International Journal of Solids and Structures 112 (2017) 83–96

Contents lists available at ScienceDirect

International Journal of Solids and Structures

journal homepage: www.elsevier.com/locate/ijsolstr

On the choice of boundary conditions for micromechanical

simulations based on 3D imaging

Modesar Shakoor a , ∗, Ante Buljac

b , c , Jan Neggers b , François Hild

b , Thilo F. Morgeneyer c , Lukas Helfen

d , e , Marc Bernacki a , Pierre-Olivier Bouchard

a

a MINES ParisTech, PSL Research University, CEMEF-Centre de mise en forme des matériaux, CNRS UMR 7635, CS 10207 rue Claude Daunesse, 06904 Sophia

Antipolis Cedex, France b Laboratoire de Mécanique et Technologie (LMT), ENS Cachan/CNRS/Université Paris-Saclay, 61 avenue du Président Wilson, 94235 Cachan, France c MINES ParisTech, PSL Research University, Centre des Matériaux, CNRS UMR 7633, BP 87, 91003 Evry, France d Institute for Photon Science and Synchrotron Radiation, Karlsruhe Institute of Technology (KIT), 76131 Karlsruhe, Germany e European Synchrotron Radiation Facility (ESRF), 38043 Grenoble, France

a r t i c l e i n f o

Article history:

Received 24 August 2016

Revised 2 December 2016

Available online 17 February 2017

Keywords:

Digital Volume correlation

Ductile fracture

Level set

Multiscale analysis

Micromechanical modeling

Microstructure meshing

X-ray laminography

a b s t r a c t

The present paper addresses the challenge of conducting Finite Element (FE) micromechanical simulations

based on 3D X-ray data, and quantifying errors between simulations and experiments. This is of great in-

terest, for example, in the study of ductile fracture as local comparisons and error indicators would help

understanding the limitations of current plasticity and damage models. Standard methods used in the lit-

erature to conduct FE simulations at the microscale are often based on multiscale schemes. Relevant me-

chanical fields computed in an FE simulation at the specimen scale are used as boundary conditions for

the micromechanical simulation, where the real microstructure is meshed from 3D X-ray images. These

methods hence rely on an identification of material behavior at the macroscale, say, using force measure-

ments and 2D surface images. In an earlier work by the authors, a method for conducting micromechan-

ical simulations using measured boundary conditions thanks to Digital Volume Correlation (DVC) was

proposed. The interest of this DVC-FE approach is that it uses solely 3D X-ray images acquired in-situ

during the experiments. Thus, FE simulations are directly conducted at the microscale, with no depen-

dence on specimen scale simulations or multiscale schemes. This approach also includes a methodology

to perform local error measurements with respect to experimental observations. In this paper, both mul-

tiscale schemes and this DVC-FE approach are applied to new experimental results on a nodular graphite

cast iron specimen with machined holes. Ductile fracture due to the nucleation, growth and coalescence

of microscopic voids between the machined holes is observed in-situ thanks to synchrotron 3D imaging.

The objective of this paper is to assess the accuracy of boundary conditions for each approach and con-

clude on the optimal choice. Based on both average and local error measurements, it is shown that void

growth is underestimated with multiscale schemes, while predictions are significantly improved with the

DVC-FE approach.

© 2017 Elsevier Ltd. All rights reserved.

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. Introduction

In the aim for understanding the physical nature of any phe-

omenon, comparison between experiments and theoretical mod-

ls is used extensively ( Avril et al., 2008 ). In the case of ductile

racture, difficulties arise due to the competition between local-

zation and softening events occurring at different scales ( Teko ̌glu

t al., 2015; Pineau et al., 2016 ). In particular, investigations of duc-

∗ Corresponding author.

E-mail address: modesar.shakoor@mines-paristech.fr (M. Shakoor).

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ttp://dx.doi.org/10.1016/j.ijsolstr.2017.02.018

020-7683/© 2017 Elsevier Ltd. All rights reserved.

ile fracture mechanisms, namely, void nucleation, growth and co-

lescence at the microscale require 3D observations at this scale.

n the numerical side, being able to simulate these mechanisms

ith comparable geometric and loading conditions is still an open

esearch topic, which is once again due to the scale at which the

eometry and the loading have to be identified.

Although 3D in-situ experimental observations ( e.g., X-ray to-

ography and laminography) have provided valuable information

or ductile fracture modeling ( Bouchard et al., 2008; Weck and

ilkinson, 2008; Cao et al., 2014; Lecarme et al., 2014; Seo et al.,

015a; Hannard et al., 2016 ), numerous questions regarding com-

84 M. Shakoor et al. / International Journal of Solids and Structures 112 (2017) 83–96

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parisons between models and experiments are still to be answered.

For instance, the well-known work of Gurson (1975) has motivated

researchers to rely extensively on homogenization theory to model

ductile fracture based on analytical ( Gurson, 1975; Gologanu et al.,

1993; Benzerga and Leblond, 2010; Scheyvaerts et al., 2011 ) or

numerical procedures ( Needleman and Tvergaard, 1984; Benzerga

and Leblond, 2010; Scheyvaerts et al., 2011; Bosco et al., 2015; Cao

et al., 2015 ). Even though interesting results can be obtained on

average values, the applicability of these homogenized models to

real 3D microstructures is quite limited due to the assumptions of

idealistic microstructures and boundary conditions on which they

are based ( Landron et al., 2010; Fansi et al., 2013; Cao et al., 2014;

Hannard et al., 2016 ).

Numerical procedures at the microscale could be applied given

that 3D X-ray images of the microstructure are acquired in-situ ,

that a microstructure meshing technique is available, and that

boundary conditions are applied. This is a critical piece of infor-

mation that will be analyzed herein. There are different ways of

performing such microscale simulations:

• The most straightforward approach is Direct Numerical Simu-

lation (DNS), where the microstructure of the whole specimen

has to be meshed, hence taking into account only microscale

constitutive models. Since the whole specimen is simulated,

boundary conditions are applied directly at pins, as in the ex-

periment. However, depending on specimen size, DNS can have

a huge cost regarding both experiments and simulations. Scan-

ning the whole specimen could require multiple scans at each

loading step, and this large set of 3D data would then have to

be meshed.

• To avoid this huge computational cost, full specimen Finite Ele-

ment (FE) simulations only partially taking into account the mi-

crostructure have been considered ( Tian et al., 2010; Hosokawa

et al., 2013a; Tang et al., 2013; Alinaghian et al., 2014; Hütter

et al., 2014; O’Keeffe et al., 2015 ). Microscale constitutive mod-

els are used in the ROI where the microstructure is meshed,

while macroscopic homogeneity is assumed in the remainder

of the specimen. Thus, an appropriate macroscale constitutive

model has to be defined and identified for this out-of-ROI ho-

mogeneous material. This approach will be referred to as strong

FE (sFE) coupling.

• Opposed to the previous approach in which the ROI mesh is

directly embedded within the specimen mesh, the two FE cal-

culations can be weakly coupled. In this weak FE (wFE) ap-

proach ( Tvergaard and Hutchinson, 2002; Bandstra et al., 2004;

Kaye et al., 2013 ), the specimen scale simulation assumes a ho-

mogeneous material in the whole domain. A relevant mechan-

ical field is then transferred from this first simulation to the

second simulation at the ROI scale, where the microstructure is

meshed. This mechanical field defines the boundary conditions

for the micromechanical simulation. These conditions can use

displacements, forces, or a combination of both.

• In a recent work ( Buljac et al., 2017 ), an alternative option

for performing microscale simulations was proposed. Digital

Volume Correlation (DVC) was used in order to measure dis-

placement fields between 3D X-ray images taken at consecu-

tive loading steps ( Roux et al., 2008; Rannou et al., 2010; Bul-

jac et al., 2016 ). The DVC technique is based on tracking the

natural image contrast, herein originating from the heteroge-

neous microstructure. FE simulations were conducted by ap-

plying measured DVC displacements to the boundaries of the

meshed ROI. The difference between FE results and DVC dis-

placement fields inside the FE domain could be assessed. Ad-

ditionally, a better quantification of the error produced by FE

models was obtained by computing gray level residuals based

on FE and DVC displacement fields. An advantage of this so-

called DVC-FE method is that no specimen scale simulation is

required, thus no macroscale constitutive model has to be iden-

tified. Yet, it requires a thorough uncertainty assessment re-

garding DVC measurements.

The present paper aims at showing the interest of the DVC-FE

pproach and compare its predictive capability with simulations at

he microscale based upon either weak or strong FE couplings. One

ey aspect is related to the boundary conditions that are applied

o the simulated ROI where the microstructure is meshed. This

nvestigation is conducted with new experimental results using

odular graphite cast iron specimens with a geometry inspired

rom the work of Weck and Wilkinson (2008) . A first test (A) was

erformed using small loading steps in order to obtain accurate

orce measurements and 2D surface images to be exploited thanks

o global Digital Image Correlation (DIC). The second test (B) was

erformed using larger loading steps since it was conducted in a

ynchrotron facility. For this second test, both 3D X-ray scans of

he ROI and 2D surface images were acquired (see Fig. 1 ).

While the DVC-FE methodology presented above uses directly

nd solely the 3D X-ray data, the two alternatives to obtain bound-

ry conditions for ROI calculations considered in this paper ( i.e.,

FE and sFE) rely on force measurements and 2D surface images.

orce measurements are used to identify the behavior of the mate-

ial at the specimen scale, as well as 2D surface images thanks to a

ecent Integrated-DIC framework ( Mathieu et al., 2015 ). In the wFE

ethod, an FE simulation of the experiment at the specimen scale

s conducted, and calculated displacement fields are used to drive a

econd FE calculation at the ROI scale. In the sFE method, the ROI

s embedded and meshed directly inside the specimen mesh and

nly one FE simulation is conducted. In this second method, the

aterial behavior is modeled using microscale constitutive models

n the ROI, while macroscale constitutive models are used in the

est of the domain.

The aforementioned approaches with weak (wFE) and strong

sFE) couplings between specimen and ROI calculations are applied

o a real 3D microstructure observed thanks to 3D in-situ laminog-

aphy experiments. To the best of the authors’ knowledge, this is

he first time such simulations are conducted and compared lo-

ally to experimental observations, this last point being possible

hanks to gray level residual computation. The results are then

ompared to those obtained using the DVC-FE approach ( Buljac

t al., 2017 ), showing the interest of this methodology. Details re-

arding the material, the experiments and the numerical method

sed for ROI calculations are presented in Section 2 , while the

echnical implementation of each approach for boundary condi-

ions is described in Section 3 . The results are presented and com-

ared based on error measurements with respect to experimental

mages in Section 4 .

. Experimental and numerical framework

.1. Experiments

The material used in this study is a commercial nodular

raphite cast iron with the serial code EN-GJS-400. It features a

erritic matrix and graphite nodules at the microscale (see Fig. 2 ),

ith no significant porosity in the initial state. Under tensile load-

ng, ductile fracture is known to be mainly driven by early debond-

ng of the nodules from the matrix and coalescence of the sub-

equent nucleated voids ( Dong et al., 1997; Hütter et al., 2015;

omi ̌cevi ́c et al., 2016 ). Previous works ( Dong et al., 1997; Zhang

t al., 1999; Bonora and Ruggiero, 2005; Hütter et al., 2015 ) sug-

est to model the nodules as voids, as their load carrying capacity

s very low under tensile loading. This assumption is made herein

M. Shakoor et al. / International Journal of Solids and Structures 112 (2017) 83–96 85

Fig. 1. Schematic view of the sample with zoomed region between the holes showing on the right: DVC (blue) and FE (cyan) meshes plotted over the corresponding cast

iron microstructure in isometric view. On the left: surface image with speckle pattern and DIC mesh from test (B). (For interpretation of the references to color in this figure

legend, the reader is referred to the web version of this article.)

Fig. 2. (a) Sample geometry with the scanned region between the pin holes; (b) section of the reconstructed volume with ROI position.

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eaning that at the microscale the material is considered as a

wo-phase microstructure with a ferritic matrix and voids.

The specimen geometry, which is inspired by the work of Weck

nd Wilkinson (2008) , is shown in Fig. 2 . The holes have been ma-

hined via Electrical Discharge Machining (EDM). The testing de-

ice applies the load by manually controlling the displacement via

crew rotation.

Test (A) is conducted as a pre-test for test (B) to study in de-

ail the particular sample behavior ( i.e., to assess the load levels for

he scanning procedure). Therefore, identical sample geometry and

aterial are used for test (B). The load/displacement curve for test

A) is shown in Fig. 5 (b). Load data acquired during test (A) refer

o peak values. The time lapses between peak loads and lamino-

raphic scans are less than 10 min for test (B).

In test (B) after applying each loading step, a scan is acquired

hile the sample is rotated about the laminographic axis ( i.e., par-

llel to the specimen thickness direction). This axis is inclined with

espect to the X-ray beam direction by an angle θ ≈ 60 °. The 3D

mages used in this work were obtained at beamline ID15A of the

uropean Synchrotron Radiation Facility (ESRF, Grenoble, France)

ith a 60 keV white beam, using 30 0 0 projections per scan. The

eries of radiographs acquired is then used to reconstruct 3D

olumes by using a filtered-back projection algorithm ( Myagotin

t al., 2013 ). The parameter optimization has been performed au-

86 M. Shakoor et al. / International Journal of Solids and Structures 112 (2017) 83–96

Fig. 3. Mid-thickness section of the reconstructed volume for three different loading steps.

Table 1

Elasto plastic properties of the ferritic matrix.

E (GPa) ν σ y (MPa) K (MPa) n

210 0 .30 290 382 0 .35

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tomatically using a GPU-accelerated implementation of this algo-

rithm ( Vogelgesang et al., 2016 ). The reconstructed volume has a

size of 1600 × 1600 × 1600 voxels. The physical size (length) of

1 cubic voxel is equal to 1 . 1 μm. After scanning the undeformed

state (0), 12 additional scans are acquired during the stepwise

loading procedure, where the last scan corresponds to the fully

opened final crack.

The scanned region incorporates two holes while the ROI em-

ployed in DVC and FE calculations is mainly concentrated in the

ligament between the holes (see Figs. 1 and 2 ). Since the two

machined holes have diameters of 500 μm, the nodule popula-

tion (treated as voids in the FE calculations) in the ligament area

with the characteristic size of 60 μm can be considered as a sec-

ondary void population. Hence, subsequent concurrent micro and

macro plasticity and damage localization phenomena can be ob-

served. Weck and Wilkinson (2008) used machined holes of mi-

crometer size, which made impossible the observation of the sec-

ondary void population (this limitation on the size of observable

voids is actually due to current imaging techniques ( Tasan et al.,

2012 )). Here, the larger size of the holes, and the large size of the

graphite nodules allow for such observations. This is illustrated in

Fig. 3 where mid-thickness sections of the reconstructed volume

for the three different load stages are shown. As a consequence

of the mentioned multiscale flow conditions, classical void coales-

cence mechanisms are accompanied by sheet coalescence between

the two machined holes that is observed in the last loading step

(deformed state (11)).

2.2. FE mesh of the microstructure

The framework used for microscale FE simulations is based on

previous developments ( Roux et al., 2013; 2014; Shakoor et al.,

2015b; 2015a; 2017 ). As explained by Buljac et al. (2017) , the ROI

used for FE simulations has to be included in all DVC domains

throughout each loading step. Otherwise displacement fields would

not be available as boundary conditions on the whole boundary

of the FE mesh. This requirement is necessary for the DVC-FE ap-

proach but also for other approaches in order to allow for compar-

isons with experimental observations. In practice, in order to be as

representative as possible, a 3D box as large as possible is chosen.

This 3D box has to be small enough to remain in all scanned re-

gions throughout loading. The image meshing technique used to

model the microstructure observed in the experimental 3D im-

ages starts with standard image processing operations ( Schindelin

et al., 2012; Schneider et al., 2012 ). They consist of smoothing the

data, applying a gray value threshold to separate matrix and voids,

and then converting these binary data into a signed distance func-

tion. The latter is interpolated to a first mesh of uniform size ( i.e.,

10 μm) of the FE ROI using trilinear FE interpolation ( i.e., the im-

ge is considered as a hexahedral grid where the voxels are nodes).

he resulting signed distance function is then regularized based

n a recently proposed parallel reinitialization algorithm ( Shakoor

t al., 2015b ), and used as an intermediary to locate the inter-

ace ( Shakoor et al., 2017 ). This mesh generation step is combined

ith an adaption step taking into account the local maximum cur-

ature of the interface ( Shakoor et al., 2015a ). These different steps

re summarized in Fig. 4 .

Parameters of the final mesh are defined to have a size of

0 μm close to matrix/void interfaces, and 50 μm at a distance of

00 μm from any interface, the transition in this layer being linear.

t can be observed qualitatively in Fig. 4 that the FE approximation

f the geometry is really close to experimental data. A more thor-

ugh analysis of the sensitivity of the results to meshing parame-

ers was assessed (not reported herein) and revealed this influence

o be negligible, as already validated in a previous study for a dif-

erent experiment on the same material ( Buljac et al., 2017 ).

.3. Constitutive laws and parameters

.3.1. Microscale model

The microscale simulations need material models for the

raphite nodules and the ferritic matrix. As stated in the intro-

uction, graphite nodules are modeled as voids ( Dong et al., 1997;

hang et al., 1999; Bonora and Ruggiero, 2005; Hütter et al., 2015 ),

hile the ferritic matrix is considered as elastoplastic with power

aw hardening

0 ( ε ) = σy + K ( ε ) n (1)

here ε is the equivalent (von Mises) plastic strain, σ y the yield

tress, K the plastic consistency and n the hardening exponent. A

articularity of the present numerical method is that voids are

eshed and defined as a purely elastic material with very low

oung’s modulus with respect to the matrix. A sensitivity analy-

is regarding this Young’s modulus was conducted ( Buljac et al.,

017 ) showing that a ratio of 1,0 0 0 between this modulus and

hat of the matrix was sufficient. The properties of the matrix are

educed from stress/strain curves ( Zhang et al., 1999 ) and pre-

ented in Table 1 . These data are based on tensile experiments

n a purely ferritic material. Thus, it is possible that the ac-

ual behavior of the matrix be more complex. An interesting per-

pective to the present work would be to use micromechanical

M. Shakoor et al. / International Journal of Solids and Structures 112 (2017) 83–96 87

Fig. 4. Image immersion and meshing. (a) Initial laminography 2D section, (b) signed distance function computed thanks to image processing, (c) signed distance function

interpolated and reinitialized on the FE mesh ( Shakoor et al., 2015b ), (d) conforming FE mesh generated and adapted to interfaces and local maximum curvature, (e) zoom

on the FE mesh, (f) comparison between initial laminography 2D section and interfaces in the final FE mesh (in white).

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alculations and local error measurements to study more appropri-

te microscale constitutive models and identify more realistic ma-

erial properties.

Mechanical solution to the equilibrium equations is based on

mixed velocity-pressure formulation solved using a P 1 + /P 1 el-

ment to avoid locking issues ( Boffi et al., 2008 ). The nonlinear

ehavior of the matrix requires a Newton–Raphson scheme both

or numerical integration of the plastic law and global equilib-

ium ( Wagoner and Chenot, 2001 ). In order to handle large de-

ormations, an updated Lagrangian scheme is used and the veloc-

ty field resulting from the mechanical solution is applied to move

esh nodes. An advanced mesh motion technique is necessary and

lays a key role in micromechanical simulations. Since the latter

ill be driven herein by boundary conditions reproducing as ac-

urately as possible what is observed in Fig. 3 , large distortions

ill occur inside the ROI. As a result, mesh quality will deterio-

ate, thus affecting the accuracy of FE solutions. Flip of elements

ay also occur. These issues are handled herein thanks to an au-

omatic mesh motion and adaption method developed in previous

orks ( Shakoor et al., 2015b; 2015a; 2017 ). This method was de-

igned to preserve at best the local distribution of volumes be-

ween the matrix and the voids while handling large void growth

nd complex topological events such as void coalescence.

.3.2. Macroscale model

Both wFE and sFE methods require the identification of mate-

ial parameters { p } based on experimental measurements acquired

uring test (A). In the case of wFE, these material parameters are

sed in the whole domain for the calculation at the specimen scale

f test (B). In the case of sFE, these parameters are only used for

he homogenized out-of-ROI material, while microscale material

arameters ( Table 1 ) are used inside the ROI. The identification of

aterial parameters can be performed only using force measure-

ents, or using both force measurements and 2D surface images.

he two approaches are detailed hereafter.

Load data

A first identification is conducted using standard global opti-

ization methods ( Storn and Price, 1997; Grédiac and Hild, 2012 ),

here the objective function is defined as ( Roux and Bouchard,

015 )

88 M. Shakoor et al. / International Journal of Solids and Structures 112 (2017) 83–96

Fig. 5. Identification of macroscopic material parameters based on test (A) data. (a) Mesh used for specimen scale calculations, (b) comparison between the simulated

force/displacement curves and experimental data.

Table 2

Elastoplastic properties of nodular graphite cast iron obtained using two inverse

analyses on test (A). The first one is only based on load data. Integrated-DIC also

uses kinematic data.

Method E (GPa) ν σ y (MPa) K (MPa) n

Load data 187 0 .28 64 520 0 .19

Integrated-DIC 136 0 .28 220 410 0 .33

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E( F ( u )) =

√ ∫ U 0

(F ( u ) − F exp ( u )

)2 d u √ ∫ U

0 F exp ( u ) 2 d u

(2)

where the forces obtained in an FE simulation F are compared with

the sum of reaction forces measured during the experiment F exp ,

with an integral on the loading path [ 0 , U ] . Therefore, only load

data acquired during test (A) are taken into account. Note that

the displacement u , and all displacements mentioned in this pa-

per, are measured via 2D DIC ( Besnard et al., 2006; Hild and Roux,

2012; Kahziz et al., 2016 ). These values were determined directly at

the upper and lower parts of the mesh presented in Fig. 5 (a). The

force/displacement curve based on measurements and simulations

is shown in Fig. 5 (b). The experimental data were acquired up to

final failure, but the last loading steps of test (B) were discarded

for this inverse analysis because no micromechanical simulation is

carried out up to final failure in Section 4 . The force/displacement

curves show that the numerical approximation is very good, but

here only the macroscopic force is compared. Local measurements

are considered in the sequel.

Load data and pictures

In order to obtain more realistic material properties more ad-

vanced identification techniques can be considered. In Integrated-

DIC or DVC ( Mathieu et al., 2015; Neggers et al., 2015; Hild et al.,

2016 ), both pictures ( i.e., 2D surface images from test (A)) and load

( i.e., force measurements from test (A)) data are taken into account.

Integrated DIC requires displacement fields to be fully mechani-

cally admissible, i.e., satisfy equilibrium for the chosen constitutive

law. Since material parameters are sought, the corresponding sen-

sitivity fields ( Tarantola, 1987 ) are also needed. The displacement

fields are parameterized with the sought corrections { δp } to the

current material parameters { ̂ p } over the whole loading history t

u ( x , t, { p } ) =

∑

t

∑

p

u p ({ p } , t) �p ( x ) (3)

where the kinematic degrees of freedom u p are linked to the

sought material parameters via sensitivities

u p ({ p } , t) = u p ({ ̂ p } , t) +

{

δu p

δ{ p } ({ ̂ p } , t) } T

{ δp } . (4)

The kinematic sensitivities are collected in matrix [ S u (t)] and eval-

uated for each loading step t via FE simulations in which measured

displacements are prescribed on the top and bottom boundaries of

he considered ROI. These Dirichlet boundary conditions are mea-

ured a priori by conducting global DIC on 2D images from test (A).

imultaneously, the load measurements gathered in vector { F exp }re compared with the resultant forces { F F E } from the correspond-

ng FE simulation. As for the kinematic part, the load sensitivity

atrix [ S F (t)] to the sought material parameters is computed to

pdate { δp } from the current estimate { F F E ({ ̂ p } ) } of the reaction

orces. Since both images and loads are used in a single approach

nd due to different physical natures of the data a Bayesian frame-

ork is considered herein, in which each piece of data is weighted

y its variance and covariance with all the other data ( Mathieu

t al., 2015; Hild et al., 2016 ).

The Integrated-DIC code used herein is an in-house Matlab im-

lementation with C ++ kernels ( i.e., Correli 3.0 ( Leclerc et al.,

015 )) while the accompanying FE simulations are performed us-

ng the commercial package Abaqus/Standard. More details about

echanical correlation can be found in ( Mathieu et al., 2015 ).

ue to very low sensitivity the Poisson’s ratio had to be set in

ntegrated-DIC to its initial value of 0.28 while the other elasto-

lastic parameters have been calibrated.

Material parameters obtained using both identification proce-

ures are given in Table 2 . Though these parameters may seem to

iffer notably, the influence of this difference is significant only at

ow ( i.e., < 0.2) and very large ( i.e., > 1) equivalent plastic strain,

he latter not being experienced herein. Hence, the results are not

xpected to strongly depend on the choice of identification proce-

ure apart from the first loading steps, as illustrated in Fig. 5 (b).

. Boundary conditions

As explained in Section 1 , three different techniques will be

sed to determine the boundary conditions of FE simulations at

he microscopic scale. Two procedures, namely, wFE and DVC-

E, consist of computing or measuring displacement fields and

hen applying them as Dirichlet boundary conditions on the whole

oundary of the FE mesh of the ROI of the microscale calculations.

M. Shakoor et al. / International Journal of Solids and Structures 112 (2017) 83–96 89

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t

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d

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t

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he third procedure ( i.e., sFE) has the ROI mesh embedded within

he sample geometry and the whole computation is run in a single

tep. The three approaches and their implementation are detailed

ereafter.

.1. Weak Finite Element (wFE) technique

An application of the wFE technique to ductile fracture with

eshed microstructure can be found in ( Kaye et al., 2013 ). The

bjective of this study was to assess the influence of macro-

copic loading conditions on the microstructure, especially regard-

ng damage localization around inclusions. Although the real mi-

rostructure was meshed from a 3D X-ray image of the material,

here was no comparison with experiments regarding microme-

hanical calculations, as proposed herein.

In the present work, a first macroscopic simulation of test (B) is

onducted at the specimen scale. The material is considered as ho-

ogeneous and modeled using the same elastoplastic model with

he power law hardening defined in Eq. (1) and the two sets of

aterial parameters given in Table 2 . Since specimens used in tests

A) and (B) had identical geometries, the same mesh as in Fig. 5 (a)

s used. The displacement field between each consecutive loading

teps where 3D X-ray scans were acquired is stored in the refer-

nce configuration. These displacements are then interpolated at

he boundaries of the ROI during a second FE simulation at the

icroscale by means of linear interpolation (both meshes being ex-

lusively composed of tetrahedra).

.2. Strong Finite Element (sFE) technique

In the sFE method, a single FE simulation is carried out. The

icrostructure is directly meshed within the specimen mesh, with

progressive mesh coarsening from the very fine mesh close to

he microstructure to the coarse mesh out of the ROI. Such mesh

an be quite complex to build with conventional meshing tools,

specially when a real microstructure is considered. This approach

as been used for small compact tension specimens ( Tian et al.,

010; Tang et al., 2013; O’Keeffe et al., 2015 ). The microstructure

f the studied high strength steel featured two populations of par-

icles and voids, namely, one of micrometer size and one of sub-

icrometer size. In these studies, only the major population was

eshed and simulated with the sFE method. Promising results

ere obtained by using a damage model inside the matrix mate-

ial in order to account for sub-micrometer size voids. However, in

he absence of local error measurements and in-situ experiment, it

as not possible to quantify locally the accuracy of this microme-

hanical model.

In order to avoid the difficulties linked to real microstructures

nd their randomness, experiments on tensile specimens with ma-

hined micrometer size holes were proposed ( Weck and Wilkin-

on, 2008 ). Simulations based on this experimental procedure were

erformed ( Hosokawa et al., 2013b; Alinaghian et al., 2014 ). Be-

ause the position of the holes and their size was part of the spec-

men geometry, generating meshes adapted to the microstructure

as simplified. However, these studies revealed an influence of a

inor void population, which could not be meshed.

The specimen geometry with two machined holes used in the

resent work is directly inspired from ( Weck and Wilkinson, 2008 ).

hanks to the millimeter size of the machined holes, and the mi-

rometer resolution of synchrotron laminography, it is proposed

erein to mesh both void populations, namely, the machined holes

nd the nodules (considered as voids herein). This is illustrated in

ig. 6 . This mesh is similar to that shown in Fig. 5 (a), with the dif-

erence that this time it incorporates the ROI with its microstruc-

ure. In the ROI, the material properties correspond to those of

able 1 , while in the rest of the specimen the two sets of mate-

ial properties given in Table 2 are considered. Compared to other

pproaches, this method adds significant computation time, as il-

ustrated by the used mesh of ≈ 1.5 million elements in Fig. 6 .

owever, it has the advantage that ductile fracture can be studied

imultaneously at two scales.

.3. Digital Volume Correlation - Finite Element (DVC-FE) technique

DVC used herein is an extension of global 2D DIC ( Besnard

t al., 2006; Hild and Roux, 2012 ). The reconstructed volume is

epresented by a discrete gray level field of the spatial (voxel) co-

rdinate x . The principle of DVC lies in matching the gray levels f

n the reference configuration with those of the deformed volume

such that their conservation is obtained

f ( x ) = g[ x + u ( x )] (5)

here u is the displacement field with respect to the reference

olume. In real experiments the strict conservation of gray lev-

ls is not satisfied, especially in laminography where deviations

ppear not just due to acquisition noise but also due to recon-

truction artifacts because of missing information ( Xu et al., 2012 ).

onsequently, the solution consists in minimizing the gray level

esidual ρ( x ) = f ( x ) − g[ x + u ( x )] by considering its L2-norm with

espect to kinematic unknowns associated with the parameteriza-

ion of the displacement field. Since a global approach is used in

his work, the whole ROI is considered, the global residual �2 c

2 c =

∑

ROI

ρ2 ( x ) (6)

s minimized with respect to the unknown degrees of freedom u p ,

he displacement field being written as

( x ) =

∑

p

u p �p ( x ) (7)

here �p ( x ) are the chosen displacement fields for the parameter-

zation of u ( x ) . Among a whole range of available fields, finite ele-

ent shape functions are particularly attractive because of the link

hey provide between the measurement of the displacement field

nd numerical models. Thus, a weak formulation based on hexahe-

ral finite elements with trilinear shape functions is chosen ( Roux

t al., 2008 ).

Conducting DVC analyses with the full size reconstructed vol-

mes is computationally too demanding. Therefore only a part of

he reconstructed volume called DVC ROI is considered herein, as

hown in Fig. 2 . Additionally, to be able to keep large DVC ROI

izes, the original reconstructed volumes are a priori coarsened, i.e.,

ach 8 neighboring voxels are averaged to form one supervoxel. By

oing this, file sizes are decreased by a factor of 8.

The DVC resolution is evaluated by correlating two scans of the

nloaded sample (0) with (denoted “rbm”) and without (denoted

bis”) a rigid body motion (RBM) applied between the acquisitions.

ue to the noise contribution and reconstruction artifacts, these

wo volumes are not identical. Therefore, the measured displace-

ent field accounts for the cumulated effects of laminography and

VC on the measurement uncertainty ( Morgeneyer et al., 2013 ).

he uncertainty values are evaluated by the standard deviation of

easured displacement fields. Figure 7 shows the standard resolu-

ion levels for different element sizes expressed in supervoxels.

ecreasing the element size is followed by an increase of the dis-

lacement resolution ( Besnard et al., 2006; Avril et al., 2008; Hild

nd Roux, 2012; Leclerc et al., 2012 ). The element size used in this

ork is = 16 supervoxels (length) for all three directions, which

ields a standard displacement resolution of 0.25 supervoxel. This

alue represents the limit below which the estimated displacement

evels are not trustworthy.

90 M. Shakoor et al. / International Journal of Solids and Structures 112 (2017) 83–96

Fig. 6. Inside view of the mesh used in sFE calculations. (a) Full specimen, (b) zoom on the ROI. The three shades of blue represent, from lighter to darker, the homogenized

out-of-ROI material, the matrix, and the voids. (For interpretation of the references to color in this figure legend, the reader is referred to the web version of this article.)

Fig. 7. Standard displacement resolutions as functions of the element size ex-

pressed in supervoxels for two different scans.

p

w

m

s

4

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a

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r

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≈

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Successful DVC registrations have been obtained for the first 9

incremental ( i.e., between Step n − 1 and Step n ) loading steps and

they have been used for DVC-FE boundary conditions. The DVC dis-

placement fields are interpolated at each loading step onto the FE

mesh of the ROI using trilinear interpolation ( i.e., identical to shape

functions of DVC measurements).

4. Results

In this section, results using the DVC-FE approach are first pre-

sented and discussed, as comparisons with experiments are quali-

tatively and quantitatively possible. Then, comparisons with other

approaches are proposed.

All micromechanical simulations (wFE, sFE, and DVC-FE) were

performed on a cluster of two nodes with a 1.2 GHz Intel Xeon

20-core processor and 64GB RAM each. These simulations included

≈ 100 voids meshed with ≈ 1 million elements ( ≈ 1.5 million el-

ements for the sFE simulation). The computation time remained

close to one hour even for the sFE simulation. This quite low com-

putation time is very promising in the perspective of applying in-

verse analyses at the microscale to identify micromechanical prop-

erties, especially regarding coalescence modeling. This is helped by

the fact that the three approaches avoid the requirement of mesh-

ing the microstructure of the whole specimen, as in DNS. In the

present case, considering the ratio between the volume of the out-

of-ROI material and the volume of the ROI in Fig. 6 , the number

of elements that a DNS calculation would require can be estimated

at ≈ 100 million elements. Although such calculation is not con-

ducted herein, the aforementioned computation time shows that it

could be in the very near future, at least regarding the numerical

part, thanks to the proposed methods and High Performance Com-

uting (HPC) capabilities demonstrated in this paper. However, the

FE, sFE, and DVC-FE approaches are, by far, more efficient, as the

icrostructure has to be modeled only inside the ROI in the FE

imulations.

.1. DVC-FE coupling

The results using the DVC-FE approach are shown in Fig. 8 ,

here both void growth and plastic strain are observed. Errors

re be assessed first qualitatively, for example by comparing the

midsection of the ROI with experimental images, as shown in

ig. 9 . Since DVC boundary conditions are expected to follow ex-

erimental images, the matrix/void interfaces in the simulation

in white in the figure) can be superimposed on these images

nd compared. This figure reveals that interfaces are overall very

ccurately meshed and tracked during the simulation, up to the

ast loading step. However, there is an irregularity in the mate-

ial, namely a non spherical void in the top left region of Fig. 9 .

n the undeformed state, this defect is already poorly captured by

he meshing technique due to its very small size, and this error ac-

umulates during loading. The same remark applies to small voids

hat can be seen in the experimental image in Fig. 9 (a), but not in

he numerical approximation. This figure shows that void growth

n the simulations compares well with what is observed in X-ray

mages.

For each pair of consecutive loading steps, the scan acquired

or the second step can be deformed back with the displacement

eld obtained by the ROI calculation and the result can be com-

ared with the scan acquired at the first step. That is, by means of

newly developed tetrahedral-DVC code ( Leclerc et al., 2015; Hild

t al., 2016 ) FE solutions with corresponding tetrahedral meshes

re directly imported in the reconstructed volumes frame where

he displacement results are interpolated voxel-wise and the cor-

esponding deformed volume g( x ) is corrected by the computed

isplacement field u F E ( x ) . The gray level residuals, i.e., differences

etween the reference volume f ( x ) and corrected deformed vol-

me g( x + u ( x )) can then be compared for DVC and FE compu-

ations, enabling quantitative and local error measurements. This

s shown in Fig. 10 as standard deviation of residual fields (nor-

alized by dynamic range of the volume, i.e., 256 gray levels) for

VC and DVC-FE calculations. These residuals remain very close to

hose observed in the resolution analysis for which no strains oc-

urred. Therefore the DVC results are deemed trustworthy.

Note that the DVC-FE curve is not expected to lie below the

VC one, since the latter is used to drive the former. The er-

or produced by micromechanical models inside the DVC-FE do-

ain is low and it slightly increases at late loading steps (from

15% initially to ≈ 20% in late loading steps). A look at the Z

idsection in Fig. 11 (a and b) reveals that these differences be-

M. Shakoor et al. / International Journal of Solids and Structures 112 (2017) 83–96 91

Fig. 8. ROI calculation results using the DVC-FE approach showing the 3D meshed voids and the equivalent plastic strain on sections when: (a) u = 0 (undeformed state),

(b) u = 83 μm, (c) u = 192 μm, (d) u = 321 μm.

Fig. 9. ROI (blue line) calculation results using the DVC-FE approach comparing the

numerical matrix/void interface (white line) with experimental images for the X

midsection when: (a) u = 0 (undeformed state), (b) u = 83 μm, (c) u = 192 μm,

(d) u = 321 μm. (For interpretation of the references to color in this figure legend,

the reader is referred to the web version of this article.)

t

t

t

s

b

n

Fig. 10. Standard deviation for the dimensionless gray level residual fields for all

loading steps. For comparison purposes, the dashed line corresponds to the resolu-

tion analysis for the so-called “bis” case (see Section 3.3 ).

e

s

w

a

h

i

e

i

w

e

fi

a

fi

p

d

e

s

ween simulations (DVC-FE) and experiments (DVC) are concen-

rated around interfaces. This is expected due to plastic localiza-

ion. The growth of a minor void population and damage at a lower

cale are certainly also an explanation. These minor voids cannot

e observed due to the resolution of the synchrotron imaging tech-

ique used herein. In ( Hütter et al.; 2015 ) they have been observed

xperimentally using scanning electron microscopy of fractured

urfaces for a similar material. Once the minor voids are imaged

ith sufficient resolution, the present DVC-FE framework would

llow to model them. Hence, the development and application of

igher resolution techniques, such as nanolaminography, would be

nteresting in order to check these assumptions. Yet, the present

rror measurements already serve as a basis for modeling and

dentification of more advanced plasticity and damage models,

hich will be considered in future work.

Once the absolute errors in terms of gray level residuals are

stimated, relative comparisons can be shown. DVC displacement

elds are applied to the boundaries of the FE domain, but they

re also available inside the domain. Hence, DVC and FE kinematic

elds can also be interpolated on the same mesh and directly com-

ared as shown in Fig. 12 (a). Again, it is confirmed that the main

ifferences are concentrated in the debond zones, while the differ-

nces close to the boundaries are mostly zero.

Although the results using the DVC-FE method are promising,

everal aspects remain to be improved. In particular, Fig. 9 as qual-

92 M. Shakoor et al. / International Journal of Solids and Structures 112 (2017) 83–96

(a) DVC (b) DVC-FE

(c) wFE (d) sFE

Fig. 11. Absolute gray level differences at the Z midsection after correction with

DVC (a), DVC-FE (b), wFE (c) and sFE (d) displacements for the ninth loading step.

Note that due to different rigid body motions for each simulation, the Z midsections

slightly differ.

(a) (b)

Fig. 12. Mid-section normal to z -direction showing absolute difference between:

(a) DVC and DVC-FE displacement fields, (b) DVC and wFE displacement fields for

the ninth loading step. The black area at the top of sub-figure (b) represents a zone

out of the DVC ROI. The color bar range is set according to the data set from sub-

figure (a) in order to have more convenient visual comparison. The displacement

difference is expressed in supervoxels (1 supervoxel ←→ 2 . 2 μm). (For interpreta-

tion of the references to color in this figure legend, the reader is referred to the

web version of this article.)

Fig. 13. Void volume change curves for all the approaches investigated in this pa-

per.

c

a

T

m

c

a

T

m

a

m

a

l

s

a

v

m

4

i

p

t

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2

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s

p

t

4

w

c

p

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w

t

t

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t

c

s

i

itative and Fig. 11 (a and b) as quantitative comparisons show that

there still is a significant gap between DVC-FE and DVC results.

This gap increases slowly by reaching the ultimate loading steps.

Similarly, the displacement difference is significantly larger than

the displacement resolution (see Fig. 7 ). Locally, the differences

are mainly concentrated around debond areas ( i.e., matrix-nodule

interfaces). It is hence important to consider more carefully the

nodules and their impact both on the load carrying capacity of

the material, and void growth. Additionally, the increase of the er-

ror at late loading steps indicates the inability of the microscale

onstitutive models used for the matrix material to capture the

cceleration of void growth, and the subsequent void coalescence.

his observation calls for more advanced plasticity models at the

icroscale that can capture the complex multiscale plastic flow lo-

alization in the matrix. The growth of a minor void population

nd damage at a lower scale could also be a possible explanation.

hus, additional material parameters should be introduced for the

atrix material, and the present procedure should be extended to

llow for the identification of these parameters.

These developments will extensively rely on the DVC-FE

ethod and its ability to provide experimentally measured bound-

ry conditions for micromechanical simulations, and then compute

ocal and relevant error estimators to assess the validity of these

imulations. The extension of the Integrated-DIC framework to 4D

nalyses ( Hild et al., 2016 ) will also be considered to conduct in-

erse analyses based on these error measurements and identify

aterial parameters at the microscale.

.2. Comparisons with wFE and sFE results

For quantitative comparisons between the DVC-FE method and

ts two alternatives used in this paper, two approaches are pro-

osed. In the first approach, as commonly carried out by most au-

hors, void growth curves are compared, giving only average quan-

ities and global error indicators ( Babout et al., 2004; Landron

t al., 2010; Fansi et al., 2013; Cao et al., 2014; Seo et al., 2015b;

015a ). The second approach aims to study which method is closer

o experimental observations based on local differences. These lo-

al errors are computed by reducing the kinematic data for each

imulation back to gray level residuals. This requires to extend the

rocedure already applied to DVC and DVC-FE kinematic data to

he wFE and sFE methods.

.2.1. Global error indicators

Void growth is defined by the following relationships

f =

void volume

ROI volume , void growth =

f

f 0 (8)

here f 0 denotes the initial void volume fraction. Void growth

urves are shown in Fig. 13 . The EXP curve is obtained by com-

uting void growth in processed laminography images ( i.e., images

ith smooth signed distance functions as in Fig. 4 (b)). The 3D box

here this experimental void growth is computed remains fixed

o the initial ROI. The wFE and sFE curves correspond to simula-

ions using material properties based only on load data (first line

f Table 2 ), while the wFE I-DIC and sFE I-DIC curves correspond

o simulations using material properties based on Integrated-DIC

alibration (second line of Table 2 ). Many of the numerical results

how an important decrease of the porosity f at the first load-

ng step, which is not observed for the experimental curve. This

M. Shakoor et al. / International Journal of Solids and Structures 112 (2017) 83–96 93

i

u

b

f

w

v

p

m

i

R

i

t

f

t

g

c

i

d

c

i

o

p

p

p

a

a

a

t

t

b

c

4

i

f

i

m

p

s

m

l

t

s

w

t

t

a

a

u

a

m

w

n

v

o

g

u

f

t

e

c

Fig. 14. Standard deviation for the dimensionless gray level residual fields for all

loading steps and investigated cases.

g

t

F

g

l

l

i

t

i

s

m

p

r

t

w

t

r

a

m

i

b

w

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S

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b

i

D

i

o

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r

o

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m

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s due to the fact that remeshing is extensively used for all sim-

lations. This remeshing has the consequence that interfaces can

e slightly smoothened, and void volume can be diffused. Apart

rom this numerical issue, the curves reveal significant void growth

ith increasing load for DVC-FE and sFE results, while nearly no

oid growth occurs in the wFE simulation. This effect can be ex-

lained by the fact that the computation used to obtain displace-

ent boundary conditions for the wFE simulation does not take

nto account damage and the subsequent volume change in the

OI. Hence, the displacement fields that are transferred to the ROI

n this case are incompressible in the plastic regime, and neither

he void volume nor the ROI volume can evolve. This effect calls

or more advanced models ( i.e., including damage) to be used at

he macroscopic scale.

Regarding comparisons with the experimental curve, void

rowth seems to be overestimated with the DVC-FE method. This

an be explained by the fact that nodules are considered as voids

n the simulations, while in reality only the voids nucleated after

ebonding of nodule/matrix interfaces grow (see Fig. 9 ). Thus, it

an be assumed that all curves would have a slightly lower slope

f nodules were taken into account. The DVC-FE method is the only

ne that shows a void growth similar to that observed in the ex-

eriment.

The comparison between DVC-FE and sFE results using the pro-

osed procedure shows that although void growth and the com-

ressibility effects induced by the presence of voids are taken into

ccount in the sFE method, this void growth is not as significant

s with the DVC-FE method. In particular, the slopes of the wFE

nd sFE curves are clearly lower than what is observed experimen-

ally. Both results are not improved when using Integrated-DIC ma-

erial parameters. The influence of material parameters seems to

e negligible in the present case. This observation remains to be

onfirmed with local error measurements.

.2.2. Local error indicators

In order to compare wFE and sFE kinematic fields with X-ray

mages, special care is taken to subtract the rigid body motions

rom wFE/sFE calculated displacement fields followed by apply-

ng the corresponding rigid body motions measured by DVC. First,

ean deformation gradients over the ROI for DVC and FE dis-

lacement fields are calculated. By employing a polar decompo-

ition on the latter, FE and DVC rotations are evaluated while the

ean values of kinematic fields represent the corresponding trans-

ations. From the initial FE displacement solutions then are sub-

racted mean FE translations and rotations and added the corre-

ponding mean DVC translations and rotations. Hence, from the

FE/sFE displacements fields, rigid body motions originating from

he FE simulations are first eliminated, and then rigid body mo-

ions associated with the experiment ( i.e., measured with DVC) are

pplied. This is performed in order to have equal conditions for

ll presented methods when reducing them to gray level resid-

al images. The resulting errors are shown in Figs. 11 (c and d)

nd 14 while the displacement difference between DVC measure-

ents and wFE calculated fields is shown in Fig. 12 (b).

Surprisingly, although sFE yields slightly better results than

FE, independently of the material parameters, this difference is

egligible. This means that although the sFE method predicts a

oid growth that is globally closer to experimental data, the shape

f these voids is inexact. Therefore, errors that do not appear in

lobal measurements are revealed by local measurements, thereby

nderlining the interest of the present methodology. The small dif-

erence between wFE and sFE results could mean that the constitu-

ive model used at the macroscale corresponds well to the homog-

nized mechanical response obtained in the ROI, where microscale

onstitutive models are used.

Regarding the comparison between sFE/wFE and DVC-FE, the

ap is increasing during the load history, ending with a deviation

hat is twice higher with the sFE/wFE methods than with DVC-

E. This confirms the tendency that was observed based on void

rowth curves. To have a more precise idea of this gap, the cumu-

ative gray level residuals distribution is shown in Fig. 15 (a) for the

ast loading step. This difference does not seem important because

t takes into account a large number of supervoxels belonging to

he matrix, where contrast is low. This observation is illustrated

n Fig. 15 (b), where only gray level residuals higher than 40 are

hown, hence discarding most supervoxels of the matrix. The cu-

ulative distribution taking into account only the remaining su-

ervoxels ( i.e., close to voids or nodules) is shown in Fig. 15 (c). It

eveals an important difference between the DVC-FE results and

hose obtained with the sFE/wFE methods. The residual is doubled

ith the latter. This underlines the inability of macro simulations

o precisely describe all the micro localization phenomena occur-

ing in the ROI both between the machined holes and the nodules,

nd between the nodules themselves with the chosen constitutive

odels. There are nearly zero differences in the center of the ROI

n Fig. 12 (b), which indicates that the position of the macroshear

and ( i.e., between the machined holes) is properly captured by

FE simulations, as well as the kinematics within that band. In

he same figure, more significant differences are observed in the

est of the ROI compared with DVC-FE simulations.

Overall, the use of Integrated-DIC material parameters for wFE

nd sFE methods slightly improves the results. As expected in

ection 2.3.2 , this gain is concentrated at lower strains while at

igher strains the influence of these material parameters is negligi-

le. Paint cracking and subsequent DIC convergence issues from 2D

mages (test (A)) could be responsible for decreasing Integrated-

IC performance at late loading steps. This observation adds to the

nterest for the DVC method, which relies directly on the contrast

f the material. Thus, the way the material parameters are identi-

ed herein improves the results but only slightly.

Using more complex macroscale constitutive models would cer-

ainly improve the results obtained using the wFE and sFE meth-

ds. This would also increase the need for identification methods

uch as Integrated-DIC that take into account field measurements.

or instance, the inability of both wFE and sFE methods to cor-

ectly predict plastic localization could be linked to an anisotropy

f the yield surface at the macroscale, due to the presence of the

odules. Damage and the subsequent volume changes of the ho-

ogenized material should also be considered, due to debonding

f the nodules from the matrix. These remarks stress out the im-

ortance of understanding the influence of the microstructure on

he mechanical response of the homogenized material. Therefore,

numerical validation procedure for micromechanical simulations

hat does not rely on any macroscopic constitutive model, such as

he DVC-FE method, is of great interest.

94 M. Shakoor et al. / International Journal of Solids and Structures 112 (2017) 83–96

(a) (b) (c)

Fig. 15. Cumulative distribution of gray level residuals at the last incremental loading step considering to (a) all supervoxels, (c) only supervoxels with gray level residuals

higher than 40. In (b), the absolute gray level differences for the sFE method with residuals higher than 40 are shown for the same Z midsection as Fig. 11 (d).

o

c

m

s

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i

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c

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I

a

5. Conclusion

The present paper discusses different choices of boundary con-

ditions for micromechanical Finite Element (FE) simulations based

on 3D X-ray data. Because it would require meshes of more than

100 million elements, Direct Numerical Simulation (DNS) is dis-

carded and methods allowing the microstructure to be meshed

only in a small Region of Interest (ROI), instead of the whole spec-

imen, are considered. The accuracy of these methods is analyzed

based on global and local error estimators relative to the 3D X-

ray data acquired at consecutive loading steps for the same mi-

crostructure. The latter corresponds to new experimental results on

a nodular graphite cast iron specimen with two machined holes.

Due to the 45 ° alignment of the machined holes with respect to

the tensile direction, a macroscopic shear band develops between

the holes. Within that macroscopic shear band, ductile fracture due

to the nucleation ( i.e., by debonding of the nodules), growth and

coalescence of microscopic voids is observed.

• In an earlier work ( Buljac et al., 2017 ), it was proposed to mea-

sure boundary conditions for micromechanical simulations via

Digital Volume Correlation (DVC) directly from consecutive 3D

X-ray images. Promising results were obtained for a tensile ex-

periment on a specimen with a central hole. The results pre-

sented herein show that this DVC-FE method remains very ac-

curate with a shear band traversing the ROI ( Fig. 8 ). Global gray

level residuals indicate an increase of the error at late loading

steps, which is due to void coalescence ( Fig. 13 ). Local residuals

show that these errors are located close to the debond areas

around the nodules, hence calling for more relevant microme-

chanical models.

• A second method, which is referred to as weak FE (wFE), con-

sists of first conducting an FE simulation at the specimen scale,

considering a totally homogeneous material, and then using the

displacement fields from this first simulation to drive a second

simulation at the ROI scale. It is proposed to identify material

parameters for the specimen scale simulation based only on

load data, and then on both load data and 2D surface images.

Independently of the identification method, results show that

the wFE method leads to a significant underestimation of void

growth, as there is no global void volume change in wFE results

( Fig. 13 ). The slight improvement of the results when using a

material parameter identification method that also takes into

account 2D surface images indicates that the investigation of

more relevant constitutive models for the specimen scale sim-

ulation is worth considering.

• Another method, which is coined strong FE (sFE), consists of

embedding the ROI mesh in the specimen mesh. Macroscale

constitutive models are used for the out-of-ROI material where

the microstructure is not meshed, while microscale constitu-

tive models are used inside the ROI where it is meshed. Void

volume change curves indicate a significantly increased void

growth compared to wFE results. However, local residuals in-

dicate no significant improvement. This proves that while voids

in the sFE simulation grow at a rate that is closer to the ex-

perimental observations than in wFE results, the shape of these

voids is not accurately predicted. Once again, a slight improve-

ment of the results is observed in the first loading steps when

using a material parameters identification method that also

takes into account 2D surface images.

As a conclusion, the dependence of both sFE and wFE methods

n a specimen scale simulation and a corresponding macroscale

onstitutive model constitute their main limitation. Although the

acroscale constitutive model used herein could be improved,

uch task is not obvious, especially as large plastic strains and

omplex damage phenomena are observed locally. This limitation

f both sFE and wFE methods is also the main advantage of the

VC-FE method since DVC measurements avoid the use of a speci-

en scale simulation and a corresponding macroscale constitutive

odel.

It is worth mentioning that since the DVC-FE method requires

n-situ 3D X-ray imaging, it is hardly applicable to thick specimens

nd industrial applications. Future work will hence lean toward the

se of this DVC-FE method for small specimens in order to study

ore relevant micromechanical models and calibrate material pa-

ameters for these models. These enhanced micromechanical mod-

ls could then be the basis for the investigation of more relevant

onstitutive models to be used at the macroscale in, say, wFE and

FE methods. Such approach is likely to be applicable to other ma-

erials where micromechanical modeling is gaining an increasing

nterest, such as metallic foams and composite materials for in-

tance.

cknowledgments

This work was performed within the COMINSIDE project funded

y the French Agence Nationale de la Recherche ( ANR-14-CE07-

034-02 grant). The authors also acknowledge the European Syn-

hrotron Radiation Facility for provision of beamtime at beamline

D15, experiment ME 1366. It is also a pleasure to acknowledge

he support of BPI France (“DICCIT” project), and of the Carnot

.I.N.E.S institute (“CORTEX” project). M. Kuna and L. Zybell from

MFD, TU Freiberg are thanked for materials supply and machining

s well as for scientific discussions.

M. Shakoor et al. / International Journal of Solids and Structures 112 (2017) 83–96 95

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