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Research Papers

Finite Element Simulation of Shielding/Intensification Effects of Primary Inclusion Clusters in High Strength Steels Under Fatigue Loading

[+] Author and Article Information
Nima Salajegheh

SNC-Lavalin Inc.,
Toronto, ON M9C 5K1, Canada
e-mail: Nima_Salajegheh@gatech.edu

R. Prasannavenkatesan, Herng-Jeng Jou

QuesTek Innovations LLC,
Evanston, IL 60201

David L. McDowell

George W. Woodruff School of
Mechanical Engineering,
Georgia Institute of Technology,
Atlanta, GA 30332;
School of Materials Science and Engineering,
Georgia Institute of Technology,
Atlanta, GA 30332

Gregory B. Olson

Department of Materials
Science and Engineering,
Robert R. McCormick School of Engineering and
Applied Science,
Northwestern University,
Evanston, IL 60208;
QuesTek Innovations LLC,
Evanston, IL 60201

1Corresponding author.

Contributed by the Materials Division of ASME for publication in the JOURNAL OF ENGINEERING MATERIALS AND TECHNOLOGY. Manuscript received May 12, 2013; final manuscript received March 30, 2014; published online April 30, 2014. Assoc. Editor: Georges Cailletaud.

J. Eng. Mater. Technol 136(3), 031003 (Apr 30, 2014) (8 pages) Paper No: MATS-13-1083; doi: 10.1115/1.4027380 History: Received May 12, 2013; Revised March 30, 2014

The change of potency to nucleate cracks in high cycle fatigue (HCF) at a primary nonmetallic inclusion in a martensitic gear steel due to the existence of a neighboring inclusion is computationally investigated using two-and three-dimensional elastoplastic finite element (FE) analyses. Fatigue indicator parameters (FIPs) are computed in the proximity of the inclusion and used to compare crack nucleation potency of various scenarios. The nonlocal average value of the maximum plastic shear strain amplitude is used in computing the FIP. Idealized spherical (cylindrical in 2D) inclusions with homogeneous linear elastic isotropic material properties are considered to be partially debonded, the worst case scenario for HCF crack nucleation as experimentally observed for similar systems (Furuya et al., 2004, “Inclusion-Controlled Fatigue Properties of 1800 Mpa-Class Spring Steels,” Metall. Mater. Trans. A, 35A(12), pp. 3737–3744; Harkegard, 1974, “Experimental Study of the Influence of Inclusions on the Fatigue Properties of Steel,” Eng. Fract. Mech., 6(4), pp. 795–805; Lankford and Kusenberger, 1973, “Initiation of Fatigue Cracks in 4340 Steel,” Metall. Mater. Trans. A, 4(2), pp. 553–559; Laz and Hillberry, 1998, “Fatigue Life Prediction From Inclusion Initiated Cracks,” Int. J. Fatigue, 20(4), pp. 263–270). Inclusion-matrix interfaces are simulated using a frictionless contact penalty algorithm. The fully martensitic steel matrix is modeled as elastic-plastic with pure nonlinear kinematic hardening expressed in a hardening minus dynamic recovery format. FE simulations suggest significant intensification of plastic shear deformation and hence higher FIPs when the inclusion pair is aligned perpendicular to the uniaxial stress direction. Relative to the reference case with no neighboring inclusion, FIPs decrease considerably when the inclusion pair aligns with the applied loading direction. These findings shed light on the anisotropic HCF response of alloys with primary inclusions arranged in clusters by virtue of the fracture of a larger inclusion during deformation processing. Materials design methodologies may also benefit from such cost-efficient parametric studies that explore the relative influence of microstructure attributes on the HCF properties and suggest strategies for improving HCF resistance of alloys.

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References

Figures

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Fig. 1

SEM micrographs of mating fracture surfaces of subsurface fatigue initiation at clusters of (a) La2O2S and (b) Al2O3 particles in Ferrium® C61 martensitic gear steel [7]

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Fig. 2

(a) Schematic of the two-dimensional finite element model, (b) inclusion spacing and orientation, and (c) detailed view of the FE mesh

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Fig. 3

Contours of plastic strain magnitude at the end of 3rd load before unloading showing FIP averaging regions (εmax = 0.8 εys, Rε = 0; d = 8 μm, both inclusions are debonded and 20 μm in diameter)

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Fig. 4

2D FE prediction of ΔγN dependence on inclusion orientation (εmax = 0.8 εys, Rε = 0; both inclusions are 20 μm in diameter)

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Fig. 5

Variation of ΔγN versus inclusion spacing for θ = 0 deg (εmax = 0.8 εys and Rε = 0)

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Fig. 6

Variation of ΔγN versus particle spacing at θ = 90 deg (εmax = 0.8 εys and Rε = 0)

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Fig. 7

Ratio of beneficial shielding gain at θ = 90 deg to the unfavorable ΔγN at θ = 0 deg versus particle spacing (εmax = 0.8 εys and Rε = 0)

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Fig. 8

ΔγN dependence on remote applied strain amplitude (20 μm inclusions at d = 10 μm, Rε = 0)

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Fig. 9

Variation of ΔΓN versus inclusion orientation (εmax = 0.8 εys and Rε = 0)

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Fig. 10

Schematic of the three-dimensional finite element model

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Fig. 11

3D FE prediction of ΔγN dependence on inclusion orientation (εmax = 0.8 εys and Rε = 0)

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