the strain distribution was non uniform. Localized high strain-concentrated areas appeared at a
fraction of the ultimate load (Figure 5a), probably due to the interaction between bundle ends and
transverse fibers [14]. As the load increases, fractures occur locally, leading to a redistribution of
strain (Figure 5b) without noticeably affecting the stress-strain response of the sample. When strain
concentration appears at the interface with the SMC and the endless fiber reinforcement, strain a
wider area is affected (Figure 5c). With increasing load, damage on continuous fibers increases
(Figure 5d) dealing with the failure of the continuous fiber reinforcement.

<figure><img src="image_508.png" /><figcaption>Figure 5: Strain distribution on the surface of a UD2/100%/TW1/100%/SMC2/80% samples at different loads. a) SMC side at 70% of ultimate load b) SMC side at 90% of ultimate load c) UD side at 70% of ultimate load d) SMC side at 90% of ultimate load</figcaption></figure>

Conclusions

A viable technological solution for the manufacturing of flat components with hybrid fiber
architecture laminates in a single step compression molding operation was proposed in this work
by integrating a layer of fabric to mitigate the waviness induced during processing. Specimens
were extracted on compression molded plates and successfully tested in tensile and flexure. Main
findings are the follow:

* Hybridization led to a 137% increase in tensile strength and up to 147% rise in flexural modulus compared to the standard SMC baseline.

* The standard deviation on tensile strengths and elastic modulus of the hybrid laminate is consistently comparable to that of a long-fiber system.

Future works will investigate the reliability of numerical simulation tools in predicting the
quality of compression molded parts with hybrid fiber architectures. Simulations based on state of
the art strategies will be used to predict wrinkles and other forming effects which can occur in the
continuous reinforcement during molding of complex parts [15]. Interactions between SMC flow
and endless fiber reinforcement will be numerically studied by means of direct fiber simulation
[16]. The process simulation results will be used for the development of a suitable simulation
model for virtual test of individual materials as well as hybrid-architectures.

References

[1] Zulueta K, Arrillaga A, Stelzer PS, Martinez I. An evaluation of the curing characteristics of thermosetting epoxy carbon fiber sheet molding compounds and validation through computer simulations and molding trials. J Appl Polym Sci 2021;138. https://doi.org/10.1002/app.50826
the strain distribution was non uniform. Localized high strain-concentrated areas appeared at a fraction of the ultimate load (Figure 5a), probably due to the interaction between bundle ends and transverse fibers [14]. As the load increases, fractures occur locally, leading to a redistribution of strain (Figure 5b) without noticeably affecting the stress-strain response of the sample. When strain concentration appears at the interface with the SMC and the endless fiber reinforcement, strain a wider area is affected (Figure 5c). With increasing load, damage on continuous fibers increases (Figure 5d) dealing with the failure of the continuous fiber reinforcement.

<figure><img src="image_519.png" /><figcaption>Figure 5: Strain distribution on the surface of a UD2/100%/TW1/100%/SMC2/80% samples at different loads. a) SMC side at 70% of ultimate load b) SMC side at 90% of ultimate load c) UD side at 70% of ultimate load d) SMC side at 90% of ultimate load</figcaption></figure>

## Conclusions

A viable technological solution for the manufacturing of flat components with hybrid fiber architecture laminates in a single step compression molding operation was proposed in this work by integrating a layer of fabric to mitigate the waviness induced during processing. Specimens were extracted on compression molded plates and successfully tested in tensile and flexure. Main findings are the follow:

* Hybridization led to a 137% increase in tensile strength and up to 147% rise in flexural modulus compared to the standard SMC baseline.

* The standard deviation on tensile strengths and elastic modulus of the hybrid laminate is consistently comparable to that of a long-fiber system.

Future works will investigate the reliability of numerical simulation tools in predicting the quality of compression molded parts with hybrid fiber architectures. Simulations based on state of the art strategies will be used to predict wrinkles and other forming effects which can occur in the continuous reinforcement during molding of complex parts [15]. Interactions between SMC flow and endless fiber reinforcement will be numerically studied by means of direct fiber simulation [16]. The process simulation results will be used for the development of a suitable simulation model for virtual test of individual materials as well as hybrid-architectures.

## References

[1] Zulueta K, Arrillaga A, Stelzer PS, Martinez I. An evaluation of the curing characteristics of thermosetting epoxy carbon fiber sheet molding compounds and validation through computer simulations and molding trials. J Appl Polym Sci 2021;138. https://doi.org/10.1002/app.50826
[2] Feraboli P, Peitso E, Cleveland T, Stickler PB, Halpin JC, Feraboli Paolo, Peitso Elof, Cleveland Tyler, Stickler HJC. Notched behavior of prepreg-based discontinuous carbon fiber/epoxy systems. Compos Part A Appl Sci Manuf 2009;40:289-99.
https://doi.org/10.1016/j.compositesa.2008.12.012

[3] Laribi MA, TieBi R, Tamboura S, Shirinbayan M, Tcharkhtchi A, Dali H Ben, et al. Sheet Molding Compound Automotive Component Reliability Using a Micromechanical Damage Approach. Appl Compos Mater 2020;27:693-715. https://doi.org/10.1007/s10443-020-09831-5

[4] Falaschetti MP, Rondina F, Maccaferri E, Mazzocchetti L, Donati L, Zucchelli A, et al. Improving the crashworthiness of CFRP structures by rubbery nanofibrous interlayers. Compos Struct 2023;311. https://doi.org/10.1016/j.compstruct.2023.116845

[5] Rondina F, Falaschetti MP, Zavatta N, Donati L. Numerical simulation of the compression crushing energy of carbon fiber-epoxy woven composite structures. Compos Struct 2023;303.
https://doi.org/10.1016/j.compstruct.2022.116300

[6] Mallick PK. Effect of fiber misorientation on the tensile strength of compression molded continuous fiber composites. Polym Compos 1986;7:14-8. https://doi.org/10.1002/pc.750070104.

[7] DG Taggart, RB Pipes, RA Blake, JW Gillespie Jr, R Prabhakaran JW. Properties of SMC composites. Cent Compos Mater Univ Delaware 1979.

[8] Trauth A, Weidenmann KA. Continuous-discontinuous sheet moulding compounds – Effect of hybridisation on mechanical material properties. Compos Struct 2018.
https://doi.org/10.1016/j.compstruct.2018.05.048

[9] Trauth A, Weidenmann KA, Altenhof W. Puncture properties of a hybrid continuous-discontinuous sheet moulding compound for structural applications. Compos Part B Eng 2019;158:46-54. https://doi.org/10.1016/j.compositesb.2018.09.035

[10] Raimondi L, Brugo TM, Zucchelli A. Fiber misalignment analysis in PCM-UD composite materials by Full Field Nodal Method. Compos Part C Open Access 2021;5.
https://doi.org/10.1016/j.jcomc.2021.100151

[11] Corbridge DM, Harper LT, De Focatiis DSA, Warrior NA. Compression moulding of composites with hybrid fibre architectures. Compos Part A Appl Sci Manuf 2017.
https://doi.org/10.1016/j.compositesa.2016.12.018

[12] Feraboli P, Peitso E, Cleveland T, Stickler PB. Modulus measurement for prepreg-based discontinuous carbon fiber/epoxy systems. J Compos Mater 2009;43:1947-65.
https://doi.org/10.1177/0021998309343028

[13] Visweswaraiah SB, 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;37:548-80. https://doi.org/10.1177/0731684418754360

[14] Johanson K, Harper LT, Johnson MS, Warrior NA. Heterogeneity of discontinuous carbon fibre composites: Damage initiation captured by Digital Image Correlation. Compos Part A Appl Sci Manuf 2015;68:304-12. https://doi.org/10.1016/j.compositesa.2014.10.014

[15] Boisse P, Akkerman R, Carlone P, Kärger L, Lomov S V., Sherwood JA. Advances in composite forming through 25 years of ESAFORM. Int J Mater Form 2022;15.
https://doi.org/10.1007/s12289-022-01682-8

[16] Qian CC, Deshpande A, Jesri M, Groves R, Reynolds N, Kendall K. A comprehensive assessment of commercial process simulation software for compression moulding of sheet moulding compound. ESAFORM 2021 - 24th Int. Conf. Mater. Form., 2021.
https://doi.org/10.25518/esaform21.2771
[2] Feraboli P, Peitso E, Cleveland T, Stickler PB, Halpin JC, Feraboli Paolo, Peitso Elof, Cleveland Tyler, Stickler HJC. Notched behavior of prepreg-based discontinuous carbon fiber/epoxy systems. Compos Part A Appl Sci Manuf 2009;40:289-99.
https://doi.org/10.1016/j.compositesa.2008.12.012

[3] Laribi MA, TieBi R, Tamboura S, Shirinbayan M, Tcharkhtchi A, Dali H Ben, et al. Sheet Molding Compound Automotive Component Reliability Using a Micromechanical Damage Approach. Appl Compos Mater 2020;27:693-715. https://doi.org/10.1007/s10443-020-09831-5

[4] Falaschetti MP, Rondina F, Maccaferri E, Mazzocchetti L, Donati L, Zucchelli A, et al. Improving the crashworthiness of CFRP structures by rubbery nanofibrous interlayers. Compos Struct 2023;311. https://doi.org/10.1016/j.compstruct.2023.116845

[5] Rondina F, Falaschetti MP, Zavatta N, Donati L. Numerical simulation of the compression crushing energy of carbon fiber-epoxy woven composite structures. Compos Struct 2023;303.
https://doi.org/10.1016/j.compstruct.2022.116300

[6] Mallick PK. Effect of fiber misorientation on the tensile strength of compression molded continuous fiber composites. Polym Compos 1986;7:14-8. https://doi.org/10.1002/pc.750070104.

[7] DG Taggart, RB Pipes, RA Blake, JW Gillespie Jr, R Prabhakaran JW. Properties of SMC composites. Cent Compos Mater Univ Delaware 1979.

[8] Trauth A, Weidenmann KA. Continuous-discontinuous sheet moulding compounds – Effect of hybridisation on mechanical material properties. Compos Struct 2018.
https://doi.org/10.1016/j.compstruct.2018.05.048

[9] Trauth A, Weidenmann KA, Altenhof W. Puncture properties of a hybrid continuous-discontinuous sheet moulding compound for structural applications. Compos Part B Eng 2019;158:46-54. https://doi.org/10.1016/j.compositesb.2018.09.035

[10] Raimondi L, Brugo TM, Zucchelli A. Fiber misalignment analysis in PCM-UD composite materials by Full Field Nodal Method. Compos Part C Open Access 2021;5.
https://doi.org/10.1016/j.jcomc.2021.100151

[11] Corbridge DM, Harper LT, De Focatiis DSA, Warrior NA. Compression moulding of composites with hybrid fibre architectures. Compos Part A Appl Sci Manuf 2017.
https://doi.org/10.1016/j.compositesa.2016.12.018

[12] Feraboli P, Peitso E, Cleveland T, Stickler PB. Modulus measurement for prepreg-based discontinuous carbon fiber/epoxy systems. J Compos Mater 2009;43:1947-65.
https://doi.org/10.1177/0021998309343028

[13] Visweswaraiah SB, 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;37:548-80. https://doi.org/10.1177/0731684418754360

[14] Johanson K, Harper LT, Johnson MS, Warrior NA. Heterogeneity of discontinuous carbon fibre composites: Damage initiation captured by Digital Image Correlation. Compos Part A Appl Sci Manuf 2015;68:304-12. https://doi.org/10.1016/j.compositesa.2014.10.014

[15] Boisse P, Akkerman R, Carlone P, Kärger L, Lomov S V., Sherwood JA. Advances in composite forming through 25 years of ESAFORM. Int J Mater Form 2022;15.
https://doi.org/10.1007/s12289-022-01682-8

[16] Qian CC, Deshpande A, Jesri M, Groves R, Reynolds N, Kendall K. A comprehensive assessment of commercial process simulation software for compression moulding of sheet moulding compound. ESAFORM 2021 - 24th Int. Conf. Mater. Form., 2021.
https://doi.org/10.25518/esaform21.2771
# Recent advances in development, characterization and joining of new sustainable materials

DOINA BANEA Mariana¹,ᵃ*,
FERNANDES MEDEIROS DE QUEIROZ Henrique²,ᵇ,
DE SOUZA E SILVA NETO Jorge²,ᶜ and
KIOSHI KAWASAKI CAVALCANTI Daniel²,ᵈ

¹CICECO - Aveiro Institute of Materials, Department of Materials and Ceramic Engineering, University of Aveiro, 3810-193, Aveiro, Portugal

²CEFET/RJ, Federal Center of Technological Education in Rio de Janeiro, R. Gen. Canabarro, 485 - Maracanã, Rio de Janeiro - RJ, 20271-204, Brazil

ᵃmdbanea@gmail.com, ᵇhenriquefmq@gmail.com, ᶜjorgesouzanetto@gmail.com, ᵈdanielkkc@gmail.com

**Keywords:** Composite Materials, Hybrid Composites, Natural Fibers, Synthetic Fibers, Sustainable Materials, Mechanical Characterization, Adhesive Bonded Joints

**Abstract.** Nowadays, in order to increase the transport efficiency and reduce fuel consumption and emissions of contaminants, a reduction in weight associated with improved safety performance of the materials in use must be achieved. On the other hand, environmental concern has generated interest in research about new materials aligned with the principles of sustainability. Thus, this need for better performing and ecological structures has resulted in the development of a new variety of materials. Among these materials, currently, are the composites produced from resources of renewable sources. Natural fiber composites have recently attracted a great deal of attention by the industry due to their many attractive benefits (e.g., high strength-to-weight ratio, sustainable characteristics and low cost). Welding is highly impractical to use in these situations (i.e., thermoset polymer composites) and rivets or screws exhibit stress concentrations and offer a low fatigue resistance. Adhesive bonding is usually the preferred method since it allows for greater flexibility in design and is more efficient in mechanical and energy aspects. In this work, recent advances in development, characterization and joining of new sustainable materials, critical challenges and future perspectives are presented. The application of sustainable green composites may further increase in many structural and non-structural applications if their joining behavior is well-known and established.

**Introduction**

Global initiatives and agreements, such as the United Nations Sustainable Development Goals (SDGs) and the Paris Agreement, have highlighted the urgency of adopting sustainable practices in all aspects of human activity. Industries are increasingly recognizing that sustainable materials not only align with ethical and environmental objectives but also offer economic advantages, including cost savings, enhanced brand reputation, and access to emerging markets with environmentally conscious consumers. As a result, researchers, scientists, and engineers are actively engaged in developing innovative sustainable materials that exhibit improved performance characteristics while minimizing their environmental footprint.

Sustainable materials such as natural fiber composites (NFRCs) can be found in applications in many industrial sectors, such as: automotive, construction, marine industry, and sports, among others [1]. Moreover, the most promising applications of NFRCs are in the automotive industry. In general, they are used in non-structural car body parts, such as: door panels, package trays, hat

<figure><img src="image_509.png" /></figure>
# Recent advances in development, characterization and joining of new sustainable materials

DOINA BANEA Mariana¹,ª*,
FERNANDES MEDEIROS DE QUEIROZ Henrique²,ᵇ,
DE SOUZA E SILVA NETO Jorge²,ᶜ and
KIOSHI KAWASAKI CAVALCANTI Daniel²,ᵈ

¹CICECO - Aveiro Institute of Materials, Department of Materials and Ceramic Engineering, University of Aveiro, 3810-193, Aveiro, Portugal

²CEFET/RJ, Federal Center of Technological Education in Rio de Janeiro, R. Gen. Canabarro, 485 - Maracanã, Rio de Janeiro - RJ, 20271-204, Brazil

ªmdbanea@gmail.com, ᵇhenriquefmq@gmail.com, ᶜjorgesouzanetto@gmail.com, ᵈdanielkkc@gmail.com

**Keywords:** Composite Materials, Hybrid Composites, Natural Fibers, Synthetic Fibers, Sustainable Materials, Mechanical Characterization, Adhesive Bonded Joints

**Abstract.** Nowadays, in order to increase the transport efficiency and reduce fuel consumption and emissions of contaminants, a reduction in weight associated with improved safety performance of the materials in use must be achieved. On the other hand, environmental concern has generated interest in research about new materials aligned with the principles of sustainability. Thus, this need for better performing and ecological structures has resulted in the development of a new variety of materials. Among these materials, currently, are the composites produced from resources of renewable sources. Natural fiber composites have recently attracted a great deal of attention by the industry due to their many attractive benefits (e.g., high strength-to-weight ratio, sustainable characteristics and low cost). Welding is highly impractical to use in these situations (i.e., thermoset polymer composites) and rivets or screws exhibit stress concentrations and offer a low fatigue resistance. Adhesive bonding is usually the preferred method since it allows for greater flexibility in design and is more efficient in mechanical and energy aspects. In this work, recent advances in development, characterization and joining of new sustainable materials, critical challenges and future perspectives are presented. The application of sustainable green composites may further increase in many structural and non-structural applications if their joining behavior is well-known and established.

**Introduction**

Global initiatives and agreements, such as the United Nations Sustainable Development Goals (SDGs) and the Paris Agreement, have highlighted the urgency of adopting sustainable practices in all aspects of human activity. Industries are increasingly recognizing that sustainable materials not only align with ethical and environmental objectives but also offer economic advantages, including cost savings, enhanced brand reputation, and access to emerging markets with environmentally conscious consumers. As a result, researchers, scientists, and engineers are actively engaged in developing innovative sustainable materials that exhibit improved performance characteristics while minimizing their environmental footprint.

Sustainable materials such as natural fiber composites (NFRCs) can be found in applications in many industrial sectors, such as: automotive, construction, marine industry, and sports, among others [1]. Moreover, the most promising applications of NFRCs are in the automotive industry. In general, they are used in non-structural car body parts, such as: door panels, package trays, hat

<figure><img src="image_520.png" /></figure>
racks, instrument panels, internal engine covers, sun visors, boot liners and oil air filters. However, they are evolving to be used in more structurally demanding parts such as seat backs, exterior underfloor paneling as well as anti-roll bars [2-4]. They can also be used in construction, and building applications like bricks, door panels, roofing sheets, furniture panels, etc.

The main natural fibers studied and used in the industry (e.g., jute, sisal, kenaf, and flax) are well established on the global market with a well-defined production line. However, new promising types of natural fibers are being discovered and studied [1, 5]. Both thermoplastic and thermoset polymers are used as matrices in natural fiber reinforced polymer composites [6]. There is an increased interest in the scientific community in using bio-based polymers in composites as combining these matrices with natural fibers produces “green composites” or “bio-composites”. An emerging trend is also found in the use of thermoplastic matrices of renewable sources, such as PLA, as a matrix material for NFRCs [7, 8]. This is crucial due to the fact that the classical thermoset matrix materials such as epoxy are dependent on fossil fuels.

As the global emphasis on sustainability grows, the demand for more sustainable (eco-friendly) materials with enhanced properties has intensified. This paper presents an overview of cutting-edge research and innovations in the development of more sustainable materials (NFRCs), their characterization methods, and the adhesive joining of these materials.

## Development and Characterization of New Sustainable Materials

Natural-fiber-reinforced composites (NFRCs) vary greatly in their mechanical properties. Mechanical properties (e.g., tensile, flexural, and impact) are highly dependent on different factors such as fiber and matrix type, interfacial bonding between fiber and matrix, fiber dispersion orientation and processing, among others. By increasing their mechanical performance, the capabilities and applications of natural-fiber-reinforced composites will be expanded.

The most common mechanical tests used to characterize sustainable composites are the tensile, flexural and impact tests. Generally, the tensile strength of composites increases with fiber content up to a maximum value, then it tends to decrease. This is due to the higher mechanical properties of the fibers when compared to the matrices [2, 9]. Furthermore, it is well known that natural fibers contain moisture and when introduced into a hydrophobic matrix, the higher water absorption tendency of the natural fibers will lead to lower interfacial fiber/matrix adhesion and consequently lower mechanical properties of the NFRCs. One possibility to tackle this issue is the use of surface treatments [10]. The flexural properties of NFRCs are sensitive to a number of parameters, such as layup, ply angles, material type and stacking sequence [11]. This is due to the fact that the stresses increase from the neutral line and reach their maximum compressive and tensile stresses on the top and bottom of the test specimen, respectively. Impact resistance is a measure used to characterize the efficiency of the material during a rapid load event and is intrinsically linked with its toughness. The energy absorption may be derived via a penetration impact, residual properties after impact, and analysis of the damaged area after a non-penetration impact. Furthermore, the stacking sequence of the composite is a significant parameter when it comes to absorbed energy as it was observed that higher energy absorbing fiber layers on the outside of the composite enhances the overall impact resistance [12]. On the other hand, the interfacial quality of the fiber/matrix is highly significant as a low-quality interface will poorly transfer the loads and decrease absorbed energy. This issue can be improved by the hybridization with synthetic fibers.

## Strategies for Improving the Performance of New Sustainable Materials

The main methods used to increase the performance of natural fiber-reinforced composites, such as: fiber treatment and modification, the use of filler materials (either on the fiber surface or into the matrix as the second reinforcing phase) and fiber hybridization) are briefly presented in this section.
racks, instrument panels, internal engine covers, sun visors, boot liners and oil air filters. However, they are evolving to be used in more structurally demanding parts such as seat backs, exterior underfloor paneling as well as anti-roll bars [2-4]. They can also be used in construction, and building applications like bricks, door panels, roofing sheets, furniture panels, etc.

The main natural fibers studied and used in the industry (e.g., jute, sisal, kenaf, and flax) are well established on the global market with a well-defined production line. However, new promising types of natural fibers are being discovered and studied [1, 5]. Both thermoplastic and thermoset polymers are used as matrices in natural fiber reinforced polymer composites [6]. There is an increased interest in the scientific community in using bio-based polymers in composites as combining these matrices with natural fibers produces “green composites” or “bio-composites”. An emerging trend is also found in the use of thermoplastic matrices of renewable sources, such as PLA, as a matrix material for NFRCs [7, 8]. This is crucial due to the fact that the classical thermoset matrix materials such as epoxy are dependent on fossil fuels.

As the global emphasis on sustainability grows, the demand for more sustainable (eco-friendly) materials with enhanced properties has intensified. This paper presents an overview of cutting-edge research and innovations in the development of more sustainable materials (NFRCs), their characterization methods, and the adhesive joining of these materials.

## Development and Characterization of New Sustainable Materials

Natural-fiber-reinforced composites (NFRCs) vary greatly in their mechanical properties. Mechanical properties (e.g., tensile, flexural, and impact) are highly dependent on different factors such as fiber and matrix type, interfacial bonding between fiber and matrix, fiber dispersion orientation and processing, among others. By increasing their mechanical performance, the capabilities and applications of natural-fiber-reinforced composites will be expanded.

The most common mechanical tests used to characterize sustainable composites are the tensile, flexural and impact tests. Generally, the tensile strength of composites increases with fiber content up to a maximum value, then it tends to decrease. This is due to the higher mechanical properties of the fibers when compared to the matrices [2, 9]. Furthermore, it is well known that natural fibers contain moisture and when introduced into a hydrophobic matrix, the higher water absorption tendency of the natural fibers will lead to lower interfacial fiber/matrix adhesion and consequently lower mechanical properties of the NFRCs. One possibility to tackle this issue is the use of surface treatments [10]. The flexural properties of NFRCs are sensitive to a number of parameters, such as layup, ply angles, material type and stacking sequence [11]. This is due to the fact that the stresses increase from the neutral line and reach their maximum compressive and tensile stresses on the top and bottom of the test specimen, respectively. Impact resistance is a measure used to characterize the efficiency of the material during a rapid load event and is intrinsically linked with its toughness. The energy absorption may be derived via a penetration impact, residual properties after impact, and analysis of the damaged area after a non-penetration impact. Furthermore, the stacking sequence of the composite is a significant parameter when it comes to absorbed energy as it was observed that higher energy absorbing fiber layers on the outside of the composite enhances the overall impact resistance [12]. On the other hand, the interfacial quality of the fiber/matrix is highly significant as a low-quality interface will poorly transfer the loads and decrease absorbed energy. This issue can be improved by the hybridization with synthetic fibers.

## Strategies for Improving the Performance of New Sustainable Materials

The main methods used to increase the performance of natural fiber-reinforced composites, such as: fiber treatment and modification, the use of filler materials (either on the fiber surface or into the matrix as the second reinforcing phase) and fiber hybridization) are briefly presented in this section.
## Fiber Treatment and Modification

Surface treatment of natural fibers is usually performed to enhance their mechanical and thermal properties, before using them to manufacture a composite material. The fiber modification techniques provide improved fiber-matrix interfacial adhesion, improved fiber roughness and wettability and depends on the particular fiber/matrix used and the composite application [13]. It is fundamental to understand the interfacial properties and bonding mechanisms of fiber-matrix and this requires significant research efforts in order to maximize natural fiber composites applications [10, 13, 14].

## Use of filler materials

Another method explored in the literature to improve the properties of natural fiber-reinforced composites is the use of filler materials (either on the fiber surface or into the matrix as the second reinforcing phase) [15-17]. Due to their high specific surface area, good chemical interaction with the hydrophobic matrix and crack arresting capabilities, a synergistic relationship is observed with NFRCs [18]. However, a careful determination of the fabrication and matrix/filler homogenization technique must be done in order to avoid filler agglomeration, which may lead to detrimental effects. The most commonly used synthetic nanofillers are: MWCNTs, metal oxides and titanium dioxide.

Neto et al. [17] investigated the effect of multi walled carbon nanotubes (MWCNTs) in the UV and water spray ageing process of natural fiber and hybrid composites. Jute and glass fiber plain weave bidirectional fabrics were used along with an epoxy resin matrix. The compression molding technique was used to fabricate the composite plates. The filler weight fraction was 0.6%. The mechanical properties were analyzed via tensile and flexural tests according to ASTM standards. The samples were aged at 500h and 1000h of exposure, respectively. They found that for the pure natural fiber composites an improvement in the mechanical properties was observed. However, for the hybrid composites, a slight decrease in tensile and flexural strength was observed. More recently, Neto et al. [19] studied the effect of the addition of the MWCNTs on the mechanical and thermal properties of curauá fiber reinforced composites. The MWCNTs were added either to the fiber surface (see Fig. 1) or into the resin matrix as the second reinforcing phase. They found that the incorporation of MWCNTs on the curauá fibers showed positive effects (an increase in properties was found for the MWCNT-modified fibers and their composites).

<figure><img src="image_510.png" /><figcaption>Figure 1. Schematic of the coating process of curauá fibers with MWCNT [19].</figcaption></figure>

## Fiber hybridization

One method to increase the mechanical performance of NFRCs to extend their applications is the fiber hybridization [1, 15, 20-22]. Generally, natural fiber reinforced hybrid composites are produced by hybridizing natural fibers with another either natural or synthetic fiber with superior properties (e.g., higher mechanical strength, chemical stability, nontoxicity, resistance to high temperatures, and thermal or acoustic insulation).

Fiber hybridization: use of intralaminar hybridization [21-23], use of interlaminar hybridization [24], use of 3D fiber-reinforcement architecture and use of multi-scale hybridization [15].

For example, for the case of intralaminar hybridization, Pereira et al. [23] used ramie, sisal and curauá fibers as natural fiber reinforcements for the hybrid composites based on sisal. They concluded that the mechanical properties are improved by the hybridization of sisal-based
## Fiber Treatment and Modification

Surface treatment of natural fibers is usually performed to enhance their mechanical and thermal properties, before using them to manufacture a composite material. The fiber modification techniques provide improved fiber-matrix interfacial adhesion, improved fiber roughness and wettability and depends on the particular fiber/matrix used and the composite application [13]. It is fundamental to understand the interfacial properties and bonding mechanisms of fiber-matrix and this requires significant research efforts in order to maximize natural fiber composites applications [10, 13, 14].

## Use of filler materials

Another method explored in the literature to improve the properties of natural fiber-reinforced composites is the use of filler materials (either on the fiber surface or into the matrix as the second reinforcing phase) [15-17]. Due to their high specific surface area, good chemical interaction with the hydrophobic matrix and crack arresting capabilities, a synergistic relationship is observed with NFRCs [18]. However, a careful determination of the fabrication and matrix/filler homogenization technique must be done in order to avoid filler agglomeration, which may lead to detrimental effects. The most commonly used synthetic nanofillers are: MWCNTs, metal oxides and titanium dioxide.

Neto et al. [17] investigated the effect of multi walled carbon nanotubes (MWCNTs) in the UV and water spray ageing process of natural fiber and hybrid composites. Jute and glass fiber plain weave bidirectional fabrics were used along with an epoxy resin matrix. The compression molding technique was used to fabricate the composite plates. The filler weight fraction was 0.6%. The mechanical properties were analyzed via tensile and flexural tests according to ASTM standards. The samples were aged at 500h and 1000h of exposure, respectively. They found that for the pure natural fiber composites an improvement in the mechanical properties was observed. However, for the hybrid composites, a slight decrease in tensile and flexural strength was observed. More recently, Neto et al. [19] studied the effect of the addition of the MWCNTs on the mechanical and thermal properties of curauá fiber reinforced composites. The MWCNTs were added either to the fiber surface (see Fig. 1) or into the resin matrix as the second reinforcing phase. They found that the incorporation of MWCNTs on the curauá fibers showed positive effects (an increase in properties was found for the MWCNT-modified fibers and their composites).

<figure><img src="image_521.png" /><figcaption>Figure 1. Schematic of the coating process of curauá fibers with MWCNT [19].</figcaption></figure>

## Fiber hybridization

One method to increase the mechanical performance of NFRCs to extend their applications is the fiber hybridization [1, 15, 20-22]. Generally, natural fiber reinforced hybrid composites are produced by hybridizing natural fibers with another either natural or synthetic fiber with superior properties (e.g., higher mechanical strength, chemical stability, nontoxicity, resistance to high temperatures, and thermal or acoustic insulation).

Fiber hybridization: use of intralaminar hybridization [21-23], use of interlaminar hybridization [24], use of 3D fiber-reinforcement architecture and use of multi-scale hybridization [15].

For example, for the case of intralaminar hybridization, Pereira et al. [23] used ramie, sisal and curauá fibers as natural fiber reinforcements for the hybrid composites based on sisal. They concluded that the mechanical properties are improved by the hybridization of sisal-based
composites. On the other hand, they state that the hybridization did not significantly affect the thermal stability of the composites studied.

Interlaminarly hybridized composites of natural/synthetic fiber have been shown in the literature to present a non-linear enhancement of the material properties as the weight fraction of synthetic fibers increases. For example, Queiroz et al. [25], demonstrated that for hybrid composites reinforced with a symmetric synthetic glass envelope around a jute fabric core, the in-plane properties increased significantly as the number of glass layers increased from 2 to 3 (i.e., an approx. 35% enhancement in tensile strength). This is due to the much higher density of synthetic fibers present in a woven fabric compared to the in-plane intralaminar hybridization. On the other hand, Queiroz et al. [26] recently demonstrated that the effect of 3D reinforcement of hybrid natural composites (i.e., z-binder fibers stitched in the transverse direction through intralaminar 2D fabrics) is not significant in the in-plane properties. The highest variation observed in the composite tensile strength (when the 2D case is compared to the 3D) was of approx. 15%, while for most of the studied cases it was below 10%. Therefore, differently from the interlaminar reinforcement, significant variations will not be observed in the in-plane composite properties but rather in the transverse properties (strength and toughness).

## 3D Printing with Natural Fibers

Another area currently developing fast is the application of natural fiber as reinforcements in composites produced by additive manufacturing (AM or 3D printing) [27-32]. This technology allows for the fabrication of complex geometries without the need for expensive tooling and molds. It also allows for precise control over material distribution, opening up new possibilities for complex geometries and customizable structures [33, 34].

The primary materials utilized in Fused Deposition Modeling (FDM) are acrylonitrile butadiene styrene (ABS) and polylactic acid (PLA). However, in their unmodified state, these materials exhibit relatively modest mechanical properties, characterized by typical bulk strengths ranging from 30 to 100 MPa and elastic modulus within the 1.3–3.6 GPa range [27]. Consequently, parts produced through FDM using these materials are not suitable for high-performance applications. To address this limitation, diverse methodologies have been investigated in the literature to enhance their properties. Examples include the incorporation of nanoparticles into the material matrix and the utilization of fibers or fabrics as reinforcements. These strategies aim to enhance the mechanical characteristics, thereby elevating the suitability of FDM-printed composites for applications demanding higher performance levels.

The research of our group in this area focuses on different key aspects: studying the effect of filament type and printing parameters on the mechanical properties of AM parts [28], the mechanical performance of AM parts reinforced with natural (jute and curauá) and synthetic (glass) fibers [30], developing natural fiber/filler-reinforced PLA and ABS composites through fused deposition modelling [31, 32] and exploring the mechanical performance of adhesive joints using various fiber reinforcements for AM parts [29].

For example, the use of natural fibers as filament reinforcement was investigated by Cavalcanti et al. [31] by using short curauá fibers with different lengths (3, 6, and 8 mm), and concentrations in terms of weight percentage (2, 3.5, and 5 wt. %) to fabricate polylactic acid (PLA) filaments which were subsequently used to fabricate composites via fused deposition modelling. They concluded that the curauá fiber-reinforced PLA composites may be a promising innovation to improve the performance of these materials, which might enable them to be used for new applications. However, some challenges remain to be solved in the fabrication of the reinforced filaments, such as difficult quality control due to voids and porosities, among others.
composites. On the other hand, they state that the hybridization did not significantly affect the thermal stability of the composites studied.

Interlaminarly hybridized composites of natural/synthetic fiber have been shown in the literature to present a non-linear enhancement of the material properties as the weight fraction of synthetic fibers increases. For example, Queiroz et al. [25], demonstrated that for hybrid composites reinforced with a symmetric synthetic glass envelope around a jute fabric core, the in-plane properties increased significantly as the number of glass layers increased from 2 to 3 (i.e., an approx. 35% enhancement in tensile strength). This is due to the much higher density of synthetic fibers present in a woven fabric compared to the in-plane intralaminar hybridization. On the other hand, Queiroz et al. [26] recently demonstrated that the effect of 3D reinforcement of hybrid natural composites (i.e., z-binder fibers stitched in the transverse direction through intralaminar 2D fabrics) is not significant in the in-plane properties. The highest variation observed in the composite tensile strength (when the 2D case is compared to the 3D) was of approx. 15%, while for most of the studied cases it was below 10%. Therefore, differently from the interlaminar reinforcement, significant variations will not be observed in the in-plane composite properties but rather in the transverse properties (strength and toughness).

## 3D Printing with Natural Fibers

Another area currently developing fast is the application of natural fiber as reinforcements in composites produced by additive manufacturing (AM or 3D printing) [27-32]. This technology allows for the fabrication of complex geometries without the need for expensive tooling and molds. It also allows for precise control over material distribution, opening up new possibilities for complex geometries and customizable structures [33, 34].

The primary materials utilized in Fused Deposition Modeling (FDM) are acrylonitrile butadiene styrene (ABS) and polylactic acid (PLA). However, in their unmodified state, these materials exhibit relatively modest mechanical properties, characterized by typical bulk strengths ranging from 30 to 100 MPa and elastic modulus within the 1.3–3.6 GPa range [27]. Consequently, parts produced through FDM using these materials are not suitable for high-performance applications. To address this limitation, diverse methodologies have been investigated in the literature to enhance their properties. Examples include the incorporation of nanoparticles into the material matrix and the utilization of fibers or fabrics as reinforcements. These strategies aim to enhance the mechanical characteristics, thereby elevating the suitability of FDM-printed composites for applications demanding higher performance levels.

The research of our group in this area focuses on different key aspects: studying the effect of filament type and printing parameters on the mechanical properties of AM parts [28], the mechanical performance of AM parts reinforced with natural (jute and curauá) and synthetic (glass) fibers [30], developing natural fiber/filler-reinforced PLA and ABS composites through fused deposition modelling [31, 32] and exploring the mechanical performance of adhesive joints using various fiber reinforcements for AM parts [29].

For example, the use of natural fibers as filament reinforcement was investigated by Cavalcanti et al. [31] by using short curauá fibers with different lengths (3, 6, and 8 mm), and concentrations in terms of weight percentage (2, 3.5, and 5 wt. %) to fabricate polylactic acid (PLA) filaments which were subsequently used to fabricate composites via fused deposition modelling. They concluded that the curauá fiber-reinforced PLA composites may be a promising innovation to improve the performance of these materials, which might enable them to be used for new applications. However, some challenges remain to be solved in the fabrication of the reinforced filaments, such as difficult quality control due to voids and porosities, among others.
<figure><img src="image_511.png" /><figcaption>Figure 2. Schematic of the fabrication of the reinforced filaments [31].</figcaption></figure>

In another study of our group [30] the mechanical properties of composite materials made by 3D printed (PLA and ABS) parts reinforced by different types of fibers was studied. Natural (jute and curauá) and synthetic (glass) fiber fabrics were laminated on the outer sides of the AM parts by using an epoxy resin via compression moulding. The reinforcement was applied to the AM parts as 1 and 2-layers for each fiber type. The mechanical characterization of the resulting composites was performed through tensile and flexural tests in accordance with their respective ASTM standards. This work provides an alternative to the more common reinforcement methods, by using a hybrid manufacturing method: 3D printing+ compression moulding (see Fig 2). The application of novel manufacturing methodologies allied with renewable biodegradable fibers can provide an interesting path forward for fabricating high-performance 3D printed thermoplastic composites.

<figure><img src="image_512.png" /><figcaption>Figure 3. Schematic of the hybrid manufacturing method: 3D printing+ compression molding.</figcaption></figure>

## Adhesive Joining of New Sustainable Materials

The joining of natural fiber composites to larger structures, such as automotive interiors, is also of crucial importance [12, 35]. This field is still in its infancy and little research has been done. One of the main issues facing the joining of sustainable materials is that natural fiber reinforced
<figure><img src="image_522.png" /><figcaption>Figure 2. Schematic of the fabrication of the reinforced filaments [31].</figcaption></figure>

In another study of our group [30] the mechanical properties of composite materials made by 3D printed (PLA and ABS) parts reinforced by different types of fibers was studied. Natural (jute and curauá) and synthetic (glass) fiber fabrics were laminated on the outer sides of the AM parts by using an epoxy resin via compression moulding. The reinforcement was applied to the AM parts as 1 and 2-layers for each fiber type. The mechanical characterization of the resulting composites was performed through tensile and flexural tests in accordance with their respective ASTM standards. This work provides an alternative to the more common reinforcement methods, by using a hybrid manufacturing method: 3D printing+ compression moulding (see Fig 2). The application of novel manufacturing methodologies allied with renewable biodegradable fibers can provide an interesting path forward for fabricating high-performance 3D printed thermoplastic composites.

<figure><img src="image_523.png" /><figcaption>Figure 3. Schematic of the hybrid manufacturing method: 3D printing+ compression molding.</figcaption></figure>

## Adhesive Joining of New Sustainable Materials

The joining of natural fiber composites to larger structures, such as automotive interiors, is also of crucial importance [12, 35]. This field is still in its infancy and little research has been done. One of the main issues facing the joining of sustainable materials is that natural fiber reinforced
composites present significantly lower mechanical properties when compared to their synthetic counterparts. This is a significant drawback, and it hinders the bonded joint performance of these materials. Therefore, it is crucial to enhance the mechanical properties of sustainable materials prior to bonding them. The reason the adherend material properties of composites are vital in a bonded joint is that due to the nature of the global geometry of the commonly used single lap joint (SLJ), an out-of-plane stress develops due to the eccentricity of the load path between the adherends (see Fig. 4). This peel stress is responsible for the crack initiation at the overlap edges where the stresses (both peel and shear) are highest. However, the propagation and path of the crack is controlled by other parameters, such as composite transverse strength and toughness. Furthermore, adherend stiffness will significantly impact its rotation tendency during the loading of the SLJ, either increasing or decreasing the peel stresses. This leads to the failure mode of delamination which is common for composite bonded joints and is characterized by the removal, partial or complete, of one or more layers of the composite. So far, the hybridization technique has been shown to be the most effective in enhancing the material properties of these more environmentally friendly materials, usually by employing a small weight fraction of synthetic fibers. This hybridization is done in the form of distinct architectural techniques, such as: intralaminar, interlaminar and 3D architectures. The intralaminar reinforcement architecture is defined by the application of two or more reinforcement fiber phases within the same fabric layer (see Fig. 5a). This architecture is more complex and labor intensive, but a highly tailored composite architecture can be achieved. On the other hand, the interlaminar reinforcement architecture is simpler, where each layer is comprised of a single fiber type, usually a fabric (see Fig. 5b). This arrangement has the advantage of being able to leverage the synthetic fiber's high water barrier properties to keep moisture from entering the composite [12, 36]. This is particularly relevant for natural fibers which have a high hydrophilic tendency.

<figure><img src="image_513.png" /><figcaption>Figure 4. Schematic of composite bonded joint stress state and relevant parameters.</figcaption></figure>
composites present significantly lower mechanical properties when compared to their synthetic counterparts. This is a significant drawback, and it hinders the bonded joint performance of these materials. Therefore, it is crucial to enhance the mechanical properties of sustainable materials prior to bonding them. The reason the adherend material properties of composites are vital in a bonded joint is that due to the nature of the global geometry of the commonly used single lap joint (SLJ), an out-of-plane stress develops due to the eccentricity of the load path between the adherends (see Fig. 4). This peel stress is responsible for the crack initiation at the overlap edges where the stresses (both peel and shear) are highest. However, the propagation and path of the crack is controlled by other parameters, such as composite transverse strength and toughness. Furthermore, adherend stiffness will significantly impact its rotation tendency during the loading of the SLJ, either increasing or decreasing the peel stresses. This leads to the failure mode of delamination which is common for composite bonded joints and is characterized by the removal, partial or complete, of one or more layers of the composite. So far, the hybridization technique has been shown to be the most effective in enhancing the material properties of these more environmentally friendly materials, usually by employing a small weight fraction of synthetic fibers. This hybridization is done in the form of distinct architectural techniques, such as: intralaminar, interlaminar and 3D architectures. The intralaminar reinforcement architecture is defined by the application of two or more reinforcement fiber phases within the same fabric layer (see Fig. 5a). This architecture is more complex and labor intensive, but a highly tailored composite architecture can be achieved. On the other hand, the interlaminar reinforcement architecture is simpler, where each layer is comprised of a single fiber type, usually a fabric (see Fig. 5b). This arrangement has the advantage of being able to leverage the synthetic fiber's high water barrier properties to keep moisture from entering the composite [12, 36]. This is particularly relevant for natural fibers which have a high hydrophilic tendency.

<figure><img src="image_524.png" /><figcaption>Figure 4. Schematic of composite bonded joint stress state and relevant parameters.</figcaption></figure>
<figure><img src="image_514.png" /><figcaption>Figure 5. Types of architectural reinforcement, a) intralaminar, b) interlaminar and c) 3D.</figcaption></figure>

Another crucial material property of a composite to be bonded is the transverse strength and toughness. Due to the much lower transverse strength, the delamination failure is directly linked to the peel stresses exceeding the strength and a crack initiating and propagating in the interlaminar space. Therefore, the technique of 3D reinforcement comes into play to solve this issue by applying a z-binder fiber phase that binds the fabric layers in the transverse direction, significantly increasing the composite transverse strength and delamination toughness (see Fig. 5c).

Our group is responsible for the most relevant research efforts to date in the effect of natural fiber composite architectures in bonded joint performance [4, 12, 24-26]. The effect of intralaminar reinforcement in the bonded joint performance of hybrid composites was analyzed by Queiroz et al. [4]. Jute and sisal fabrics were used as a base fiber phase for intralaminar reinforcement via weaving with glass fiber to fabricate hybrid composites. Neat jute, sisal and glass fiber reinforced composites were also fabricated via compression molding. A bi-component epoxy matrix was used to fabricate the composites. Single lap joints, both similar and dissimilar, were fabricated and tested in quasi-static tension. An automotive grade epoxy adhesive as well as the epoxy matrix were used as adhesives. It was found that it is possible to reach up to approx. 77% of the bonded joint failure load of the neat glass composite (GFRP) via the glass fiber intralaminar hybridization of jute fabrics.

In a follow-up study by Queiroz et al. [25], the interlaminar reinforcement technique was used to improve the bonded joint performance of jute/glass fiber composites. Glass and jute fiber fabrics were used and the glass layers were placed on the outside of a jute fiber core of 5 layers. The number of exterior glass fiber layers varied from 2 to 3. The architecture was symmetric. An epoxy matrix was used along with a structural automotive bi-component epoxy adhesive. Single lap joints were fabricated and tested. They found that a hybrid interlaminar glass/jute composite with 3 glass layers was able to reach approx. 100% of the bonded joint failure load of a fully synthetic glass composite (GFRP). This was linked to the optimum balance of improved adherend material properties.

Finally, in a recent research effort, Queiroz et al. [26], investigated the effect of a novel 3D reinforcement architecture in the bonded joint performance of hybrid composites. Jute bidirectional fabrics were used as well as sisal, curauá, hemp and glass fibers in order to fabricate intralaminarly reinforced 2D fabrics. These fabrics were employed with the aid of a mold to fabricate 3D fiber-reinforced preforms by weaving a z-binder fiber through 5 2D fabrics. The composites were fabricated via the compression molding technique with an epoxy matrix. Single lap joints were fabricated with a modern automotive grade epoxy adhesive. It was found that the 3D reinforcement successfully enhanced the bonded joint performance of the novel sustainable composite materials for both the sisal and curauá specimens. This also marks the first time in the literature where a fully natural fiber reinforced composite joint matched approx. 100% the bonded joint performance of a fully synthetic joint (i.e., GFRP and CFRP).
<figure><img src="image_525.png" /><figcaption>Figure 5. Types of architectural reinforcement, a) intralaminar, b) interlaminar and c) 3D.</figcaption></figure>

Another crucial material property of a composite to be bonded is the transverse strength and toughness. Due to the much lower transverse strength, the delamination failure is directly linked to the peel stresses exceeding the strength and a crack initiating and propagating in the interlaminar space. Therefore, the technique of 3D reinforcement comes into play to solve this issue by applying a z-binder fiber phase that binds the fabric layers in the transverse direction, significantly increasing the composite transverse strength and delamination toughness (see Fig. 5c).

Our group is responsible for the most relevant research efforts to date in the effect of natural fiber composite architectures in bonded joint performance [4, 12, 24-26]. The effect of intralaminar reinforcement in the bonded joint performance of hybrid composites was analyzed by Queiroz et al. [4]. Jute and sisal fabrics were used as a base fiber phase for intralaminar reinforcement via weaving with glass fiber to fabricate hybrid composites. Neat jute, sisal and glass fiber reinforced composites were also fabricated via compression molding. A bi-component epoxy matrix was used to fabricate the composites. Single lap joints, both similar and dissimilar, were fabricated and tested in quasi-static tension. An automotive grade epoxy adhesive as well as the epoxy matrix were used as adhesives. It was found that it is possible to reach up to approx. 77% of the bonded joint failure load of the neat glass composite (GFRP) via the glass fiber intralaminar hybridization of jute fabrics.

In a follow-up study by Queiroz et al. [25], the interlaminar reinforcement technique was used to improve the bonded joint performance of jute/glass fiber composites. Glass and jute fiber fabrics were used and the glass layers were placed on the outside of a jute fiber core of 5 layers. The number of exterior glass fiber layers varied from 2 to 3. The architecture was symmetric. An epoxy matrix was used along with a structural automotive bi-component epoxy adhesive. Single lap joints were fabricated and tested. They found that a hybrid interlaminar glass/jute composite with 3 glass layers was able to reach approx. 100% of the bonded joint failure load of a fully synthetic glass composite (GFRP). This was linked to the optimum balance of improved adherend material properties.

Finally, in a recent research effort, Queiroz et al. [26], investigated the effect of a novel 3D reinforcement architecture in the bonded joint performance of hybrid composites. Jute bidirectional fabrics were used as well as sisal, curauá, hemp and glass fibers in order to fabricate intralaminarly reinforced 2D fabrics. These fabrics were employed with the aid of a mold to fabricate 3D fiber-reinforced preforms by weaving a z-binder fiber through 5 2D fabrics. The composites were fabricated via the compression molding technique with an epoxy matrix. Single lap joints were fabricated with a modern automotive grade epoxy adhesive. It was found that the 3D reinforcement successfully enhanced the bonded joint performance of the novel sustainable composite materials for both the sisal and curauá specimens. This also marks the first time in the literature where a fully natural fiber reinforced composite joint matched approx. 100% the bonded joint performance of a fully synthetic joint (i.e., GFRP and CFRP).
## Conclusion

The increasing importance of sustainable materials is a response to the pressing need to address global environmental challenges. This background sets the stage for a deeper exploration of recent advances in the development, characterization, and joining of new sustainable materials, as industries seek transformative solutions to create a more sustainable and resilient future. This paper reported the recent advances that have been made in this field. The main take away can be summarized as follows:

* The application of natural fiber composites, particularly in automotive, civil industry and sports is quite promising and will continually increase in the future due to the significant pressure on reducing costs and increasing demand for material recyclability.

* The future of natural fiber composites is likely to witness an increased focus on hybridization, combining different types of natural fibers or natural/synthetic fibers to create synergistic composites. This approach can enhance mechanical properties, durability, and tailor the materials for specific applications.

* The hybridization technique has clearly demonstrated its ability to improve the natural fiber composites adherend properties and the bonded joint efficiency at relatively low fabrication complexity. In addition, synthetic interlaminar hybridization also provides benefits by acting as a barrier to moisture, increasing the service life of natural/synthetic fiber reinforced composite structures. Finally, NFRC joints can be a viable replacement for synthetic composite joints at no load-bearing loss.

## References

[1] J. Neto, H. Queiroz, R. Aguiar, R. Lima, D. Cavalcanti, M. D. Banea, A review of recent advances in hybrid natural fiber reinforced polymer composites, J. Renew. Mater. 10, (2022) 561-589. https://doi.org/10.32604/jrm.2022.017434

[2] K. L. Pickering, M. A. Efendy, T. M. Le, A review of recent developments in natural fibre composites and their mechanical performance, Compos. - A: Appl. Sci. 83, (2016) 98-112.

[3] B. Ravishankar, S. K. Nayak, M. A. Kader, Hybrid composites for automotive applications – A review, J. Reinf. Plast. Compos. 38, (2019) 835-845. https://doi.org/10.1177/0731684419849708

[4] H. F. M. de Queiroz, M. D. Banea, D. K. K. Cavalcanti, Experimental analysis of adhesively bonded joints in synthetic- and natural fibre-reinforced polymer composites, J. Compos. Mater. 54, (2020) 1245-1255. https://doi.org/10.1177/0021998319876979

[5] J. S. S. Neto, H. F. M. de Queiroz, R. A. A. Aguiar, M. D. Banea, A review on the thermal characterisation of natural and hybrid fiber composites, Polym. 13, (2021) https://doi.org/10.3390/polym13244425

[6] D. K. K. Cavalcanti, M. D. Banea, J. S. S. Neto, R. A. A. Lima, Comparative analysis of the mechanical and thermal properties of polyester and epoxy natural fibre-reinforced hybrid composites, J. Compos. Mater. 55, (2021) 1683-1692. https://doi.org/10.1177/0021998320976811

[7] V. Chauhan, T. Kärki, J. Varis, Review of natural fiber-reinforced engineering plastic composites, their applications in the transportation sector and processing techniques, J. Thermoplast. Compos. Mater. 35, (2022) 1169-1209.

[8] U. K. Sanivada, G. Mármol, F. P. Brito, R. Fangueiro, PLA Composites Reinforced with Flax and Jute Fibers—A Review of Recent Trends, Processing Parameters and Mechanical Properties, Polym. 12, (2020) 2373.
## Conclusion

The increasing importance of sustainable materials is a response to the pressing need to address global environmental challenges. This background sets the stage for a deeper exploration of recent advances in the development, characterization, and joining of new sustainable materials, as industries seek transformative solutions to create a more sustainable and resilient future. This paper reported the recent advances that have been made in this field. The main take away can be summarized as follows:

* The application of natural fiber composites, particularly in automotive, civil industry and sports is quite promising and will continually increase in the future due to the significant pressure on reducing costs and increasing demand for material recyclability.

* The future of natural fiber composites is likely to witness an increased focus on hybridization, combining different types of natural fibers or natural/synthetic fibers to create synergistic composites. This approach can enhance mechanical properties, durability, and tailor the materials for specific applications.

* The hybridization technique has clearly demonstrated its ability to improve the natural fiber composites adherend properties and the bonded joint efficiency at relatively low fabrication complexity. In addition, synthetic interlaminar hybridization also provides benefits by acting as a barrier to moisture, increasing the service life of natural/synthetic fiber reinforced composite structures. Finally, NFRC joints can be a viable replacement for synthetic composite joints at no load-bearing loss.

## References

[1] J. Neto, H. Queiroz, R. Aguiar, R. Lima, D. Cavalcanti, M. D. Banea, A review of recent advances in hybrid natural fiber reinforced polymer composites, J. Renew. Mater. 10, (2022) 561-589. https://doi.org/10.32604/jrm.2022.017434

[2] K. L. Pickering, M. A. Efendy, T. M. Le, A review of recent developments in natural fibre composites and their mechanical performance, Compos. - A: Appl. Sci. 83, (2016) 98-112.

[3] B. Ravishankar, S. K. Nayak, M. A. Kader, Hybrid composites for automotive applications – A review, J. Reinf. Plast. Compos. 38, (2019) 835-845. https://doi.org/10.1177/0731684419849708

[4] H. F. M. de Queiroz, M. D. Banea, D. K. K. Cavalcanti, Experimental analysis of adhesively bonded joints in synthetic- and natural fibre-reinforced polymer composites, J. Compos. Mater. 54, (2020) 1245-1255. https://doi.org/10.1177/0021998319876979

[5] J. S. S. Neto, H. F. M. de Queiroz, R. A. A. Aguiar, M. D. Banea, A review on the thermal characterisation of natural and hybrid fiber composites, Polym. 13, (2021) https://doi.org/10.3390/polym13244425

[6] D. K. K. Cavalcanti, M. D. Banea, J. S. S. Neto, R. A. A. Lima, Comparative analysis of the mechanical and thermal properties of polyester and epoxy natural fibre-reinforced hybrid composites, J. Compos. Mater. 55, (2021) 1683-1692. https://doi.org/10.1177/0021998320976811

[7] V. Chauhan, T. Kärki, J. Varis, Review of natural fiber-reinforced engineering plastic composites, their applications in the transportation sector and processing techniques, J. Thermoplast. Compos. Mater. 35, (2022) 1169-1209.

[8] U. K. Sanivada, G. Mármol, F. P. Brito, R. Fangueiro, PLA Composites Reinforced with Flax and Jute Fibers—A Review of Recent Trends, Processing Parameters and Mechanical Properties, Polym. 12, (2020) 2373.
[9] H. Ku, H. Wang, N. Pattarachaiyakoop, M. Trada, A review on the tensile properties of natural fiber reinforced polymer composites, Compos. B. Eng. 42, (2011) 856-873.
https://doi.org/10.1016/j.compositesb.2011.01.010

[10] R. de Araujo Alves Lima, D. Kawasaki Cavalcanti, J. de Souza e Silva Neto, H. Meneses da Costa, M. D. Banea, Effect of surface treatments on interfacial properties of natural intralaminar hybrid composites, Polym. Compos. 41, (2020) 314-325. https://doi.org/10.1002/pc.25371

[11] Y. Swolfs, L. Gorbatikh, I. Verpoest, Fibre hybridisation in polymer composites: A review, Compos. - A: Appl. Sci. 67, (2014) 181-200. https://doi.org/10.1016/j.compositesa.2014.08.027

[12] H. F. M. de Queiroz, M. D. Banea, Methods to increase the mechanical performance of composite adhesive joints: An overview with focus on joints with natural fibre composite adherends, Journal of Composite Materials. 56, (2022) 3993-4010.
https://doi.org/10.1177/00219983221125906

[13] M. D. Banea, J. S. S. Neto, D. K. K. Cavalcanti, "Recent Trends in Surface Modification of Natural Fibres for Their Use in Green Composites," in Materials Horizons: From Nature to Nanomaterials, ed, 2021, pp. 329-350.

[14] J. S. S. Neto, R. A. A. Lima, D. K. K. Cavalcanti, J. P. B. Souza, R. A. A. Aguiar, M. D. Banea, Effect of chemical treatment on the thermal properties of hybrid natural fiber-reinforced composites, J. Appl. Polym. Sci. 136, (2019). https://doi.org/10.1002/app.47154

[15] H. Fernandes Medeiros de Queiroz, M. D. Banea, D. Kioshi Kawasaki Cavalcanti, J. de Souza e Silva Neto, The effect of multiscale hybridization on the mechanical properties of natural fiber-reinforced composites, J. Appl. Polym. Sci. 138, (2021) 10.1002/app.51213

[16] J. S. S. Neto, M. D. Banea, D. K. K. Cavalcanti, H. F. M. Queiroz, R. A. A. Aguiar, Analysis of mechanical and thermal properties of epoxy multiwalled carbon nanocomposites, J. Compos. Mater. 54, (2020) 4831-4840. https://doi.org/10.1177/0021998320939232

[17] J. S. S. Neto, H. F. M. de Queiroz, D. K. K. Cavalcanti, R. A. A. Aguiar, A. B. Pereira, M. D. Banea, Effect of ultraviolet radiation and water spraying on the mechanical properties of multi-walled carbon nanotubes reinforced natural fiber and hybrid composites, J. Appl. Polym. Sci. 139, (2022) 10.1002/app.51915

[18] V. Prasad, K. Sekar, M. A. Joseph, Mechanical and water absorption properties of nano TiO2 coated flax fibre epoxy composites, Constr Build Mater. 284, (2021) 122803.
https://doi.org/10.1016/j.conbuildmat.2021.122803

[19] J. S. S. Neto, D. K. K. Cavalcanti, L. E. da Cunha Ferro, H. F. M. de Queiroz, R. A. A. Aguiar, M. D. Banea, Effect of Multi-Walled Carbon Nanotubes on the Mechanical and Thermal Properties of Curaua&acute; Natural-Fiber-Reinforced Composites, C. 9, (2023) 102.

[20] M. D. Banea, Natural Fibre Composites and Their Mechanical Behaviour, Polym. 15, (2023).
https://doi.org/10.3390/polym15051185

[21] D. K. K. Cavalcanti, M. D. Banea, J. S. S. Neto, R. A. A. Lima, L. F. M. da Silva, R. J. C. Carbas, Mechanical characterization of intralaminar natural fibre-reinforced hybrid composites, Compos. B. Eng. 175, (2019). https://doi.org/10.1016/j.compositesb.2019.107149

[22] A. L. Pereira, M. D. Banea, A. B. Pereira, Effect of intralaminar hybridization on mode I fracture toughness of natural fiber-reinforced composites, J BRAZ SOC MECH SCI. 42, (2020).
https://doi.org/10.1007/s40430-020-02525-w

[23] A. L. Pereira, M. D. Banea, J. S. S. Neto, D. K. K. Cavalcanti, Mechanical and thermal characterization of natural intralaminar hybrid composites based on sisal, Polym. 12, (2020).
https://doi.org/10.3390/POLYM12040866
[9] H. Ku, H. Wang, N. Pattarachaiyakoop, M. Trada, A review on the tensile properties of natural fiber reinforced polymer composites, Compos. B. Eng. 42, (2011) 856-873.
https://doi.org/10.1016/j.compositesb.2011.01.010

[10] R. de Araujo Alves Lima, D. Kawasaki Cavalcanti, J. de Souza e Silva Neto, H. Meneses da Costa, M. D. Banea, Effect of surface treatments on interfacial properties of natural intralaminar hybrid composites, Polym. Compos. 41, (2020) 314-325. https://doi.org/10.1002/pc.25371

[11] Y. Swolfs, L. Gorbatikh, I. Verpoest, Fibre hybridisation in polymer composites: A review, Compos. - A: Appl. Sci. 67, (2014) 181-200. https://doi.org/10.1016/j.compositesa.2014.08.027

[12] H. F. M. de Queiroz, M. D. Banea, Methods to increase the mechanical performance of composite adhesive joints: An overview with focus on joints with natural fibre composite adherends, Journal of Composite Materials. 56, (2022) 3993-4010.
https://doi.org/10.1177/00219983221125906

[13] M. D. Banea, J. S. S. Neto, D. K. K. Cavalcanti, "Recent Trends in Surface Modification of Natural Fibres for Their Use in Green Composites," in Materials Horizons: From Nature to Nanomaterials, ed, 2021, pp. 329-350.

[14] J. S. S. Neto, R. A. A. Lima, D. K. K. Cavalcanti, J. P. B. Souza, R. A. A. Aguiar, M. D. Banea, Effect of chemical treatment on the thermal properties of hybrid natural fiber-reinforced composites, J. Appl. Polym. Sci. 136, (2019). https://doi.org/10.1002/app.47154

[15] H. Fernandes Medeiros de Queiroz, M. D. Banea, D. Kioshi Kawasaki Cavalcanti, J. de Souza e Silva Neto, The effect of multiscale hybridization on the mechanical properties of natural fiber-reinforced composites, J. Appl. Polym. Sci. 138, (2021) 10.1002/app.51213

[16] J. S. S. Neto, M. D. Banea, D. K. K. Cavalcanti, H. F. M. Queiroz, R. A. A. Aguiar, Analysis of mechanical and thermal properties of epoxy multiwalled carbon nanocomposites, J. Compos. Mater. 54, (2020) 4831-4840. https://doi.org/10.1177/0021998320939232

[17] J. S. S. Neto, H. F. M. de Queiroz, D. K. K. Cavalcanti, R. A. A. Aguiar, A. B. Pereira, M. D. Banea, Effect of ultraviolet radiation and water spraying on the mechanical properties of multi-walled carbon nanotubes reinforced natural fiber and hybrid composites, J. Appl. Polym. Sci. 139, (2022) 10.1002/app.51915

[18] V. Prasad, K. Sekar, M. A. Joseph, Mechanical and water absorption properties of nano TiO2 coated flax fibre epoxy composites, Constr Build Mater. 284, (2021) 122803.
https://doi.org/10.1016/j.conbuildmat.2021.122803

[19] J. S. S. Neto, D. K. K. Cavalcanti, L. E. da Cunha Ferro, H. F. M. de Queiroz, R. A. A. Aguiar, M. D. Banea, Effect of Multi-Walled Carbon Nanotubes on the Mechanical and Thermal Properties of Curaua&acute; Natural-Fiber-Reinforced Composites, C. 9, (2023) 102.

[20] M. D. Banea, Natural Fibre Composites and Their Mechanical Behaviour, Polym. 15, (2023).
https://doi.org/10.3390/polym15051185

[21] D. K. K. Cavalcanti, M. D. Banea, J. S. S. Neto, R. A. A. Lima, L. F. M. da Silva, R. J. C. Carbas, Mechanical characterization of intralaminar natural fibre-reinforced hybrid composites, Compos. B. Eng. 175, (2019). https://doi.org/10.1016/j.compositesb.2019.107149

[22] A. L. Pereira, M. D. Banea, A. B. Pereira, Effect of intralaminar hybridization on mode I fracture toughness of natural fiber-reinforced composites, J BRAZ SOC MECH SCI. 42, (2020).
https://doi.org/10.1007/s40430-020-02525-w

[23] A. L. Pereira, M. D. Banea, J. S. S. Neto, D. K. K. Cavalcanti, Mechanical and thermal characterization of natural intralaminar hybrid composites based on sisal, Polym. 12, (2020).
https://doi.org/10.3390/POLYM12040866
[24] H. F. M. de Queiroz, R. H. A. Rosa, J. S. S. Neto, D. K. K. Cavalcanti, M. D. Banea, Mechanical characterization of short Curauá/glass fiber-reinforced interlaminar hybrid composite adhesive joints, J Adhes Sci Technol. 2023). https://doi.org/10.1080/01694243.2023.2261263

[25] H. F. M. de Queiroz, M. D. Banea, D. K. K. Cavalcanti, Adhesively bonded joints of jute, glass and hybrid jute/glass fibre-reinforced polymer composites for automotive industry, Appl. Adhes. Sci. 9, (2021) 2. https://doi.org/10.1186/s40563-020-00131-6

[26] H. F. M. de Queiroz, J. S. S. Neto, D. K. K. Cavalcanti, M. D. Banea, A novel 3D fibre-reinforcement architecture for high performance natural fibre reinforced composite adhesively bonded joints, J Adhes. 100, (2024) 118-138. https://doi.org/10.1080/00218464.2023.2200173

[27] D. K. K. Cavalcanti, M. D. Banea, H. F. M. de Queiroz, Mechanical Characterization of Bonded Joints Made of Additive Manufactured Adherends, Ann. Dunarea Jos" Univ. Galati Fascicle XII Weld. Equip. 30, (2019) 27-33. https://doi.org/10.35219/awet.2019.04

[28] D. K. K. Cavalcanti, M. D. Banea, H. F. M. de Queiroz, Effect of material on the mechanical properties of additive manufactured thermoplastic parts, Ann. Dunarea Jos" Univ. Galati Fascicle XII Weld. Equip. 31, (2020) 5-12. https://doi.org/10.35219/awet.2020.01

[29] D. K. K. Cavalcanti, H. F. M. de Queiroz, M. D. Banea, Performance enhancement of adhesive joints of additive manufactured parts by using different types of fibre reinforcements, Int. J. Adhes. 124, (2023). https://doi.org/10.1016/j.ijadhadh.2023.103371

[30] D. K. K. Cavalcanti, H. F. M. de Queiroz, J. S. S. Neto, M. D. Banea, Strengthening of additive manufactured parts by using different type of fibre reinforcements, Int. J. Adv. Manuf. Technol. 123, (2022) 1889-1903. https://doi.org/10.1007/s00170-022-10327-8

[31] D. K. K. Cavalcanti, J. S. S. Neto, H. F. M. D. Queiroz, Y. Wu, V. F. S. Neto, M. D. Banea, Development and Mechanical Characterization of Short Curauá Fiber-Reinforced PLA Composites Made via Fused Deposition Modeling, Polym. 14, (2022). https://doi.org/10.3390/polym14225047

[32] A. C. de Mendonça, D. K. K. Cavalcanti, H. F. M. de Queiroz, J. S. S. Neto, F. J. P. Chaves, M. D. Banea, MECHANICAL CHARACTERIZATION OF FILLER MODIFIED ABS 3D PRINTED COMPOSITES MADE VIA FUSED FILAMENT FABRICATION, Ann. Dunarea Jos" Univ. Galati Fascicle XII Weld. Equip. 33, (2022) 27-34. https://doi.org/10.35219/awet.2022.02

[33] N. V. Santos, D. C. T. Cardoso, 3D printing of vegetable yarn-reinforced polymer components, J. Clean. Prod. 415, (2023) 137870. https://doi.org/10.1016/j.jclepro.2023.137870

[34] P. Boisse, R. Akkerman, P. Carlone, L. Kärger, S. V. Lomov, J. A. Sherwood, Advances in composite forming through 25 years of ESAFORM, Int. J. Mater. Form. 15, (2022) 39. 10.1007/s12289-022-01682-8

[35] Y. Singh, J. Kumar, I. Singh, P. K. Rakesh, "Joining behavior of natural fiber reinforced polymer composites," in *Joining Processes for Dissimilar and Advanced Materials*, ed, 2021, pp. 33-63.

[36] G. B. de Seixas, H. F. M. de Queiroz, J. S. S. Neto, M. D. Banea, Effect of water on the mechanical and thermal properties of natural fibre reinforced hybrid composites, J. Compos. Mater. 57, (2023) 1941-1958. https://doi.org/10.1177/00219983231166570
[24] H. F. M. de Queiroz, R. H. A. Rosa, J. S. S. Neto, D. K. K. Cavalcanti, M. D. Banea, Mechanical characterization of short Curauá/glass fiber-reinforced interlaminar hybrid composite adhesive joints, J Adhes Sci Technol. 2023). https://doi.org/10.1080/01694243.2023.2261263

[25] H. F. M. de Queiroz, M. D. Banea, D. K. K. Cavalcanti, Adhesively bonded joints of jute, glass and hybrid jute/glass fibre-reinforced polymer composites for automotive industry, Appl. Adhes. Sci. 9, (2021) 2. https://doi.org/10.1186/s40563-020-00131-6

[26] H. F. M. de Queiroz, J. S. S. Neto, D. K. K. Cavalcanti, M. D. Banea, A novel 3D fibre-reinforcement architecture for high performance natural fibre reinforced composite adhesively bonded joints, J Adhes. 100, (2024) 118-138. https://doi.org/10.1080/00218464.2023.2200173

[27] D. K. K. Cavalcanti, M. D. Banea, H. F. M. de Queiroz, Mechanical Characterization of Bonded Joints Made of Additive Manufactured Adherends, Ann. Dunarea Jos" Univ. Galati Fascicle XII Weld. Equip. 30, (2019) 27-33. https://doi.org/10.35219/awet.2019.04

[28] D. K. K. Cavalcanti, M. D. Banea, H. F. M. de Queiroz, Effect of material on the mechanical properties of additive manufactured thermoplastic parts, Ann. Dunarea Jos" Univ. Galati Fascicle XII Weld. Equip. 31, (2020) 5-12. https://doi.org/10.35219/awet.2020.01

[29] D. K. K. Cavalcanti, H. F. M. de Queiroz, M. D. Banea, Performance enhancement of adhesive joints of additive manufactured parts by using different types of fibre reinforcements, Int. J. Adhes. 124, (2023). https://doi.org/10.1016/j.ijadhadh.2023.103371

[30] D. K. K. Cavalcanti, H. F. M. de Queiroz, J. S. S. Neto, M. D. Banea, Strengthening of additive manufactured parts by using different type of fibre reinforcements, Int. J. Adv. Manuf. Technol. 123, (2022) 1889-1903. https://doi.org/10.1007/s00170-022-10327-8

[31] D. K. K. Cavalcanti, J. S. S. Neto, H. F. M. D. Queiroz, Y. Wu, V. F. S. Neto, M. D. Banea, Development and Mechanical Characterization of Short Curauá Fiber-Reinforced PLA Composites Made via Fused Deposition Modeling, Polym. 14, (2022). https://doi.org/10.3390/polym14225047

[32] A. C. de Mendonça, D. K. K. Cavalcanti, H. F. M. de Queiroz, J. S. S. Neto, F. J. P. Chaves, M. D. Banea, MECHANICAL CHARACTERIZATION OF FILLER MODIFIED ABS 3D PRINTED COMPOSITES MADE VIA FUSED FILAMENT FABRICATION, Ann. Dunarea Jos" Univ. Galati Fascicle XII Weld. Equip. 33, (2022) 27-34. https://doi.org/10.35219/awet.2022.02

[33] N. V. Santos, D. C. T. Cardoso, 3D printing of vegetable yarn-reinforced polymer components, J. Clean. Prod. 415, (2023) 137870. https://doi.org/10.1016/j.jclepro.2023.137870

[34] P. Boisse, R. Akkerman, P. Carlone, L. Kärger, S. V. Lomov, J. A. Sherwood, Advances in composite forming through 25 years of ESAFORM, Int. J. Mater. Form. 15, (2022) 39. 10.1007/s12289-022-01682-8

[35] Y. Singh, J. Kumar, I. Singh, P. K. Rakesh, "Joining behavior of natural fiber reinforced polymer composites," in *Joining Processes for Dissimilar and Advanced Materials*, ed, 2021, pp. 33-63.

[36] G. B. de Seixas, H. F. M. de Queiroz, J. S. S. Neto, M. D. Banea, Effect of water on the mechanical and thermal properties of natural fibre reinforced hybrid composites, J. Compos. Mater. 57, (2023) 1941-1958. https://doi.org/10.1177/00219983231166570
# Investigation of the friction behavior of uni- and bidirectional non-crimp fabrics

SCHÄFER Bastian¹, a*, NAOUAR Naim² and KÄRGER Luise¹, b*

¹Karlsruhe Institute of Technology (KIT), Institute of Vehicle System Technology (FAST), Lightweight Engineering, Karlsruhe, Germany

²Université de Lyon, LaMCoS CNRS, Lyon, F-69621, France

a bastian.schaefer@kit.edu, b luise.kaerger@kit.edu

**Keywords:** Fabrics/Textiles, Experimental Characterization, Friction, UD-NCF, Biax-NCF

**Abstract.** The friction behavior of engineering textiles directly affects the forming quality during composite molding processes. In forming tests of dry engineering textiles large relative slip between plies and the tools is observed. The resulting tangential contact stresses influence the material's membrane stresses, which in turn impact the fabric's deformation and potentially lead to forming defects such as gapping or ruptures of the textile. The characterization of friction is commonly conducted via relative motion between a fabric ply and either another fabric ply (ply-ply) or a tool (tool-ply) under controlled transverse pressure. The resulting behavior of a textile reinforcement depends on the mesoscopic structure of its unit cell and the material of its constituents. In this work, the tangential friction behavior at interfaces between ply and tooling and between plies of a unidirectional and a bidirectional non-crimp fabric are investigated in sled pull-over-tests. The behavior is analyzed with respect to the applied normal forces, the relative velocity and the relative fiber orientation. A generally rate-independent behavior is observed. Tool-ply friction is only slightly affected by the applied pressure, while ply-ply friction is strongly influenced by the stitching pattern at the contact interface.

## Introduction

The final fiber orientation in components manufactured from liquid composite molding processes is significantly influenced by relative slip between individual fabric layers as well as the tools. The relative movement results in tangential friction, which induces in-plane stresses and impacts the fabric's deformation. This influences the material draw-in [1] and can potentially result in defects such as gapping or ruptures of the textile [2]. Non-crimp fabrics (NCFs) have straight fibers compared to woven fabrics with undulated fibers and therefore provide a higher lightweight potential. However, they are more susceptible to defects due to the low stiffness and strength of the stitching.

The characterization of friction is commonly conducted via relative motion between a fabric ply and either another fabric ply (ply-ply) or a tool (tool-ply) under controlled transverse pressure. Usually, Coulomb's friction is assumed for the analysis of dry textiles

$$f_f = \mu N, \qquad (1)$$

where $f_f$ is the tangential friction force, $N$ is the applied normal load and $\mu$ is the coefficient of friction (CoF).

In a benchmark study, Sachs et al. [3] compared different experimental setups from three design categories for the same material, i.e. pull-out-tests, pull-through-tests and sled pull-over-tests. Comparable results were achieved for all setups, but systematic errors were observed and attributed to edge effects and uneven pressure distributions. Thus, sled pull-over- or pull-through-

<figure><img src="image_515.png" /></figure>

Content from this work may be used under the terms of the Creative Commons Attribution 3.0 license. Any further distribution of this work must maintain attribution to the author(s) and the title of the work, journal citation and DOI. Published under license by Materials Research Forum LLC.
# Investigation of the friction behavior of uni- and bidirectional non-crimp fabrics

SCHÄFER Bastian¹, a*, NAOUAR Naim² and KÄRGER Luise¹, b*

¹Karlsruhe Institute of Technology (KIT), Institute of Vehicle System Technology (FAST), Lightweight Engineering, Karlsruhe, Germany

²Université de Lyon, LaMCoS CNRS, Lyon, F-69621, France

a bastian.schaefer@kit.edu, b luise.kaerger@kit.edu

**Keywords:** Fabrics/Textiles, Experimental Characterization, Friction, UD-NCF, Biax-NCF

**Abstract.** The friction behavior of engineering textiles directly affects the forming quality during composite molding processes. In forming tests of dry engineering textiles large relative slip between plies and the tools is observed. The resulting tangential contact stresses influence the material's membrane stresses, which in turn impact the fabric's deformation and potentially lead to forming defects such as gapping or ruptures of the textile. The characterization of friction is commonly conducted via relative motion between a fabric ply and either another fabric ply (ply-ply) or a tool (tool-ply) under controlled transverse pressure. The resulting behavior of a textile reinforcement depends on the mesoscopic structure of its unit cell and the material of its constituents. In this work, the tangential friction behavior at interfaces between ply and tooling and between plies of a unidirectional and a bidirectional non-crimp fabric are investigated in sled pull-over-tests. The behavior is analyzed with respect to the applied normal forces, the relative velocity and the relative fiber orientation. A generally rate-independent behavior is observed. Tool-ply friction is only slightly affected by the applied pressure, while ply-ply friction is strongly influenced by the stitching pattern at the contact interface.

## Introduction

The final fiber orientation in components manufactured from liquid composite molding processes is significantly influenced by relative slip between individual fabric layers as well as the tools. The relative movement results in tangential friction, which induces in-plane stresses and impacts the fabric's deformation. This influences the material draw-in [1] and can potentially result in defects such as gapping or ruptures of the textile [2]. Non-crimp fabrics (NCFs) have straight fibers compared to woven fabrics with undulated fibers and therefore provide a higher lightweight potential. However, they are more susceptible to defects due to the low stiffness and strength of the stitching.

The characterization of friction is commonly conducted via relative motion between a fabric ply and either another fabric ply (ply-ply) or a tool (tool-ply) under controlled transverse pressure. Usually, Coulomb's friction is assumed for the analysis of dry textiles

$$f_f = \mu N, \qquad (1)$$

where $f_f$ is the tangential friction force, $N$ is the applied normal load and $\mu$ is the coefficient of friction (CoF).

In a benchmark study, Sachs et al. [3] compared different experimental setups from three design categories for the same material, i.e. pull-out-tests, pull-through-tests and sled pull-over-tests. Comparable results were achieved for all setups, but systematic errors were observed and attributed to edge effects and uneven pressure distributions. Thus, sled pull-over- or pull-through-

<figure><img src="image_526.png" /></figure>

Content from this work may be used under the terms of the Creative Commons Attribution 3.0 license. Any further distribution of this work must maintain attribution to the author(s) and the title of the work, journal citation and DOI. Published under license by Materials Research Forum LLC.
tests with a sufficiently large contact area and chamfered edges were recommended. However, the sled-test was originally developed for thin plastic film materials (ASTM standard D1894-14 [4]) and therefore mainly suitable for lower transverse pressure to prevent balancing problems and stick-slip effects [2,5].

The friction behavior of a textile reinforcement depends on the mesoscopic structure of its unit cell and the material of its constituents. Protruding high points of the material are especially notable during ply-ply contact due to stick-slip or shock effects. These result from undulations in woven fabrics [6,7,8] or the stitching in non-crimp fabrics [2,9]. Additionally, pressure- and rate-dependency as well as an impact of a layer's relative orientation and deformation state (shear) are reported for dry textiles.

Tool-ply friction for dry textiles is usually found to be nearly independent of the relative velocity [3,10], of the ply's relative orientation [9,11,12] and for woven fabrics independent of shear deformation applied to the textile before testing [13]. An increasing normal pressure reduces the CoF [10,14,15,16], because of its effect on the surface roughness and true contact area between fibrous material as demonstrated by Avgoulas et al. [16] based on Hertzian theory. However, this correlation applies mainly for high pressures (> 10 kPa), while for low pressures often no or even a slightly positive influence is measured [3,16,17].

Ply-ply friction for dry textiles is also reported as nearly independent of the relative velocity [10,18], but for woven fabric increasing under shear deformation applied to the textile before testing [13]. The effective CoF between plies decreases for high transverse pressure [10,14,18], due to the lateral spreading of the fabric under compaction that results in smaller undulations and thus less shock effects [18]. The reported influence of the relative orientation is most heavily dependent on the specific material being tested. The most common observation for woven fabrics is a lower CoF for relative orientations unequal to 0° or 90° [10,19], with the lowest value often, but not always measured for 45° [6,13]. In contrast, the behavior of NCFs does not follow a distinct trend and depends on the precise architecture [2,9,12]. However, Quenzel et al. [9] recently measured an increasing CoF by decreasing the stitching length of Biax-NCFs manufactured from the same materials as it increases the number of stitch yarn loops.

In this work, the tangential friction behavior at interfaces between ply and tooling (TP) and between plies (PP) of a unidirectional (UD) and a bidirectional (Biax) NCF is investigated in sled pull-over-tests to identify the most relevant factors that influence the friction behavior. The behavior is analyzed with respect to the applied normal forces, relative velocity and relative fiber orientation as well as fabric side at the interfaces. Both NCFs are stitched in a tricot pattern, resulting in an additional influence of the fabric side that has not been investigated in previous studies on NCFs because they applied bidirectional fabrics with similar chain stitching patterns on both sides [2,9,12].

## Experimental test setup and procedure

*Materials.* In this work, a unidirectional (UD300) and a bidirectional (MD600) non-crimp fabric, both without binder, are used. The fabrics are manufactured by Zoltek and produced from the same PX35-50K continuous carbon fiber (CF) heavy tows. Both fabrics are stitched together with a 76 dtex PES yarn in a tricot pattern. The UD-NCF consists of a single layer of aligned CF rovings with thin glass fibers (GF) on the back for improved handleability and the Biax-NCF of two layers in a 0°/90° orientation. Both fabrics have a similar number of CFs with about 300 g/m² in their respective main reinforcing directions.

*Sled pull-over tests.* The inter-ply behavior is investigated with a sled pull-over setup based on ASTM standard D1894-14 [4] mounted to a universal testing machine, cf. Figure 1. All edges on the sled are chamfered and relatively light weights are used to minimize stick-slip effects and comply with the findings of the benchmark of Sachs et al. for dry textiles [3]. The surface
tests with a sufficiently large contact area and chamfered edges were recommended. However, the sled-test was originally developed for thin plastic film materials (ASTM standard D1894-14 [4]) and therefore mainly suitable for lower transverse pressure to prevent balancing problems and stick-slip effects [2,5].

The friction behavior of a textile reinforcement depends on the mesoscopic structure of its unit cell and the material of its constituents. Protruding high points of the material are especially notable during ply-ply contact due to stick-slip or shock effects. These result from undulations in woven fabrics [6,7,8] or the stitching in non-crimp fabrics [2,9]. Additionally, pressure- and rate-dependency as well as an impact of a layer's relative orientation and deformation state (shear) are reported for dry textiles.

Tool-ply friction for dry textiles is usually found to be nearly independent of the relative velocity [3,10], of the ply's relative orientation [9,11,12] and for woven fabrics independent of shear deformation applied to the textile before testing [13]. An increasing normal pressure reduces the CoF [10,14,15,16], because of its effect on the surface roughness and true contact area between fibrous material as demonstrated by Avgoulas et al. [16] based on Hertzian theory. However, this correlation applies mainly for high pressures (> 10 kPa), while for low pressures often no or even a slightly positive influence is measured [3,16,17].

Ply-ply friction for dry textiles is also reported as nearly independent of the relative velocity [10,18], but for woven fabric increasing under shear deformation applied to the textile before testing [13]. The effective CoF between plies decreases for high transverse pressure [10,14,18], due to the lateral spreading of the fabric under compaction that results in smaller undulations and thus less shock effects [18]. The reported influence of the relative orientation is most heavily dependent on the specific material being tested. The most common observation for woven fabrics is a lower CoF for relative orientations unequal to 0° or 90° [10,19], with the lowest value often, but not always measured for 45° [6,13]. In contrast, the behavior of NCFs does not follow a distinct trend and depends on the precise architecture [2,9,12]. However, Quenzel et al. [9] recently measured an increasing CoF by decreasing the stitching length of Biax-NCFs manufactured from the same materials as it increases the number of stitch yarn loops.

In this work, the tangential friction behavior at interfaces between ply and tooling (TP) and between plies (PP) of a unidirectional (UD) and a bidirectional (Biax) NCF is investigated in sled pull-over-tests to identify the most relevant factors that influence the friction behavior. The behavior is analyzed with respect to the applied normal forces, relative velocity and relative fiber orientation as well as fabric side at the interfaces. Both NCFs are stitched in a tricot pattern, resulting in an additional influence of the fabric side that has not been investigated in previous studies on NCFs because they applied bidirectional fabrics with similar chain stitching patterns on both sides [2,9,12].

## Experimental test setup and procedure

*Materials.* In this work, a unidirectional (UD300) and a bidirectional (MD600) non-crimp fabric, both without binder, are used. The fabrics are manufactured by Zoltek and produced from the same PX35-50K continuous carbon fiber (CF) heavy tows. Both fabrics are stitched together with a 76 dtex PES yarn in a tricot pattern. The UD-NCF consists of a single layer of aligned CF rovings with thin glass fibers (GF) on the back for improved handleability and the Biax-NCF of two layers in a 0°/90° orientation. Both fabrics have a similar number of CFs with about 300 g/m² in their respective main reinforcing directions.

*Sled pull-over tests.* The inter-ply behavior is investigated with a sled pull-over setup based on ASTM standard D1894-14 [4] mounted to a universal testing machine, cf. Figure 1. All edges on the sled are chamfered and relatively light weights are used to minimize stick-slip effects and comply with the findings of the benchmark of Sachs et al. for dry textiles [3]. The surface
roughness R_a of the sled is, similar to the roughness of a mold applied during resin transfer molding or wet compression molding, about 0.617 µm [11].

<figure><img src="image_516.png" /><figcaption>Fig. 1. Schematic illustration of the experimental friction setup and test bench on tensile machine.</figcaption></figure>

A dry specimen is clamped to the bottom surface of the setup and a sled with a surface area of 65 x 65 mm² is pulled over a distance of 100 mm at a constant velocity vᵢ over the fabric to investigate tool-ply friction (TP). The influence of the applied normal force is investigated with three different sled weights of mᵢ ∈ [218 g, 720 g, 1222 g]. Two different velocities are applied during the tests with v₁ = 50 mm/min and v₂ = 150 mm/min. An additional specimen is wrapped around the sled to investigate ply-ply friction (PP). The clamps used to secure the specimen to the sled and the fabric itself add weight, which is measured before each repetition and is approximately ~30-35g. The front and back of both fabrics are investigated separately due to the different visible stitching patterns. Each test is repeated at least three times for each configuration with new specimens. The fiber orientation of the visible side of each fabric is used for a consistent naming convention to investigate different fabric orientations, cf. Figure 2.

<figure><img src="image_517.png" /><figcaption>Fig. 2. Naming convention for the inter-ply characterization tests.</figcaption></figure>

The measured pulling force of the sled f_s is used to calculate the resulting coefficient of friction (CoF) according to Equation 1 with:

$$ \mu = \frac{f_s}{g m_i} \qquad (2) $$
roughness $R_a$ of the sled is, similar to the roughness of a mold applied during resin transfer molding or wet compression molding, about 0.617 µm [11].

<figure><img src="image_527.png" /><figcaption>Fig. 1. Schematic illustration of the experimental friction setup and test bench on tensile machine.</figcaption></figure>

A dry specimen is clamped to the bottom surface of the setup and a sled with a surface area of 65 x 65 mm² is pulled over a distance of 100 mm at a constant velocity $v_i$ over the fabric to investigate tool-ply friction (TP). The influence of the applied normal force is investigated with three different sled weights of $m_i \in [218 \text{ g}, 720 \text{ g}, 1222 \text{ g}]$. Two different velocities are applied during the tests with $v_1 = 50 \text{ mm/min}$ and $v_2 = 150 \text{ mm/min}$. An additional specimen is wrapped around the sled to investigate ply-ply friction (PP). The clamps used to secure the specimen to the sled and the fabric itself add weight, which is measured before each repetition and is approximately ~30-35g. The front and back of both fabrics are investigated separately due to the different visible stitching patterns. Each test is repeated at least three times for each configuration with new specimens. The fiber orientation of the visible side of each fabric is used for a consistent naming convention to investigate different fabric orientations, cf. Figure 2.

<figure><img src="image_528.png" /><figcaption>Fig. 2. Naming convention for the inter-ply characterization tests.</figcaption></figure>

The measured pulling force of the sled $f_s$ is used to calculate the resulting coefficient of friction (CoF) according to Equation 1 with:

$$ \mu = \frac{f_s}{g m_i} \qquad (2) $$
The static CoF $\mu_s$ is obtained as the maximum value, which occurs in all tests within the first 10 mm of sled displacement due to a slight initial slag in the wire. The dynamic CoF $\mu_d$ is calculated as the average between a displacement of 40 mm and 90 mm where a steady state behavior was observed for all tests. The calculation of both CoF is demonstrated in Figure 3. The wavelength of the oscillations in the PP-tests agrees closely with the stitching pitch (7.2 mm) of UD-NCF. This behavior is similar to stick-slip and shock effects observed for other NCFs with chain stitching patterns [2,9] as well as the effect of undulations in woven fabrics [6,7,8]. The averages and standard deviations for the dynamic CoF are calculated based on the raw CoF-displacement curves of all individual tests in the desired range, rather than the standard deviation of the average CoF for each individual, to account for the magnitude of the oscillations during PP-friction. The static CoF is 5 - 20 % higher for TP-friction and 20 – 40 % for PP-friction than the dynamic CoF in all tests. However, the static and dynamic CoF show very similar tendencies in comparisons for different configurations. The results are therefore discussed below on the basis of the dynamic CoF.

<figure><img src="image_518.png" /><figcaption>Fig. 3. Schematic illustration of the calculation of the static $\mu_s$ and dynamic $\mu_d$ coefficient of friction. The $\pm$ values give the standard deviation of all tests in the desired range and for the dynamic CoF represent the magnitude of the observed oscillations.</figcaption></figure>

## Results

*Tool-ply.* The influence of the sled weight and velocity on the tool-ply behavior are investigated for the 0°-configuration of both materials, cf. Figure 4 a. The dynamic CoF is independent of the sled velocity for both materials (same colored bars). A higher sled weight results in a slight increase of the measured $\mu_d$ (each column) for UD-NCF and reduces the standard deviation. This behavior is similar to observations for other engineering textiles at low pressure [3,16,17]. For Biax-NCF, the increase of $\mu_d$ for a higher sled weight is only significant for the higher sled velocity, but with overlapping scatter.

The influence of the fiber orientation and fabric side on the TP behavior of both materials is investigated at a constant sled weight $m_2$ and velocity $v_1$, cf. Figure 4 b. The dynamic CoF is slightly higher for the back side of UD-NCF, but no clear influence of the fiber orientation is observed. For Biax-NCF, no clear trend is observed for the ply's side or orientation.
The static CoF $\mu_s$ is obtained as the maximum value, which occurs in all tests within the first 10 mm of sled displacement due to a slight initial slag in the wire. The dynamic CoF $\mu_d$ is calculated as the average between a displacement of 40 mm and 90 mm where a steady state behavior was observed for all tests. The calculation of both CoF is demonstrated in Figure 3. The wavelength of the oscillations in the PP-tests agrees closely with the stitching pitch (7.2 mm) of UD-NCF. This behavior is similar to stick-slip and shock effects observed for other NCFs with chain stitching patterns [2,9] as well as the effect of undulations in woven fabrics [6,7,8]. The averages and standard deviations for the dynamic CoF are calculated based on the raw CoF-displacement curves of all individual tests in the desired range, rather than the standard deviation of the average CoF for each individual, to account for the magnitude of the oscillations during PP-friction. The static CoF is 5 - 20 % higher for TP-friction and 20 – 40 % for PP-friction than the dynamic CoF in all tests. However, the static and dynamic CoF show very similar tendencies in comparisons for different configurations. The results are therefore discussed below on the basis of the dynamic CoF.

<figure><img src="image_529.png" /><figcaption>Fig. 3. Schematic illustration of the calculation of the static $\mu_s$ and dynamic $\mu_d$ coefficient of friction. The $\pm$ values give the standard deviation of all tests in the desired range and for the dynamic CoF represent the magnitude of the observed oscillations.</figcaption></figure>

## Results

*Tool-ply.* The influence of the sled weight and velocity on the tool-ply behavior are investigated for the 0°-configuration of both materials, cf. Figure 4 a. The dynamic CoF is independent of the sled velocity for both materials (same colored bars). A higher sled weight results in a slight increase of the measured $\mu_d$ (each column) for UD-NCF and reduces the standard deviation. This behavior is similar to observations for other engineering textiles at low pressure [3,16,17]. For Biax-NCF, the increase of $\mu_d$ for a higher sled weight is only significant for the higher sled velocity, but with overlapping scatter.

The influence of the fiber orientation and fabric side on the TP behavior of both materials is investigated at a constant sled weight $m_2$ and velocity $v_1$, cf. Figure 4 b. The dynamic CoF is slightly higher for the back side of UD-NCF, but no clear influence of the fiber orientation is observed. For Biax-NCF, no clear trend is observed for the ply's side or orientation.
Overall, Biax-NCF has a slightly higher CoF in the TP-tests despite the similar surface architecture of both fabrics. The measured dynamic CoF for both materials remains relatively small for all tests $\mu_d < 0.17$, which agrees with measurements of other dry textiles [3,9,11,13].

<figure><img src="image_519.png" /><figcaption>Fig. 4. Experimental results of the tool-ply (TP) dynamic CoF | (a) Influence of the applied mass $\mathbf{m}_i$ and velocity $\mathbf{v}_i$ in a $0^\circ_f$-configuration for both materials; (b) Influence of the relative interface orientations with an applied mass of $\mathbf{m} = \mathbf{m}_2$ and velocity $\mathbf{v} = \mathbf{v}_1$. The error bars indicate the standard deviation of all tests.</figcaption></figure>

**Ply-Ply.** The influence of the sled weight and velocity on the ply-ply behavior are investigated for the $0^\circ_f/0^\circ_f$-configuration of UD-NCF, cf. Figure 5, under the assumption of a similar behavior for the $0^\circ_f/0^\circ_f$-configuration of Biax-NCF. The dynamic CoF is independent of the sled velocity but decreases for an increasing normal pressure. The higher sled weight flattens the roving and stitching, reducing shock effects, which is reflected in the lower amplitude of the observed oscillations as indicated by the standard deviation [2,9].
Overall, Biax-NCF has a slightly higher CoF in the TP-tests despite the similar surface architecture of both fabrics. The measured dynamic CoF for both materials remains relatively small for all tests $\mu_d < 0.17$, which agrees with measurements of other dry textiles [3,9,11,13].

<figure><img src="image_530.png" /><figcaption>Fig. 4. Experimental results of the tool-ply (TP) dynamic CoF | (a) Influence of the applied mass $\mathbf{m}_i$ and velocity $\mathbf{v}_i$ in a $0^\circ_f$-configuration for both materials; (b) Influence of the relative interface orientations with an applied mass of $\mathbf{m} = \mathbf{m}_2$ and velocity $\mathbf{v} = \mathbf{v}_1$. The error bars indicate the standard deviation of all tests.</figcaption></figure>

**Ply-Ply.** The influence of the sled weight and velocity on the ply-ply behavior are investigated for the $0^\circ_f/0^\circ_f$-configuration of UD-NCF, cf. Figure 5, under the assumption of a similar behavior for the $0^\circ_f/0^\circ_f$-configuration of Biax-NCF. The dynamic CoF is independent of the sled velocity but decreases for an increasing normal pressure. The higher sled weight flattens the roving and stitching, reducing shock effects, which is reflected in the lower amplitude of the observed oscillations as indicated by the standard deviation [2,9].
<figure><img src="image_520.png" /><figcaption>Fig. 5. Influence of the applied mass $m_i$ and velocity $v_i$ on the dynamic CoF for the ply-ply behavior of UD-NCF in a $0^\circ_f/0^\circ_f$-configuration. The error bars indicate the standard deviation of all tests.</figcaption></figure>

The influence of the fabric side and relative fiber orientation at the interface on the PP behavior of both materials is investigated at a constant sled weight $m_2$ and velocity $v_1$, cf. Figure 6.

First, configurations with different sides of the fabrics and the same fiber orientations at the interface are compared, i.e. $0^\circ_f/0^\circ_b$, $0^\circ_f/90^\circ_b$ and $90^\circ_f/90^\circ_b$. The dynamic CoF decreases for configurations with the back side of one or two plies at the interface for both fabrics (e.g. $0^\circ_f/0^\circ_f > 0^\circ_f/0^\circ_b > 0^\circ_b/0^\circ_b$). This effect is observed independent of the relative fiber orientation at the interface. The stitching on the front side of the fabric with the tricot pattern covers a larger area of the ply and thus offers more opportunities for interlocking with another ply, presumably causing the higher $\mu_d$. A particularly high amplitude of the oscillations is measured for the $0^\circ_b/0^\circ_b$-configuration of Biax-NCF, which is presumably caused by significant shock effects due to the linear stitching pattern on the back side. This behavior is not observed for UD-NCF because the linear stitch pattern is recessed between the rovings instead of crossing rovings on the back as for Biax-NCF, cf. Fig. 2.

Second, configurations with similar sides of the fabric and different fiber orientations are compared, i.e. $\square^\circ_f/\square^\circ_f$ for both materials and additional $\square^\circ_f/\square^\circ_b$ for UD-NCF. A decrease of $\mu_d$ is measured for more $90^\circ$ plies when both tricot stitch patterns on the front are in contact (e.g. $0^\circ_f/0^\circ_f > 0^\circ_f/90^\circ_f > 90^\circ_f/90^\circ_f$). This effect is notably stronger for Biax-NCF and only very small for UD-NCF. It is not observed in the $\square^\circ_f/\square^\circ_b$-configurations of UD-NCF.

Overall, the CoF for PP-contact is significantly higher compared to the TP-contact and strongly influenced by the specific configuration of the stitching at the interface.
<figure><img src="image_531.png" /><figcaption>Fig. 5. Influence of the applied mass $m_i$ and velocity $v_i$ on the dynamic CoF for the ply-ply behavior of UD-NCF in a $0^\circ_f/0^\circ_f$-configuration. The error bars indicate the standard deviation of all tests.</figcaption></figure>

The influence of the fabric side and relative fiber orientation at the interface on the PP behavior of both materials is investigated at a constant sled weight $m_2$ and velocity $v_1$, cf. Figure 6.

First, configurations with different sides of the fabrics and the same fiber orientations at the interface are compared, i.e. $0^\circ_f/0^\circ_b$, $0^\circ_f/90^\circ_b$ and $90^\circ_f/90^\circ_b$. The dynamic CoF decreases for configurations with the back side of one or two plies at the interface for both fabrics (e.g. $0^\circ_f/0^\circ_f > 0^\circ_f/0^\circ_b > 0^\circ_b/0^\circ_b$). This effect is observed independent of the relative fiber orientation at the interface. The stitching on the front side of the fabric with the tricot pattern covers a larger area of the ply and thus offers more opportunities for interlocking with another ply, presumably causing the higher $\mu_d$. A particularly high amplitude of the oscillations is measured for the $0^\circ_b/0^\circ_b$-configuration of Biax-NCF, which is presumably caused by significant shock effects due to the linear stitching pattern on the back side. This behavior is not observed for UD-NCF because the linear stitch pattern is recessed between the rovings instead of crossing rovings on the back as for Biax-NCF, cf. Fig. 2.

Second, configurations with similar sides of the fabric and different fiber orientations are compared, i.e. $\square^\circ_f/\square^\circ_f$ for both materials and additional $\square^\circ_f/\square^\circ_b$ for UD-NCF. A decrease of $\mu_d$ is measured for more $90^\circ$ plies when both tricot stitch patterns on the front are in contact (e.g. $0^\circ_f/0^\circ_f > 0^\circ_f/90^\circ_f > 90^\circ_f/90^\circ_f$). This effect is notably stronger for Biax-NCF and only very small for UD-NCF. It is not observed in the $\square^\circ_f/\square^\circ_b$-configurations of UD-NCF.

Overall, the CoF for PP-contact is significantly higher compared to the TP-contact and strongly influenced by the specific configuration of the stitching at the interface.
<figure><img src="image_521.png" /><figcaption>Fig. 6. Influence of the relative interface orientations on the dynamic CoF for the ply-ply behavior with an applied mass of **m** = **m**₂ and velocity **v** = **v**₁. The error bars indicate the standard deviation of all tests.</figcaption></figure>

## Summary and discussion

The friction behavior of a UD- and a Biax-NCF is investigated for different configurations at the tool-ply as well as ply-ply interfaces. A comprehensive analysis is conducted based on the dynamic CoF to evaluate the influences of the applied normal forces, relative sled velocity, fabric side and relative fiber orientation at the interfaces.

The inter-ply behavior of UD- and Biax-NCFs is significantly different for ply-ply (PP) compared to tool-ply (TP) contact. The tangential friction is rate-independent for both cases and significantly higher CoF are measured during PP-contact compared to TP-contact. For an increased sled weight, the CoF increases slightly for TP-contact and notably decreases for PP-contact. This effect can be explained for PP-contact due to flattening of the rovings and stitching for increased normal pressure. The positive pressure-dependency during TP-contact is also observed for other engineering textiles at low pressure [3,16,17], but no conclusive reason is known.

The relative fiber orientation between two plies has no significant influence, except for a slight reduction of $\mu_d$ for Biax-NCF, in contrast to the behavior of woven fabrics [6,10,13,19]. Instead, the most relevant factor for the PP-behavior of both materials is the architecture of the contacting side of the fabric. Configurations with plies in contact with the zigzag pattern of the tricot stitching on the front have a notably higher CoF due to the increased opportunities for interlocking. Additionally, the stitching pattern results in strong oscillations in the measured behavior due to stick-slip and shock effects, similar to the observations for undulations in woven fabrics. The measured $\mu_d$ is in all tests higher for Biax-NCF than UD-NCF despite similar materials and areal weight of the constituents in the respective directions. This potentially results from the smaller stitching pitch of the investigated Biax-NCF, but would require further studies for a conclusive claim.

The results of this study allow a better understanding of the relevant factors influencing the friction behavior of both investigated non-crimp fabrics. They can be used in forming simulation models to analyze the impact on the resulting slippage between individual plies and between plies and tools.
<figure><img src="image_532.png" /><figcaption>Fig. 6. Influence of the relative interface orientations on the dynamic CoF for the ply-ply behavior with an applied mass of **m** = **m**₂ and velocity **v** = **v**₁. The error bars indicate the standard deviation of all tests.</figcaption></figure>

## Summary and discussion

The friction behavior of a UD- and a Biax-NCF is investigated for different configurations at the tool-ply as well as ply-ply interfaces. A comprehensive analysis is conducted based on the dynamic CoF to evaluate the influences of the applied normal forces, relative sled velocity, fabric side and relative fiber orientation at the interfaces.

The inter-ply behavior of UD- and Biax-NCFs is significantly different for ply-ply (PP) compared to tool-ply (TP) contact. The tangential friction is rate-independent for both cases and significantly higher CoF are measured during PP-contact compared to TP-contact. For an increased sled weight, the CoF increases slightly for TP-contact and notably decreases for PP-contact. This effect can be explained for PP-contact due to flattening of the rovings and stitching for increased normal pressure. The positive pressure-dependency during TP-contact is also observed for other engineering textiles at low pressure [3,16,17], but no conclusive reason is known.

The relative fiber orientation between two plies has no significant influence, except for a slight reduction of $\mu_d$ for Biax-NCF, in contrast to the behavior of woven fabrics [6,10,13,19]. Instead, the most relevant factor for the PP-behavior of both materials is the architecture of the contacting side of the fabric. Configurations with plies in contact with the zigzag pattern of the tricot stitching on the front have a notably higher CoF due to the increased opportunities for interlocking. Additionally, the stitching pattern results in strong oscillations in the measured behavior due to stick-slip and shock effects, similar to the observations for undulations in woven fabrics. The measured $\mu_d$ is in all tests higher for Biax-NCF than UD-NCF despite similar materials and areal weight of the constituents in the respective directions. This potentially results from the smaller stitching pitch of the investigated Biax-NCF, but would require further studies for a conclusive claim.

The results of this study allow a better understanding of the relevant factors influencing the friction behavior of both investigated non-crimp fabrics. They can be used in forming simulation models to analyze the impact on the resulting slippage between individual plies and between plies and tools.
## Acknowledgments

The authors would like to thank the Deutsche Forschungsgemeinschaft (DFG, German Research Foundation) and the French National Research Agency (ANR) for funding the collaborative project "Composite forming simulation for non-crimp fabrics based on generalized continuum approaches" (DFG: 431354059, ANR: ANR-19-CE06-0031-012), which the presented work is carried out for. This work is also part of the Heisenberg project "Digitalization of fiber-reinforced polymer processes for resource-efficient manufacturing of lightweight components", funded by the DFG (project no. 798455807141).

## References

[1] W.-R. Yu, P. Harrison, A. Long, Finite element forming simulation for non-crimp fabrics using a non-orthogonal constitutive equation, Composites Part A: Applied Science and Manufacturing 36 (2005) 1079-1093. https://doi.org/10.1016/j.compositesa.2005.01.007

[2] G.D. Lawrence, S. Chen, N.A. Warrior, L.T. Harper, The influence of inter-ply friction during double-diaphragm forming of biaxial NCFs, Composites Part A: Applied Science and Manufacturing 167 (2023) 107426. https://doi.org/10.1016/j.compositesa.2023.107426

[3] U. Sachs, R. Akkerman, K. Fetfatsidis, E. Vidal-Sallé, J. Schumacher, G. Ziegmann, S. Allaoui, G. Hivet, B. Maron, K. Vanblooster, S.V. Lomov, Characterization of the dynamic friction of woven fabrics, Composites Part A: Applied Science and Manufacturing 67 (2014) 289-298. https://doi.org/10.1016/j.compositesa.2014.08.026

[4] ASTM Standard, Test Method for Static and Kinetic Coefficients of Friction of Plastic Film and Sheeting: D1894-14, ASTM International, West Conshohocken, PA, 2014

[5] N. Fulleringer, J.-F. Bloch, Forced stick-slip oscillations allow the measurement of the friction force: Application to paper materials, Tribology International 91 (2015) 94-98. https://doi.org/10.1016/j.triboint.2015.06.021

[6] S. Allaoui, G. Hivet, A. Wendling, P. Ouagne, D. Soulat, Influence of the dry woven fabrics meso-structure on fabric/fabric contact behavior, Journal of Composite Materials 46 (2012) 627-639. https://doi.org/10.1177/0021998311424627

[7] B. Cornelissen, U. Sachs, B. Rietman, R. Akkerman, Dry friction characterisation of carbon fibre tow and satin weave fabric for composite applications, Composites Part A: Applied Science and Manufacturing 56 (2014) 127-135. https://doi.org/10.1016/j.compositesa.2013.10.006

[8] S. Allaoui, C. Cellard, G. Hivet, Effect of inter-ply sliding on the quality of multilayer interlock dry fabric preforms, Composites Part A: Applied Science and Manufacturing 68 (2015) 336-345. https://doi.org/10.1016/j.compositesa.2014.10.017

[9] P. Quenzel, H. Kröger, B. Manin, K. Ngoc Vu, T.X. Duong, T. Gries, M. Itskov, R.A. Sauer, Material characterisation of biaxial glass-fibre non-crimp fabrics as a function of ply orientation, stitch pattern, stitch length and stitch tension, Journal of Composite Materials 56 (2022) 3971-3991. https://doi.org/10.1177/00219983221127005

[10] C. Poppe, D. Dörr, F. Kraus, L. Kärger, Experimental and numerical investigation of the contact behavior during FE forming simulation of continuously reinforced composites in wet compression molding, AIP Conference Proceedings 2113 (2019) 020002. https://doi.org/10.1063/1.5112507

[11] J. Hüttl, F. Albrecht, C. Poppe, F. Lorenz, B. Thoma, L. Kärger, P. Middendorf, F. Henning, Investigations on friction behaviour and forming simulation of plain woven fabrics for wet compression moulding, Proceedings of Sampe Europe, Stuttgart (2017).
## Acknowledgments

The authors would like to thank the Deutsche Forschungsgemeinschaft (DFG, German Research Foundation) and the French National Research Agency (ANR) for funding the collaborative project "Composite forming simulation for non-crimp fabrics based on generalized continuum approaches" (DFG: 431354059, ANR: ANR-19-CE06-0031-012), which the presented work is carried out for. This work is also part of the Heisenberg project "Digitalization of fiber-reinforced polymer processes for resource-efficient manufacturing of lightweight components", funded by the DFG (project no. 798455807141).

## References

[1] W.-R. Yu, P. Harrison, A. Long, Finite element forming simulation for non-crimp fabrics using a non-orthogonal constitutive equation, Composites Part A: Applied Science and Manufacturing 36 (2005) 1079-1093. https://doi.org/10.1016/j.compositesa.2005.01.007

[2] G.D. Lawrence, S. Chen, N.A. Warrior, L.T. Harper, The influence of inter-ply friction during double-diaphragm forming of biaxial NCFs, Composites Part A: Applied Science and Manufacturing 167 (2023) 107426. https://doi.org/10.1016/j.compositesa.2023.107426

[3] U. Sachs, R. Akkerman, K. Fetfatsidis, E. Vidal-Sallé, J. Schumacher, G. Ziegmann, S. Allaoui, G. Hivet, B. Maron, K. Vanblooster, S.V. Lomov, Characterization of the dynamic friction of woven fabrics, Composites Part A: Applied Science and Manufacturing 67 (2014) 289-298. https://doi.org/10.1016/j.compositesa.2014.08.026

[4] ASTM Standard, Test Method for Static and Kinetic Coefficients of Friction of Plastic Film and Sheeting: D1894-14, ASTM International, West Conshohocken, PA, 2014

[5] N. Fulleringer, J.-F. Bloch, Forced stick-slip oscillations allow the measurement of the friction force: Application to paper materials, Tribology International 91 (2015) 94-98. https://doi.org/10.1016/j.triboint.2015.06.021

[6] S. Allaoui, G. Hivet, A. Wendling, P. Ouagne, D. Soulat, Influence of the dry woven fabrics meso-structure on fabric/fabric contact behavior, Journal of Composite Materials 46 (2012) 627-639. https://doi.org/10.1177/0021998311424627

[7] B. Cornelissen, U. Sachs, B. Rietman, R. Akkerman, Dry friction characterisation of carbon fibre tow and satin weave fabric for composite applications, Composites Part A: Applied Science and Manufacturing 56 (2014) 127-135. https://doi.org/10.1016/j.compositesa.2013.10.006

[8] S. Allaoui, C. Cellard, G. Hivet, Effect of inter-ply sliding on the quality of multilayer interlock dry fabric preforms, Composites Part A: Applied Science and Manufacturing 68 (2015) 336-345. https://doi.org/10.1016/j.compositesa.2014.10.017

[9] P. Quenzel, H. Kröger, B. Manin, K. Ngoc Vu, T.X. Duong, T. Gries, M. Itskov, R.A. Sauer, Material characterisation of biaxial glass-fibre non-crimp fabrics as a function of ply orientation, stitch pattern, stitch length and stitch tension, Journal of Composite Materials 56 (2022) 3971-3991. https://doi.org/10.1177/00219983221127005

[10] C. Poppe, D. Dörr, F. Kraus, L. Kärger, Experimental and numerical investigation of the contact behavior during FE forming simulation of continuously reinforced composites in wet compression molding, AIP Conference Proceedings 2113 (2019) 020002. https://doi.org/10.1063/1.5112507

[11] J. Hüttl, F. Albrecht, C. Poppe, F. Lorenz, B. Thoma, L. Kärger, P. Middendorf, F. Henning, Investigations on friction behaviour and forming simulation of plain woven fabrics for wet compression moulding, Proceedings of Sampe Europe, Stuttgart (2017).
[12] F. Yu, S. Chen, L.T. Harper, N.A. Warrior, Investigation into the effects of inter-ply sliding during double diaphragm forming for multi-layered biaxial non-crimp fabrics, Composites Part A: Applied Science and Manufacturing 150 (2021) 106611. https://doi.org/10.1016/j.compositesa.2021.106611

[13] F.N. Nezami, T. Gereke, C. Cherif, Analyses of interaction mechanisms during forming of multilayer carbon woven fabrics for composite applications, Composites Part A: Applied Science and Manufacturing 84 (2016) 406-416. https://doi.org/10.1016/j.compositesa.2016.02.023

[14] K.A. Fetfatsidis, D. Jauffrès, J.A. Sherwood, J. Chen, Characterization of the tool/fabric and fabric/fabric friction for woven-fabric composites during the thermostamping process, International Journal of Material Forming 6 (2013) 209-221. https://doi.org/10.1007/s12289-011-1072-5

[15] W. Najjar, C. Pupin, X. Legrand, S. Boude, D. Soulat, P. Dal Santo, Analysis of frictional behaviour of carbon dry woven reinforcement, Journal of Reinforced Plastics and Composites 33 (2014) 1037-1047. https://doi.org/10.1177/0731684414521670

[16] E.I. Avgoulas, D.M. Mulvihill, A. Endruweit, M.P.F. Sutcliffe, N.A. Warrior, D.S.A. de Focatiis, A.C. Long, Frictional behaviour of non-crimp fabrics (NCFs) in contact with a forming tool, Tribology International 121 (2018) 71-77. https://doi.org/10.1016/j.triboint.2018.01.026

[17] Y. Liu, Z. Xiang, Z. Wu, X. Hu, Experimental investigation of friction behaviors of glass-fiber woven fabric, Textile Research Journal 93 (2023) 18-32. https://doi.org/10.1177/00405175221115468

[18] L. Montero, S. Allaoui, G. Hivet, Characterisation of the mesoscopic and macroscopic friction behaviours of glass plain weave reinforcement, Composites Part A: Applied Science and Manufacturing 95 (2017) 257-266. https://doi.org/10.1016/j.compositesa.2017.01.022

[19] E. Vidal-Sallé, F. Massi, Friction Measurement on Dry Fabric for Forming Simulation of Composite Reinforcement, KEM 504-506 (2012) 319-324. https://doi.org/10.4028/www.scientific.net/KEM.504-506.319
[12] F. Yu, S. Chen, L.T. Harper, N.A. Warrior, Investigation into the effects of inter-ply sliding during double diaphragm forming for multi-layered biaxial non-crimp fabrics, Composites Part A: Applied Science and Manufacturing 150 (2021) 106611. https://doi.org/10.1016/j.compositesa.2021.106611

[13] F.N. Nezami, T. Gereke, C. Cherif, Analyses of interaction mechanisms during forming of multilayer carbon woven fabrics for composite applications, Composites Part A: Applied Science and Manufacturing 84 (2016) 406-416. https://doi.org/10.1016/j.compositesa.2016.02.023

[14] K.A. Fetfatsidis, D. Jauffrès, J.A. Sherwood, J. Chen, Characterization of the tool/fabric and fabric/fabric friction for woven-fabric composites during the thermostamping process, International Journal of Material Forming 6 (2013) 209-221. https://doi.org/10.1007/s12289-011-1072-5

[15] W. Najjar, C. Pupin, X. Legrand, S. Boude, D. Soulat, P. Dal Santo, Analysis of frictional behaviour of carbon dry woven reinforcement, Journal of Reinforced Plastics and Composites 33 (2014) 1037-1047. https://doi.org/10.1177/0731684414521670

[16] E.I. Avgoulas, D.M. Mulvihill, A. Endruweit, M.P.F. Sutcliffe, N.A. Warrior, D.S.A. de Focatiis, A.C. Long, Frictional behaviour of non-crimp fabrics (NCFs) in contact with a forming tool, Tribology International 121 (2018) 71-77. https://doi.org/10.1016/j.triboint.2018.01.026

[17] Y. Liu, Z. Xiang, Z. Wu, X. Hu, Experimental investigation of friction behaviors of glass-fiber woven fabric, Textile Research Journal 93 (2023) 18-32. https://doi.org/10.1177/00405175221115468

[18] L. Montero, S. Allaoui, G. Hivet, Characterisation of the mesoscopic and macroscopic friction behaviours of glass plain weave reinforcement, Composites Part A: Applied Science and Manufacturing 95 (2017) 257-266. https://doi.org/10.1016/j.compositesa.2017.01.022

[19] E. Vidal-Sallé, F. Massi, Friction Measurement on Dry Fabric for Forming Simulation of Composite Reinforcement, KEM 504-506 (2012) 319-324. https://doi.org/10.4028/www.scientific.net/KEM.504-506.319
# Experimental permeability measurement of different reinforcement types for virtual permeability determination validation

BOUBAKER Mouadh<sup>1,a*</sup>, WIJAYA Willsen<sup>2,b</sup>, CANTAREL Arthur<sup>1,c</sup>, DEBENEST Gérald<sup>3,d</sup>, BICKERTON Simon<sup>2,e</sup>

**Keywords:** Composites, Experimental Permeability, Virtual Permeability, CFD, Porous Media

**Abstract.** Permeability measurement of engineering textiles is a key step in preparing composites manufacturing processes. A radial flow experimental set-up is used in this work to characterize the unsaturated and saturated in-plane permeabilities of different types of textiles. In order to identify fabrics in which the dual-scale flow effect is stronger, comparisons are made between the measured saturating and saturated permeabilities. In parallel, the delayed tow saturation during the oil injection stage of the saturating measurement is observed visually. In addition, virtual permeability of porous media is studied using the numerical implementation of Darcy-Brinkman equation in a finite volume method (FVM) open-source software (OpenFOAM). A numerical method is proposed to determine the permeability of a given geometry at mesoscale. The method is used to determine the permeability of a realistic geometry acquired using an X-ray micro tomography (µCT) scanner and the results are compared to experimental values obtained with the proposed experimental set-up on the same plain weave textile.

**Introduction**

Composite materials are increasingly and widely applied in automobile, shipbuilding, aeronautics and aerospace applications. These materials are fabricated from fibrous reinforcements that can take the form of woven fabrics, non-crimp fabrics, knits, braids or random mats [1]. The fabrics are transformed into composite parts using different manufacturing processes. One of the most used and developed manufacturing methods is the Liquid Composite Molding (LCM) due to its various advantages, being low-cost, clean and more efficient than other methods. This fabrication process is based on the impregnation of a dry reinforcement with a low viscosity resin. During the impregnation step, the liquid resin flows through the channels of the reinforcement which can be considered as a porous medium. This flow is a key phenomenon in the modeling of the manufacturing process and has a significant effect on the quality of the final part. An important parameter to be considered at this step is the permeability of the preform, which quantifies the ability to be crossed by a fluid. The permeability influences the resin flow front advancement inside the pores and as a result it controls the mold filling time which is a crucial parameter of composite fabrication optimization.

In order to meet the growing demand for composites, different types of reinforcements are being developed with different properties and architectures. Therefore, an accurate and efficient method

<figure><img src="image_522.png" /></figure>

<footnote n="1"><sup>1</sup> Institut Clément Ader (ICA), CNRS UMR 5312, University of Toulouse, IUT of Tarbes, UPS, France</footnote>

<footnote n="2"><sup>2</sup> Centre for Advanced Materials Manufacturing and Design, Department of Mechanical Engineering, The University of Auckland, Auckland, New Zealand</footnote>

<footnote n="3"><sup>3</sup> INPT, UPS, IMFT (Institut de Mécanique des Fluides de Toulouse), Université de Toulouse, Allée Camille Soula, 31400, Toulouse, France</footnote>

<footnote><sup>a</sup> MOUADH.BOUBAKER@iut-tarbes.fr, <sup>b</sup> willsen.wijaya@auckland.ac.nz, <sup>c</sup> arthur.cantarel@iut-tarbes.fr, <sup>d</sup> gerald.debenest@toulouse-inp.fr, <sup>e</sup> s.bickerton@auckland.ac.nz</footnote>
# Experimental permeability measurement of different reinforcement types for virtual permeability determination validation

BOUBAKER Mouadh<sup>1,a*</sup>, WIJAYA Willsen<sup>2,b</sup>, CANTAREL Arthur<sup>1,c</sup>, DEBENEST Gérald<sup>3,d</sup>, BICKERTON Simon<sup>2,e</sup>

**Keywords:** Composites, Experimental Permeability, Virtual Permeability, CFD, Porous Media

**Abstract.** Permeability measurement of engineering textiles is a key step in preparing composites manufacturing processes. A radial flow experimental set-up is used in this work to characterize the unsaturated and saturated in-plane permeabilities of different types of textiles. In order to identify fabrics in which the dual-scale flow effect is stronger, comparisons are made between the measured saturating and saturated permeabilities. In parallel, the delayed tow saturation during the oil injection stage of the saturating measurement is observed visually. In addition, virtual permeability of porous media is studied using the numerical implementation of Darcy-Brinkman equation in a finite volume method (FVM) open-source software (OpenFOAM). A numerical method is proposed to determine the permeability of a given geometry at mesoscale. The method is used to determine the permeability of a realistic geometry acquired using an X-ray micro tomography (µCT) scanner and the results are compared to experimental values obtained with the proposed experimental set-up on the same plain weave textile.

**Introduction**

Composite materials are increasingly and widely applied in automobile, shipbuilding, aeronautics and aerospace applications. These materials are fabricated from fibrous reinforcements that can take the form of woven fabrics, non-crimp fabrics, knits, braids or random mats [1]. The fabrics are transformed into composite parts using different manufacturing processes. One of the most used and developed manufacturing methods is the Liquid Composite Molding (LCM) due to its various advantages, being low-cost, clean and more efficient than other methods. This fabrication process is based on the impregnation of a dry reinforcement with a low viscosity resin. During the impregnation step, the liquid resin flows through the channels of the reinforcement which can be considered as a porous medium. This flow is a key phenomenon in the modeling of the manufacturing process and has a significant effect on the quality of the final part. An important parameter to be considered at this step is the permeability of the preform, which quantifies the ability to be crossed by a fluid. The permeability influences the resin flow front advancement inside the pores and as a result it controls the mold filling time which is a crucial parameter of composite fabrication optimization.

In order to meet the growing demand for composites, different types of reinforcements are being developed with different properties and architectures. Therefore, an accurate and efficient method

<figure><img src="image_533.png" /></figure>

<footnote n="1"><sup>1</sup> Institut Clément Ader (ICA), CNRS UMR 5312, University of Toulouse, IUT of Tarbes, UPS, France</footnote>

<footnote n="2"><sup>2</sup> Centre for Advanced Materials Manufacturing and Design, Department of Mechanical Engineering, The University of Auckland, Auckland, New Zealand</footnote>

<footnote n="3"><sup>3</sup> INPT, UPS, IMFT (Institut de Mécanique des Fluides de Toulouse), Université de Toulouse, Allée Camille Soula, 31400, Toulouse, France</footnote>

<footnote><sup>a</sup> MOUADH.BOUBAKER@iut-tarbes.fr, <sup>b</sup> willsen.wijaya@auckland.ac.nz, <sup>c</sup> arthur.cantarel@iut-tarbes.fr, <sup>d</sup> gerald.debenest@toulouse-inp.fr, <sup>e</sup> s.bickerton@auckland.ac.nz</footnote>
to characterize these reinforcements is needed, errors in permeability prediction for example can lead to an inefficient process design and the appearance of defects in parts such as voids and dry spots.

In the scope of this paper, we are interested in characterizing the permeability of a given reinforcement. Different permeability determination methods can be found in the literature and can be classified into three main groups: analytical methods, experimental methods and numerical methods.

Different analytical models relate the permeability to the fibre volume fraction and are based on simplifying assumptions where the fibres are modeled as cylinders arranged in regularly ordered networks (quadratic and hexagonal) [2]. These models are based on idealistic assumptions and do not accurately represent the complexity of the real preforms, in addition they are only valid for a certain range of fibre volume fractions. Alternatively, many researchers used experimental procedures to measure the permeability of fibrous domains. These experiments give more accurate results but have the disadvantage of being time and material-consuming and labor-intensive since they involve many sensors, transducers and digital cameras [3]. Another drawback is the absence of a standard approach. Several international benchmark exercises [4-6] were launched in order to study the variation in the experimental measurements on different geometries (non-crimp and woven fabrics) and found a coefficient of variation of up to 44% for the woven fabric for example.

A promising alternative is the use of numerical methods which combine textile modeling and computational fluid dynamics techniques. The potential of numerical methods is increasing with the increasing computing capabilities. The virtual permeability prediction has many advantages like the ability to explore a large range of parameters efficiently without needing to reproduce the experiment or without using big equipment and generating large amounts of waste. With these advantages come some challenges from the complex structure of reinforcements like the geometrical variability and the dual-scale porosity. This dual-scale porosity makes it necessary to use multiscale approaches in order to predict virtual permeability since macro-scale models are not accurate enough and micro-scale models are not efficient and need too much information. The meso-scale models seem to give a good compromise between accuracy and efficiency.

The flow of resin in a fibrous network is usually considered as a creeping flow because high velocities cannot be achieved which results in low values of the Reynolds number. In this case, Darcy's law (Eq. 1) can be used to model the resin flow. Being based on experimental observations, this law is a phenomenological macroscopic model [7] which gives a relation between the volume-averaged velocity and the volume-averaged pressure gradient and is used to characterize porous media at the macro scale

$$u = -\frac{K}{\mu}\nabla P \qquad (1)$$

u is the volume averaged velocity, K is the permeability tensor of the reinforcement, μ is the viscosity of the fluid and P is the volume-averaged pressure.

In the first section of this paper, a 2D radial experimental procedure to determine the in-plane permeability tensor of different textiles will be presented. The experimental set-up was designed in order to measure the unsaturated (transient) and saturated (steady-state) permeability in a single run. Different types of textiles were explored and one of them was scanned using an X-ray μCT scanner [3] and a virtual permeability prediction method was applied on the 3D virtual geometry obtained. The results of the numerical simulation will be compared to the experimental ones at different fibre volume fractions in order to test the accuracy of the virtual permeability determination method.
to characterize these reinforcements is needed, errors in permeability prediction for example can lead to an inefficient process design and the appearance of defects in parts such as voids and dry spots.

In the scope of this paper, we are interested in characterizing the permeability of a given reinforcement. Different permeability determination methods can be found in the literature and can be classified into three main groups: analytical methods, experimental methods and numerical methods.

Different analytical models relate the permeability to the fibre volume fraction and are based on simplifying assumptions where the fibres are modeled as cylinders arranged in regularly ordered networks (quadratic and hexagonal) [2]. These models are based on idealistic assumptions and do not accurately represent the complexity of the real preforms, in addition they are only valid for a certain range of fibre volume fractions. Alternatively, many researchers used experimental procedures to measure the permeability of fibrous domains. These experiments give more accurate results but have the disadvantage of being time and material-consuming and labor-intensive since they involve many sensors, transducers and digital cameras [3]. Another drawback is the absence of a standard approach. Several international benchmark exercises [4-6] were launched in order to study the variation in the experimental measurements on different geometries (non-crimp and woven fabrics) and found a coefficient of variation of up to 44% for the woven fabric for example.

A promising alternative is the use of numerical methods which combine textile modeling and computational fluid dynamics techniques. The potential of numerical methods is increasing with the increasing computing capabilities. The virtual permeability prediction has many advantages like the ability to explore a large range of parameters efficiently without needing to reproduce the experiment or without using big equipment and generating large amounts of waste. With these advantages come some challenges from the complex structure of reinforcements like the geometrical variability and the dual-scale porosity. This dual-scale porosity makes it necessary to use multiscale approaches in order to predict virtual permeability since macro-scale models are not accurate enough and micro-scale models are not efficient and need too much information. The meso-scale models seem to give a good compromise between accuracy and efficiency.

The flow of resin in a fibrous network is usually considered as a creeping flow because high velocities cannot be achieved which results in low values of the Reynolds number. In this case, Darcy's law (Eq. 1) can be used to model the resin flow. Being based on experimental observations, this law is a phenomenological macroscopic model [7] which gives a relation between the volume-averaged velocity and the volume-averaged pressure gradient and is used to characterize porous media at the macro scale

$$u = -\frac{K}{\mu}\nabla P \qquad (1)$$

u is the volume averaged velocity, K is the permeability tensor of the reinforcement, μ is the viscosity of the fluid and P is the volume-averaged pressure.

In the first section of this paper, a 2D radial experimental procedure to determine the in-plane permeability tensor of different textiles will be presented. The experimental set-up was designed in order to measure the unsaturated (transient) and saturated (steady-state) permeability in a single run. Different types of textiles were explored and one of them was scanned using an X-ray μCT scanner [3] and a virtual permeability prediction method was applied on the 3D virtual geometry obtained. The results of the numerical simulation will be compared to the experimental ones at different fibre volume fractions in order to test the accuracy of the virtual permeability determination method.
## Experimental set-up for the permeability characterisation of fibrous textiles

A 2D radial in-plane permeability determination experimental set-up (Fig. 1) developed at the Centre for Advanced Materials Manufacturing and Design CAMMD of The University of Auckland is used to determine the unsaturated and saturated permeability of 5 different E-glass fabrics: quad axial stitched fabric, twill weave fabric, 0/90 biaxial stitched fabric, unidirectional reinforcement and a plain weave fabric. The experimental facility including the different tools used and the details of the experimental procedure will be presented and the results and the key findings will be discussed.

<figure><img src="image_523.png" /><figcaption>Fig. 1: Schematic representation of the experimental facility [3]</figcaption></figure>

This setup was used by Zhang and Comas-Cardona et al. [8,9] to investigate transient permeability of Chopped Strand Mat (CSM) and other fabrics. This 2D in-plane radial flow permeability test rig has been developed over the years in order to efficiently characterize the in-plane permeability of fibrous textiles. This rig complies with the international standards found in the literature [4-6,10].

The experimental procedure involves the compaction between two rigid flat plates of the textile in order to get the desired fibre volume fraction through the control of the sample thickness. It involves also the injection of a test fluid via a circular port of a diameter of 15 mm located at the center of the sample, the test fluid used is a Newtonian Mobil DTE series HEAVY mineral oil ISO VG 100 which viscosity as a function of temperature was determined using a Parr Physica UDS200 rheometer. This mineral oil was used instead of a thermoset resin since the resin cure phase of the LCM does not affect the permeability measurements. The oil is supplied from an enclosed metallic pot driven by the pressurized air regulated by a FESTO proportional pressure controller. After being injected, the fluid will flow radially in the textile.

The radial flow technique was particularly chosen because of the relative simplicity of the setup compared to other techniques in addition to the ability to compute the complete in-plane unsaturated permeability tensor with a single test which minimizes the effort and the cost.

## Experimental facility

The 2D radial testing facility is placed on an Instron 1186 Universal Testing Machine (UTM) that provides an accurate control of cavity thickness. The whole facility was assembled in an environmentally controlled (a constant temperature of 22°C) room to get consistent measurements. The sample is set between two plates: a glass lower plate made of 55 mm thick glass with a side length of 350 mm that allows flow front detection through optical monitoring in order to better understand the flow behavior, and a circular upper platen of a diameter of 250 mm and 30 mm thick that is aligned to the lower platen with a spherical alignment unit and a manufactured strain-
## Experimental set-up for the permeability characterisation of fibrous textiles

A 2D radial in-plane permeability determination experimental set-up (Fig. 1) developed at the Centre for Advanced Materials Manufacturing and Design CAMMD of The University of Auckland is used to determine the unsaturated and saturated permeability of 5 different E-glass fabrics: quad axial stitched fabric, twill weave fabric, 0/90 biaxial stitched fabric, unidirectional reinforcement and a plain weave fabric. The experimental facility including the different tools used and the details of the experimental procedure will be presented and the results and the key findings will be discussed.

<figure><img src="image_534.png" /><figcaption>Fig. 1: Schematic representation of the experimental facility [3]</figcaption></figure>

This setup was used by Zhang and Comas-Cardona et al. [8,9] to investigate transient permeability of Chopped Strand Mat (CSM) and other fabrics. This 2D in-plane radial flow permeability test rig has been developed over the years in order to efficiently characterize the in-plane permeability of fibrous textiles. This rig complies with the international standards found in the literature [4-6,10].

The experimental procedure involves the compaction between two rigid flat plates of the textile in order to get the desired fibre volume fraction through the control of the sample thickness. It involves also the injection of a test fluid via a circular port of a diameter of 15 mm located at the center of the sample, the test fluid used is a Newtonian Mobil DTE series HEAVY mineral oil ISO VG 100 which viscosity as a function of temperature was determined using a Parr Physica UDS200 rheometer. This mineral oil was used instead of a thermoset resin since the resin cure phase of the LCM does not affect the permeability measurements. The oil is supplied from an enclosed metallic pot driven by the pressurized air regulated by a FESTO proportional pressure controller. After being injected, the fluid will flow radially in the textile.

The radial flow technique was particularly chosen because of the relative simplicity of the setup compared to other techniques in addition to the ability to compute the complete in-plane unsaturated permeability tensor with a single test which minimizes the effort and the cost.

## Experimental facility

The 2D radial testing facility is placed on an Instron 1186 Universal Testing Machine (UTM) that provides an accurate control of cavity thickness. The whole facility was assembled in an environmentally controlled (a constant temperature of 22°C) room to get consistent measurements. The sample is set between two plates: a glass lower plate made of 55 mm thick glass with a side length of 350 mm that allows flow front detection through optical monitoring in order to better understand the flow behavior, and a circular upper platen of a diameter of 250 mm and 30 mm thick that is aligned to the lower platen with a spherical alignment unit and a manufactured strain-
gauged compressive alignment specimen. Through these alignment tools, parallelism and uniform cavity thickness are achieved with a tolerance of 0.005 mm. These plates are designed in order to minimize deflection during the compaction step allowing a maximum deflection of 0.07 mm under an average compaction pressure of 2 MPa [9], the range of target thickness cavity was around 3mm which respects the ISO 4410 recommendations of target thickness being under 2%.

A Navitar (SV5C10 0238) digital camera was attached to the set-up via an L-bend with its lens pointing to an angled mirror of 45° with respect to the horizontal plate used to capture the flow front evolution during the test fluid infusion. Two touch sensors of type Mitutoyo IDU25 with 0.01 mm resolution were used to measure the distance between the parallel plates; they were attached to both sides of the top plate with their probe touching the upper surface of the bottom plate. The different measurement accuracies of the experimental facility are summarized in the following table:

<table><caption>Table 1: Measurement accuracies</caption><thead><tr><td>Measurement</td><td>Accuracy</td></tr></thead><tbody><tr><td>Sample mass</td><td>±0.01 g</td></tr><tr><td>Injected oil mass</td><td>±0.1 g</td></tr><tr><td>Cavity thickness</td><td>±0.005 mm</td></tr><tr><td>Fluid pressure</td><td>±1%</td></tr></tbody></table>

## Experimental procedure

Using a dedicated circular cutting knife in a hydraulic press machine, circular stacks of 200 mm diameter of the fibrous textiles were prepared. A hole of 15 mm diameter was punched at the center of each sample in order to serve as an injection port. The mass of each sample was measured before starting the experiment using a digital balance. Then, the sample is placed on the bottom glass plate. The placement must be accurately centered in order to have a pure in-plane flow during the injection step. The position of this bottom plate is then monitored in order to achieve the target fibre volume fraction $V_f$ given by equation (Eq. 2)

$$V_f = \frac{m}{\rho A h} \qquad (2)$$

with *m* the mass of the sample, *ρ* is the fibre density (characteristic of the used fabric), *A* the sample's planar area and *h* is the target compaction thickness.

The oil is then injected at a fixed nominal injection pressure of 50 kPa and the flow front progression is recorded as a series of images at the frequency of 1 Hz (Fig. 2).

<figure><img src="image_524.png" /><figcaption>Fig. 2: Flow front progression with time.</figcaption></figure>

In addition to the flow front levels, other parameters need to be recorded during the experimental run in order to be used in the permeability computation using the formula given by
gauged compressive alignment specimen. Through these alignment tools, parallelism and uniform cavity thickness are achieved with a tolerance of 0.005 mm. These plates are designed in order to minimize deflection during the compaction step allowing a maximum deflection of 0.07 mm under an average compaction pressure of 2 MPa [9], the range of target thickness cavity was around 3mm which respects the ISO 4410 recommendations of target thickness being under 2%.

A Navitar (SV5C10 0238) digital camera was attached to the set-up via an L-bend with its lens pointing to an angled mirror of 45° with respect to the horizontal plate used to capture the flow front evolution during the test fluid infusion. Two touch sensors of type Mitutoyo IDU25 with 0.01 mm resolution were used to measure the distance between the parallel plates; they were attached to both sides of the top plate with their probe touching the upper surface of the bottom plate. The different measurement accuracies of the experimental facility are summarized in the following table:

<table><caption>Table 1: Measurement accuracies</caption><thead><tr><td>Measurement</td><td>Accuracy</td></tr></thead><tbody><tr><td>Sample mass</td><td>±0.01 g</td></tr><tr><td>Injected oil mass</td><td>±0.1 g</td></tr><tr><td>Cavity thickness</td><td>±0.005 mm</td></tr><tr><td>Fluid pressure</td><td>±1%</td></tr></tbody></table>

## Experimental procedure

Using a dedicated circular cutting knife in a hydraulic press machine, circular stacks of 200 mm diameter of the fibrous textiles were prepared. A hole of 15 mm diameter was punched at the center of each sample in order to serve as an injection port. The mass of each sample was measured before starting the experiment using a digital balance. Then, the sample is placed on the bottom glass plate. The placement must be accurately centered in order to have a pure in-plane flow during the injection step. The position of this bottom plate is then monitored in order to achieve the target fibre volume fraction $V_f$ given by equation (Eq. 2)

$$V_f = \frac{m}{\rho A h} \qquad (2)$$

with *m* the mass of the sample, *ρ* is the fibre density (characteristic of the used fabric), *A* the sample's planar area and *h* is the target compaction thickness.

The oil is then injected at a fixed nominal injection pressure of 50 kPa and the flow front progression is recorded as a series of images at the frequency of 1 Hz (Fig. 2).

<figure><img src="image_535.png" /><figcaption>Fig. 2: Flow front progression with time.</figcaption></figure>

In addition to the flow front levels, other parameters need to be recorded during the experimental run in order to be used in the permeability computation using the formula given by
Weitzenbock et al. [11] in equations (Eq. 3 and Eq. 4). These parameters are the static inlet pressure measured by a pressure sensor located at the inlet and the oil temperature measured by a thermo-couple in the pressure pot.

$$K_{11} = (a^2[2\ln(\frac{a}{r_i}) - 1] + r_i^2) \frac{\mu(1-V_f)}{4tP} \quad (3)$$

$$K_{22} = (b^2[2\ln(\frac{b}{r_i}) - 1] + r_i^2) \frac{\mu(1-V_f)}{4tP} \quad (4)$$

where $K_{11}$ and $K_{22}$ are respectively the major and minor in-plane principal permeability components, $a$ and $b$ are the major and minor radii of the fitted flow front ellipse, $r_i$ is the radius of the injection hole, $P$ the total pressure drop which can be approximated to the inlet pressure based on the atmospheric flow front pressure assumption, $\mu$ is the test oil viscosity deduced from the measured oil temperature and finally $t$ is the time from the start of the injection. For each test, the permeability tensor is computed for each flow front image and plotted against the time steps. The reported permeability values are given by the average of the final one-third of the computed values to eliminate any flow stabilization effects at the start of the experiment. This setup also allows the prediction of isotropic saturated permeability. An additional quantity of oil is injected inside the sample after being fully saturated, which gives the saturated mass flow rate (by fitting a linear regression to the mass-time curve) and the steady-state permeability is deduced using Darcy's law through this equation (Eq. 5) [12]

$$K_{\text{saturated}} = \frac{Q \cdot \mu \cdot \ln \left( \frac{r_0}{r_i} \right)}{2\pi \cdot \rho \cdot h \cdot P_{\text{inj}}} \quad (5)$$

where $Q$ is the mass flow rate, $\mu$ is the oil viscosity, $r_0$ is the sample's outer radius, $r_i$ is the internal radius taken as the radius of the punched hole, $\rho$ is the oil density, $h$ is the cavity thickness and $P_{\text{inj}}$ is the injection pressure.

## Results

These experiments were carried out in order to observe the flow front advancement for different types of textiles used to manufacture composite parts and to quantify the difference between the transient and steady-state permeabilities of these textiles. The tests were carried out at three different fibre volume fractions $V_f$ (0.5, 0.55 and 0.6) except for the twill-weave fabric for which only two $V_f$ were targeted. The layers were oriented in a way that gives an isotropic flow in both directions (the flow fronts are almost circular) and the unsaturated permeability used for the comparison with the saturated value is an average of the values found for the principal directions $K_{\text{unsat}} = \sqrt{K_{11} \cdot K_{22}}$. The ratios $K_{\text{unsat}}/K_{\text{sat}}$ were found to vary between 0.9 and 1.42 which lies in the range found in literature [13,14]. Researchers found ratios between 0.25, and 4 but for most of them [15,16] the ratio was around 1. Table 2 gives an overview of the results, each test was repeated 3 to 4 times and the relative standard deviation was between 3% and 16% which gives confidence on the repeatability of our results.
Weitzenbock et al. [11] in equations (Eq. 3 and Eq. 4). These parameters are the static inlet pressure measured by a pressure sensor located at the inlet and the oil temperature measured by a thermo-couple in the pressure pot.

$$K_{11} = (a^2[2\ln(\frac{a}{r_i}) - 1] + r_i^2) \frac{\mu(1-V_f)}{4tP} \quad (3)$$

$$K_{22} = (b^2[2\ln(\frac{b}{r_i}) - 1] + r_i^2) \frac{\mu(1-V_f)}{4tP} \quad (4)$$

where $K_{11}$ and $K_{22}$ are respectively the major and minor in-plane principal permeability components, $a$ and $b$ are the major and minor radii of the fitted flow front ellipse, $r_i$ is the radius of the injection hole, $P$ the total pressure drop which can be approximated to the inlet pressure based on the atmospheric flow front pressure assumption, $\mu$ is the test oil viscosity deduced from the measured oil temperature and finally $t$ is the time from the start of the injection. For each test, the permeability tensor is computed for each flow front image and plotted against the time steps. The reported permeability values are given by the average of the final one-third of the computed values to eliminate any flow stabilization effects at the start of the experiment. This setup also allows the prediction of isotropic saturated permeability. An additional quantity of oil is injected inside the sample after being fully saturated, which gives the saturated mass flow rate (by fitting a linear regression to the mass-time curve) and the steady-state permeability is deduced using Darcy's law through this equation (Eq. 5) [12]

$$K_{\text{saturated}} = \frac{Q \cdot \mu \cdot \ln \left( \frac{r_0}{r_i} \right)}{2\pi \cdot \rho \cdot h \cdot P_{\text{inj}}} \quad (5)$$

where $Q$ is the mass flow rate, $\mu$ is the oil viscosity, $r_0$ is the sample's outer radius, $r_i$ is the internal radius taken as the radius of the punched hole, $\rho$ is the oil density, $h$ is the cavity thickness and $P_{\text{inj}}$ is the injection pressure.

## Results

These experiments were carried out in order to observe the flow front advancement for different types of textiles used to manufacture composite parts and to quantify the difference between the transient and steady-state permeabilities of these textiles. The tests were carried out at three different fibre volume fractions $V_f$ (0.5, 0.55 and 0.6) except for the twill-weave fabric for which only two $V_f$ were targeted. The layers were oriented in a way that gives an isotropic flow in both directions (the flow fronts are almost circular) and the unsaturated permeability used for the comparison with the saturated value is an average of the values found for the principal directions $K_{\text{unsat}} = \sqrt{K_{11} \cdot K_{22}}$. The ratios $K_{\text{unsat}}/K_{\text{sat}}$ were found to vary between 0.9 and 1.42 which lies in the range found in literature [13,14]. Researchers found ratios between 0.25, and 4 but for most of them [15,16] the ratio was around 1. Table 2 gives an overview of the results, each test was repeated 3 to 4 times and the relative standard deviation was between 3% and 16% which gives confidence on the repeatability of our results.
<table><caption>Table 2: Summary of the experimental results</caption><thead><tr><th>Textile</th><th>V<sub>f</sub></th><th>K<sub>sat</sub>[10<sup>-10</sup>m<sup>2</sup>]</th><th>K<sub>unsat</sub>[10<sup>-10</sup>m<sup>2</sup>]</th><th>K<sub>unsat</sub>/K<sub>sat</sub></th></tr></thead><tbody><tr><td rowspan="3">E-glass Plain weave (800 g/m<sup>2</sup>)</td><td>0.5</td><td>2.916 ± 0.19</td><td>3.16 ± 0.24</td><td>1.08</td></tr><tr><td>0.55</td><td>1.12 ± 0.11</td><td>1.16 ± 0.14</td><td>1.04</td></tr><tr><td>0.6</td><td>0.482 ± 0.06</td><td>0.458 ± 0.054</td><td>0.95</td></tr><tr><td rowspan="3">0/90 Biaxial stitched fabric (825 g/m<sup>2</sup>)</td><td>0.5</td><td>1.39 ± 0.2</td><td>1.6 ± 0.3</td><td>1.15</td></tr><tr><td>0.55</td><td>0.672 ± 0.097</td><td>0.729 ± 0.11</td><td>1.08</td></tr><tr><td>0.6</td><td>0.326 ± 0.044</td><td>0.34 ± 0.011</td><td>1.04</td></tr><tr><td rowspan="3">Quad Axial Stitched Fabric (851 g/m<sup>2</sup>)</td><td>0.5</td><td>0.62 ± 0.026</td><td>0.649 ± 0.038</td><td>1.045</td></tr><tr><td>0.55</td><td>0.27 ± 0.012</td><td>0.286 ± 0.007</td><td>1.05</td></tr><tr><td>0.6</td><td>0.171 ± 0.016</td><td>0.177 ± 0.0178</td><td>1.03</td></tr><tr><td rowspan="3">Unidirectional reinforcement fabrics (850 g/m<sup>2</sup>)</td><td>0.5</td><td>3.5 ± 0.22</td><td>4.97 ± 0.8</td><td>1.42</td></tr><tr><td>0.55</td><td>1.91 ± 0.057</td><td>2.47 ± 0.097</td><td>1.29</td></tr><tr><td>0.6</td><td>1.12 ± 0.16</td><td>1.33 ± 0.218</td><td>1.19</td></tr><tr><td rowspan="2">Twill weave (295 g/m<sup>2</sup>)</td><td>0.5</td><td>0.398 ± 0.004</td><td>0.38 ± 0.045</td><td>0.96</td></tr><tr><td>0.55</td><td>0.2 ± 0.022</td><td>0.183 ± 0.023</td><td>0.91</td></tr></tbody></table>

Observing the flow front advancement in the different textiles, it was noticed that the dual-scale effect, i.e. the difference of fluid velocity between the inter-tow and intra-tow space, is visible in the unidirectional (UD) reinforcement (Fig. 3(a)) and the Biaxial stitched fabrics (Fig. 3(b)), less visible for the plain weave textile (Fig. 3(c)), whereas, the flow front was nearly uniform for the twill weave and quad axial stitched fabrics and no impregnation delay was observed.

<figure><img src="image_525.png" /><figcaption>Fig. 3: Oil flow fronts during injection for different textiles.</figcaption></figure>

## Numerical method for permeability prediction

Numerical simulation is widely used at the component scale to simulate the LCM process but its usage for material characterization is not very well developed despite its increasing potential [17]. For complicated and realistic flow fields, numerical simulation is convenient. By imposing a
<table><caption>Table 2: Summary of the experimental results</caption><thead><tr><th>Textile</th><th>V<sub>f</sub></th><th>K<sub>sat</sub>[10<sup>-10</sup>m<sup>2</sup>]</th><th>K<sub>unsat</sub>[10<sup>-10</sup>m<sup>2</sup>]</th><th>K<sub>unsat</sub>/K<sub>sat</sub></th></tr></thead><tbody><tr><td rowspan="3">E-glass Plain weave (800 g/m<sup>2</sup>)</td><td>0.5</td><td>2.916 ± 0.19</td><td>3.16 ± 0.24</td><td>1.08</td></tr><tr><td>0.55</td><td>1.12 ± 0.11</td><td>1.16 ± 0.14</td><td>1.04</td></tr><tr><td>0.6</td><td>0.482 ± 0.06</td><td>0.458 ± 0.054</td><td>0.95</td></tr><tr><td rowspan="3">0/90 Biaxial stitched fabric (825 g/m<sup>2</sup>)</td><td>0.5</td><td>1.39 ± 0.2</td><td>1.6 ± 0.3</td><td>1.15</td></tr><tr><td>0.55</td><td>0.672 ± 0.097</td><td>0.729 ± 0.11</td><td>1.08</td></tr><tr><td>0.6</td><td>0.326 ± 0.044</td><td>0.34 ± 0.011</td><td>1.04</td></tr><tr><td rowspan="3">Quad Axial Stitched Fabric (851 g/m<sup>2</sup>)</td><td>0.5</td><td>0.62 ± 0.026</td><td>0.649 ± 0.038</td><td>1.045</td></tr><tr><td>0.55</td><td>0.27 ± 0.012</td><td>0.286 ± 0.007</td><td>1.05</td></tr><tr><td>0.6</td><td>0.171 ± 0.016</td><td>0.177 ± 0.0178</td><td>1.03</td></tr><tr><td rowspan="3">Unidirectional reinforcement fabrics (850 g/m<sup>2</sup>)</td><td>0.5</td><td>3.5 ± 0.22</td><td>4.97 ± 0.8</td><td>1.42</td></tr><tr><td>0.55</td><td>1.91 ± 0.057</td><td>2.47 ± 0.097</td><td>1.29</td></tr><tr><td>0.6</td><td>1.12 ± 0.16</td><td>1.33 ± 0.218</td><td>1.19</td></tr><tr><td rowspan="2">Twill weave (295 g/m<sup>2</sup>)</td><td>0.5</td><td>0.398 ± 0.004</td><td>0.38 ± 0.045</td><td>0.96</td></tr><tr><td>0.55</td><td>0.2 ± 0.022</td><td>0.183 ± 0.023</td><td>0.91</td></tr></tbody></table>

Observing the flow front advancement in the different textiles, it was noticed that the dual-scale effect, i.e. the difference of fluid velocity between the inter-tow and intra-tow space, is visible in the unidirectional (UD) reinforcement (Fig. 3(a)) and the Biaxial stitched fabrics (Fig. 3(b)), less visible for the plain weave textile (Fig. 3(c)), whereas, the flow front was nearly uniform for the twill weave and quad axial stitched fabrics and no impregnation delay was observed.

<figure><img src="image_536.png" /><figcaption>Fig. 3: Oil flow fronts during injection for different textiles.</figcaption></figure>

## Numerical method for permeability prediction

Numerical simulation is widely used at the component scale to simulate the LCM process but its usage for material characterization is not very well developed despite its increasing potential [17]. For complicated and realistic flow fields, numerical simulation is convenient. By imposing a
pressure gradient $\Delta P$ across the domain length $L_{flow}$, the steady state volumetric flow rate $Q$ through a cross-section area $A_{flow}$ can be numerically approximated by solving Stokes equation assuming creeping flow, i.e incompressible non-inertial characterized by a low Reynolds number, of a virtual incompressible Newtonian fluid of viscosity $\mu$. The saturated permeability (capillary effects are not considered) is deduced using equation (Eq. 6) which is a simplification of Darcy's law (Eq. 1)

$$K = \frac{Q \cdot \mu \cdot L_{flow}}{A_{flow} \cdot \Delta P} \qquad (6)$$

This numerical method is used to predict the saturated permeability of a fibrous reinforcement based on real images acquired from a 4 layers stack E-glass plain weave textile used in the previous experiment. The physical dimensions of the scanned specimen were 25.14 mm × 26.04 mm with a variable height for each target fibre volume fraction (between 2.18 mm and 3.28 mm). The voxel dimensions were 2012 × 2084 pixels in the weft and warp directions and a variable number of pixels in the through thickness direction (175 for the highest volume fraction 0.6 and 211 for the smallest one 0.5). The procedure of acquisition is described in detail by Wijaya et al. [18].

## Single-scale flow

In this case, the tows are considered as solid impermeable zones and in order to optimize the simulation time and allocated computational resources, only the inter-tow space was meshed. The test fluid used has a viscosity of $\mu = 1$ mPa.s and a pressure gradient $\Delta P = p_{in} - p_{out} = 1000 - 100 = 900$ Pa is imposed between the inlet and the outlet.

The results of the numerical permeability in the weft direction compared to the experimental values for different fibre volume fractions are given in figure 4(a). The numerical method underestimates the experimental values for the different volume fractions which was expected since the flow inside the tows is neglected.

<figure><img src="image_526.png" /><figcaption>Fig. 4: Comparison between the numerical simulation results and the experimental values.</figcaption></figure>

We observe, in figure 4(b) a large scatter between the numerical simulation and the experimental results for the warp direction (60% for $V_f = 0.55$) compared to the weft direction (maximum of 25% for $V_f = 0.5$) which can be explained by the small size of the flow channels in the warp direction, therefore these narrow geometries are more difficult to be meshed and can be eventually removed if they are badly shaped. As a result, the numerical simulation underestimates the permeability in the warp direction. This problem can be solved if the flow inside the tows is taken into consideration.
pressure gradient $\Delta P$ across the domain length $L_{flow}$, the steady state volumetric flow rate $Q$ through a cross-section area $A_{flow}$ can be numerically approximated by solving Stokes equation assuming creeping flow, i.e incompressible non-inertial characterized by a low Reynolds number, of a virtual incompressible Newtonian fluid of viscosity $\mu$. The saturated permeability (capillary effects are not considered) is deduced using equation (Eq. 6) which is a simplification of Darcy's law (Eq. 1)

$$K = \frac{Q \cdot \mu \cdot L_{flow}}{A_{flow} \cdot \Delta P} \qquad (6)$$

This numerical method is used to predict the saturated permeability of a fibrous reinforcement based on real images acquired from a 4 layers stack E-glass plain weave textile used in the previous experiment. The physical dimensions of the scanned specimen were 25.14 mm × 26.04 mm with a variable height for each target fibre volume fraction (between 2.18 mm and 3.28 mm). The voxel dimensions were 2012 × 2084 pixels in the weft and warp directions and a variable number of pixels in the through thickness direction (175 for the highest volume fraction 0.6 and 211 for the smallest one 0.5). The procedure of acquisition is described in detail by Wijaya et al. [18].

## Single-scale flow

In this case, the tows are considered as solid impermeable zones and in order to optimize the simulation time and allocated computational resources, only the inter-tow space was meshed. The test fluid used has a viscosity of $\mu = 1$ mPa.s and a pressure gradient $\Delta P = p_{in} - p_{out} = 1000 - 100 = 900$ Pa is imposed between the inlet and the outlet.

The results of the numerical permeability in the weft direction compared to the experimental values for different fibre volume fractions are given in figure 4(a). The numerical method underestimates the experimental values for the different volume fractions which was expected since the flow inside the tows is neglected.

<figure><img src="image_537.png" /><figcaption>Fig. 4: Comparison between the numerical simulation results and the experimental values.</figcaption></figure>

We observe, in figure 4(b) a large scatter between the numerical simulation and the experimental results for the warp direction (60% for $V_f = 0.55$) compared to the weft direction (maximum of 25% for $V_f = 0.5$) which can be explained by the small size of the flow channels in the warp direction, therefore these narrow geometries are more difficult to be meshed and can be eventually removed if they are badly shaped. As a result, the numerical simulation underestimates the permeability in the warp direction. This problem can be solved if the flow inside the tows is taken into consideration.
## Dual-scale flow

In this case, the tows are considered permeable and the fluid will flow both between and inside the tows. The same boundary conditions of the single-scale test are used. The permeability of the tows can be computed using the Gebart analytical model assuming hexagonal arrangement of fibers and using Gebart model (the fibers have a 12 µm diameter) [2].

<figure><img src="image_527.png" /><figcaption>Fig. 5: Comparison between the numerical simulation results and the experimental values considering dual scale.</figcaption></figure>

As can be noticed in figure Fig. 5, the permeability values given by the numerical simulation considering porous tows are closer to the experimental values compared to single-scale tests which gives more confidence in the numerical method proposed. The deviation from the experimental values is reduced to 11% for $V_f = 0.5$ and 8% for $V_f = 0.6$ in the weft direction and it is around 20% in the warp direction which is a significant reduction compared to the single-scale case.

## Conclusions

We used a 2D radial flow experimental set-up to investigate the discrepancy between the unsaturated and saturated permeability and to observe the flow front advancement during oil injection of different fabrics. The ratio between both permeabilities was between 0.9 and 1.42 which lies in the range found in the literature and was coherent with the flow front observations: the ratios were greater and different from 1 for the UD and biaxial textiles for example.

A numerical method for permeability prediction was used to estimate the permeability of the plain-weave textile which was already measured experimentally. A good agreement was found between the experimental and numerical values especially when the intra-tow flow is considered in the numerical simulation.

## References

[1] V. Michaud, A Review of Non-saturated Resin Flow in Liquid Composite Moulding processes, Transp Porous Med, vol. 115, n 3, p. 581-601, déc. 2016. https://doi.org/10.1007/s11242-016-0629-7

[2] B. R. Gebart, « Permeability of Unidirectional Reinforcements for RTM », Journal of Composite Materials, vol. 26, n° 8, p. 1100-1133, août 1992. https://doi.org/10.1177/002199839202600802

[3] Willsen Wijaya. Permeability of 2D Woven Composite Textile Reinforcements: Textile Geometry and Compaction, and Flow Modelling. PhD thesis, 2020, The University of Auckland.

[4] D. May et al., In-plane permeability characterization of engineering textiles based on radial flow experiments: A benchmark exercise, Compos. Part A Appl. Sci. Manuf. 121 (2019). https://doi.org/10.1016/j.compositesa.2019.03.006
## Dual-scale flow

In this case, the tows are considered permeable and the fluid will flow both between and inside the tows. The same boundary conditions of the single-scale test are used. The permeability of the tows can be computed using the Gebart analytical model assuming hexagonal arrangement of fibers and using Gebart model (the fibers have a 12 µm diameter) [2].

<figure><img src="image_538.png" /><figcaption>Fig. 5: Comparison between the numerical simulation results and the experimental values considering dual scale.</figcaption></figure>

As can be noticed in figure Fig. 5, the permeability values given by the numerical simulation considering porous tows are closer to the experimental values compared to single-scale tests which gives more confidence in the numerical method proposed. The deviation from the experimental values is reduced to 11% for $V_f = 0.5$ and 8% for $V_f = 0.6$ in the weft direction and it is around 20% in the warp direction which is a significant reduction compared to the single-scale case.

## Conclusions

We used a 2D radial flow experimental set-up to investigate the discrepancy between the unsaturated and saturated permeability and to observe the flow front advancement during oil injection of different fabrics. The ratio between both permeabilities was between 0.9 and 1.42 which lies in the range found in the literature and was coherent with the flow front observations: the ratios were greater and different from 1 for the UD and biaxial textiles for example.

A numerical method for permeability prediction was used to estimate the permeability of the plain-weave textile which was already measured experimentally. A good agreement was found between the experimental and numerical values especially when the intra-tow flow is considered in the numerical simulation.

## References

[1] V. Michaud, A Review of Non-saturated Resin Flow in Liquid Composite Moulding processes, Transp Porous Med, vol. 115, n 3, p. 581-601, déc. 2016. https://doi.org/10.1007/s11242-016-0629-7

[2] B. R. Gebart, « Permeability of Unidirectional Reinforcements for RTM », Journal of Composite Materials, vol. 26, n° 8, p. 1100-1133, août 1992. https://doi.org/10.1177/002199839202600802

[3] Willsen Wijaya. Permeability of 2D Woven Composite Textile Reinforcements: Textile Geometry and Compaction, and Flow Modelling. PhD thesis, 2020, The University of Auckland.

[4] D. May et al., In-plane permeability characterization of engineering textiles based on radial flow experiments: A benchmark exercise, Compos. Part A Appl. Sci. Manuf. 121 (2019). https://doi.org/10.1016/j.compositesa.2019.03.006
[5] N. Vernet et al, Experimental determination of the permeability of engineering textiles: Benchmark II, Compos. Part A Appl. Sci. Manuf. 61 (2014) 172-184.
https://doi.org/10.1016/j.compositesa.2014.02.010

[6] R. Arbter et al. Experimental determination of the permeability of textiles: A benchmark exercise, Compos. Part A Appl. Sci. Manuf. 42 (2011) 1157-1168.
https://doi.org/10.1016/j.compositesa.2011.04.021

[7] Darcy, Henry. "Les fontaines publiques de la ville de Dijon, Dalmont." Paris, France (1856).

[8] F. Zhang, S. Comas-Cardona, et C. Binetruy, Statistical modeling of in-plane permeability of non-woven random fibrous reinforcement, Composites Science and Technology, vol. 72, n 12, p. 1368-1379, 2012. https://doi.org/10.1016/j.compscitech.2012.05.008

[9] S. Comas-Cardona, B. Cosson, S. Bickerton, et C. Binetruy, An optically-based inverse method to measure in-plane permeability fields of fibrous reinforcements, Composites Part A: Applied Science and Manufacturing, vol. 57, p. 41-48, 2014.
https://doi.org/10.1016/j.compositesa.2013.10.020

[10] Sousa Pedro, Lomov Stepan V., Ivens Jan, Hurdles and limitations for design of a radial permeameter conforming to the benchmark requirements. Frontiers in Materials, 9, 2022.
https://doi.org/10.3389/fmats.2022.871235

[11] J. R. Weitzenbock, R. A. Shenoi, et P. A. Wilson, Radial flow permeability measurement. Part A: Theory, Composites Part A: Applied Science and Manufacturing, vol. 30, n 6, p. 781-796, 1999. https://doi.org/10.1016/S1359-835X(98)00183-3

[12] R. Pomeroy, S. Grove, J. Summerscales, Y. Wang, et A. Harper, Measurement of permeability of continuous filament mat glass-fibre reinforcements by saturated radial airflow, *Composites Part A: Applied Science and Manufacturing*, vol. 38, n° 5, p. 1439-1443, 2007.
https://doi.org/10.1016/j.compositesa.2006.11.011

[13] K. M. Pillai, Modeling the Unsaturated Flow in Liquid Composite Molding Processes: A Review and Some Thoughts, *Journal of Composite Materials*, vol. 38, n° 23, p. 2097-2118, 2004.
https://doi.org/10.1177/0021998304045585

[14] M. F. Foley et T. Gutowski, The effect of process variables on permeability in the flexible resin transfer molding (FRTM) process, p. 326-340, 1991.

[15] B. R. Gebart et P. Lidström, Measurement of in-plane permeability of anisotropic fiber reinforcements, *Polymer Composites*, vol. 17, n° 1, p. 43-51, 1996.
https://doi.org/10.1002/pc.10589

[16] M. Pollard, Permeabilities of Fiber Mats Used in Resin Transfer Molding, 24th International SAMPE Technical Conference, pp. T408-T420, 1992.

[17] T. Schmidt, D. May, M. Duhovic, A. Widera, M. Humbert, et P. Mitschang, A combined experimental-numerical approach for permeability characterization of engineering textiles, Polymer Composites, vol. 42, n 7, p. 3363-3379, 2021. https://doi.org/10.1002/pc.26064

[18] W. Wijaya, M. A. Ali, R. Umer, K. A. Khan, P. A. Kelly, et S. Bickerton, An automatic methodology to CT-scans of 2D woven textile fabrics to structured finite element and voxel meshes, Composites Part A: Applied Science and Manufacturing, vol. 125, p. 105561, 2019.
https://doi.org/10.1016/j.compositesa.2019.105561
[5] N. Vernet et al, Experimental determination of the permeability of engineering textiles: Benchmark II, Compos. Part A Appl. Sci. Manuf. 61 (2014) 172-184.
https://doi.org/10.1016/j.compositesa.2014.02.010

[6] R. Arbter et al. Experimental determination of the permeability of textiles: A benchmark exercise, Compos. Part A Appl. Sci. Manuf. 42 (2011) 1157-1168.
https://doi.org/10.1016/j.compositesa.2011.04.021

[7] Darcy, Henry. "Les fontaines publiques de la ville de Dijon, Dalmont." Paris, France (1856).

[8] F. Zhang, S. Comas-Cardona, et C. Binetruy, Statistical modeling of in-plane permeability of non-woven random fibrous reinforcement, Composites Science and Technology, vol. 72, n 12, p. 1368-1379, 2012. https://doi.org/10.1016/j.compscitech.2012.05.008

[9] S. Comas-Cardona, B. Cosson, S. Bickerton, et C. Binetruy, An optically-based inverse method to measure in-plane permeability fields of fibrous reinforcements, Composites Part A: Applied Science and Manufacturing, vol. 57, p. 41-48, 2014.
https://doi.org/10.1016/j.compositesa.2013.10.020

[10] Sousa Pedro, Lomov Stepan V., Ivens Jan, Hurdles and limitations for design of a radial permeameter conforming to the benchmark requirements. Frontiers in Materials, 9, 2022.
https://doi.org/10.3389/fmats.2022.871235

[11] J. R. Weitzenbock, R. A. Shenoi, et P. A. Wilson, Radial flow permeability measurement. Part A: Theory, Composites Part A: Applied Science and Manufacturing, vol. 30, n 6, p. 781-796, 1999. https://doi.org/10.1016/S1359-835X(98)00183-3

[12] R. Pomeroy, S. Grove, J. Summerscales, Y. Wang, et A. Harper, Measurement of permeability of continuous filament mat glass-fibre reinforcements by saturated radial airflow, *Composites Part A: Applied Science and Manufacturing*, vol. 38, n° 5, p. 1439-1443, 2007.
https://doi.org/10.1016/j.compositesa.2006.11.011

[13] K. M. Pillai, Modeling the Unsaturated Flow in Liquid Composite Molding Processes: A Review and Some Thoughts, *Journal of Composite Materials*, vol. 38, n° 23, p. 2097-2118, 2004.
https://doi.org/10.1177/0021998304045585

[14] M. F. Foley et T. Gutowski, The effect of process variables on permeability in the flexible resin transfer molding (FRTM) process, p. 326-340, 1991.

[15] B. R. Gebart et P. Lidström, Measurement of in-plane permeability of anisotropic fiber reinforcements, *Polymer Composites*, vol. 17, n° 1, p. 43-51, 1996.
https://doi.org/10.1002/pc.10589

[16] M. Pollard, Permeabilities of Fiber Mats Used in Resin Transfer Molding, 24th International SAMPE Technical Conference, pp. T408-T420, 1992.

[17] T. Schmidt, D. May, M. Duhovic, A. Widera, M. Humbert, et P. Mitschang, A combined experimental-numerical approach for permeability characterization of engineering textiles, Polymer Composites, vol. 42, n 7, p. 3363-3379, 2021. https://doi.org/10.1002/pc.26064

[18] W. Wijaya, M. A. Ali, R. Umer, K. A. Khan, P. A. Kelly, et S. Bickerton, An automatic methodology to CT-scans of 2D woven textile fabrics to structured finite element and voxel meshes, Composites Part A: Applied Science and Manufacturing, vol. 125, p. 105561, 2019.
https://doi.org/10.1016/j.compositesa.2019.105561
# Automatic planning strategy for robotic lay-up of prepregs on a complex-shaped mold

GAMBARDELLA Antonio¹, a*, ESPERTO Vitantonio¹, b, CARLONE Pierpaolo¹, c

¹Department of Industrial Engineering, University of Salerno, 132 Via Giovanni Paolo II, 84084 Fisciano, Italy

aangambardella@unisa.it, bvesperto@unisa.it, cpcarlone@unisa.it

**Keywords:** Robotic Layup, Composites Manufacturing, Prepreg Forming, Process Planning

**Abstract.** The evolution of composite component production has been driven by a constant quest for improvements in process efficiency, precision, and repeatability. The eventual transition from traditional hand layup to robotic layup represents a significant step in this evolution. The implementation of robotic layup systems has become increasingly prevalent in the manufacturing industry, particularly in the aerospace and automotive sectors, where lightweight, strength, and precision are mandatory requirements. Ideally, the goal is the development of processes where a highly precise robotic arm could automate the deposition of composite materials onto the mold, providing a certain reduction of human errors, and minimizing material waste and associated costs. In this context, this paper proposes a computational tool that is able to provide automatic layup planning for the robotic layup process. The implemented algorithm incorporates the knowledge of a professional laminator: it can automatically analyze a generic mold surface of complex shape, work out the correct strategies for lamination, and generate instructions for robot movements.

## Introduction

In the ever-evolving landscape of advanced manufacturing, the aerospace and automotive industries are continually pushing the boundaries of materials and design to enhance performance and efficiency [1]. In this framework, advanced process monitoring, control and prediction is fundamental for the realization of robust manufacturing of high-performances components [2-4]. One main aspect of this evolution is the integration of automated programming for the robotic layup of prepreg onto complex mold shapes [5]. Prepreg, a composite material consisting of reinforcing fibers pre-impregnated with a resin matrix, offers exceptional strength-to-weight ratios, making it a preferred choice for high-performance applications [6,7]. The utilization of robotics in the layup process brings forth unprecedented precision, repeatability, and efficiency, ultimately revolutionizing the way composite structures are manufactured [8].

Manual layup is often labor-intensive, time-consuming, and error-prone, making it a bottleneck in producing these high-performance materials [9]. Robotic layup emerges as a promising solution to overcome these drawbacks and enhance the efficiency and quality of the layup process.

Robotic layup involves the automated placement of prepreg plies onto a mold surface according to a predefined layup sequence. This process requires precise control of the robot's toolpath and the ability to handle the delicate prepreg materials without damage [10]. Automated programming plays a crucial role in enabling robotic layup, as it transforms the layup sequence into a series of instructions that the robot can execute [8].

The automation of the layup process not only addresses the challenges posed by intricate mold shapes but also significantly reduces human error, resulting in consistently high-quality composite structures [11]. Through the integration of advanced algorithms and robotic systems, manufacturers can achieve unparalleled accuracy in the placement of prepreg layers, optimizing

<figure><img src="image_528.png" /></figure>
# Automatic planning strategy for robotic lay-up of prepregs on a complex-shaped mold

GAMBARDELLA Antonio¹, a*, ESPERTO Vitantonio¹, b, CARLONE Pierpaolo¹, c

¹Department of Industrial Engineering, University of Salerno, 132 Via Giovanni Paolo II, 84084 Fisciano, Italy

aangambardella@unisa.it, bvesperto@unisa.it, cpcarlone@unisa.it

**Keywords:** Robotic Layup, Composites Manufacturing, Prepreg Forming, Process Planning

**Abstract.** The evolution of composite component production has been driven by a constant quest for improvements in process efficiency, precision, and repeatability. The eventual transition from traditional hand layup to robotic layup represents a significant step in this evolution. The implementation of robotic layup systems has become increasingly prevalent in the manufacturing industry, particularly in the aerospace and automotive sectors, where lightweight, strength, and precision are mandatory requirements. Ideally, the goal is the development of processes where a highly precise robotic arm could automate the deposition of composite materials onto the mold, providing a certain reduction of human errors, and minimizing material waste and associated costs. In this context, this paper proposes a computational tool that is able to provide automatic layup planning for the robotic layup process. The implemented algorithm incorporates the knowledge of a professional laminator: it can automatically analyze a generic mold surface of complex shape, work out the correct strategies for lamination, and generate instructions for robot movements.

## Introduction

In the ever-evolving landscape of advanced manufacturing, the aerospace and automotive industries are continually pushing the boundaries of materials and design to enhance performance and efficiency [1]. In this framework, advanced process monitoring, control and prediction is fundamental for the realization of robust manufacturing of high-performances components [2-4]. One main aspect of this evolution is the integration of automated programming for the robotic layup of prepreg onto complex mold shapes [5]. Prepreg, a composite material consisting of reinforcing fibers pre-impregnated with a resin matrix, offers exceptional strength-to-weight ratios, making it a preferred choice for high-performance applications [6,7]. The utilization of robotics in the layup process brings forth unprecedented precision, repeatability, and efficiency, ultimately revolutionizing the way composite structures are manufactured [8].

Manual layup is often labor-intensive, time-consuming, and error-prone, making it a bottleneck in producing these high-performance materials [9]. Robotic layup emerges as a promising solution to overcome these drawbacks and enhance the efficiency and quality of the layup process.

Robotic layup involves the automated placement of prepreg plies onto a mold surface according to a predefined layup sequence. This process requires precise control of the robot's toolpath and the ability to handle the delicate prepreg materials without damage [10]. Automated programming plays a crucial role in enabling robotic layup, as it transforms the layup sequence into a series of instructions that the robot can execute [8].

The automation of the layup process not only addresses the challenges posed by intricate mold shapes but also significantly reduces human error, resulting in consistently high-quality composite structures [11]. Through the integration of advanced algorithms and robotic systems, manufacturers can achieve unparalleled accuracy in the placement of prepreg layers, optimizing

<figure><img src="image_539.png" /></figure>
the mechanical properties of the final composite part [12,13]. Automated robotic layup offers
several advantages over manual layup, including:

* increased efficiency: robots can layup prepreg at a much faster rate than humans, reducing production time and cost;

* improved accuracy: robots can precisely place prepreg plies with minimal error, ensuring the manufactured part's consistent quality and dimensional accuracy[14];

* reduced labor costs: By automating the layup process, manufacturers can reduce reliance on skilled labor and associated costs;

* improved worker safety: manual layup can be hazardous due to the handling of sharp prepreg materials and the exposure to fumes and solvents. Robots can perform layup tasks safely without exposing workers to these risks.

The use of automated programming for robotic layup has the potential to revolutionize the manufacturing of carbon fiber composites. By automating this critical step, manufacturers can achieve significant improvements in efficiency, accuracy, and cost-effectiveness, paving the way for the widespread adoption of these high-performance materials in various applications [13-15].

The idea behind this work is the creation of a tool to improve the efficiency of the robotic layup process and reduce human involvement even in the coding phase, simply providing the discretized surface of the mold as input for the algorithm, in charge of the definition of the layup strategy. This step is supported by a forming simulation, where prepregs are modelled as a virtual grid in which cell edges represent the fiber yarns, and the cell nodes represent their crossing points.

## Materials and methods

The tool mentioned above has been implemented in MATLAB environment. The algorithm starts by importing the STL file of the mold surface to be processed. "Gmsh", an open-source CAE software widely used for FEM analysis, has been used for mesh generation.

Several surface geometries have been created to test the algorithm under different conditions. CADs of the molds have been made with CATIA V5 software and then loaded into the MATLAB environment.

The implemented algorithm starts by loading the discretized surface of a mold. In order to understand the geometry loaded by the user, the algorithm first needs to process the mesh and its elements. After that, following a series of iterations, the choice of the best drafting strategy is provided. The best strategy is defined in terms of the order of areas of the surface to be processed during the robotic layup process in order to incur as little deformation of the prepreg as possible. In fig. 1, a description of the implemented algorithm is provided via flowchart describing its main steps.

After loading the discretized surface, the first calculation performed is to identify the normal vectors for each mesh element. The surface is divided into sub-surfaces, and the curvature for each sub-surface is estimated using a Gaussian curvature calculation function. Next, the user is prompted to specify which part of the surface will be involved in the layup process. The user has the option to choose whether to lay up the entire mold surface or only a portion of it. These steps aim to reduce the amount of data as the algorithm progresses. Once the user selects the sub-surfaces to be processed, a new discretized surface is created, consisting only of the parts involved in the layup process.
the mechanical properties of the final composite part [12,13]. Automated robotic layup offers
several advantages over manual layup, including:

* increased efficiency: robots can layup prepreg at a much faster rate than humans, reducing production time and cost;

* improved accuracy: robots can precisely place prepreg plies with minimal error, ensuring the manufactured part's consistent quality and dimensional accuracy[14];

* reduced labor costs: By automating the layup process, manufacturers can reduce reliance on skilled labor and associated costs;

* improved worker safety: manual layup can be hazardous due to the handling of sharp prepreg materials and the exposure to fumes and solvents. Robots can perform layup tasks safely without exposing workers to these risks.

The use of automated programming for robotic layup has the potential to revolutionize the manufacturing of carbon fiber composites. By automating this critical step, manufacturers can achieve significant improvements in efficiency, accuracy, and cost-effectiveness, paving the way for the widespread adoption of these high-performance materials in various applications [13-15].

The idea behind this work is the creation of a tool to improve the efficiency of the robotic layup process and reduce human involvement even in the coding phase, simply providing the discretized surface of the mold as input for the algorithm, in charge of the definition of the layup strategy. This step is supported by a forming simulation, where prepregs are modelled as a virtual grid in which cell edges represent the fiber yarns, and the cell nodes represent their crossing points.

## Materials and methods

The tool mentioned above has been implemented in MATLAB environment. The algorithm starts by importing the STL file of the mold surface to be processed. "Gmsh", an open-source CAE software widely used for FEM analysis, has been used for mesh generation.

Several surface geometries have been created to test the algorithm under different conditions. CADs of the molds have been made with CATIA V5 software and then loaded into the MATLAB environment.

The implemented algorithm starts by loading the discretized surface of a mold. In order to understand the geometry loaded by the user, the algorithm first needs to process the mesh and its elements. After that, following a series of iterations, the choice of the best drafting strategy is provided. The best strategy is defined in terms of the order of areas of the surface to be processed during the robotic layup process in order to incur as little deformation of the prepreg as possible. In fig. 1, a description of the implemented algorithm is provided via flowchart describing its main steps.

After loading the discretized surface, the first calculation performed is to identify the normal vectors for each mesh element. The surface is divided into sub-surfaces, and the curvature for each sub-surface is estimated using a Gaussian curvature calculation function. Next, the user is prompted to specify which part of the surface will be involved in the layup process. The user has the option to choose whether to lay up the entire mold surface or only a portion of it. These steps aim to reduce the amount of data as the algorithm progresses. Once the user selects the sub-surfaces to be processed, a new discretized surface is created, consisting only of the parts involved in the layup process.
<figure><img src="image_529.png" /><figcaption>Figure 1 - Description of the implemented algorithm using a flowchart describing its main steps.</figcaption></figure>

From this point on, the algorithm incorporates concepts from another work described in the article "A simple MATLAB draping code for fiber-reinforced composites with application to optimization of manufacturing process parameters" [16]. The processes outlined in this article are implemented to provide data better suited for the automatic programming of a robotic layup process.

For this new surface, a layup simulation is performed. The prepreg is simulated by creating a grid on the surface mold from a specific starting point (see Fig. 2). The choice of the starting point is crucial as it represents the initial contact between the prepreg sheet and the mold during the actual layup. The simulation yields different results for different starting points, mirroring the real process where the first point of contact influences the entire lay-up outcome. The curvatures encountered during layup deform the prepreg based on the starting point and layup directions.

The objective is to simulate layup from various points, studying the deformation of the virtual prepreg sheet, and defining the optimal layup approach that results in the best layup strategy for the specific mold surface to be machined. Finally, all information generated in the previous steps is collected and provided to an algorithm for generating robotic routes and movements.

<figure><img src="image_530.png" /><figcaption>Figure 2 - The fabric modeled as a grid of pin-joined cells. The origin node is chosen as starting point for weft and warp generation. On the right, an example of an origin node for a sub-surface from which the warp and weft lines start.</figcaption></figure>