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Effect of single initial overload and mean load on the low-cycle fatigue life of normalized 300 M alloy steel
International Journal of Fatigue 130 (2020) 105273
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International Journal of Fatigue journal homepage: www.elsevier.com/locate/ijfatigue
Effect of single initial overload and mean load on the low-cycle fatigue life of normalized 300 M alloy steel
T
Chris Bassindale, Ronald E. Miller, Xin Wang
⁎
Department of Mechanical and Aerospace Engineering, Carleton University, Ottawa, Ontario K1S 5B6, Canada
ARTICLE INFO
ABSTRACT
Keywords: Fatigue Residual stress Experimental testing Finite element analysis
In this work the effect of single overload and the resulting residual stress on the low-cycle fatigue life of 300 M alloy steel was experimentally investigated. Notched fatigue specimens were designed to localize the residual stress/strain and provide a known location for fatigue damage initiation. A common design assumption, whereby a residual stress is approximated as a mean stress was tested through comparing the fatigue life of specimens with residual stresses to the fatigue life of specimens with the same level of stress, except applied as a mean stress during testing. The residual stress was determined numerically through examining a finite element model of the test coupon. ABAQUS 2017x was used to generate the model and numerically solve the analysis. All simulations and tests were examined under quasi-static loading conditions. It was found that the residual stress generated by a single initial overload (tensile and compressive) had minimal effect on the low-cycle fatigue life of 300 M steel. The initial overload appeared to increase the life slightly. Moreover, it was shown that approximating the residual stress as a mean stress for calculating the expected life of a component is highly conservative with tensile residual stresses and highly non-conservative with compressive residual stresses when dealing with low-cycle fatigue.
1. Introduction Fatigue analysis techniques are significant to design engineers to ensure that a component will not fail before the end of the intended life. In the design of aircraft components, a high degree of importance is placed on the reliability of design methodologies to accurately predict the life of a component. One of the great challenges of designing components for a certain fatigue life is accurately predicting the loads which the component will encounter in its service life. One of the most unpredictable loading scenarios is a component overload, which can be caused by an aircraft experiencing high winds during flight, performing a strenuous maneuver, or a hard landing. The result of a component overload is the generation of a residual stress in load critical components, such as landing gear. Residual stresses are stresses which exist in a component after said component experiences a load which is higher than expected (designed for), causing yielding [1]. The cause of these stresses can be due to mechanical and thermal loading. In this work the focus will be on examining mechanically induced residual stresses. The effect of residual stress has been extensively studied for highcycle fatigue (HCF) loading conditions. In the work of Xu et al [2] steel specimens with surface-rolled induced residual stresses were tested to examine the effect of residual stresses on the endurance limit. A ⁎
summary of the effect of residual stress on the fatigue behaviour of vnotched components was presented in the work of Ferro et al [3]. In the work of Ramos et al [4], the effect of single and multiple overloads on the residual fatigue life of a structural steel was examined for a HCF loading condition using a compact tension specimen. It was observed that a single tensile overload increased the fatigue life of the specimen. This increase in fatigue life was attributed to the residual stresses in the vicinity of the crack tip following the overload cycle. This result suggests that in the presence of a crack tip (or notch) that an overload can be beneficial to the fatigue life. However, there has not been much attention paid to the effect of residual stress on the low-cycle fatigue (LCF) behaviour of engineering materials. This in large part is due to the fact that the residual stresses tend to be drowned out after the first few cycles due to the large amplitude of oscillating plastic strains [5]. The loss of residual stress in LCF testing was demonstrated in the work of Altenberger et al [6] for AISI 304 stainless steel. In their work shot peened specimens and deep rolled specimens were examined at half fatigue life to determine the amount of residual stress remaining. It was found that the residual stress was reduced by approximately 70%. This result of residual stress relaxation was further exemplified in the work of Martin et al [7] for AISI 1045 steel. Both studies only examined specimens in strain-controlled tests
Corresponding author. E-mail address: [email protected] (X. Wang).
https://doi.org/10.1016/j.ijfatigue.2019.105273 Received 5 July 2019; Received in revised form 23 August 2019; Accepted 13 September 2019 Available online 14 September 2019 0142-1123/ © 2019 Elsevier Ltd. All rights reserved.
International Journal of Fatigue 130 (2020) 105273
C. Bassindale, et al.
Nomenclature
n'
cyclic strain hardening exponent, fatigue strength coefficient, Pa
Symbol b c C L d D Kt
fatigue ductility coefficient, number of cycles to failure, true strain range, R radius of notch tip, mm E Young’s modulus, Pa Notch Angle angle of machined notch, Degrees
' f
Meaning, Units fatigue strength exponent, fatigue ductility exponent, number of circumferential elements, number of length elements, diameter at notch tip, mm nominal diameter of specimen, mm stress concentration factor, -
' f
Nf
with compressive residual stresses from deep rolling and shot peening. With the inclusion of stress concentrators in mechanical components an understanding of the effect of residual stresses is essential to confidently design structures for LCF. The objective of this work was to experimentally investigate the effect of residual stresses on the LCF life of 300 M alloy steel under load control fatigue testing and evaluate the assumption of equating the residual stress in an overloaded component to that of a mean stress for fatigue life predictions. This assumption was directly assessed through examining the overall life of notched fatigue specimens.
The fatigue tests were performed under load-control. The load levels for each test are shown in Table 5. All specimens were loaded at an average loading rate of 2 cycles per minute, this translates to an average approximate strain rate of 2 × 10 3 s 1 at the notch root. This loading rate was chosen to ensure the strain rate was in the realm of quasi-static and remove any effect of strain rate on the material strength. The specimens were placed in the wedge grips and gripped with a clamping pressure of 0.2 MPa. The residual stress (tensile or compressive) was introduced to the specimens through overloading the specimen to the required (compression or tension) load level and then releasing the load. The overloaded specimens were loaded over a time interval of 30 s and then released at the same loading rate. The equivalent mean stress was simply implemented during testing using the MTS control software.
2. Experimental methods 2.1. Material characterization The material examined in this work was 300 M alloy steel in the normalized state. The chemical composition and the quasi-static mechanical properties are shown in Tables 1 and 2, respectively. The stock material was acquired in 12.7 mm (1/2″) diameter round stock from the supplier. The mechanical properties were determined through uniaxial tensile tests. The stress and strain data can be seen plotted in Fig. 1. The ultimate tensile strength presented in Table 2 was a lower bound value, as the extensometer was removed prior to failure to prevent damaging the instrumentation. The specimens used for the uniaxial tensile test were manufactured according to the ASTM standard for tensile testing metallic materials [8]. The specimens were tested in the ‘as received’ state, meaning no additional heat treatment was performed after machining. The cyclic material parameters were estimated based on the criteria outlined in [9] and are presented in Table 3. Details of these parameters will be discussed later in Section 3.2.
3. Numerical model A two-dimensional, axis-symmetric finite element model was developed to estimate the level of residual stress at the root of the notch due to overloading. The finite element (FE) code ABAQUS 2017x [13] was used to generate the model and perform the analysis. The FE model is shown in Fig. 4. A quarter symmetry model was generated by replacing the bottom half of the notch with a boundary condition (BC). A symmetry BC was placed on the centreline, and a vertical constraint was placed on the symmetry plane around the notch. A reference point was assigned a tie constraint with the grip section of the model. This allowed for the simulation to be load controlled, while applying a constant stress along the gauge section. The material properties of 300 M, Table 2, were used as direct input into the model. The true stress and true strain data, Fig. 1, were used to describe the strain hardening of the material. The mesh of the model was generated around the root of the notch using a swept meshing technique [14]. This was done to better capture the stress and strain distribution around the notch root. The remainder of the model was assigned a structured mesh. A mesh sensitivity study was performed on the swept mesh surround the notch root. The number of elements along the lengths L and C in Fig. 5 were incrementally increased until the maximum stress was within 1% of the previous mesh. The implementation of the material model was verified through comparing the stress-strain data from the FE model to the experimental data obtained in Section 2.1. The numerical data was within 0.5% of the experimental data which validates the implementation of the material model. The model was further verified through comparing the Kt value calculated from the model to that extracted from [12]. A 0.5% difference was achieved between the model and the theoretical value.
2.2. Experimental procedure 2.2.1. Notched specimen design The shape of the fatigue specimens used in this work were cylindrical, conforming to dimensional guidelines presented in the ASTM standards for fatigue testing, [10] and [11]. The geometry of the specimen is shown in Fig. 2. The notch radius of the specimen was chosen to be the smallest radius achievable with the turning operation used to machine the specimens. The stress concentration factor was calculated using the charts in Pilkey et al. [12]. To maximize the Kt value a ratio of d/D equal to 0.75 was chosen. The notch angle was chosen to be 90° to allow for ease of manufacturing. The other critical design criterion for the specimen was to be able to withstand a high compressive load without buckling. Thus, it was also critical to maximize D and d. The final fatigue specimen notch parameters are summarized in Table 4.
Table 1 Chemical composition of 300 M alloy steel. 300 M
2.2.2. Experimental testing parameters The fatigue specimens were tested at room temperature (~20 °C) using an MTS 810 tensile testing system with a ± 25 kN load capacity. The experimental set-up is shown in Fig. 3.
Chemical Composition (wt.%) C 0.42
2
Cr 0.8
Fe Bal.
Mn 0.75
Mo 0.4
Ni 1.8
Va 0.07
Si 1.65
International Journal of Fatigue 130 (2020) 105273
C. Bassindale, et al.
Table 2 Quasi-static tensile properties of 300 M steel. (Average of 3 tests).
Average
Young’s Modulus (GPa)
Yield Stress (0.2% method) (MPa)
Ultimate Tensile Stress (MPa)
% Reduction of Area
201.6
935
1051*
74
* The ultimate tensile strength recorded is a lower bound value. Table 4 Fatigue specimen notch parameters. Specimen
R (mm)
d (mm)
D (mm)
R/D
d/D
Kt
Fatigue
0.2032
4.13
5.51
0.0369
0.75
3.0
Fig. 1. Tensile stress-strain data for 300 M alloy steel (Extensometer removed prior to fracture). Table 3 Estimated fatigue parameters for 300 M. Fatigue Parameter ' f ' f
c b
Value 1222 MPa 0.74 −0.55 −0.068
3.1. Determination of residual stress at the notch root The residual stress at the root of the notch was calculated using the FE model presented in Section 3. It is important to note that when the residual stress is mentioned in this work it is referring to the uniaxial normal stress component. The residual stress was treated as a one-dimensional stress component for the purposes of this work. The residual stress was calculated from the FE model by applying an overload and then releasing the load. The residual stress was then the remaining stress under zero external load. The residual stress as a function of the distance from the notch root for a tensile overloaded model can be seen in Fig. 6(A). Fig. 6(B) illustrates the residual stress distribution from the FEA (for compressive overload case). The Von Mises stress as a function of distance from the notch root is plotted in Fig. 7. As anticipated the tensile overload produced a compressive residual stress at the notch root, with a gradual increase to a tensile residual
Fig. 3. Instrumented fatigue specimen.
stress as the distance from the notch root is increased. The maximum stress located near the notch root was determined to be 1130 MPa from the FE model. The stress level exceeds the material yield stress; however, surrounding the notch root is a state of complex stress. To simplify the treatment of the stress in the notch root, the residual stress was approximated as the uniaxial normal component. This residual stress was induced through applying a 15 kN load. The equivalent mean stress was approximated through taking the 1130 MPa residual stress and calculating a load which would generate that stress at the notch root. It
Fig. 2. Geometry for notched fatigue specimen (units: mm). 3
International Journal of Fatigue 130 (2020) 105273
C. Bassindale, et al.
Table 5 Loading levels for fatigue specimens. Specimen Number
Overload (kN)
Mean load (kN)
Maximum load (kN)
Minimum load (kN)
1–3 4–6 7–9 10 – 12 13 – 15
0 −15 0 15 0
0 0 +5 0 −5
+10 +10 +15 +10 +5
−10 −10 −5 −10 −15
was found that a 5 kN mean load would generate the required stress at the notch root. 3.2. Notch root strain amplitude With the geometry of the specimen designed, the strain amplitude was assessed to cause the specimen to fail within a reasonable number of cycles. Low-cycle fatigue testing (number of cycles to failure (2Nf ) < 104 cycles) was chosen to be examined, therefore a strain based relation between the strain amplitude and the number of cycles at failure was used [9,11]. The relation, known as the strain-life model, is defined as
2
=
where
' f
E 2
(2Nf )b +
' f
(2Nf ) c
Fig. 5. FE model mesh.
specimen. The number of cycles at failure for each specimen are tabulated in Table 6. Three specimens were tested for each mean/residual stress case. The specimens tested with no residual stress and no mean stress (specimens 1–3) failed in the designed range between 900 and 1000 cycles. The average life of the specimens was 943 cycles. This verifies the design methodology for the fatigue specimens and sets the baseline for fatigue performance of the specimens. Though this is a small set of specimens, this result encourages the notion that the strain-life model can be used to model load-controlled testing; however, further validation is required.
(1)
is the strain amplitude, E is the material Young’s modulus,
' f
is the fatigue strength coefficient, is the fatigue ductility coefficient, and b and c are the fatigue strength and ductility exponents, respectively [15]. All the parameters for 300 M alloy steel are listed in Table 3. The strain amplitude calculated from Eq. (1) is for a straincontrolled test in which the strain amplitude is constant throughout the test. However, in the case of this study the strain-life model is used based on the initial strain amplitude of the first cycle. When the test is load-controlled the level of strain at the notch root can change throughout the test which is not reflected in the strain life model; however, this study will provide some insight into using the strain life model for load-controlled testing in which the specimen experiences plastic deformation. The desired number of cycles at failure was chosen to be approximately 1000 for a baseline specimen with no residual stress and no mean stress. From Eq. (1) the specimen would fail at a strain amplitude of approximately 0.015. The nominal strain at the notch root as a function of the applied load was examined to determine the required load to produce a strain level at the notch root. The strain amplitude at the root of the notch is plotted as a function of the applied load in Fig. 8. Thus, for the given notch geometry a load amplitude of approximately 10 kN would produce the desired initial strain amplitude to cause the specimen to fail in approximately 1000 cycles. ' f
4.1. Effect of mean stress on fatigue life The tensile mean stress specimens (specimens 7–9) showed a significant decrease in the specimen life compared to the baseline case (specimens 1–3). The average life of the specimens with the tensile mean stress was 496 cycles. This result was anticipated as increasing the maximum tensile stress increases the maximum stress/strain at the notch root in each cycle thus reducing the life of the specimen. The specimens which were tested under an equivalent compressive mean stress (specimens 13–15) presented a drastic increase in the number of cycles at failure. The average fatigue life of the specimens with the compressive mean stress was 3086 cycles. The increase in the life of the specimens with the compressive mean stress was anticipated as the lower mean stress (load) reduces the maximum strain amplitude at the notch tip and increases the life of the specimen. The inclusion of a mean stress for fatigue life calculations has been well studied for many engineering materials. Two prominent mean stress models are the Smith-Watson-Topper (SWT) model and the Morrow mean stress correction method. The SWT model was developed in the 1970′s by Smith et al. [16]. The model proposes a correction to both the elastic and plastic components of the strain-life equation outlined as
4. Experimental results and discussion All specimens were tested with a load amplitude of 10kN. The specific load cases are summarized in Table 5. The main result examined in this work was the number of cycles at failure for each
Fig. 4. FE model with boundary conditions. 4
International Journal of Fatigue 130 (2020) 105273
C. Bassindale, et al.
Fig. 6. Residual stress as a function of the distance from the notch root.
max
2
=
( f' )2 E
(2Nf ) 2b +
' ' bc f f (2Nf )
(2)
where max is the maximum tensile stress experienced during a cycle. The SWT model has had extensive success with grey cast iron [17], hardened carbon steels [18], and a micro-alloyed steel [19]. The SWT model is regarded as more promising for practical use [9]. The Morrow mean stress correction model [20], defined as
2
=
' f
m
E
(2Nf )b +
' f
(2Nf ) c
(3)
Applies a correction to the elastic portion of the strain-life equation through the inclusion of the mean stress, m . In the case of a tensile stress the strain amplitude is lowered and in the case of a compressive mean stress the strain amplitude is increased for a given life. The Morrow mean stress model has been extensively cited for its applicability to steels in the long-life regime, where plasticity is of little significance [9]. Both the Morrow model and the SWT model were developed assuming constant strain amplitude throughout the life of the component. In the current study, load-controlled experiments were performed with specimens experiencing high plasticity. As such the correction models are not explicitly applicable; however, this allows the examination of the applicability of the two models to force-controlled loading conditions. In the work of Fash and Socie [17], the SWT model was applied to grey cast iron under load-control with good agreement. The results of the current study were compared with the predictions
Fig. 8. Notch root strain amplitude as a function of applied load.
of both empirical models to assess their applicability to predicting force-controlled experimental results. The correction methods are compared with the average of the three tests performed at each mean stress level. The fatigue parameters used in the comparison are presented in Table 3. The SWT model is plotted with the experimental data in Fig. 9. Note the error bars on the graphs are to represent the
Fig. 7. Residual Von Mises stress as a function of the distance from the notch root. 5
International Journal of Fatigue 130 (2020) 105273
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are higher than the strain amplitude of the applied cyclic load. This in turn appears to apply the same effect as notch tip blunting which occurs during fatigue crack propagation. The larger strain field around the notch root acts to delay the onset of fatigue crack initiation, and thus delays the failure of the specimen. Examining the compressive residual stress specimens (specimens 10–12), a compressive residual stress does affect the number of cycles at failure. The average life of the specimens was 1120. Similar to the specimens with the tensile residual stress a slight increase in number of cycles at failure was observed. Interestingly, the nature of the residual stress did not appear to affect the fatigue life of the specimen. This is due to the inclusion of residual stress/strain at the notch root, which was higher than the cyclic amplitude stress/strain, which in turn delays the fracture of the specimen. Of course, the argument of the residual stress relaxation around the notch root under high plastic strain remains an issue to assess; however, the current results suggest that if the residual stress/strain around the notch root, as a result of the overload is sufficiently high, that enough residual stress remains to impact the overall life of the component. The results of the current effort support this notion strongly and it is the intention of the authors’ to further explore this topic.
Table 6 Experimental results. Specimen Number
Nominal Mean Stress (Notch Root) (MPa)
Residual Stress (MPa)
Fatigue life (Cycles)
Average Life (Cycles)
1 2 3
0 0 0
0 0 0
1010 907 910
943
4 5 6
0 0 0
1130 1130 1130
1228 1037 1058
1108
7 8 9
376 (1130) 376 (1130) 376 (1130)
0 0 0
581 467 439
496
10 11 12
0 0 0
−1130 −1130 −1130
1159 971 1229
1120
13 14 15
−376 (−1130) −376 (−1130) −376 (−1130)
0 0 0
3325 3156 2776
3086
4.3. Comparison of fatigue life with residual stress and mean stress A common design assumption when attempting to design for residual stress is to apply the level of residual stress as a mean stress for a fatigue life calculation. The fatigue life is plotted as a function of the residual/mean stress for all stress cases in Fig. 11. Examining the tensile mean stress specimens and the tensile residual stress specimens a 50% difference in fatigue life was observed. The specimens tested with a tensile mean stress failed after half the cycles of the specimens with the tensile residual stress. The assumption of equating the tensile residual stress to a mean stress is a conservative assumption as it reduces the expected life. Comparing the life of the compressive mean stress specimens to the specimens with the compressive residual stress a 200% difference is observed. Equating the compressive residual stress to a compressive mean stress is shown to be highly non-conservative as the specimens with a compressive residual stress failed earlier than the specimens with the compressive mean stress.
Fig. 9. Comparison of SWT mean stress model with experimental data.
dispersion of the experimental data. The experimental data is in fair agreement with the SWT model. The average tensile mean stress fatigue life results are within 15% of the SWT model, while the compressive mean stress data presented a 30% difference from the SWT model. The experimental data is shown plotted against the Morrow model in Fig. 10. The negative mean stress experimental fatigue life data presented a 4% difference from the Morrow model, while the tensile mean stress data presented a 12% difference with the Morrow model. Of course, the current study can only compare two experimental data points with the two models; however, the strong agreement with both models is promising for their applicability to the chosen material. It is the authors’ intention to conduct further experimental investigations to properly assess the best model for 300 M in the future.
4.4. Fracture surface analysis An overview of the fracture surface of the baseline fatigue specimen is shown in Fig. 12. The crack initiated from the surface of the specimen (notch root) and propagated inward. The cracks initiated from surface defects and voids produced during fabrication. The crack propagation
4.2. Effect of residual stress on fatigue life The effect of residual stress was systematically assessed by mechanically inducing tensile and compressive residual stresses in specimens and then comparing the fatigue life to specimens with no residual stress. The specimens with a tensile residual stress (specimens 4–6) show a slight increase in their fatigue life. The average life of the specimens with tensile residual stresses was 1108 cycles. This slight increase in fatigue performance is due to the residual stress and strain in the notch root. The residual stress and strain surrounding the notch root
Fig. 10. Comparison of Morrow mean stress model with experimental data. 6
International Journal of Fatigue 130 (2020) 105273
C. Bassindale, et al.
region of fatigue crack propagation, Fig. 12(B), presented an inlay of microscopic cracks and striations. The region of final failure, Fig. 12(A), was immediately next to the propagation region and presented many voids of varying size. The final failure region was elongated, and fracture resembled that of the cup-cone shape indicating a ductile failure mode. The transition region from fatigue crack propagation to final fracture is visibly evident for each specimen. The most notable comparison comes from comparing the tensile mean stress and compressive mean stress specimens depicted in Fig. 13(A) and (B), respectively. The tensile mean stress specimen failed with a larger fracture region and smaller crack propagation region. This makes sense as the tensile mean stress specimen experiences a larger maximum tensile load which would cause the specimen to fail with a larger cross section. The compressive mean stress specimen presented a smaller failure region since it experienced a smaller maximum tensile load. The smaller maximum tensile load resulted in a larger fatigue crack propagation region and ultimately required the final cross-section to be significantly smaller to cause fracture. Furthermore, the compressive mean stress specimen would experience a larger compressive load which would cause the fatigue propagation region to contact itself under a larger amount of force and damage the surface showing it to be smooth and flat.
Fig. 11. Effect of residual stress and mean stress (notch root) on the number of cycles at failure.
region is flat in all specimens, and then it transitions to a rough elongated fracture region in the centre of the specimen. These two regions are shown by the areas marked (B) and (A) in Fig. 12, respectively. The
Fig. 12. SEM images of baseline fatigue specimen. (A) Microscopic image of inner failure region. (B) Microscopic image of fatigue crack propagation region. 7
International Journal of Fatigue 130 (2020) 105273
C. Bassindale, et al.
Fig. 13. Macroscopic SEM images of the fatigue specimen fracture surfaces. (A) Tensile mean stress specimen. (B) Compressive mean stress specimen. (C) Tensile residual stress specimen. (D) Compressive residual stress specimen.
The SEM images of the tensile and compressive residual stress specimens are shown in Fig. 13(C) and (D), respectively. The fatigue crack propagation regions of the residual stress specimens are of a comparable size. This indicates that the residual stress mainly affected the crack initiation and had little to no effect on the propagation or the final failure of the specimen.
Acknowledgements
5. Conclusions
References
The authors gratefully acknowledge funding provided by the Natural Sciences and Engineering Research Council (NSERC) of Canada. The authors would also like to acknowledge the assistance from Mr. Steve Truttmann for his valuable insight and advice with fatigue testing.
In conclusion, the effect of residual stress on the fatigue life of 300 M alloy steel was experimentally examined. It was observed that 300 M alloy steel behaves according to the strain-life model under LCF loading conditions. Furthermore, good agreement was achieved through comparing the experimental data with the Morrow mean stress model and the SWT model. The assumption of approximating the residual stress in a component as a mean stress for a fatigue life prediction was experimentally investigated. The results demonstrate that for the stress states tested, both compressive or tensile residual stresses had a minimal effect on the low-cycle fatigue life of a component manufactured from 300 M alloy steel. Moreover, it has also been shown that the assumption of approximating a tensile residual stress as a tensile mean stress is a very conservative assumption. On the other hand, approximating the compressive residual stress as a mean stress is highly non-conservative and should not be used.
[1] Juvinall R, Marshek K. Fundamentals of Machine Component Design. Danvers: John Wiley & Sons Inc.; 2012. [2] Xu K, He J, Zhou H. Effect of residual stress on fatigue behaviour of notches. Int J Fatigue 1994;16(5):337–43. [3] Ferro P, Berto F, Razavi S, Ayatollahi M. The effect of residual stress on fatigue behavior of V-notched components: a review. Procedia Struct Integr 2017;3:119–25. [4] Ramos M, Pereira M, Darwish F, Motta S, Carneiro M. Effect of single and multiple overloading on the residual fatigue life of a structural steel. Fatigue Fract Eng Mater Struct 2003;26(2):115–21. [5] Withers P. Residual stress and its role in failure. Manchester: IOP Publishing Ltd; 2007. [6] Altenberger I, Scholtes B, Martin U, Oettel H. Cyclic deformation and near surface microstructures of shot peened or deep rolled austenitic stainless steel AISI 304. Mater Sci Eng:A 1999;264(1–2):1–16. [7] Martin U, Altenberger I, Scholtes B, Kremmer K, Oettel H. Cyclic deformation and near surface microstructures of normalized shot peened steel SAE 1045. Mater Sci Eng, A 1998;246(1–2):69–80. [8] ASTM, ASTM /E8M-16a Standard Test Methods for Tension Testing of Metallic Materials, West Conshohocken: ASTM International, 2018.
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C. Bassindale, et al. [9] Lee Y-L, Pan J, Hathaway R, Barkley M. Fatigue Testing and Analysis. Burlington: Elsevier Inc.; 2005. [10] ASTM, ASTM E466-15 Standard Practice For Conducting Force Controlled Constant Amplitude Axial Fatigue Tests of Metallic Materials, West Conshohocken: ASTM International, 2018. [11] ASTM, ASTM E606/E606M-12 Standard Practice For Strain Controlled Fatigue Testing, West Conshohocken: ASTM International, 2018. [12] Pilkey W, Pilkey D. Peterson's Stress Concentration Factors. New Jersey: John Wiley and Sons Inc; 2008. [13] Dassault Systemes, “ABAQUS Version 2017,” ABAQUS Inc., Providence, USA, 2017. [14] ABAQUS, ABAQUS 2017, 2017 - Analysis User’s Manual, 2017. [15] Morrow J. Cyclic plastic strain energy and fatigue of metals. Internal Friction,
Damping and Cyclic Plasticity, West Conshohocken, PA, ASTM. 1965. p. 45–68. [16] Smith K, Watson P, Topper T. A Stress-strain function for the fatigue of metals. J Mater 1970;5(4):767–78. [17] Fash J, Socie D. Fatigue behaviour and mean effects in grey cast iron. Int J Fatigue 1982;4(3):137–42. [18] Koh S, Stephens R. Mean stress effects on low-cycle fatigue for a high strength steel. Fatigue Fract Eng Mater Struct 1991;14(4):413–28. [19] Forsetti P, Blasarin A. Fatigue behaviour of microalloyed steels for hot-forged mechanical components. Int J Fatigue 1988;10(3):153–61. [20] Morrow M. Fatigue Design Handbook, Section 3.2, SAE Advances in Engineering. Warrendale, PA: Society of Automotive Engineers. 1968. p. 21–9.
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International Journal of Fatigue journal homepage: www.elsevier.com/locate/ijfatigue
Effect of single initial overload and mean load on the low-cycle fatigue life of normalized 300 M alloy steel
T
Chris Bassindale, Ronald E. Miller, Xin Wang
⁎
Department of Mechanical and Aerospace Engineering, Carleton University, Ottawa, Ontario K1S 5B6, Canada
ARTICLE INFO
ABSTRACT
Keywords: Fatigue Residual stress Experimental testing Finite element analysis
In this work the effect of single overload and the resulting residual stress on the low-cycle fatigue life of 300 M alloy steel was experimentally investigated. Notched fatigue specimens were designed to localize the residual stress/strain and provide a known location for fatigue damage initiation. A common design assumption, whereby a residual stress is approximated as a mean stress was tested through comparing the fatigue life of specimens with residual stresses to the fatigue life of specimens with the same level of stress, except applied as a mean stress during testing. The residual stress was determined numerically through examining a finite element model of the test coupon. ABAQUS 2017x was used to generate the model and numerically solve the analysis. All simulations and tests were examined under quasi-static loading conditions. It was found that the residual stress generated by a single initial overload (tensile and compressive) had minimal effect on the low-cycle fatigue life of 300 M steel. The initial overload appeared to increase the life slightly. Moreover, it was shown that approximating the residual stress as a mean stress for calculating the expected life of a component is highly conservative with tensile residual stresses and highly non-conservative with compressive residual stresses when dealing with low-cycle fatigue.
1. Introduction Fatigue analysis techniques are significant to design engineers to ensure that a component will not fail before the end of the intended life. In the design of aircraft components, a high degree of importance is placed on the reliability of design methodologies to accurately predict the life of a component. One of the great challenges of designing components for a certain fatigue life is accurately predicting the loads which the component will encounter in its service life. One of the most unpredictable loading scenarios is a component overload, which can be caused by an aircraft experiencing high winds during flight, performing a strenuous maneuver, or a hard landing. The result of a component overload is the generation of a residual stress in load critical components, such as landing gear. Residual stresses are stresses which exist in a component after said component experiences a load which is higher than expected (designed for), causing yielding [1]. The cause of these stresses can be due to mechanical and thermal loading. In this work the focus will be on examining mechanically induced residual stresses. The effect of residual stress has been extensively studied for highcycle fatigue (HCF) loading conditions. In the work of Xu et al [2] steel specimens with surface-rolled induced residual stresses were tested to examine the effect of residual stresses on the endurance limit. A ⁎
summary of the effect of residual stress on the fatigue behaviour of vnotched components was presented in the work of Ferro et al [3]. In the work of Ramos et al [4], the effect of single and multiple overloads on the residual fatigue life of a structural steel was examined for a HCF loading condition using a compact tension specimen. It was observed that a single tensile overload increased the fatigue life of the specimen. This increase in fatigue life was attributed to the residual stresses in the vicinity of the crack tip following the overload cycle. This result suggests that in the presence of a crack tip (or notch) that an overload can be beneficial to the fatigue life. However, there has not been much attention paid to the effect of residual stress on the low-cycle fatigue (LCF) behaviour of engineering materials. This in large part is due to the fact that the residual stresses tend to be drowned out after the first few cycles due to the large amplitude of oscillating plastic strains [5]. The loss of residual stress in LCF testing was demonstrated in the work of Altenberger et al [6] for AISI 304 stainless steel. In their work shot peened specimens and deep rolled specimens were examined at half fatigue life to determine the amount of residual stress remaining. It was found that the residual stress was reduced by approximately 70%. This result of residual stress relaxation was further exemplified in the work of Martin et al [7] for AISI 1045 steel. Both studies only examined specimens in strain-controlled tests
Corresponding author. E-mail address: [email protected] (X. Wang).
https://doi.org/10.1016/j.ijfatigue.2019.105273 Received 5 July 2019; Received in revised form 23 August 2019; Accepted 13 September 2019 Available online 14 September 2019 0142-1123/ © 2019 Elsevier Ltd. All rights reserved.
International Journal of Fatigue 130 (2020) 105273
C. Bassindale, et al.
Nomenclature
n'
cyclic strain hardening exponent, fatigue strength coefficient, Pa
Symbol b c C L d D Kt
fatigue ductility coefficient, number of cycles to failure, true strain range, R radius of notch tip, mm E Young’s modulus, Pa Notch Angle angle of machined notch, Degrees
' f
Meaning, Units fatigue strength exponent, fatigue ductility exponent, number of circumferential elements, number of length elements, diameter at notch tip, mm nominal diameter of specimen, mm stress concentration factor, -
' f
Nf
with compressive residual stresses from deep rolling and shot peening. With the inclusion of stress concentrators in mechanical components an understanding of the effect of residual stresses is essential to confidently design structures for LCF. The objective of this work was to experimentally investigate the effect of residual stresses on the LCF life of 300 M alloy steel under load control fatigue testing and evaluate the assumption of equating the residual stress in an overloaded component to that of a mean stress for fatigue life predictions. This assumption was directly assessed through examining the overall life of notched fatigue specimens.
The fatigue tests were performed under load-control. The load levels for each test are shown in Table 5. All specimens were loaded at an average loading rate of 2 cycles per minute, this translates to an average approximate strain rate of 2 × 10 3 s 1 at the notch root. This loading rate was chosen to ensure the strain rate was in the realm of quasi-static and remove any effect of strain rate on the material strength. The specimens were placed in the wedge grips and gripped with a clamping pressure of 0.2 MPa. The residual stress (tensile or compressive) was introduced to the specimens through overloading the specimen to the required (compression or tension) load level and then releasing the load. The overloaded specimens were loaded over a time interval of 30 s and then released at the same loading rate. The equivalent mean stress was simply implemented during testing using the MTS control software.
2. Experimental methods 2.1. Material characterization The material examined in this work was 300 M alloy steel in the normalized state. The chemical composition and the quasi-static mechanical properties are shown in Tables 1 and 2, respectively. The stock material was acquired in 12.7 mm (1/2″) diameter round stock from the supplier. The mechanical properties were determined through uniaxial tensile tests. The stress and strain data can be seen plotted in Fig. 1. The ultimate tensile strength presented in Table 2 was a lower bound value, as the extensometer was removed prior to failure to prevent damaging the instrumentation. The specimens used for the uniaxial tensile test were manufactured according to the ASTM standard for tensile testing metallic materials [8]. The specimens were tested in the ‘as received’ state, meaning no additional heat treatment was performed after machining. The cyclic material parameters were estimated based on the criteria outlined in [9] and are presented in Table 3. Details of these parameters will be discussed later in Section 3.2.
3. Numerical model A two-dimensional, axis-symmetric finite element model was developed to estimate the level of residual stress at the root of the notch due to overloading. The finite element (FE) code ABAQUS 2017x [13] was used to generate the model and perform the analysis. The FE model is shown in Fig. 4. A quarter symmetry model was generated by replacing the bottom half of the notch with a boundary condition (BC). A symmetry BC was placed on the centreline, and a vertical constraint was placed on the symmetry plane around the notch. A reference point was assigned a tie constraint with the grip section of the model. This allowed for the simulation to be load controlled, while applying a constant stress along the gauge section. The material properties of 300 M, Table 2, were used as direct input into the model. The true stress and true strain data, Fig. 1, were used to describe the strain hardening of the material. The mesh of the model was generated around the root of the notch using a swept meshing technique [14]. This was done to better capture the stress and strain distribution around the notch root. The remainder of the model was assigned a structured mesh. A mesh sensitivity study was performed on the swept mesh surround the notch root. The number of elements along the lengths L and C in Fig. 5 were incrementally increased until the maximum stress was within 1% of the previous mesh. The implementation of the material model was verified through comparing the stress-strain data from the FE model to the experimental data obtained in Section 2.1. The numerical data was within 0.5% of the experimental data which validates the implementation of the material model. The model was further verified through comparing the Kt value calculated from the model to that extracted from [12]. A 0.5% difference was achieved between the model and the theoretical value.
2.2. Experimental procedure 2.2.1. Notched specimen design The shape of the fatigue specimens used in this work were cylindrical, conforming to dimensional guidelines presented in the ASTM standards for fatigue testing, [10] and [11]. The geometry of the specimen is shown in Fig. 2. The notch radius of the specimen was chosen to be the smallest radius achievable with the turning operation used to machine the specimens. The stress concentration factor was calculated using the charts in Pilkey et al. [12]. To maximize the Kt value a ratio of d/D equal to 0.75 was chosen. The notch angle was chosen to be 90° to allow for ease of manufacturing. The other critical design criterion for the specimen was to be able to withstand a high compressive load without buckling. Thus, it was also critical to maximize D and d. The final fatigue specimen notch parameters are summarized in Table 4.
Table 1 Chemical composition of 300 M alloy steel. 300 M
2.2.2. Experimental testing parameters The fatigue specimens were tested at room temperature (~20 °C) using an MTS 810 tensile testing system with a ± 25 kN load capacity. The experimental set-up is shown in Fig. 3.
Chemical Composition (wt.%) C 0.42
2
Cr 0.8
Fe Bal.
Mn 0.75
Mo 0.4
Ni 1.8
Va 0.07
Si 1.65
International Journal of Fatigue 130 (2020) 105273
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Table 2 Quasi-static tensile properties of 300 M steel. (Average of 3 tests).
Average
Young’s Modulus (GPa)
Yield Stress (0.2% method) (MPa)
Ultimate Tensile Stress (MPa)
% Reduction of Area
201.6
935
1051*
74
* The ultimate tensile strength recorded is a lower bound value. Table 4 Fatigue specimen notch parameters. Specimen
R (mm)
d (mm)
D (mm)
R/D
d/D
Kt
Fatigue
0.2032
4.13
5.51
0.0369
0.75
3.0
Fig. 1. Tensile stress-strain data for 300 M alloy steel (Extensometer removed prior to fracture). Table 3 Estimated fatigue parameters for 300 M. Fatigue Parameter ' f ' f
c b
Value 1222 MPa 0.74 −0.55 −0.068
3.1. Determination of residual stress at the notch root The residual stress at the root of the notch was calculated using the FE model presented in Section 3. It is important to note that when the residual stress is mentioned in this work it is referring to the uniaxial normal stress component. The residual stress was treated as a one-dimensional stress component for the purposes of this work. The residual stress was calculated from the FE model by applying an overload and then releasing the load. The residual stress was then the remaining stress under zero external load. The residual stress as a function of the distance from the notch root for a tensile overloaded model can be seen in Fig. 6(A). Fig. 6(B) illustrates the residual stress distribution from the FEA (for compressive overload case). The Von Mises stress as a function of distance from the notch root is plotted in Fig. 7. As anticipated the tensile overload produced a compressive residual stress at the notch root, with a gradual increase to a tensile residual
Fig. 3. Instrumented fatigue specimen.
stress as the distance from the notch root is increased. The maximum stress located near the notch root was determined to be 1130 MPa from the FE model. The stress level exceeds the material yield stress; however, surrounding the notch root is a state of complex stress. To simplify the treatment of the stress in the notch root, the residual stress was approximated as the uniaxial normal component. This residual stress was induced through applying a 15 kN load. The equivalent mean stress was approximated through taking the 1130 MPa residual stress and calculating a load which would generate that stress at the notch root. It
Fig. 2. Geometry for notched fatigue specimen (units: mm). 3
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Table 5 Loading levels for fatigue specimens. Specimen Number
Overload (kN)
Mean load (kN)
Maximum load (kN)
Minimum load (kN)
1–3 4–6 7–9 10 – 12 13 – 15
0 −15 0 15 0
0 0 +5 0 −5
+10 +10 +15 +10 +5
−10 −10 −5 −10 −15
was found that a 5 kN mean load would generate the required stress at the notch root. 3.2. Notch root strain amplitude With the geometry of the specimen designed, the strain amplitude was assessed to cause the specimen to fail within a reasonable number of cycles. Low-cycle fatigue testing (number of cycles to failure (2Nf ) < 104 cycles) was chosen to be examined, therefore a strain based relation between the strain amplitude and the number of cycles at failure was used [9,11]. The relation, known as the strain-life model, is defined as
2
=
where
' f
E 2
(2Nf )b +
' f
(2Nf ) c
Fig. 5. FE model mesh.
specimen. The number of cycles at failure for each specimen are tabulated in Table 6. Three specimens were tested for each mean/residual stress case. The specimens tested with no residual stress and no mean stress (specimens 1–3) failed in the designed range between 900 and 1000 cycles. The average life of the specimens was 943 cycles. This verifies the design methodology for the fatigue specimens and sets the baseline for fatigue performance of the specimens. Though this is a small set of specimens, this result encourages the notion that the strain-life model can be used to model load-controlled testing; however, further validation is required.
(1)
is the strain amplitude, E is the material Young’s modulus,
' f
is the fatigue strength coefficient, is the fatigue ductility coefficient, and b and c are the fatigue strength and ductility exponents, respectively [15]. All the parameters for 300 M alloy steel are listed in Table 3. The strain amplitude calculated from Eq. (1) is for a straincontrolled test in which the strain amplitude is constant throughout the test. However, in the case of this study the strain-life model is used based on the initial strain amplitude of the first cycle. When the test is load-controlled the level of strain at the notch root can change throughout the test which is not reflected in the strain life model; however, this study will provide some insight into using the strain life model for load-controlled testing in which the specimen experiences plastic deformation. The desired number of cycles at failure was chosen to be approximately 1000 for a baseline specimen with no residual stress and no mean stress. From Eq. (1) the specimen would fail at a strain amplitude of approximately 0.015. The nominal strain at the notch root as a function of the applied load was examined to determine the required load to produce a strain level at the notch root. The strain amplitude at the root of the notch is plotted as a function of the applied load in Fig. 8. Thus, for the given notch geometry a load amplitude of approximately 10 kN would produce the desired initial strain amplitude to cause the specimen to fail in approximately 1000 cycles. ' f
4.1. Effect of mean stress on fatigue life The tensile mean stress specimens (specimens 7–9) showed a significant decrease in the specimen life compared to the baseline case (specimens 1–3). The average life of the specimens with the tensile mean stress was 496 cycles. This result was anticipated as increasing the maximum tensile stress increases the maximum stress/strain at the notch root in each cycle thus reducing the life of the specimen. The specimens which were tested under an equivalent compressive mean stress (specimens 13–15) presented a drastic increase in the number of cycles at failure. The average fatigue life of the specimens with the compressive mean stress was 3086 cycles. The increase in the life of the specimens with the compressive mean stress was anticipated as the lower mean stress (load) reduces the maximum strain amplitude at the notch tip and increases the life of the specimen. The inclusion of a mean stress for fatigue life calculations has been well studied for many engineering materials. Two prominent mean stress models are the Smith-Watson-Topper (SWT) model and the Morrow mean stress correction method. The SWT model was developed in the 1970′s by Smith et al. [16]. The model proposes a correction to both the elastic and plastic components of the strain-life equation outlined as
4. Experimental results and discussion All specimens were tested with a load amplitude of 10kN. The specific load cases are summarized in Table 5. The main result examined in this work was the number of cycles at failure for each
Fig. 4. FE model with boundary conditions. 4
International Journal of Fatigue 130 (2020) 105273
C. Bassindale, et al.
Fig. 6. Residual stress as a function of the distance from the notch root.
max
2
=
( f' )2 E
(2Nf ) 2b +
' ' bc f f (2Nf )
(2)
where max is the maximum tensile stress experienced during a cycle. The SWT model has had extensive success with grey cast iron [17], hardened carbon steels [18], and a micro-alloyed steel [19]. The SWT model is regarded as more promising for practical use [9]. The Morrow mean stress correction model [20], defined as
2
=
' f
m
E
(2Nf )b +
' f
(2Nf ) c
(3)
Applies a correction to the elastic portion of the strain-life equation through the inclusion of the mean stress, m . In the case of a tensile stress the strain amplitude is lowered and in the case of a compressive mean stress the strain amplitude is increased for a given life. The Morrow mean stress model has been extensively cited for its applicability to steels in the long-life regime, where plasticity is of little significance [9]. Both the Morrow model and the SWT model were developed assuming constant strain amplitude throughout the life of the component. In the current study, load-controlled experiments were performed with specimens experiencing high plasticity. As such the correction models are not explicitly applicable; however, this allows the examination of the applicability of the two models to force-controlled loading conditions. In the work of Fash and Socie [17], the SWT model was applied to grey cast iron under load-control with good agreement. The results of the current study were compared with the predictions
Fig. 8. Notch root strain amplitude as a function of applied load.
of both empirical models to assess their applicability to predicting force-controlled experimental results. The correction methods are compared with the average of the three tests performed at each mean stress level. The fatigue parameters used in the comparison are presented in Table 3. The SWT model is plotted with the experimental data in Fig. 9. Note the error bars on the graphs are to represent the
Fig. 7. Residual Von Mises stress as a function of the distance from the notch root. 5
International Journal of Fatigue 130 (2020) 105273
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are higher than the strain amplitude of the applied cyclic load. This in turn appears to apply the same effect as notch tip blunting which occurs during fatigue crack propagation. The larger strain field around the notch root acts to delay the onset of fatigue crack initiation, and thus delays the failure of the specimen. Examining the compressive residual stress specimens (specimens 10–12), a compressive residual stress does affect the number of cycles at failure. The average life of the specimens was 1120. Similar to the specimens with the tensile residual stress a slight increase in number of cycles at failure was observed. Interestingly, the nature of the residual stress did not appear to affect the fatigue life of the specimen. This is due to the inclusion of residual stress/strain at the notch root, which was higher than the cyclic amplitude stress/strain, which in turn delays the fracture of the specimen. Of course, the argument of the residual stress relaxation around the notch root under high plastic strain remains an issue to assess; however, the current results suggest that if the residual stress/strain around the notch root, as a result of the overload is sufficiently high, that enough residual stress remains to impact the overall life of the component. The results of the current effort support this notion strongly and it is the intention of the authors’ to further explore this topic.
Table 6 Experimental results. Specimen Number
Nominal Mean Stress (Notch Root) (MPa)
Residual Stress (MPa)
Fatigue life (Cycles)
Average Life (Cycles)
1 2 3
0 0 0
0 0 0
1010 907 910
943
4 5 6
0 0 0
1130 1130 1130
1228 1037 1058
1108
7 8 9
376 (1130) 376 (1130) 376 (1130)
0 0 0
581 467 439
496
10 11 12
0 0 0
−1130 −1130 −1130
1159 971 1229
1120
13 14 15
−376 (−1130) −376 (−1130) −376 (−1130)
0 0 0
3325 3156 2776
3086
4.3. Comparison of fatigue life with residual stress and mean stress A common design assumption when attempting to design for residual stress is to apply the level of residual stress as a mean stress for a fatigue life calculation. The fatigue life is plotted as a function of the residual/mean stress for all stress cases in Fig. 11. Examining the tensile mean stress specimens and the tensile residual stress specimens a 50% difference in fatigue life was observed. The specimens tested with a tensile mean stress failed after half the cycles of the specimens with the tensile residual stress. The assumption of equating the tensile residual stress to a mean stress is a conservative assumption as it reduces the expected life. Comparing the life of the compressive mean stress specimens to the specimens with the compressive residual stress a 200% difference is observed. Equating the compressive residual stress to a compressive mean stress is shown to be highly non-conservative as the specimens with a compressive residual stress failed earlier than the specimens with the compressive mean stress.
Fig. 9. Comparison of SWT mean stress model with experimental data.
dispersion of the experimental data. The experimental data is in fair agreement with the SWT model. The average tensile mean stress fatigue life results are within 15% of the SWT model, while the compressive mean stress data presented a 30% difference from the SWT model. The experimental data is shown plotted against the Morrow model in Fig. 10. The negative mean stress experimental fatigue life data presented a 4% difference from the Morrow model, while the tensile mean stress data presented a 12% difference with the Morrow model. Of course, the current study can only compare two experimental data points with the two models; however, the strong agreement with both models is promising for their applicability to the chosen material. It is the authors’ intention to conduct further experimental investigations to properly assess the best model for 300 M in the future.
4.4. Fracture surface analysis An overview of the fracture surface of the baseline fatigue specimen is shown in Fig. 12. The crack initiated from the surface of the specimen (notch root) and propagated inward. The cracks initiated from surface defects and voids produced during fabrication. The crack propagation
4.2. Effect of residual stress on fatigue life The effect of residual stress was systematically assessed by mechanically inducing tensile and compressive residual stresses in specimens and then comparing the fatigue life to specimens with no residual stress. The specimens with a tensile residual stress (specimens 4–6) show a slight increase in their fatigue life. The average life of the specimens with tensile residual stresses was 1108 cycles. This slight increase in fatigue performance is due to the residual stress and strain in the notch root. The residual stress and strain surrounding the notch root
Fig. 10. Comparison of Morrow mean stress model with experimental data. 6
International Journal of Fatigue 130 (2020) 105273
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region of fatigue crack propagation, Fig. 12(B), presented an inlay of microscopic cracks and striations. The region of final failure, Fig. 12(A), was immediately next to the propagation region and presented many voids of varying size. The final failure region was elongated, and fracture resembled that of the cup-cone shape indicating a ductile failure mode. The transition region from fatigue crack propagation to final fracture is visibly evident for each specimen. The most notable comparison comes from comparing the tensile mean stress and compressive mean stress specimens depicted in Fig. 13(A) and (B), respectively. The tensile mean stress specimen failed with a larger fracture region and smaller crack propagation region. This makes sense as the tensile mean stress specimen experiences a larger maximum tensile load which would cause the specimen to fail with a larger cross section. The compressive mean stress specimen presented a smaller failure region since it experienced a smaller maximum tensile load. The smaller maximum tensile load resulted in a larger fatigue crack propagation region and ultimately required the final cross-section to be significantly smaller to cause fracture. Furthermore, the compressive mean stress specimen would experience a larger compressive load which would cause the fatigue propagation region to contact itself under a larger amount of force and damage the surface showing it to be smooth and flat.
Fig. 11. Effect of residual stress and mean stress (notch root) on the number of cycles at failure.
region is flat in all specimens, and then it transitions to a rough elongated fracture region in the centre of the specimen. These two regions are shown by the areas marked (B) and (A) in Fig. 12, respectively. The
Fig. 12. SEM images of baseline fatigue specimen. (A) Microscopic image of inner failure region. (B) Microscopic image of fatigue crack propagation region. 7
International Journal of Fatigue 130 (2020) 105273
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Fig. 13. Macroscopic SEM images of the fatigue specimen fracture surfaces. (A) Tensile mean stress specimen. (B) Compressive mean stress specimen. (C) Tensile residual stress specimen. (D) Compressive residual stress specimen.
The SEM images of the tensile and compressive residual stress specimens are shown in Fig. 13(C) and (D), respectively. The fatigue crack propagation regions of the residual stress specimens are of a comparable size. This indicates that the residual stress mainly affected the crack initiation and had little to no effect on the propagation or the final failure of the specimen.
Acknowledgements
5. Conclusions
References
The authors gratefully acknowledge funding provided by the Natural Sciences and Engineering Research Council (NSERC) of Canada. The authors would also like to acknowledge the assistance from Mr. Steve Truttmann for his valuable insight and advice with fatigue testing.
In conclusion, the effect of residual stress on the fatigue life of 300 M alloy steel was experimentally examined. It was observed that 300 M alloy steel behaves according to the strain-life model under LCF loading conditions. Furthermore, good agreement was achieved through comparing the experimental data with the Morrow mean stress model and the SWT model. The assumption of approximating the residual stress in a component as a mean stress for a fatigue life prediction was experimentally investigated. The results demonstrate that for the stress states tested, both compressive or tensile residual stresses had a minimal effect on the low-cycle fatigue life of a component manufactured from 300 M alloy steel. Moreover, it has also been shown that the assumption of approximating a tensile residual stress as a tensile mean stress is a very conservative assumption. On the other hand, approximating the compressive residual stress as a mean stress is highly non-conservative and should not be used.
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