Slant Crack Growth Research Articles

Crack tip flipping under mode I tearing: Investigated by X-ray tomography

The fracture surface morphology that results from mode I tearing of ductile plate metals depends heavily on both the elastic-plastic material properties and the microstructure. Severe tunneling of the advancing crack tip (resulting in cup-cup, or bath-tub like fracture surfaces) can take place in a range of materials, often of low strength, while tearing of high strength metals typically progress by the shear band failure mechanism (slanting). In reality, however, most fracture surfaces display a mixture of morphologies. For example, slant crack propagation can be accompanied by large shear lips near the outer free plate surface or a complete shear band switch - seemingly distributed randomly on the fracture surface. The occasionally observed shear band switch of mode I slant cracks, related to ductile plate tearing, is far from random as the crack can flip systematically from one side to the other in roughly 45° shear bands. This "flipping" action of a slanted crack remains to be fully understood, and the present study serves to share details on the phenomenon by exploiting X-ray tomography scanning to access the plate interior and the very crack tip. Throughout, the focus is on a crack tip where the flip is underway. Extensive growth of single edge cracks under mode I loading is achieved in a purpose build test set-up. Here, considering a 4mm plate of normal strength / high strain hardening steel which has been found to display successive flipping of the slant crack face. While undergoing a shear band switch, such that the flipping mechanism is active, the plate tearing test is interrupted and the crack tip extracted for further investigation. The conducted X-ray tomography scans reveal the failure process ahead of the advancing crack tip to resemble the ductile slant crack growth governed by local thinning and moderate crack tip tunneling. However, small shear lips form at the outer free plate surface, well behind the 45° slant (tunneling) crack tip, as the flipping action engages. Upon further loading, the shear lips subsequently grow to form a set of secondary crack fronts at an angle to the primary tunneling slant crack. Eventually, these secondary crack fronts catch up on the primary slant crack front and overtake the growth to complete the shear band switch. Once the crack slants, an out-of-plane action occurs due to the loss of symmetry in the system. It is this out-of-plane action which is believed to set-off the flipping mechanism.

Fatigue crack growth under cyclic torsional loading

In the research herein presented, mode III and mode I fatigue loading tests were carried out on standard compact tension (CT) specimens using a bi-axial servo-hydraulic Instron 8874 machine. Thin sheet CT specimens were made of two types of austenitic stainless steels, namely an AISI 316L stainless steel and a high-strength Cr-Mn stainless steel. Specimens were pre-cracked in mode I, at room temperature, and crack growth paths were experimentally determined. Fracture surfaces of rectangular sections were observed using optical devices and crack growth revealed to be higher at the external surfaces than at the mid plane of the specimens tested under mode III fatigue loading; in addition, smooth slant crack growth occurred, as well as crack branching. Fracture surfaces obtained under mode I loading revealed that crack grew flat, coplanar and normal to the tensile axis, as expected. Additionally, stress intensity factors present at tip of each branch crack loaded under mode III were calculated through finite element analyses (FEA). Hence, although an out-of-plane torsional loading was remotely applied, a mixture of modes was foreseen to be applied locally at crack tip during crack growth, namely mode I, mode III and mode II loading, in that order of relevance, and equivalent stress intensity factor values at the crack tip were calculated using Pook, Richard or Schöllmann et al. criterion.

FEM Evaluation of the Dissipated Energy in Front of a Crack Tip under 2D Mixed Mode Loading Condition

Numerical and experimental studies of fatigue crack growth under both pure mode I and combined mode I and II were performed. A CT-like specimen with an additional hole for mixed-mode loading made from an Al-alloy D16CT1 was used. A 2D elastic-plastic FEM simulation of a growing fatigue crack allowed us to determine the value of the total plastic energy dissipation per cycle throughout the cyclic (reversed) plastic zone directly from the elastic-plastic stress and strain fields in the vicinity of a crack tip. According to the Klingbeil's theory, the total plastic energy dissipated ahead of a crack tip is related to the fatigue crack growth rate. To review this theory, experimentally observed fatigue crack growth rates were correlated to the computed energies for both straight and slant crack growth. Moreover, a comparison according to the conventional concept of an equivalent stress intensity factor is presented.

Application of a strip-yield model to predict crack growth under variable-amplitude and spectrum loading – Part 1: Compact specimens

Fatigue-crack-growth tests were conducted on compact, C( T), specimens made of D16Cz (clad) aluminum alloy under constant-amplitude loading, a single spike overload, and simulated aircraft spectrum loading. Constant-amplitude tests were conducted to generate crack-growth-rate data from threshold to near fracture over a wide range of stress ratios ( R = P min/ P max = 0.1–0.75) using the new compression pre-cracking test methods. Comparisons were made between test data generated on the C( T) specimens with test data from the literature on middle-crack-tension, M( T), specimens machined from the same sheet. A crack-closure analysis was used to collapse the rate data from both specimen types into a narrow band over many orders of magnitude in rates using proper constraint factors. The constraint factors were established from constant-amplitude (CA) and single-spike overload tests. The life-prediction code, FASTRAN, which is based on the strip-yield model concept, was used to calculate crack-length-against-cycles under CA loading and a single-spike overload (OL) test, and to predict crack growth under simulated aircraft spectrum loading tests on C( T) specimens. The calculated crack-growth lives under CA loading were generally within about ±25% of the test results, but slower crack growth under the double-shear fatigue mode, unlike the single-shear mode (45 o slant crack growth), may be the reason for some of the larger differences. The predicted results under the single-spike overload and the Mini-Falstaff+ spectrum were within 10% of the test data.

Energy-based assessment of ductile tearing in a thin sheet aluminium alloy

This paper deals with several issues in ductile tearing assessment. One of them is the contrasting of energy-based parameters, namely, the Essential Work of Fracture (EWF) [B. Cotterell, J.K. Reddel, The essential work of plane stress ductile fracture, International Journal of Fracture 13 (1977) 267–277], wet, and the Energy Dissipation Rate (EDR) [C.E. Turner, A re-assessment of ductile tearing resistance (Part I and II). In: Fracture Behaviour and Design of Materials and Structures 1990; Proc. ECF 8; 2: 933–949 and 951–968], R. We intend to answer the following questions: (i) Is there a reasonable correlation between wet and R? (ii) Is there any correlation between their constituents? and, finally, (iii) Are the energy- and displacement-based characterizations of slant crack growth under uncontained yielding consistent with each other?

フレッティング疲労における傾斜き裂進展に及ぼすき裂位置の影響

It is well known that fretting fatigue cracks usually initiate in the contact area near contact edge and grow obliquely to the normal direction of the contact surface. In order to determine the fretting fatigue crack growth mechanism, the stress intensity factors KI and KII have been widely used. The authors have established a scheme and software to obtain the values of KI and KII of the oblique fretting fatigue cracks which are initiated at the contact edge and subjected to arbitrarily distributed contact pressure with friction. In this paper, the solution has been expanded to the fretting fatigue cracks initiated near contact edge. The calculated results of stress intensity factors show that the distance between crack position and contact edge markedly affects on the values of KI and KII. Crack propagation direction has been estimated by using maximum tensile stress (MTS) criterion according to the calculated values of KI and KII. The effect of mode II stress intensity factor KII due to normal and tangential stress distributions near the contact edge on the crack propagation direction has been discussed. It has been concluded from the present study that the slant crack growth in fretting fatigue results from crack initiation near contact edge.

CTOA and crack-tunneling measurements in thin sheet 2024-T3 aluminum alloy

The stable tearing behavior of 2.3-mm thick sheets of 2024-T3 aluminum alloy was experimentally investigated for middle crack tension specimens having initial flaws that were: (a) flat fatigue cracks (low fatigue stress) and (b) 45-deg through-thickness slant cracks (high fatigue stress). The critical CTOA values during stable tearing were measured by two independent methods, optical microscopy and digital-image correlation. Results from the two methods agree well. The CTOA measurements and observations of the fracture surfaces have shown that the initial stable tearing behavior of low and high fatigue stress tests is significantly different. The cracks in the low fatigue stress tests underwent a transition from flat-to-slant crack growth, during which the CTOA values were high and significant crack tunneling occurred. After crack growth equal to about the thickness (Δa>B), CTOA reached a constant value of 6 deg and after crack growth equal to about twice the thickness (Δa>2B), crack tunneling stabilized. The cracks in the high fatigue stress tests reach the same constant CTOA value after crack growth equal to about the thickness, but produced only slightly higher CTOA values during initial crack growth. The amount of tunneling in the high fatigue stress tests was about the same as that in the low fatigue stress tests after the flat-to-slant transition. This study indicates that stress history has an influence on the initial portion of the stable tearing behavior. The initial high CTOA values, in the low fatigue crack tests, coincided with large three-dimensional crack front shape changes due to a variation in the through-thickness crack-tip constraint. The measured CTOA reached a constant value of 6 deg for crack growth of about the specimen thickness. This coincided with the onset of 45-deg slant crack growth and a stabilized, slightly tunneled (about 20 percent of the thickness) crack-front shape. For crack growth on the 45-deg slant, the crack front and local field variables are still highly three dimensional.

Photoelastic analysis of friction-grip bolted joints: Czerkas, K., Kratena, J., Bartos, F., Junkova, S. and Kratenova, M. Acta Tech. CSAV 1990 35, (1), 93–102

Fatigue-crack-growth tests were conducted on compact, C(T), specimens made of D16Cz (clad) aluminum alloy under constant-amplitude loading, a single spike overload, and simulated aircraft spectrum loading. Constant-amplitude tests were conducted to generate crack-growth-rate data from threshold to near fracture over a wide range of stress ratios (R = Pmin/Pmax = 0.1–0.75) using the new compression pre-cracking test methods. Comparisons were made between test data generated on the C(T) specimens with test data from the literature on middle-crack-tension, M(T), specimens machined from the same sheet. A crack-closure analysis was used to collapse the rate data from both specimen types into a narrow band over many orders of magnitude in rates using proper constraint factors. The constraint factors were established from constant-amplitude (CA) and single-spike overload tests. The life-prediction code, FASTRAN, which is based on the strip-yield model concept, was used to calculate crack-length-against-cycles under CA loading and a single-spike overload (OL) test, and to predict crack growth under simulated aircraft spectrum loading tests on C(T) specimens. The calculated crack-growth lives under CA loading were generally within about ±25% of the test results, but slower crack growth under the double-shear fatigue mode, unlike the single-shear mode (45o slant crack growth), may be the reason for some of the larger differences. The predicted results under the single-spike overload and the Mini-Falstaff+ spectrum were within 10% of the test data.

Effect of oxidation on high-temperature, high-cycle fatigue behaviour of engineering steels: Kanazawa, K., Sato, M. and Kimura, M. J. Soc. Mater. Sci., Japan Feb. 1990 39, (437), 175–180 (in Japanese)

Fatigue-crack-growth tests were conducted on compact, C(T), specimens made of D16Cz (clad) aluminum alloy under constant-amplitude loading, a single spike overload, and simulated aircraft spectrum loading. Constant-amplitude tests were conducted to generate crack-growth-rate data from threshold to near fracture over a wide range of stress ratios (R = Pmin/Pmax = 0.1–0.75) using the new compression pre-cracking test methods. Comparisons were made between test data generated on the C(T) specimens with test data from the literature on middle-crack-tension, M(T), specimens machined from the same sheet. A crack-closure analysis was used to collapse the rate data from both specimen types into a narrow band over many orders of magnitude in rates using proper constraint factors. The constraint factors were established from constant-amplitude (CA) and single-spike overload tests. The life-prediction code, FASTRAN, which is based on the strip-yield model concept, was used to calculate crack-length-against-cycles under CA loading and a single-spike overload (OL) test, and to predict crack growth under simulated aircraft spectrum loading tests on C(T) specimens. The calculated crack-growth lives under CA loading were generally within about ±25% of the test results, but slower crack growth under the double-shear fatigue mode, unlike the single-shear mode (45o slant crack growth), may be the reason for some of the larger differences. The predicted results under the single-spike overload and the Mini-Falstaff+ spectrum were within 10% of the test data.

Slant fatigue crack growth in A1 2024 sheet material

Fatigue crack growth experiments have been performed on centre cracked specimens of A1 2024-T351. For two plate thicknesses (6 and 10.3 mm) experiments under constant stress intensity and under constant stress amplitude have been carried out. It appeared that at higher ΔK shear lips develop at the edges of the specimen, giving rise to slant crack growth. For constant ΔK tests the width of the shear lips reaches an equilibrium state of width. It requires some crack growth before this steady state is reached. Both the steady state and non-steady state will be described. The results obtained in the constant ΔK tests are used to explain the behaviour under constant stress amplitude loading.

On slant (45°) crack growth in fatigue

to be expected for slant crack growth,4 as confirmed by the Ritchie, Smith, and Knott tests. Buckling involves out-of-plane deformations, but these should not be confused with those caused by throughthickness yielding associated with plane-stress conditions at a crack tip. When pronounced, buckling in very thin sheets can reduce strength through an increase in K I due to the inflexion which appears at a crack tip,5 not by the appearance of Mode III .deformations; in this case slant growth is Mode I, by the small-scale argument. Buckling would not normally be expected in 'thin' specimens of the type used by Ritchie, Smith, and Knott, but if it did occur it would obviously cause a particular 45 plane to be favoured. Observation of fracture surfaces shows6 that the transition to slant crack growth takes place by the development of shear lips (Fig. 1), which increase in width until they reach a maximum size or, in 'thin' specimens, they meet, completing the transition. (Tunnelling of the crack in the centre of the specimen while the transition is taking place is more likely to be the result of the complex three-dimensional stress system in the transition region than evidence that square crack growth is faster.) The length of this transition zone of mixed square and slant fracture depends on precise loading conditions.6 The problem of what causes slant growth therefore reduces to finding the conditions for a shear lip to develop. The appearance of shear lips in thick specimens rules out buckling as a cause. Shear lips appear on both sides of a specimen at about the same time, so their appearance must be due to the satisfaction of some condition. In thick specimens shear lips appear randomly on parallel and perpendicular planes (Fig. 1), whereas in thin specimens they usually, but not invariably,6 appear on parallel planes. Material textures which suppress through-thickness yielding also suppress shear lips,' as can changes in work-hardening behaviour.! Slant crack growth only Qccurs in those materials in which the static tensile fracture of a piece of the same material of the same thickness also occurs on a 450 through-the-thickness plane.6 RITCHIE, SMITH, and KNOTT have recently published! some data on the fatigue crack growth properties of mild steel. Loading conditions were such that medium rates of crack growth (5 x 10-6 to 5 X 10-4 mm/cycle) were obtained. This is above the region of the threshold for fatigue-crack growth, but below the region where gross plasticity effects are likely. Data were analysed in terms of !:::.K I, the range of opening mode stress-intensity factor K I, in a fatigue cycle. As would be expected, results f~r thick (13, 15, 20, and 48 mm) singleedge-notched (SEN) specimens tested in pure bending agreed reasonably well with results for 20 mm thick SEN specimens tested in tension. The relatively minor inconsistencies observed were probably rightly attributed to shortcomings in the use of !:::.K I in the analysis. An interesting feature of the results for thin specimens is that for 2 mm SEN bend specimens the crack -growth rates were about half those for thick specimens, with 'square' growth on a 90 plane, whereas for 1 and 2 mm SEN tension specimens the crack -growth rates were about twice those for thick specimens with 'slant' growth on a 45 plane through the specimen thickness. Possible reasons for these differences were discussed in general terms, but some further clarification of the fracture mechanics involved is desirable; various factors involved -are therefore briefly _reviewed and discussed.