Composites Science and Technology 198 (2020) 108238

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Composites Science and Technology
journal homepage: http://www.elsevier.com/locate/compscitech

Experimental and numerical investigations on push-out delamination in
drilling of composite laminates
R. Higuchi a, *, S. Warabi b, W. Ishibashi b, T. Okabe b, cabc

Department of Aeronautics and Astronautics, The University of Tokyo, 7-3-1 Hongo, Bunkyo-ku, Tokyo, 113-8656, Japan
Department of Aerospace Engineering, Tohoku University, 6-6-01, Aramaki-Aza-Aoba, Aoba-ku, Sendai, Miyagi, 980-8579, Japan
Department of Materials Science and Engineering, University of Washington, 302 Roberts Hall, Seattle, WA, 98195, USA

A R T I C L E I N F O

A B S T R A C T

Keywords:
Drilling
Delamination
Damage mechanics
Finite element analysis (FEA)
Non-destructive testingThis study conducts experimental and numerical studies on the progressive damage occurring during the drillingof composite laminates. In particular, this work focuses on push-out delamination, which occurs at the interface
around the drill-exit and is the most critical damage suffered during the drilling process. To investigate thedamage progression mechanisms, penetration and interruption drilling tests are performed on composite laminate consisting of quasi-isotropic CFRP plies and fabric GFRP ply (bottom side). After drilling tests, the damage
evolution is evaluated using X-ray computed tomography and optical microscopy. Based on the experimental
results, a simplified simulation model is established, and damage progression simulation is performed using an
explicit dynamic finite element method. The results show that the bending deformation in the bottom two plies
triggers the propagation of push-out delamination. Therefore, the extent of delamination is significantly affectedby the thickness and the material properties of the bottom plies.

1. Introduction
Drilling of composite laminates and the associated internal damages
are critical issues in aircraft manufacturing. Although drilling is an
essential process during the assembly of several parts, the thrust force
from the drill tip and the peeling up from the drill edge often trigger
various kinds of internal damage (e.g., fiber pull-out, matrix crack, andpeel-up/push-out delamination) around the hole. Most of these damages
occurred within the diameter of the drill and are removed by the drilling
process. However, inappropriate drilling conditions can cause significant damage, and the damage left around the hole can negatively affect
the quality of the hole [1–3].
Over the past few decades, several experimental studies have been
conducted to investigate the internal damage during the drilling process[4–14]. These studies have revealed that push-out delamination is themost critical type of internal damage [5–8]. Liu et al. [4] reported that
this delamination is induced by a combination of the thrust force and the
reduction of the bending stiffness of the un-processed part due to adecrease in thickness by drilling. Therefore, the initiation and propagation of this delamination during the drilling process depends not only
on drilling conditions (feed rate [6–13], spindle speed [6–12], drill
shape [10–13], drill wear [14], etc.), but also on the material property[9,15], stacking sequence [9], ply thickness, and other such parameters.
These findings have been obtained empirically based on many parametric studies with different drilling conditions and are not based on thetheoretical background. Although many experimental studies haveinvestigated delamination, all of these studies have only evaluated thedelamination in completely drilled specimens; the detailed initiationand propagation mechanism of push-out delamination during the drilling process has not been completely clarified yet. To the best of theauthors’ knowledge, no previous work has investigated the damage
evolution during the drilling process. The authors believe that it is
impossible to determine the drilling condition theoretically (notempirically) without a comprehensive understanding of the mechanismof push-out delamination. In fact, practical drilling conditions in aircraft
manufacturing are primarily determined based on empirically derived
criteria or by trial and error.
Numerical studies [2,9,16–18] that use the finite element method(FEM) have recently increased as a result of improvements in computational power and the development of simulation techniques. Models
used in previous numerical studies can be separated into two categories;(i) complete simulation models [16–18] and (ii) simplified simulation
models [2,9,18]. Complete simulation models consider almost all phenomena in the drilling process. In these models, a drill model with an* Corresponding author.
E-mail address: higuchi@aastr.t.u-tokyo.ac.jp (R. Higuchi).
https://doi.org/10.1016/j.compscitech.2020.108238
Received 6 February 2020; Received in revised form 9 May 2020; Accepted 13 May 2020
Available online 14 June 2020
0266-3538/© 2020 Elsevier Ltd. All rights reserved.

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Fig. 1. Schematic diagram for the drilling experiment system.

accurate shape is introduced, and both feed rate and spindle rate areimposed on it. In contrast, simplified simulation models neglect the effect of torque and simplify the drilling process into a quasi-static
bending problem. This is because it has been reported that the thrustforce has a greater effect on the push-out delamination than torque [4].
This simplification enables us to avoid performing a dynamic simulation
with high computational cost due to the numerous eroding elements[18]. However, it is difficult to predict the delamination area quantitatively under various drilling conditions, even when a complete simulation model is used. To improve the prediction accuracy, we believe thatthe damage modelling strategy should be verified by detailed comparisons with experimental results in terms of both the final damage left
around the completely drilled holes and the intermediate damage during
the drilling process.
Therefore, the aims of this study are (i) to clarify the damage initiation and propagation mechanisms using experiments, and (ii) to
develop and verify an experiment-based mesoscale damage model for
drilling simulation. First, penetration and interruption drilling tests andtwo kinds of damage observations on the specimens are conducted toinvestigate the damage mechanisms during drilling. The experimental
setup, procedures, and results are summarized in Section 2. Second, wedevelop a simplified simulation scheme with mesoscale damage models
that is based on experimental observations. The details of the damage
modelling strategy, verification against the experimental results, andsome sensitivity studies are described in Section 3. Finally, we present
the conclusions of the present work in Section 4.

2.1. Testing procedure
The experimental setup is depicted in Fig. 1 (a). The experimental
setup and conditions were determined on the basis of previous studies[13,14]. All drilling experiments were conducted using an NC millingmachine (PSF550–CNC, Prospec Industry Co., Ltd.) with a maximum
spindle speed of 3200 rpm. The thrust force and torque were measured
using a 6-axis force sensor (PFS100YS302U6, Leptrino Co. Ltd.) that wasplaced behind the specimen. The specimens were prepared using
T800S/3900-2B (Toray) unidirectional tape. The basic stacking
sequence was as follows: 16 plies of quasi-isotropic laminates½0= " 45=90=45#2S were stacked, and a woven GFRP ply ½0; 90# wasplaced at the bottom. The addition of this ply at the bottom is standard
procedure in the aircraft industry to prevent push-out delamination. The
specimen dimensions were 200 mm (length) $ 30.0 mm (width) $ 3.06
mm (thickness) (CFRP: 3.00 mm and GFRP: 0.06 mm). In all, nine holes
were drilled per specimen at intervals of 20.0 mm along the fiber direction of 0∘ ply. Specimen geometries are shown in Fig. 1 (b). The
specimen was clamped with two knock pins on a backup plate with ahole. The diameter of the hole was large enough (20.0 mm) to mitigate
the effect of the backup plate [18]. In fact, drilling in the aircraft industry is conducted without any backup plate. A crystal diamond coatedtwist drill with a diameter of 6.35 mm (DCDCF03650, Nachi-Fujikoshi
Corp.) was used in this study. The shape and dimensions of the drill
are illustrated in Fig. 1 (c).
Two kinds of observations were made on all drilled holes for damage
characterization. First, three-dimensional damage distribution wasobserved using X-ray computed tomography (ScanXmate-D225RSS270,
Comscantecno Co., Ltd.). The X-ray voltage was 70 kV, and the current
was 90 μA. Moreover, cross-sectional observation was performed using
an optical microscopy (VHX-900, Keyence Corp.). The workpiece wascut parallel to 0∘ plies using a fine cutter and embedded in epoxy. Then,the surface of the cross-section was polished by buffing with diamondabrasive grains (φ ¼ 1, 3 μm).

2. Experiments
In this study, two kinds of experiments were conducted; penetrationand interruption tests. In the former, complete holes were drilled at
different feed rates to examine their effects on the push-out damage
behavior. In the latter, the drilling process was interrupted at several
characteristic points on the thrust force-displacement response. For adetailed investigation of internal damage progression, all manufactured
holes were investigated using X-ray Computed Tomography (CT) andcross-section observations by optical microscopy. The details of the test
procedure are summarized below.

2.2. Penetration test and results
Penetration tests were carried out at a constant spindle speed (3200rpm) and different feed rates (120, 375, 500, and 750 mm/min) toinvestigate the effect of feed rate on the thrust history and damage left
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radius, and 90∘ from 1.3 mm to 3.175 mm radius. This means that themore central region of the drill can impart a higher reaction force in thefeeding direction per unit radius. Therefore, the gradient of thrust history increases when the central region contacts the specimen or penetrates the specimen. Consequently, it can be concluded that the trend ofthrust history is independent of the feed rate, but dependent on thecontact condition between the drill and the specimen. This contact
condition can be characterized by the specimen thickness, the drill geometry, and the drill position.
Fig. 3 summarizes the back views of the specimen and the X-ray CT
images of the 16-plies of CFRP at different feed rates. Push-put delamination occurs at all feed rates (even at rates below the recommended
value of 375 mm/min). The delamination mainly propagates in the fiber
direction of the outermost CFRP layer, but its shape is asymmetricalowing to the effect of drill rotation. Basically, delamination increases
with feed rate. In particular, at a feed rate of 750 mm/min, the outermost plies exhibited considerable push-out and showed significant
delamination. As a result, the quality of the drilled hole shape was very
poor. It is worth noting that very few little damage occurred in the plies,
except for the drill entry and exit, even at the highest feed rate, as shownin Fig. 4. Fig. 4 indicates that the push-out delamination at the drill exit
is much more significant than the peel-up delamination at the drill entry,
which agrees with results obtained in previous studies [5–8]. Thus, weonly focus on the push-out delamination in this work; the peel-up
delamination is not included within the scope of this paper.
The cross-section image along with the 0∘ direction at the hole periphery at a feed rate of 750 mm/min is shown in Fig. 5. It is worth
noting that the push-out delamination does not progress at the ply
interface, but inside the CFRP 16-ply. Strictly speaking, this type ofcomposite fracture is called ”splaying”, not delamination. However, theauthors believe that the mechanism of splaying in this study is the sameas that of the push-out delamination reported in previous studies—both
fractures are mainly caused by the thrust force. The difference in fracture location is thought to be caused by the difference between thecritical energy release rate at the ply interface of the material used in thisstudy and the materials used in previous studies. In conventional CFRP
laminates, the resin-rich region around the ply interface is the weakest
location and easily suffers delamination under out-of-plane deformation. However, the material used in this study (T800S/3900-2B)

Fig. 2. Thrust force-displacement curves with feed rates of 120, 375, 500, and
750 mm/min.

around the completed holes. The drilling conditions were determinedbased on the factory-recommended values (spindle speed of 3200 rpm
and feed rate of 375 mm/min) to explore drilling behaviour underrealistic manufacturing conditions. To ensure repeatability, two specimens were tested at each feed rate. Fig. 2 shows the relationships between the thrust force and drill tip displacement. At all feed rates, thethrust force reached its maximum around 3 mm displacement, i.e.,
where the drill tip penetrated the lowermost layer. Thrust history can bedivided into four stages. During Stages 1 and 2, thrust force linearly
increases with the high and low gradients, respectively. Then, it decreases with the high and the low gradients during Stages 3 and 4. Thischange in thrust history corresponds to the change in the point angle ofthe drill. As shown in Fig. 1 (c), the point angle of the drill is not constantfrom the center to the edge. The point angles are 180∘ at the chisel edge
region from the center to 0.1 mm radius, 108∘ from 0.1 mm to 1.3 mm

Fig. 3. Damage behaviors at the hole periphery with feed rates of 120, 375, 500, and 750 mm/min.
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Fig. 4. Through-thickness damage distribution at a feed rate of 750 mm/min.

contains particles made of a thermoplastic resin at the ply interface toenhance the fracture toughness. The critical energy release rate at theply interface is larger than that inside the ply. As a result, the fracture
location is changed from the ply interface to the interior of the ply.
Therefore, it was inferred that the push-out fracture occurs at the
weakest location around the drill exit, and that there is no difference
between the mechanisms of splaying and delamination. To maintainconsistency in terminology with previous studies, splaying is referred toas “push-out delamination” in this work, although it occurs in theinterior of the specimen.

2.3. Interruption tests and results
The interrupted drilling tests were conducted at a spindle speed of3200 rpm and a feed rate of 375 mm/min. In this test, the feed was
stopped at drill tip displacements of 3.9, 5.0, and 6.0 mm which werenamed points A, B, and C, respectively (as illustrated in Fig. 2). The drill
was pulled out from the workpiece immediately after the feed was
stopped. The interrupted displacement was set at characteristic points inthe thrust force history obtained from penetration tests with the same
drilling conditions shown in Fig. 2. This is because the characteristic
point in the thrust force history is considered as the point where the4

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stiffness reduces by internal damages or where the contact condition
between the drill and laminate changes.
Fig. 6 shows the back and front views of the specimen and X-ray CT
images of the CFRP 16-ply at each interrupted displacement. At point A,
immediately after the drill penetrated the final layer, a cross-shaped
crack was formed. This consisted of the transverse crack and fiber
breakage in the CFRP 16-ply and fiber breakage in the GFRP ply. This
crack propagated along the fiber direction of the CFRP 16-ply rather
than the transverse direction. This is because less energy is required toform transverse cracks than to form fiber breakages in the CFRP 16-ply.
Furthermore, the formation of this cross-shaped crack was accompanied
by push-out delamination. At point B, although the drill tip was located
1.1 mm away from the penetration point, only a small portion of the twooutermost plies were removed. Most regions were not drilled but bent by
the thrust-force since the cross-shaped crack made around point A
reduced the bending stiffness of the unprocessed region. At this point,
several transverse cracks formed in the CFRP 16-ply and divided theremaining region into several strips. This lowered the bending stiffnessof the unprocessed region. Until point C, the push-out delamination was
continuously propagated by the out-of-plane bending of the unprocessedregion. At this point, the delamination reached almost the same size as
the final delamination. Therefore, it was assumed that the drillingFig. 5. Cross-section image along with the 0-degree direction at a feed rate of
750 mm/min.

Fig. 6. Damage process observations at each drill tip displacement obtained in the interruption test with a feed rate of 375 mm/min.

Fig. 7. Conceptual figure of the mechanism of damage propagation at drill-exit.
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Composites Science and Technology 198 (2020) 108238& The extent of delamination may be controlled by adjusting the material properties of the additional ply. This is numerically examinedin Section 3.2.
3. Numerical simulation
3.1. FEM model setup
This section develops a simplified simulation model using a mesoscale damage model to predict the push-out delamination during thedrilling of composite laminates. The developed scheme is based on thesimplified simulation model proposed by Durao et al. [2] and Park [9].
In the simplified simulation model, the drill is approximated by a circular cone, and only the feed rate is imposed on it. This means that thedrilling is simplified into the quasi-static out-of-plane bending problem.
The authors believe that this assumption is not too drastic for the prediction of the push-out delamination since the feed rate is the dominantfactor and the spindle speed does not have a significant effect on it [4].
Fig. 8 (a) illustrates the simplified simulation model and boundary
conditions. The simulation was conducted using Abaqus/Explicit [19].
In the simulation, only a 30 mm square region around the hole wasmodelled and all degrees of freedom on the nodes located on the sides ofthe specimen were fixed. The drill and the back-up plate were modelled
as a rigid body. Contact between the drill and the laminate and between
the back-up plate and the laminate was simulated using the generalcontact algorithm in Abaqus/Explicit [19]. In order to reduce thecomputational cost, the mass density of the laminate was scaled up 1000
times and the feed rate was scaled up 10 times while keeping the percentage of the kinetic energy under 2% of the strain energy.
As discussed in Section 2, the reduction in the bending stiffness of theunprocessed outermost plies due to the geometrical change caused by
the drilling appears to be an important phenomenon. Thus, the element
deletion method [9] was introduced to reproduce the actual contact
condition between the drill and laminate. In this method, the elementsin the processed region are removed when they are applied to the
maximum force determined based on the sectional thrust force obtained
from the experiment. The details of the calculation procedure of thrust
distribution are described in Appendix A. In the complete model[16–18], the element deletion was done based on the strength and energy to consider the local energy balance when a new fracture plane is
generated in one element. However, it is impossible to use these criteria
in the simplified model where the actual fracture planes cut by the drill
edge are not directly modelled. Alternatively, the simplified model
considers the global energy balance between the total work done by the
drill feed, stored strain energy in the laminates, and energy dissipated by
the fracture at drill exit.
Additionally, the experiment-based mesoscale damage modelling
approach [20], which combines continuous and discrete damage models
based on experimental observations, was utilized to predict the various
internal damages. Generally, the continuous damage model such as thesmeared crack model (SCM) is computationally efficient and robust but
it can easily underestimate the propagation of large cracks. In thecomposite laminates, the deterioration of the prediction accuracy oftransverse cracks directly affects that of the delamination since thetransverse crack and the delamination strongly interact with each other[21,22]. On the other hand, like the cohesive zone model (CZM), thediscrete damage model is able to capture damage propagation including
the interaction between the transverse cracks and delamination precisely, but is computationally expensive. Therefore, theexperiment-based hybrid modelling strategy enables reasonably
computationally-efficient modelling. As observed in the drilling experiments, relatively large cracks formed only in the two bottom plies,
which consisted of several transverse cracks and fiber tensile failure in
the outermost CFRP ply and fiber tensile failure in the additional ply.
These cracks were accompanied by the push-out delamination. In orderto model their interactions, we employed CZM for these cracks and theFig. 8. (a) FEM model setup and boundary conditions for simplified drilling
simulation, (b) Schematics of the multiple crack (MC) model, and (c) Schematics of the no crack (NC) model.

process after point C did not cause propagation of the delamination but
removed the remaining region in the two outermost plies. This isbecause the thrust force applied only from outside the drill was not highenough to cause the delamination to progress.
Consequently, the bending behavior and the internal damages in
additional GFRP ply significantly affect on the push-out delamination
behavior. Therefore, the improvement of material properties (stiffness,
strength, and fracture toughness) of additional ply has the great potential to suppress the extent of delamination.
The findings about the mechanisms of the push-out delamination
obtained from the penetration and interruption tests are summarized
below and illustrated in Fig. 7:& The cross-shaped crack, consisting of the transverse crack in the
CFRP 16-ply and the fiber breakages in the CFRP 16-ply and GFRP
ply, formed immediately after the drill tip penetration (i.e. after the
maximum thrust force).
& As drilling progressed, several transverse cracks formed in the CFRP
16-ply and divided the two unprocessed outermost plies into several
strips.
& These transverse cracks and fiber breakages lowered the bending
stiffness of the unprocessed region in the outermost plies. The out-ofplane bending of the unprocessed outermost plies caused the propagation of the push-out delamination.

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Table 1
Material properties of T800S-3900-2B.
T800S/3900-2B[25]

GFRP [29]2500 kg/m3
26 GPa26 GPa8 GPa
3.8 GPa
2.8 GPa

Densityρ1800 kg/m3

Longitudinal Young’s modulus
Transverse Young’s modulus
Out-of-plane Young’s modulus
In-plane shear modulus
Out-of-plane shear modulus

E11
E22
E33
G12
G23

Out-of-plane shear modulus
In-plane Poisson’s ratio

G31
v12

153 GPa
8 GPa
8 GPa
4.03 GPa2.75 GPa [26,27]
4.03 GPa
0.34

v23

0.45

0.25 [30]

v13

0.34

0.25 [30]3100 MPa
1242 MPa [27]193.3N/mm

850 MPa
720 MPa
50N/mm

GC1C

25.9N/mm

50N/mmt10

Out-of-plane Poisson’s ratio
Out-of-plane Poisson’s ratio

2.8 GPa
0.10 [30]Longitudinal tensile strengthLongitudinal compressive strength
Fracture toughness for fiber tensile
failureFracture toughness rate for fiber
compressive failure
Mode I maximum traction

XT
XC
GT1C

66.9 MPa

214 MPa

Mode II and III maximum traction

t20 ; t30

100 MPa

38 MPaGCI

0.2N/mm [28]

GCII ;
GCIII

1.0N/mm [28]3.47N/
mm
1.0N/
mma

Mode I fracture toughness
Mode II and III fracture toughnessa

GCII and GCIII in GFRP are estimated values.

push-out delamination. In contrast, no significant damage propagation
was observed in the inner plies (Fig. 4). These damages were thus
modelled by SCM. The peel up delamination is beyond the scope of theprediction. The schematic of the mesoscale damage model is shown in
Fig. 8 (b). The laminate was divided into reduced integration solid elements (C3D8R). All plies other than the CFRP 16-ply were divided into
one element in the thickness direction. The CFRP-16 ply where thepush-out delamination was observed was divided into two elements inthe thickness direction. To model the large multiple cracks in the twobottom plies and the push-out delamination, the surface-based CZM wasintroduced. While the CZM for splittings in the CFRP-16 ply wereinserted parallel to the fiber direction, those for fiber failure in the
CFRP-16 ply and woven GFRP ply were inserted perpendicular to thefiber direction. Additionally, the crack distance was set to 0.8 mm basedon experimental observation. This value is consistent with the saturated
crack density found in the tensile test of similar material reported in [23]and analytically predicted by Onodera et al. [24]. Therefore, when thismodel is applied to other materials and stackings, the crack distance canbe determined based on experiment or analytical models. To examine
the importance of modelling the transverse cracks for the prediction of
push-out delamination, an additional simulation model with no cracks
was prepared (Fig. 8 (c)). Hereafter, the simulation models with andwithout multiple cracks are named as MC and NC models, respectively.
The details of SCM and CZM can be found in Refs. [19,20]. All material
properties used in the simulation are summarized in Table 1 [25–30]. As
discussed in Section 2.1, the push-out delamination happened inside theply and not at the interface toughened by thermoplastic particles. In fact,
the critical energy release rates at the toughened interface (GCI ¼ 0:54
N/mm and GCII ¼ 1:54 N/mm [25]) are relatively higher than those ofgeneral CFRP interface (e.g., GCI ¼ 0:2 N/mm and GCII ¼ 1:00 N/mm[28]). It was assumed that the critical energy release rates inside the ply
in which there are no toughening particles are similar to those of thegeneral CFRP interface since it is difficult to obtain them.

Fig. 9. Comparisons of the thrust history and damage propagation behavior
with a feed rate of 375 mm/min between experiment and simulations using theMC and NC models.

model were carried out. Then, sensitivity studies were conducted to
explore ways to mitigate the push-out delamination.
Firstly, the simulations using the MC and NC models were comparedwith the experiment under the condition of a spindle speed of 3200 rpm
and a feed rate of 375 mm/min. Fig. 9 compares the thrust history andthe damage growth between the experiment and simulations. As shownin Fig. 9, the length of the push-out delamination predicted by the NC
model was shorter than that of the MC model. In the MC model, thedelamination length at the center was the longest because the centertransverse crack propagated easier than other cracks because of Mode I
deformation and triggered the delamination. On the other hand, in the
NC model, the delamination front became flat since no crack triggered
the delamination. As a result, the delamination shape of the MC model iscloser to the experimental result than the NC model. This result enhances the importance of modelling the transverse cracks in the prediction of push-out delamination. Additionally, Fig. 9 confirms that thepredicted thrust history in the MC model is in good agreement with theexperiment. One exception is the overestimation of thrust force at thedisplacement of 4.0–5.0 mm. As shown in the back view at a displacement of 5.0 mm, several regions of the outermost plies were removed in
the experiment. However, this removal could not be reproduced in thesimulation since these regions were cut by drill rotation, which was not
actually reproduced in the simulation. Therefore, the simplified simulation overestimated the area of the unprocessed region and the resulting reaction force from it. This mismatch is the limitation of the

3.2. Results and discussion
In this section, first, several verifications of the proposed simplified
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Fig. 10. Comparisons of predicted results by the MC model with experiments for each feed rate; (a) Maximum thrust force, (b) Delamination area.

configurations were prepared as follows.
& Stacking A CFRP½0= " 45=90=45#2S þ GFRP½0; 90# ðoriginalÞ
& Stacking B CFRP½02 = " 452 =902 =452 #2S þ GFRP½0; 90#& Stacking C CFRP½0= " 45=90=45#2S þ CFRP½0; 90#

By comparing Stacking A with B, it is possible to understand theeffect of ply thickness on the drilling damage. This is one of the size
effect well-known in the polymers and polymer composites [28,31–35].
On the other hand, the comparison between Stackings A and C will
provide us the knowledge needed to control the drilling damage by
adjusting the material properties of the additional ply on the bottom
side. In the woven CFRP ply of Stacking C, the longitudinal elastic andfracture properties of T800S/3900-2B were assumed in both in-plane
directions. The predicted delamination areas and maximum
out-of-plane displacements are compared in Fig. 11. Here, the maximum
out-of-plane displacement is obtained at the center nodes of additionalply before their elements were removed by the drilling. It can be seenfrom Fig. 11 that there is a clear size effect of ply thickness in the drilling
damage. The thicker ply laminate (Stacking B) resulted in a larger
delamination area than the thinner ply laminate (Stacking A). This isbecause the transverse cracks that trigger the push-out delamination
have an in-situ effect [36,37]. As discussed in the references [36,37], thetransverse cracks propagate easier in the thick ply than in the thin ply.
Consequently, applying a thinner CFRP ply at the bottom side is useful in
reducing the drilling damage. Furthermore, by comparing Stackings A
and C, it was confirmed that the stiffer and stronger additional ply
successfully reduced the out-of-plane deflection and resultant delamination area. These results show that delamination can be controlled byadjusting the material properties of the additional ply on the bottom
side.

Fig. 11. Comparisons of the delamination area and maximum out-of-plane
displacement with a feed rate of 375 mm/min between three different specimen configurations.

simplified simulation model. However, the simulation was able toreproduce the sequence of the internal damages (cross-shaped crack,
several transverse cracks, and the push-out delamination) and theresultant residual delamination area. The authors thus believe that thismismatch is not a severe disadvantage from the perspective of engineering use of the simulation tool. In fact, the simplified simulation
required only 3–5 h when using the 32 CPU workstation (Intel Xeon E52680 v3, 2.50 GHz). This CPU time is a few percent of that for thecomplete simulation model [16,18].
For more detailed verification of the MC model, simulations wereconducted under the conditions of a spindle speed of 3200 rpm and feed
rates of 120, 375, 500, and 750 mm/min. The comparisons of the predicted and experimental maximum thrust force and push-out delamination areas are summarized in Fig. 10(a) and (b). Here, the predictedmaximum thrust force is obtained from the moving average of 20 pointsof raw data from the simulation because numerical oscillation is inevitable in the drilling simulation where discontinuous contact happens. In
all feed rates, the proposed model can accurately predict the maximum
thrust force and the delamination area. From these results, it was
concluded that the proposed simplified simulation with the experimentbased mesoscale damage model is useful for the computationally efficient prediction of the thrust history, damage scenario, maximum thrustforce, and residual delamination area.
Finally, the drilling simulations with different specimen configurations were performed to explore the way to reduce push-out delamination. In addition to the original specimen configuration, two specimen

4. Conclusions
This study conducted an experimental and numerical investigation
into the push-out delamination during the drilling of composite laminates. In the experiment, the penetration and interruption tests werecarried out to determine the damage progression mechanism during thedrilling process. From these experiments, the following damage scenario
was revealed; (i) after the drill tip penetration, several damages in theoutermost plies (transverse cracks in CFRP and fiber breakage in CFRP
and GFRP) reduce the bending stiffness, (ii) bending (not drilling) results in propagation of delamination. From these findings, theexperiment-based mesoscale damage model was developed. Transverse
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Fig. A1. Definition of sectional thrust force.

to be the dominant events against the push-out delamination, weremodelled by CZM. The fiber breakages in the other regions, which did
not propagate largely, were modelled by SCM. These damage models
were implemented in the simplified drilling simulation model, in which
the drilling was approximated as an out-of-plane bending problem by
the conical. The developed simulation scheme was able to predict thethrust history and delamination propagation in all feed rates. It was thus
confirmed that the developed tool is a useful engineering tool todetermine the drilling conditions for low residual damage. Furthermore,
based on the sensitivity studies conducted using the proposed model, it
was concluded that the extent of delamination can be controlled byadjusting the thickness and material properties of the bottom plies.

T. Okabe: Supervision, Conceptualization.
Declaration of competing interest
The authors declare that they have no known competing financial
interests or personal relationships that could have appeared to influence
the work reported in this paper.
Acknowledgements
This study was partially supported by the New Energy and Industrial
Technology Development Organization (NEDO) (Project No. P08024,
No. P15006). This study was also supported by the Council for Science,
Technology and Innovation (CSTI), the Cross-ministerial Strategic
Innovation Promotion Program (SIP), and the ’Materials Integration’ forrevolutionary design system of structural materials (Funding agency:
Japan Science and Technology Agency (JST)).

CRediT authorship contribution statement
R. Higuchi: Conceptualization, Methodology, Software, Writing original draft, Writing - review & editing. S. Warabi: Software, Validation, Investigation. W. Ishibashi: Software, Validation, Investigation.

Appendix A. Calculation procedure of thrust distribution
In this study, the drilling is simplified to an out-of-plane bending problem caused by the feeding of a conical, whose shape is the same as therotating shape of the drill. In the simulation, the drilling process is modelled using the element deletion method. In this method, the elements are
removed when they are subjected to maximum stress σ c ðrÞ. By considering the increment of thrust force dF during the infinitesimal feed of drill, as
shown in Fig. A1 (a), the maximum stress can be determined as follows.
dF ¼ σ c ðrÞdSsinθ ¼ σ c ðrÞ2πrdr ⇔ σc ðrÞ ¼

1 dF
⋅ ;
2πr dr(A.1)

where, dF=dr can be calculated from the relationship between the thrust force and radial position as shown in Fig. A1 (b). Here, the radial position is
calculated using the drill tip position and drill geometry. From Fig. A1 (b), dF=dr is discontinuous at Rkink where the point angle changes from 108∘ to
90∘ . Additionally, we assumed that the specimen at the chisel edge region is subjected to pure transverse compression. Consequently, the distributionof maximum stress σ c ðrÞ results in8
ðr * Rchisel Þ
Yc
>
>
>
!
>
>
!
1
dF
>
< karea ⋅ ⋅ !
ðRchisel * r * Rkink Þ
2π r dr !r*Rkink;(A.2)
σ c ðrÞ ¼
>
>
!
>
>
!
1 dF
>
>
: karea ⋅ ⋅ !!
ðRkink * rÞ
2π r dr
Rkink *rwhere, Yc is the transverse compressive strength, Rchisel is the chisel edge radius, and karea is the correction factor of the contact area. In the drilling
simulation, remeshing is necessary to ensure the processed surface is along the drill surface, however, this results in high computational cost.
Therefore, in this study, the laminate is divided by the general structural mesh. In this case, the drilled surface of the specimen becomes step-wise andthe contact area is underestimated. In order to correct the contact area in the simulation with the structural mesh to the ideal contact area between theconical and laminate, a correction factor karea was introduced. karea was determined based on the simulation with a certain feed rate; the same factor isused for the other feed rate as it only depends on the FEM mesh and the drill shape. The determined values are as follows.

9R. Higuchi et al.

karea ¼"

Composites Science and Technology 198 (2020) 108238

9:0 ðRchisel * r * Rkink Þ:
3:0 ðRkink * rÞ(A.3)

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