Computational Fracture Mechanics

Transcription

1 Computational Fracture Mechanics Students Project Final report Wolfgang Brocks Diya Arafah January 214 Dipartimento di Meccanica Politecnico di Milano report_brocks-arafah, wbrocks, 3 March 214, - 1 -

2 Exercise #1: Determination of a true stress-strain curve for ABAQUS input from tensile test data. Geometry: Material round tensile bar diameter d = 6 mm (nominal) measuring length L = 3 mm 26NiCrMoV115 YOUNG s modulus E = 244 MPa POISSON s ratio ν =.3.2% proof stress R p.2 = 79 MPa (average) 26NiCrMoV11 force [kn] YA-1-26 YA-1-34 YA-1-28 YA-1-3 YA-1-33 YB elongation [mm] Test results by Dip. Meccanica, PoliMi, data in tensile_tests.xls Note that ABAQUS requires true stresses,, in dependence on logarithmic plastic strains,, as input for an elasto-plastic analysis, a σ(ε p ) curve beyond uniform elongation of the tensile bar is needed for the fracture mechanics analysis, which necessitates an extrapolation of the stress-strain curve by a power law where, are normalisation values, commonly taken as yield limit and. report_brocks-arafah, wbrocks, 3 March 214, - 2 -

3 Check the stress-strain curve by simulation of the tensile test beyond uniform elongation (maximum force) and comparing simulation with test data in terms of force, F, vs. elongation, L. Analysis of tensile test axisymmetric model, account for symmetries, displacement controlled, geometrically nonlinear ( large deformations ). Strain and stress measures for large deformations Linear strain Logarithmic (HENCKY) strain!1# $ Nominal stress % & ' 4& )* + True (CAUCHY) stress,-./ & ' Provided the existence of a uniaxial stress state for % % & %12 ' and a uniform elongation, the requirement of isochoric plastic deformation yields ' 4', assuming that the elastic elongation is negligible, true stresses can be calculated as 5,,1 / 4,,1,,-./ & ' 4 & ',-./!1# $ report_brocks-arafah, wbrocks, 3 March 214, - 3 -

4 1.1. Evaluation of test data specimen YA-1-26 YA-1-34 YA-1-28 YA-1-3 YA-1-33 YB-1-28 d [mm] The force elongation curves, F( L), measured in the tests are converted into engineering stress-strain curves, σ %!ε $, as presented in Fig. 1.1, and true stress-strain curves, σ,-./ ε, next, Fig. 1.2, according to the formulas given above. The engineering nominal stress-strain curves reflect the measured load-elongation data. Except YA-1-28, which will be excluded in the following, the curves practically coincide up to maximum load or % %. Scatter of the test results occurs after uniform elongation when necking of the specimens starts. The conversion to true stresses holds only up to R m, as for larger elongations, deformations are not uniform and the stress state is not uniaxial any more. 1 1 σ nom [MPa] YA-1-26 YA-1-34 YA-1-28 YA-1-3 YA-1-33 YB-1-28 σ [MPa] YA-1-26 sig_true sig_nom,5,1,15 ε lin [-],5,1,15 ε [-] Fig. 1.1: Nominal stress vs linear strain (engineering stress-strain curves) Fig. 1.2: Conversion to true stress vs logarithmic strain (specimen YA- 1-26) The curves of true stress vs. logarithmic plastic strain, 5,-./, practically coincide for the five test specimens excluding YA-1-28, on the scale of Fig Whereas the technical definition of yielding by the.2% proof stress has been determined as MPa in average, the physical yield strength, i.e. first occurrence of plastic strain,?, varies between 64 MPa and 67 MPa, Fig σ true [MPa] YA-1-26 YA-1-34 YA-1-3 YA-1-33 YB-1-28 σ true [MPa] YA-1-26 YA-1-34 YA-1-3 YA-1-33 YB-1-28,2,4,6,8 ε p [-] 62,5,1,15,2 ε p [-] Fig. 1.3: True stress vs logarithmic plastic strain up to uniform elongation Fig. 1.4: Determination of the yield strength report_brocks-arafah, wbrocks, 3 March 214, - 4 -

5 A value of 65 MPa is chosen for the following. Extrapolation beyond uniform elongation can be based on any of the stress strain curves of Fig 1.3, as they do not differ much. They cannot be fitted over the whole range of the logarithmic plastic strain by a unique power law σ σ true p = α ε n ε, (1.1) as their curvatures obviously change. Therefore, only the final part is used to define the power law. Normalisation is recommended in order to obtain a dimensionless equation so that α does not depend on the specific dimension used for the stresses. The normalising values can be arbitrarily selected, commonly taken as the yield-limit values, here 65 MPa and.381. Depending on the range of ε p used in the power-law fitting, different values of the parameters α and n are obtained (Figs. 1.5, 1.6). The effect on the extrapolation is negligible, however (Fig. 1.7). 1,38 1,38 σ / σ 1,37 1,36 y = 1,721x,786 σ / σ 1,37 y = 1,112x,698 1,35 1,36 1,34 YA ,33 Pot.(YA-1-26) 1, ε p / ε 1,35 YA-1-26 Pot.(YA-1-26) 1, ε p / ε Fig. 1.5: Power law fit #1 of true stress vs logarithmic plastic strain, see extrapol(1) in Fig. 1.7 Fig. 1.6: Power law fit #2 of true stress vs logarithmic plastic strain, see extrapol(2) in Fig. 1.7 σ true [MPa] YA-1-26 YA-1-34 YB-1-28 extrapol(1) extrapol(2),1,2,3,4 ε p [-] σ [MPa] ABAQUS ,2,4,6,8 1 ε p [-] Fig. 1.7: Extrapolation of true stress vs logarithmic plastic strain curve. Fig. 1.8: True stress vs. logarithmic plastic strain data as input for an elastoplastic analysis with ABAQUS The data points used to describe the elasto-plastic material behaviour in ABQUS, Fig. 1.8, are established as report_brocks-arafah, wbrocks, 3 March 214, - 5 -

6 first point:, second point 5.+.2, 5 points from specimen YA-1-28 up to.39, 1 equidistant points for,6 1. from an extrapolation by power law fit #2 with α = 1.1, n = Numerical simulation of the tensile test An axisymmetric model of the tensile bar is applied, Fig The symmetry to the x 1 -axis is considered to reduce the computational effort; it guarantees that necking occurs at x 2 =, in addition. Therefore only a quarter of the specimen is modelled, which is realised by applying the boundary conditions, E +!F G,$, along this line. The radial displacements, E G!F G,$, have to be free (roller support condition). The specimen is loaded in displacement control, as the axial force decreases under still increasing elongation due to local necking. A prescribed displacement is applied at the upper cross section of the model by connecting all respective nodes to some reference point, Fig. 1.1, using the coupling constraint option in ABAQUS. This point moves in the x 2 - direction with displacement u 2, whereas the boundary condition u 1 = is applied. The corresponding external force is obtained as reaction force at the reference point. Fig. 1.9: Axisymmetric model of tensile bar Fig. 1.1: Application of prescribed displacement 1 Fig. 1.11: Mesh of tensile bar 1 Fig. 1.12: Iso-contours of radial displacement, u 1, showing localisation of deformation after uniform elongation In order to realise a homogeneous stress and strain state over the measuring length, L, up to uniform elongation, the fixation of the specimen with a larger radius at the upper end has 1 courtesy of GIORGIA GOBBI, PhD student, Dipartimento di Meccanica report_brocks-arafah, wbrocks, 3 March 214, - 6 -

7 to be incorporated in the model. Fig For a cylinder of constant radius, necking might occur in the section where the prescribed displacement is applied instead of the centre section. The correct location of necking has to be checked, Figs to Fig. 1.13: Deformed mesh in final load-step displaying necking with isocontours of axial displacement, u 2 Fig. 1.14: 3D visualisation of necking, grey: plastic zone, 5 H.2 Fig 1.14 shows a comparison of the simulation results with the test data. The external force decreases due to local necking of the bar in the mid-section (Figs. 1.12, 1.13), which of course occurs only if a geometrically nonlinear analysis is performed. The corresponding value of uniform elongation is affected by any small imperfection of the bar, e.g. a reduction of diameter by 1% or a mesh refinement in the mid-section. As the bar is a geometrically ideal cylinder in the present model, necking is delayed compared to the tests. force [kn] NiCrMoV11 YA-1-26 YA-1-34 YA-1-33 YB-1-28 ABAQUS Fig. 1.14: Comparison of test data with numerical simulation elongation [mm] The maximum force reached in the simulation is within the scatter of the test results, Tab The coincidence of the FE simulation with the tensile test data finally confirms the stress-strain input data. Specimen YA-1-26 YA-1-34 YA-1-28 YA-1-3 YA-1-33 YB-1-28 FE F max [kn] Table 1.1: Maximum tensile force report_brocks-arafah, wbrocks, 3 March 214, - 7 -

8 Exercise #2: Numerical analysis of a side-grooved C(T) specimen with stationary crack. Geometry: width W = 5 mm height H = 1.2 W = 6 mm initial crack length a = 33.2 mm thickness B = 25 mm net thickness B n = 2 mm (2% side grooved) Material: 26NiCrMoV115 as above for exercise #1 YOUNG s modulus E = 244 MPa POISSON s ratio ν =.3.2% proof stress R p.2 = 79 MPa yield limit F [kn] YYB-19-1 YYB-19-3 YYB-19-4 YYB-19-6 J [kj/m 2 ] YYB-19-1 YYB-19-3 YYB-19-4 YYB-19-6 YYB-19-9 YYB YYB-19-9 YYB ,5 1 1,5 2 2,5 3,5 1 1,5 2 2,5 3 Test results by Dip. Meccanica, PoliMi, data in CT-tests.xls 2 Compare tests with FE simulation in terms of force vs. load-line displacement, F(V LL ), up to approx. maximum force, i.e. I JJ 41.5KK (LM4.5KK) Evaluate J-integral and plot J(V LL ), Investigate path dependence of J. Analysis of C(T) plane strain, account for symmetry, displacement controlled, geometrically nonlinear ( large deformations ). 2 Note that the test results include crack extension, which is not considered in the present FE analysis! J is evaluated according to ASTM E 182. report_brocks-arafah, wbrocks, 3 March 214, - 8 -

9 Mesh: Mesh data in CT_mesh 2.1. FE model The FE model disregards all technical details of the real C(T) specimen, as there are the pinhole for application of the force, the basically stress-free material volume left of the applied force and the machined notch used as starter for fatigue cracking and fitting for the clip gauge measuring V LL in testing. It is thus reduced to a simple rectangle of size W H/2. The mesh is refined towards the ligament, M F G N, reaching a minimum element size of mm 2. The regular arrangement of square elements in the ligament is specific for simulations of crack extension (see exercise #3) but will yield strongly distorted elements at the crack tip and is hence inappropriate for calculating reliable stress and strain values at a stationary crack. Since the present analysis is focused on global values, namely load-line displacement and J-integral, the meshing at the crack tip is uncritical. The FE-model accounts for symmetry with respect to the x-axis, which is realised by fixing all vertical displacements in the ligament, E +!M F G N,F + $. The specimen is loaded in displacement control mode by applying a vertical displacement u 2 to a reference point as in exercise #1, which is connected either just to the node representing the midpoint of the pinhole or to some additional nodes around. The respective force is picked up as reaction force at the reference point. In order to avoid localisation of plastic deformation in the points of load application, the area around should be defined as elastic only (Fig. 2.1). The load-line displacement, V LL, is picked up at the lower left edge of the model, I JJ 2E +!,$, which is slightly different from the clip position in the test. Due to the high stiffness and stresses close to zero in this part of the specimen, the respective effect on the V LL values is negligible, however. The analysis is performed as plane strain with the specimen s net thickness B n. report_brocks-arafah, wbrocks, 3 March 214, - 9 -

10 Fig. 2.1: FE mesh of C(T) (half model) Fig. 2.2: Automatic option of contour definition for J-integral calculation: 1 st contour = crack tip, last contour # J-Integral evaluation The J-integral is defined as an integral over a closed contour, Γ, around the crack tip, ( 2 ij j i,1 ) J = wdx σ n u ds, (2.1) Γ where w is the strain energy density, σ ij the stress tensor and n j the outer normal to the contour. It is equal to the energy release rate in a plane body of thickness, B, due to a crack extension, a, in mode I, U J =, (2.2) B a V L which is exploited in the virtual crack extension (vce) method to calculate J numerically as well as in the experimental determination 3 of J. Applying the divergence theorem, any contour integral can be converted into an area (domain) integral in two dimensions or a volume integral in three dimensions, over a finite domain surrounding the crack. ABAQUS/Standard uses this domain integral method to evaluate the J-integral and automatically finds the elements that form each ring from the regions defined as the crack tip or crack line; commonly focused mesh design is used to provide such rings of elements. Each contour provides an evaluation of the contour integral. The number, n, of contours must be specified in the history output request. *CONTOUR INTEGRAL, CONTOURS=n, TYPE=J By definition, J is a path-independent integral, claiming that each contour should yield the same J-value. The derivation of its path-independence requires (1) time independent processes, (2) absence of volume forces and surface forces on the crack faces, (3) small strains (geometrically linear), (4) homogeneous hyper-elastic material (deformation theory of plasticity), however. A geometrically nonlinear elasto-plastic analysis applying incremental theory of plasticity violates the two last-mentioned conditions, so that the values obtained for the various contours will differ and the J-integral becomes path-dependant, more precisely, J 3 ASTM E 182-6: Standard test method for measurement of fracture toughness, Annual book of ASTM Standards, Vol 3.1, American Society for Testing and Materials, Philadelphia. report_brocks-arafah, wbrocks, 3 March 214, - 1 -

11 increases with increasing size of the contour. The ABAQUS User s Manual provides the following statement 4 to this problem: Domain dependence The J-integral should be independent of the domain used provided that the crack faces are parallel to each other, but J-integral estimates from different rings may vary because of the approximate nature of the finite element solution. Strong variation in these estimates, commonly called domain dependence or contour dependence, typically indicates an error in the contour integral definition. Gradual variation in these estimates may indicate that a finer mesh is needed or, if plasticity is included, that the contour integral domain does not completely include the plastic zone. If the equivalent elastic material is not a good representation of the elastic-plastic material, the contour integrals will be domain independent only if they completely include the plastic zone. Since it is not always possible to include the plastic zone in three dimensions, a finer mesh may be the only solution. If the first contour integral is defined by specifying the nodes at the crack tip, the first few contours may be inaccurate. To check the accuracy of these contours, you can request more contours and determine the value of the contour integral that appears approximately constant from one contour to the next. The contour integral values that are not approximately equal to this constant should be discarded. Most of this is simply rubbish! The domain dependence has physical reasons and is not due to the approximate nature of the finite element solution. The derivation of the pathindependence of the contour integral does not only require that the crack faces are parallel but also the assumptions mentioned above. 5 Contour (or domain) dependence thus is physically significant due to the dissipative character of plastic deformation and not only some error or inaccuracy of the numerical calculation, and not only the first few contours display contour dependence of J. Suggesting a finer mesh at the crack tip is counterproductive as it just increases the number of contours automatically generated by ABAQUS and thus increases the number of contour dependent J-values. The manual is right suggesting you can request more contours and determine the value of the contour integral that appears approximately constant from one contour to the next. What is needed in a fracture mechanics analysis is a saturation value or far-field value of J corresponding to that evaluated from the mechanical work done by the external force at the load-line displacement, eq. (2.2). For further details see BROCKS & SCHEIDER 6. The user must define the crack front and the virtual crack extension in ABAQUS to request contour integral output; this can be done as follows: Defining the crack front The user must specify the crack front, i.e., the region that defines the first contour. ABAQUS/Standard uses this region and one layer of elements surrounding it to compute the first contour integral. An additional layer of elements is used to compute each subsequent contour. The crack front can be equal to the crack tip in two dimensions or it can be a larger region surrounding the crack tip, in which case it must include the crack tip. By default ABAQUS defines the crack tip as the node specified for the crack front in the so called CRACK TIP NODES option of the contour integral command: *CONTOUR INTEGRAL, CONTOURS=n, CRACK TIP NODES Contour Integral Evaluation, ABAQUS User s Manual, Section The violation of condition (4) is apparently addressed in the obscure clause if the equivalent elastic material is not a good representation of the elastic-plastic material W. BROCKS and I. SCHEIDER: Reliable J-values - Numerical aspects of the path-dependence of the J-integral in incremental plasticity, Materialprüfung 45 (23), report_brocks-arafah, wbrocks, 3 March 214,

12 Specify the crack front node set name and the crack tip node number or node set name. Alternatively, a user-defined node set which include the crack tip can be provided to ABAQUS manually by omitting the CRACK TIP NODES option: *CONTOUR INTEGRAL, CONTOURS=n Specify the crack front node set name and the crack tip node number or node set name. Specifying the virtual crack extension direction The direction of virtual crack extension must be specified at each crack tip in two dimensions or at each node along the crack line in three dimensions by specifying either the normal to the crack plane, n, or the virtual crack extension direction, q, see Fig Fig. 2.3: Left: Crack front tangent, t, normal to the crack plane, n, and virtual crack extension direction, q, at a curved crack; Right: Typical focused mesh for fracture mechanics evaluation [ABAQUS user s manual]. The normal to the crack plane, n, can be defined in ABAQUS using the so called (normal) option 7 in the contour integral command: *CONTOUR INTEGRAL, CONTOURS=n, NORMAL n x -direction cosine, n y -direction cosine, n z -direction cosine(or blank), crack front node set name (2-D) or names (3-D) ABAQUS will calculate a virtual crack extension direction, q, that is orthogonal to the crack front tangent, t, and the normal, n, see Fig. 2.3 (left). Alternatively, the (normal) option can be omitted and the virtual crack extension direction, q, can be defined directly: *CONTOUR INTEGRAL, CONTOURS=n crack front node set name, q x -direction cosine, q y -direction cosine, q z -direction cosine (or blank) Note, that the symmetry option (SYMM) must be used to indicate that the crack front is defined on a symmetry plane: *CONTOUR INTEGRAL, CONTOURS=n, SYMM 7 The normal option can only be used if the crack plane is flat report_brocks-arafah, wbrocks, 3 March 214,

13 The change in potential energy calculated from the virtual crack front advance is doubled to compute the correct contour integral values. Different combination of the above mentioned options can be used to request contour integral within ABAQUS, more details about this could be found in the ABAQUS documentation. Fig. 2.4: Manual option of contour definition for J-integral calculation: 1 st contour = node set Fig. 2.5: Manual option of contour definition for J-integral calculation: last contour #12 Two options of defining the contours are used in the present study, The CRACK TIP NODES option and a sufficient number of contours 8 to approach a saturated (or far-field) value of J are chosen, Fig. 2.2, hereafter called automatic option ; A first contour around the crack tip of finite size (omitting the CRACK TIP NODES option) is defined via a crack tip node set, Fig. 2.4, hereafter called manual option. The number of contours is much lower in the latter case. In the present mesh, n = 24 if the automatic option is used and the first contour is the first element ring around the crack tip, Fig The domain left of the crack tip is obviously unnecessarily large. In the manual option, n = 12 to reach the largest possible contour, Fig Simulation Results Fig. 2.6 compares the &!I JJ $ curves of the test specimen YYB-19-3 and the simulation. The simulation results meet the test data up to approximately maximum force, F = 41.6 kn, reached at a load line displacement V LL = 1.36 mm; the respective difference,!& O% &,/O, $/ &,/O,, is less than 3%. The force keeps increasing in the simulation as no crack extension is modelled. Whereas the load decline in the tensile test was due to plastic necking, it is crack extension here which reduces the ligament area. The coincidence between test and simulation results holds much longer for Q!I JJ $, Fig. 2.7, as the opposing effects of decreasing force and increasing a on J test cancel each other. At V LL = 2. mm, the difference!q O% Q,/O, $/Q,/O, is just 1.8%. The curves J-24 and J-12 in Fig. 2.7 denote the far-field J-values obtained from contours #24 (Fig. 2.2) and #12 (Fig.2.5). There is obviously no difference between the two. 8 Note that the last contour must not touch the outer surface of the specimen, as the elements outside this contour are taken for calculating the energy release rate. report_brocks-arafah, wbrocks, 3 March 214,

14 5 8 F [kn] simulation test YYB-19-3 J [kj/m 2 ] J 24 (automatic) J 12 (manual) test YYB Fig. 2.6: Comparison of test data with simulation: force vs. load-line displacement Fig. 2.7: Comparison of test data with simulation: J-integral vs. load-line displacement As described above, the first contours in the automatic option are element rings in the near field around the crack tip, where large plastic strains with corresponding rearrangement of stresses occur. These stress-strain fields violate the requirements of deformation theory of plasticity. In order to obtain meaningful far-field values of J, the corresponding contours should imbed the plastic zone emanating from the crack tip, which is not completely possible, however, as crack initiation occurs under gross yielding where the plastic zone extends from the crack tip to the free surface of the specimen. For V LL = 1 mm and 3 mm, the 11 th and the 16 th J-contour, respectively, are inclosing the plastic zone as far as possible, see Figs. 2.8 and 2.1. Fig. 2.8: Plastic zone at V LL = 1 mm and 11 th J-contour of automatic option Fig. 2.9: Plastic zone at V LL = 1 mm and 3 rd J-contour of manual option In the manual option, the 1 st contour is a user-defined node set including a comparably large area of elements (Fig. 2.4). Therefore, already the 3 rd contour is large enough to embed the plastic zone, see Figs. 2.9 and Fig. 2.1: Plastic zone at V LL = 3 mm and 16 th J-contour of automatic option Fig. 2.11: Plastic zone at V LL = 3 mm and 3 rd J-contour of manual option report_brocks-arafah, wbrocks, 3 March 214,

15 Due to plastic deformations, a significant domain dependence of J occurs close to the crack tip, which increases with increasing V LL, see curves J-1 to J-1 in Fig It needs about 1 to 12 contours in the automatic option to obtain a saturation (far-field) value of J. In the manual option starting from a 1 st contour according to Fig. 2.3, domain independence of J is immediately achieved, Fig This is also demonstrated in Fig. 2.14, where the J-values at V LL = 3. mm are plotted in dependence on the contour number. Note that J must increase with increasing domain size 9., i.e. Q RSTGU HQ RSU, where k is the contour number. J [kj/m 2 ] "automatic" J 24 J 12 J 1 J 8 J 7 J 6 J 5 J 4 J 3 J [kj/m 2 ] "manual" J 12 J 9 J 7 J 5 J 3 J 1 (node set),5 1 1,5 2 2,5 3 J 2 J 1 (crack tip),5 1 1,5 2 2,5 3 Fig. 2.12: Contour dependence of J: automatic option, 1 st contour = crack tip Fig. 2.13: Contour dependence of J: manual option, 1 st contour = node set J [kj/m 2 ] automatic manual J [kj/m 2 ] J_ASTM J 24 (automatic) J 12 (manual) 2 2 V LL = 3. mm contour,5 1 1,5 2 2,5 3 Fig. 2.14: Contour dependence of J Fig. 2.15: J from far-field domain integral and ASTM 182 The J-values calculated by a contour (or domain) integral, eq. (2.1), in the far-field should equal the energy release rate calculated from the work of external forces, eq. (2.2) as it is done in the tests based on ASTM E The basic equations for the latter are as follows. J is split into an elastic and a plastic part, where the elastic part is calculated from the stress intensity factor, K, and the plastic part from the area under the &!I JJ $ curve: K ( 1 ν ) 2 2 Ap = e + p = + η, (2.3) E Bnb J J J K F = f BB W n a W, (2.4) 9 W. BROCKS and H. YUAN: On the J-integral concept for elastic-plastic crack extension, Nuclear Engineering and Design 131 (1991), report_brocks-arafah, wbrocks, 3 March 214,

16 , (2.5) A = F dv FV = F dv C F 1 e 1 2 p LL 2 LL LL 2 LL η = b, (2.6) W where a is the initial crack length, B n the specimen net thickness 1 and b the uncracked ligament, V NM, W the width of the specimen, WX Y Z \ is a geometry function and C LL [ the elastic compliance, both given in ASTM E 182 as polynomials of!m N$ for C(T) specimens. Fig. 2.16: Definition of area A p for J calculation. Fig 2.15 shows that the far-field values of the contour integral and the J values calculated from the F(V LL ) data indeed coincide. If analytical formulas are available as in the case of fracture mechanics specimens, this comparison is a convenient method to check either the calculated contour integral values or the analytical formula 11. The above equations (2.3) to (2.6) hold for constant crack length, a, i.e. a stationary crack. Since crack length, MM # M, and ligament width, VNM, change with crack extension, a, they must be continuously updated during the J evaluation. An incremental procedure for calculating J( a) curves is given in ASTM 182: J η A A a a. (2.7) p p p p ( i) ( i) ( i 1) ( i) ( i 1) ( i) = J( i 1) + 1 γ ( i 1) b( i) Bn b( i 1) 1 11 Note that in the plane-strain analysis B = B n = 2 mm W. BROCKS, P. ANUSCHEWSKI and I. SCHEIDER: Ductile tearing resistance of metal sheets. Engineering Failure Analysis 17 (21), report_brocks-arafah, wbrocks, 3 March 214,

17 Exercise #3: Numerical analysis of a side-grooved C(T) specimen accounting for crack growth by applying cohesive elements Geometry: width W = 5 mm initial crack length thickness net thickness a = 33.2 mm B = 25 mm B n = 2 mm (2% side grooved) Material: 26NiCrMoV115 as above in exercise #1 YOUNG s modulus POISSON s ratio ν =.3 E = 244 MPa.2% proof stress R p.2 = 79 MPa yield limit F [kn] , YYB-19-1 YYB-19-3 YYB-19-4 YYB-19-6 YYB-19-9 YYB ,5 1,5 YYB-19-1 YYB-19-3 YYB-19-4 YYB-19-6 YYB-19-9 YYB-19-1,5 1 1,5 2 2,5 3,5 1 1,5 2 2,5 3 a [mm] Test results by Dip. Meccanica, PoliMi,data in CT-tests.xls Compare tests with FE simulations in terms of force vs. load-line displacement, F(V LL ), and load-line displacement vs. crack extension, V LL ( a) Investigate the effect of cohesive parameters σ c, Γ c Analysis of C(T) plane strain, account for symmetry, displacement controlled, geometrically nonlinear ( large deformations ), cohesive elements (UEL by SCHEIDER [21]) report_brocks-arafah, wbrocks, 3 March 214,

18 Cohesive law (traction-separation law) For details of the implementation as UEL in ABAQUS and description of input see The Cohesive Model Foundation and Implementation Report by INGO SCHEIDER, Release 2..3, Helmholtz Centre Geesthacht, 211 czm-doku-v2..3 Suggested cohesive parameters: (1) (2) (3) cohesive strength σ c MPa critical separation δ c mm separation energy Γ 12 c kj/m shape parameter δ 1 /δ c shape parameter δ 2 /δ c Mesh: ABAQUS input file CT_input 12 regard the relation ]^ _`abc b XGd` f_ Tf` efb fb \ report_brocks-arafah, wbrocks, 3 March 214,

19 3.1. FE model The FE model is basically the same as in exercise #2, see Figs. 2.1 and 2.2. Instead of fixing the vertical displacements in the ligament, M F G N,F +, cohesive elements 13 are applied in the ligament to allow for crack extension. Symmetry 14 is realised by applying linear constraint equations, E +!d$ E+!T$, Fig. 3.1, which guarantee a symmetric opening of the cohesive element, g E +!T$ E+!d$ 2E+!T$, implying delta_n = δc. These constraint equations induce extra nodal forces 15, however, which reduce the stresses in the respective element to one half, so that the cohesive strength traction_n = σ c /2. constraint equations E +!d$ E+!T$ E G!d$ E G!T$ cohesive parameters delta_n = δ c, traction_n = σ c /2 Fig. 3.1: Cohesive elements at symmetry line: Constraint equations and input variables The variables of user elements cannot be displayed by ABAQUS/Viewer. They are hence transferred to the adjacent continuum elements as user output variables by providing an element mapping between cohesive and continuum elements which is given in the.cohes file, *ELEMENT MAP cont_elem_1, coh_elem_1 cont_elem_2, coh_elem_2 cont_elem_n, coh_elem_n The cohesive law of SCHEIDER 13 is applied δn δn ( δ ) ( δ ) 2 2 for δ 1 1 n δ1 σ ( δ ) = σ 1 for δ < δ δ, (3.1) n n c 1 n δn δ2 δn δ ( δ ) 3 ( ) for c δ2 δc δ δ 2 2 δn δc I. SCHEIDER: The cohesive model Foundation and implementation, Release 2..3, Report, Helmholtz Centre Geesthacht, 211. W. BROCKS, D. ARAFAH, M. MADIA: Exploiting symmetries of FE models and application to cohesive elements, Report Milano / Kiel, November Linear constraint equations introduce constraint forces at all degrees of freedom appearing in the equations. These forces are considered external, but they are not included in reaction force output. Therefore, the totals provided at the end of the reaction force output tables may reflect an incomplete measure of global equilibrium. - ABAQUS Analysis User's Manual (6.11): Linear constraint equations. report_brocks-arafah, wbrocks, 3 March 214,

20 σ n /R [-] (1) (2) (3) 2 Fig. 3.2: Traction-separation laws for the three suggested parameter sets (R = 65 MPa) 1,2,4,6,8,1 δ n [mm] with the cohesive energy δc 1 2 δ1 δ 2 c = n ( n ) d n = c c δ c δc Γ σ δ δ σ δ. (3.2) The J-integral is calculated as described above in Exercise #2 for the contour #12 of the manual option, Fig Simulation Results - Comparison with Test Data. F [kn] , test YYB-19-3 FE param set (1) FE param set (2) FE param set (3) 2 1,5 1,5 test YYB-19-3 FE param set (1) FE param set (2) FE param set (3),5 1 1,5 2 2,5 3,5 1 1,5 2 2,5 3 a [mm] Fig. 3.3: Force vs. load-line displacement Fig. 3.4: Load-line displacement vs. crack extension Parameter set (1) results in a reasonably good coincidence of test and FE results in terms of force vs. load-line displacement, F(V LL ), Fig. 3.3; the simulation results for V LL ( a) do not meet the test data at all, however, Fig Initiation of crack extension occurs much too late, which is due to the large values of the critical separation, δ c, see Fig. 3.2, and hence the separation energy, Γ c. The graphs strongly indicate that a comparison of test and simulation data solely in terms of global quantities like F(V LL ) can result in wrong conclusions with respect to the quality of the parameters. Parameter sets (2) and (3) with a ten times smaller δ c and a nearly six times smaller Γ c 16 yield a reasonable estimate for the initiation of crack extension but underestimate the maximum force. The initial curvature of V LL ( a) is better matched by parameter set (1). Increasing the cohesive strength, σ c, obviously results in a higher slope of the V LL ( a) curve compared to that of parameter set (1). There is little difference between the parameter sets 16 regard the relation between δ c, Γ c and σ c according to eq. (3.2) report_brocks-arafah, wbrocks, 3 March 214, - 2 -

21 (2) and (3) as one might expect looking on Fig 3.2 and no conclusion is possible, which one is the better. As both the cohesive strength, σ c, and the shape parameter, δ 2, are varied, it is difficult to uniquely separate the respective effects on the simulation results. Firm conclusions on the sensitivity of the simulation results to parameter variations require more systematic studies varying no more than one parameter at each time, see chapter 3.3. Risks with missing convergence may show up, however. The J R -curve, Fig. 3.5, provides basically the same information as the plot V LL ( a), Fig. 3.4, since the relation between J and the load-line displacement is linear over a wide range, see Fig It allows for relating the separation energy to a macroscopic fracture parameter, namely the J-value at initiation of crack extension, J i. Parameter set (1) with Γ c = 137 kj/m 2 yields J i = 215 kj/m 2, whereas parameter sets (2) and (3) with a six times lower Γ c = kj/m 2 result in a value of J i = 36 kj/m 2, which is also six times lower. The J R - curves for the parameter sets (2) and (3) are nearly straight lines after initiation. The plot J(V LL ) in Fig. 3.6 shows nearly no sensitivity to the cohesive parameters and hence does not provide any significant information. From Fig. 2.7 which has been obtained for constant crack length no substantial effect of crack extension could be expected in any case. J [kj/m 2 ] J [kj/m 2 ] test YYB-19-3 FE param set (1) FE param set (2) FE param set (3) 2 test YYB-19-3 FE param set (1) 1 FE param set (2) FE param set (3),5 1 1,5 2 2,5 3 a [mm] Fig. 3.5: J R -curves 2 1,5 1 1,5 2 2,5 3 Fig. 3.6: J-integral vs. load-line displacement A systematic trial and error procedure based on an understanding of the effect of individual parameter variations on the mechanical response of the structure helps finding optimised parameters. Based on the sensitivity study described in the following section, a new parameter set (4) has been identified which yields a reasonable approximation of the test data for both global and local quantities, see Figs F [kn] ,5 3 2 YYB-19-3 FE par-set (4) 2 1,5 1 YYB-19-3 FE par-set (4) 1,5,5 1 1,5 2 2,5 3 Fig. 3.7: Force vs. load-line displacement,5 1 1,5 2 2,5 3 a[mm] Fig. 3.8: Load-line displacement vs. crack extension report_brocks-arafah, wbrocks, 3 March 214,

22 J [kj/m2] YYB-19-3 FE par-set (4) parameter set (4): σ c = 213 MPa, δ c =.5 mm, Γ c = 94.4 kj/m 2, δ 1 /δ c =.4, δ 1 /δ c =.8 1,5 1 1,5 2 2,5 3 a [mm] Fig. 3.9: J R -curves 3.3 Sensitivity Study for Cohesive Parameters The following graphs 17 show the effects of a systematic variation of the four parameters cohesive strength, σ c, critical separation, δ c, and shape parameters, δ 1 /δ c, δ 2 /δ c, on the global response, F(V LL ), and the R-curve in terms of V LL ( a). One parameter is varied at a time, whereas the three others are kept constant. The respective separation energy, Γ c, results from eq. (3.2). The sensitivity with respect to the cohesive parameters σ c and δ c is considerably more pronounced than that to the shape parameters δ 1 and δ Cohesive Strength The cohesive strength is the maximum stress which the cohesive element can endure. It thus determines the maximum load which a structure can bear and the R-curve, see Fig If σ c is low, e.g. σ c = 15 MPa 2.3 R, fracture occurs quasi-brittle, the R-curve is nearly a constant and the specimen will fail unstable in small-scale yielding. If σ c is high, e.g. σ c = 3 MPa 4.6 R, the stresses in the ligament are limited by the yield condition and hardly reach σ c so that nearly no crack extension occurs. The respective load-displacement curve of the specimen approaches that of Fig The parameter study suggests a range of 3 R < σ c < 4 R for the present (plane strain) problem. F [kn] YYB-19-1 sigma_c=15 2 sigma_c=2 sigma_c=21 sigma_c=213 1 sigma_c=3,5 1 1,5 2 2, ,5 2 1,5 1,5 YYB-19-1 sigma_c=15 sigma_c=2 sigma_c=21 sigma_c=213 sigma_c=3,5 1 1,5 2 2,5 a [mm] Fig. 3.1: Effect of σ c [MPa], δ c =.1 mm, δ 1 /δ c =.2, δ 2 /δ c =.3 17 The study presented in this paragraph has been performed by CAN IÇÖS, PhD student, Dipartimento di Meccanica, Note that the stress-strain curve input for ABAQUS is derived from the same tensile test data but is not identical to that shown in Fig report_brocks-arafah, wbrocks, 3 March 214,

23 3.3.2 Critical separation The critical displacement at total local failure governs the ductility of the separation process. The effects of the two parameters σ c and δ c interact, however, via the separation energy, Γ c ~ σ c δ c, according to eq. (3.2). Γ c controls the initiation value of J as well as the slope of the R-curve. J i can hence be taken as first approximation for Γ c. If the slope of the R-curve is too high, F(V LL ) approaches Fig. 2.6 again. In the present parameter study, σ c = 213 MPa is kept unchanged, so that together with the chosen shape parameters, Γ c.68 σ c δ c = 145 MPa δ c. The range of.1 mm δ c.1 mm in Fig thus corresponds to 14.5 kj/m 2 Γ c 145 kj/m 2. F [kn] , YYB-19-1 delta_c=.1 delta_c=.2 delta_c=.5 delta_c=.1 1,5 1,5 YYB-19-1 delta_c=.1 delta_c=.2 delta_c=.5 delta_c=.1,5 1 1,5 2 2,5 3,5 1 1,5 2 2,5 a [mm] Fig. 3.11: Effect of δ c [mm], σ c = 213 MPa, δ 1 /δ c =.2, δ 2 /δ c = Shape parameter δ 1 The sensitivity of the simulation with respect to δ 1 is minor but perceivable, see Fig However varying the shape parameter δ 1 should not be part of the parameter identification. It determines the initial elastic stiffness, K coh, of the cohesive element, which is inevitable for the numerical stability of the solution but should be significantly larger than the elastic stiffness of the continuum elements in the ligament in order to avoid numerical artefacts, 18 K resulting in dσ 2σ E n c coh = = >, (3.3) dδn δ1 h δ elem n 2σ c δ 1 < helem =.25 mm (3.4) E for σ c = 213 MPa and h elem =.125 mm. All values of δ 1 /δ c in Fig meet this condition. 18 M. ELICES, G.V. GUINEA, J. GÓMEZ AND J. PLANAS: The cohesive zone model: advantages, limitations and challenges. Eng. Fract. Mech. 69 (22), report_brocks-arafah, wbrocks, 3 March 214,

24 5 3 F [kn] 4 2, YYB-19-1 d1=.15 d1=.2 d1=.25 d1=.5 d1=.1 1,5 1,5 YYB-19-1 d1=.15 d1=.2 d1=.25 d1=.5 d1=.1,5 1 1,5 2 2,5 3,5 1 1,5 2 2,5 a [mm] Fig. 3.12: Effect of d 1 = δ 1 /δ c, σ c = 213 MPa, δ c =.1 mm, δ 2 /δ c = Shape parameter δ 2 The shape parameter δ 2 affects the value of Γ c for fixed σ c and δ c according to eq. (3.2). It can hence be used for a fine tuning of the slope of the R-curve, see Fig F [kn] ,5 2 3 YYB ,5 2 1 d2=.2 d2=.3 d2=.5 d2=.7 1,5 YYB-19-1 d2=.2 d2=.3 d2=.5 d2=.7,5 1 1,5 2 2,5 3,5 1 1,5 2 2,5 a [mm] Fig. 3.12: Effect of d 2 = δ 2 /δ c, σ c = 213 MPa, δ c =.1 mm, δ 1 /δ c =.2 Epilogue and Acknowledgement The report presents the results of an exercise for the students of a PhD course on Computational Fracture Mechanics held at the Dipartimento di Meccanica of the Politecnico di Milano in October and November 213. Beyond the direct results of the assignment which have been used as benchmark for the evaluation of the students reports, additional information and results are provided which have not been part of the definition of the project but may serve as background explanations, particularly the comments on the ABAQUS theory manual with respect to J calculation and the detailed sensitivity study for the cohesive parameters. The authors gratefully acknowledge the contributions of GIORGIA GOBBI (Figs. 1.1, 1.11, 1.13), CAN IÇÖS, (Figs ), PhD students, and STEFANO FOLETTI (test data), professor at the Dipartimento di Meccanica. The report can hopefully also give guidance to engineers and scientists working in the field of numerical simulations of fracture problems how to achieve reliable results. report_brocks-arafah, wbrocks, 3 March 214,