![(a) Schematic of probe configuration for EC testing; (b) electric and dielectric behaviour of CFRP. Carbon fibre materials can be simplified as a network of resistors and capacitors, adapted from [24].](https://www.researchgate.net/publication/365661009/figure/fig1/AS:11431281099012896@1669173684918/a-Schematic-of-probe-configuration-for-EC-testing-b-electric-and-dielectric_Q320.jpg)
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J. Manuf. Mater. Process. 2022, 6, 151. https://doi.org/10.3390/jmmp6060151 www.mdpi.com/journal/jmmp
Article
Quality Analysis of Weld-Line Defects in Carbon Fibre
Reinforced Sheet Moulding Compounds by Automated Eddy
Current Scanning
Nessa Fereshteh-Saniee
1,
*, Neil Reynolds
1
, Danielle Norman
1
, Connie Qian
1
, David J. Armstrong
2
,
Paul Smith
3
, Richard Kupke
3
, Mark A. Williams
1
and Kenneth Kendall
1
1
WMG, University of Warwick, Coventry CV4 7AL, UK
2
Department of Physics, University of Warwick, Coventry CV4 7AL, UK
3
SURAGUS GmbH, 01109 Dresden, Germany
* Correspondence: n.fereshteh-saniee@warwick.ac.uk
Abstract: Discontinuous fibre reinforced composites enable the manufacture of integrated structural
components via the complex flow process of compression moulding. However, such processes can
lead to the formation of detrimental weld-lines. Here, the meso-structure of carbon fibre sheet
moulding compounds (C-SMC) was analysed using conventional non-destructive techniques and
automated eddy current (EC) scanning, as well as destructive methods, in an attempt to identify
defects such as weld-lines in this class of materials. Compression-moulded plaques with forced
weld-lines in two different configurations (adjacent and opposing flow joints) were analysed, show-
ing up to 80% strength reduction versus a defect-free plaque. The EC-determined local fibre orien-
tation and elucidated local microstructure matched those obtained using conventional techniques,
showing a dramatic fibre tow alignment parallel to the weld-lines. It was found that failure occurred
in proximity to the “non-uniformity” defect regions identified by EC analyses, demonstrating the
use of robot-guided EC for successful defect detection in C-SMC structures.
Keywords: carbon fibres; discontinuous reinforcement; compression moulding; non-destructive
testing; eddy current
1. Introduction
In an endeavour to reduce vehicle weight in the automotive industry, discontinuous
fibre composites (DFCs) such as carbon fibre sheet moulding compounds (C-SMC) are
under consideration. C-SMCs are ideal light-weighting candidates, due to their high de-
gree of formability and consequent potential to replace highly integrated structural auto-
motive metal components. C-SMCs offer rapid high-volume processing and excellent
mass-specific performance. Owing to their discontinuous fibre reinforcement, and hence
their ability to flow well, they can be used to mould complex 3D geometries in a one-shot
high volume manufacturing route, which is not possible with their continuous fibre rein-
forced counterparts [1,2].
The mechanical properties of “forced defect free” (monolithic) DFCs in a two-dimen-
sional format have previously been studied in the literature [3–6]. A high degree of heter-
ogeneity has been observed in the microstructure of SMCs, with randomly oriented
strands (ROS) yielding near quasi-isotropic in-plane stiffness [3,5]. However, local varia-
tions in fibre orientation and fibre volume fraction, arising because of the process-induced
anisotropic microstructure, affect tensile strength due to the stress concentrations caused
by discontinuities in the fibre architecture [4]. During the compression moulding process,
a flow-induced meso-structure arises that can result in critical material flaws, such as resin
rich pockets between the fibre tows, swirls, and trapped air (voids) [4,7]. In addition, a
Citation: Fereshteh-Saniee, N.;
Reynolds, N.; Norman, D.; Qian, C.;
Armstrong, D.J.; Smith, P.; Kupke,
R.; Williams, M.A.; Kendall, K.
Quality Analysis of Weld-Line
Defects in Carbon Fibre Reinforced
Sheet Moulding Compounds by
Automated Eddy Current Scanning.
J
. Manuf. Mater. Process. 2022, 6, 151.
https://doi.org/10.3390/jmmp6060151
Academic Editor: Steven Y. Liang
Received: 25 October 2022
Accepted: 17 November 2022
Published: 22 November 2022
Publisher’s Note: MDPI stays neu-
tral with regard to jurisdictional
claims in published maps and institu-
tional affiliations.
Copyright: © 2022 by the authors. Li-
censee MDPI, Basel, Switzerland.
This article is an open access article
distributed under the terms and con-
ditions of the Creative Commons At-
tribution (CC BY) license (https://cre-
ativecommons.org/licenses/by/4.0/).
J. Manuf. Mater. Process. 2022, 6, 151 2 of 25
combination of typically complex part design, high volume fraction, and characteristic
fibre length in advanced C-SMCs, promotes specific flow conditions that lead to process-
induced complex anisotropic microstructures [8], such as the formation of weld-line de-
fects (WLDs), where two flow fronts meet.
Very limited research has been done on the formation of weld-lines in SMCs, whereas
the effect of WLDs on short fibre reinforced (thermoplastic) composites has been studied
to a great extent. In these cases, the effect of WLDs on the mechanical and physical prop-
erties was shown to depend on how homogeneously the two flow fronts unite [9]. Such
previous studies showed that discontinuous fibres tend to reorient parallel to the WLD,
decreasing the cohesion at the weld-region [10,11]. Mechanical strength can be reduced
significantly after the introduction of WLDs, by up to 90% [12–14]. Lower fibre content
has been reported, as well as reorientation of the fibres along the WLD, suggesting that
fibres from each side are not entangled at the weld-line, as they are in the rest of the part.
The direction of travel of the flow fronts also has an impact on the reduction of mechanical
properties: opposing flow fronts can have a bigger detriment than flow fronts running
adjacent and parallel to each other. It has, therefore, been demonstrated that the failure in
short fibre reinforced thermoplastics at a WLD is governed by the nature of the mecha-
nism that formed the interface [15].
Likewise, there has been limited research on the effect of WLDs on the performance
of DFCs. LeBlanc et al. manufactured flat plaques using chopped carbon fibre tow-based
composites in two configurations: “forced defect free”, and with a central WLD. It was
shown that a strength reduction of between 60% and 80% occurred for the plaques with
the WLD [16]. A thick compression moulded component (15 mm) was made by Martulli
et al., with weld surfaces formed at different locations when the charge pattern was
changed [17]. It was shown that the choice of the charge pattern configuration could play
a significant role in the final part performance, particularly via the formation of weld sur-
faces. A recent study investigated the effect of WLD on the morphology and mechanical
properties of the C-SMC [18]. It was shown that the different charge configuration resulted
in strength reductions of between 48% and 88% at the WLD.
Despite the severe performance reduction caused by the WLDs, none of the above-
mentioned studies applied truly non-destructive testing (NDT) methods, in an attempt to
detect them; other than expensive and highly specialised techniques, such as micro-CT.
At the high spatial resolutions required to characterise defects and attendant physical
sample size limitations, micro-CT is considered a destructive technique and would not be
applicable for in-line automated inspection at component level.
Simultaneously, accurate prediction of processing-induced microstructure using cur-
rent numerical simulation methods has several limitations, such as the inability to model
fibre–fibre or fibre–strand interactions, anisotropic viscosity, and simple one-way cou-
pling [19].
Feraboli et al. reported that the use of conventional NDT methods, such as ultrasonic
C-scan inspections, on DFCs proved to be difficult, due to the heterogeneity of the fibre
tow distribution and orientation [7]. Typical defects that corresponded to “hot spots” in
ultrasonic C-scans were voids (entrapped air at the resin-starved regions), resin rich areas,
and swirls (areas of high flow, leading to fibre kinks). However, they found that failure
was unlikely to occur in proximity of these hot spots, and there was no correlation be-
tween the location of the detected defects (using conventional ultrasound inspection) and
the location of the final failure. Therefore, using conventional inspection methods may
result in an overestimated rejection rate of parts [7].
One hitherto overlooked NDT method in the field of DFCs is eddy current (EC) test-
ing; it has, however, been used for quality analysis and for detecting major flaws such as
fibre misalignment and waviness in continuous carbon fibre reinforced polymers (CFRPs)
(unidirectional, woven, and non-crimped fabrics) [14,20–26]. EC characterisation of such
CFRPs relies on a high aspect ratio and the resultant anisotropic conductivity of carbon
fibre tows. Due to the relatively low conductance of carbon fibre (compared to most
J. Manuf. Mater. Process. 2022, 6, 151 3 of 25
metals), higher frequencies (in the ~MHz range) are used in EC testing of CFRPs. Due to
the skin depth effect, this higher frequency will result in a reduced excitation depth within
the sample, and the induced voltage, according to Faraday’s law, is proportional to rate
of change in the magnetic flux. Therefore, a higher frequency results in an increased volt-
age being measured by the receiver coil, giving an insight into the fibre tow orientation
and the stacking configuration. A schematic of this process is depicted in Figure 1a. Heuer
et al. [24] attributed the changes seen in the complex impedance signal to the three main
parameters that are shown schematically in Figure 1b. First, the average electrical conduc-
tivity of the material through-thickness is a reliable measure of carbon fibre volume frac-
tion, varying according to the amount of conductive carbon fibre (relative to the polymeric
resin) in the CFRP. Moreover, the bulk conductivity of the CFRP laminate depends on the
nature of the interface and connection between fibre tows, which defines the number of
current paths produced by overlapping fibre bundles.
Figure 1. (a) Schematic of probe configuration for EC testing; (b) electric and dielectric behaviour of
CFRP. Carbon fibre materials can be simplified as a network of resistors and capacitors, adapted
from [24].
In the original work by Schulze and Heuer, a multi-frequency eddy current device
was developed that was able to reveal defects underneath the surface of CFRPs, such as
missing carbon fibre bundles, suspensions, fibre angle errors, and waviness [26]. Heuer et
al. later used this technique to evaluate 3D parts using a radio frequency eddy current
sensor [24]. It was demonstrated that both the permittivity and conductivity of the carbon
fibre tows in CFRP determine the complex impedance, as measured by EC.
Whereas the EC testing of continuous carbon fibre fabrics (both dry and consolidated
within a laminate) is considered an established technique, usage of the technique to ex-
amine and measure fibre orientation in chopped tow-based DFCs (C-SMC) has not previ-
ously been investigated. It is expected that localised trends in fibre orientation and fibre
volume fraction within a moulded DFC will strongly influence the EC response, and par-
ticularly, a micro/meso-structural moulding defect (such as a WLD) might be detected
through the changing characteristics of the EC signal. This paper investigates a new NDT
approach, using automated EC scanning to examine compression moulded vinyl ester C-
SMC flat plaques, with and without WLDs. This resin system was selected to meet the
requirements for the C-SMC material for use in high volume manufacturing applications
(automotive), and as such the material system was formulated to deliver the best balance
of “mouldability”, cost, and performance.
To demonstrate the robustness of the EC scanning method, fibre volume fraction and
fibre analysis results were compared between various methods, including eddy current,
J. Manuf. Mater. Process. 2022, 6, 151 4 of 25
X-ray micro-computer tomography (micro-CT) scanning, radiograph, and process simu-
lation. The authors propose that the novelty of this work lies in not only using EC tech-
niques to detect WLDs in tow-based C-SMC, but also in measuring the main fibre orien-
tation in such materials, using a collaborative robot that can safely be placed within a
manufacturing environment.
2. Materials and Methods
2.1. Materials and Manufacturing
Commercially available Polynt 24 CF60-12K SMC material (25 mm-long chopped car-
bon fibre-12K/vinyl ester resin fibre, weight ratio of 60%) was used in this study. All of
the materials were moulded at WMG, using a flat plaque mould tool (550 × 550 mm) on a
V Duo 1700 tonne press (Engel, Austria). Moulding conditions used were a tool tempera-
ture of 140 °C, clamping force of 3000 kN, closure speed of 3 mm/s, and a 180 s cure cycle.
Three different moulding charge patterns were used to create varying WLDs, as per
the following (shown schematically in Figure 2):
Figure 2. Charge lay-up configuration and mechanical test specimen cutting location in a full flat
plaque for (a) mono, (b) butt, and (c) flow. Charge lay-up configuration and mechanical test speci-
men cutting location in a full flat plaque for (a) mono, (b) butt, and (c) flow.
Monolithic plaque, (Figure 2a): This was manufactured using a centrally-placed
standard (continuous) 60% coverage charge, selected to eliminate the occurrence of
WLDs. The charge size was 426 × 426 mm, consisting of 4 layers of SMC sheet, and it was
placed in the centre of the mould.
Butt jointed plaque, (Figure 2b): Based on the standard 60% coverage charge (426 ×
426 mm and 4 layers). Before moulding, a horizontal cut was made, to create a disconti-
nuity in the charge. The split charge halves were placed immediately adjacent to each
other without a visible gap (butt joint condition). This was designed to create a “low flow”
joining condition across the centre of the plaque.
Flow jointed plaque, (Figure 2c): Also based on the standard 60% coverage charge
(426 × 426 mm and 4 layers). Before moulding, a horizontal cut was made, to create a
discontinuity in the charge. The split charge halves were then placed at opposing sides of
the mould, leaving a 124 mm gap between them. This was designed to create a “high flow”
joining condition across the centre of the plaque, resulting in the formation of a WLD at
the meeting of linear flow fronts from each charge.
Hereafter, these three configurations are termed Mono, Butt, and Flow. In Figure 2,
the yellow dashed lines show the cutting configuration for tensile specimens across all
three configurations, and the pre-test EC measurement area is highlighted with a red
dashed line.
2.2. NDT Tests and Microstructural Analysis
J. Manuf. Mater. Process. 2022, 6, 151 5 of 25
2.2.1. Robot-Based EC Measurement
A high-frequency eddy current sensor from SURAGUS GmbH, guided by a collabo-
rative robot (Kuka-LBR-iiwa-14), was used for the EC measurements. Similar sensor tech-
nology has been used for quality assessment of carbon fibre materials [21,24,26,27]. This
system permits data acquisition of the phase shift and amplitude of the EC signal in a
complex plane, as a C-scan or a time plot, by holding the EC sensor precisely at a pre-
scribed distance during the entire scan. The scanning height is critical, as when measuring
low-conductive materials such as CFRP, any distance between the test coil and the speci-
men (lift-off) can cause an adverse disturbance to the received complex signal, and here it
was fixed at 0.1 mm.
The sensor was moved along equally spaced tracks over the surface of the moulded
plaques in a raster scanning pattern. It has been shown previously that the scanning speed
has no significant influence on the EC test results [28]; and in this study, a relative robot
speed of 25% was used to acquire high resolution EC scans. This speed setting is related
to the maximum rotational speed of the individual axes and the 25% velocity translates to
a linear scan speed of 30 mm/s. Identical EddyCus sensors are already employed for in-
line fibre areal weight measurements at much higher speeds, with a maximum of 300
mm/s [24,29]. A photograph of the system is shown in Figure 3. This system was inte-
grated and supplied by Expert Tooling and Automation Ltd., Allesley, UK.
Figure 3. Photographs of the Kuka-iiwa collaborative robot equipped with the EddyCus EC sensor.
The result of the rasterised scan operation is a 2D map of the sample’s electrical con-
ductivity relative to the orientation of the sensor head during the scan. By rotating the EC
sensor and performing repeated scan operations, it is possible to combine the electrical
conductivity maps, to calculate variations in uniformity and main fibre orientation. There-
fore, four repeated scan operations were performed for each region, and for each scan, the
sensor orientation was changed from −45° to 90°, in 45° increments. Table 1 summarises
the scanning parameters for the experiments, including the errors in measurements.
J. Manuf. Mater. Process. 2022, 6, 151 6 of 25
Table 1. Eddy current scanning parameters for the experiments.
Parameter Value
Robot scanning speed 30 mm/s
Robot positioning accuracy (ISO 9283) ±0.1 mm
Sensor type SURAGUS 48
-Coil diameter
-Shape of coil
-Coil configuration
3 mm
Helical
Half-transmission
Sample rate 100 Hz
Eddy current frequency 6 MHz
Average Distance between data points 0.25 mm
In this study, all of the raw or calculated images are plotted using the magnitude of
complex impedance signal defined as: |Z|= √(Z_Re)2 + (Z_Im)2); where Z_Re is the real
and Z_Im is the imaginary part of the complex impedance signal. The magnitude is inde-
pendent of the phase change, and therefore no signal manipulation or phase rotation (as
mentioned in [27]) was required to enhance the features of the C-SMC EC time plot im-
ages. After scanning, the imaginary and real arrays were used to plot a 2D image of the
conductivity using a Python script.
Total impedance strength, also called uniformity, directly relates to the conductivity
of a region, and hence is proportional to the (in-plane) carbon fibre volume fraction. Due
to the inhomogeneous nature of the C-SMC, it is not practical to convert the magnitude of
a complex impedance signal precisely into a fibre areal weight. As discussed, uniformity
(u) (directly related to fibre areal weight) can be calculated using the magnitude (m) of the
complex impedance signal measure in each sensor orientation (n = different sensor direc-
tions (−45°, 0°, 45°, 90°)):
𝑢(𝑥,𝑦)= (𝑚 (𝑥,𝑦));
(1)
where fraction is defined as f, magnitude as m, uniformity as u, and fibre angle as fa. The
ratio of material orientated towards the sensor direction (for example 0°) is defined as:
𝑓
°(𝑥,𝑦)= 𝑚° (𝑥,𝑦)
∑𝑚 (𝑥,𝑦)
(2)
The direction of the prominent fibre angle at a particular region is given by:
𝑓
𝑎(𝑥,𝑦)=arctan(∑(𝑐𝑜𝑠2∗𝑛).
𝑓
(𝑥,𝑦)
∑(𝑠𝑖𝑛2∗𝑛).
𝑓
(𝑥,𝑦)
) (3)
SURAGUS GmbH uses this above approach to determine fibre areal weight meas-
urement and to characterise anisotropy in thin films, and the formulation was hereby ap-
plied to characterise fibre angle in C-SMC [29]. The results of anisotropy measured from
sheet resistance for each rotation angle of −45°, 0°, 45°, and 90° are shown in the supple-
mentary dataset.
In an attempt to understand the variations arising in the EC uniformity data, 2D ra-
diography and micro-CT was carried out for the regions, as indicated in Figure 4. These
areas showed the most complex signal attenuation, corresponding to where the material
experienced the most complex flow conditions.
J. Manuf. Mater. Process. 2022, 6, 151 7 of 25
Figure 4. (a,b) DIC from tensile testing (dashed orange), eddy current (dashed blue), µCT, and ra-
diography (green box) area are shown for the Butt and Flow plaques. The green box highlights the
location of µCT and radiography. The purple box highlights the RoI from which all NDT and DIC
images are compared; (c) DIC area is shown with a dashed orange line and the EC measurement
with a dashed blue line on the actual three different plaques.
2.2.2. Micro CT and Radiography
The samples were micro-CT scanned using a Tescan Unitom XL X-ray Computed
Tomography (CT) scanner at CiMAT, WMG, University of Warwick, UK. The sample is
penetrated by X-rays from a cone-beam X-ray source, which are digitised by a detector
into 16-bit grey value pixels, representing the X-ray absorption of the sample. Many radi-
ograph projections are taken as the sample rotates through 360 degrees, thereby allowing
reconstruction of digital 3D volumes. Alternatively, a single projection (radiograph) can
be taken, where grey values are averaged through the thickness of the sample. When a
sample is too large to achieve the desired scan resolution and field of view, as was the case
in this work, multiple sub-scans and/or radiographs can be taken before stitching the im-
ages together. Five radiographs and three micro-CT scans of the Flow and Butt samples
were conducted; X-ray parameters and sample sizes are detailed in Table 2. The focal spot
size of this micro-focused X-ray source was 15 µm. Individual radiographs were stitched
together using ImageJ, with contrast and brightness adjusted to maximise the grey scale
range. The micro-CT scans were stitched together and reconstructed using Tescan’s pro-
prietary Acquila reconstruction software. A first order beam hardening correction was
applied during reconstruction, to normalise the grey value gradient across the width of
the samples (BHC = 0.1). The reconstructed micro-CT scans were imported into VGStudio
Max 2.2 (Volume Graphics GmbH, Heidelberg, Germany), a commercially available soft-
ware for the analysis of CT and voxel data. The resin and fibre structures were segmented
using a fixed threshold and the sample was aligned using a best-fit plane on the sample
front face. VGStudio’s Fibre Analysis Module was used to determine the fibre orientation,
tensor angle, and relative fibre volume fraction (the latter two were averaged through the
thickness of the samples and within a 2 × 2mm mesh). This fibre module has been demon-
strated as an effective tool for analysing composite structures from micro-CT images [19].
J. Manuf. Mater. Process. 2022, 6, 151 8 of 25
Table 2. Tescan UnitomXL micro-CT parameters used to scan and radiograph the flow and butt flat
plaques pre and post testing.
Tescan Unitom XL
Parameters
Radiographs 3D Micro-CT Scans
Flow Butt 50 mm Flow 50 mm Butt 25 mm Flow
and Butt 50 mm Flow 50 mm Butt 25 mm Flow
and Butt
Sample size (mm) 250 × 250 250 × 250 50 × 250 50 × 250 2 × (25 × 250) 50 × 250 50 × 250 2 × (25 × 250)
Field of view (mm) 297 × 262 297 × 262 49 × 49 49 × 49 54 × 54 13 × 187 13 × 192 74 × 74
Image stacking 2 × 2 2 × 2 1 × 6 1 × 6 1 × 1 1 × 5 1 × 5 1 × 1
No. sub-scans 4 4 6 6 1 5 5 1
Detector size (pixels) 2856 × 2856 2856 × 2856 2856 × 2856 2856 × 2856 2856 × 2856 1920 × 1896 1920 × 1896 2856 × 2856
Image size (pixels) 5034 × 5712 5034 × 5712 2856 × 2856 2856 × 2856 2856 × 2856 1920 × 7181 1920 × 7112 2856 × 2856
Resolution (µm) 52 52 17 18 19 26 27 26
Voltage (kV) 40 40 40 40 40 40 40 40
Power (W) 15 15 15 15 15 30 30 30
Grey range (×1000) 29–52 33–52 29–52 33–52 29–52 10–52 10–52 10–52
2.3. Process Simulation
Compression moulding process simulation was performed using 3D TIMON Com-
posites PRESS developed by Toray Engineering D Solutions. 3D TIMON is a commercially
available flow simulation software with an in-built Direct Fibre Simulation (DFS) solver,
which models the movement of individual fibres in the bulk flow. The closure profile was
modelled on the speed versus time data recorded by the press during the compression
moulding process. The Euler method (fixed mesh) was selected in this study, and the Eu-
ler domain was meshed using an in-plane seed size of 2 mm and an out-of-plane seed size
of 0.5 mm for all three cases. For the DFS analysis, a 2D random fibre architecture was
generated within the volume of the initial charge using a density of 2 fibres per unit vol-
ume, and each fibre was divided into 6 linked beam elements.
2.4. Mechanical Testing with 3D DIC
Quasi-static mechanical testing was performed using an Instron test machine with a
250 kN load cell. Tensile testing was performed according to ASTM D3039/3039M [1].
Specimens were cut using a Compcut 200 composite cutting machine to a standard sample
size of 250 × 25 mm and tested at 2 mm/min with a starting grip distance of 140 mm. This
cutter was designed for cutting composites, hence it results in a smooth surface, as it uses
a diamond abrasive wheel.
During the tensile tests, the surface strain distribution was obtained using a three-
dimensional digital image correlation (3D DIC) system. DIC has been successfully used in
previous studies of discontinuous C-SMC [4,30]. Here, a GOM 12M system with GOM
ARAMIS software was used to capture and process stereoscopic DIC images at 1 Hz, and
the system enabled corresponding data acquisition. Four static frames were recorded and
analysed prior to commencing testing for each sample, to check the effectiveness of the
applied stochastic pattern and to determine the noise level. The parameters of the meas-
urements are listed in Table 3. Mean bias and precision for 3D DIC system calculated from
static frames are shown in Table 4.
J. Manuf. Mater. Process. 2022, 6, 151 9 of 25
Table 3. DIC setup parameters.
Sensor GOM 12 M with Titanar 100 mm Lens
Image window 4000 pixel × 3000 pixel
Measurement area 150 mm × 100 mm
Calibration plate used CP20 90 × 72
Facet size 19 pixels (16 pixels step size)
Depth of field 39 mm
Frame rate 1 Hz
Table 4. Mean bias and precision for 3D DIC system calculated from static frames.
εx (%) dx (mm) dy (mm)
Mean
Bias Mean Precision Mean Bias Mean Precision Mean Bias Mean Precision
−0.0005 0.0195 0.0034 0.0005 −0.0002 0.0006
3. Results
3.1. Effect of Different Flow Patterns on Microstructure
To better understand the effect of flow on the resultant microstructure, for the Flow
and Butt samples, a 25 mm-wide tensile specimen was EC scanned and then sectioned
into three pieces across the WLD region, and through-thickness micrographs were pre-
pared using a Zeiss optical microscope. Corresponding EC real impedance maps (in-
plane, top image) and optical micrographs (out-of-plane/through thickness, lower three
images) are shown in Figure 5. The location of each optical micrograph section is shown
on the EC maps with a coloured dotted line. It can be seen that the weld line defect extends
between 5 and 10 mm, with visible cracks resulting from both the in- and out-of- plane
movement of the tows.
The high-flow mechanism orients the fibres parallel to the weld line and, as a result,
the reinforcement across the weld line is very poor. In this case, CF tows are no longer
oriented in-plane, and the typical in-plane microstructure is locally disrupted through the
thickness. The flow fronts seem to be crushed and distorted, also resulting in the for-
mation of cracks. Similar results were shown in works by Martulli et al. [17] and Evans et
al. [31]. It can be seen that the WLD region is not discrete and varies between 10 and 20
mm in width. In this case, the opposing flow front acts as a barrier that results in out-of-
plane flow and alignment of the fibre tows.
Figure 5. mm-wide in-plane EC (magnitude) image (a) Butt tensile coupon (top); (b) Flow tensile
coupon (top) and three successive through-thickness (out-of-plane) optical micrographs from the
same specimen region. The coloured dashed lines show the location of the cross-section depicted in
the same colour dashed box.
J. Manuf. Mater. Process. 2022, 6, 151 10 of 25
Unlike the Flow plaque (which has the most of out-of-plane distortion), the micro-
graph of the Butt sample shows a more localised defect, and the typical SMC in-plane
layered structure is restored within 10 mm of the WLD. However, the magnitude of dis-
turbance in the EC uniformity signal seems to be much higher for this type of disturbance
in the fibre architecture. It was therefore demonstrated that the action of simply splitting
the charge by mechanical cutting, and subsequently compression moulding this split
charge whilst immediately adjacent, had a significant effect on the in-plane EC uniformity.
The in-plane electrical conductivity remained disturbed by the original cutting of the fibre
tows.
3.2. Mechanical Properties and Comparing DIC/EC
Fourteen specimens each were tested from the Mono, Butt, and Flow plaques. All
specimens with forced defects (from the Butt and Flow plaques) failed at the WLD, with
no exceptions. Specimens from the Mono plaque all failed at different locations within the
gauge length, as might be expected. Figure 6 shows the mean strength for each of the
materials, along with the reduction in strength arising from the WLD in the Butt and Flow
versus the Mono plaques. The nominal strength given by the manufacturer’s datasheet of
the monolithic material at full coverage is 130 MPa. Mean tensile strengths are displayed,
and the relative strength reduction is shown for the samples with induced WLDs in Figure
6.
Figure 6. Tensile strength results of the three different charge patterns. Strength reductions shown
for the plaques with WLDs are compared to that of the Mono plaque (WLD-free).
Figure 7 presents the EC uniformity maps for the three types of specimen (a1 (Mono);
b1 (Butt); c1 (Flow)), the DIC surface strain maps obtained prior to ultimate failure during
tensile testing (a2-b2-c2), and the EC principal fibre angle data (a3-b3-c3). This side-by-
side comparison of the EC uniformity data/DIC data from the same regions for each ma-
terial type shows that there was some correlation between the morphology of the EC uni-
formity map and the local strain distributions present at the surface of the sample at fail-
ure. Local strain variations were largely associated with the non-homogeneous nature of
the C-SMC, not only at the surface, but also within the underlying micro/meso-structure
[30]. This is also exactly how the EC signature is governed, as the depth penetration of EC
was expected to be greater than the thickness of the samples used in this study; this ex-
plains the close correlation between these EC uniformity maps and the DIC strain maps.
The “hot spots” in the strain map indicate an area of the lowest modulus for the specimen,
and for the Butt/Flow materials these lower modulus regions correspond exactly to the
lower EC magnitude signal, as expected from resin-rich regions/locally reduced in-plane
fibre orientation.
J. Manuf. Mater. Process. 2022, 6, 151 11 of 25
Figure 7. Comparing EC mapping (before testing) with the DIC/mechanical strength of a (a) contin-
uous (Mono) plaque; (b) Butt (low-flow) WLD plaque, (c) Flow (high-flow) WLD plaque. (a,b,c-1)
EC uniformity; (a,b,c-2) major strain (Ɛy), fixed scale 1.5%; (a,b,c-3) main EC fibre angle with overlaid
uniformity colours.
In particular, it can be seen from Figure 7(a-1–a-3) that the Mono plaque does not
show the same linear discontinuity in the strain field and uniformity map as observed in
the Butt and Flow plaques; there is no WLD present in the middle of the plaque, nonethe-
less several strain concentrations can be detected in the DIC strain maps. Considering that
the original chopped tows in the SMC were of 25 mm length, large modulus variations
are expected in a tensile specimen of 25 mm width. For a 25 mm fibre length SMC, the
critical RVE size is ~100 mm. The macro-scale modulus with such a small coupon size
cannot be correctly captured. This gives rise to greater macroscopic heterogeneities and
measurable discontinuities in the modulus [30]. The EC uniformity map from the Mono
plaque confirms such a macroscopic variation in the laminate. The low intensity (blue
colour) regions are ~50 mm. It is, therefore, proposed that the distribution of chips at the
mesoscopic level affects the EC signal measurement at a macroscopic scale. The local fibre
orientations calculated from the four-orientation EC scans of the Butt and Flow plaques
(Figure 7(b-3,c-3)) indicate a high degree of transverse orientation of the fibres at the WLD
(transverse to applied uni-axial load, parallel to WLD), as also reported in previous stud-
ies [15,18]. Figure 8 focuses on the results from three representative tensile specimens
taken from the same location in each plaque, showing the EC local fibre angle data
J. Manuf. Mater. Process. 2022, 6, 151 12 of 25
(superposed with the EC uniformity map) alongside DIC strain maps for each specimen
at the maximum applied load before failure (σmax). The dimensions of area shown are 25
mm × 140 mm, and the failure locations are highlighted with a dashed box. An EC map
data was taken prior to cutting and testing. Note that the correlation of locations is only
approximate, due to the elimination of material during tensile specimen extraction and as
the DIC strain maps were obtained from deformed specimens (at σMax). The DIC ƐY (axial
direction) strain maps were scaled automatically (relative scaling), with the ƐY (Local Max)
given for each map shown (corresponding to red). Figure 8g shows the distribution of
surface strains across all calculated DIC elements for each specimen, at a fixed maximum
of 1.5% to facilitate comparison. More data for the rest of the specimens can be found in
the supplementary data.
Figure 8. Comparing tensile coupons from the Mono plaque (M2, (a,b)), Butt (B2, (c,d)), Flow (F2,
(e,f)). (a,c,e) DIC measurement at σMax; (b,d,f) main EC fibre angles; (g) Comparing specimen strain
distributions at σMax. (For interpretation of the references to colour in this figure legend, the reader
is referred to the web version of this article). The width of each test specimen and the image is 2.5
cm.
Again, there is close agreement between the patterns seen in the EC uniformity and
the strain map for each specimen. One common feature among all ultimate failure loca-
tions is the relatively lower EC magnitude, corresponding to an increased overall resin
content/lower fibre volume fraction and broken fibres. Furthermore, the calculated prin-
cipal fibre angles (EC) indicate that a transverse fibre alignment (to the applied load) is
J. Manuf. Mater. Process. 2022, 6, 151 13 of 25
dominant in the high local strain areas (DIC) and in many lower EC magnitude regions.
Where these two factors combine, it is more likely that the specimen will fail, seen most
clearly in the Flow example (Figure 8f(F2)). To understand if EC can provide insights
about the failure location in WLD-free samples, four individual specimens are shown in
Figure 9.
Taking the case of the M1, M2, and M4 specimens (Figure 9), where the failure oc-
curred at ~45° angle in the various sections (seen clearly from the DIC strain map), it was
also seen that the uniformity signal is low for this 45° section and the principal fibre ori-
entation is also at 45°. Whilst the EC uniformity maps reveal some information about po-
tentially weaker regions, to better predict the failure location, the local fibre orientations
should also be considered. Therefore, for the defect-free (Mono) material, it is shown that
EC scanning can reveal candidate locations for failure. This contrasts with what has been
reported for other NDT methods (such as ultrasound C-scans), where the detected “hot
spots” in scans prior to testing did not correlate with the failure location [7].
Figure 9. EC uniformity/principal fibre angle measurements along with the DIC major strain maps
for tensile coupons of 2.5 cm width from the Mono sample. The maximum strain in DIC maps was
set to 1.5% to enable data comparison.
In previous studies, it was shown that strong transverse orientations parallel to the
weld line can exist in high flow C-SMC parts [18]. As already discussed, this parallel ori-
entation and a lack of intermingling leads to the observed high local surface strains and
results in a highly visible disruption of EC uniformity. Furthermore, the blue resin-rich
“branch-like” features seen in the EC uniformity maps for the Flow material are attributed
to this local transverse orientation and poor intermingling of the CF tows. When compar-
ing the induced WLD plaques, a clear pattern is revealed from EC and DIC, whereby the
weakest part of the sample is the WLD, and this can be clearly detected by EC prior to
destructive mechanical testing. The weld-line region at the centre of each specimen shows
a lower EC uniformity signal and a transverse fibre orientation at the weld-line (parallel
to WLD, transverse to applied stress). The combination of these two factors will certainly
reduce the mechanical performance and hence the elongation to failure and strength. Ta-
ble 5 summarises the mechanical properties of the three specimen types (M2, B2, and F2),
as shown in Figure 8.
J. Manuf. Mater. Process. 2022, 6, 151 14 of 25
Table 5. Mechanical properties of three specimen M2, B2, and F2 shown in Figure 8. * Please note
that the modulus of the WL specimens was obtained from the “virtual” extensometer applied in the
central gauge region in the DIC analyses.
Configuration Tensile Strength (MPa)
Apparent
Modulus
(GPa)
Max Global εy (%) Max Local εy
(%)
Mono (M2) 90.4 ± 0.1 16.5 ± 0.1 0.569 ± 0.001 6.208 ± 0.001
Butt (B2) 40.5 ± 0.1 * 17.2 ± 0.1 0.242 ± 0.001 3.425 ± 0.001
Flow (F2) 12.3 ± 0.1 * 16.3 ± 0.1 0.073 ± 0.001 1.732 ± 0.001
It can be seen that the induced WLD did not decrease the global modulus (130 mm
gauge length, calculated from DIC and across the WLD). However, the maximum tensile
strength was dramatically reduced for these chosen samples for B2 and F2, by 56% and
86%, respectively. This decrease is also reflected in the changes in the DIC surface strain
distributions for all specimens at maximum stress before failure (σMax), as shown in Figure
8g. For the samples with WLDs, not only is the centre of the peak of the distribution shifted
to lower strains, but also the full width-half maximum of the peak has narrowed. This
indicates a more discrete localised defect in the WLD specimens than for the Mono. For
the Mono sample, the peak in the distribution is at a higher value of major strain than for
the WLD-induced plaques (reflecting the higher σMax and also greater elongation to fail-
ure) and is also much broader. It is, therefore, harder to predict the location of failure in
the Mono samples prior to failure; multiple localised strain hot-spots continue to develop
across the sample during loading, up to the final failure.
3.3. Microstructure Analysis Using Micro-CT
It can be seen from Figures 10 and 11 that irregularities in the fibre tows exist, both
in-plane and also in the through-the-thickness direction near the WLD for both Butt and
Flow moulding configurations. Given the localised nature of the WLD irregularities, it is
expected that the microstructure influences the local conductivity and reduces the EC im-
pedance complex magnitude signal near the WLD. In Figure 10a, it can be seen that a tow
of carbon fibre has an out-of-plane orientation and a crack has propagated along this tow,
avoiding the nearby void in the tows. The same explanation can be used to understand
the similarities seen between the high strain pattern and the low conductivity detected at
the WLD. The typical mesoscopic nature of the C-SMC is no longer recognisable within
the WLD region of the Flow sample. The cross-section shows that the typical defects such
as macroscopic voids, swirling, and fibre kinking (and resin rich areas) are common de-
fects at these WLDs. It appears that the failure in tow-based SMC is insensitive not only
to macroscale notches [4,32], but also in smaller voids between tows near the WLD. The
failure seems to run around the tow ends at the WLD and circumvents the intra tow cracks
and voids, suggesting that the failure is matrix dominated. If there is a void in the weld-
line, the failure crack goes through it; otherwise the voids in the bundles are unaffected
and the failure is of a fibre–matrix debonding nature.
J. Manuf. Mater. Process. 2022, 6, 151 15 of 25
Figure 10. Flow tensile coupon micro-CT, in plane (upper) out-of-plane (lower), (a) before testing;
(b) after testing; (c,d) larger sections of both scans for (c) in-plane and (d) out-of-plane, also before
and after failure.
The micro-CT in-plane sections for the Flow sample (Figure 10a,b upper images)
clearly show tows aligned in the direction of the WLD; in this region, it is not necessarily
a horizontal line. However, the out-of-plane cross-sections (Figure 10c,d) show evidence
of turbulent flow where the opposing flow fronts met. In contrast, similar out-of-plane
sections for the Butt coupon (Figure 11c,d) show less out of plane movement of the fibre
tows away from the WLD, as was discussed earlier for the optical micrograph. Therefore,
the cross-section from the micro-CT shows similar phenomena to those seen in the optical
micrograph, confirming that the out of plane flow of fibre tows are contributing to the loss
of EC signal, leading to the lower intensity in the EC uniformity maps.
Figure 11. Butt tensile coupon micro-CT, in plane (upper) out-of-plane (lower); (a) before testing,
(b) after testing; (c,d) larger sections of both scans for (c) in-plane and (d) out-of-plane, also before
and after failure.
From the microscopy, micro-CT, and reduced EC signal (uniformity magnitude) at
the WLDs, it can be concluded that the two opposing flow fronts act as an impermeable
barrier, similarly to the walls of the moulding cavity, as observed by Evans et al. [31]. The
out-of-plane tow distortions caused by the meeting of the opposing flow fronts are as pre-
viously shown using micro-CT [18] (characteristic bending or crushing). The discontinui-
ties observed from the EC uniformity maps in this study at a WLD (Butt or Flow), arising
from the lack of conductivity across the unbridged resin gap, confirm that there is scant
intermingling happening between fibres from opposing flow fronts. In particular, in the
case of the induced cut in the Butt plaque, the broken electrical circuit was not “healed”
during subsequent compression moulding of the SMC. Therefore, it can be concluded that
the detrimental effects of introducing a highly localised low-flow WLD in C-SMC by
J. Manuf. Mater. Process. 2022, 6, 151 16 of 25
placing immediately adjacent split charges remain following a high pressure co-moulding
process.
In Figure 12, various features of the weld-line are examined in further detail using
micro-CT. It is clear that the 2D-planar nature of the tow-based SMC was disturbed by the
introduction of weld-lines. The perturbations away from an in-plane orientation cannot
have been caused solely by the flow, as the orientation distribution observed by previous
work for tow-based SMC showed that, for thin flat plaques, the 2D planar nature of the
SMC sheets can be preserved, even in high flow conditions.
The micro-CT images of the strips from the Flow plaque shows that the tows are
dramatically re-aligned away from the in-plane direction, transverse to the weld-line (Fig-
ure 12). None of the characteristic rectangular-shaped “chips”, as present in unprocessed
C-SMC, are recognisable in this sample. Figure 12c shows that some undistorted “square”
ends of cut tows can be found away from the weld-line. In contrast, the carbon fibre tows
in the weld-line region have been distorted in-plane and lost all shape. It is shown below
that the mechanical properties of the Flow material were lower than those of the Butt. The
wrinkled/distorted tows can be seen at the flow front, explaining the significant reduction
in the mechanical strength for these specimens. In the Flow sample, away from the weld-
line, it appears that the material kept its mesoscopic structure. However, the high local
distortion at the weld-line affected the global mechanical strength. The micro-CT images
in Figure 12 were taken from the mid-section of the specimen. When examining the im-
ages in the sequence “through sample thickness”, it can be observed that higher distortion
exists for the mid-plane section when compared to the outer layer in both specimens. This
implies a skin–core effect resulting from a plug–flow behaviour. Darker resin rich areas
and voids are seen to be more prevalent in the swirls of the distorted tows (both out-of-
plane and in-plane), causing fibre waviness in all three principal axes.
Figure 12. Micro-CT images of (a) Flow sample strip, showing two regions: (b) weld-line showing
high distortion of the tow shapes, (c) away from the weld-line the tows with rectangular ends are
encircled in c. The slice through is at x = 20 mm. ((d) Butt sample strip, showing two regions: (e)
weld-line region where some tows with originals rectangular shape can be seen, (f) away from the
weld-line high distortion is observed. The slice through is at x = 20 mm, showing the planar behav-
iour away from the weld-line.
J. Manuf. Mater. Process. 2022, 6, 151 17 of 25
In Figure 13, out-of-plane CT micrographs from the B1 Butt tensile specimen for two
similar sections are shown, before and after testing for the same regions. It is clear that the
failure has run through-thickness in the resin-rich region where the fibre tows met as op-
posing flow fronts at the WLD. This discontinuity within in-plane fibre architecture ex-
plains the measured drop in the EC signal as the electrical circuit of the bundle tows are
broken because of the non-conductive resin rich areas [24]. Minor defects such as a small
round void are seen in Figure 13 and these could be as a result of air trapped between the
flow fronts. Interestingly, the failure crack has propagated very close but not through the
void. Similar behaviour can be seen in (b) when a crack/void in the fibre bundle is high-
lighted. It is therefore seen that the 2D-planar architecture of the SMC is interrupted in
the WLD region of the Butt sample. The voids are more commonly seen in the resin rich
areas, whilst cracks observed in the tow bundle suggest that two flow fronts do not inter-
mingle and instead act as opposing impermeable barriers to flow. Tows become distorted
as a result of this opposition. Nonetheless, away from the WLD, the typical ‘in-plane lay-
ered’ SMC structure is still visible. Comparison of the results from the process simulation
with the EC uniformity scan of larger regions are presented in Figure 14.
Figure 13. CT Micrograph slices of a Butt tensile specimen before and after failure. (a) A small cir-
cular void is highlighted in the resin rich area between the tows at the WLD. Even though that the
crack proceeds through the resin rich area, it misses the void. (b) Another section of the same spec-
imen, before and after testing. A larger void in the fibre tow is encircled, which remained isolated
after failure, indicating that a matrix dominated failure occurred.
Figure 14. Comparing simulation with 3D-Timon with EC fibre angle and fibre volume fraction:
(right) Flow (left) Butt.
J. Manuf. Mater. Process. 2022, 6, 151 18 of 25
4. Discussion
The NDT methods of advanced radiography and micro-CT, as well as DIC, were
compared with EC scans for the Butt and Flow plaques for outer regions in both, where it
was certain that material had flowed i.e., outside of the original charge location, as shown
in Figure 4. Fibre angle analysis was performed using micro-CT, to evaluate the accuracy
of the EC principal fibre angle measurements. Furthermore, the compression moulding
process simulation results for these plaques are shown for the region of interest (to high-
light the efficacy of EC scanning in detecting WLDs in C-SMC, which is not possible oth-
erwise. The results for the Butt and Flow samples from the specified region of interest are
compared in Figures 15 and 16.
J. Manuf. Mater. Process. 2022, 6, 151 19 of 25
Figure 15. Comparing fibre angle and fibre volume fraction results for the Flow specimen, (a) micro
CT image (at z = 1.48 mm), (b) micro-CT relative FVF, (c) DIC longitudinal strain map of two adja-
cent tensile coupons (max strain at 1%), (d) eddy current uniformity (fibre areal weight), (e) simu-
lated fibre areal weight, (f) radiograph, (g) fibre tensors (micro-CT), (h) micro-CT fibre orientation
(at z = 1.48 mm), (i) eddy current principal fibre angle, and (j) simulated fibre orientation. The WLD
areas are highlighted with a red dashed box. Dimensions of area shown are 39 mm × 135 mm.
As before, there is some correlation between the DIC strain measurement (c) and EC
uniformity maps (d), as the microstructure and local fibre volume fraction (FVF) directly
influenced the stress–strain distribution in C-SMC. The EC sensor picks up the superpo-
sition of magnetic field contributions through the entire thickness, coming from the flow
of current within the C-SMC. Each chopped bundle/layer will contribute differently to the
overall magnetic field, and thus the impedance measured by the EC coils. Cheng et al.
predicted, using finite element analysis, that the current density is at a local maximum
when fibre tows cross [23]. The WLDs are seen to provide the opposite effect to that wit-
nessed at interfaces between cross-plies, due to the demonstrated lack of intermingling
where the two flow fronts meet. Therefore, a local minimum of current density can be
expected for all layers of the C-SMC. A measurable difference in the complex contribution
is shown when looking at the total impedance strength scans from the combined four-
direction EC scans (d). This is further emphasised when looking at the microstructure in
the micro-CT data (a). To replicate the fibre angle measured by EC using micro-CT, a vol-
ume of 2 × 2 × 2.5 mm was chosen to calculate the fibre tensors shown in (g) and also to
measure the relative FVF through the thickness (b).
J. Manuf. Mater. Process. 2022, 6, 151 20 of 25
Figure 16. Comparing the fibre angle and fibre volume fraction results for the Butt specimen, (a)
micro CT image (at z = 1.45 mm), (b) micro-CT relative FVF, (c) DIC longitudinal strain map of two
adjacent tensile coupons (max strain at 0.5%), (d) eddy current uniformity (fibre areal weight), (e)
simulated fibre areal weight, (f) radiograph, (g) fibre tensors (micro-CT), (h) micro-CT fibre orien-
tation (at z = 1.45 mm), (i) eddy current principal fibre angle, and (j) simulated fibre orientation. The
WLD area is highlighted with a red dashed box. Dimensions of area shown are 39 mm × 135 mm.
However, this is not the only factor that leads to such a reduction in the ultimate
strength. Away from the weld-line (also in “forced defect-free” materials), a complex EC
uniformity map was obtained that correlates with the radiographs (f) of the same part
showing a lower intensity. It was shown earlier, in Figures 7 and 9, that the EC uniformity
J. Manuf. Mater. Process. 2022, 6, 151 21 of 25
map in “forced defect free” material has an associated strain distribution map with corre-
sponding minima and maxima throughout the sample, showing that the observed varia-
tion in local modulus is “material inherent” to DFC. It was previously shown that these
materials are virtually notch insensitive at the macro-scale [4,32].
It can be seen from Figures 15 and 16 that the weld-line region shows a lower in-
plane/FVF carbon content (lower uniformity EC signal), thus agreeing with the simulation
results (e), micro-CT FVF (b) and the radiographs (f). However, the morphology of varia-
tions in microstructure is not in agreement when comparing EC uniformity (d) and pre-
dicted FVF from simulation (e) in Figure 12. It demonstrates that the simulation cannot
replicate the complex flow behaviour at the meeting flow fronts and hence cannot predict
the resultant microstructure.
In Figure 15, in the region adjacent to the WLD, a 45° principal fibre angle is predicted
by the simulation (j) and this is in agreement with both micro-CT tensor (g), fibre orienta-
tion data (h) and EC principal fibre angle (i). However, the process simulation predicted
higher level of fibre alignment along the WLD, due to the inappropriate constraints im-
posed to the fibres at the flow front. The imprint of the WLD can be seen in radiographs
(Figure 15f) and the micro-CT images (Figure 15a). These correlate with the DIC strain
maps (Figure 15c) highlighting that the failure occurred at the weld-line where the EC
uniformity showed the weakest signal (prior to testing). Some correlation between the
relative FVF measured by micro-CT (b) and the EC uniformity signal (d) can be seen in
Figure 16, indicating unexpected WLD formations (at 45°) in this region near the edge of
the mould. Process simulation fails to replicate such process-induced features possibly
due to the flow front effects caused by the DFS model. The fibre angle measurements ob-
served in Figure 13 again reveal that the fibres are oriented parallel to the weld plane
(Figure 16g) and (Figure 16i). Therefore, using the EC principal fibre angle data and EC
uniformity map, it is proposed that failure location could readily be predicted.
Tables 6 and 7 show the mean orientation tensors measured using Micro-CT, Simu-
lation and Eddy Current for Flow and Butt specimen of scanned areas marked as 1, 2 and
3. The area that these tensors were measured and the coordinate system are shown in
Figures 15 and 16 for Flow and Butt samples respectively. The red dashed area corre-
sponds to Area 2 (weld-line region) in both specimens.
Table 6. Mean orientation tensor measured for the Flow specimen of scanned area, marked as 1, 2,
and 3, using Micro-CT, simulation, and eddy current. The failure location is highlighted in red.
Flow Joint Micro-CT Simulation EC (2D Meas-
urement)
Area 1: 40 mm above
the WL 0.55 0 0
𝑠𝑦𝑚. 0.44 0
𝑠𝑦𝑚. 𝑠𝑦𝑚. 0.01 0.46 −0.09 0
𝑠𝑦𝑚. 0.54 0
𝑠𝑦𝑚. 𝑠𝑦𝑚. 0 0.59 .
.0.41
Area 2: ±20 mm
around the WL 0.64 0.03 0
𝑠𝑦𝑚. 0.33 0
𝑠𝑦𝑚. 𝑠𝑦𝑚. 0.03 0.61 −0.03 0
𝑠𝑦𝑚. 0.39 0
𝑠𝑦𝑚. 𝑠𝑦𝑚. 0 0.67 .
.0.33
Area 3: 40 mm below
the WL 0.52 0.03 0
𝑠𝑦𝑚. 0.47 0
𝑠𝑦𝑚. 𝑠𝑦𝑚. 0.01 0.44 0.08 0
𝑠𝑦𝑚. 0.56 0
𝑠𝑦𝑚. 𝑠𝑦𝑚. 0 0.38 .
.0.61
Table 7. Mean orientation tensor measured for the Butt specimen of scanned area marked as 1, 2,
and 3 using Micro-CT, simulation, and eddy current. The failure location is highlighted in red. Scan
locations and coordinate system are shown in Figures 15 and 16. The red dashed area corresponds
to Area 2.
Adjacent Joint Micro-CT Simulation
EC 2D Measurement
(Mean Vector
Vx = sin2α, Vy = cos2α)
J. Manuf. Mater. Process. 2022, 6, 151 22 of 25
Area 1: 40 mm
above the WL 0.51 0.01 0
𝑠𝑦𝑚. 0.49 0
𝑠𝑦𝑚. 𝑠𝑦𝑚. 0.01 0.46 0.02 0
𝑠𝑦𝑚. 0.54 0
𝑠𝑦𝑚. 𝑠𝑦𝑚. 0 0.60 .
.0.40
Area 2: ±20 mm
around the WL 0.54 0.03 0
𝑠𝑦𝑚. 0.46 0
𝑠𝑦𝑚. 𝑠𝑦𝑚. 0.01 0.51 0.01 0
𝑠𝑦𝑚. 0.49 0
𝑠𝑦𝑚. 𝑠𝑦𝑚. 0 0.64 .
.0.37
Area 3: 40 mm be-
low the WL 0.52 0.04 0
𝑠𝑦𝑚. 0.47 0
𝑠𝑦𝑚. 𝑠𝑦𝑚. 0.01 0.47 −0.03 0
𝑠𝑦𝑚. 0.53 0
𝑠𝑦𝑚. 𝑠𝑦𝑚. 0 0.62 .
.0.39
The measured out-of-plane orientation is higher in area 2 of both specimens accord-
ing to Micro-CT and simulation results, and this is in agreement with results reported in
[18]. This outcome is connected to the tows' high out-of-plane distortion seen in the Micro-
CT slices shown in Figures 10, 11 and 13.
The values of both A11 and A22 for both specimens are close to 0.5 in the two scans
(Areas 1 and 3 in Figures 15 and 16) below the weld lines, this is more characteristic of a
2D uniform random tow orientation. The main difference can be seen in the A11 values of
the scan results of Area 2. A11 of the Flow samples is higher than that of the Butt specimen
at the weld-line area that demonstrate a stronger dominating orientation parallel to the
weld-line when compared with the Butt sample, as would be predicted given the consid-
erable in-mould flow. Particularly, Table 7 shows A11 components of 0.64 for Flow and
0.54 for Butt, which are comparable to the measured values for the orientated specimens
in [11].
The orientation state measured by EC are showing similar trend of higher A11 com-
ponent at the weld-line region, however higher A11 values to the values measured by
Micro-CT in all 3 scan areas. To resolve this issue, further targeted studies are required to
calibrate the EC method by using unidirectional carbon fibre tows rotated at known an-
gles. This slight error can be a systematic error in alignment of the robot head to the define
x-axis coordinate system. The results match in terms of higher A11 values for the Flow
when compared to the A11 values of the Butt sample in scan area 2.
The results shown in this work are an important step towards realising the capability
of EC testing in detecting defects for quality assurance in C-SMC. An example of an EC
uniformity map obtained from a 3D generic beam geometry (versus a partially filled man-
ufactured part) is shown in Figure 17. The dark regions on the upper face indicate out-of-
plane flow and subsequent alignment of the carbon fibre tows into an underlying (out-of-
plane) internal rib structure. Unexpectedly, the EC mapping taken from the side walls of
this component also shows the presence of WLDs/resin-rich areas and out-of-plane fibre
alignments that are not predicted by simulation models. The pattern obtained from EC
testing matches the flow pattern seen in the partially filled counterpart.
J. Manuf. Mater. Process. 2022, 6, 151 23 of 25
Figure 17. (a) A photo of a partially filled 3D structural CFRP part containing two characteristic “u-
profile beams”. (b) EC real raw image obtained from the outer surfaces of a fully moulded twin
characteristic “u-profile beam” 3D part with an underlying internal rib structure.
5. Conclusions
The effect of WLDs on the morphology and mechanical properties of the selected
compression-moulded CF-SMC was shown using a combination of EC scanning, micro-
CT, and mechanical testing with DIC strain mapping. In particular, it was shown that the
introduction of a forced WLD can lead to a severe reduction in tensile strength (up to
80%). The opposing fronts were seen to behave as impermeable boundaries (analogous to
the mould cavity walls), with no evidence of fibre intermingling across the fronts at the
WLD. It was shown that this lack of fibre bridging across the WLD eliminates the contacts
between adjacent carbon fibres, forming an electrical discontinuity that breaks the EC
loops in the composite part. Carbon fibre tows were shown to be highly distorted at the
weld-line, due to the material flow phenomena, with the highest distortions occurring
near the edges of the mould. Defects such as voids and cracks were observed at the WLDs
using multiple NDT methods, such as radiography, micro-CT, and EC, and compared to
the process simulation results.
Furthermore, the technique can also successfully identify and characterise the critical
variations in fibre content and alignment inherent in “forced defect free” carbon DFCs.
Failure in DFCs occurs when a stress concentration (due to discontinuities in the micro-
scale and meso-scale structure) and material weakness coincide. Weakness in DFCs
J. Manuf. Mater. Process. 2022, 6, 151 24 of 25
originates from discontinuities, such as in the fibre–matrix interface, and variation in the
resin content in meso-scale structure. The discontinuities that lead to failure in DFCs sig-
nificantly alter the EC complex impedance signal strength, making the EC technique
highly sensitive to their presence and uniquely suited to their detection. Consequently,
the total current density in the failed region is considered to decay for three reasons, (a)
an overall reduction in the volume fraction of conductive fibre (resin-rich regions); (b) out-
of-plane orientation of the fibres, thus reducing the in-plane current density; and (c) a
continuous, linear electrical discontinuity, created due to the lack of fibre crossing. Using
this phenomenon and performing EC scans at multiple sensor orientations, it was possible
to elucidate both the local fibre-matrix content and local principal fibre angles at appro-
priate spatial resolutions for defect characterisation. Comparisons revealed that the prin-
cipal fibre angle distributions determined from such four-directional EC scans correlated
strongly with the distributions obtained by micro-CT for samples with WLDs.
As such, robot-guided EC scanning methods constitute a potential solution for de-
tecting process-induced anisotropic microstructures non-destructively. Charge placement
during the development of the moulding process for a given part can be optimised to
eliminate or move defects away from critical locations, to reduce material waste and per-
form routine quality control during production.
Author Contributions: Conceptualization, N.F.-S. and N.R.; methodology N.F.-S., N.R., D.N., C.Q.,
P.S., and R.K.; software, D.J.A., P.S., and R.K.; validation, N.F.-S., D.J.A., C.Q., and D.N.; formal
analysis, N.F.-S., N.R., D.J.A., C.Q., D.N., and P.S.; investigation, N.F.-S., N.R., P.S., R.K., and D.N.;
resources, N.R., C.Q., and K.K.; data curation, N.F.-S., C.Q., D.N., and N.R.; writing—original draft
preparation, N.F.-S.; writing—review and editing, N.F.-S., N.R., C.Q., D.N., D.J.A., and P.S.; visual-
ization, N.F.-S., D.N., and C.Q.; funding acquisition, C.Q., K.K., and M.A.W. All authors have read
and agreed to the published version of the manuscript.
Funding: This research was funded by APC10 TUCANA grant no. 113198, EPSRC Future Metrology
Hub (grant number EP/P006930)1 and the EPSRC Future Composites Manufacturing Hub (grant
number EP/P006701/1). D.J.A acknowledges support from the STFC via an Ernest Rutherford Fel-
lowship (ST/R00384X/1).
Institutional Review Board Statement: Not applicable.
Informed Consent Statement: Not applicable.
Data Availability Statement: Not applicable.
Acknowledgments: The materials used in this work were moulded as part of an APC-funded CHA-
MAELEON at WMG by Richard Groves, and the authors would like to acknowledge the APC-
funded TUCANA project consortium for their support in completing the study. We would like to
thank Expert Tooling and Automation Ltd. for providing access to the Kuka robot and Adam Wal-
ton for robot integration.
Conflicts of Interest: The authors declare no conflict of interest.
References
1. Fereshteh-Saniee, N.; Reynolds, N.; Kelly, C.A.; Wilson, P.R.; Jenkins, M.J.; Kendall, K.N. Introducing cryomilling for reliable
determination of resin content and degree of cure in structural carbon fibre reinforced thermoset composites. Compos. Part A
Appl. Sci. Manuf. 2018, 107, 197–204.
2. Gupta, J.; Reynolds, N.; Chiciudean, T.; Kendall, K. A comparative study between epoxy and vinyl ester CF-SMC for high
volume automotive composite crash structures. Compos. Struct. 2020, 244, 112299.
3. Centea, T.; Hubert, P. Measuring the impregnation of an out-of-autoclave prepreg by micro-CT. Compos. Sci. Technol. 2011, 71,
593–599.
4. Qian, C.; Harper, L.; Turner, T.; Warrior, N. Notched behaviour of discontinuous carbon fibre composites: Comparison with
quasi-isotropic non-crimp fabric. Compos. Part A Appl. Sci. Manuf. 2011, 42, 293–302.
5. Selezneva, M.; Lessard, L. Characterization of mechanical properties of randomly oriented strand thermoplastic composites. J.
Compos. Mater. 2016, 50, 2833–2851.
6. Visweswaraiah, S.B.; Selezneva, M.; Lessard, L.; Hubert, P. Mechanical characterisation and modelling of randomly oriented
strand architecture and their hybrids—A general review. J. Reinf. Plast. Compos. 2018, 37, 548–580.
J. Manuf. Mater. Process. 2022, 6, 151 25 of 25
7. Feraboli, P.; Cleveland, T.; Ciccu, M.; Stickler, P.; DeOto, L. Defect and damage analysis of advanced discontinuous car-
bon/epoxy composite materials. Compos. Part A Appl. Sci. Manuf. 2010, 41, 888–901.
8. Favaloro, A.J.; Tseng, H.-C.; Pipes, R.B. A new anisotropic viscous constitutive model for composites molding simulation. Com-
pos. Part A Appl. Sci. Manuf. 2018, 115, 112–122.
9. Chookaew, W.; Mingbunjurdsuk, J.; Jittham, P.; Ranong, N.N.; Patcharaphun, S. An Investigation of Weldline Strength in In-
jection Molded Rubber Parts. In Energy Procedia, Proceedings of the 10th Eco-Energy and Materials Science and Engineering Sympo-
sium, Ubon-Ratchathani, Thailand, 5–8 December 2012; Yupapin, P., PivsaArt, S., Ohgaki, H., Eds.; Elsevier: Amsterdam, The Neth-
erlands, 2013; pp. 767–774.
10. Chrysostomou, A.; Hashemi, S. Mechanical properties of injection moulded styrene maleic anhydride (SMA) Part II Influence
of short glass fibres and weldlines. J. Mater. Sci. 1998, 33, 4491–4501.
11. Debondue, E.; Fournier, J.-E.; Lacrampe, M.-F.; Krawczak, P. Weld-line sensitivity of injected amorphous polymers. J. Appl.
Polym. Sci. 2004, 93, 644–650.
12. Fellahi, S.; Meddad, A.; Fisa, B.; Favis, B.D. Weldlines in injection-molded parts: A review. Adv. Polym. Technol. 1995, 14, 169–
195.
13. Hashemi, S. Thermal effects on weld and unweld tensile properties of injection moulded short glass fibre reinforced ABS com-
posites. Express Polym. Lett. 2007, 1, 688–697.
14. Baradi, M.B.; Cruz, C.; Riedel, T.; Régnier, G. Mechanical and microstructural characterization of flowing weld lines in injection-
molded short fiber-reinforced PBT. Polym. Test. 2019, 74, 152–162.
15. Quintana, M.C.; Frontini, P. Weld line strength factors in a reinforced injection molded part: Relationship with predicted fiber
orientation. J. Reinf. Plast. Compos. 2019, 39, 219–230.
16. Leblanc, D.; Landry, B.; Levy, A.; Hubert, P.; Roy, S.; Yousefpour, A.; Quinlan, E. Study of Processing Conditions on the Forming
of Ribbed Features Using Randomly Oriented Strands Thermoplastic Composites. J. Am. Helicopter Soc. 2015, 60, 1–9.
17. Martulli, L.; Creemers, T.; Schöberl, E.; Hale, N.; Kerschbaum, M.; Lomov, S.V.; Swolfs, Y. A thick-walled sheet moulding com-
pound automotive component: Manufacturing and performance. Compos. Part A Appl. Sci. Manuf. 2020, 128, 105688.
18. Martulli, L.; Kerschbaum, M.; Lomov, S.V.; Swolfs, Y. Weld lines in tow-based sheet moulding compounds tensile properties:
Morphological detrimental factors. Compos. Part A Appl. Sci. Manuf. 2020, 139, 106109.
19. Teuwsen, J.; Hohn, S.K.; Osswald, T.A. Direct Fiber Simulation of a Compression Molded Ribbed Structure Made of a Sheet
Molding Compound with Randomly Oriented Carbon/Epoxy Prepreg Strands—A Comparison of Predicted Fiber Orientations
with Computed Tomography Analyses. J. Compos. Sci. 2020, 4, 164.
20. Auld, B.A.; Moulder, J.C. Review of Advances in Quantitative Eddy Current Nondestructive Evaluation. J. Nondestruct. Eval.
1999, 18, 3–36.
21. Bardl, G.; Nocke, A.; Hübner, M.; Gereke, T.; Pooch, M.; Schulze, M.; Heuer, H.; Schiller, M.; Kupke, R.; Klein, M.; et al. Analysis
of the 3D draping behavior of carbon fiber non-crimp fabrics with eddy current technique. Compos. Part B Eng. 2018, 132, 49–60.
22. Berger, D.; Lanza, G. Development and Application of Eddy Current Sensor Arrays for Process Integrated Inspection of Carbon
Fibre Preforms. Sensors 2017, 18, 4.
23. Cheng, J.; Ji, H.; Qiu, J.; Takagi, T.; Uchimoto, T.; Hu, N. Role of interlaminar interface on bulk conductivity and electrical
anisotropy of CFRP laminates measured by eddy current method. NDT E Int. 2014, 68, 1–12.
24. Heuer, H.; Schulze, M.; Pooch, M.; Gäbler, S.; Nocke, A.; Bardl, G.; Cherif, C.; Klein, M.; Kupke, R.; Vetter, R.; et al. Review on
quality assurance along the CFRP value chain—Non-destructive testing of fabrics, preforms and CFRP by HF radio wave tech-
niques. Compos. Part B Eng. 2015, 77, 494–501.
25. Hughes, R.R.; Drinkwater, B.W.; Smith, R.A. Characterisation of carbon fibre-reinforced polymer composites through radon-
transform analysis of complex eddy-current data. Compos. Part B Eng. 2018, 148, 252–259.
26. Schulze, M.H.; Heuer, H.; Küttner, M.; Meyendorf, N. High-resolution eddy current sensor system for quality assessment of
carbon fiber materials. Microsyst. Technol. 2010, 16, 791–797.
27. Bardl, G.; Nocke, A.; Cherif, C.; Pooch, M.; Schulze, M.; Heuer, H.; Schiller, M.; Kupke, R.; Klein, M. Automated detection of
yarn orientation in 3D-draped carbon fiber fabrics and preforms from eddy current data. Compos. Part B Eng. 2016, 96, 312–324.
28. Khan, A.M.; Bardl, G.; Nocke, A.; Cherif, C. Quality analysis of 2D and 3D-draped carbon preforms by eddy current scanning.
Compos. Part B Eng. 2019, 176, 107110.
29. Suragus GmbH. Available online: https://www.suragus.com/en/products/carbon-fiber-testing/portable-devices/portable-
handheld-for-carbon-fiber-testing/ (accessed on 15 October 2019).
30. Feraboli, P.; Peitso, E.; Cleveland, T.; Stickler, P.B. Modulus Measurement for Prepreg-based Discontinuous Carbon Fiber/Epoxy
Systems. J. Compos. Mater. 2009, 43, 1947–1965.
31. Evans, A.; Qian, C.; Turner, T.; Harper, L.; Warrior, N. Flow characteristics of carbon fibre moulding compounds. Compos. Part
A Appl. Sci. Manuf. 2016, 90, 1–12.
32. Feraboli, P.; Peitso, E.; Cleveland, T.; Stickler, P.B.; Halpin, J.C. Notched behavior of prepreg-based discontinuous carbon fi-
ber/epoxy systems. Compos. Part A Appl. Sci. Manuf. 2009, 40, 289–299.
