Abstract. Adhesion of thin film multilayers deposited on glass is a crucial issue for many industrial applications. Thus, it becomes of great interest to measure and also to increase the adhesion. Many mechanisms of toughening a brittle solid can be found in literature; but few of them can be applied to thin film layer. By introducing a heterogeneous interfacial toughness field, it should be possible to increase adhesion. This toughness modification would be the consequence of the existence of a pinning regime due to a local change of the toughness. To experimentally validate this new approach of adhesion modification, we investigate the **crack** **front** pinning by performing cleavage tests on multilayer coated samples with a heterogeneous interfacial toughness. We have tested different patterns of pinning region. The **crack** **front** morphology was nicely described in the framework of the perturbative approach initially developed by Gao and Rice and allowed us to determine the local value of the energy release rate (~adhesion).

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This paper investigates an approach for calculating the cyclic J-Integral through a new industrial application. A previously proposed method is investigated further with the extension of this technique through a new application of a practical 3D notched component containing a semi-elliptical surface **crack**. Current methods of calculating the cyclic J-Integral are identified and their limitations discussed. A modified monotonic loading concept is adapted to calculate the cyclic J-integral of this 3D Semi Elliptical Surface **Crack** under cyclic loading conditions. Both the finite element method (FEM) and the Extended Finite Element Method (XFEM) are discussed as possible methods of calculating the cyclic J-Integral in this investigation. Different loading conditions including uniaxial tension and out of plane shear are applied, and the relationships between the applied loads and the cyclic J-integral are established. In addition, the variations of the cyclic J-integral along the **crack** **front** are investigated. This allows the critical load that can be applied before **crack** propagation occurs to be determined as well as the identification of the critical **crack** direction once propagation does occur.

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The behavior of a surface **crack** with straight **front** in a round bar under tension cyclic loading has been analyzed. The effects of supposed cracks on the fatigue behavior of round bar were simulated as well. The finite element method is employed in the present simulation technique because of its versatility and generality for complicated cracked structures. Also, the 1/4-point displacement method was used for evaluating the stress intensity factor by strain/stress singularity at the corner of a 20-node isoparametric element. It is noted that a few parameters, namely the initial **crack** aspect ratio has an influence on the **crack** **front** evolution, provide that the **crack** geome- try is represented by relative dimensions with respect to the bar diameter. Results were shown, under pure cyclic tension loading, it can be seen that the **crack** propagation paths differ with diverse initial flaw depths, but con- verge to the same configuration when the **crack** depth ratio b/D is larger than about 0.5. The functions of aspect ratio and relative **crack** depth are obtained and by means of the function, the **crack** **front** shape and **crack** growth rate can be predicted well. This observation is obtained by means of the computational model. By using extra- polate method equations were offered for K 1C and **crack** growth rate estimations. Based on these equations, the

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The material investigated is a German 22 NiMoCr 37 (A 508 Class 2) rector pressure vessel steel. The fracture toughness of the material is determined using a variety of fracture mechanics specimens including Charpy sized SE(B)10x10 specimens with three different **crack** depths (a/w ≈ 0.51, a/w ≈ 0.18 and a/w ≈ 0.13), C(T) specimens with different sizes (C(T)25 and C(T)50, both with a/w ≈ 0.51) as well as center cracked CC(T)100 specimens with 2a/w ≈ 0.51. All specimens are tested displacement controlled under quasi static loading. The test temperatures are varied between T = -120°C and T = 0°C. The fracture toughness is determined according to ASTM Standard E 1921-03 (2003). For the shallow cracked SE(B)10x10 specimens, the plastic correction factor in the approximate formulae given in ASTM E1921 is adjusted such that the obtained approximation for the average J-integral matches with the value obtained from a finite element analysis of the test (see Section 3) using its definition as a path independent integral around the **crack** **front**. The fracture toughness for the CC(T) specimens is obtained in a similar manner. All toughness results are normalized to a **crack** **front** length of B = 25 mm as required in ASTM E 1921 (2003).

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In order to calculate fracture mechanics parameters of the **crack** in the specific position of the reactor pressure vessel, the **crack** model is established by the finite element method. Due to the existence of the singular field at the **crack** tip, the fine meshing is the prerequisite for obtaining accurate stress intensity factor. In this paper, boundary release and mesh relaxation techniques are used to ensure the mesh quality of the **crack** **front**. The example of boundary release and mesh relaxation is shown in Figure 2.

In brittle solids, stress concentration at **crack** tips makes the macroscopic fracture properties very sensitive to heterogeneities at the scale of the micro structure. So, the macroscopic resistance of a solid depends strongly on the resistance fluctuations at the microscopic scale. To describe quantitatively this phenomenon in disordered materials, we model first the behavior of the **crack** by a stochastic equation of motion taking into account the role of the microstructure. Our approach is first validated by comparing our theoretical predictions with recent experimental observations made on the dynamics and morphology of a **crack** **front**. We show how to use this approach to determine the e ﬀ ective resistance of a brittle material from the characteristics of its micro structure.

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At VTT, development work has been in progress for 15 years to develop and validate testing and analysis methods applicable for fracture resistance determination from small material samples. The VTT approach is a holistic approach by which to determine static, dynamic and **crack** arrest fracture toughness properties either directly or by correlations from small material samples. The development work has evolved a testing standard for fracture toughness testing in the transition region. The standard, known as the Master Curve standard is in a way ”first of a kind”, since it includes guidelines on how to properly treat the test data for use in structural integrity assessment. No standard, so far, has done this. The standard is based on the VTT approach, but presently, the VTT approach goes beyond the standard. Key components in the standard are statistical expressions for describing the data scatter, and for predicting a specimen’s size (**crack** **front** length) effect and an expression (Master Curve) for the fracture toughness temperature dependence. The standard and the approach it is based upon can be considered to represent the state of the art of small specimen fracture toughness characterization. Normally, the Master Curve parameters are determined using test specimens with "straight" **crack** fronts and comparatively uniform stress state along the **crack** **front**. This enables the use of a single K I

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node, whereas the axial force is directly applied in the direction-x on the cross-sectional area of the bar. At the other end, the component is constrained appropriately. In order to obtain a suitable finite element model, it is necessary to compare the proposed model with other published models [11, 16, 17]. In this work SIFs results are used for the validation purposes. Since, it is hard to find the result of J-integral results for these particular Abstract: This paper numerically discusses the role of J-integral along the surface **crack** **front** in cylindrical bar under combined mode I loading. It is also verified the analytical model derived from the first part of this paper by comparing the results obtained numerically using ANSYS finite element program. It is found that the proposed model capable to predict the J-integral successfully along the **crack** **front** but not for the area away from the deepest **crack** depth. This is probably due to the fact that the problem of singularity.

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Abstract. A new remote nondestructive inspection technique, based on thermoelastic temperature measurement by infrared thermography, is developed for detection and evaluation of fatigue cracks propagating from welded joints in steel bridges. Fatigue cracks are detected from localized high thermoelastic temperature change at **crack** tips due to stress singularity under variable loading from traffics on the bridge. Self- reference lock-in data processing technique is developed for the improvement of signal/noise ratio in the **crack** detection process. The technique makes it possible to perform correlation processing without an external reference signal. It is very difficult to detect through-deck type fatigue cracks in steel decks by the conventional NDT technique, since they are not open to the inspection. In this paper, self-reference lock-in thermography is applied for detection of through-deck type fatigue cracks. Experiments are carried out to steel deck sample, which simulates an actual steel bridge, during **crack** propagation test. It is found that significant stress concentration zone can be observed near the **crack** **front**, which enabled us to detect through-deck type fatigue cracks and to estimate its size.

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* CEDI(i) Coefficients for each of the 6 stress-work density terms in the* * calculation of the equivalent domain integral. * * DSHAP(i,j) Array for the derivative of the i-th shape function w.r.t. the * * local coordinate j. j=1:x1; j=2:x2; j=3:x3. * * DSRST(i,j) Array for the derivatives of the s-function w.r.t. the natural * * coordinates j at the i-th integration point. j=1:xi; j=2:eta; * * j=3:zeta. * * DSX(i,j) Array for the derivatives of the s-function w.r.t. the local * * coordinates j at the i-th integration point. j=1:x1; j=2:x2; * * j=3:x3. * * DUX1(i,j) Array for the derivatives of the j-th displacement component * * w.r.t. the local x1 coordinate at the i-th integration point. * * j=1:u1; j=2:u2; j=3:u3. * * KE The element number in the ring. * * KR The ring number. * * KS The element layer(segment) number along the **crack** **front**. * * RX(i,j) Inverse of the Jacobian. i=1..3; j=1..3 * * SF(i) s-function at the i-th integration point, i=1..8. * * SIGMA(i,j) 3x3 array for the stress tensor of the element. * * TERMEDI(i,j) Array for the terms in the expression of the equivalent * * domain integral. i=1: 1st term; i=2: 2nd term. * * TERMI2(i,j) Array for the terms in the expression of the interaction * * integral I(2). i=1: 1st term; i=2: 2nd term. * * TERMII(i,j) Array for the terms in the expression of the interaction * * integral I(1). i=1: 1st term; i=2: 2nd term

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Abstract: In this paper, the ductile fracture mechanism is discussed. The results of the numerical and experimental analyses are used to estimate of the onset of the **crack** **front** growth . It is assumed that the ductile fracture in **front** of the **crack** starts at the location along the **crack** **front** where the accumulated effective plastic strain reaches a critical value. It is also assumed that the critical effective plastic strain depends on the stress triaxiality and the Lode angle. The experimental programme was performed using five different specimen geometries, three different materials and three different temperatures of +20°C, -20°C and -50°C. Using the experimental data and the results of the finite element computations, the critical effective plastic strains are determined for each material and each temperature. However, before the critical effective plastic strain is determined, a careful calibration of the stress–strain curves was performed after modification of the Bai– Wierzbicki procedure. Finally, by analysing the experimental results recorded during the interrupted fracture tests and scanning microscopy observations, the research hypothesis is verified.

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This thesis deals with issues related to the experimental determination of the fracture toughness resistance curves, i.e. the J-integral (J)-R curve and **crack** tip opening displacement (CTOD)-R curves, using the single-edge bend (SE(B)) and single-edge tension (SE(T)) specimens. First, the impact of the **crack** **front** curvature on the J-R curve measured from the SE(B) specimen is investigated through systematic linear-elastic and elastic-plastic three-dimensional (3D) finite element analyses (FEA) of SE(B) specimens containing both straight and curved **crack** fronts. Three average relative **crack** lengths are considered, namely 0.3, 0.5 and 0.7, and three specimen width-to-thickness ratios are considered: 0.25, 0.5 and 1. The curved **crack** fronts are characterized by a power-law expression. The analysis results suggest that the **crack** length evaluated from the CMOD compliance of the SE(B) specimen is insensitive to the **crack** **front** curvature and that the impact of the **crack** **front** curvature on the experimentally-evaluated J values varies with the specimen configurations. For a given specimen configuration, as the **crack** **front** curvature increases, the value of J evaluated based on the test standard ASTM E1820-11 without considering the **crack** **front** curvature becomes less conservative and tends to overestimate the actual J. New **crack** **front** straightness criteria that are in most cases less stringent than ASTM E1820-11, are recommended. The accuracy of the double clip-on gauge method for experimentally determining CTOD is examined through systematic 3D elastic-plastic large-strain FEA of clamped SE(T) specimens. The relative **crack** lengths of the specimens range from 0.3 to 0.7, and the thickness-to-width ratios are 0.5, 1 and 2. It is observed that the CTOD values determined from the double clip-on gauge method may involve significant errors. This error primarily depends on the **crack** length, the material property and the loading level. Based on the analysis results, a modified CTOD evaluation equation is developed to improve the accuracy of CTOD evaluated using the double-clip on gauge method.

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The specimens were manufactured from artificially aged material of decommissioned (not operated) reactor pressure vessel of WWER 440 Type NPP Nord. Each of the specimens contained through-thickness **crack** embedded in the base material (with approx. 3 mm ligament separating the upper **crack** **front** from cladding), and was loaded by 4PB loading at room temperature. During loading, majority of specimens exhibited pop-ins followed by ductile tearing of cladding and final failure; only 3 specimens fractured through suddenly, without preceding pop-ins. Evaluation of the experiments concentrated on both process of cleavage fracture in base material and process of ductile tearing in cladding.

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fatigue tests were conducted on a closed-loop, servo-electric machine with a water circulation loop under load control and at a frequency of 0.02 Hz. The constant load amplitude was periodically reduced to maintain a constant stress intensity factor range to characterize the fatigue **crack** growth rates in the weld, interface and base metal. The load amplitude was set at an R ratio of 0.2 by an inner load cell control, which deducted the friction force between the pulling rod and the sealing material. The external load cell measurements including the friction force were also monitored for a comparison with the ones taken by the inner load cell. The ﬁltered air was injected into the water tank to maintain a saturated oxygen level in the water environment. The conditions of the water environment are summarized in Table 4. The **crack** length was measured by an ACPD technique. The AC current leads and PD probes were made of SS 316L wires 0.5 mm in diameter. The wires were spot-welded to the specimens. ACPD signals were measured by a Matelect model CGM5R **crack** growth monitor. In the present study, an alternating current of 230 mA was applied to the specimens at a frequency of 3 kHz. An ampliﬁcation gain of 70 dB was selected for all tests. 6) The ﬁnal fatigue **crack** length measurement was further calibrated against the average value of ﬁve measure- ments taken along the **crack** **front** on the fracture surface by a microscope at a magniﬁcation of 20 according to ASTM E 647.

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As a **crack** model, this proposed model has adopted the physical characteristic of **crack**. It is similar to the idea in other existing researches in **crack** modeling. In this proposed model, the higher value of **crack** initial energy makes the **crack** grows larger. In the other hand, the **crack** resistance factor reduces the **crack** growth probability [3]. The energy reduction variable has the same function with the energy release rate terminology that is used in the existing research [3]. The **crack** segment length adopts stochastic approach while in some existing research, **crack** segment length adopted deterministic approach [3]. The **crack** propagation direction uses continuous approach and it is similar to other existing researches [3,4]. In this proposed model, material surface can be viewed as solid material while other existing research viewed materials as a set of disjoint elements [19].

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From the literature study we found that although geometric imperfection plays dominant role in load carrying capacity, but there are few studies conducted on imperfect cones with **crack** introduced. This experimental project work aims at examining the effect of **crack** orientation (i.e., axial **crack**, circumferential **crack** and angle **crack**) with different **crack** length on the load carrying capacity of conical shells subjected to axial compression.

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PLATE MODEL: The plate model as shown in Figure 2and3 is modelled using ANSYS with two holes at the centre of the place. The plate is found to have symmetry along the centre of the hole and the **crack** is modelled in the horizontal plane at the ends of the hole. The symmetry condition means that only one half of the plate is modelled to introduce the **crack** and find the stress intensity factor for the whole model. The visibility of the **crack** at the centre of the model is more in the symmetric modelling of the plate. The plate shown in Figure 2 has two clamped portions on either ends of the model, these portions will not be modelled in ANSYS as it is used to clamp the plate while experimental analysis are carried out.The dimensions of the plate are determined by the following relationship. r/t = 0.075, 0.1, 0.2, 0.333, 0.5, 1.0, 2.0, 3.0, 6.0

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An analytical and experimental approach by H. Nahvi and M. Jabbari et al. to the **crack** detection in cantilever beams by vibration analysis. Sensibility analysis of the inverse problem of the **crack** parameters (location and depth) determined by M. B. Rosales, C P Filipich and F S Buezas et al. An efficient numerical technique is necessary to obtain significant results.

1-D and 2-D reactive transport simulations were conducted to determine the degree of concrete dissolution and pH change that may occur as boric acid solution diffuses into the matrix or flows in the **crack** of a reinforced concrete structure. The depth of concrete leaching by boric acid solution derived from the 1-D model agree relatively well with the measured leaching depths in concrete specimens that were immersed in boric acid solutions for up to 300 days. The 1-D simulation results indicate that leaching by boric acid solution diffusing into concrete is mitigated by the acid- neutralizing capacity of the cement minerals such that reinforcement steel with a 5.1-cm [2-in] concrete cover is unlikely to undergo corrosion for at least 77 years. The 2-D simulation results indicate that concrete provides significant chemical reactivity to neutralize the acidic pH of borated water that may leak from an SFP, flow into a **crack** in the SFP concrete structure, and diffuse into the concrete matrix. However, reinforcement steel close to the **crack** inlet, whether exposed in the **crack** or covered by concrete, can become susceptible to corrosion depending on the **crack** aperture, solution flow rate and duration, and boric acid concentration.

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The S-N curves based fatigue designs of welded components and structures do not fully represent the weld details like effects of geometry, welding process and material defects. Experimental fatigue life prediction of welded joints is time consuming and costly. These limitations have prompted the need for adopting other ways for modeling the fatigue process rather than simply relating applied stress and fatigue life as in the S-N curves. A recent trend in this regard is the use of local approaches like fracture mechanics based attempts to model the whole fatigue process by considering the above effects. The stress at the **crack** tip can be accurately evaluated by means of linear elastic fracture mechanics by using suitable singular **crack** tip elements and proper mesh density. This can be further used to predict the fatigue life of a welded joint by using Paris’s law. The final **crack** length of a welded joint can be found out by virtual **crack** extension method (VCEM), where the final **crack** length is that **crack** length at which for a given loading the stress intensity factor of a **crack** approaches or exceeds an upper limit known as critical stress intensity factor.

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