Composite Structures 349-350 (2024) 118509
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Composite Structures
journal homepage: www.elsevier.com/locate/compstruct
Identification of mode I fracture toughness in GFRP/Al and GFRP/Cu joints
for structural batteries
Maryam Niazi a,b , Federico Danzi b , Ricardo Carbas b,c , Pedro P. Camanho b,c ,∗
a
b
c
Engineering Physics Department, Engineering Faculty, University of Porto, Rua Dr. Roberto Frias, Porto, 465-4200, Portugal
LAETA-INEGI, Institute of Science and Innovation in Mechanical and Industrial Engineering, Campus da FEUP, Rua Dr. Roberto Frias, Porto, 400-4200, Portugal
Mechanical Engineering Department, Engineering Faculty, University of Porto, Rua Dr. Roberto Frias, Porto, 465-4200, Portugal
ARTICLE
INFO
ABSTRACT
Keywords:
Hybrid-interface
Structural batteries
Fracture toughness
VCCT
Roughness
Surface free energy
Fiber metal laminates
Fiber metal laminates (FMLs) have been proposed as components of structural batteries, yet their elastic mismatch can lead to interface cracks, compromising structural integrity and both mechanical and electrochemical
efficiency. To this end, the bonding between the metal and the composite layer is of utmost importance. In
this study, the effect of various metal surface treatments on the mode I interlaminar fracture toughness of
two FML configurations suitable for structural batteries — Glass Fiber Reinforced Polymer (GFRP)/aluminum
laminate and GFRP/copper laminate — was examined. The surface treatments included sulfo-ferric etching,
NaOH/HNO3 etching, and Sol–Gel anodizing for aluminum 2024-T3, as well as FeCl3 /HCl/Glycerol treatment
and Sol–Gel anodizing for copper alloy. The results were compared with untreated conditions and with the
baseline GFRP/GFRP configuration. GFRP/metal coupons were designed to achieve pure mode I interlaminar
fracture toughness in double cantilever beam (DCB) tests, and the designs were verified using the Virtual Crack
Closure Technique (VCCT). The surface characterization of the metals was performed using contact angle tests
to estimate the surface free energy, while Coherence Scanning Interferometry (CSI) was used to measure the
surface roughness and topography.
1. Introduction
and 1d). In this concept, the core components and metallic current
collectors need to be enclosed with an insulating protective shell for
safety reasons. In both configurations, GFRP stands out as a promising
candidate, either as the structural shell for embedding the battery or
as an insulating cover for hosting the multifunctional system, due to
its excellent mechanical and insulating properties [11]. However, the
bonding between GFRP and the current collectors is a concern due to
the material discontinuity and the stiffness mismatch that make the
interface layer prone to crack nucleation and propagation. This issue
can compromise both the structural integrity and the electrochemical efficiency of the system [12–14]. In the realm of electrochemical
performance, delamination between the composite shell and current
collectors causes the system to lose its compactness and strength,
compromising the support for the core components of the battery
against internal loads caused by charge/discharge cycles [15–17]. This
may affect the bonding between internal layers, such as electrodes
and electrolytes, thereby degrading the cell electrochemical efficiency.
Therefore, identifying effective methods to enhance the interlaminar
Extensive efforts have been dedicated to enhance the specific capacity of energy storing devices concerning both mass and volume, leading
to the development of multifunctional structures [1]. In this regard, the
design of structural batteries with the capability of carrying mechanical
loads and storing electric energy at the same time, promises to be a
breakthrough in energy storage devices [2–6]. In the last decade, two
different approaches have emerged for designing weightless structural
batteries. In the first approach, the core components of the battery
(cathode, electrolyte, and anode), along with the current collectors
(mostly aluminum and copper) are integrated inside a structural shell
made of fiber-reinforced polymers. This design concept is typically
addressed as ‘‘embedded batteries’’ [7,8] and it is shown in Figs. 1a
and 1b. In this configuration, the battery core components handle
the electrochemical functions, while the structural shell carries the
mechanical load. In the second approach, called ‘‘multi-functional materials’’, the fibers and matrix of the composite not only serve as the
core components of the battery and perform electrochemical functions
but also carry mechanical loads simultaneously [9,10] (see Figs. 1c
∗ Corresponding author at: LAETA-INEGI, Institute of Science and Innovation in Mechanical and Industrial Engineering, Campus da FEUP, Rua Dr. Roberto
Frias, Porto, 400-4200, Portugal.
E-mail address: pcamanho@fe.up.pt (P.P. Camanho).
https://doi.org/10.1016/j.compstruct.2024.118509
Received 12 April 2024; Received in revised form 5 July 2024; Accepted 18 August 2024
Available online 22 August 2024
0263-8223/© 2024 The Author(s). Published by Elsevier Ltd. This is an open access article under the CC BY license (http://creativecommons.org/licenses/by/4.0/).
Composite Structures 349-350 (2024) 118509
M. Niazi et al.
Fig. 1. Different types of structural batteries, a. Flat embedded structural batteries, b. Coaxial embedded structural batteries [4], c. Multi-functional materials with matrix
cathode [10] (reprinted with permission from Elsevier-2024), and d. Multi-functional materials with coated carbon fibers cathode [18].
fracture toughness between current collectors and the composite shell
is crucial for achieving an efficient energy storage system.
Nowadays, aluminum and copper are essential current collectors in
batteries, valued in commercial markets due to their electrochemical
stability, low cost, abundance, and excellent electrical properties [19,
20]. Aluminum serves as the current collector for the cathode, while
copper serves this role for the anode in many commercial battery packs,
and these assignments are not interchangeable [19]. When assessing
the fracture toughness of bonding between these metals and composite
laminate for structural battery applications, simultaneous study of both
current collectors is required for evaluating the performance, integrity,
and identifying potential weaknesses. Due to the poor adhesive properties of the bare metals, surface treatment prior to bonding is required
to enhance interfacial adhesion through mechanical interlocking and by
increasing the number of active chemical sites at the metal–composite
interface [21]. Regarding the aluminum alloy, to date, several methods have been used to improve the surface morphology: mechanical
abrasion and grit-blasting, alkaline etching, acidic etching, and anodizing [22]. Mechanical abrasion and grit-blasting increase the surface
roughness, improving the mechanical interlocking between the metal
layer and the composite laminate [23,24]. The alkaline and acidic
etching enhance the adhesion by altering the surface chemical composition and by creating a strong adherent porous oxide layer [25–27]. It
has been reported that chemical treatments with NaOH/HNO3 cause
a chemical attack by dissolving the native oxide layer and particles
of the aluminum, providing more appropriate layer for bonding with
composite [28,29]. In addition, some electrochemical methods such
as Forest Product Laboratory etching (FPL-etching) [30], sulfo-ferric
etching (P2-etching) [23,31], chromic acid anodizing (CAA) [32,33],
and phosphoric acid anodizing (PAA) [34,35] have been referred as
effective treatments that enhance interfacial bonding of composite and
aluminum through making porous nanostructure on the metal surface.
For several years after 1950, FPL-etching was very popular in the
aerospace industry, however, the use of toxic chromium compounds
(e.g. hexavalent chromium) in the treatment process restricted its application in many countries. Therefore, sulfo-ferric etching (P2-etching)
was developed in the 1970s as a less toxic alternative with a similar
performance [36]. During the last ten years, a user-friendly Sol–Gel
based technique, with good chemical stability, has been widely used
in aerospace field as a potential technique to mitigate the negative
effects that are associated with service surface treatments, and to
protect metals from common problems such as corrosion, fouling, and
wearing [37–39]. Regarding the copper, E 407–99 standard suggests
several chemical treatments. For example, Ferric chloride (FeCl3 ) is
indicated for small-scale etching process [40]. But less information are
currently available in the literature.
The interlaminar fracture of bonded joints is usually a combination
of three fracture modes: opening mode (Mode I), shearing mode (Mode
II) and tearing mode (Mode III). The Double Cantilever Beam (DCB)
test has been extensively used to experimentally measure the mode I
critical strain energy release rate (𝐺𝐼𝐶 ) for symmetric adherents (same
material and same thickness) [41–43]. Due to the high application
of FMLs in many areas such as aerospace [44–47] automotive [48],
and multi-functional structural energy devices [5], the interlaminar
fracture properties between composite and metal adherents (usually
with different thickness) has attracted recent studies [49–53].
This study focuses on assessing the impact of various surface treatments on the mode I interlaminar fracture toughness of two FML
configurations deemed suitable for structural batteries: GFRP/Al and
GFRP/Cu. The treatments applied to aluminum 2024-T3 included sulfoferric acidic etching (P2-etching), NaOH/HNO3 chemical etching, and
Sol–Gel anodizing. For copper, the treatments comprised FeCl3 /HCl/
Glycerol chemical etching and Sol–Gel anodizing. The results were
subsequently compared with untreated conditions and the symmetric GFRP/GFRP laminates. The hybrid DCB configurations were initially designed using theoretical approaches found in the literature
for bi-material systems. These designs were numerically verified by
studying the mode mixity at the crack tip using the Virtual Crack
Closure Technique (VCCT). For each treatment, the surface free energy
was measured using contact angle tests. Moreover, surface roughness and topography were obtained using the Coherence Scanning
Interferometry (CSI) method.
2. Materials and method
2.1. Material
The high-strength aluminum alloy Al2024-T3 and copper alloy
Cu110-H02, both with a thickness of 2 mm, were sourced from
McMaster-Carr. E-glass PPG-600T-15 𝜇-25 gsm-300 mm-TP736LT was
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Composite Structures 349-350 (2024) 118509
M. Niazi et al.
Table 1
Mechanical properties of E-Glass Lamina.
𝐸1 (GPa)
𝐸2 (GPa)
𝜈12
𝐺12 (GPa)
40
6.43
0.298
2.735
The reaction continues with ferric sulfate of P2-paste, where ferric
sulfate make pitting of the alloy surface, leading to non-uniform attack
at specific areas [36].
Table 2
Mechanical properties of metals.
Material
𝐸 (GPa)
𝜈
𝜎𝑦 (MPa)
𝜎𝑢 (MPa)
Aluminum 2024-T3
Copper 110-H02
73
117
0.33
0.3
316
261
457
289
Aluminum
Copper
(3)
2Fe3+ + Cu → 2Fe2+ + Cu2+
(4)
Afterwards, the substrates were cleaned with deionized water and
dried at room temperature.
2.2.3. Chemical etching of aluminum with NaOH-HNO3
Following the procedure specified in [28,29], after pretreatment,
the aluminum substrates were immersed in NaOH (5%) solution for
5 min, forming sodium aluminate (NaAlO2 ) and hydrogen gas according to etching reaction:
Table 3
The selected surface treatments for metals.
Metals
3Fe3+ + Al → Al2+ + 3Fe2+
Surface treatments
Without treatment
P2-etching
NaOH/HNO3
Sol–Gel
2Al + 2NaOH + 2H2 O → 2NaAlO2 + 3H2
Afterward, the substrates were cleaned by rinsing with deionized water.
During this step, NaAlO2 hydrolyzed in the presence of water, resulting
in Al(OH)3 as shown in Eq. (6).
Without treatment
FeCl3 /HCl/Glycerol
Sol–Gel
NaAlO2 + 2H2 O → NaOH + Al(OH)3
2Al(OH)3 → Al2 O3 + 3H2 O
After cutting the aluminum and copper sheets to the specified
dimensions of the DCB specimens (200 mm × 25 mm) using a 5axis DMG-DMU-60eVo CNC machine, surface treatments as outlined in
Table 2 were applied. The procedures for each treatment are detailed
in the following subsections (see Table 3).
2.2.4. Sol–Gel anodizing of aluminum and copper
Sol–Gel anodizing was performed using the 3M-AC-130-2 aerospace
sealant, which consists of Part A — a water solution containing propyl
alcohol, zirconium n-propoxide, acetic acid, and water — and Part
B, GTMS—a blend of 3-glycidoxy propyl-trimethoxysilane and methyl
alcohol. The Sol–Gel paste was prepared by mixing Part A and Part
B in compliance with the supplier’s recommended ratio of 49.2:1.06,
until a transparent and homogeneous solution was achieved. Following
pretreatment of aluminum and copper substrates, the Sol–Gel solution
was applied uniformly to cover their entire surface area. Subsequently,
the substrates were air-dried at room temperature for approximately
one hour. This treatment involves the growth of metal-oxo polymers
through hydrolysis (Eq. (8)) and condensation (Eq. (9)), resulting in
the formation of a thin, inorganic zirconium oxide film incorporated
with epoxy-silane on the metal surface [38,39,55].
2.2.1. Pretreatment
Before starting any treatment, all metal substrates were cleaned
using a soft sponge, soap, and water to remove dust and contaminants
originating from manufacturing processes. This was followed by rinsing
with deionized water and drying at 40 ◦ C in an oven for 1 h. Each
substrate surface was then cleaned with acetone and dried at room
temperature. Subsequently, the substrates were manually sanded using
800-grit sandpaper, followed by cleaning with acetone and drying at
room temperature. It is worth noting that all configurations, including
those labeled ‘‘without treatment’’, incorporate a preliminary step for
surface pretreatment to ensure proper cleaning of the metal substrates.
2.2.2. P2-etching of aluminum
Following the procedure specified in [36,54], P2-paste was made by
mixing 5 g of concentrated sulfuric acid, 2 g of 75% ferric sulfate, and
10 g deionized water. After pretreatment of the aluminum substrates,
the P2 solution was applied on their surfaces and then left at room
temperature for 25–30 min. The sulfuric acid in the P2 paste resulted
in the following chemical reactions, on the aluminum and copper
components of the aluminum substrates:
2+
Cu + 4H+ + SO−
+ SO2 + H2 O
4 → Cu
(2)
(7)
Finally, the aluminum substrates were immersed in a 30% weight
nitric acid (HNO3 ) solution for 45 s to clean the metallic unstable
particles on the surface, resulting in a new rough layer of Al2 O3 that is
appropriate to be bonded to the composite [28,29]. Subsequently, the
substrates were rinsed with deionized water at room temperature. The
substrates were then dried in an oven at 40 ◦ C for one hour.
2.2. Surface treatments
(1)
(6)
Then, during the drying step (at 40 ◦ C for one hour), aluminum quickly
oxidized to aluminum oxide (Al2 O3 ), as shown in Eq. (7), forming a thin
(approximately 10 nm), unstable layer that can result in poor bonding
if cured with the composite at this stage.
used as the thin-ply lamina. The mechanical properties of the Eglass fiber lamina were calculated using the rule of mixtures and are
shown in Table 1, where 𝐸1 is Young’s modulus in the longitudinal
(fiber) direction, 𝐸2 is Young’s modulus in the transverse and thickness
direction, and 𝜈12 and 𝐺12 are the Poisson’s ratio and shear modulus,
respectively. The mechanical properties of the aluminum alloy and
copper alloy are shown in Table 2, where 𝐸 and 𝜈 represent the
Young’s modulus and Poisson’s ratio of the metals. The parameters 𝜎𝑦
and 𝜎𝑢 represent the yield stress and tensile strength of the metals,
respectively.
2Al + 6H → 2Al3+ + 3H2
(5)
Si(OEt)4 + 2H2 O → Si(OH)4 + 4EtOH (hydrolysis)
(8)
Si(OH)4 → SiO4 + 2H2 O (condensation)
(9)
2.2.5. Chemical etching of copper with FeCl3 /HCl/Glycerol
According to ASTM E 407–99 standard, a mixture was prepared
by combining 1 g of ferric chloride (FeCl3 ), 2 mL of HCl, 10 mL of
glycerol, and 3 g of deionized water. The components were stirred until
a homogeneous solution was achieved. Following the pretreatment
stage of copper substrates, the solution was applied to their surface
using a swab for 60 s. Subsequently, the treated substrates were rinsed
with warm deionized water and then air-dried at room temperature for
approximately one hour.
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Composite Structures 349-350 (2024) 118509
M. Niazi et al.
Fig. 2. The dimensions of the hybrid DCBs.
Table 4
Configuration of GFRP/Al and GFRP/Cu laminates.
Lay up of GFRP
ℎ𝑐 (mm)
ℎ𝑚 (mm) [theoretical thickness]
𝐸𝑥𝑓 (MPa)
GFRP/Al
GFRP/Cu
[0◦ /45◦ /90◦ /−45◦ ]20𝑆
4
2 [1.99]
18 106
[0◦ /45◦ /90◦ /−45◦ ]25𝑆
5
2 [1.96]
18 022
the specimen, the mesh size was varied from 0.25 mm near the crack
tip to 2 mm in regions farther from the crack tip. For the thickness,
except for the first 8 plies, the mesh size was varied from 0.25 mm
to 0.5 mm. The element size through the width of the specimens was
fixed at 0.25 mm. The total number of elements in the model was
314000. By obtaining the displacement of the nodes adjacent to the
crack tip and using the nodal forces at the crack tip nodes, along with
the equations of VCCT shown in Appendix B of the supplementary
data [60], the contribution of each mode of interlaminar fracture toughness was determined. The energy release rates for mode I, mode II, and
mode III components, denoted as 𝐺𝐼 , 𝐺𝐼𝐼 , and 𝐺𝐼𝐼𝐼 , respectively, were
calculated for GFRP/Al and GFRP/Cu laminates. Afterwards, the energy
release rates were divided by the total energy release rate (the sum of
all components) to calculate the relative contribution of each mode.
Fig. 4 shows the contribution of each mode for both configurations with
a crack length of 50 mm and an opening displacement of 4 mm. Results
for other crack lengths, including 60 mm, 70 mm, 80 mm, 90 mm, and
100 mm, are provided in Appendix B of the supplementary data. For
all crack lengths investigated, the findings indicated that the design of
both hybrid laminates predominantly exhibits mode I behavior (around
97%), with mode II contributing approximately 3% and mode III less
than 0.01%.
2.3. Designs of hybrid DCBs and VCCT verification
As discussed in the literature, for asymmetric DCBs, achieving pure
mode I interlaminar fracture toughness typically requires matching
the flexural stiffness of the two arms. Otherwise, shear stress in the
interface of dissimilar adherents results in a mixed-mode loading scenario [50,56,57]. Ouyang et al. [56] developed an accurate method
to estimate the thickness of each arm in bi-material DCB systems
under pure mode-I loading. Their theoretical approach suggests that the
condition 𝐸ℎ2𝑚 = 𝐸𝑥𝑓 ℎ2𝑐 , where 𝐸 is the Young’s modulus of the metal
arm, ℎ𝑚 and ℎ𝑐 are the thicknesses of the metal arm and composite arm,
respectively, and 𝐸𝑥𝑓 is the longitudinal flexural stiffness of the composite arm, minimizes the effects of shear stresses and results in a pure
mode-I loading scenario. This method was validated by numerical and
experimental studies [50,56]. In this study, by applying this condition
(𝐸ℎ2𝑚 = 𝐸𝑥𝑓 ℎ2𝑐 ) and using the classical laminate theory, the thickness
of the composite arms for both hybrid configurations (GFRP/Al and
GFRP/Cu) was determined based on the specified thickness of the metal
arms. The general equations to obtain 𝐸𝑥𝑓 are presented in the Appendix
A of the supplementary data [50]. The resulting configurations were
numerically tested in Abaqus [58] to ensure there was no plastic
yielding in the metal arms during quasi-static fracture testing. The final
designed configurations are presented in Table 4, and shown in Fig. 2.
The 3D Virtual Crack Closure Technique (VCCT) was employed to
calculate mode-mixity and determine the energy release rate profiles
at the crack tip line [59]. The model was defined as linear elastic.
Both the GFRP and metal arms were modeled by properly partitioning
a single part. To represent the initial crack, a seam was assigned to
the relevant partition of pre-existing delamination. A contour integral
was used to define the crack tip line, and a singularity of 0.25 was
specified to accurately capture the stress intensity factors around the
crack tip. Fig. 3a illustrates the boundary conditions applied in the
numerical models: on the crack side of the DCB, the outer edge of each
adherend along the width were subjected to an opening displacement
in out-of-plane direction. Fig. 3b shows the mesh configuration of the
numerical models. C3D8I (8-node linear brick, incompatible modes)
were employed for the finite element mesh. The first eight plies of
the GFRP were modeled separately using partitions and each ply was
meshed with one element through its thickness. Through the length of
2.4. Manufacturing DCB specimens and testing procedure
The GFRP laminates were manufactured through hand layup and
cut using a roller cutter, aligning it with metal sheets as a reference
and adopting the dimensions of the DCB specimen (200 mm × 25 mm).
The insert was manufactured from an Upilex film with a thickness of
25 μm, simulating an initial artificial crack of approximately 50 mm.
The metal adherends were co-cured with the GFRP laminate using a hot
press at a pressure of 25 bar and temperatures specified by the supplier
(1 h at 70 ◦ C and 3 h at 100 ◦ C). After curing, the thickness of the
hybrid laminate was measured to be 5.7 mm for GFRP/Al and 6.6 mm
for GFRP/Cu. To prepare the specimens for DCB tests, parallelepipedal
steel blocks were used. These blocks measured 25 mm × 20 mm ×
20 mm and had a central hole with a diameter of 6.2 mm. They
were placed on the pre-crack sides of the specimens using an epoxybased adhesive and a mould with appropriate reference pins to ensure
alignment. The DCB tests were performed in compliance with the
ASTM-D5528 standard test method. Quasi-static tests were conducted
using a 1 kN load-cell tensile testing machine under displacement
control at a rate of 1 mm/min. Digital photos were made every two
seconds by using Canon EOS 80D camera. The pre-crack was extended
by 3 to 5 mm from the insert prior to testing, as recommend in the
standard for inserts thicker than 13 μm. The crack length was calculated
as the distance between the crack tip and the center-line of the pins of
the blocks.
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Composite Structures 349-350 (2024) 118509
M. Niazi et al.
Fig. 3. Finite element model of hybrid DCBs: a.Boundary conditions, b. Mesh overview.
Fig. 4. Mode mixity of the designed hybrid laminates with a crack length of 50 mm and an opening displacement of 4 mm: a. GFRP/Al, b. GFRP/Cu.
Table 5
Contact angles (𝜃) of metal substrates with different treatments (degrees).
3. Results and discussion
Metal
3.1. Contact angle test
Contact angle tests were conducted on the surfaces of aluminum
and copper substrates, each with dimensions of 25 mm × 40 mm, and
subjected to various surface treatments as discussed in the previous
section. The experiments were carried out utilizing an OCA15 Neurtek
goniometer, employing three distinct liquids: water, ethylene glycol
55%, and n-hexadecane. After adjusting the tester, camera, and the
substrate, a small droplet (around 5 μL) of each liquid was placed on
the surface of the substrates, by using the dosing unit. The image of the
droplet was captured using an optical camera. The static contact angle
was measured by the SCA15 software after defining the base line of
the surface and droplet shape of the liquid droplet. The Owens, Wendt,
Rabel, and Kaelble method (OWRK-model) was used as a standard
method for calculating the Surface Free Energy (𝜎 s ) of metal surfaces
from the contact angle (𝜃) [61,62]. The formula of this method and
the properties of the mentioned liquids are presented in Appendix C of
the supplementary data. The experimentally determined contact angle
Al
Cu
Surface treatment
Liquid
Water
Ethylene glycol
n-hexadecane
Without treatment
P2 etching
NaOH/HNO3
Sol–Gel
97.6
84.7
56.4
71.5
66.9
55.4
39.6
82.5
14
0
0
0
Without treatment
FeCl3 /HCl/Glycerol
Sol–Gel
107.7
100.9
83.5
74.7
41.3
53.9
16
0
0
(𝜃), which is the angle between the metal–liquid interface and the
liquid–gas interface, is shown in Table 5.
Accordingly, 𝜃 = 0◦ indicates the liquid completely wetted the
substrate, 0◦ < 𝜃 < 90◦ shows high wettability, 90◦ < 𝜃 < 180◦ indicates
low wettability of the surface.
Fig. 5 shows the ethylene glycol droplet and its contact angles
on aluminum and copper substrates treated with Sol–Gel anodizing.
The image reveals that the copper substrate exhibited a lower contact
5
Composite Structures 349-350 (2024) 118509
M. Niazi et al.
Fig. 5. Ethylene glycol droplet on substrates with Sol–Gel anodizing treatment: a. Aluminum substrate, b. Copper substrate.
Table 6
Surface free energy of different substrates, 𝜎 s (mJ/m2 ).
Table 7
Surface roughness of different substrates, 𝑆𝑎 (μm).
Surface treatment
Al
Cu
Surface treatment
Al
Cu
Without treatment
P2-etching
NaOH/HNO3
Sol–Gel
FeCl3 /HCl/Glycerol
27.8
31.8
43
30
–
27
–
–
32.3
34.8
Without treatment
P2-etching
NaOH/HNO3
Sol–Gel
FeCl3 /HCl/Glycerol
0.45
0.43
0.56
0.62
–
0.41
–
–
0.71
0.56
angle, indicating higher wettability compared to the aluminum substrate under the same treatment condition. This enhanced wettability
suggests potentially stronger adhesive bonding between the copper and
the GFRP laminate compared to the one between aluminum and GFRP
laminate.
The surface free energy (𝜎 s ) of various substrates, as presented in
Table 6, reveals notable variations due to different chemical treatments. Chemical etching with NaOH/HNO3 significantly affected the
surface free energy of aluminum (43 mJ/m2 ), compared to untreated
substrate (27.8 mJ/m2 ), Sol–Gel anodizing (30 mJ/m2 ), and P2-etching
(31.8 mJ/m2 ). For copper, chemical etching with FeCl3 /HCl/Glycerol
resulted in a higher surface free energy of 34.8 mJ/m2 , compared to
untreated substrates (27 mJ/m2 ) and Sol–Gel anodizing (32.3 mJ/m2 ).
roughness of both substrates, especially the copper substrate with a
peak of 0.71 μm, whereas untreated copper displayed the lowest roughness of 0.41 μm. NaOH/HNO3 of aluminum and FeCl3 /HCl/Glycerol of
copper substrates both resulted in similar surface roughness of 0.56 μm.
In the context of P2-etching of aluminum substrates, a noticeable reduction in roughness from 0.45 μm to 0.43 μm was observed in comparison
to untreated specimens. This phenomenon arose from the chemical
action induced by ferric sulfate and sulfuric acid during the etching
process. This chemical interaction targeted and diminished the height
of peak points, which were originally formed during the preliminary
cleaning step with sanding paper in the pretreatment stage.
3.2. Surface roughness
The mode I interlaminar fracture toughness of each specimen was
measured using the Modified Beam Theory (MBT) based on the ASTM
D5528 standard. Accordingly, the fracture toughness is given by:
3.3. Interlaminar fracture toughness
Coherence Scanning Interferometry (CSI) was used to measure the
surface roughness and topography with nanometric precision. The 3D
surface characterization of the specimens was performed using a noncontact measurement system NPFLEX from Bruker. It is an optical
unit with an XLWD Industrial Objective with 10 times magnification,
a working distance of 34 mm, an aperture of 0.28 and an optical
resolution based on Sparrow Criteria at 535 nm of 1 μm. The vertical
resolution was less than 15 nm.
The surface roughness measurement, 𝑆𝑎 , was determined as the
arithmetic mean of the absolute heights within a sampling area, 4.8 mm
× 4.8 mm square region extracted from the center of the specimen:
𝑆𝑎 =
1
|𝑧(𝑥, 𝑦)|𝑑𝑥𝑑𝑦
𝐴∫
𝐺𝐼 =
3𝑃 𝛿
2𝑏(𝑎 + |𝛥|)
(11)
where 𝑃 is the applied opening load, 𝛿 is the load-point vertical displacement, and 𝑏 is the specimen width. In the context of the MBT, |𝛥|
is used to define an effective crack length which accounts for non-zero
rotations of the crack tip. The fracture toughness in each DCB test was
determined by analyzing the stable part of the each resistance curve (Rcurve). Three separate samples were examined for each configuration
to assess setup repeatability. In the following subsections, the averages
of the R-curve and load–displacement are presented for each group
of specimens. Detailed information for each individual specimen can
be found in Appendix E of the supplementary data. Additionally, the
coefficient of variation (CV) was calculated for each dataset to illustrate
the extent of variation within each group of data, focusing on the stable
plateau of the R-curve.
(10)
where 𝐴 is the sampling area in the xy plane, and z is the height of
each point of the sample. Surface roughness parameter, 𝑆𝑎 , directly
influence the contact area and strength of mechanical interlocking
between the metal and composite layers. Higher 𝑆𝑎 values typically
result in increased micro-asperities and contact points, enhancing the
mechanical interlocking capability. This, in turn, improves the load
transfer across the interface and contributes to the overall fracture
toughness by resisting crack propagation [63]. Fig. 6 depicts the 2D
surface profiles of aluminum and copper substrates with Sol–Gel anodizing. The 2D surface profiles for other treatments can be found in
Appendix D of the supplementary data.
The surface roughness of different substrates is presented in Table 7.
Accordingly, Sol–Gel anodizing had the strongest effect on the surface
3.3.1. Fracture toughness of GFRP laminate
Thin-ply E-glass laminates were tested in DCB experiments with
two distinct configurations: unidirectional[0]72S and quasi-isotropic
[−45◦ /90◦ /45◦ /0◦ ]18S . The quasi-isotropic configuration, strategically
designed within the hybrid laminate to mitigate thermal residual stress,
was selected for comparison with the ASTM D5528-standardized unidirectional configuration. This comparison aims to explore variations
in crack propagation behavior between the two configurations. The Rcurve and load–displacement data for both configurations are depicted
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M. Niazi et al.
Fig. 6. The 2D surface profiles of substrates with Sol–Gel anodizing: a. Aluminum substrate, b. Copper substrate.
Fig. 7. The fracture toughness results of GFRP laminate, a. R-Curve, b. Load–displacement.
Fig. 8. The lateral view of delamination propagation of GFRP laminate: a. Unidirectional laminate, b. Quasi-isotropic laminate.
Table 8
Average 𝐺𝐼𝐶 and CV of the GFRP laminate.
𝐺𝐼𝐶 (N/mm)
CV (%)
UD ([0]72S )
QI ([−45◦ /90◦ /45◦ /0◦ ]18S )
0.86
6
0.52
4
3.3.2. Fracture toughness of GFRP/Al laminate
The R-curve and load–displacement behavior of GFRP/Al with different surface treatments on the aluminum are shown in Fig. 9. The
results show how different treatments of aluminum significantly influenced the R-curve shapes, indicating varied resistance to crack initiation and propagation at the interface. The average 𝐺𝐼𝐶 values of different treatments, along with their relative CVs, are presented in Table 9.
Configurations with Sol–Gel anodizing treatment achieved the highest
fracture toughness, with an average 𝐺𝐼𝐶 of 1.73 N/mm. In contrast,
untreated samples had an average 𝐺𝐼𝐶 of 0.37 N/mm. Specimens with
NaOH/HNO3 treatment yielded a 𝐺𝐼𝐶 of 0.73 N/mm, while the ones
with P2-etching resulted in 𝐺𝐼𝐶 of 0.66 N/mm. Notably, P2-etching
and Sol–Gel anodizing configurations exhibited higher repeatability,
with average CVs of 4.2% and 7%, respectively. In comparison, untreated samples and those subjected to NaOH/HNO3 etching showed
lower repeatability, with average CVs of approximately 40% and 47%,
respectively.
Fig. 10 provides a visual representation of the surface of both arms
after debonding, as well as the lateral fracture morphology observed
during the test, for each configuration of GFRP/Cu. In this figure,
in Fig. 7. The average 𝐺𝐼𝐶 values of both configurations, along with
their relative CVs, are presented in Table 8. The unidirectional (UD)
GFRP exhibited an average value of 0.86 N/mm for 𝐺𝐼𝐶 with a CV
of 6%, while the quasi-isotropic (QI) configuration showed an average
𝐺𝐼𝐶 of 0.52 N/mm with a CV of 4%. A similar study conducted by [64]
reached the same conclusion, finding that unidirectional laminates
result in higher 𝐺𝐼𝐶 values compared to multidirectional laminates.
This difference can be attributed to the higher amount of fiber bridging
observed in the unidirectional configuration, which originated not only
from the first ply but also involved fiber bridging from adjacent plies
near the interface (see Fig. 8a). In contrast, for the quasi-isotropic laminate, fiber bridging primarily arose from the first ply at the interface
(see Fig. 8b).
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Composite Structures 349-350 (2024) 118509
M. Niazi et al.
Fig. 9. The fracture toughness results of GFRP/Al with different surface treatment on aluminum, a. R-Curve, b. Load–displacement.
Table 9
Average 𝐺𝐼𝐶 and CV of hybrid GFRP/Al with different surface treatments.
𝐺𝐼𝐶 (N/mm)
CV (%)
Table 10
Average 𝐺𝐼𝐶 and CV of hybrid GFRP/Cu with different surface treatments.
Without treatment
P2-etching
NaOH/HNO3
Sol–Gel
0.37
40
0.66
4.2
0.73
47
1.73
7
𝐺𝐼𝐶 (N/mm)
CV (%)
Without treatment
FeCl3 /HCl/Glycerol
Sol–Gel
0.26
42
0.12
15
1.8
4
for untreated specimens and NaOH/HNO3 etched samples, two different examples are presented due to lower repeatability, showcasing
different possible phenomena. Conversely, for P2-etched and Sol–Gel
anodized samples, where there was higher repeatability and similar
surface morphology, only one example is shown. Fig. 10a and b display
untreated specimens, highlighting different interface phenomena across
samples. While the specimen presented in Fig. 10a exhibited poor
bonding and straightforward crack propagation, resulting in lower 𝐺𝐼𝐶 ,
the specimen presented in Fig. 10b showed occasional crack jumping
and significant fiber bridging, leading to higher 𝐺𝐼𝐶 . These differences
explain the inconsistent behavior and repeatability observed. Fig. 10c
depicts P2-etched specimens with increased fiber bridging and crack
jumping, resulting in higher 𝐺𝐼𝐶 consistently among treated samples.
Fig. 10d and e illustrate NaOH/HNO3 treated specimens, showing
mixed results. While one sample had poor bonding at the interface,
another showed extensive fiber bridging and crack jumping, contributing to higher 𝐺𝐼𝐶 . This variation indicated differences in treatment
effectiveness and specimen behavior. Fig. 10f shows the outcomes of
Sol–Gel anodizing treatment, highlighting features such as extensive
fiber bridging, bifurcation, and widespread crack jumping across the
entire surface of treated specimens. A similar fracture mechanism was
observed in [65]. This treatment showed superior fracture toughness
with more plies involved in failure compared to other treatments,
indicating strong bonding between metal and GFRP arm when using
this treatment.
as well as the lateral fracture morphology observed during the test,
for each configuration of GFRP/Cu. Fig. 12a shows that the specimens
without treatment exhibited small areas of crack jumping, with the
remaining areas showing poor bonding. The lateral view also reveals
occasional instances of fiber bridging amidst predominantly straightforward crack propagation. Fig. 12b shows the results of specimens
treated with FeCl3 /HCl/Glycerol. From the residual color traces of
FeCl3 on the GFRP surface, it is evident that this treatment resulted
in the formation of an unstable layer on the copper surface. This
instability contributed to the poor bonding. In this configuration, no
fiber bridging or other secondary fracture mechanisms were observed.
Fig. 12c depicts the results of Sol–Gel anodizing. Widespread crack
jumping, fiber bridging, and bifurcation mechanisms across the entire
area of the specimen indicate that this treatment created a superior
bond between the metal and the composite. Consequently, it offered
superior interlaminar fracture toughness by inhibiting crack initiation
and propagation. In this particular configuration, when the total fracture toughness exceeded 1.8 N/mm, the metal underwent yielding,
dissipating energy on plastic deformation. The fracture toughness continued to increase without stabilization due to metal yielding, the total
𝐺𝐼𝐶 value for this treatment was identified as 1.8 N/mm. In terms
of repeatability, specimens without treatment exhibited a high CV of
42%, indicating low consistency. In contrast, specimens treated with
FeCl3 /HCl/Glycerol had a CV of 15%, while those treated with Sol–Gel
demonstrated the highest repeatability with a CV of 4%.
3.3.3. Fracture toughness of GFRP/Cu laminate
The R-curve and load–displacement behavior of GFRP/Cu with
various surface treatments on the copper are shown in Fig. 11. As
illustrated in this figure, different treatments applied to copper alloys
markedly affect the shapes of the R-curves and the load–displacement
behavior of GFRP/Cu, highlighting varying levels of resistance to crack
initiation and propagation at the interface.
Based on the observation of the stable part of the R-curve in Fig. 11a
(compiled in Table 10), specimens without surface treatment exhibited
an average fracture toughness (𝐺𝐼𝐶 ) of 0.26 N/mm. However, the
application of FeCl3 /HCl/Glycerol treatment to the copper resulted in
a notable reduction of 𝐺𝐼𝐶 to 0.12 N/mm compared to the untreated
configuration. Conversely, Sol–Gel anodizing of copper significantly
enhanced the mode I fracture toughness of the hybrid laminate. Fig. 12
illustrates the surfaces of the copper and GFRP arms after debonding,
3.3.4. Key insights and patterns across treatments
In the previous subsections, the impact of various treatments on the
surface free energy and surface roughness of aluminum and copper, as
well as the interlaminar fracture toughness of GFRP/Al and GFRP/Cu
laminates, were discussed in detail. In this subsection, the possible
correlations between these parameters are discussed. The total results
for the different configurations are shown in Fig. 13. To highlight
variations in different parameters within a single image and facilitate
comparison, the focus has shifted from absolute surface free energy
to considering the ratio of surface energy. This ratio was determined
by dividing the surface energy of each treatment by that observed in
the benchmark (without treatment) configuration. In general, surface
roughness appeared to correlate with interlaminar fracture toughness.
However, there was an exception: in the case of P2-etching, the roughness was slightly lower than that without treatment; however, the
8
Composite Structures 349-350 (2024) 118509
M. Niazi et al.
Fig. 10. The surface morphology of GFRP and aluminum arms of GFRP/Al laminate with different surface treatment on aluminum (left side: aluminum, right side: GFRP), along
with the lateral view of delamination propagation: a and b. Without treatment, c. P2-etching, d and e. NaOH/HNO3 treatment, f. Sol–Gel anodizing.
Fig. 11. The fracture toughness results of GFRP/Cu with different surface treatment on copper, a. R-Curve, b. Load–displacement.
𝐺𝐼𝐶 of P2-etching was approximately 2 times higher than that of
the untreated samples. This can be attributed to the fact that both
treatments were cleaned with sandpaper, and the P2-based treatment
possibly removing some of the surface roughness peaks. For specimens
with Sol–Gel treatment, both the surface roughness and interlaminar
fracture toughness were the highest in both configurations (GFRP/Al
and GFRP/Cu), whereas the surface free energy was lower compared to
other treatments. This indicates that surface free energy, which reflects
the wettability of the surface, did not correlate with interlaminar fracture toughness. Another example is the FeCl3/HCl/Glycerol treatment,
which, despite having the highest surface free energy among copper
treatments, exhibited the lowest interlaminar fracture toughness. Although this treatment enhanced wettability, it created an unstable layer
on the copper surface. This instability resulted in poor bonding with
the GFRP, leading to rapid delamination propagation during the DCB
test without fiber bridging, consequently resulting in a low 𝐺𝐼𝐶 . This
shows that achieving good wettability through metal surface treatment alone does not guarantee strong bonding in hybrid laminates.
The stability and strength of the created layer on the metal surface
after treatment are also crucial factors. The high mechanical interlocking and chemical bonding, along with the stability of the created
layer on the metal surface provided by Sol–Gel anodizing, significantly enhanced the Mode I interlaminar fracture toughness of GFRP/Al
and GFRP/Cu composites by inhibiting crack initiation and propagation. The R-curve analysis revealed a substantial increase in fracture
toughness from 0.37 N/mm to 1.72 N/mm and from 0.26 N/mm to
1.8 N/mm for GFRP/Al and GFRP/Cu, respectively, in comparison to
without treatment configuration. This treatment outperformed other
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Composite Structures 349-350 (2024) 118509
M. Niazi et al.
Fig. 12. The surface morphology of GFRP and copper arms of GFRP/Cu laminate with different surface treatment on copper (left side: copper, right side: GFRP), along with the
lateral view of delamination propagation: a. Without treatment, b. FeCl3 /HCl/Glycerol, c. Sol–Gel anodizing.
Fig. 13. Comparison of different obtained parameters: a. GFRP/Al, b.GFRP/Cu.
chemical treatments such as P2-etching and NaOH/HNO3 etching, and
FeCl3 /HCl/Glycerol treatment, underscoring its efficacy in improving
the structural integrity of hybrid laminates for structural batteries
applications. Furthermore, high repeatability was observed in Sol–Gel
anodizing and P2-etching treatments, demonstrating the reliability of
these methodologies. In contrast, specimens without treatment and
NaOH/HNO3 etching exhibited lower repeatability.
It should be noted that the calculated mode I fracture toughness
values were often affected by multiple secondary fracture mechanisms
such as crack jumping and fiber bridging. Collectively, these factors
indicate that the obtained values reflect structural properties rather
than purely material properties, especially regarding effective treatments such as Sol–Gel anodizing. Despite achieving consistent interlaminar fracture toughness across different samples with Sol–Gel treatment, these complexities in interlaminar fracture behavior may pose
challenges for numerical verification.
treatment of copper, despite enhancing surface free energy and thereby
improving wettability, resulted in poor bonding with GFRP due to the
formation of an unstable layer on the copper surface. These findings underscore that increased metal surface wettability alone does not always
guarantee effective bonding with composite laminate. Beside metal surface roughness, stability and strength of the created layer on the surface
of metal are also critical factors for achieving robust bonding in hybrid
laminates. Sol–Gel anodizing emerged as the best metal treatment,
increasing the mode I fracture toughness of GFRP/Al and GFRP/Cu
by about five and seven times, respectively, compared to the baseline
configurations without treatment. This technique outperformed the
chemical treatment methods such as P2-etching, NaOH/HNO3 etching,
and FeCl3 /HCl/Glycerol treatment. The remarkable repeatability observed in Sol–Gel anodizing and P2-etching treatments enhanced the
reliability of these techniques. In contrast, untreated specimens and
NaOH/HNO3 etching presented challenges with lower repeatability,
warranting caution when using these methods.
4. Conclusions
A comprehensive analysis of GFRP/Al and GFRP/Cu hybrid laminates with varying surface treatments on the metal was presented.
Critical parameters such as surface free energy, surface roughness, and
Mode I interlaminar fracture toughness were analyzed. The findings
revealed that surface roughness of the metal generally correlates with
interlaminar fracture toughness of hybrid laminates, although exceptions exist, such as with P2-etching. Sol–Gel treatment resulted in the
highest surface roughness and superior interlaminar fracture toughness
despite lower surface free energy. Conversely, the FeCl3 /HCl/Glycerol
CRediT authorship contribution statement
Maryam Niazi: Writing – original draft, Methodology, Formal analysis, Conceptualization. Federico Danzi: Writing – review & editing,
Supervision, Formal analysis, Conceptualization. Ricardo Carbas: Validation, Investigation. Pedro P. Camanho: Writing – review & editing,
Supervision, Methodology.
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Composite Structures 349-350 (2024) 118509
M. Niazi et al.
Declaration of competing interest
[21] Ochoa-Putman C, Vaidya UK. Mechanisms of interfacial adhesion in metal–
polymer composites–effect of chemical treatment. Composites A 2011;42(8):906–
15.
[22] Park SY, Choi WJ, Choi HS, Kwon H, Kim SH. Recent trends in surface treatment
technologies for airframe adhesive bonding processing: a review (1995–2008). J
Adhes 2010;86(2):192–221.
[23] Aghamohammadi H, Abbandanak SNH, Eslami-Farsani R, Siadati SH. Effects
of various aluminum surface treatments on the basalt fiber metal laminates
interlaminar adhesion. Int J Adhes Adhes 2018;84:184–93.
[24] Lawcock G, Ye L, Mai Y-W, Sun C-T. The effect of adhesive bonding between aluminum and composite prepreg on the mechanical properties of
carbon-fiber-reinforced metal laminates. Compos Sci Technol 1997;57(1):35–45.
[25] Hu Y, Yuan B, Cheng F, Hu X. NaOH etching and resin pre-coating treatments
for stronger adhesive bonding between CFRP and aluminium alloy. Composites
B 2019;178:107478.
[26] Joshi S, Fahrenholtz WG, O’Keefe MJ. Effect of alkaline cleaning and activation
on aluminum alloy 7075-T6. Appl Surf Sci 2011;257(6):1859–63.
[27] Jin Z, Cai C, Hashimoto T, Yuan Y, Kang D, Hunter J, Zhou X. Alkaline etching
and desmutting of aluminium alloy: The behaviour of Mg2Si particles. J Alloys
Compd 2020;842:155834.
[28] Gonzalez-Canche N, Flores-Johnson E, Cortes P, Carrillo J. Evaluation of surface
treatments on 5052-H32 aluminum alloy for enhancing the interfacial adhesion
of thermoplastic-based fiber metal laminates. Int J Adhes Adhes 2018;82:90–9.
[29] Zain NM, Ahmad SH, Ali ES. Effect of surface treatments on the durability
of green polyurethane adhesive bonded aluminium alloy. Int J Adhes Adhes
2014;55:43–55.
[30] McGarel W, Farnham H, McGuckin R. Influence of pretreatment rinse waters
on the durability of adhesive-bonded aluminium alloys. Int J Adhes Adhes
1986;6(2):89–92.
[31] Prolongo S, Ureña A. Effect of surface pre-treatment on the adhesive strength of
epoxy–aluminium joints. Int J Adhes Adhes 2009;29(1):23–31.
[32] İlhan Ç, Gürbüz R. Effects of chromic acid anodizing on fatigue behavior of 7050
T7451 aluminum alloy. Mater Test 2021;63(9):805–10.
[33] Critchlow GW, Yendall KA, Bahrani D, Quinn A, Andrews F. Strategies for the
replacement of chromic acid anodising for the structural bonding of aluminium
alloys. Int J Adhes Adhes 2006;26(6):419–53.
[34] Davis G, Sun T, Ahearn J, Venables J. Application of surface behaviour diagrams
to the study of hydration of phosphoric acid-anodized aluminium. J Mater Sci
1982;17:1807–18.
[35] Hardwick D, Ahearn J, Desai A, Venables J. Environmental durability of
phosphoric acid anodized aluminium adhesive joints protected with hydration
inhibitors. J Mater Sci 1986;21:179–87.
[36] Barkhimer J, Erich M, Nair G. Effect of time delay between etching and
adhesive bonding (‘‘outlife’’ time) on lap-shear strength of aluminum alloys using
environmentally-friendly P2 etch. 2015.
[37] Seifzadeh D, Golmoghani-Ebrahimi E. Formation of novel and crack free
nanocomposites based on sol gel process for corrosion protection of copper. Surf
Coat Technol 2012;210:103–12.
[38] Park SY, Choi WJ, Choi HS. A review of the recent developments in surface
treatment techniques for bonded repair of aluminum airframe structures. Int J
Adhes Adhes 2018;80:16–29.
[39] Voevodin N, Grebasch N, Soto W, Kasten L, Grant J, Arnold F, Donley M.
An organically modified zirconate film as a corrosion-resistant treatment for
aluminum 2024-T3. Prog Org Coat 2001;41:287–93.
[40] Çakır O, Temel H, Kiyak M. Chemical etching of Cu-ETP copper. J Mater Process
Technol 2005;162:275–9.
[41] Cabello M, Turon A, Zurbitu J, Renart J, Sarrado C, Martínez F. Progressive
failure analysis of DCB bonded joints using a new elastic foundation coupled
with a cohesive damage model. Eur J Mech A Solids 2017;63:22–35.
[42] Shi W, Zhang C, Wang B, Li M, Zhang C. Mode I interlaminar fracture toughness
of two-dimensional continuous fiber reinforced ceramic matrix composites using
wedge-loaded double cantilever beam method. Composites A 2023;168:107466.
[43] Todo M, Jar P-Y. Study of mode-I interlaminar crack growth in DCB specimens
of fibre-reinforced composites. Compos Sci Technol 1998;58(1):105–18.
[44] Prasad NE, Wanhill RJ. Aerospace materials and material technologies, vol. 1,
Springer; 2017.
[45] Asundi A, Choi AY. Fiber metal laminates: an advanced material for future
aircraft. J Mater Process Technol 1997;63(1–3):384–94.
[46] Sinmazçelik T, Avcu E, Bora MÖ, Çoban O. A review: Fibre metal laminates, background, bonding types and applied test methods. Mater Des
2011;32(7):3671–85.
[47] Niazi M, Sadr MH. Numerical shear buckling investigation of GLAREs with initial
delamination. Adv Mater Sci Eng 2022;2022(1):9059917.
[48] Heggemann T, Homberg W. Deep drawing of fiber metal laminates for
automotive lightweight structures. Compos Struct 2019;216:53–7.
[49] Banea M, Rosioara M, Carbas R, Da Silva L. Multi-material adhesive joints for
automotive industry. Composites B 2018;151:71–7.
[50] Wang W, Fernandes RL, De Freitas ST, Zarouchas D, Benedictus R. How pure
mode I can be obtained in bi-material bonded DCB joints: a longitudinal
strain-based criterion. Composites B 2018;153:137–48.
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.
Data availability
Data will be made available on request.
Acknowledgments
The authors would like to acknowledge the support of the Portuguese Foundation for Science and Technology under the FCT grant for
the UIDP/50022/2020 LAETA project. Additionally, MN expresses gratitude to LAETA and FCT for the doctoral scholarship UI/BD/151558/
2021.
Appendix A. Supplementary data
Supplementary material related to this article can be found online
at https://doi.org/10.1016/j.compstruct.2024.118509.
References
[1] Lim GJH, Chan KK, Sutrisnoh NAA, Srinivasan M. Design of structural batteries:
carbon fibers and alternative form factors. Mater Today Sustain 2022;20:100252.
[2] Kalnaus S, Asp LE, Li J, Veith GM, Nanda J, Daniel C, Chen XC, Westover A,
Dudney NJ. Multifunctional approaches for safe structural batteries. J Energy
Storage 2021;40:102747.
[3] Danzi F, Salgado RM, Oliveira JE, Arteiro A, Camanho PP, Braga MH. Structural
batteries: A review. Molecules 2021;26(8):2203.
[4] Danzi F, Camanho PP, Braga MH. An all-solid-state coaxial structural battery
using sodium-based electrolyte. Molecules 2021;26(17):5226.
[5] Fu Y, Chen Y, Yu X, Zhou L. Fiber metal laminated structural batteries with multifunctional solid polymer electrolytes. Compos Sci Technol
2022;230:109731.
[6] Chen J, Zhou Y, Islam MS, Cheng X, Brown SA, Han Z, Rider AN, Wang CH.
Carbon fiber reinforced Zn–MnO2 structural composite batteries. Compos Sci
Technol 2021;209:108787.
[7] Pattarakunnan K, Galos J, Das R, Mouritz A. Tensile properties of multifunctional composites embedded with lithium-ion polymer batteries. Composites A
2020;136:105966.
[8] Ye J, Ji X, Liu Z, Liu K, Li J, Wang R, Wang J, Lei Q. Carbon fiber reinforced structural battery composites: Progress and challenges toward industrial
application. Composites B 2024;111411.
[9] Asp LE, Johansson M, Lindbergh G, Xu J, Zenkert D. Carbon fibre composite
structural batteries: A review. Funct Compos Struct 2019.
[10] Xu J, Varna J. Matrix and interface cracking in cross-ply composite structural
battery under combined electrochemical and mechanical loading. Compos Sci
Technol 2020;186:107891.
[11] Qiu Q. Brittle fracture mechanisms of glass-fiber reinforced polymer insulators.
Oregon Graduate Institute of Science and Technology; 1995.
[12] Moyer K, Boucherbil NA, Zohair M, Eaves-Rathert J, Pint CL. Polymer reinforced
carbon fiber interfaces for high energy density structural lithium-ion batteries.
Sustain Energy Fuels 2020;4(6):2661–8.
[13] Di Mauro G, Russo P, Guida M. Impact response of a composite structural battery
in low-velocity tests. J Mater Eng Perform 2023;32(9):3881–7.
[14] Chen Y, Wang Y, Wang H. Research progress on interlaminar failure behavior
of fiber metal laminates. Adv Polym Technol 2020;2020(1):3097839.
[15] McDowell MT, Lee SW, Nix WD, Cui Y. 25Th anniversary article: understanding
the lithiation of silicon and other alloying anodes for lithium-ion batteries. Adv
Mater 2013;25(36):4966–85.
[16] McDowell MT, Xia S, Zhu T. The mechanics of large-volume-change
transformations in high-capacity battery materials. Extreme Mech Lett
2016;9:480–94.
[17] Zhang Z, Wang R, Zeng J, Shi K, Zhu C, Yan X. Size effects in sodium ion
batteries. Adv Funct Mater 2021;31(52):2106047.
[18] Carlstedt D, Runesson K, Larsson F, Tu V, Jänicke R, Asp LE. Computational
modelling of structural batteries accounting for stress-assisted convection in the
electrolyte. Int J Solids Struct 2022;238:111343.
[19] Jeong H, Jang J, Jo C. A review on current collector coating methods for
next-generation batteries. Chem Eng J 2022;446:136860.
[20] Zhu P, Driscoll EH, Dong B, Sommerville R, Zorin A, Slater PR, Kendrick E.
Direct reuse of aluminium and copper current collectors from spent lithium-ion
batteries. Green Chem 2023;25(9):3503–14.
11
Composite Structures 349-350 (2024) 118509
M. Niazi et al.
[58] Hibbitt, Karlsson, Sorensen. Abaqus/CAE user’s manual. Hibbitt, Karlsson &
Sorensen, Incorporated; 2002.
[59] Rybicki E, Kanninen M. A finite element calculation of stress intensity factors
by a modified crack closure integral. Eng Fract Mech 1977;9(4):931–8. http:
//dx.doi.org/10.1016/0013-7944(77)90013-3, URL https://www.sciencedirect.
com/science/article/pii/0013794477900133.
[60] Krueger R. Virtual crack closure technique: History, approach, and applications.
Appl Mech Rev 2004;57:109–43.
[61] Owens DK, Wendt R. Estimation of the surface free energy of polymers. J Appl
Polym Sci 1969;13(8):1741–7.
[62] Loeb GI, Schrader ME. Modern approaches to wettability: theory and
applications. Springer Science & Business Media; 2013.
[63] Gent A, Lin C-W. Model studies of the effect of surface roughness and mechanical
interlocking on adhesion. J Adhes 1990;32(2–3):113–25.
[64] Shokrieh M, Heidari-Rarani M. Effect of stacking sequence on R-curve behavior
of glass/epoxy DCB laminates with 0//0 crack interface. Mater Sci Eng A
2011;529:265–9.
[65] Lima R, Tao R, Bernasconi A, Carboni M, Carrere N, De Freitas ST. Uncovering
the toughening mechanisms of bonded joints through tailored CFRP layup.
Composites B 2023;263:110853.
[51] Mujika F, Tsokanas P, Arrese A, Valvo PS, da Silva LF. Mode decoupling in
interlaminar fracture toughness tests on bimaterial specimens. Eng Fract Mech
2023;290:109454.
[52] Shanmugam L, Naebe M, Kim J, Varley RJ, Yang J. Recovery of mode I selfhealing interlaminar fracture toughness of fiber metal laminate by modified
double cantilever beam test. Compos Commun 2019;16:25–9.
[53] Koord J, Völkerink O, Petersen E, Hühne C. Effect of low temperature on mode
I and mode II interlaminar fracture toughness of CFRP-steel hybrid laminates.
Composites B 2023;262:110773.
[54] Wegman RF, Van Twisk J. Surface preparation techniques for adhesive bonding.
William Andrew; 2012.
[55] Rider A, Williams I, Shum E, Mirabella L. Environmental durability trial of
bonded composite repairs to metallic aircraft structure. DTIC Document; 2005.
[56] Ouyang Z, Ji G, Li G. On approximately realizing and characterizing pure mode-I
interface fracture between bonded dissimilar materials. 2011.
[57] Gong X, Hernandez F, Verchery G. Fracture toughness of adhesive bonded
composite joints under mixed mode loading. In: ICCM12 conference. 1999.
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