ij = KII √ 2πrf (II) ij (θ) (2.3) limr→0σ (III) ij = KIII √ 2πrf (III) ij (θ) (2.4)

For modes I, II, and III, respectively.

In a mixed-mode problem (i.e., when more than one loading mode is present), the individual contri-butions to a given stress can be summed:

σij(T otal)= σ(I)ij + σij(II)+ σij(III) (2.5)

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Considering this thesis will focus on loading Mode I, it is shown below both the stress and displace-ment field ahead a crack tip,

σxx= KI √ 2πrcos(θ)  1 − sin(θ 2)sin( 3θ 2 )  (2.6) σyy = KI √ 2πrcos( θ 2)  1 + sin(θ 2)sin( 3θ 2 )  (2.7) τxy= KI √ 2πrcos( θ 2)sin( θ 2)cos( 3θ 2 ) (2.8)

For plane stress,

σzz= 0 (2.9)

For plane strain,

σzz= ν(σxx+ σyy) (2.10)

Where (r, θ) are the polar coordinates, ν the Poisson ratio. The remaining components of the stress tensor are zero. For the displacement field,

ux= KI 2µ r r 2πcos( θ 2)  κ − 1 + 2sin2(θ 2)  (2.11) uy= KI 2µ r r 2πsin( θ 2)  κ + 1 − 2cos2(θ 2)  (2.12)

Where u is the displacement, µ is the shear modulus and κ = 3 − 4ν for plane strain and κ = 3−ν1+ν for plane stress.

2.1.4

The Griffith Energy Balance

According to the first law of thermodynamics, when a system goes from a non-equilibrium state to equilibrium, there is a net decrease in energy [7].

In 1920, Griffith applied this idea to the formation of a crack [12]:

”It may be supposed, for the present purpose, that the crack is formed by the sudden annihi-lation of the tractions acting on its surface. At the instant following this operation, the strains,

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and therefore the potential energy under consideration, have their original values; but in gen-eral, the new state is not one of equilibrium. If it is not a state of equilibrium, then, by the theorem of minimum potential energy, the potential energy is reduced by the attainment of equilibrium; if it is a state of equilibrium, the energy does not change.”

The total energy must decrease or remain constant to form a crack or to allow its propagation. Thus the critical conditions for fracture can be defined as the point where the crack growth occurs under equilibrium conditions, with no net change in the total energy.

Consider a wide plate subjected to a constant stress load with a crack 2a long (figure 2.4). In order for this crack to increase in size, sufficient potential energy must be available in the plate to overcome the surface energy γs of the material. The Griffith energy balance for an incremental increase in the

crack area dA, under equilibrium conditions, can be expressed as follow:

dΦ dA = dΠ dA + dWs dA = 0 (2.13) Or, dΠ dA = − dWs dA (2.14)

Where Φ is the total energy, Π the potential energy supplied by the internal strain energy and external forces and Wsthe work required to create new surfaces.

Figure 2.4: A through-thickness crack in an infinitely wide plate subjected to a remote tensile stress [7]

From Anderson [7], for the cracked plate illustrated in figure 2.4, Griffith [12] used the stress analysis of Inglis to show that,

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Π = Π0−

πσ2a2B

E (2.15)

Where Π0 is the potential energy of an uncracked plate and B is the plate thickness. Since the

formation of a crack requires the creation of two surfaces, Wsis given by

Ws= 4aBγs (2.16)

With γsthe surface energy of the material. Thus,

−dΠ dA = πσ2a E (2.17) And, dWs dA = 2γs (2.18)

Solving for the fracture stress,

σf = (

2Eγs

πa )

1/2

(2.19)

It is important to have in mind the distinction between crack area and surface area. The crack area is defined as the projected area of the crack (2aB in the present example), but since a crack includes two matching surfaces, the surface area is 2A .

2.1.5

The Energy Release Rate

Irwin [8], in 1956, proposed an energy approach for fracture that is essentially equivalent to the Griffith model, except that Irwin approach is in a form more convenient for solving engineering problems. Irwin defined an energy release rate G, which is a measure of the energy available for an increment of crack extension:

G = −dΠ

dA (2.20)

The term rate, as it is used in this context, does not refer to a derivative with respect to time; G is the rate of change in potential energy with the crack area. Since G is obtained from the derivative of a potential, it is also called the crack extension force or the crack driving force.

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According to equation 2.17, the energy release rate for a wide plate in plane stress with a crack of length 2a (figure 2.4) is given by,

G = πσ

2a

E (2.21)

Referring to the previous section, the crack extension occurs when G reaches a critical value,

Gc=

dWs

dA = 2γs (2.22)

Where Gcis also a measure of the fracture toughness of the material.

At this moment, it must be said the energy release rate is extremely important for this thesis. This is justified by its direct relationship with the stress intensity factor.

Irwin’s showed that for linear elastic materials, under loading Mode I, it may be written,

G =K

2 I

E0 (2.23)

Where for plane stress,

E0 = E (2.24)

And for plane strain,

E0 = E

1 − ν2 (2.25)

Nevertheless, the energy release rate is still not enough and practical to get the value of the stress intensity factor.

2.1.6

The J-Integral

In the previous section, it was showed the basis behind the energy release rate, with a direct relation with the stress intensity factor. Even the energy release rate is a simple concept, it is not obvious how to deduce it with the finite element method. Fortunately, there is another concept in the LEFM theory, called the J-Integral, which may be calculated numerically and reveals itself very useful because in the context of LEFM, the J-Integral is equal to the energy release rate G.

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J-Integral Calculation

As said before, the stress intensity factor can be calculated by the energy release rate G, which in this thesis context is equal to the J-Integral.

The J-Integral is a contour integral for bi-dimensional geometries (see figure 2.5). Its definition is easily extended to three-dimensional geometries, and it is used to extract the stress intensity factors.

Figure 2.5: a) 2D contour integral, b) 2D closed contour integral [6] For the two-dimensional case, the J-Integral is given by,

J = limΓ→0

Z

Γ

n.H.qdΓ (2.26)

Where Γ is the contour containing the crack tip, n is the exterior normal to the contour, and q is the unitary vector within the virtual displacement direction of the crack.

The function H is defined by,

H = W I − σ.∂u

∂x (2.27)

Where W is the elastic strain energy1, I the identity tensor, σ the stress tensor and u the vector of

displacements.

The contour Γ connects the two crack faces and encloses the crack tip. This is shown in figure 2.5 a). The contour tends to zero, until it only contains the crack tip (equation 2.26). The exterior normal n moves along the integration while q stands fixed in the crack tip.

1The strain energy definition may also include the elastic-plastic effects, which are not presented, considering the fact that they

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It is very important to note the J-Integral is independent of the chosen path for elastic materials in the absence of imposed forces in the body or tension applied on the crack, so the contour does not need to contract itself on the crack, but it has only to enclose the crack tip.

The two dimensional integral may be rewritten as a closed bi-dimensional contour integral as the following [13], J = − I C+C−+C++Γ m.H.¯qdΓ − Z C++C− t.∂u ∂x.qdΓ¯ (2.28)

Where the line integrals are preformed in a closed contour, which is an extension of Γ. C+and C−

are contours along the crack faces, enclosed by C. The normal m had to be introduced as the unitary exterior normal to the contour C, respecting m = −n. The function ¯qhad also to be introduced, being a unitary vector applied in the direction of the virtual extension of the crack tip, which respects ¯q = qin Γ and vanishes in C.

In equation (2.28), t is the tension on the crack faces. Crack tension is also a subject not considered in this thesis. The second term may be erased from equation 2.28.

The J-Integral may be now transformed in a surface integral by the divergence theorem properties, yielding to, J = Z S ( ∂ ∂x).(H ¯.q)dS (2.29)

Where S is the area in the closed domain. The equilibrium forces equation is,

( ∂

∂x).σ + f = 0 (2.30)

Where σ is the tension tensor, and f the volume forces. And the energy strain gradient, for an homogeneous material with constant properties is,

(∂W () ∂x ) = ∂W ∂ ∂ ∂x = σ ∂ ∂x (2.31)

Where  is the mechanical strain.

Considering these two previous equations, the J-Integral may now be written as,

J = − Z S (  H∂ ¯q ∂x+ (f. ∂u ∂x).¯q  )dS (2.32)

The bi-dimensional equation for the J-Integral is easily extended to a three dimensional formulation. The J-Integral has to be defined in order to a parametric variable s, in the crack front, in such manner

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J (s)is defined by a function which characterizes the bi-dimensional J-Integral for each point placed in the path defining the crack front, which is also described parametrically in order to s (figure 2.6).

Figure 2.6: a) Local coordinates system, b) 3D J-Integral [6]

The local system of cartesian coordinates is placed in the crack front. See figure 2.6 a). The axis x3, runs tangentially the crack, x2is defined perpendicular to the crack front. In this formulation, x1will

always be directed forward at the crack front. x1and x2define a perpendicular plane to the crack front.

J(s) is so described in the x1x2plane.

From figure 2.6, it is obvious that for the three-dimensional case, each infinitesimal 2D contour must be integrated, for each position of s, along the path described by the crack front in order to obtain a volume J-Integral.

Stress Intensity Factors Extraction

Having defined the procedure to obtain the J-Integral, for both, bi-dimensional and three-dimensional crack geometries, it becomes necessary to extract the stress intensity factors. Consulting Abaqus R

Documentation [6], for a linear elastic material, the J-Integral is related to the stress intensity factors by the following relation,

J = 1 8πK

TP−1K (2.33)

With K = [KI, KII, KIII] T

and P the pre-logarithmic energy factor tensor [14, 15, 16, 17]. For homogeneous and isotropic materials this equation may be simplified in the form,

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J = 1 E0(K 2 I + K 2 II) + 1 2GK 3 III (2.34)

Where, E0is given by equations 2.24 and 2.25.

At last, for pure Mode I loading, the relation between the J-Integral and the stress intensity factor is given by,

J = (K

2 I

E0) (2.35)

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2.2

The Finite Element Method

2.2.1

System of Equations

In this section, are presented the governing equations of the finite elements, used for the analyses.

Figure 2.7: FEM domain and boundary condtions [5]

Considering the domain Ω of figure 2.7, the border Λ may be divided in four independent borders: Λt

with tension applied, Λuwith imposed displacements, and the last two domains,Λ+c and Λ−c, representing

the crack faces.

The equilibrium equations and boundary conditions are,

∇σ + f = 0 in Ω (2.36)

σ.n = ¯t in Λt (2.37)

σ.n = 0 in Λ+c (2.38)

σ.n = 0 in Λ−c (2.39)

u = uimpin Λu (2.40)

Where f represents the volume forces, n the outer normal, ¯tthe superficial forces, uimpthe imposed

displacements. Finally, considering an infinitesimal deformation δv, the weak formulation is,

Z Ω σ∇δv = Z Λt tδvdΓ + Z Λ f δvdΩ (2.41)

2.2.2

Constitutive Relations

Although the fact the weak formulation has always the same form, the element quality depends on the constitutive relations, as well of the selected shape forms.

This thesis is limited to the linear elastic fracture mechanics, implying that only small strains will be considered. The material model could not be different than the presented next, which is a limitation imposed by the commercial software of analysis Abaqus R, allowing only this model for XFEM aid.

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Respecting the elasticity theory, the tension obeys to the following,               σxx σyy σzz σxy σxz σyz               = D               xx yy zz 2xy 2xz 2yz               (2.42)

Being D given by,

D = E(1 − ν) (1 + ν)(1 − ν)               1 1−νν 1−νν 0 0 0 ν 1−ν 1 ν 1−ν 0 0 0 ν 1−ν ν 1−ν 1 0 0 0 0 0 0 2(1−ν)1−2ν 0 0 0 0 0 0 2(1−ν)1−2ν 0 0 0 0 0 0 2(1−ν)1−2ν               (2.43)

Where σijand ij are the tension and strain components, E the Young’s modulus, and ν the Poisson

coefficient.

2.2.3

Element Types

In Abaqus R, the geometries under analysis can be modelled with two types of volumetric

ele-ments: the tetrahedral and hexahedral, which remains the isoparametric element most used for three-dimensional elasticity [18].

Abaqus R admits two formulations of this element. The linear element of 8 nodes, identified as C3D8,

and the quadratic of 20 nodes, C3D20 (figure 2.8 a and b).

At each node, for both elements, there are three degrees of freedom, corresponding to three possible displacements. Thus, the element of 8 nodes, has 24 degrees of freedom, a number three times lower than the 60 degrees of freedom of the element of 20 nodes [6].

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As for the tetrahedral, there is a linear element of 4 nodes, C3D4 and a quadratic element with 10 nodes, the C3D10 (figure 2.8 c).

Any of the four elements allows two kinds of numerical integration; reduced or full integration. The reduced integration is identified by a R in the element code. For example, C3D20R, indicates a three-dimensional element of 20 nodes, being a quadratic with reduced integration.

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2.3

The Classical Approach to the Stress Intensity Factor

Calcu-lation

In order to have a full understanding of the XFEM, it is necessary to evaluate the geometries with the classical approach for the stress intensity factor calculation.

In the classical approach, according to [7], in two-dimensional problems quadrilateral elements are collapsed to triangles where three nodes occupy the same point in space, like what is shown on figure 2.9. For three dimensions problems, a brick element is degenerated to a wedge (figure 2.10).

Figure 2.9: Degeneration of a quadrilateral element into a triangle at the crack tip [7]

Figure 2.10: Degeneration of a brick element into a wedge [7]

In elastic problems, the nodes at the crack tip are normally tied, and the mid-side nodes moved to the 1/4 points. This modification is necessary to introduce a 1/√r strain singularity in the element, which brings numerical accuracy due to the fact that the analytical solution contains the same term.

A similar result can be achieved by moving the midside nodes to 1/4 points in non collapsed quadri-lateral elements, but the singularity would exist only on the element edges; collapsed elements are preferable in this case because the singularity exists within the element as well as on the edges.

When a plastic zone forms, the singularity no longer exists at the crack tip. Consequently, elastic singular elements are not appropriate for elastic-plastic analyses. Figure 2.11 shows an element that exhibits the desired strain singularity under fully plastic conditions.

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Figure 2.11: Crack-tip elements for elastic and elastic-plastic analyses. Element (a) produces a 1/√r strain singularity, while (b) exhibits a 1/r strain singularity (a) Elastic singularity element and (b) plastic singularity element [7]

According to [1], [6] and [7], for typical problems, the most efficient mesh design for the crack-tip region is the “spider-web” configuration (figure 2.12), consisting of concentric rings of quadrilateral ele-ments that are focused toward the crack tip. The eleele-ments in the first ring are degenerated to triangles, as described above. Since the crack tip region contains stress and strain gradients, the mesh refinement should be greater at the crack-tip. The spider-web design allows a smooth transition from a fine mesh at the tip to a coarser. In addition, this configuration results in a series of smooth, concentric integration domains (contours) for the J-Integral calculation.

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2.4

The eXtended Finite Element Method

The extended finite element method, XFEM, is an evolution of the classical finite element method based on the concept of partition unit, i.e. the sum of shape functions is equal to one.

This method was initially developed by Ted Belytschko [3] and his colleagues in 1999. The XFEM based on the concept of partition of unity [19], adds a priori known information about the solution of a given problem, to the FEM formulation, making possible, for example, to represent discontinuities and singularities, independently of the mesh. This particular feature makes this method very robust and attractive to simulate the propagation of cracks, since it is no longer necessary to have a continual updating of the mesh. The XFEM is then referenced as aMeshless method.

In the XFEM, enrichment functions are added to additional nodes, in order to include information about discontinuities and singularities around the crack. These functions are the asymptotic near-tip solutions, which are sensitive to singularities, and the Jump function, which simulates the discontinuity when the crack opens.

2.4.1

XFEM Enrichment: Jump Function

To explain the form how the discontinuities are added to the FEM, consider a simple bi-dimensional geometry (figure 2.13), with four elements and an edge crack.

Figure 2.13: Bi-dimensional geometry for the XFEM deduction The solution for the displacement is typically given by,

u(x, y) =

10

X

i=1

Ni(x, y)ui (2.44)

Where Ni(x, y)is the shape function on node i with coordinates (x, y), and ui is the displacement

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Defining c as a middle point between u9and u10and d as the distance between the two nodes, it is possible to write, c = u9+ u10 2 (2.45) d = u9− u10 2 (2.46) But also, u9= c + d (2.47) u10= c − d (2.48)

Manipulating the expression 2.44,

u(x, y) =

8

X

i=1

Ni(x, y)ui+ c(N9+ N10) + d(N9+ N10)H(x, y) (2.49)

Where the Jump function was added obeying to,

H(x, y) =      1 y > 0 −1 y < 0 (2.50)

We may substitute (N9+ N10)by N11and c by u11.

Figure 2.14: Bi-dimensional geometry, the final mesh, with the new node The displacement approximation becomes,

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u(x, y) =

8

X

i=1

(Ni(x, y)ui+ N11u11) + (dN11H(x, y)) (2.51)

The first term of the equation is the conventional approximation by the FEM, and the second term cor-responds to the Jump enrichment, associated to a new node, with displacement u11, and consequently

new degrees of freedom.

The equation 2.51, reveals that the geometry crack can then be represented by a mesh which does not contain any discontinuity, since this is included in the equation due to the presence of the enrichment term reproducing the crack discontinuity.

2.4.2

XFEM Enrichment: Asymptotic Near-Tip Singularity Functions

In the previous section, it was shown which enrichment function can capture the discontinuity due to the presence of a crack or notch. It remains to illustrate how to pick up the singularities existing in this type of problem. The enrichment function appointed above, will be introduced in all the elements around the crack. However, the following asymptotic solutions will be introduced only at the crack tips.

Let all nodes be represented by the set S, the nodes that comprise the tips are the set Sc and the

other, comprising the physical discontinuity, are the set Sh.

The approximation becomes,

u(x, y) = 8 X i∈S Ni(x, y)(ui+ H(x, y)ai i∈Sh + 4 X i=1 ψi(x, y)bi i∈Sc ) (2.52)

Where uiis the nodal displacement, airepresents the vector of enriched nodes with the discontinuity

function, and bi are the nodes enriched with the crack tip asymptotic solutions.

The enrichment functions at the crack tips, for linear elastic isotropic materials are given by,

{ψi(x, y)}4i=1= (

rsin(α),√rcos(α/2),√rsin(α/2)sin(α),√rcos(α/2)sin(α)) (2.53)

Where (r, α) correspond to the polar coordinates of point with Cartesian coordinates (x, y).

2.4.3

XFEM Limitations

Although the XFEM has been developed for the study of static or propagating cracks, this work will focus mainly on the static case, since crack propagation is even complex and heavy in computational terms, with numerical convergence issues. Moreover, for propagating cracks, the asymptotic near-tip singularity functions are not included in the enrichment scheme (figure 2.15).

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Figure 2.15: Abaqus R enrichment schemes for both, stationary and propagating cases [20]

Another limitation relates to the fact that all the geometries presented are three-dimensional since the considered Abaqus R version (CAE 6.12) does not allow the calculation of the stress intensity factor

in two-dimensional geometries with XFEM use.

All simulations will assume linear elastic materials and the theory of linear elastic fracture. This limitation is imposed by the Abaqus R, which only allows stationary analysis with linear elastic materials,

since Abaqus R 6.12 only contemplates the near-tip asymptotic solutions for the stress field at the crack’s

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Chapter 3

Case Studies

This thesis purpose is to validate the XFEM in the context of mechanical behaviour, most precisely, the fracture mechanics. The most interesting parameter to consider is the stress intensity factor; a value very useful to quantify the crack importance on the stress distribution. In order to emphasize the possible advantages or disadvantages brought by the XFEM, it must be investigated the quality of the solutions with XFEM aid. From Anderson [7], it is very hard to find the closed-forms or analytical solutions for the stress intensity factor of a given loaded geometry. Nevertheless, there are still some solutions for very simple configurations. For the more complex geometries, the stress intensity factor must be derived numerically, by finite elements. For example, one configuration for which a closed-form solution exists is a through crack in an infinite plate subjected to a remote tensile stress (figure 2.4). It can be shown the solution in this case, is given by,

KI = Y (a)σ

πa (3.1)

Y (a) = 1 (3.2)

Thus the amplitude of the crack-tip singularity for this configuration is proportional to the remote stress and the square root of the crack size. Note this solution is for a infinite body, turning it useless for this thesis. There is also a related solution for a semi-infinite plate with an edge crack, figure 3.1, from [7].

This configuration can be obtained by slicing the plate in figure 2.4 through the middle of the crack. The stress intensity factor for the edge crack is given by,

KI = 1.12σ

πa (3.3)

This is very close to equation 3.1. The 12% increase in KI for the edge crack is caused by different

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Figure 3.1: Edge crack in a semi-infinite plate subject to a remote tensile stress [7]

is less restrained than the through crack, which forms an elliptical shape when loaded. Even though, a semi-infinite body is not simple to be simulated in the considered Abaqus R version. Resuming, most

configurations for which there is a closed-form K solution consist of a crack with a simple shape in an infinite or semi-infinite plate. Presented in another way, the crack dimensions are small compared to the size of the plate; the crack-tip conditions are not influenced by external boundaries. In all the remaining cases, where the body has finite dimensions, a solution is not possible. For these cases, the existing solutions were obtained by finite element analysis through the last decades. Solutions of this type are usually fit to a polynomial expression. Several handbooks devoted solely to stress intensity solutions have been published,[21, 22].

The most common cases are abridged by Mohammadi [4] and also by Anderson [7]. They will be the main support for this thesis analyses. The presented solutions will be assumed as the reference solutions of the considered cases. They are known and approved by the scientific community, making them very reliable.

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3.1

The SENT Specimen

The first geometry to be considered for the analysis is known as the SENT specimen, an acronym for Single Edge Notched Tension plate. It consists of a parallelepiped specimen with a corner crack in its longitudinal symmetry axis. The choice of this specimen to the first analysis is due to its popularity in experimental studies of fracture mechanics. Due to this fact, this sample is likely to have an accurate empirical solution for the stress intensity factor.

Figure 3.2: The SENT specimen

As can be seen in the figure 3.2, the sample has a height H, width W , an edge crack of width a, and thickness B. The solution for the stress intensity factor of this specimen, accessed in [4], is given by,

KI = Y (a)σ √ πa (3.4) Y (a) = [1.12 − 0.23(a W) + 10.56( a W) 2− 21.74(a W) 3+ 30.42( a W) 4] (3.5)

This specimen is the basis of study for most of the analyses presented in this thesis. The results obtained for the different analysis performed in this specimen, allow the study of the others geometries, with superior knowledge about XFEM.

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3.2

The CCT Specimen

The second specimen considered for the numerical investigation is another one very popular among the scientific community identified as CCT, an acronym for Center Crack Tension plate. The stress intensity factor for a through crack 2a long, at right angles, in an infinite plane, with an uniform stress field applied, is given by the equation 3.1.

Figure 3.3: The CCT specimen

Considering the figure 3.3, the specimen has a height H, width W , and a centre crack 2a long. Consulting [4], the stress intensity factor is,

KI = Y (a)σ √ πa (3.6) Y (a) = [1 + 0.256( a W) − 1.152( a W) 2+ 12.2( a W) 3] (3.7)

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3.3

The SENB Specimen

The last considered specimen is the Single Edge Notch Bending specimen, identified by the acronym SENB. This one, unlike the others presented, is not under traction load but in bending, caused by a three point load apply.

Figure 3.4: The SENB specimen [7]

The specimen is horizontally positioned, with three forces applied; P on the bottom, and P/2 on each of the two top corners. The span is S = 4W , contrasting with the height W . The thickness is B.

From Bower [23], KI = 4P B r π W h 1.6( a W) 1/2− 2.6( a W) 3/2+ 12.3( a W) 5/2− 21.2( a W) 7/2+ 21.8( a W) 9/2i (3.8)

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3.4

Cylindrical Pressure Vessel

After the validation of the XFEM with the previous specimens, with the acquired knowledge, it will be deducted a closed-form solution for a cylindrical vessel under pressure.

Cylindrical vessels, under internal pressure, are commonly used in many engineering applications. They may be used as industrial compressed air receivers and domestic hot water storage tanks. Other examples of pressure vessels are the distillation towers, pressure reactors, autoclaves. They are also present in oil refineries and petrochemical plants, nuclear reactor vessels, submarine, aircraft and space ship habitats, and storage vessels for liquified gases such as ammonia, chlorine, propane, butane, and LPG.

Due to fatigue loads, vibrations and thermal stresses, the vessels are affected by crack initiation and propagation until the inevitable fracture (figure 3.5).

Figure 3.5: Broken pressure vessel <http://www.pjedwards.net/>

The existence of closed-forms for the stress intensity factors in vessels are of extreme importance. With this structure, the goal is to preform a set of analyses, producing a closed-form under certain conditions, and compare it with one existing solution.

For a semicircular crack (figure 3.6 with c = a), from [24], the closed-form solution is,

KI = 1.12

2 πσh

πa (3.9)

Where σh is the hoop stress, and a the crack radius. The hoop stress, is related with the internal

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Figure 3.6: Vessel with semicircular crack <http://www.afgrow.net/applications/DTDHandbook>

stresses: the hoop and the longitudinal. The hoop stress in a thin vessel is given by,

σh=

pintri

t (3.10)

Where pintis the internal pressure, rithe inner radius and t the thickness (figure 3.7).

It was noted that equation 3.5 is actually the closed-form for a semicircular crack on a infinite plate. For a cylinder, the closed-form should be function of θ. No other theoretical expression was found for this case study. Nevertheless, it was decided to pursuit with this equation. The reason is that for elliptic cracks without eccentricity in a pressure vessel it may be assumed that K is not function of θ [25]. In short, with the vessel study, it is intended to show if the XFEM may be used to investigate closed-forms for any structure.

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Chapter 4

Numerical Study

4.1

The SENT Specimen Study

The study and the validation of this new method have not proven to be an easy task because of the large number of parameters which need to be evaluated. Organize and structure the agenda proved to be a complex point. In this section, using the SENT specimen, it is presented and discussed the results for different aspects of this validation.

4.1.1

Mesh Construction

Bearing in mind the XFEM, it is known in advance, from the bibliographic research [26], that the mesh should be orthonormal around the crack. Moreover, it should have a high density, leading to more accurate results [26]. Thus, considering hexahedral elements in the first instance, a perfectly structured mesh is chosen as first approach (figure 4.1).

Figure 4.1: Perfectly structured mesh.

Although not evident, the mesh represented on figure 4.1, is quite poor. If it was a problem of displacement calculation, it would probably suffice. However, it is intended here to calculate the stress

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intensity factor, which requires a much finer mesh around the crack.

Imposing a better refinement, leading to a sufficiently fine mesh, the solving process becomes quite heavy in time. Running a very fine mesh, appears very difficult for a common user laptop, as the one that was used for this thesis realization. It is then necessary to find a mesh that obey to all the criteria for the analyses and which is not too expensive in computational time.

It was chosen to partition the mesh [6]. In other words, the mesh is defined by blocks, where each block has a proper refinement. However, after preliminary testing for convergence, the mesh defined by blocks may pose some problems. Sometimes, for a given partition scheme, the Abaqus R meshing

algorithms are forced to build unstructured meshes with degenerated elements (figure 4.2), which makes the mesh weak with a consequent loss of accuracy.

Figure 4.2: Unstructured mesh

A structured mesh, with high density around the crack, is really necessary for good results. It should be noted at this point that for the Abaqus R, a structured mesh, is not necessarily a mesh with

orthonor-mal elements. In Abaqus R, a structured mesh, must verify that all the elements have a parallelogram

shape in 2D and a hexahedral shape in 3D.

After consultation of [7] and an optimisation study, the partition scheme presented on figure 4.3 was derived. This partition allows a fine refinement near the crack, and a coarser mesh for the remaining points.

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This partition scheme generated structured meshes in all simulations. It is noted however that when the specimen has its height H or too low or too high, the distorted elements begin to increase. Those elements do not introduce a significant error since they are situated in the far-field.

In addition to this partition scheme, two edges sets were defined: the Global Refinement, which assigns to all the edges in the plane the introduced number of divisions, and the Depth Refinement, which controls the refinement of all the elements responsible of the three-dimensionality. It can be appreciated in figure 4.4 the consequence of these parameters variation.

Figure 4.4: Mesh refinement control: a) Global Refinement of 10 with Depth Refinement of 1 b) Global Refinement of 10 with Depth Refinement of 10

The crack block is the block around the crack. It is always used a square as block, with a characteristic dimension corresponding to the double of the crack length. This block definition has been validated from several analysis, proving to be the one with less oscillations in the majority of the different solutions.

The greatest advantage of all, with this mesh design, it is logically the high density around the crack which is essential, since it is a point of accumulation of stresses, with high gradients, requiring a higher number of nodes. For instance if the Global Refinement is fixed at a value of 20 there will be in the crack block 400 elements in the plane.

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4.1.2

Requested Contours and GRef Influence

As mentioned in section 2, the J-Integral, that allows the stress intensity factor deduction, is calcu-lated by 2D closed contour integrals. In theory, the J-Integral is path independent, but numerically (or computationally) this is not true. Thus, different contours give different solutions for the stress intensity factor.

Abaqus R allows the user to introduce the number of contours to use for the stress intensity factor

calculation, but what number should be introduced? Furthermore, it was observed that the package of requested contours is calculated for each of the generated points in the crack’s front.

The number of generated points, appears to be related to the number of mesh divisions in depth. For structured geometries, if the geometry is a parallelepiped and has i divisions in depth, there will be i + 1 point, which implies i + 1 packages of j contours, each one with a stress intensity factor estimation.

Against this backdrop, it was necessary to design a set of tests that could clarify various doubts. A large sample of analysis was needed in order to be able to understand how to tune the parameters to obtain good results.

Choosing then a SENT geometry, which was set with a height H = 90 mm, width W = 10 mm, crack length a = 2 mm, it was made the study of the computational burden associated to the mesh variations. In terms of thickness, B = 1 mm, and only 2 divisions in depth were considered. For the elements, it was used hexahedral elements with first order accuracy, and reduced integration, since as it is shown later on this thesis, these characteristics are the best choice.

As previously stated, two sets of refinement were defined, Global Refinement and Depth Refinement, referred to from now on as GRef and DRef. In this set of analyses, the first will vary and the second will be fixed in 2 divisions. The partition scheme remains constant, so the increase of GRef , increases the mesh density around the crack.

Another question arises now. What will be the composition of the solution since each contour pro-vides a different value for the stress intensity factor? And which contours should be chosen? All of them or only some?

In all the existing Computed Assisted Engineering (CAE) solutions, regarding the stress intensity factor calculation, the number of contours for the stress intensity factor calculation is asked to the user. In the classical method, 5 contours is always desirable [1]. As the XFEM is still a recent method, it was decided to also study the impact of the requested number of contours.

Hence, considering the aforementioned geometry, with GRef = 80, DRef = 2, and the traction tension unitary, different numbers of contours for the stress intensity factor calculation were requested. The considered range of contours was [5, 10, 15, 20, 25] and for each contour, the error was calculated.

For the errors calculations, it has always been used a relative error expression as the written in equation 4.1, where Kref is the reference stress intensity factor obtained from literature.

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relative=

KI− Kref

Kref

(4.1)

The results can be appreciated on figure 4.5 where the first contour is always neglected.

Figure 4.5: Error evolution for different number of requested contours

It may be said that all the different plot lines oscillate in turn of the same final value. As can also be observed, it is confirmed that each contour gives a different value for the stress intensity factor. It is necessary to define how to calculate a final value. The fairest way, it is to make an average.

It was thought that this average could be a controlled one, by the exclusion of the values that affect to much negatively the solution, but it was decided to consider all the values for the average except the one for the first contour, which is always far from the solution. This decision is related to the fact, that if a random geometry is considered, for which the correlation for the stress intensity factor is unknown, it will be very hard to know what values can be excluded. In order to develop a strategy, it is important to work with the same rules that if a random geometry was the case under evaluation.

Having the average in mind, it was possible to derive table 4.1.

Table 4.1: Impact in the average error of the requested number of contours

Number of Requested Contours [#] Average Error [%]

5 0.46

10 0.53

15 0.42

20 0.42

25 0.54

As the sample of contours growth, the average error becomes lower and a minimum is observed for 15and 20 requested contours. The variations in the error are very small (0.1%) but it was decided to ask 20contours in all further analyses, to observe the error evolution with the contour number.

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More contours could be requested but it was verified that sometimes, for higher number of requested contours, the outer contours may give values for the stress intensity factor completely wrong. This makes sense because if a too high number of contours is requested, the outer ones, will have a radius relatively to the crack’s tip so big, that the solution is no more accurate.

For the data treatment, it is important to refer that each Abaqus R analysis, produces a very long

output file, with all the kind of information, such as the elements and nodes number, the time taken by the analysis, among others, which makes the data treatment time consuming. Thus, it was programmed in C + + a code that takes as input the output file from Abaqus R, and processes all the information,

isolating what is needed, and in the end, writes a new file. This feature proved to be a big jump in this thesis since it has saved hours of data analysis. The treated data were then entered into excel files, allowing the graphs and tables production.

After the study of what should be the number of requested contours, 6 different meshes were considered, consequential to 6 variations in the GRef value, which took one of the following values: [10, 20, 40, 80, 160, 200]. For all these analyses, DRef = 2.

This set of values assigned to the global refinement conducted to the analyses characteristics pre-sented in table 4.2.

Table 4.2: SENT analyses characteristics

GRef [#] Elements [#] Nodes [#] Time [s]

10 1200 4176 3 20 4800 15276 6 40 19200 59076 20 80 76800 233076 44 160 307200 926676 1258 200 480000 1446276 1657

As can be seen in figure 4.6, the time evolution with the parameter GRef is exponential, indicating that the GRef parameter is time sensitive.

From table 4.2, it may be observed that the number of elements and nodes, increase very fast. From GRef = 160, the analyses already has many elements and are time consuming, becoming heavy since the analyses should be expedite for this particular case, which is a simple geometry. To more complex geometries, it is reasonable to expect rather longer runs, associated to heavier meshes.

For each value of GRef it is calculated 3 packs with 20 stress intensity factors. The 3 packages are consequential of the 2 divisions in depth considered for the DRef value. Each one of the 20 stress intensity factors is consequent of a J-Integral calculation, in one of the 20 considered contours.

Table 4.3, shows part of an excel file for a given refinement.

For each GRef value, it is extracted the time1taken by the analysis, the number of nodes as well the

1Abaqus R gives different times for each analysis; the more important are the CPU time and the Wall-clock time. It is important

to note that the chosen reference time for this thesis was the Wall-clock. It corresponds to the human perception of the passage of time from the start to the completion of a task. In this thesis, it was used a processor Intel i7, with 8 logic units and a CPU of 3.6 GHz, associated with a RAM memory of 8 GB

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Figure 4.6: Time evolution with the GRef parameter

Table 4.3: Numerical study, files appearance

number of elements.

For each of the 3 generated points, the 20 stress intensity factors estimations are introduced and the respective error is calculated.

To each point, it is associated an average error. At last, the absolute average error is calculated, corresponding to the average of the three averages errors. In practise, this absolute average error, corresponds to the average error verified for a given GRef value.

It is important to note, that in the case of 2 divisions in depth, which produces 3 points, it was always the centre one who gave the best results. So far, the explanation has not been found, but it is the main reason which support the fact that all important validations in this thesis are done with DRef = 2.

Another interesting point is the evolution of the error, with the number of contour, for the different analyses (figure 4.7). The plotted values are extracted for each analysis from the second point, the more accurate values.

The more eye-catching point is the fact that the second contour produces the biggest error for all GRef values.

Then, the solutions for GRef = 10 and GRef = 20 divisions appear unstable. Their behaviour invalidates them completely. From GRef = 40, all solutions are cleaner.

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Figure 4.7: SENT contour error evolution

and GRef = 200 divisions are already very high. The difference between the solution for GRef = 80 and GRef = 160 is 0.02% (table 4.4), however in computational cost (Wall-clock time), GRef = 160 is 31 times more expensive than GRef = 80 (table 4.2), which makes the analysis of GRef = 80 a reference for subsequent analysis.

Ultimately, the solution is converging, as can be appreciated on table 4.4. In short,the results are very good revealing a very good start for this XFEM evaluation and validation.

Table 4.4: SENT GRef influence results

GRef [#] Absolute Average Error [%]

10 3.16 20 1.92 40 0.91 80 0.48 160 0.50 200 0.51

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4.1.3

Influence of the Ratio a/W

Continuing with the analyses, the new goal is to understand the influence of the a/W ratio on the solution quality.

Keeping the SENT geometry, with H = 90 mm, for a constant width W = 10 mm, the ratio a/W is varied between 0.1 and 0.35, with a step of 0.05.

From the previous analysis, GRef = 80 and DRef = 2. The integration is reduced type, linear interpolation for the shape functions, and 8 nodes hexahedral elements are considered. The results are resumed in table 4.5.

Table 4.5: Ratio a/W influence for GRef =80 divisions

a [m] a/W [#] Absolute Average Error [%]

1 0.10 1.43 1.5 0.15 0.81 2 0.20 0.48 2.5 0.25 0.60 3 0.30 0.97 3.5 0.35 1.49

Figure 4.8: Error behaviour versus a/W ratio

The variations of the error are not greater than 1% (figure 4.8). All solutions are very good considering the fact that GRef = 80, corresponds to a coarse mesh. The error seems to have a parabolic behaviour, with a clear minimum for the ratio of a/W = 0.2. In short, the ratio a/W is not particularly related to the solution accuracy.

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4.1.4

DRef Influence

It was set the ratio a/W = 0.2, with W = 10 mm, W = 90 mm, and keeping GRef = 80, the refinement was varied in depth, DRef , between 1 and 6.

The range is chosen in a so short interval because as can be observed in the results, DRef makes the analyses quickly become very heavy (time speaking).

It is important to recall here that for a structured geometry like this one, with DRef = i, there will be i + 1points for the stress intensity factor calculation. Furthermore, it is always requested 20 contours. The consequence is that if DRef = 1, there are 2 points, and therefore, 40 estimates for the stress intensity factor. Incrementing, if DRef = 2, there are 60 stress intensity factors.

The analyses characteristics are presented in table 4.6 and the errors plotted on figure 4.9.

Table 4.6: SENT DRef influence results

DRef [#] Elements [#] Nodes[#] Time [s]

1 38400 155384 22 2 76800 233076 48 3 115200 310768 74 4 153600 388460 210 5 192000 466152 306 6 230400 543844 747

Figure 4.9: DRef influence

Looking at the results, it can be said that the refinement in depth does not improve the solution. It was decided to consider DRef = 2 for further analyses, which proved to be a good decision, since during this thesis realization, it was investigated the effect of the DRef parameter, and the results were moreover the same, showing that DRef = 2 is always the best choice for the depth refinement.

References

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