# Machinability analysis of UD-GFRP composites in edge trimming with diamond-coated burr tools at various fiber orientations

Benoît Schrab, Anne Collaine, Jean-Marie Freyburger, Michel Tourlonias

► To cite this version:

Benoît Schrab, Anne Collaine, Jean-Marie Freyburger, Michel Tourlonias. Machinability analysis of UD-GFRP composites in edge trimming with diamond-coated burr tools at various fiber orientations. CIRP Journal of Manufacturing Science and Technology, 2025, 59, pp.194-206. 10.1016/j.cirpj.2025.03.008. hal-05162679

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Submitted on 15 Jul 2025

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# Machinability analysis of UD-GFRP composites in edge trimming with diamond-coated burr tools at various fiber orientations

Benoît Schrab *, Anne Collaine, Jean-Marie Freyburger, Michel Tourlonias

Université de Haute-Alsace, Laboratoire de Physique et Mécanique Textiles (LPMT, UR 4365), École Nationale Supérieure d'Ingénieurs Sud Alsace, 11 rue Alfred Werner, 68093 Mulhouse, France

<figure><img src="image_2.png" /></figure>

## ARTICLE INFO

**Keywords:**

GFRP

Edge trimming

Burr tools

Fiber orientation

Cutting forces

Temperature

Machining quality

## ABSTRACT

Machining operations that are usually required during the manufacturing process of fiber-reinforced polymers (FRPs) are likely to generate numerous specific defects on the machined parts. Cutting tools with specific geometries, such as burr tools with many pyramidal teeth, help to reduce this type of defect. Previous studies in milling with this kind of tool focused mainly on multidirectional (MD) FRP composites. However, it is well known that fiber orientation is a critical parameter in FRP machining. Accordingly, this paper presents a machinability analysis of unidirectional (UD) FRP composites in edge trimming at various fiber orientations, for the specific case of burr tools. Edge trimming experiments are conducted using glass fiber reinforced polymer (GFRP) composites and diamond-coated burr tools. For each machining test, cutting forces, machined surface temperature and surface quality are quantified, as well as defects on external plies, which are characterized with a device specifically developed for this purpose. The results demonstrate that all the parameters are affected greatly by fiber orientation. The maximum temperature rise and the resultant force generally move in the same direction. The surface irregularities, and the maximum rise in temperature, move globally in opposite directions. Uncut fibers on the lower face of the workpiece are related to axial forces and may be facilitated by high temperatures. Moreover, the effect of fiber orientation on cutting forces in the workpiece plane, machined surface temperature and surface quality is analogous to previous literature on orthogonal cutting.

## 1. Introduction

Fiber-reinforced polymers (FRPs) are widely used in a large range of applications because they offer high stiffness and strength for a lower weight than metals [1]. Glass and carbon fibers are the main reinforcement materials, with glass fiber reinforced polymers (GFRPs) accounting for over 90 % of the market in volume in 2022 [2]. Machining operations are generally necessary in the manufacturing process of FRP parts: these include the use of edge trimming to finish the periphery of the part and drilling for its fastening [3]. FRP machining is challenging due to the heterogeneous and anisotropic nature of the material. Indeed, a range of specific defects may occur, such as cracks, fiber pullout or matrix degradation on the machined surface, as well as delamination or uncut fibers on the surface plies [4].

In order to understand the complex phenomena involved in FRP machining, several authors have performed orthogonal cutting experiments, which revealed *inter alia* the critical role of fiber orientation. This

parameter was reported to affect cutting forces [5,6], surface quality [7, 8] and the temperature during cutting [9,10]. FRP cutting mechanisms in orthogonal cutting were identified, and were found to depend on fiber orientation together with cutting edge geometry [5,11]. Five different cutting mechanisms were summarized by Sheikh-Ahmad [3]: delamination (type I), fiber buckling (type II), fiber cutting (types III and IV) and macrofracture (type V). These mechanisms were identified primarily in the case of carbon fiber reinforced polymers (CFRPs). However, they were also mentioned in the case of GFRP composites [12], consistently with the fact that both carbon and glass are inorganic fibers that fracture in a brittle manner [3]. In more recent work, Meng *et al.* [13] have proposed a similar classification with tearing type, crushing type and bending type.

Although orthogonal cutting experiments give us a fundamental understanding of FRP machining, they do not take into account all the aspects involved in a real milling operation. In fact, these operations induce variable chip thickness and cutting speed direction, as well as

<footnote>* Corresponding author.</footnote>

<footnote>E-mail addresses: benoit.schrab@uha.fr (B. Schrab), anne.collaine@uha.fr (A. Collaine), jean-marie.freyburger@uha.fr (J.-M. Freyburger), michel.tourlonias@uha.fr (M. Tourlonias).</footnote>
teeth entry and exit phenomena. Moreover, complex tool geometries with several cutting edges may be used. For all these reasons, it is also necessary to study milling operations. In this paper, we focus on multi-tooth tools with pyramidal teeth, often referred to as burr tools in previous literature [14], which are widespread in the industry for FRP machining [14,15]. This geometry was developed to avoid the axial force generated by conventional helical geometries, which may lead to vibrations and surface ply delamination [16,17] in particular. Burr tools can feature different tooth sizes and may include an additional cutting edge on the top of each tooth [18].

Numerous studies have investigated CFRP milling with burr tools. Many authors have worked on cutting parameters, especially the impact of cutting speed and feed speed, feed per revolution or feed per tooth on surface roughness [15,19], defect occurrence [15,19], cutting forces [20] and temperature [21,22]. Tool wear has also been studied, and two wear patterns were reported by Janardhan *et al.* [23], namely tooth tip fracture and flank wear. Moreover, dedicated direct and indirect wear characterization or monitoring approaches for burr tools have been put forward [16,24]. In the case of GFRP composites, almost no contribution was found about milling with burr tools. Ducobu *et al.* [25] performed slot milling on multidirectional (MD) GFRP composites with different tool geometries and materials, including uncoated carbide burr tools. However, only the end of the tool, which had a different design, was used for milling experiments, without taking full advantage of the multi-tooth geometry. As a result, the tool performance was relatively poor in terms of delamination and cutting forces. Girot *et al.* [26] also studied GFRP milling with a burr tool and observed low wear resistance compared to polycrystalline diamond (PCD) tools. This can be explained by the fact that the material of the burr tool was uncoated carbide. Additionally, burr tools provided good results in the hole milling of sheet molding compound (SMC) composites reinforced with chopped glass fibers [27].

Earlier studies on milling with burr tools focused mainly on MD-FRP composites, where fiber orientation influence is complex. For a better understanding of the effect of this parameter, which has been found to govern FRP cutting mechanisms in orthogonal cutting, it is required to work on unidirectional (UD) FRP. The aim of this paper, therefore, is to study the effect of fiber orientation in the edge trimming of UD-GFRP composites with diamond-coated burr tools. During the experiments, machining is monitored for the cutting forces and machined surface temperature, while the machining quality is characterized after each test. This data enables a multi-parameter approach of the fiber orientation effect, which is absent in the current literature for the specific case of milling with burr tools.

In the following sections, we first introduce the experimental setup and the measurement devices and methods. We then discuss the impact of fiber orientation on the different parameters, as well as the correlations between these parameters.

## 2. Materials and methods

### 2.1. Edge trimming experiments

#### 2.1.1. Workpiece material

A UD-GFRP laminate is manufactured from HexPly 913/28%/192/ EC9756 prepregs (24 layers) and cured in a hydraulic press. The main reinforcement and matrix properties are given in Table 1. Thermal

properties are given at room temperature. After curing, the fiber architecture is the same throughout the thickness of the workpiece. The main specifications of the final material are given in Table 2. The plates for edge trimming experiments (90 *40 *3.15 mm³) are cut in the laminate in order to obtain four different fiber orientations, namely 0°, 45°, 90° and 135°.

#### 2.1.2. Machining experiments

A Mécanuméric Charlyrobot 2U three-axis milling machine with a spindle power of 1.1 kW, a maximum spindle speed of 24,000 rpm and a maximum feed speed of 6,000 mm/min is used for the experiments. An overview of the experimental setup is shown in Fig. 1.

Each edge trimming test is conducted with the following procedure. First, the composite plate is positioned laterally using a positioning guide, and the cantilever length is roughly fixed. The plate is then held in position by a clamping element with two screws tightened to 3 Nm. Subsequently, the cantilever length is adjusted precisely to 15.5 mm with a first edge trimming pass using a dedicated tool in order to guarantee that the radial engagement of the final pass is the same for all the experiments. The final edge trimming pass consists of down-milling the full plate thickness following a straight path parallel to the front of the clamping system. For all the experiments, this machining pass is conducted with a SECO 871060.0-DURA burr tool, which is diamond coated to improve wear resistance. Its geometry is described in Fig. 2 using the terminology proposed by Gara *et al.* [18], and its properties are given in Table 3. Each tooth tip includes an additional cutting edge, which is referred to as the main cutting edge.

Cutting parameters are the same for all the experiments and are given in Table 4. A small radial engagement is used compared to the tool diameter, as edge trimming operations on FRPs are generally finishing operations. The cutting speed and the feed speed values are chosen to limit vibrations and thus prevent tool coating breakage. Indeed, preliminary tests with higher values have led to heavy vibrations when milling perpendicular to the fibers. It should be noted that the cutting speed and feed speed values are lower than those recommended by the tool supplier for the selected radial engagement. However, these recommended values do not take into account material properties like fiber fraction or fiber orientation.

Fiber orientation during cutting is defined similarly to Sheikh-Ahmad *et al.* [29] by the $Ψ'$ angle with respect to feed speed, and the $θ$ angle with respect to cutting speed (Fig. 3). Hereafter, $Ψ'$ angle is called “workpiece fiber orientation”, which takes the four different values mentioned in Section 2.1.1 (0°, 45°, 90° and 135°), while $θ$ angle is called “instantaneous fiber orientation”. The different instantaneous fiber orientation $θ$ ranges associated with each workpiece fiber orientation $Ψ'$, in the case of the selected radial engagement, are given in Table 5.

Five identical edge trimming tests are performed for each workpiece fiber orientation $Ψ'$. To avoid any wear-related effect, a different new cutting tool is used for each orientation, and the five tests are then carried out with the same tool. Between each test, the condition of the cutting tool is checked (especially the diamond coating integrity) and the tool is cleaned. Besides, the time between two tests on the same plate is sufficient to allow the workpiece to return to the room temperature.

<table><caption>**Table 1**
Reinforcement and matrix properties.</caption><thead><tr><th></th><th>Reinforcement: E-glass</th><th>Matrix: epoxy</th></tr></thead><tbody><tr><td>Density (g/cm³)</td><td>2.63</td><td>1.23</td></tr><tr><td>Area weight (g/m²)</td><td>192</td><td>-</td></tr><tr><td>Thermal conductivity (W/m K)[28]</td><td>1.3</td><td>0.19</td></tr><tr><td>Specific heat (J/kg K)[28]</td><td>810</td><td>1050</td></tr></tbody></table>

<table><caption>**Table 2**
UD-GFRP laminate specifications (with the sample standard deviations as error values).</caption><tr><td>Thickness (mm)</td><td>3.15 ± 0.01</td></tr><tr><td>Glass transition temperature<br>Tg onset (DMA - 5 °C/min - 1 Hz - 15 µm) (°C)</td><td>127.8 ± 2.3</td></tr><tr><td>Fiber weight fraction (%)</td><td>72 ± 1</td></tr><tr><td>Longitudinal flexural modulus (GPa)</td><td>42.3 ± 0.8</td></tr><tr><td>Transverse flexural modulus (GPa)</td><td>13.3 ± 0.1</td></tr></table>
<figure><img src="image_3.png" /><figcaption>Fig. 1. overview of the experimental setup.</figcaption></figure>

<figure><img src="image_4.png" /><figcaption>Fig. 2. cutting tool geometry.</figcaption></figure>

<table><caption>**Table 3**
Cutting tool properties.</caption><tr><td>Diameter (mm)</td><td>6</td></tr><tr><td>Right helixes number</td><td>10</td></tr><tr><td>Left helixes number</td><td>12</td></tr><tr><td>Measured helix angle (°)</td><td>29.3</td></tr><tr><td>Measured main cutting edge radius (μm)</td><td>≈ 20</td></tr><tr><td>Measured coating thickness (μm)</td><td>≈ 9</td></tr></table>

<table><caption>**Table 4**
Cutting parameters.</caption><tr><td>Cutting speed Vc (m/min)</td><td>115</td></tr><tr><td>Feed speed Vf (mm/min)</td><td>732</td></tr><tr><td>Depth of cut a<sub>p</sub> (mm)</td><td>3.15</td></tr><tr><td>Radial engagement a<sub>ε</sub> (mm)</td><td>0.5</td></tr></table>

## 2.2. Online monitoring

### 2.2.1. Cutting forces

Cutting forces during the edge trimming experiments are measured using a Kistler 9129AA piezoelectric dynamometer, which is located under the clamping system (Fig. 4). Fx is the force measured perpendicular to the feed direction, Fy corresponds to the force in the feed direction, and Fz is the force in the tool axis direction. The signals are then processed with a Kistler LabAmp type 5167A charge amplifier and recorded with a Brüel & Kjaer 3053-B-120 input module, using a 65.536 kHz sampling frequency. Instantaneous force data for a given tooth is not available due to the complex geometry of the cutting tool. Thus, cutting forces are analyzed globally in the dynamometer fixed measurement directions. For each edge trimming test, the average forces in the three directions Fx<sub>avg</sub>, Fy<sub>avg</sub>, Fz<sub>avg</sub> are calculated in the steady state machining region (Fig. 5). Additionally, the resultant of average forces Fr (N) is also calculated according to Eq. 1:

$$F_r = \sqrt{F_{x\text{avg}}^2 + F_{y\text{avg}}^2 + F_{z\text{avg}}^2} \quad (1)$$

where Fx<sub>avg</sub>, Fy<sub>avg</sub> and Fz<sub>avg</sub> are the average forces in each direction (N).

### 2.2.2. Machined-surface temperature

The temperature of the machined surface is measured with an Optris CTLT20CF infrared (IR) sensor facing the surface (Fig. 4). The measuring spot has a diameter of around 2.3 mm and is centered approximately on the surface by laser sight. The sensor is connected to an electronic box, which carries out signal averaging and makes it possible to adjust the emissivity ε of the surface. The emissivity measurement of composite machined surfaces is relatively challenging due to the heterogeneity and anisotropy of the material. As a result, this parameter is maximized to ε = 1, which means that the true temperature is higher than the measured temperature. The temperature signal is then recorded synchronously with cutting forces.

## 2.3. Post-machining characterization

### 2.3.1. Surface quality

Surface quality is characterized by areal parameters on 3D surfaces captured with a Keyence VHX-6000 digital microscope, which is also used for qualitative observation. The machined surface is cleaned prior to characterization, using a vacuum cleaner and a fine brush. Analysis is carried out at two scales: the small scale refers to fiber (or a few fibers)

<table><caption>**Table 5**
Instantaneous fiber orientation θ ranges associated with each workpiece fiber orientation Ψ.</caption><thead><tr><td>Ψ (°)</td><td>θ<sub>entry</sub> (°)</td><td>θ<sub>exit</sub> (°)</td></tr></thead><tbody><tr><td>0</td><td>33.6</td><td>0</td></tr><tr><td>45</td><td>168.6</td><td>135</td></tr><tr><td>90</td><td>123.6</td><td>90</td></tr><tr><td>135</td><td>78.6</td><td>45</td></tr></tbody></table>

<figure><img src="image_5.png" /><figcaption>Fig. 3. workpiece fiber orientation Ψ and instantaneous fiber orientation θ.</figcaption></figure>
<figure><img src="image_6.png" /><figcaption>Fig. 4. cutting forces and temperature measurement.</figcaption></figure>

<figure><img src="image_7.png" /><figcaption>Fig. 5. average force in the steady state machining region.</figcaption></figure>

scale irregularities, and the large scale includes larger scale irregularities (Fig. 6). These two scales are obtained by Gaussian filtering with the nesting indexes Nis, Nic and Nif on the basis of ISO 25178–3 standard [30].

The fiber diameter is measured to approximately 10 µm, and the small-scale region is set up to five fiber diameters. The ISO standard recommends setting the nesting index value to five times the largest considered scale, which gives Nis = 0.25 mm. Following the ISO standard, the small-scale evaluation area is then a square with a side length of Nis value. Nif value is set to the highest standardized value not exceeding the composite plate thickness, because Nif is also the side length of the large-scale square evaluation area. Lastly, Nis value is chosen based on the recommendation of the microscope manufacturer, which is to take a greater value than three times the measurement resolution in the surface plane. For each edge trimming test, the arithmetic mean height Sa and the maximum height Sz standardized parameters [31] are evaluated in four zones at large scale and height zones at small scale.

### 2.3.2. Delamination and uncut fibers

Previous studies reported that delamination and uncut fibers are related [32,33] and may be considered as one single defect [32].

<figure><img src="image_8.png" /><figcaption>Fig. 6. scales for surface quality characterization.</figcaption></figure>

Consequently, this study focuses on uncut fibers for the assessment of surface plie defects. An optical device consisting of a Triton TRI050S-PC camera with a 3.45 µm/px resolution sensor and a 0.35 magnification Opto Engineering TC23024 telecentric lens is used for image acquisition. This device was specifically developed at the laboratory. The camera is positioned at a 45° angle relative to the composite plate, which makes it possible to separately characterize uncut fibers on the upper and lower surface plies. Image acquisition and processing is controlled by a Matlab code, whose main steps are presented in Fig. 7. After acquisition, image thresholding using Otsu’s method [34] and binarization are performed to extract the composite plate with uncut fibers from the background. The fibers/machined surface boundary is then defined. The image may subsequently undergo angle correction and filtering to remove any unwanted residues from the background. Finally, the average uncut fiber length L<sub>avg</sub> (mm) and the maximum uncut fiber length L<sub>max</sub> (mm) are determined. L<sub>avg</sub> is calculated according to Eq. 2, similarly to previous authors [35]:

$$L_{\text{avg}} = \frac{2 A_{\text{fibers}}}{W_{\text{plate}}} \qquad (2)$$

where A<sub>fibers</sub> is the uncut fibers area (mm²), determined taking into account the camera inclination angle (Fig. 7), and W<sub>plate</sub> is the width of the composite plate (mm). These parameters are calculated on half of the composite plate width (W<sub>plate</sub>/2) because the entire plate width cannot be characterized at once. Therefore, for each face of the composite plate (upper or lower), two measurements are necessary.

The measurement accuracy of the device is estimated horizontally over a length of 20 mm (W<sub>plate</sub>/2), and vertically over a length of 4 mm, which corresponds approximately to the depth of field. The results are given in Table 6. Besides, the repeatability of the device over 10
<figure><img src="image_9.png" /><figcaption>Fig. 7. main image processing steps.</figcaption></figure>

measurements is also quantified under different conditions. The co-efficients of variation (CV) obtained for $L_{avg}$ and $L_{max}$ under the most unfavorable conditions (i.e. the highest CV values) are given in Table 7.

## 3. Results and discussion

### 3.1. Cutting forces

The average cutting forces $Fx_{avg}$, $Fy_{avg}$, $Fz_{avg}$ and the resultant force $Fr$ are given for each workpiece fiber orientation $Ψ$ in Fig. 8. On these graphs and throughout this paper, error bars correspond to twice the sample standard deviation.

For all the edge trimming tests, it is observed that average cutting forces are positive in the three measurement directions. In the workpiece plane, this means that $Fx_{avg}$ is directed towards the inside of the material, while $Fy_{avg}$ is in the opposite direction to the feed speed. This is consistent with the direction of the cutting speed during the whole tool engagement in the workpiece (Fig. 3). In the tool axis direction ($Fz_{avg}$), it means that the cutting tool generates a downward axial force on the workpiece. Considering the tooth geometry (Fig. 2), it appears that secondary cutting edges induce both upward and downward axial forces depending on their helix direction, while the main cutting edge produces a downward axial force, which has a dominant effect. Moreover, the number of left helices inducing downward forces is higher than the number of right helices inducing upward forces. This downward axial force corresponds to the tool specifications provided by the manufacturer [36].

Irrespective of workpiece fiber orientation $Ψ$, it is observed that the highest values are obtained for $Fx_{avg}$ and the lowest for $Fz_{avg}$. In particular, it is noted that $Fx_{avg}$ is always higher than $Fy_{avg}$, which indicates that, contrary to expectations, cutting forces are lower in the feed direction than in the direction perpendicular to the feed. The same phenomenon has also been reported previously in edge trimming of MD-CFRP composites with burr tools [37]. This phenomenon, which is particularly pronounced for $Ψ = 0°$ and $Ψ = 135°$, probably reflects a significant pushing action of the workpiece against the clearance faces of the teeth.

Fiber orientation in the workpiece has a major impact on our results with different trends depending on the measurement direction. $Fx_{avg}$ and

<table><caption>**Table 6**
Measurement error in horizontal and vertical directions.</caption><thead><tr><td>Measurement direction</td><td>Measured value (mm)</td><td>Relative error (%)</td></tr></thead><tbody><tr><td>Horizontal</td><td>20.03</td><td>0.1</td></tr><tr><td>Vertical</td><td>3.99</td><td>-0.3</td></tr></tbody></table>

<table><caption>Table 7<br>Repeatability estimation over 10 measurements.</caption><thead><tr><th>Parameter</th><th>CV (%)</th></tr></thead><tbody><tr><td>L<sub>avg</sub></td><td>11.26</td></tr><tr><td>L<sub>max</sub></td><td>2.75</td></tr></tbody></table>

$Fz_{avg}$ are the highest for $Ψ = 135°$, while $Fy_{avg}$ is the largest for $Ψ = 90°$. Irrespective of the measurement direction, the lowest values are all obtained for the same orientation $Ψ = 45°$. Hence, fiber orientation in the workpiece, which is defined in the workpiece plane, has an impact on cutting forces in this plane; but it also affects the axial forces perpendicular to this plane. In overall terms, the resultant force is the lowest for $Ψ = 45°$ and the highest for $Ψ = 135°$. An increase of 293 % is obtained between these two values, which highlights the strong effect of fiber orientation on cutting forces.

In terms of instantaneous fiber orientation $θ$, these results are similar to previous literature on orthogonal cutting of UD-FRP composites. Indeed, the evolution of $Fy_{avg}$ is consistent with the trend observed for the cutting force in orthogonal cutting. In the case of UD-CFRP composites, several studies showed that the cutting force generally increases up to $θ = 90°$, where it reaches its maximum, and then tends to decrease for $θ > 90°$ [5,13]. Some research into the orthogonal cutting of UD-GFRP composites in the range $0° ≤ θ ≤ 90°$ reported similar trends [6,7]. Moreover, the evolution of $Fx_{avg}$ is consistent with the trend obtained for the normal force in orthogonal cutting. In the case of UD-CFRP composites, the normal force was generally found to increase up to a maximum value at around $θ = 45°$ [5,13], and then decrease to a minimum value at around $θ = 150°$ before increasing again [13,38]. Some studies have also observed a maximum normal force for $θ = 10°$ [38] and a minimum for $θ = 165°$ [5]. For UD-GFRP composites, a similar trend has been observed by some authors up to $θ = 90°$ [6], but not by others [7].

### 3.2. Machined surface temperature

A typical temperature signal recorded during edge trimming experiments is shown in Fig. 9 for a machining test performed at lower feed speed (200 mm/min). After cutting tool entry in the composite plate, the temperature rises slowly as the tool approaches the measurement zone due to heat conduction in the workpiece. The tool then gradually enters the sensor field and in the central zone, the sensor only measures the tool temperature, which increases. Additional tests with an IR camera show that it keeps rising over the entire feed length, which indicates that a
<figure><img src="image_10.png" /><figcaption>Fig. 8. average cutting forces (a) and resultant force (b) for each workpiece fiber orientation $Ψ$.</figcaption></figure>

<figure><img src="image_11.png" /><figcaption>Fig. 9. typical temperature signal.</figcaption></figure>

temperature steady state could not be reached in these tests. Thus, the temperature analysis is only a comparative analysis. The cutting tool subsequently leaves the sensor field and the temperature then decreases as it moves away. This temperature decrease is governed by two phenomena: heat conduction from the cutting zone and natural cooling by convection. Finally, a temperature drop is observed as the tool exits from the material.

An apparently linear decrease is observed after the peak temperature, while it is the superposition of different nonlinear phenomena. Indeed, Mehnen et al. [39] obtained a nonlinear temperature decrease considering only heat conduction, and heat convection results in an exponential decay, according to Newton's law of cooling. However, the tool temperature keeps rising in our case. Besides, our temperature decrease may actually be nonlinear over a longer period of time.

In the following, temperature analysis is carried out using the maximum temperature rise $\Delta T_{max}$ and the cooling rate $CR$. The temperature measured by the sensor is greater on the machined surface close to the cutting tool than on the cutting tool itself, and thus $\Delta T_{max}$ represents the maximum temperature rise on the machined surface. It is a crucial parameter since thermal defects occur for excessive peak temperatures. $CR$ describes how quickly we leave this potentially critical peak temperature. It is defined as the absolute value of the temperature

slope in a fixed zone between the peak temperature and the tool exit.

The maximum temperature rise $\Delta T_{max}$ is given for each workpiece fiber orientation $Ψ$ in Fig. 10. It is observed that the maximum temperature rise $\Delta T_{max}$ is the lowest for $Ψ = 45°$, and the highest for $Ψ = 90°$, with a slightly lower result for $Ψ = 135°$, while $Ψ = 0°$ gives an intermediate value. Thus, fiber orientation in the workpiece has a great

<figure><img src="image_12.png" /><figcaption>Fig. 10. maximum temperature rise $\Delta T_{max}$ for each workpiece fiber orientation $Ψ$.</figcaption></figure>
impact on the machined surface temperature in the immediate vicinity of the cutting zone, with an increase of 260 % between the minimum and the maximum value.

Comparing these results with cutting forces (Fig. 8), it is observed that $\Delta T_{max}$ and Fr follow globally a similar trend, which indicates that the mechanical energy required for cutting is mainly dissipated as heat, as stated by other authors [40]. However, $\Delta T_{max}$ is higher for $\Psi' = 90^\circ$ than for $\Psi' = 135^\circ$, while the highest Fr value is obtained for $\Psi' = 135^\circ$. Thus, other phenomena related to fiber orientation are involved. In particular, friction may have a significant impact on heat generation. Indeed, Nayak et al. [41] reported, in a limited fiber orientation range, a great influence of fiber orientation on the friction coefficient. Other authors observed a slighter impact of fiber orientation on the apparent friction coefficient when measured close to the cutting process [42]. Their results tend to show that friction could be slightly higher for $\Psi' = 90^\circ$ than for $\Psi' = 135^\circ$.

In terms of instantaneous fiber orientation $\theta$, the evolution of the maximum temperature rise $\Delta T_{max}$ is similar to the trend reported by some authors in the orthogonal cutting of UD-CFRP composites. Indeed, An et al. [9] measured the machined surface temperature and reported an increase up to $\theta = 90^\circ$, followed by a decrease with a small secondary peak for $\theta = 165^\circ$. Wang et al. [10] obtained a similar trend by measuring the maximum tool temperature. Our results are also partly consistent with those of Zitoune et al. [43], who reported in the range $0^\circ \le \theta \le 90^\circ$ a maximum tool temperature increase for $\theta = 90^\circ$, but a minimum rise for $\theta = 45^\circ$.

The cooling rate CR of the machined surface is given for each workpiece fiber orientation $\Psi'$ in Fig. 11. The cooling rate CR is also affected by the workpiece fiber orientation and follows almost the same trend as $\Delta T_{max}$. This is consistent with Newton's law of cooling, according to which the cooling rate of a body is greater for a large temperature increase with respect to the temperature of its environment. However, similar CR values are obtained for $\Psi' = 90^\circ$ and $\Psi' = 135^\circ$, while $\Delta T_{max}$ is clearly higher for $\Psi' = 90^\circ$. It follows that other phenomena related to fiber orientation are involved. On the one hand, heat convection may be impacted by fiber orientation. Indeed, previous literature reported that fiber orientation has an influence on machined surface quality, as mentioned previously (Section 1), and it is known that surface geometry affects the convection heat transfer coefficient [44]. On the other hand, heat dissipation in the workpiece is also affected by fiber orientation. Indeed, E-glass fibers have a higher thermal conductivity than epoxy resin (Table 1), which favors heat dissipation in the fiber direction. This phenomenon is illustrated in Fig. 12, where the thick arrows represent heat transfer in the fibers, and thin arrows heat transfer in the resin.

For $\Psi' = 135^\circ$ (Fig. 12d), a significant pre-heating ahead of the cutting tool is observed, which may contribute to the high $\Delta T_{max}$ value, while the temperature decreases rapidly after the tool has passed with a high CR value. This is explained by the fact that the heat generated by the cutting process is mainly transferred forwards by conduction in the fibers. For $\Psi' = 90^\circ$ (Fig. 12c), no significant pre-heating is observed but

<figure><img src="image_13.png" /><figcaption>Fig. 11. cooling rate CR for each workpiece fiber orientation $\Psi'$.</figcaption></figure>

again a high CR value is obtained, because the heat transfer in the fibers occurs perpendicularly to the feed direction. In these two cases, no heat transfer in the fibers can occur backwards from the cutting zone, which induces high CR values. However, heat dissipation in the fibers is more efficient for $\Psi' = 135^\circ$, because the free edge of the workpiece is reached more quickly. This can explain why the CR value is the same for $\Psi' = 90^\circ$ and $\Psi' = 135^\circ$, even if $\Delta T_{max}$ is higher for $\Psi' = 90^\circ$. For $\Psi' = 45^\circ$ (Fig. 12b), a significantly lower CR value is observed without preheating, because the heat is mainly transferred backwards. Lastly, when $\Psi' = 0^\circ$ (Fig. 12a), an intermediate CR value is reached, because the heat is primarily transferred in the forward direction but also in the backward direction. In this case, heat conduction in the fibers occurs directly towards the free edges of the workpiece and some fibers may be exposed. This results in efficient heat dissipation and thus low preheating of the workpiece.

Hence, in our experimental conditions, heat conduction in the workpiece and natural cooling by convection induce concordant effects on the cooling rate CR of the machined surface depending on workpiece fiber orientation $\Psi'$.

### 3.3. Surface quality

Typical machined surfaces are presented in Fig. 13 for each workpiece fiber orientation $\Psi'$. We observe great surface irregularities for $\Psi' = 45^\circ$, whereas rather smooth surfaces are achieved for the other fiber orientations. Machining marks tend to appear in the direction perpendicular to feed for $\Psi' = 90^\circ$ and $\Psi' = 135^\circ$, whereas exposed glass fibers can be noticed on the surface for $\Psi' = 0^\circ$.

$Sa$ and $Sz$ parameters at large and small scale are also given for each workpiece fiber orientation $\Psi'$ in Fig. 14. We observe similar trends for both parameters, which indicates that surface irregularities are relatively uniform, without significant localized defects. Moreover, the results are very similar at both scales. Thus, fiber orientation in the workpiece has almost the same impact on surface irregularities from the fiber scale to the macro-scale.

Furthermore, quantitative parameters confirm the qualitative observation of workpiece fiber orientation effect. Indeed, the highest $Sa$ and $Sz$ values are clearly obtained for $\Psi' = 45^\circ$, at large as well as at small scale, which means that the surface has significant irregularities down to the fiber scale. In contrast, $Sa$ and $Sz$ are far smaller at both scales for the three other fiber orientations with minimum values for $\Psi' = 135^\circ$.

Comparing these results with cutting forces (Fig. 8), it is noted that great surface irregularities are associated with a low resultant force Fr, while smooth surfaces are rather obtained for high Fr values. Thus, high cutting forces do not lead to substantial defects on the machined surface. Moreover, great surface irregularities are also associated with low maximum temperature rise $\Delta T_{max}$ (Fig. 10), while smooth surfaces tend to be obtained for higher $\Delta T_{max}$ values. This might be explained by the adhesive friction between the teeth clearance faces and the workpiece, which might be greater in the case of smooth surfaces because of a larger real contact area. According to the literature [41,42], fiber orientation has an impact on the friction coefficient, even if its influence is more or less high, as already mentioned in Section 3.2.

In terms of instantaneous fiber orientation $\theta$, the evolution of the $Sa$ and $Sz$ parameters at both measurement scales is consistent with the trend observed for linear surface roughness parameters in the orthogonal cutting of UD-CFRP composites. Using different roughness parameters, various authors have reported maximum values at around $\theta = 135^\circ$ or $\theta = 150^\circ$, while very small variations have been noted in the range $0^\circ \le \theta \le 90^\circ$ [8,13]. In the case of UD-GFRP composites, a minimum surface roughness was found between $\theta = 30^\circ$ and $\theta = 60^\circ$ for measurements in the range $0^\circ \le \theta \le 90^\circ$ [7].

This analysis is illustrated by scanning electron microscope (SEM) images in Fig. 15. These images correspond to edge trimming tests performed with the same composite plates, the same tool geometry and
<figure><img src="image_14.png" /><figcaption>Fig. 12. heat transfer in the workpiece Q and temperature evolution for $Ψ = 0°$ (a), $Ψ = 45°$ (b), $Ψ = 90°$ (c) and $Ψ = 135°$ (d).</figcaption></figure>

the same procedure than described in Section 2.1. However, after a given machining length, cutting is interrupted and the tool exits from the workpiece perpendicular to the machined surface (Fig. 15a). This allows to observe the rounded area corresponding to tool engagement. The same experiment is conducted for each workpiece fiber orientation $Ψ$ (Fig. 15b - Fig. 15e).

We observe that surface texture changes with instantaneous fiber orientation $θ$, between tooth entry and tooth exit. However, this is less pronounced for $Ψ = 45°$ and even less for $Ψ = 135°$. For a given $Ψ$ value, the surface texture at tooth exit (right side of each image) corresponds to the image in Fig. 13 for the same fiber orientation.

The evolution of surface texture with instantaneous fiber orientation reflects the continuous sequence of cutting mechanisms over tool engagement. Overall, these mechanisms are consistent with the orthogonal cutting mechanisms mentioned previously (Section 1), which are indicated for $θ_{entry}$ and $θ_{exit}$ at the bottom of each image. Indeed, between $θ = 34°$ and $θ = 0°$ ($Ψ = 0°$), the machined surface is consistent with an evolution from type III/IV mechanism (shearing of the fibers) to type I (delamination). Between $θ = 124°$ and $θ = 90°$ ($Ψ = 90°$), the surface texture evolution can be explained by a change from type V mechanism (macrofracture) to type III/IV. Between $θ = 169°$ and $θ = 135°$ ($Ψ = 45°$), the randomly distributed hollow areas on the surface reflect a type V mechanism over the entire tool engagement. In this case, the evolution of surface texture is due to the change of instantaneous fiber orientation for the same cutting mechanism. Finally, between $θ = 79°$ and $θ = 45°$ ($Ψ = 135°$), the surface has

fewer irregularities, which is consistent with a type III/IV mechanism over the entire tool engagement. Less surface texture variations are observed between tooth entry and tooth exit, because of the intrinsic nature of this cutting mechanism.

These cutting mechanisms are also consistent with the resultant force $Fr$ (Fig. 8). Indeed, type III/IV mechanism, that involves shearing of the fibers, induces a high resultant force, compared to type I and type V mechanisms, where the material fracture is initiated by delamination. This is the reason why the highest $Fr$ value is obtained for $Ψ = 135°$ (type III/IV mechanism) and the lowest for $Ψ = 45°$ (type V mechanism), whereas intermediate values are reached for $Ψ = 0°$ (combination of type III/IV and type I) and $Ψ = 90°$ (combination of type III/IV and type V).

### 3.4. Uncut fibers

Typical uncut fibers observed on the lower and on the upper face of the workpiece are presented in Fig. 16 and Fig. 17 respectively, for each workpiece fiber orientation $Ψ$. On the lower face, we observe uncut fibers for all workpiece fiber orientations $Ψ$, while on the upper face, uncut fibers are noticed only for $Ψ = 135°$ as well as a few isolated fibers for $Ψ = 90°$. Furthermore, it is obvious that uncut fibers for $Ψ = 0°$ have a specific nature with long fibers, which are more or less bonded to the workpiece.

The average uncut fiber length $L_{avg}$ and the maximum uncut fiber length $L_{max}$, on the lower and on the upper face, are also given in Fig. 18
<figure><img src="image_15.png" /><figcaption>Fig. 13. typical machined surfaces for each workpiece fiber orientation $Ψ$.</figcaption></figure>

<figure><img src="image_16.png" /><figcaption>Fig. 14. $Sa$ (a) and $Sz$ (b) parameters at large and small scale for each workpiece fiber orientation $Ψ$.</figcaption></figure>

for each workpiece fiber orientation $Ψ$. Because of their specific nature, uncut fibers for $Ψ = 0°$ could not be adequately quantified and thus are not included in these results. It should be noted that $L_{max}$ values can be slightly higher than the radial engagement, due to the effect of the first edge trimming pass (Section 2.1.2).

Irrespective of the workpiece fiber orientation $Ψ$, it is observed that $L_{avg}$ and $L_{max}$ values are higher on the lower face than on the upper face. On the lower face, the highest $L_{avg}$ value is obtained for $Ψ = 90°$ with a large dispersion, and the lowest for $Ψ = 45°$, while $Ψ = 135°$ gives an intermediate result. On the contrary, $L_{max}$ on the lower face does not vary considerably, even if $Ψ = 135°$ gives a lower value than $Ψ = 45°$ and $Ψ = 90°$. On the upper face, the highest $L_{avg}$ value is reached for $Ψ = 135°$, whereas very small values, which are not significant, are obtained for the two other fiber orientations. A similar trend is obtained for $L_{max}$ on the upper face, with a large dispersion for $Ψ = 90°$, because of irregular uncut fibers occurrence in this case. These results are consistent with qualitative observation.

The systematic occurrence of uncut fibers on the lower face, as well as the larger $L_{avg}$ and $L_{max}$ values than on the upper face can be explained

by the downward axial forces produced by the cutting tool, as described in the cutting forces analysis in Section 3.1. While $L_{max}$ on the lower face is only slightly affected by fiber orientation, $L_{avg}$, which reflects the amount of uncut fibers, can be correlated with cutting forces and temperature. Indeed, $L_{avg}$ values on the lower face and average axial forces $Fz_{avg}$ globally move in the same direction (Fig. 8). Even with the burr tool multi-tooth geometry, axial forces are sufficient to delaminate the lower plies, which allows the fibers to bend without being cut. Moreover, $L_{avg}$ on the lower face follows the same trend as the maximum temperature rise $\Delta T_{max}$ (Fig. 10). Thus, high temperatures may contribute to the occurrence of uncut fibers on the lower face by favoring the lower plies delamination because of resin softening. The great dispersion observed for $L_{avg}$ on the lower face in the case of $Ψ = 90°$ is due to the fact that uncut fibers appear massively only after a certain machining distance for some of the tests. This is perhaps related to the progressive heating of the tool due to friction.

On the upper face, both qualitative and quantitative results show that the occurrence of uncut fibers tends to increase with the workpiece fiber orientation $Ψ$. Axial forces being now directed towards the inside of
<figure><img src="image_17.png" /><figcaption>Fig. 15. principle of interrupted cutting tests (a), SEM images of the rounded area corresponding to tool engagement for Ψ = 0° (b), Ψ = 45° (c), Ψ = 90° (d) and Ψ = 135° (e).</figcaption></figure>

<figure><img src="image_18.png" /><figcaption>Fig. 16. typical uncut fibers on the lower face.</figcaption></figure>

the workpiece, they cannot lead to delamination on the upper face and may have a limited impact on uncut fibers occurrence in this case. However, uncut fibers on the upper face can still bend somewhat in the workpiece plane due to tool motion. Assuming that the fibers bend only in the workpiece plane, they undergo great bending deformation for small Ψ angles, while this deformation is not sufficient to cause fiber fracture for higher Ψ angles. This is highlighted by the bending angle of

<figure><img src="image_19.png" /><figcaption>Fig. 17. typical uncut fibers on the upper face.</figcaption></figure>

the fibers at teeth exit (Fig. 19).

In terms of instantaneous fiber orientation θ, the occurrence of uncut fibers tends to decrease on the lower face and even disappear on the upper face for θ > 90°. This is consistent with the results of Hintze et al. [45] in the slot milling of UD-CFRP composites with zero helix angle tools.
<figure><img src="image_20.png" /><figcaption>Fig. 18. average uncut fiber length $L_{avg}$ (a) and maximum uncut fiber length $L_{max}$ (b) for each workpiece fiber orientation $\Psi$.</figcaption></figure>

<figure><img src="image_21.png" /><figcaption>Fig. 19. bending angle of the fibers at teeth exit for $\Psi = 45^{\circ}$ (a), $\Psi = 90^{\circ}$ (b) and $\Psi = 135^{\circ}$ (c).</figcaption></figure>

## 4. Conclusions

In this study, we conducted edge trimming of UD-GFRP composites with diamond-coated burr tools for different workpiece fiber orientations $\Psi$ using constant cutting parameters. Cutting forces and machined surface temperature were recorded during machining, while surface quality and surface plies defects were characterized after each test. The following conclusions can be drawn:

- Fiber orientation in the workpiece has a great impact on cutting forces, machined surface temperature, surface quality and the occurrence of uncut fibers on surface plies.

- The effect of workpiece fiber orientation $\Psi$ on cutting forces depends on the measurement direction. Overall, the resultant of average forces is the lowest for $\Psi = 45^{\circ}$ and the highest for $\Psi = 135^{\circ}$. Moreover, average cutting forces are systematically higher in the direction perpendicular to the feed than in the feed direction, particularly for $\Psi = 0^{\circ}$ and $\Psi = 135^{\circ}$.

- The maximum temperature rise at the center of the machined surface is the lowest for $\Psi = 45^{\circ}$ and the highest for $\Psi = 90^{\circ}$. It tends to be greater for fiber orientations that induce high resultant forces. The cooling rate at the same location on the surface follows almost the same trend. It is justly explained by natural convection and heat conduction in the fibers.

- Surface irregularities are relatively uniform and are affected in the same way by fiber orientation in the workpiece from the fiber scale to the macro-scale. $Sa$ and $Sz$ parameters are the highest for $\Psi = 45^{\circ}$, while they are far lower for the other orientations, the minimum values being obtained for $\Psi = 135^{\circ}$. Great surface irregularities are not associated with high resultant forces, while surface irregularities, and the maximum rise in temperature, move globally in opposite directions.

- Uncut fibers systematically occur on the lower face of the workpiece due to the downward axial forces generated by the cutting tool. Uncut fibers also occur on the upper face of the workpiece, mainly for $\Psi = 135^{\circ}$. Indeed, considering their bending deformation in the workpiece plane, it may not be sufficient to cause breakage.

Moreover, the occurrence of uncut fibers on the lower face may be facilitated by high temperatures during cutting.

- In the thickness of the workpiece, the effect of fiber orientation on cutting forces in the workpiece plane, machined surface temperature and surface quality is very similar to previous studies in orthogonal cutting. It follows that the main cutting mechanisms with the complex burr tool geometry used in this work may be very similar in overall terms to those identified in orthogonal cutting. This analysis is supported by the qualitative observation of the machined surface, in the rounded area corresponding to tool engagement. On surface layers, the results for uncut fibers are also consistent with previous literature.

The results obtained in this work can be used to optimize machining operations with burr tools on MD-FRP composites, making it possible to identify the fiber orientations to favor depending on the criteria to be optimized. To supplement this study, additional experiments could be performed with a modified burr tool consisting of one single tooth to obtain instantaneous data. This would enable an in-depth analysis of the interaction between the tooth and the workpiece. In addition, more precise temperature mapping and further investigation into thermal phenomena are required to improve our understanding of their impact on surface quality and external plies defects. The tool temperature could also be measured by adding an IR sensor at the spindle level. Finally, the influence of cutting parameters needs to be studied, in particular the radial engagement, which affects the instantaneous fiber orientation range during cutting.

## CRediT authorship contribution statement

**Schrab Benoît:** Writing – original draft, Visualization, Validation, Software, Methodology, Investigation, Conceptualization.
**Collaine Anne:** Writing – review & editing, Supervision, Project administration, Methodology, Investigation, Conceptualization.
**Freyburger Jean-Marie:** Writing – review & editing, Methodology, Investigation, Conceptualization.
**Tourlonias Michel:** Writing – review & editing, Supervision, Project administration, Methodology, Investigation,
Funding acquisition, Conceptualization.

Declaration of Competing Interest

The authors declare that they have no known competing financial interests or personal relationships that could have appeared to influence the work reported in this paper.

Acknowledgements

The authors would like to thank the Hexion company for providing the raw material necessary for this study, the Laboratoire de Photochimie et d'Ingénierie Macromoléculaires (LPIM) for its contribution to the composites characterization, and the Institut de Recherche en Informatique, Mathématiques, Automatique et Signal (IRIMAS) for its contribution to the uncut fibers characterization device development. The authors also give special thanks to Philippe Blondel, Christian Pidancier and Jean-Jacques Bockstaller from LPMT and the Ecole Nationale Supérieure d'Ingénieurs Sud Alsace (ENSISA).

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Benoît Schrab is a mechanical engineer. He graduated from the Ecole Nationale Supérieure d'Ingénieurs Sud-Alsace (Mulhouse, France) in 2020. Benoît has been working as a PhD student at the Laboratoire de Physique et Mécanique Textiles (Université de
Haute-Alsace, Mulhouse, France) since 2021. His research focuses on the optimization of milling operations performed on fiber-reinforced polymers after curing.

**Anne Collaine** is an assistant professor in mechanical engineering at the Université de Haute-Alsace (France). She teaches on bachelor and master courses at ENSISA (Ecole Nationale Supérieure d'Ingénieurs Sud-Alsace) in the field of machining. Anne's research addresses the machining of composite materials on milling in particular. Her latest work focuses on the milling of flax fiber-reinforced composites, with the aim of optimizing their cutting properties.

**Jean-Marie Freyburger** is an assistant professor at the Université de Haute-Alsace (France). He teaches on bachelor and master courses at ENSISA (Ecole Nationale Supérieure d'Ingénieurs Sud-Alsace) in the field of electronics and signal processing. His research focuses on the machining of composite materials, particularly milling.

**Michel Tourlonias** is a textile engineer (2002 - Mulhouse). He obtained his doctorate at the Université de Haute-Alsace (France) in 2005. Since 2021, Michel has been a professor at ENSISA (Ecole Nationale Supérieure d'Ingénieurs Sud-Alsace). His research focuses on textile materials at the Laboratoire de Physique et Mécanique Textiles in tribology (at the fiber, yarn and textile-surface scale) and the machining of composite materials. This work is carried out using an experimental approach.