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ISBN print: 978-3-0364-0985-6
ISBN eBook: 978-3-0364-1985-5
DOI book: 10.4028/b-bC9QOP
© 2026 Trans Tech Publications Ltd, All Rights Reserved

Non-Conventional Processes

Edited by
Denise Bellisario

Non-Conventional Processes

Special topic volume with invited peer-reviewed papers only

Edited by
Denise Bellisario

Scientific.Net

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Solid State Phenomena, ISSN: 1662-9779, Vol. 391, pp -1-0

© 2026 The Author(s). Published by Trans Tech Publications Ltd, Switzerland.

Accepted:
2026-04-15

Online:
2026-04-20

Preface

The original version of this paper is available on https://www.scientific.net/SSP.391.-1.pdf

The main topics of this issue is electrochemical and electro-physical machining processes, including electrical discharge machining (EDM), precision / electrochemical machining (ECM / PECM), electromagnetic forming (EMF), electrohydraulic forming (EHF), materials processing by plasma, electron beam or laser beam. Our issue ranges from surface engineering to joining, forming, subtractive and additive manufacturing, single point forming for polymers or composites, or creep age forming. To maintain this characteristic of individuating new processes of scientific interest, the papers regarding innovative processes or forming processes of non-conventional materials (innovative materials or hybrid structures) are presented in this issue too.

Editor

The original version of this paper is available on https://www.scientific.net/SSP.391.-1.pdf

Bellisario Denise
University of Rome Tor Vergata, Italy

Co-editors

The original version of this paper is available on https://www.scientific.net/SSP.391.-1.pdf

Table of Contents

Preface

Mechanical Performance of Incrementally Formed Flax Fiber-Reinforced Polypropylene Caps

Antonio Formisano, Luca Boccarusso, Dario de Fazio, Giuseppe Irace, Massimo Durante

Modulated Pulsed Electrochemical Machining of Tungsten Carbide-Cobalt: An Experimental Study According to DIN SPEC 91399

Alexander Thielecke, Pascal Clauß, Gunnar Meichsner, Philipp Damm, Guido Jungblut, Marco Schäfer, Matthias Hackert-Oschätzchen

Mining Critical Raw Materials from Waste Printed Circuit Boards by Thermo-Mechanical Disassembly

Alice Proietti, Denise Bellisario, Leandro Iorio, Dounia Noqra, Giorgio Patrizii, Fabrizio Quadrini, Loredana Santo

Preliminary Comparative Assessment of Hole Quality and Geometrical Tolerances in AM-Built and CNC-Drilled PLA Components

Martina Panico, Guido Di Bella, Massimo Durante, Mohamed Chairi, Antonio Langella, Luca Boccarusso

Classification and Detection Method for Discharges in Electrical Discharge Machining through Unsupervised Learning

Philipp Damm, Oliver Bederitzky, Jakob Vogel, Arthur Riefer, Philipp Plänitz, Christoph Lerez, Gunnar Meichsner, Matthias Hackert-Oschätzchen

Concurrent Effect of Laser Texturing and Resin Pre-Coating on the Performance of CFRP Single Lap Joints

Gianluca Parodo, Luca Sorrentino, Sandro Turchetta, Giuseppe Moffa

Enabling the Reuse of Single-Lap Adhesive-Bonded Joints through Laser Cleaning and Texturing

Vincenzina Siciliani, Matteo Benvenuto, Chiara Mandolfino, Barbara Reggiani, Leonardo Orazi

Unconventional Recycling of Aeronautical Ground Carbon Fiber Composite Wastes

Pietro Russo, Francesco Napolitano, Fabrizio Quadrini, Denise Bellisario

Electrodeposited Metallic Coatings as a Manufacturing Route to Enhance the Durability of NdFeB Magnets

Fabio Scherillo, Alessia Serena Perna, Antonio Viscusi

Study of the Effect of Foam Substrate Material on Hydrophobic Coating Efficiency

Aikaterini Baxevani, Eleni Lamprou, Fani Stergioudi, Nikolaos Michailidis

Origami-Inspired Design for Manufacture of Composite Parts

Sutharsanan Navaratnarajah, Vincent K. Maes, Mark Schenk, James Kratz

Surface Roughness Control in Incoloy 800HT via Laser Engraving Parameters

Evangelos Nikolidakis, Panagiotis Karmiris-Obratański, Emmanouil Lazaros Papazoglou, Lisa Fatyga, Angelos P. Markopoulos

Non-Conventional Method of Evaluation of the Abrasion Resistance of a Polymeric Material Using a Hole Enlargement Process by Abrasion

Roxana Gabriela Hobjâlă, Razvan Mititelu, Vasile Ermolai, Ştefan Jureschi, Ioan Surugiu, Vlada Ţisari, Margareta Coteata, Laurenţiu Slătineanu

Non-Conventional Recycling Process of Hemp/Carbon Hybrid Laminates via Thermo-Mechanical Disassembly

Luca Boccarusso, Denise Bellisario, Alessandro Canneva, Dario de Fazio, Massimo Durante, Fabrizio Quadrini, Loredana Santo

Keywords Index

Authors Index

The original version of this eBook is available on https://www.scientific.net/book/non-conventional-processes/978-3-0364-1985-5

Solid State Phenomena, ISSN: 1662-9779, Vol. 391, pp 1-9

doi: 10.4028/p-HmIX8y

© 2026 The Author(s). Published by Trans Tech Publications Ltd, Switzerland.

Submitted:
2025-11-28

Revised:
2026-02-09

Accepted:
2026-02-17

Online:
2026-04-20

Mechanical Performance of Incrementally Formed Flax Fiber-Reinforced Polypropylene Caps

The original version of this paper is available on https://www.scientific.net/SSP.391.1.pdf

Antonio Formisano 1,a*, Luca Boccarusso 1, b, Dario De Fazio 1,c, Giuseppe Irace 1, d and Massimo Durante 1,e

¹Department of Chemical, Materials and Production Engineering, University of Naples Federico II, P.le V. Tecchio 80, 80125 Napoli, Italy

a aformisa@unina.it, b luca.boccarusso@unina.it, c dario.defazio@unina.it, d giuseppe.irace@unina.it, emdurante@unina.it

Keywords: Incremental Forming, Natural Fiber-Reinforced Plastics, Mechanical Performance.

Abstract

Incremental forming represents a versatile and cost-effective alternative to conventional sheet forming processes. In recent years, its application has been extended to polymers and composite materials. Among these, natural fiber-reinforced thermoplastics offer several advantages, as natural fibers are widely available, contribute to the semi-biodegradability of the composites, and serve as effective reinforcements for polymer matrices. This experimental study examines the mechanical properties of flax woven fabric-reinforced polypropylene composites, fabricated via compression molding, and their suitability for producing spherical caps through cold incremental forming. A range of features was investigated to assess the effectiveness of incremental forming on these biobased composites and to compare the mechanical performance of undeformed and deformed laminates.

Introduction

The original version of this paper is available on https://www.scientific.net/SSP.391.1.pdf

Incremental sheet forming is a manufacturing technique originally developed as an alternative to conventional forming methods, particularly for shaping flat metal sheets into complex threedimensional profiles [1]. The core principle of this technique involves the progressive deformation of a clamped sheet of material by a forming tool, which is controlled by a computerized numerical control (CNC) machine. The tool follows a programmed path, gradually shaping the sheet into its final geometry [2]. In recent years, its scope has expanded to include polymers and composite materials, broadening its applicability to lightweight and sustainable engineering solutions.

Composite materials are used in various fields, thanks to their outstanding properties [3]; the use of natural fibers as an alternative to conventional synthetic reinforcements (particularly glass and carbon fibers) has garnered increasing attention in both research and industrial contexts, owing to their low-cost, biodegradability, renewability, and non-toxicity [4-6].

Hemp and flax are among the most widely available and commonly used natural fibers in Europe. Their complex microstructure [7,8] contributes to their superior mechanical properties, making them the strongest and stiffest natural fibers [9], and thus ideal candidates to produce biocomposites [10].

Polypropylene, alongside polyethylene and polyvinyl chloride, is among the most used polymer matrices for natural fiber composites [11]. As the world's second-most widely produced synthetic polymer, polypropylene is extensively employed across various industrial sectors, including the fabrication of advanced composite materials [12, 13].

A previous study by the authors [14] demonstrated the feasibility of incremental forming (conducted at room temperature and without the use of dedicated dies) for natural fibre-reinforced polymer composite laminates. This enabled the fabrication of components that are typically challenging to produce due to the intrinsic characteristics of the materials involved. Moreover, the low forming force requirements suggest promising implications for energy efficiency, aligning with sustainable manufacturing objectives.

This study presents an experimental campaign focused on the manufacture of spherical caps via incremental forming with a partial counter die, starting from flax fiber-reinforced polypropylene laminates. The forming process was conducted without localized heating [15] to preserve its
operational flexibility and ease of implementation. The composite laminates were produced through compression molding using flax woven fabrics, without the application of fiber treatments or coupling agents, thereby maintaining the environmental advantages associated with natural fibers.

Following the mechanical characterization of the laminates, the study investigates key aspects of the incremental forming process applied to biocomposites, including geometrical accuracy, forming forces, power, and energy consumption, as well as the post-forming mechanical performance of the spherical caps. These components are considered for potential applications such as the shaping of stiffening ribs in structural panels in the automotive, aerospace, and marine sectors [16].

Materials and Methods

The original version of this paper is available on https://www.scientific.net/SSP.391.1.pdf

Laminates consisting of polypropylene reinforced with flax fibers in fabric form, with a thickness of 2.2 mm , were manufactured using neat polypropylene films (supplied by GDC S.r.l.; thickness: 0.5 mm ) and a woven flax fabric (supplied by FIDIA S.r.l. - Technical Global Services; areal density: 320 g/m2 ). The fabric, whose main properties are detailed in [14], was not subjected to any chemical or surface treatment prior to processing. To eliminate residual moisture, it was dried at 60∘C for 12 hours before the molding procedure.

The laminates were produced via compression molding using a conventional hydraulic press. The stacking sequence consisted of five layers: the outermost layers (two on each side) were polypropylene films, while the central layer was a woven flax fabric. The molding process was conducted at a temperature of 200∘C, with a total cycle time of 300 s (comprising an initial dwell time of 120 s ) followed by the application of 4 MPa pressure for 180 s through the press plates.

The resulting reinforced laminates, containing 15.6wt% flax fiber, were characterized by tensile testing. Compared to neat polypropylene [17], the reinforced laminates exhibited enhanced tensile properties, as summarized in Fig. 1.

[picture]

Fig. 1. Tensile properties of unreinforced and reinforced laminates

Fig. 1. Tensile properties of unreinforced and reinforced laminates

Spherical caps with a base radius of 40 mm and a polar angle of 50∘ were fabricated from the flaxreinforced polypropylene laminates using cold negative incremental sheet forming. This kind of geometry, with decreasing wall angle and deformation states, was chosen because it can be used for applications such as shaping stiffening ribs for different applications and does not show formability problems (the ISF process applied to the material under investigation highlighted a maximum wall angle of less than .50∘[14]). To mitigate bending defects near the base of the caps, a partial counter die was employed, consisting of a hollow cylinder with internal and external diameters of 80 and 100 mm , respectively.

The forming process was carried out using a non-rotating stainless-steel stylus with a hemispherical head of 10 mm in diameter, which progressively deformed the laminate. It was constrained by a clamping frame with a square working area of 100×100 mm2. The forming tool was operated by a C.B. Ferrari high-speed four-axis vertical machining center at a nominal feed rate of 1000 mm/min, following an alternating helical toolpath with 1∘ angular step-down (i.e. the angular increment per full revolution). This strategy significantly reduced the likelihood of twisting [18], as previously observed in the incremental forming of metallic [2] and polycarbonate sheets [19]. The process was conducted under lubricated conditions using Boelube 70104 (100A) lubricant (supplied by Orelube) to minimize the risk of failure and defects.

A schematic representation of the forming setup is provided in Fig. 2.

[picture]

Fig. 2. Schematic of the incremental sheet forming process

Fig. 2. Schematic of the incremental sheet forming process

Several parameters were evaluated, including geometrical accuracy, forming forces, power and energy consumption.

Geometrical accuracy was assessed by comparing the shape of the formed caps to the target geometry using a Zeiss DuraMax coordinate measuring machine (measurement accuracy: 2.4μ m ) and Calypso software.

Forming forces were recorded at a sampling rate of 50 Hz using a K-MCS10 multicomponent sensor, integrated with the QuantumX MX840B data acquisition system and Catman Easy AP software. These force measurements were also used to calculate power and energy consumption, following the methodology described in a previous study by the authors on the incremental forming of flax and hemp fiber-reinforced polypropylene composites [20]. This approach enables estimation of the actual energy required for the forming process, which constitutes only a small fraction of the total energy consumed (most of which is attributed to auxiliary functions of the equipment [21]).

To compare the mechanical performance of undeformed and post-formed laminates, flexural and penetration tests were also conducted.

Flexural tests, including both simply supported and fixed-end configurations, were performed using an MTS Alliance RT/50 universal testing machine. Loading and supporting pins with a diameter of 10 mm were employed. For the fixed-end tests, the supporting pins were replaced by the clamping frame used during the incremental forming process, with only the two edges parallel to the loading pins being constrained.

Penetration tests were conducted using a spherical cap with a base radius of 20 mm and a polar angle of 70∘, employed as the loading tool to test specimens constrained along their periphery. The
remaining equipment was identical to that used for the incremental forming tests. Failure zones resulting from these tests were examined using a Hirox RX-100 digital microscope.

All tests were performed on specimens with a square working area of 100×100 mm2. The testing machines imposed a vertical displacement on the loading tools at a nominal speed of 60 mm/min.

Results and Discussion

The original version of this paper is available on https://www.scientific.net/SSP.391.1.pdf

Two tests for each different case were performed; Given the limited variability observed among repetitions, only representative curves and average values of the investigated features are reported for the sake of conciseness. The incremental forming process was successfully conducted without the occurrence of typical defects such as twisting and wrinkling. An example of a formed spherical cap is shown in Fig. 3a, while Fig. 3b presents a comparison between the actual and target profiles. Two key geometrical features are highlighted: the deviation in maximum height and the gap observed at the intersection between the base of the cap and the flange. The components exhibited good surface quality and satisfactory geometrical accuracy, especially considering the cold nature of the process and the use of a partial counter die rather than a dedicated forming die.

[picture]

Fig. 3. Spherical cap (a) and comparison between actual and target profiles (b)

Fig. 3. Spherical cap (a) and comparison between actual and target profiles (b)

The incremental forming process was characterized by low forming forces, with a maximum value of 536 N . Consequently, both the peak power and the energy consumption were limited, amounting to 2 W and 563 J , respectively. These values indicate non-severe operating conditions, which enable the use of non-dedicated tools and machinery for the incremental forming of biobased composite materials.

Simply supported flexural tests were performed on undeformed laminates and deformed laminates with specimens oriented both concave side up and concave side down (see schematic in Fig. 4a). As shown in Fig. 4b, the concave side up configuration resulted in superior mechanical performance, with a maximum load of approximately 120 N , compared to about 70 N for both the concave side down and undeformed configurations.

[picture]

Fig. 4. Simply supported flexural tests: Schematic (a) and load-stroke curves (b)

Fig. 4. Simply supported flexural tests: Schematic (a) and load-stroke curves (b)

Based on the previous findings, fixed-end flexural tests on the deformed laminates were conducted exclusively on spherical caps oriented with the concave side up (see schematic in Fig. 5a). As shown in Fig. 5b, this configuration proved effective, with the maximum flexural force increasing from approximately 3200 N for undeformed laminates to about 4000 N for the deformed ones.

Fig. 6 presents the results of the penetration tests, which were performed on undeformed and deformed specimens, with the caps clamped in the concave side up configuration (see schematic in Fig. 6a). Fig. 6b highlights the enhanced strength of the spherical caps: the penetration force (corresponding to the peak of the load-stroke curve) resulted more than doubled compared to the undeformed laminate, rising from approximately 2000 N to 4000 N . In terms of penetration energies, the values increased from 17 J for the undeformed specimens to 115 J for the deformed ones.

[picture]

Fig. 5. Fixed-end flexural tests: Schematic (a) and load-stroke curves (b) undeformed

Fig. 5. Fixed-end flexural tests: Schematic (a) and load-stroke curves (b) undeformed

deformed
(concave side up)

[picture]

Fig. 6. Penetration tests: Schematic (a) and load-stroke curves (b)

Fig. 6. Penetration tests: Schematic (a) and load-stroke curves (b)

Fig. 7 presents a 50× magnification of the failure surface obtained from a penetration test on a deformed laminate. The image reveals poor fiber/matrix adhesion, which is characteristic of this composite system, particularly in the absence of fiber surface treatments and coupling agents [22]. Consequently, stress transfer between the matrix and the fibers is primarily governed by mechanical interlocking, facilitated by the relatively large mesh size of the flax fabric.

[picture]

Fig. 7. Failure surface from a penetration test on a deformed laminate

Fig. 7. Failure surface from a penetration test on a deformed laminate

Conclusions and Future Work

The original version of this paper is available on https://www.scientific.net/SSP.391.1.pdf

This study presents an experimental investigation on cold incremental forming with a partial counter die, applied to flax woven fabric-reinforced polypropylene laminates produced via compression molding, without the use of fiber treatments or coupling agents. The mechanical characterization of the resulting spherical caps was also performed.

The results demonstrate that incremental sheet forming is effective for biobased composites, offering satisfactory geometrical accuracy while requiring relatively low forming forces and power levels.

Furthermore, the incrementally formed spherical caps exhibited improved mechanical performance compared to the undeformed laminates.

Future research could focus on the remolding potential of incrementally formed laminates and the evaluation of their residual mechanical properties. Additionally, the incremental forming of fully natural composite laminates may provide further insights into sustainable manufacturing strategies.

Acknowledgements

The original version of this paper is available on https://www.scientific.net/SSP.391.1.pdf

This work was supported by the "Programma per il Finanziamento della Ricerca di Ateneo (FRA) 2022" of the University of Naples Federico II.

References

The original version of this paper is available on https://www.scientific.net/SSP.391.1.pdf

M.K. Agrawal, P. Singh, P. Mishra, R.K. Deb, K.A. Mohammed, S. Kumar, G. Kumar, A brief review on the perspective of a newer incremental sheet forming technique and its usefulness, Adv. Mater. Process. Technol. 10 (2024) 506-516. https://doi.org/10.1080/2374068X.2023.2168288

J. Jeswiet, F. Micari, G. Hirt, A. Bramley, J. Duflou, J. Allwood, Asymmetric single point incremental forming of sheet metal, CIRP Ann. 54 (2005) 88-114.

L. Nele, A. Caggiano, I. Improta, 4 - Machining of composite materials, in: K. Joseph, K. Oksman, G. George, R. Wilson, S. Appukuttan (Eds.), Woodhead Publishing Series in Composites Science and Engineering, Fiber Reinforced Composites, Woodhead Publishing, 2021, pp. 83-111.

I. Taha, A. El-Sabbagh, G. Ziegmann, Modelling of strength and stiffness behaviour of natural fibre reinforced polypropylene composites, Polym. Polym. Compos. 16 (2008) 295-302.

M. Alhijazi, B. Safaei, Q. Zeeshan, S. Arman, M. Asmael, Prediction of elastic properties of thermoplastic composites with natural fibers, J. Text. Inst. 114 (2023) 1488-1496.

P. Luthra, K.K. Vimal, V. Goel, R. Singh, G.S. Kapur, Biodegradation studies of polypropylene/natural fiber composites, SN Appl. Sci. 2 (2020) 1-13.

L. Pil, F. Bensadoun, J. Pariset, I. Verpoest, Why are designers fascinated by flax and hemp fibre composites? Compos. Part A Appl. Sci. Manuf. 83 (2016) 193-205.

T. Väisänen, P. Kilpeläinen, V. Kitunen, R. Lappalainen, L. Tomppo, Effect of steam treatment on the chemical composition of hemp (Cannabis sativa L.) and identification of the extracted carbohydrates and other compounds, Ind. Crops Prod. 131 (2019) 224-233.

J.P. Manaia, A.T. Manaia, L. Rodriges, Industrial hemp fibers: An overview, Fibers 7 (2019) 106.

W. Frącz, G. Janowski, Ł. Bąk, Influence of the alkali treatment of flax and hemp fibers on the properties of PHBV based biocomposites. Polymers (Basel) 13 (2021) 1965.

R.U. Arinze, E. Oramah, E.C. Chukwuma, N.H. Okoye, A.N. Eboatu, P.I. Udeozo, P.U. ChrisOkafor, M.C. Ekwunife, Reinforcement of polypropylene with natural fibers: Mitigation of environmental pollution, Environ. Challenges 11 (2023) 100688.

K.A. Al-Ghamdi, Spring back analysis in incremental forming of polypropylene sheet: An experimental study, J. Mech. Sci. Technol. 32 (2018) 4859-4869.

Z. Wang, G. Zhu, Development of the temperature-dependent constitutive model of glass fiber reinforced polypropylene composites, Mater. Manuf. Process. 38 (2023) 295-305.

A. Formisano, L. Boccarusso, D. De Fazio, G. Irace, M. Durante, Considerations on the incremental forming of natural fibre-reinforced polypropylene composites, Mater. Res. Proc. 54 (2025) 2238-2245.

S. Torres, R. Ortega, P. Acosta, E. Calderón, Hot incremental forming of biocomposites developed from linen fibres and a thermoplastic matrix, Stroj. Vestn./J. Mech. Eng. 67 (2021) 123-132.

K.Hariprasad, K. Ravichandran, V. Jayaseelan, T. Muthuramalingam, Acoustic and mechanical characterisation of polypropylene composites reinforced by natural fibres for automotive applications, J. Mater. Res. Technol. 9 (2020) 14029-14035.

L. Boccarusso, D. De Fazio, M. Durante, Production of PP composites reinforced with flax and hemp woven mesh fabrics via compression molding, Inventions 7 (2022) 5.

A. Formisano, L. Boccarusso, D. De Fazio, M. Durante, Effects of toolpath on defect phenomena in the incremental forming of thin polycarbonate sheets, Int. J. Adv. Manuf. Technol. 133 (2024) 4957-4966.

M. Durante, A. Formisano, F. Lambiase, Incremental forming of polycarbonate sheets, J. Mater. Process. Technol. 253 (2018) 57-63.

A. Formisano, D. De Fazio, G. Irace, M. Durante, Incremental forming of natural fiberreinforced polypropylene composites: Considerations on formability limits and energy consumption, Mater. 18 (2025) 2688.

I. Bagudanch, M.L. Garcia-Romeu, M. Sabater, Incremental forming of polymers: Process parameters selection from the perspective of electric energy consumption and cost, J. Clean. Prod. 112 (2016) 1013-1024.

L. Sisti, G. Totaro, M. Vannini, P. Fabbri, S. Kalia, A. Zatta, A. Celli, Evaluation of the retting process as a pre-treatment of vegetable fibers for the preparation of high-performance polymer biocomposites, Ind. Crops Prod. 81 (2016) 56-65.

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The original version of this eBook is available on https://www.scientific.net/book/non-conventional-processes/978-3-0364-1985-5

Solid State Phenomena, ISSN: 1662-9779, Vol. 391, pp 11-20

doi: 10.4028/p-wUhTP5

© 2026 The Author(s). Published by Trans Tech Publications Ltd, Switzerland.

Submitted:
2025-12-05

Revised:
2026-02-26

Accepted:
2026-02-26

Online:
2026-04-20

Modulated Pulsed Electrochemical Machining of Tungsten Carbide-Cobalt: An Experimental Study According to DIN SPEC 91399

The original version of this paper is available on https://www.scientific.net/SSP.391.11.pdf

Alexander Thielecke 1,a*, Pascal Clauß 1, b, Gunnar Meichsner 1,c, Philipp Damm 1, d, Guido Jungblut 2,e, Marco Schäfer 2,f Matthias Hackert-Oschätzchen 1, g

¹Chair of Manufacturing Technology with Focus Machining, Faculty of Mechanical Engineering, Otto von Guericke University Magdeburg, Universitätsplatz 2, 39106 Magdeburg, Germany

2 Pro-Com A. Jungblut GmbH, Gewerbegebiet Heiligenwies 17, 66663 Merzig, Germany

a  alexander.thielecke@ovgu.de, b  pascal.clauss@ovgu.de, c  gunnar.meichsner@ovgu.de, d  philipp.damm@ovgu.de, eajungblut@pro-com-web.de,  fmschaefer@pro-com-web.de, g  matthias.hackert-oschaetzchen@ovgu.de (*corresponding author)

Keywords: Modulated Pulsed Electrochemical Machining, Cemented Carbides

Abstract

The following study examines the experimental determination of the electrochemical machinability of cemented carbides with a neutral electrolyte solution. In previous research, a removal mechanism of tungsten carbide-cobalt was demonstrated applying pulsed electrochemical machining. This work focuses on the material removal of tungsten carbide-cobalt with the further developed modulated pulsed electrochemical machining process according to DIN SPEC 91399. Compared to the preceding investigation, the modulated pulsed technique demonstrates a similar removal mechanism with higher current efficiencies.

Introduction

The original version of this paper is available on https://www.scientific.net/SSP.391.11.pdf

Cemented carbides. Cemented carbides are composite materials containing a hard material phase (e. g. tungsten carbide - WC) embedded in a metallic binder matrix (e. g. cobalt - Co), which leads to excellent mechanical properties [1]. On the other hand, its mechanical properties make WC-Co to a difficult-to-cut material [2]. Therefore, electrochemical machining (ECM) is chosen as an alternative machining process, since ECM is capable of machining metals regardless of their mechanical properties like hardness and toughness [3]. The current state of the art is that WC-Co is electrochemically machined with modified, highly corrosive electrolytes [4,5,6]. If the pH ranges at which dissolution of tungsten and cobalt is thermodynamically possible are plotted on a pH scale, there is no overlap - which is displayed in Fig. 1 [7].

[picture]

Fig. 1. Areas of active dissolution of tungsten and cobalt in relation to the pH value of the electrolyte, adapted from [7]

Fig. 1. Areas of active dissolution of tungsten and cobalt in relation to the pH value of the electrolyte, adapted from [7]

The anodic oxidation reaction of tungsten in acidic and weakly alkaline solutions can be expressed as in Eq. 1 [8]:

W+3H2O=W03+6H++6e−(1)

Accordingly, the tungsten component forms a solid, non-conductive oxide layer ( WO3 ) when WC-Co is machined with neutral electrolytes such as sodium nitrate solution (NaNO3), which in turn prevents the electrochemical removal process.

Previous experiments. In previous experiments, a removal mechanism of WC-Co with NaNO3 as the electrolyte was demonstrated by applying pulsed electrochemical machining (PECM). In this work, anodic dissolution of cobalt, an oxidation of tungsten followed by a mechanical separation of the tungsten oxide layer, and the separation of tungsten carbide after the dissolution of the cobalt were shown. With suitable parameter sets, material removal was generated at slow feed rates (0.02 mm/min). Nevertheless, permanent material removal could not be maintained because after a certain distance between the electrodes, the working gap became too small due to the formation of the passivating oxide layer, which led to the experiments being terminated [9].

Modulated pulsed electrochemical machining. The modulated pulsed electrochemical machining (MPECM) process is based on the PECM process which is characterised by an oscillating cathode and a pulsed direct current. PECM itself is an advanced machining process based on the ECM sinking process, where the tool is lowered towards the workpiece and transfers its negative shape into the workpiece. PECM is characterized by an oscillating cathode and a pulsed direct current to reduce the working gap to obtain higher accuracy in consequence of a higher localized current density [10]. In MPECM, the cathode is controlled by a programmable linear axis rather than by a mechanical oscillation. This technical feature allows switching several current pulses at the bottom dead center of the cathode. Fig. 2 illustrates this principle.

[picture]

Fig. 2. Principle of MPECM with multiple modulated direct current pulses [11]

Fig. 2. Principle of MPECM with multiple modulated direct current pulses [11]

The schematic illustration of the MPECM process shown in Fig. 2 is divided into three phases. Phase 0 represents the initial state of the process. In this phase, the anode (workpiece) and cathode (tool electrode) are positioned at the predefined working gap, and no electrochemical machining takes place. In phase I, the machining process (anodic dissolution of the workpiece material) takes place applying a defined number of current pulses. During this phase, the cathode is located at the bottom dead center while maintaining the specified working gap. Phase II describes the flushing phase. No machining current is applied in this phase. The cathode moves to top dead center and the working gap is cleared of reaction products, gas bubbles, and dissolved material. Phases I and II are repeated cyclically until the desired depth of removal is achieved. MPECM therefore enables processing with several short pulses between the flushing phases. Shorter current pulses during machining reduce the working gap and thus also the ohmic resistance. This leads to a reduction in the dispersion of the electric current and an increase in the local current density, which enhances the precision of the machining process. Previous experiments with PECM showed that longer pulses resulted in a less stable removal process with WC-Co. It appears that the longer pulses allow the oxide layer more time to grow, leading to a more homogeneous and thicker oxide layer that interrupts the material removal
process [9]. Accordingly, this study experimentally investigates the material removal of WC-Co applying MPECM. The material-specific removal characteristics, surface appearances and microstructure were evaluated.

Material

The original version of this paper is available on https://www.scientific.net/SSP.391.11.pdf

The removal experiments were conducted using a commercially available CKi ®12 from Gerhard Ihle Hartmetalle Werkzeuge e.K., a tungsten carbide-cobalt type. Images of the base material were taken using a scanning electron microscope (SEM) Hirox SH 5500. Fig. 3 shows the surface of the material.

[picture]

Fig. 3. Material contrast image of a ground, etched sample

Fig. 3. Material contrast image of a ground, etched sample

The microstructure shows that a large part of the material consists of the hard and wear-resistant WC phase, while the cobalt binder is evenly distributed. The average grain size of the tungsten carbide grains corresponds to the manufacturer's specifications of 0.5μ m. The distribution of the elements was analyzed by applying an energy dispersive X-ray analysis (EDX). The chemical composition is shown in Table 1.

Table 1. Chemical composition of CKi®12

Element | WC | W | C | Co
Manufacturer specification | pi [%] | 88 | - | - | 12
EDX | pi [%] | - | 76.98 | 12.40 | 10.63

The batch of test specimens used corresponds to that in [9].

Design of Experiments

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DIN SPEC 91399. To ensure the comparability of the results from the PECM and the MPECM removal experiments, the experiments in this work were carried out according to the DIN SPEC 91399 [9, 12]. The DIN SPEC 91399 specifies a methodology with which process input parameters for PECM processes can be systematically determined. For this analysis, the aim was to obtain the material specific removal characteristic. For this purpose, the removal rate va and the current efficiency η as a function of the current density J, as well as the removed mass ma as a function of the flowed electrical charges Q, were experimentally determined. The analyzed parameters are described in Table 2.

Table 2. Analyzed parameters to describe material-specific removal characteristics

Parameter | Description
Removal rate vA(J) [mm/min] | Anodic material removal rate as a function of the average current density flowing on the removal surface
Current efficiency η [%] | Proportion of the applied electrical current that is effectively used for the material removal, excluding losses due to side reactions
Removed mass ma(Q) [g] | The mass removed from the anode in relation to the current flowing over a specific period of time.

Experimental setup. The removal experiments were performed applying a frontal removal, where the electrode surfaces had the same diameter ( ACathode =AAnode =12 mm ). A lateral flushing concept was chosen. The experiments were carried out on a commercially available PEM 800 S ECM machine tool from PEMTec SNC. Fig. 4 shows the machine tool and the removal device.

[picture]

Fig. 4. Experimental setup; left: PEM 800 S , right: Removal device

Fig. 4. Experimental setup; left: PEM 800 S , right: Removal device

Process parameters. The applied process parameters are largely derived from the previous PECM experiments. In [9], various process parameters such as electrical voltage Uq, feed rate vf, pulse width tp and pulse/oscillation frequency fp/z were varied in a total of 70 removal experiments. Based on this, a parameter set was created that enabled material removal from WC-Co. In MPECM, the process input parameters pulse break tb and number of pulses np have to be added. The selection of the MPECM-specific parameters was determined based on preliminary experiments. Table 3 summarize the final parameter set for the performed removal experiments in this study.

Table 3. Experimental parameters

Symbol | Description | Value
Constant parameters
Z | Oscillation amplitude | 100 μm
fz | Oscillation frequency | 50 Hz
SF-A | Initial front working distance | 120 μm
Uq | Electrical voltage | 10 V
np | Number of pulses | 100
tp | Pulse width | 0.5 ms
tb | Pulse break | 0.5 ms
fp | Pulse frequency | 50 Hz
Variable parameters
vf | Feed rate | 0.01 mm/min – 0.06 mm/min
Electrolyte parameters
w | Mass fraction NaNO₃ | 8 %
- | ph value electrolyte | 8 ± 0.5
pEl | Electrical conductivity | 70 mS/cm ± 2 mS/cm
σ | Electrolyte inlet pressure | 350 kPa
T | Electrolyte temperature | 20 °C ± 1 °C

Procedure of the experiments. As can be seen from Table 3, the feed rate vf of the cathode is the variable parameter of this series of experiments. Experiments with higher feed rates >0.06 mm/min were conducted but led to a premature manual stop of the experiments. Manual termination serves to protect the machine tool and the removal device from damage and is necessary since the formation of the passivating oxide layer prevents reliable short-circuit detection. Furthermore, the machine's internal parameter analysis showed a decreasing electrolyte flow rate as the experiment progressed indicating a closing working gap. It can be assumed that with this parameter set, an equilibrium between the feed rate vf and the removal rate va will not be established above a certain feed rate vf.

The current efficiency η is calculated as the quotient of the effectively removed mass meff  and the specific removal mass msp  [13].

η=meffmsp(2)

The specific removal mass msp is defined by the material constants molar mass M, the electrochemical valence zA and by Faraday's constant F.

msp=MzA·F(3)

If the material to be machined consists of several alloying elements, the specific removal mass of the alloy msp-alloy  is calculated according to Eq. 4, taking into account the mass fraction pi and the respective density ρi :

msp-alloy =1F·∑i=1nMi·ρizi(4)

The effective removal mass meff  of a removal experiment corresponds to the quotient of the removal mass ma and the amount of flowed charge Q.

meff=maQ(5)

The removal mass ma can be determined by weighing the workpiece before and after the removal experiment. For direct current processing, the amount of charge flowing Q during material removal is calculated by integrating the process current between the start time t0 and end time t1 of a multiple pulse.

Q=∫t0t1I(t)dt(6)

As can be seen from Table 3, the feed rate vf is the variable parameter for the removal experiments. The removal rate νf was varied from 0.01 mm/min to 0.06 mm/min in 0.01 mm/min steps to describe the material-specific removal parameters. First, the initial weight of all specimens was weighed using a precision balance Radwag XA 210.4Y PLUS. The removal experiments were carried out at the respective feed rate until a constant current flow was established. After completion of the removal experiments, the surfaces of the specimens were photographed and the final weight weighed. Subsequently, the samples were separated to a defined length for microstructure analysis using wire electrical discharge machining (WEDM), cleaned in an ultrasonic bath, and the surface was analyzed accordingly using SEM.

Results

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Surface appearance. Fig. 5 presents the machined test specimens after the experiments.

[picture]

Fig. 5. Initial state surface and resulting surfaces (top view of the test specimen)

Fig. 5. Initial state surface and resulting surfaces (top view of the test specimen)

The top row shows the state of the machined surfaces immediately after the experiments. The formation of an oxide layer is visible on all surfaces. A firmer, homogeneously distributed oxide layer, compared to the other test specimens, can be seen on the surface machined with the lowest feed rate. The resulting oxide layer becomes less dense with increasing feed rate, but is nevertheless distributed evenly over the entire surface. The bottom row shows the surface of the test specimens after machined by WEDM. It can be seen that the oxide layer has largely flaked off. The surfaces
shown were created solely by resting of the specimens in demineralized water for approximately 10minutes during the WEDM process.

Removal characteristics. Fig. 6 shows the feed rate vf of the cathode and the removal rate va of the workpiece as a function of the effective current density J for the MPECM parameter set shown in Table 3.

[picture]

Fig. 6. Resulting feed rate vf and removal rate va as a function of the current density J

Fig. 6. Resulting feed rate v f and removal rate v a as a function of the current density J

Each dot represents a removal experiment. It can be seen that the feed rate J and removal rate va increase linearly with increasing current density

va(J)=Vm·J+v0(7)

. Due to the pulse processing, the functions differ in their slope. The existing correlation between the removal rate Vm in z -direction and the current density can be described by a linear correlation function according to Eq. 7 [13].

CKi®12

The parameter NaNO3 describes the removal volume of WC-Co J in a 5 A/cm2 solution with direct current processing. The range of current density 16 A/cm2 from Vm=0.0011( mm/min)/(A/cm2) to Jmin=2.57 A/cm2 results in removal volume v0. The interpolated intersection of the x -axis of the linear function represents the theoretical starting point of material removal and corresponds to a minimum current density of msp . The parameter msp[g/C] describes the intersection point with the ordinate axis. For this removal function, this intersection lies in the negative range. A negative vo therefore indicates a transpassive material removal, as a minimum current density can be determined. Table 4 summarizes the parameters.

Table 4. Parameters of correlation function for the removal rate vf

Parameter | J | Vm | v₀ | Correlation coefficient R²
Unit | A/cm² | (mm/min)/(A/cm²) | mm/min | -
Value | 5 - 16 | 0.0011 | -0.0028 | 0.993

Removal mass and current efficiency. The specific removal mass ma-theoretical [g][9] was determined by Faraday's law of electrolysis and describes the mass of the removed material that is theoretically removed by the complete conversion of the electrical charges, provided that the entire electrical charge is fully utilized for the removal. Accordingly, the specific removal msp=281μ g/C mass can be used to derive a theoretical removal mass ma-theoretical . The specific removal mass for the material WC-Co with a proportion of chemical elements according to Table 1 and calculated according to Eq. 4 is Q.

Fig. 7 shows a comparison of the calculated theoretical removal mass ma-MPECM  with the removal mass per electrical charge flow ma-PECM  in the MPECM ±0.044mg and PECM Q [9] experiments.

[picture]

Fig. 7. Calculated and experimental removal mass ma-theretical  for the PECM and MPECM process as a function of flowed electrical charge Q

Fig. 7. Calculated and experimental removal mass m a-theretical for the PECM and MPECM process as a function of flowed electrical charge Q

The red and blue dots represent the respective values of the difference weighing. The maximum measurement deviation of the differential weighing's was 4,ma-theoretical . In both experimental variants, an increase in the removal weight is visible with increasing electrically flowed charge η. It can also be seen that slightly more mass per flowed charge was removed in the MPECM removal experiments compared to the PECM removal experiments. While the MPECM experiments showed a stable electrolyte flow rate, the flow rate decreased in the PECM experiments, which, as already mentioned in the sub-chapter "Procedure of the experiments", indicates a closing working gap. Based on the experiments, it can be assumed that processing with MPECM enables a more stable removal process of the WC-Co material. Furthermore, it is noticeable that the slope of the MPECM function is almost identical to the calculated mass removal function 100%. Since, according to Eq. 97%(PECM=86%[9]) corresponds to a current efficiency η of J, the MPECM experiment series with the selected parameter set has an average current efficiency of η. In the next step, the current efficiency J was examined via the current density 100%. Fig. 8 shows the course of the current efficiency meff  as a function of the current density msp  for the MPECM experiments.

[picture]

Fig. 8. Current efficiency vf as a function of current density J for different feed rates η

Fig. 8. Current efficiency v f as a function of current density J for different feed rates η

Based on Eq. 2, the current efficiency can be determined from the differential weighing measurements. Current efficiencies exceeding WO3 are possible if the effectively NaNO3 removed mass is greater than that calculated mathematical formula based on Faraday's law. This can be explained by the erosion of tungsten carbides and oxides, which are not electrochemically dissolved in the ECM process but are detected by the weight loss measurement of the specimen [14]. As the feed rate mathematical formula increases, the
current density mathematical formula increases and the current efficiency mathematical formula decreases. It appears that at higher current densities, more electrical charge is consumed in side reactions such as the oxidation of the material.

Microstructure. Fig. 9 shows surfaces of tests specimens after the MPECM and PECM experiments.

[picture]

Fig. 9. Topographic SEM images of a cleaned surface after machining, left: MPECM, right: PECM [9]

Fig. 9. Topographic SEM images of a cleaned surface after machining, left: MPECM, right: PECM [9]

After ultrasonic bath cleaning, the surface exhibits a homogeneous structure (marked area 1) with oxide particles (marked area 2). Compared to the PECM experiments, the surfaces of the test specimens after the MPECM experiments were covered only with small, scattered oxide residues. In contrast, after the PECM experiments, the surfaces were covered with larger areas of oxide. Table 5 shows the distribution of elements in the marked areas.

Table 5. Chemical composition of test sample compared with the base material

Element | W | C | Co | O
Base material CKi®12 | pi [%] | 76.98 | 12.40 | 10.63 | -
Area 1 MPECM | pi [%] | 82.46 | 13.30 | 0.44 | 3.80
Area 1 PECM [9] | pi [%] | 82.56 | 11.92 | 0.72 | 4.80
Area 2 MPECM | pi [%] | 52.52 | 28.28 | 1.44 | 18.76
Area 2 PECM [9] | pi [%] | 66.02 | 11.24 | 4 | 18.74

EDX analysis of the machined surfaces shows significant removal of the binder metal cobalt for both process variants. The formation of an oxide layer is corroborated by the EDX analysis, which shows a significantly increased oxygen signal in the corresponding areas. These findings are consistent with the reaction described in Eq. 1 and suggest that the formed oxide present in the image can be identified as tungsten trioxide ( mathematical formula ). The causes of the observed carbon enrichment and its distribution are subjects of further investigations.

Conclusion

The original version of this paper is available on https://www.scientific.net/SSP.391.11.pdf

In this study, the electrochemical removal characteristics of WC-Co was analyzed applying the MPECM process with a commercially available ECM machine tool. A neutral sodium nitrate solution was applied as the electrolyte. Based on the results, it can be concluded that the general removal mechanism in the multiple pulse process does not differ from that in the PECM process. An anodic dissolution of the cobalt content and an oxidation of the tungsten content is followed by a mechanical separation of the tungsten oxide layer by electrolyte flow conditions, gas bubbles or cavitation effects. This results in a removal process that, with the selected parameter set, only runs stably at low cathode feed rates. If a certain feed rate is exceeded, the working gap becomes too small, and removal and byproducts, including the tungsten oxide layer, can no longer be reliably flushed out. The consequence is a decreasing flow rate and, corresponding to the increasing gap resistance, a decreasing process flow and an unstable process. Nevertheless, it has been shown that the short pulsepause sequence with a resting cathode of the MPECM process stabilizes the aforementioned removal
mechanism in comparison to the sinusoidal processing of the PECM process. For future experiments, the integration of ultrasound into the PECM and MPECM processes is being intended in order to suppress the formation of the passivating oxide layer.

Acknowledgements

The original version of this paper is available on https://www.scientific.net/SSP.391.11.pdf

This project is funded by the Federal Ministry for Economic Affairs and Climate Action, following a decision of the German Bundestag. The machine tool PEM 800 S is funded by the German Research Foundation (DFG) with project number 467011871.

References

The original version of this paper is available on https://www.scientific.net/SSP.391.11.pdf

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A. Thielecke, P. Clauß, R. Petermann, G. Meichsner, P. Damm, L. Berg, G. Jungblut, M. Schäfer, M. Hackert-Oschätzchen, Experimental investigation of electrochemical precision machining process for tungsten carbide-cobalt applying sodium nitrate electrolyte, Procedia CIRP 137 (2025) 368-373.

A.Schubert, G.Meichsner, M.Hackert-Oschätzchen, A.Martin, J.Edelmann, Precision and Micro ECM with Localized Anodic Dissolution, ICIT&MPT 8 (2011). ISBN 978-961-6692-02-1

R. Petermann, P. Clauß, G. Meichsner, J. Vogel, M. Hackert-Oschätzchen, Method for the Characterisation of the Material Removal in Modulated Pulsed Electrochemical Machining, Procedia CIRP 137 (2025) 134-139.

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G. Meichsner, M. Hackert-Oschätzchen, M. Krönert, J. Edelmann, A. Schubert, M. Putz, Fast Determination of the Material Removal Characteristics in Pulsed Electrochemical Machining, Procedia CIRP 46 (2016) 123-126.

T. Haisch, E. Mittemeijer, J.W. Schultze, Electrochemical machining of the steel 100Cr6 in aqueous NaCl and solutions: microstructure of surface films formed by carbides, Electrochim. Acta 47 (2001) 235-241.

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The original version of this eBook is available on https://www.scientific.net/book/non-conventional-processes/978-3-0364-1985-5

Solid State Phenomena, ISSN: 1662-9779, Vol. 391, pp 21-27

doi: 10.4028/p-jdQ1RW

© 2026 The Author(s). Published by Trans Tech Publications Ltd, Switzerland.

Submitted:
2025-12-17

Revised:
2026-02-12

Accepted:
2026-02-20

Online:
2026-04-20

Mining Critical Raw Materials from Waste Printed Circuit Boards by Thermo-Mechanical Disassembly

The original version of this paper is available on https://www.scientific.net/SSP.391.21.pdf

Alice PROIETTI 1,a*, Denise BELLISARIO 1,b, Leandro IORIO 2,b, Dounia NOQRA 1, d, Giorgio PATRIZII 1,e, Fabrizio QUADRINI 1,f and Loredana SANTO 1, g

Keywords: thermo-mechanical disassembly, critical raw materials, urban mining.

Abstract

EU dependence on critical raw materials poses a serious risk of fragility and vulnerability, mainly in strategic sectors such as energy and defense. Electronic boards contain high concentrations of metals, particularly copper and lead, comprising critical raw materials. Urban mining offers the possibility to recover these metals, but continuous innovation is necessary to make extraction processes more efficient, productive and cheap. Contamination is an issue which reduces the extraction performance and increases pollution and costs. Thermo-mechanical disassembly can be used to separate electronic components from boards, by removing the soldering alloy. A thermomechanical disassembly process, currently under patenting (namely "impact desoldering"), allows the separation of electronic components and soldering alloys from the boards by impact. It has been applied to 4 kg of waste boards to separate the printed circuit board from the electronic components, and the soldering alloy has been melted to make an ingot. A further thermo-mechanical process has been also applied to extract copper sheets from the disassembled printed circuit boards, by rolling.

1 University of Rome Tor Vergata, via del Politecnico 1, Rome, Italy2 Mercatorum University, p.zza Mattei 10, Rome, Italya  alice.proietti@uniroma2.it, b  denise.bellisario@uniroma2.it, cleandro.iorio@unimercatorum.it, d  dounia.noqra@uniroma2.it, e  giorgio.patrizii@uniroma2.it, f  fabrizio.quadrini@uniroma2.it, gloredana.santo@uniroma2.it

Introduction

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The ever-increasing development of advanced technologies, renewable energies and digitalization requires a huge availability of high-valued raw materials, which can be difficult to collect because of several geopolitical issues. Such materials are commonly called Critical Raw Materials (CRMs) and, since 2011, they are grouped into a list which is updated every 3 years because of economic and social factors. At first 14 CRMs were identified and, since then, their number increased with an average rate of 1.67 until now, with the current list including 34 materials [1]. According to the European Commission, a classification of raw materials can be done based on their supply risk (SR) and economic importance (EI); in this frame, CRMs are those materials that have a SR ≥1 and an EI ≥2.8 [2]. The Critical Raw Materials Act (CRMA), promoted by the European Union, aims to regulate CRMs management and to support industries during their acquisition. In this regard, new strategies are developed in such a way to mitigate the supply risk, fostering the recovery of CRMs from urban waste [3]. This process is called urban mining. Indeed, CRMs are typically used in electronic systems, whose end-of-life represent a serious environmental challenge [4]. In fact, in 2022 the global electronic waste (e-waste) reached 62 million metric tons (MMTs), representing a huge source of materials since it included 31 MMTs of metals, 17 MMTs of plastics and 14 MMTs of other materials and only the 22.3% of e-waste is currently recycled [5]. Wasted electronic boards are the 7% of the total amount of e-waste; they are complex systems [6] that integrate high-value precious metals such as gold ( Au ), palladium ( Pd ) and silver ( Ag ) as well as strategic metals like copper ( Cu ) and tin ( Sn ) [7]. Electronic boards consist of a non-conducting fiberglass laminate that integrates conducting Cu traces, which is named printed circuit board (PCB), and the electric components that are secured on the substrate. Among these devices there are chips, connectors and capacitors, that contain precious and high-value metals [8]. Figure 1 shows a typical PCB with its components. To recycle such
systems, several steps are required. First, mechanical processing like dismantling, upgrading and refining are necessary [9]. The former permits the separation of materials and components into different categories; it is performed using hammers, tongs, screwdrivers and conveyors. In alternative, thermal treatments, as infra-red heating, and chemical reagents, as a solution of fluoroboric acid and hydrogen peroxide (HBF4−H2O2) or nitric acid (HNO3), can be used. After dismantling, physical processes are carried out, that involve the use of crushers and grinders to obtain small parts, and a subsequent step of separation is carried out to divide metallic and non-metallic parts. Magnetic, eddy current and density separation techniques can be used for this purpose. Chemical processes are also used to separate organic and metal parts with the 4 most relevant techniques being pyrolysis, gasification, depolymerization using supercritical fluids [10] and hydrogenolytic. Subsequently, refining is performed; in this step metals are melted by heating (pyrometallurgical processes) or dissolved by a liquid (hydrometallurgical processes) as well as by a combination of them [11,12]. Hydrometallurgical treatments involve acid, cyanide or caustic leaching of solid material. In the end, high-value metals are isolated and concentrated. An innovative approach is based on biohydrometallurgical leaching processes, which use microorganisms to extract metals by generating weaker organic acids. These processes can be used to produce ferric iron generated by bacterial oxidation of Cu and zinc ( Zn ) from PCB. This solution permits to overcome the limits of pyrometallurgical and hydrometallurgical processes by increasing the amount of PCB that can be processed [13]. Finally, the purification of the dissolved metals is carried by liquid/liquid extraction (solvent extraction), precipitation/cementation and electrolyte refinements (electrowinning/electrorecovery). Hazardous materials such as cadmium ( Cd ), lead ( Pb ) and chromium ( Cr ) must be opportunely managed whereas for rare earth metals, such as cerium (Ce), yttrium (Y) and neodymium (Nd), bio-metallurgy and solvent extraction methods are commonly used [14,15].

[picture]

Fig. 1. An example of a board with both the PCB and the electronic components.

Fig. 1. An example of a board with both the PCB and the electronic components.

Several strategies can be adopted to recover the most high-value materials in electronic boards. Chemical and bio-leach liquor obtained by conventional approaches can be followed by precipitation to recover iron ( Fe ) and Zn , and cementation to recover Cu from the dismissed PCB [16]. Valuable constituents such as Au,Cu and glass fibers (GFs) can be recovered by zero-waste loop strategies which combine swelling in a closed reactor to obtain a mixture that has to be purified by leaching and physical processes (sieving or magnetic separation) [17]. For precious metals such as Au the commonly adopted practices are based on obtaining the metallic phase from salts through inorganic materials that works as adsorbers or nano structural materials (graphene oxide, organosilica frameworks); they have a high adsorbent efficiency thanks to their intrinsic characteristics, such as high surface area and large functional groups density compared to bulk materials [18]. There are also innovative approaches which use an eutectic solvent as a novel and green carrier within an emulsion liquid membrane system [19]. Moreover, gas assisted microflow can also be used during purification
to recover Au and Ag [20]. Due to its use in photovoltaic applications recovery of Ga is also important even though it is present in a very low content into electronic boards [21]. New trends are based on automatic systems; in particular, a facility for disassembly and off-gas purification has been used to separate the electronic components from the substrate [22] avoiding hazardous volatile substances.

In managing PCBs desoldering or melting, the removal of the soldered joints is the first essential step to proceed with further recovery of metals from electric and electronic components. It can be carried out by physical desoldering which mainly includes heating a medium (as in thermal convection), infrared and laser heating, or chemical methods [23]. Moreover, the separation of the multi-layer structures of the PCB, after the removal of electrical and electronic components is crucial. Good results have been reached in terms of reduced time and enhanced process efficiency through microwave bath using solvents such as dimethyl formamide or dimethyl acetamide if compared to commonly adopted ultrasonic bath and thermostatic bath [24].

Following the recent trends in the fields of urban mining, this work uses an innovative approach to recover electric and electronic components from PCBs through desoldering by preheating in oven and further impact (Figure 2). Moreover, a thermo-mechanical disassembly strategy, already adopted in previous studies for carbon fiber reinforced composites recycling [25-27], was used for the multilayer structure of the PCBs frame, the copper traces and the soldering alloy recovery. Finally, the recycled Sn was successfully processed to manufacture a bulk ingot.

[picture]

Fig. 2. Desoldering and thermo-mechanical disassembly to recover the constituents of the electronic board.

Fig. 2. Desoldering and thermo-mechanical disassembly to recover the constituents of the electronic board.

Materials and Methods

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Printed circuit boards. Different kinds of electronic boards were used in the experimental phase with a total amount of 4 kg . They have been collected from electronic devices such as personal computers. All of them were made of 3 main categories of materials: electric and electronic components, soldering alloys and copper traces, and fiberglass laminates for the substrate. In Figure 1 an example of the processed bords is shown.

Desoldering and thermo-mechanical disassembly of PCBs. A two-step process has been carried out to recover the main constituents of the electronic boards (Figure 2). In the first step, the desoldering one, they were heated in a muffle furnace Nabertherm L45/12/B140 (Lilienthal, Germany) for 15 min at 250∘C. During heating, the soldering alloy melted and components were removed through impacts. This procedure allowed us to recover not only electronic components but also the soldering alloy. A maximum of 3 heating cycles in oven were applied to each PCB. In the end, an ingot was obtained by pouring the recovered soldering alloy. Subsequently, the PCBs were disassembled by heating in oven with the same process parameter already adopted in the previous step and then, deformed by a roll bending machine, the rollers of which had a nominal diameter of

30 mm . The applied deformation permitted the separation of the fiberglass layers, exposing the copper traces. Each PCB underwent several heating and deformation cycles.

Results and Discussion

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The experimentation has been carried out on a total of 20 boards, different in size (from the minimum of 10×6.5 mm2 to the maximum of 160×100 mm2 ). An oven has been used for PCB heating, at the temperature of 250∘C to reduce the processing time for each PCB, and to compensate for the fact that the oven door was continuously opened. To maximize the mass of disassembled components, up to 3 consecutive impacts have been carried out on each single PCB. The impact for desoldering has been produced on a cold metallic plate to have the immediate solidification of the soldering alloy after its removal from the PCB. In Figure 3, the result of the desoldering process is shown in the case of the best component separation. When desoldering is optimal, the alloy is fully removed from all the holes of the PCB, where the electronic components were initially placed. Machining cannot produce this result, leaving consistent contamination on the PCB, as well as on the components.

[picture]

Fig. 3. Separation of electronic components and soldering alloy by impact.

Fig. 3. Separation of electronic components and soldering alloy by impact.

Not all the boards have been optimally desoldered, also after 3 impacts. In the end, 40% of the boards are almost perfectly desoldered, and another 35% are sufficiently treated, but one in four boards, comprising the largest, is only partially desoldered. In badly desoldered PCBs, there are large areas where the components have been completely removed as well, but such portions (from 20 to 50% ) are almost unaffected. In the case of large boards, this issue depends on the PCB deformability after heating, and the difficulty of transferring the impact energy on the components. In the end of the desoldering campaign, the processed PCBs have a weight of 2.3 kg,58.7% of the initial mass. The disassembled components are 1460 g(37.1%), and the soldering alloy is 106 g(2.7%). A mass of about 60 g(1.5%) has been lost during the process for several reasons, such as the ejections of very small components and alloy droplets, or the degradation of such plastics. This mass loss could be reduced by using an automated impact desoldering machine, that is currently under patenting; it would permit an accurate calibration of process parameters, such as force, distance and speed. However, the recovered soldering alloy is particularly clean as shown in Figure 3.

The soldering alloy, separated by impact from the board, has been cast into a sand mold to extract a small ingot ( 72.4 g ). After pouring, a residual of 21.8 g resulted because of slag, impurities, and embedded wires (Figure 4). To conclude the extraction phase by thermo-mechanical processes, copper has been separated from clean waste PCBs by another innovative technique, namely "disassembly", originally applied to the recycling of carbon fiber reinforced laminates. This is a thermo-mechanical process, as well as the impact desoldering, and can separate single fiberglass plies of the PCBs, and copper layers from them. A hand-operated roll bending machine was used for this purpose. In the future, an automated machine will be prototyped in such a way to optimize process parameters like rolls' speed. Good results have been obtained for such PCBs, as large ones were difficult to separate because of difficulties in having homogeneous heating in the laboratory procedure. When possible, copper layers were fully separated from the inside of the waste PCBs as well as copper tracks from the top and the bottom (Figure 5).

[picture]

Fig. 4. Manufacturing of an ingot with the recovered soldering alloy.

Fig. 4. Manufacturing of an ingot with the recovered soldering alloy.

[picture]

Fig. 5. Copper extraction from waste PCBs by thermo-mechanical disassembly.

Fig. 5. Copper extraction from waste PCBs by thermo-mechanical disassembly.

Summary

The original version of this paper is available on https://www.scientific.net/SSP.391.21.pdf

Experimental results show that impact desoldering is able to separate waste PCBs from electronic components, even at laboratory scale and by hand operation. Contamination from the soldering alloy is minimal both on the PCBs and the components, thus enhancing successive recycling steps to recover CRMs or other high value substances. Secondary raw materials have been extracted in the form of an ingot for the soldering alloy, and copper sheets. Further studies will be performed in the future in such a way to increase the level of components separation. Moreover, in-line control will be implemented to increase the amount of recovered CRMs.

Acknowledgements

The original version of this paper is available on https://www.scientific.net/SSP.391.21.pdf

The authors are grateful to Ing. Fabrizio Betti for the support given in the experimentation. The activities have been carried out in the frame of the MICRAM project (Sustainable Circularity and CRM Recovery for MW Radiation Protection and Stealth Defence Applications), co-funded by the LIFE Programme of the European Union and the European Defence Agency, under the Incubation Forum for Circular Economy in European Defence (IF CEED).

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The original version of this paper is available on https://www.scientific.net/SSP.391.21.pdf

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Solid State Phenomena, ISSN: 1662-9779, Vol. 391, pp 29-42

doi: 10.4028/p-aFH2MU

© 2026 The Author(s). Published by Trans Tech Publications Ltd, Switzerland.

Submitted:
2025-12-18

Revised:
2026-02-21

Accepted:
2026-02-23

Online:
2026-04-20

Preliminary Comparative Assessment of Hole Quality and Geometrical Tolerances in AM-Built and CNC-Drilled PLA Components

The original version of this paper is available on https://www.scientific.net/SSP.391.29.pdf

Martina Panico 1,a*, Guido Di Bella 2, b, Massimo Durante 1,c, Mohamed Chairi 2,3, d, Antonio Langella 1,e and Luca Boccarusso 1,f

1 Dept. of Chemical, Materials and Production Engineering, University of Naples "Federico II", P.le Tecchio 80, 80125, Naples, Italy

2 Department of Engineering, University of Messina, Contrada di Dio, 98166 Messina, Italy

3 CNR ITAE, Salita S. Lucia sopra Contesse 5, 98126 Messina, Italy

a*martina.panico@unina.it, b guido.dibella@unime.it, c mdurante@unina.it, d mohamedchairi@cnr.it, eantgella@unina.it, fluca.boccarusso@unina.it

Keywords: Fused Deposition Modelling (FDM), Hole Quality, CNC Drilling, AM-Built Holes.

Abstract

The increasing adoption of polymer additive manufacturing (AM) in functional applications has intensified the need for reliable hole-making strategies, particularly when printed components require mechanical fastening or post-processing. However, the layered architecture of fused-filament fabrication (FFF) parts typically leads to dimensional inaccuracy and thermally induced defects during drilling. This study examines the quality of holes made directly with FFF technology and those obtained through traditional drilling on polylactic acid (PLA) components produced with two different layer heights ( 0.20 and 0.28 mm ). Additively manufactured holes were analysed via least-squares circle fitting and roundness evaluation, revealing high repeatability but substantial undersizing ( >1 mm deviation from the 6.35 mm nominal diameter). CNC drilling tests were then performed under a full-factorial combination of spindle speed and feed rate, assessing thrust force, torque, infrared-based cutting temperature, and burr formation for both top and bottom sample side. Results show that CNC drilling can restore near-nominal dimensional accuracy, but only within a restricted process parameters window. Low feed rates promote severe thermal-viscoplastic deformation, while specific combinations of spindle speed and entry surface strongly influence burr formation.

Introduction

The original version of this paper is available on https://www.scientific.net/SSP.391.29.pdf

Additive manufacturing (AM) is progressively extending its application domain from rapid prototyping to the production of structural and functional components for industrial use [1,2]. The increasing maturity of polymer-based additive processes, especially those relying on Fused Filament Fabrication (FFF) [3], has enabled the fabrication of lightweight and customized geometries with reduced material waste and shortened lead time. Despite these advantages, AM still faces significant limitations when strict geometric tolerances and performance requirements must be met, particularly in scenarios where the manufactured components are subjected to assembly operations or loadbearing conditions [4-6]. In such cases, even small deviations in local features, such as dimensional accuracy or surface integrity of holes, can critically affect mechanical performance, functional fitting and long-term reliability of the joined structures.

Holes represent a key design feature, especially in components destined for assembly via mechanical fastening. In conventional manufacturing chains, hole production is typically performed through drilling, allowing accurate control of positioning, dimensional tolerance and surface quality. Conversely, in AM, holes can be directly generated during the layer-by-layer deposition process, potentially eliminating secondary machining operations. While this approach is attractive in terms of production time and sustainability, literature highlights that printed holes often suffer from dimensional inaccuracy, poor surface finish, a higher tolerance variability due to staircase effects, process-induced defects, and accumulation of anisotropic printing deviation [7-9]. Nevertheless, AM-built holes may avoid the localized thermal and mechanical stresses associated with CNC
drilling, representing a potential advantage in terms of reduced burr formation and mitigation of heataffected or plastically deformed regions [10,11].

Previous studies have predominantly focused on the dimensional deviations induced by AM at component scale, whereas only limited work has systematically addressed the comparison between holes fabricated directly via AM and those produced through drilling on AM-built substrates. In particular, recent investigations have examined drilling performance on 3D-printed PLA in terms of thrust force, torque and surface integrity, but without benchmarking these outcomes against the characteristics of AM-built holes [11,12]. Furthermore, little attention has been paid to the influence of drilling parameters on post-process geometrical tolerances in additively manufactured polymers, nor to the assessment of positional repeatability, an aspect that is essential in applications where multi-part alignment or functional integration is required. This gap hinders the definition of reliable design criteria for integrating AM with conventional machining, especially for assembly-critical component or hybrid process chains.

In this context, the present work proposes a comparative experimental investigation on the quality of holes fabricated during the AM process or via CNC drilling after deposition phase. PLA plates were produced via FFF using two distinct values of layer thickness ( 0.20 mm and 0.28 mm ), and two configurations were examined: (i) specimens containing a row of AM-built holes, and (ii) full-solid specimens subsequently drilled under different combinations of rotational speed and feed rate. CNC drilling tests were also performed by varying the material entry surface, distinguishing between the top layer ( T -side) and the surface directly in contact with the printing bed (B-side), in order to assess the influence of print-induced surface morphology and process-induced microstructural anisotropy on drilling response. Hole quality was assessed in terms of dimensional accuracy, geometrical and positional tolerances, as well as surface features and burr height, while drilling process monitoring included thrust force, torque and temperature evolution. Based on the experimental outcomes, the optimal drilling condition was identified and employed to replicate a sequence of drilled holes for direct comparison with AM-built ones. Therefore, the objective of this study is to quantitatively evaluate the limitations and opportunities associated with both fabrication strategies, providing useful insights for integrating AM and drilling technologies in functional polymer components and supporting their potential deployment in assembly-critical applications.

Materials and Methods

The original version of this paper is available on https://www.scientific.net/SSP.391.29.pdf

PLA specimens with nominal dimensions of 200 mm×30 mm×3.2 mm were manufactured via Fused Filament Fabrication (FFF) using two layer thickness configurations, namely 0.20 mm (PLA_020) and 0.28 mm (PLA_028). A commercial black PLA filament (PLA Basic, Filament2Print, Spain; 1.75 mm diameter, density 1.24 g/cm3,1 kg spool) was employed for all prints. Printing was carried out at nozzle and bed temperatures within the ranges recommended by the manufacturer (190230∘C and 50−70∘C, respectively). For each layer condition, two sets of samples were produced: the first consisting of specimens with holes directly integrated during the AM process with a nominal diameter of 6.35 mm and a centre-to-centre spacing of 12.7 mm . Each AM-drilled specimen incorporated a sequence of 15 holes, and three specimens were fabricated per layer configuration, resulting in a total of 45 integrated holes for PLA_020 and 45 for PLA_028. The second set comprised solid specimens intended for subsequent CNC drilling. The printing process was performed using Bambu Lab A1 3D printer, maintaining identical build orientation and tool path strategy for all tests; all process parameters (e.g., nozzle temperature, print speed, bed temperature, infill strategy, cooling conditions) were held constant and are summarized in Table 1.

Table 1 Parameters of 3D printing process.

Parameter | Value/Description
Printing technology | Fused Filament Fabrication (FFF)
Printer model | Bambu Lab A1
Nozzle diameter | 0.4 mm
Nozzle temperature | 220 °C
Bed temperature | 65 °C
Layer thickness | 0.20 mm / 0.28 mm
Build orientation | Flat, XY-plane
Walls / Shells | 2 loops
Infill density | 100%
Infill pattern | Rectilinear
Infill direction | 45°
Infill-to-wall overlap | 15%
Initial layer speed (walls) | 50 mm/s
Initial layer speed (infill) | 105 mm/s
Print speed – external wall | 200 mm/s
Print speed – internal wall | 300 mm/s
Print speed – infill | 270 mm/s
Cooling | Active cooling after first layers

CNC drilling tests were performed in dry conditions on a five-axis machining centre (C.B Ferrari, Italy), according to a full-factorial parameter matrix combining different levels of spindle rotation and feed rate. The selected levels for the spindle speed were 800rpm,1600rpm, and 2400 rpm , while the corresponding values for the feed rate were 0.01 mm/rev,0.05 mm/rev, and 0.1 mm/rev. Based on the evidence reported in the literature [11,12], these cutting parameters were selected with the aim of exploring a sufficiently wide process window. In this case, the feed range ( 0.01−0.10 mm/rev ) vary from a condition characterised by high contact time and greater compressive interaction between the tool and the material, to more aggressive values, where chip formation and evacuation are facilitated. At the same time, the spindle speed levels (800−2400rpm, corresponding to a cutting speed of approximately 16−48 m/min ) cover a transition from conservative low-heat conditions to high-energy regimes, where the effects of thermal softening become increasingly significant. All CNC tests were performed using a conventional HSS twist drill bit with a diameter of 6.35 mm and a point angle of 140∘ (supplied by HTT High Tech Tools, Italy). For each condition, three holes were drilled. As an additional experimental variable, the entry surface for drilling was varied to capture the influence of deposition-induced surface characteristics. Specifically, drilling was performed from both the top surface (last deposited layer side, T -side) and the bottom surface (in contact with the printing bed, B-side), which exhibit different thermal histories and surface conditions. Surface roughness measurements, conducted by using a Confocal microscope (Sensofar S neox, 20x magnification) on an area of 25 mm2, highlighted that PLA_020 samples has an Sa of 8.415μ m (top) and 21.770μ m (bottom), while PLA_028 samples has values of 7.860μ m (top) and 20.114μ m (bottom), confirming that the bed-side layers are significantly rougher and potentially more affected by build-induced irregularities. In addition, the first layer deposited, in order to promote adhesion to the printing plate, has a height of 0.30 mm (regardless of the design height of the layers). Process monitoring included thrust force and torque acquisition using a 6-axis force/torque sensor (model KMCS10, HBK World, Southfield, MI, USA) positioned between the workpiece and machine table, with a sampling frequency of 1200 Hz , whereas thermal data were collected during selected trials for qualitative assessment of heat generation during chip evacuation. The experimental setup is shown in Fig. 1. Additionally, infrared thermal imaging was acquired using an Avio R300SR infrared camera on a subset of drilling tests, specifically for selected spindle speed-feed rate combinations where significant thermal effects were expected. The emissivity was set to 0.95 , consistent with typical values for opaque thermoplastic polymers such as PLA. Thermal images were recorded at a frame
rate of 60 Hz . The monitored wavelength range was centred at approximately 10μ m within the longwave infrared region. The camera was oriented to directly observe the hole entry surface (tool entry side) throughout the drilling process.

[picture]

Fig.1. Experimental drilling setup.

Fig.1. Experimental drilling setup.

The hole quality of both CNC drilled and additively pre-integrated holes was evaluated through multiple metrics. For drilled hole, burr height was measured by non-contact 3D optical profiling using a confocal microscope (Sensofar, S, 20x magnification), scanning the entire hole circumference and extracting the maximum burr height as the peak out-of-plane deviation from the original surface [13,14]. Roundness, for both drilled and AM-built holes, was assessed using a Zeiss DuraMax coordinate measuring machine (CMM) with a 1 mm spherical touch probe, programmed via Zeiss Calypso software to ensure repeatable measurement routines [13,14]. For each hole, the probe was initially centred on the nominal hole axis and subsequently acquired 400 discrete measurement points along the circumference at mid-depth of the hole, following a clockwise path. The acquired coordinates were processed through least-squares circle (LSC) fitting to determine the best-fitting diameter. Maximum and minimum local diameters were also extracted to calculate the roundness error ( Dmax−Dmin ), in accordance with ISO 12181. Fig. 2 shows the procedure carried out.

[picture]

Fig.2 Example of roundness assessment.

Fig.2. Example of roundness assessment.

The experimental workflow can therefore be summarised as follows: specimen fabrication via FFF for both layer thicknesses, CNC drilling according to parameter matrix and entry surface variation, process monitoring (force, torque and thermography where applicable), and post-process assessment
of hole quality including dimensional, geometrical and surface-related indicators for comparison between AM-built and CNC-drilled holes.

Results and Discussion

The original version of this paper is available on https://www.scientific.net/SSP.391.29.pdf

Analysis of CNC-drilled holes. For illustrative purpose, Fig. 3 reports the evolution of thrust force (a) and torque (b) for the representative drilling condition at n=1600rpm and f=0.05 mm/rev. Similar to the trends observed across all tested combinations of process parameters, no substantial differences were detected in the shape of thrust force, torque profiles and their peak values when varying either the entry surface (top versus bottom) or the layer thickness ( 0.20 mm versus 0.28 mm ). A further observation can be made regarding the evolution of the torque signal after the maximum peak, particularly in the phase leading to complete tool exit. For B-side drilling, for both investigated layer heights, the post-peak decay does not follow a monotonic trend toward zero. Instead, once the torque has dropped to approximately half of its maximum value, the signal enters a temporary plateau whose duration depends on the process parameters. This plateau precedes the final collapse to zero. A cautious interpretation of this behaviour may relate to the intrinsic structure of the last deposited layer: when drilling from the B-side, the tool eventually engages what corresponds to the topmost printing layer, characterized by a continuous morphology for mainly aesthetic reasons, which may temporarily provide an additional resistance to penetration, thereby sustaining the torque at an intermediate level before full breakthrough.

[picture]

Fig. 3. Time evolution of thrust force (a) and torque (b) during CNC drilling of PLA specimens under different layer thickness and entry surface configurations.

Fig. 3. Time evolution of thrust force (a) and torque (b) during CNC drilling of PLA specimens under different layer thickness and entry surface configurations.

As shown in Fig. 4a and Fig. 4b, the feed rate plays the dominant role in determining the peak thrust force. Notably, f=0.01 mm/ rev leads to the highest force values for all spindle speeds and both entry conditions. This behaviour is attributed to the extremely reduced thickness of the uncut chip, which limits the efficiency of chip formation. The material remains in contact with the cutting edges for a longer time, causing greater mechanical compression and greater local friction, whit a consequent increase in thrust force. Under such conditions, a clear increasing trend with rotational speed can be observed. Such trend aligns with the thermos-sensitive behaviour of PLA. While increasing spindle speed generally promotes cutting temperature and, consequently, material softening (leading to a reduction of resistance to penetration), two concurrent effects may arise at higher speeds. On one side, the increase in strain rate may enhance the material's apparent stiffness, counterbalancing softening to some extent and resulting in slightly higher force levels. On the other side, when the local thermal input exceeds a critical threshold, the polymer transitions from viscoplastic state to a semi-fluid condition. In this regime, the material trends not to fracture or detach cleanly under the cutting action but rather undergoes local deformation and smearing around the tool edges. This behaviour hinders efficient chip formation and increases the energy required for material
displacement, which may lead to local increases in cutting forces despite the nominal thermal softening. No substantial differences emerged between T -side and B -side drilling entries. The two layer thickness configurations generally exhibit comparable thrust force levels, with only marginal differences except at the lowest feed rate ( 0.01 mm/rev ). In this specific condition, the specimens printed with 0.28 mm layer height showed higher peak values compared to those produced with 0.20 mm . As shown in Fig. 4c and Fig. 4d, torque presents a progressive increment with spindle speed up to 1600 rpm , followed by a stabilisation or slight decrease at 2400 rpm . This behaviour reflects the combined effect of strain rate stiffening and thermally induced softening of PLA. At low feed rate (0.01 mm/rev), torque values remain limited due to reduced chip thickness and deformation-dominant removal mechanisms, while the highest torque levels are recorded at f=0.1 mm/rev, where chip formation becomes more efficient and the cutting plane area increases. Only marginal differences emerge between the two layer heights, except at the lowest feed setting, where the coarser stratification of PLA_028 favours local softening and results in slightly lower torque peaks.

[picture]

Fig. 4. (a, b) Maximum thrust force with increasing feed rate for the investigated spindle speeds and drilling direction, for (a) PLA specimens printed with 0.20 mm layer height and (b) PLA specimens printed with 0.28 mm layer height. (c, d) Corresponding maximum torque trends under the same drilling conditions.

Fig. 4. (a, b) Maximum thrust force with increasing feed rate for the investigated spindle speeds and drilling direction, for (a) PLA specimens printed with 0.20 mm layer height and (b) PLA specimens printed with 0.28 mm layer height. (c, d) Corresponding maximum torque trends under the same drilling conditions.

In order to better understand the influence of thermos-mechanical interaction on drilling response, thermographic monitoring was performed for selected combinations of process parameters, i.e., those expected to generate significant temperature rise based on the thrust force and torque analysis. Representative results are reported in Fig. 5, which shows the surface temperature distribution and corresponding temperature profile extracted along a horizontal section passing through the hole centre for the 2400rpm−0.10 mm/rev combination. Under these conditions, the temperature peak reached approximately 54.5∘C. When analysing the combined effect of spindle speed and feed rate, a clear thermal trend emerges. At 0.01 mm/ rev , surface temperatures increased from 59∘C ( 800 rpm ) to 68.6∘C(1600rpm), reaching 87.2∘C at 2400 rpm , in agreement with the observed rise in thrust force. In this configuration, the extremely thin chip promotes prolonged contact between the tool and the workpiece, enhancing friction and heat generation while limiting heat evacuation via material removal. As torque remains relatively low due to inefficient chip formation, most of the generated heat is dissipated into the hole walls, resulting in a strong increase of temperature. Conversely, at 2400 rpm with increasing feed rate ( 0.01,0.05,0.1 mm/rev ), peak temperatures gradually decreased from 87.2∘C to 70.8∘C and finally to 54.5∘C, despite the corresponding rise in torque. This behavior is consistent with the finding that as the thickness of the uncut chip increases, material removal per revolution becomes more effective, allowing the chip to play a key role in heat evacuation. In addition, the reduced exposure time of the material to the cutting edge limits local heat accumulation, preventing excessive softening. These observations align with the dual response of PLA to thermomechanical loading. At low chip thickness, the combination of low material removal rate (MRR, volume of material removed per time unit) and prolonged local deformation leads to notable heating and increased thrust force, exacerbated by viscoelastic deformation of the softened polymer. At higher feed rates, the process transitions towards a chip-forming regime with improved energy distribution, thereby reducing thermal input at the cutting interface despite the higher torque required for thicker chip deformation.

[picture]

Fig. 5. Thermographic analysis of the PLA drilling process under the 2400rpm−0.10 mm/rev combination.

Fig. 5. Thermographic analysis of the PLA drilling process under the 2400 r p m − 0.10 m m / r e v combination.

Fig. 6 illustrates the evolution of burr height at the hole entry as a function of feed rate and spindle speed for specimens manufactured with 0.20 mm layer height, distinguishing between drilling from the extrusion surface (T-side) and from the build-plate contact surface (B-side). At f=0.01 mm/rev, burr height progressively increases from 239μ m at 800 rpm to 463μ m at 1600 rpm , reflecting the combined action of increased thrust force and temperature measured at the tool entry section. At 2400 rpm , the burr height appears to decrease to 159μ m; however, visual inspection confirms the presence of severe local edge rounding ('chamfer-like' deformation, see Fig. 7a), indicating that the
reduced measured value is the result of material ejection rather than actual burr minimisation (for this reason, the burr height data of 159μ m is not shown in Fig. 6). This condition is associated with the highest heat input (cutting temperatures previously measured up to 87.2∘C ), which causes extensive lateral deformation of the material. At an intermediate feed rate ( f=0.05 mm/rev ), the burr height decreases at 800rpm(208μ m) but remains high at 1600rpm(346μ m). At 2400 rpm , chip accumulation was observed around the edge of the hole ( ≫1 mm ), making quantitative measurement of the burr meaningless (see Fig. 7b). This can be attributed to an unfavourable combination of increased chip thickness, exceeding the glass transition temperature of PLA, and limited chip evacuation, which favours local adhesion rather than fracture-based removal. At f=0.10 mm/rev, the lowest values are recorded at 800 and 2400 rpm ( 90μ m and 109μ m respectively), with a pronounced peak at 1600rpm(383μ m). The high-speed reduction of the spindle is associated with lower cutting temperatures (around 55∘C ) that fall below the viscoplastic regime, restoring a more efficient chip formation mechanism.

When drilling from the surface in contact with the printing bed, both magnitude and morphology of burr formation change significantly. At f=0.01 mm/rev, the burr height reaches 307μ m at 800rpm(+28.39% compared to the burr measured for the B-side case), while at 1600 and 2400 rpm the adhesion of chips to the edge of the hole renders the measurement meaningless. At f=0.05 mm/rev, the burr height is limited to 236μ m at 800 rpm but exceed 1 mm at 1600 rpm and return to chip adhesion at 2400 rpm . This highlights that the choice of entry surface (as well as the surface finish and thickness) becomes critical when the cutting temperature exceeds the polymer transition range, while the chip thickness is sufficient to determine lateral friction. At f=0.10 mm/rev, a general downward trend is observed as the rotation speed increases: from 184μ m(800rpm) to 174μ m ( 1600 rpm ) and down to 106μ m at 2400 rpm , closely matching the minimum values measured for T-side configuration under the same conditions. This is consistent with the decrease observed in the cutting temperature at 2400 rpm , which restores an effective chip cutting regime.

Unlike entry-side behaviour, no consistent trend was observed for burr height at hole exit, with severe burr occurrences spread across different parameter combinations. As shown by the qualitative observations, under the most critical conditions (as well as at minimum feed rate), the material did not generate a conventional burr but formed an accumulation that replicated the geometry of the tool tip (see Fig. 7c, 800rpm−0.01 mm/rev ). This indicates sever localised softening and compressive flow of the PLA, with the material being plastically displaced around the drill tip without effective fracture propagation. Under these conditions, the polymer transitions to a semi-fluid state due to prolonged contact at low feed and high compressive stress, leading to stagnation and moulding of the material rather than chip separation induced by cutting. At intermediate feed ( 0.05 mm/rev ), burr formation at the hole exit remains significant for both drilling entry configurations, particularly under low and medium spindle speeds. A substantial reduction is observed when increasing the spindle speed to 2400 rpm , where burr height decreases to 126.25μ m for T-side, corresponding to a reduction of approximately 12% compared with 1600 rpm , and to 106.10μ m for B-side, i.e. a reduction of approximately 89%. Although thermal measurements taken at the hole entrance indicate that local temperatures exceeded the glass transition threshold (approximately 70.8∘C ), the shorter toolmaterial contact time at 2400 rpm limits plastic flow near the exit surface, thereby reducing the extent of deformation and material accumulation. Even when the feed rate increases to 0.10 mm/rev, in both T -side and B-side drilling configurations, there is a decrease in burr height as the rotation speed increases. However, at maximum rotation speed ( 2400 rpm ), after the burr height decreases from 0.01 mm/rev to 0.05 mm/rev, it increases again to 0.10 mm/rev. This non-monotonic behaviour at high rotational speeds confirms that burr formation is determined by the combined effect of thermal and mechanical loads. Infrared measurements at the hole inlet showed that, for PLA_020 at 2400 rpm , the surface temperature decreases when the feed rate is increased from 0.05 to 0.10 mm/rev (from approximately 72∘C to approximately 55∘C ), moving away from the peak thermal softening regime. However, the thrust force increases simultaneously at f=0.10 mm/rev, leading to a more severe mechanical load during penetration. Under these conditions, the reduced dwell time of the tool in the material limits a further increase in temperature, but the increased penetration force promotes
more intense plastic deformation, which ultimately results in new burr height growth below 2400 rpm for both feed directions.

The burr formation behaviour observed for PLA_028 follows the same qualitative trends already described for PLA_020 (see Fig. 6b). In both entry configurations (T-side and B-side), the lowest feed rate ( 0.01 mm/rev ) results in the most severe burr formation, particularly at 1600 rpm , where burr height reaches 849μ m on the T-side. Conversely, higher feed rates ( 0.10 mm/rev ) generally result in lower burr levels, in line with lower tool-material interaction. Although the general trend corresponds to that described for samples printed with a layer height of 0.20 mm , samples printed with a layer height of 0.28 mm consistently exhibit higher top burr height under comparable process conditions. For example, at 800 rpm and 0.01 mm/rev, PLA_ 028 T-side produces a burr height of 382.8μ m, compared to 239.5μ m for PLA_020 T-side, corresponding to an approximate increase of 60%. Similarly, at 0.05 mm/rev and 1600 rpm , the burr height in PLA_028 T-side reaches 507.9μ m, compared to 346.1μ m for PLA_020 T-side ( +47% ). This can be attributed to the fact that, at the exit of the hole, the layer directly affected by the cutting action is thicker in the case of PLA_028 and behaves as a single deformable block, rather than being defined by several thinner layers as in PLA_020. One softened locally, this larger volume layer tends to undergo plastic deformation and accumulate around the cutting edges, thus promoting more severe burr development. Similarly to PLA_020, no monotonic trend can be identified for the burr generated at the hole exit in PLA_028 as a function of either feed rate or spindle speed. In contrast to the entry burr, no substantial differences emerge in absolute terms between the two entry configurations.

[picture]

Fig.6. Mean burr height measured at the hole entry as a function of feed rate and spindle speed for the two layer height configurations: (a) PLA_020 ( 0.20 mm layer height) and (b) PLA_028 ( 0.28 mm layer height).

Fig.6. Mean burr height measured at the hole entry as a function of feed rate and spindle speed for the two layer height configurations: (a) PLA_020 ( 0.20 mm layer height) and (b) PLA_028 ( 0.28 mm layer height).

[picture]

Fig. 7. Representative images of burr formation under the most critical CNC drilling conditions:
(a) 2400rpm−0.01 mm/rev (T-side); (b) 2400rpm−0.05 mm/rev (T-side) and (c) 800rpm−0.01 mm/rev.

Fig. 7. Representative images of burr formation under the most critical CNC drilling conditions: (a) 2400 r p m − 0.01 m m / r e v (T-side); (b) 2400 r p m − 0.05 m m / r e v (T-side) and (c) 800 r p m − 0.01 m m / r e v .

The analysis of thrust force, torque, and burr formation reveals that the drilling response of FFFprinted PLA does not exhibit a single condition that simultaneously minimise all indicators. Instead, the results highlight the need for a process window-based approach. Across all configurations, the lowest thrust force values were obtained at an intermediate feed rate ( 0.05 mm/rev ) combined with moderate-to-high spindle speeds ( 1600−2400rpm ). Lower feed levels ( 0.01 mm/rev ) produced markedly higher forces due to deformation-dominated removal, whereas higher feed ( 0.10 mm/rev ) increased the cutting load through a larger chip cross-section. Torque exhibited a distinct optimum. The lowest values consistently occurred at f=0.01 mm/rev and n=2400rpm, but this condition cannot be considered optimal from a machining standpoint. As previously discussed, extremely low feed rates suppress chip formation and promote a deformation-dominated removal mechanism. In this regime, the cutting edges predominantly plough the material rather than producing a clean chip. This results in a reduction of torque (because the cutting engagement area is minimal) but at the expense of a severe increase in thrust force, extensive local plasticisation, poor evacuation of softened material, and pronounced burr formation. Burr height introduces an additional constraint that further refines the selection of suitable parameters. Across both layer heights (PLA_020 and PLA_028), the most favourable feed rates for burr minimisation lie in the interval f=0.05−0.10 mm/rev. However, the preferred spindle speed differs depending on the drilling entry configuration. For T-side drilling, the lowest burr levels consistently occur at low spindle speeds ( 800 rpm ), while for B-side drilling burr height is minimised under high-speed conditions ( 2400 rpm ). Considering all performance indicators together, the most robust operating region for drilling FFF-printed PLA is defined by feed rates between 0.05 mm/rev and 0.10 mm/rev, combined with: low spindle speed ( 800 rpm ) for T-side entry and high spindle speed ( 2400 rpm ) for B-side entry.

Analysis of Additively Manufactured holes. The dimensional evaluation of the drilled AM holes was conducted by analysing the least-squares circle diameter (LSC), calculated on 45 holes for each layer configuration (PLA_020 and PLA_028). Statistical analysis showed that samples produced with 0.20 mm layer thickness exhibited an average LSC diameter of 5.311 mm with a standard deviation of 0.02056 mm , while 0.28 mm layer samples reported a mean value of 5.214 mm and a standard deviation of 0.01724 mm . To better evaluate dimensional repeatability, the coefficient of variation ( CoV=σ/ mean ) was computed, resulting in 0.387% for PLA_020 and 0.331% for PLA_028. This indicates slightly more consistent dimensional performance when using the thicker layer. Fig. 8 reports the distribution of LSC diameter values through a histogram-based representation for both layer configurations. The PLA_020 configuration shows a marginally higher average diameter but also greater dispersion, whereas PLA_028 exhibits a more concentrated distribution, in agreement with statistical findings (lower CoV).

[picture]

Fig. 8. Distribution of the LSC diameter values measured on the as-printed holes for both layer thickness configurations: (a) PLA samples manufactured with 0.20 mm layer thickness, (b) PLA samples manufactured with 0.28 mm layer thickness.

Fig. 8. Distribution of the LSC diameter values measured on the as-printed holes for both layer thickness configurations: (a) PLA samples manufactured with 0.20 mm layer thickness, (b) PLA samples manufactured with 0.28 mm layer thickness.

Despite these differences, all measured diameters are well below the nominal design value of 6.35 mm , with average deviations of approximately -1.04 mm (PLA_020) and -1.14 mm (PLA_028). According to existing literature [15-17], this dimensional undersizing is intrinsic to the FFF process and mainly attributed to: (i) thermal shrinkage during material cooling; (ii) partial lack of interlayer consolidation and reduced effective wall thickness and (iii) slicing approximation, where the layer-by-layer discretisation of curved geometries leads to local rounding errors. The slightly higher dimensional accuracy of PLA_020 may be associated with the increased discretisation fidelity (more layers per unit thickness), which improves contour resolution. Conversely, PLA_028 benefits from improved thermal stability due to fewer thermal-cycle transitions per unit height, resulting in lower variability despite higher nominal undersizing.

The geometric deviation of the printed holes was quantified by analysing the roundness error. Fig. 9 shows the comparative box plot of the roundness values obtained for both layer thickness configurations and Table 2 shows the corresponding statistical parameters. Overall, the analysis revealed comparable roundness levels between the two PLA sample categories, with mean values of 0.2216 mm for PLA built with layer height of 0.20 mm and 0.2083 mm for PLA built with layer height of 0.28 mm , thus indicating limited influence of layer thickness on the circularity defect. However, PLA_028 exhibited a slightly smaller standard deviation and lower coefficient of variation ( 15.02% ) compared to 18.11% in PLA_020, suggesting improved consistency and repeatability of the hole geometry when a thicker layer is adopted. This behaviour can be associated with the enhanced intra-layer bonding and reduced number of heat cycles occurring in PLA_028, which promote higher local thermal stabilisation during deposition. Conversely, the thinner layer provides finer geometric discretisation, but increases the frequency of deposition and cooling events, introducing higher variability.

[picture]

Fig. 9. Box plot of the roundness values obtained for both layer thickness.

Fig. 9. Box plot of the roundness values obtained for both layer thickness.

Table 2. Roundness evaluation of AM-built holes obtained with different layer thickness configurations.

| PLA_020 | PLA_028
Mean [mm] | 0.22161 | 0.20827
St. dev | 0.04014 | 0.03128
CoV [%] | 18.11 | 15.02

Beyond the dimensional characterisation of AM-built holes, an analogous evaluation was carried out for the CNC-drilled counterparts. In this case, measurements were only taken under drilling conditions that produced minimal levels of burr and an acceptable quality of the internal surface of the hole. The LSC diameter was therefore evaluated exclusively for holes produced within the operational window previously identified as suitable in terms of burr formation (i.e., f=0.05−0.10 mm/rev, and spindle speeds of 800 rpm for T-side drilling and 2400 rpm for B-side drilling). For T-side drilling, both PLA_020 and PLA_028 specimens exhibited their lowest burr levels at 800 rpm and 0.10 mm/rev. Under these conditions, the CNC-drilled holes showed near-nominal dimensional accuracy, with PLA_020 reaching an average LSC diameter of 6.348 mm (standard deviation 0.007 mm ) and PLA_028 a mean value of 6.345 mm (standard deviation 0.009 mm ). The dimensional deviation from the nominal tool diameter ( 6.35 mm ) was therefore minimal ( -0.002 mm and -0.005 mm , respectively). For B-side drilling, the smallest top burr was obtained at 2400rpm−0.10 mm/rev, whereas the most acceptable bottom burr occurred at 2400rpm−0.05 mm/rev, for both PLA_020 and PLA_028. Accordingly, LSC measurements were carried out at both parameter combinations. At 2400rpm−0.10 mm/rev, PLA_020 exhibited an average LSC diameter of 6.343 mm (st. dev. 0.010 mm ), while PLA_028 stabilised at 6.339 mm (st. dev. 0.011 mm ). At 2400rpm−0.05 mm/rev, values remained similarly close to nominal: 6.346 mm for PLA_020 and 6.341 mm for PLA_028, both with sub- 0.012 mm standard deviations. Across all admissible conditions, the CNC-drilled holes consistently show dimensional errors z<0.01 mm, with no meaningful difference between the PLA_020 and PLA_028 configurations. This stands in sharp contrast with AM-built holes, where LSC diameters were under sized by over 1 mm and exhibited significantly larger dispersion.

Conclusions

The original version of this paper is available on https://www.scientific.net/SSP.391.29.pdf

This study clarifies the fundamental differences between holes produced by fused-filament fabrication and those obtained through subsequent CNC drilling. Additively manufactured holes showed good repeatability but a systematic dimensional under sizing exceeding 1 mm relative to the nominal diameter. Both layer configurations ( 0.20 mm and 0.28 mm ) displayed a quasi-Gaussian
dispersion of LSC diameter, confirming consistent process output but also revealing that FFF, in its native form, lacks the precision required for functional through-holes. Roundness deviations were also significantly higher than typical machining tolerances, and their magnitude was intrinsically linked to the staircase effect, layer morphology, and the thermal-geometric limitations of the slicing process. These results indicate that holes made with AM technology cannot be used as printed for applications requiring strict dimensional compliance, unless compensation or oversizing strategies are introduced during the design phase. In contrast, CNC drilling applied to AM components has been shown to restore near-nominal geometry, provided that the process parameters fall within the "acceptable for burr" operating window. When burr formation remains controlled (i.e., f=0.05−0.10 mm/rev, and spindle speeds of 800 rpm for T-side drilling and 2400 rpm for B-side drilling), the resulting LSC diameters differ from the nominal value of 6.35 mm by less than 10μ m, with minimal variability. However, the work also shows that the interaction between the tool and the component introduces specific sensitivities depending on the combination of process parameters. Therefore, there is no single optimal combination; instead, it is necessary to adhere to a limited parameter window to balance primarily the thrust force (which affects tool wear) and minimisation of burr. Although this investigation was conducted on PLA using a single drilling tool geometry, the observed dimensional behaviour of the FFF holes is primarily determined by intrinsic process-related phenomena, such as layer deposition morphology, scale effect and thermal shrinkage. These mechanisms are common to most thermoplastics processed by FFF and therefore suggest that similar trends in terms of dimensional undersizing and deviation from roundness may occur in other polymers, although to an extent that depends on the material. On the contrary, the effectiveness of CNC drilling and the definition of an acceptable operating window depend on the thermomechanical properties of the specific polymer and the geometry of the drill bit used. Further investigations are therefore necessary, including different polymers and tool geometries, in order to quantitatively extend the window of identified parameters.

References

The original version of this paper is available on https://www.scientific.net/SSP.391.29.pdf

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Zaborniak M, Bremek M, Budzik G, Kluczyński J. Analysis of the Dimensional and Shape Accuracy and Repeatability of Models Produced in the Process of Additive Extrusion of Thermoplastic Polymers Using Fused Filament Fabrication Technology. Applied Sciences 2024;14:6404. .

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Gómez-Gras G, Pérez MA, Fábregas-Moreno J, Reyes-Pozo G. Experimental study on the accuracy and surface quality of printed versus machined holes in PEI Ultem 9085 FDM specimens. Rapid Prototyp J 2021;27:1-12. .

Popescu D, Gheorghe Amza C, Marinescu R, Cristiana Iacob M, Luminița Căruţaşu N. Investigations on Factors Affecting 3D-Printed Holes Dimensional Accuracy and Repeatability. Applied Sciences 2022;13:41. .

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Vishwadarshan, Shetty G, Shetty R, J P S, V B, Hegde A. Comprehensive analysis of drilling responses in additively manufactured PLA using a regression-based statistical learning approach. Mater Res Express 2025;12:055302. .

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This ePub file was created using machine-generated methods and automated software to convert the source text into an ePub format and may contain formatting inconsistencies from the original version. Formatting may vary depending on the device used.

The original version of this eBook is available on https://www.scientific.net/book/non-conventional-processes/978-3-0364-1985-5

Solid State Phenomena, ISSN: 1662-9779, Vol. 391, pp 43-53

doi: 10.4028/p-71ES2b

© 2026 The Author(s). Published by Trans Tech Publications Ltd, Switzerland.

Submitted:
2025-12-19

Revised:
2026-02-16

Accepted:
2026-02-18

Online:
2026-04-20

Classification and Detection Method for Discharges in Electrical Discharge Machining through Unsupervised Learning

The original version of this paper is available on https://www.scientific.net/SSP.391.43.pdf

DAMM Philipp 1,a *, BEDERITZKY Oliver 1, b, VOGEL Jakob 1,c, RIEFER Arthur 1, d, PLÄNITZ Philipp 1,e, LEREZ Christoph 1,f, MEICHSNER Gunnar 1,g, HACKERT-OSCHÄTZCHEN Matthias 1,h

1 Chair of Manufacturing Technology with Focus Machining, Faculty of Mechanical Engineering,

Otto von Guericke University Magdeburg, Universitätsplatz 2, 39106 Magdeburg, Germany

a  philipp.damm@ovgu.de, b  oliver.bederitzky@st.ovgu.de, cjakob.vogel@ovgu.de, d  arthur.riefer@ovgu.de, e  philipp.plaenitz@ovgu.de, f  christoph.lerez@ovgu.de, g  gunnar.meichsner@ovgu.de, h  matthias.hackert-oschaetzchen@ovgu.de

Keywords: Die Sinking EDM, Discharges, Detection Method, Machine Learning

Abstract

This paper aims to extend the evaluation of the process of electric discharge machining by analysing the discharges. Therefore, a method for detecting and classifying discharges was developed. To detect different discharge types, experiments were conducted with varying of technology parameters, such as peak current or duty factor. During the experiments, the voltages and the currents were measured via an oscilloscope. For the classification, an unsupervised machine learning method was applied, to cluster and classify the detected discharges and compare them with the measured material removal rate and the measured tool wear rate.

(*corresponding author)

Introduction

The original version of this paper is available on https://www.scientific.net/SSP.391.43.pdf

Electrical Discharge machining (EDM) is a non-conventional machining process to machine electric conductive materials independently of their mechanical properties. The material removal is achieved through a sequence of discrete electrical discharges. Each discharge turns the electrical energy in thermal energy. The thermal energy generates a plasma channel at temperature range from 8000∘C to 12000∘C that melts the surface of electrode and workpiece [1]. Based on the impact of the discharges, the conditions of the machining process alter as a result of particle debris, pollution and temperature fluctuations [2].

To evaluate the performance of the machining process, the material removal rate (MRR), the tool wear rate (TWR) and the surface roughness are typically considered [3]. However, the EDM is an inherently stochastic process influenced by multiple interacting effects such as electric, magnetic, thermal, mechanic, dynamic or hydraulic effects [4, 5]. According to Newman, electrical and nonelectrical variables must be considered to optimise the stochastic nature of the EDM process [1]. Nonelectrical variables are e. g. rotational movement and flushing of dielectric fluid. Electrical variables are e.g. pulse parameters and time domain to characterise discharge properties. Therefore, the discharges are analysed and compared with the measured MRR and TWR.

A characteristic sequence of a discharge is shown in Fig. 1. First, the electrode tool approaches the workpiece. An open-circuit voltage u^i is applied between the electrode and the workpiece and initiating the pulse duration ti. During the ignition time delay td, the circuit voltage generates an electric field [6]. Once, the plasma channel is formed, the voltage collapses. Based on the configured process parameters, an average discharge voltage u¯e and an average discharge current l¯e and a maximum discharge current i^e is applied during the discharge duration te.

[picture]

Fig. 1. Characteristic voltage and current sequence of a discharge adapted from [7]

Fig. 1. Characteristic voltage and current sequence of a discharge adapted from [7]

Depending on the gap conditions, there are different types of discharges occurring such as normal discharge, arcing discharge, short circuit or open circuit as shown in Fig. 2 [8].

[picture]

Fig. 2. Classification of discharge pulses according to [8]

Fig. 2. Classification of discharge pulses according to [8]

However, a reliable identification of discharge types remains a significant challenge. Voltage and current waveforms exhibit large variability due to debris concentration, gap instabilities, and sensor noise, making deterministic threshold-based classification unreliable. Existing approaches often depend on supervised learning, which require extensive manual labelling and offer limited robustness across changing machining conditions. In contrast, the unsupervised learning method removes the requirement for labelled data and can be applied to examine raw data [9]. In addition, the unsupervised learning method includes pattern recognition without the inclusion of a target attribute. Consequently, all variables in the analysis are utilised as inputs, and this approach is suitable for clustering [10].

To address these limitations, this study presents a discharge detection method and an unsupervised classification approach based on voltage and current measurements. The proposed methodology identifies individual discharge events, extracts characteristic signal features, and groups pulses into distinct categories without requiring labelled data or predefined thresholds. This enables a more robust interpretation of EDM performance and supports enhanced process monitoring, optimisation, and control.

Experimental Setup

The original version of this paper is available on https://www.scientific.net/SSP.391.43.pdf

The experiments were conducted on a commercially available SG12R die-sinking EDM machine from Mitsubishi Electric Europe. The utilised dielectric was an oil-based fluid IonoPlus IME-MH from Oelheld. Copper was chosen as the electrode material with a diameter of 1 mm . Solid carbide WC-6Co was chosen as the workpiece material with dimensions of 10 mm×100 mm and a thickness of 2 mm . In each experiment, a through bore was machined into the workpiece. The experimental setup for voltage and current measurement is shown in Fig. 3. A current probe Tektronix TCP312A and an amplifier Tektronix TCPA300 were utilised to measure the current signals. Since several current-conducting cables lead to the spindle, the current probe was placed around the electrode. For this purpose, a holder for the current probe was designed and mounted on the chuck holder.

[picture]

Fig. 3. Experimental setup for voltage and current measurement

Fig. 3. Experimental setup for voltage and current measurement

The voltage was measured between the chuck holder and the workpiece. A differential probe Testec TT-SI-9110 was applied to reduce the measured voltage between the electrodes. The current signals and voltage signals were then measured on the oscilloscope Tektronix MDO34. For this purpose, a measurement with measurement settings, as seen in Table 1, was recorded every 5 seconds.

Table 1. Measurement settings

Setting | Record Length | Sample Rate | Sample Interval
Value | 100,000 Points | 2.5 MS/s | 25 ms

In Table 2 the parameters of the conducted experiments are listed. Each parameter set was repeated 5 times. The parameters are based on the manufacturer's specification. Therefore, the duty cycle DC varies from 2.5% to 9.8% with an input current IP of 19.6 A and 30.8 A and a negative polarity of the electrode.

Table 2. Design of experiments

Parameter set | Input current IP [A] | Pulse on time tON [μs] | Pulse off time tOFF [μs] | Duty cycle DC [%]
01 | 19.6 | 5.2 | 192 | 2.6
02 | 19.6 | 10.4 | 192 | 5.1
03 | 19.6 | 15.2 | 192 | 7.3
04 | 19.6 | 20.8 | 192 | 9.8
05 | 30.8 | 5.2 | 192 | 2.6
06 | 30.8 | 10.4 | 192 | 5.1
07 | 30.8 | 15.2 | 192 | 7.3
08 | 30.8 | 20.8 | 192 | 9.8

Results

The original version of this paper is available on https://www.scientific.net/SSP.391.43.pdf

The detection of discharge pulses is based on van Brakel's robust peak detection algorithm [11]. To this aim, the algorithm generates a signal if a new data point is a certain multiple of a number of standard deviations away from a moving mean. The algorithm was applied to the voltage and current signals. The algorithm for detecting discharges for voltage signals is shown in Algorithm 1.

The algorithm requires a data frame ℂ with the voltage signals, the lag value L and the threshold value T as an input. A lag L=100 and a threshold T=10 were selected for the detection of the discharges. The lag L is the number of historical data points from which the mean value and standard deviation were calculated. The threshold T is the multiplier at which the algorithm signals. Specifically, if the distance between a new data point is greater than the threshold multiplied by the moving standard deviation of the data, the algorithm provides a signal 𝔹Ui. The signal 𝔹Ui is equal to 1 if the difference between the current data point 𝔻i and the previous mean 𝕄i−1 is greater than the threshold multiplied by the previous standard deviation 𝕊i−1. Otherwise, 𝔹Ui is equal to 0 . The resulting signals 𝔹U, the mean values 𝕄 and the standard deviations 𝕊 were returned as an Output.

Algorithm 1 - Discharge detection for voltage signals
    Function: VoltagePulseDetection
        Input: Data Frame \(\mathbb{C}\), Lag \(L\), Threshold \(T\)
        Output: Binary Signal \(\mathbb{B}_{U}\), Moving Mean MI, Standard Deviation \(\mathbb{S}\)
    1: \(\mathbb{D} \leftarrow\) Initiation an Array of Voltage Signals out of \(\mathbb{C}\)
    2: \(\quad \mathbb{F} \leftarrow\) Initiation an Array of Voltage Signals out of \(\mathbb{C}\)
    3: \(\quad \mathbb{B}_{U}, \mathbb{M}, \mathbb{S} \leftarrow\) Initiation an Array of zeros in length of \(\mathbb{C}\)
    4: \(\quad \mathbb{M}_{L-1}=\operatorname{mean}\left(\mathbb{F}_{0} \ldots \mathbb{F}_{L a g-1}\right)\)
    5: \(\quad \mathbb{S}_{L-1}=\) StandardDeviation \(\left(\mathbb{F}_{0} \ldots \mathbb{F}_{\text {Lag }-1}\right)\)
    For each \(i \in\{L, \ldots\), length of \((\mathbb{F})-1\}\) DO
        If \(\left|\mathbb{D}_{i}-\mathbb{M}_{i-1}\right|>T \cdot \mathbb{S}_{i-1}\) Then
            \(\mathbb{B}_{U_{i}}=1\)
            \(\mathbb{F}_{i}=\mathbb{F}_{i-\text { lag }}\)
            \(\mathbb{M}_{i}=\operatorname{mean}\left(\mathbb{F}_{i-L} \ldots \mathbb{F}_{i}\right)\)
            \(\mathbb{S}_{i}=\) StandardDeviation \(\left(\mathbb{F}_{i-L} \ldots \mathbb{F}_{i}\right)\)
        Else
            \(\mathbb{F}_{i}=\mathbb{D}_{i}\)
            \(\mathbb{B}_{U_{i}}=0\)
            \(\mathbb{M}_{i}=\operatorname{mean}\left(\mathbb{F}_{i-L} \ldots \mathbb{F}_{i}\right)\)
            \(\mathbb{S}_{i}=\) StandardDeviation \(\left(\mathbb{F}_{i-L} \ldots \mathbb{F}_{i}\right)\)
    Output: \(\mathbb{B}_{U}, \mathbb{M}, \mathbb{S}\)

Excerpt of the voltage waveform and current waveform, as seen in Fig.4, are shown with the corresponding moving mean and threshold. The waveform was measured by conducting the experiment with parameter set 02 . With this parameter set, an input current ℂ of 19 A and a duty cycle L of T were applied. As described above, a signal is generated when the threshold, highlighted by the green curve, is exceeded. The voltage waveform shows that the pulse detection is not affected by the noise, as the noise is adjusted by the moving mean, which is highlighted by the red curve. This demonstrates that the algorithm is robust and can be used for real time monitoring.

[picture]

Fig. 4. Excerpt from the measurement of top: voltage waveform and bottom: current waveform with corresponding moving mean and threshold value

Fig. 4. Excerpt from the measurement of top: voltage waveform and bottom: current waveform with corresponding moving mean and threshold value

In order to analyse and classify the discharges, an analysis of an entire pulse is required. For this reason, the voltage signals and current signals were combined and the positions of the pulse start and pulse end were output. The following algorithm PulsCombination, as seen in Algorithm 2, was applied for this purpose.

The algorithm requires the voltage signals 𝔹U and the current signals 𝕊. First, an empty array for the pulse start 𝔻← and pulse end ℂ is generated, as well as an array 𝔽← of zeros with a length of ℂ and a Boolean variable 𝔹U,𝕄,𝕊←. The algorithm checks for each element ℂ in 𝕄L−1=mean(𝔽0…𝔽Lag−1) and for each element 𝕊L−1= in (𝔽0…𝔽Lag −1) in parallel, if element i∈{L,… or element (𝔽)−1} is equal to 1 . If |𝔻i−𝕄i−1|>T·𝕊i−1 or 𝔹Ui=1 is equal to 1 the combined signal 𝔽i=𝔽i− lag  is equal to 1 , otherwise 𝕄i=mean(𝔽i−L…𝔽i) is equal to 0 . Subsequently, 𝕊i= is iterated and the pulse start is appended to the (𝔽i−L…𝔽i) array and the pulse end to the 𝔽i=𝔻i array. The output is an array combining 𝔹Ui=0 and 𝕄i=mean(𝔽i−L…𝔽i).

Algorithm 2 - Combining the voltage and current signals and output the position of the discharges

Function: PulseCombination
    Input: Binary signal for voltage signal \(\mathbb{B}_{U}\), binary signal for current Signal \(\mathbb{B}_{I}\)
    Output: PulsePosition \(\mathbb{P}\)
    \(\mathbb{P}_{\text {ON }} \leftarrow\{ \}\)
    \(\mathbb{P}_{\text {OFF }} \leftarrow\{ \}\)
    \(\mathbb{B} \leftarrow\) Initiation an Array of zeros in length of \(\mathbb{B}_{U}\),
    \(s\) ← False
    For each \(\{m, n\} \in\left\{\mathbb{B}_{U}, \mathbb{B}_{I}\right\}\) DO
        If \(m=1 \vee n=1\) Then
            \(\mathbb{B}_{i}=1\)
        Else
            \(\mathbb{B}_{i}=0\)
    For each \(i \in\{0,1, \ldots\), length of \((\mathbb{B})-1\}\)
        If \(\mathbb{B}_{i}=1 \wedge s=\) False
            \(\mathbb{P}_{\text {ON }} \leftarrow \mathbb{P}_{\text {ON }} \cup\{i\}\)

    \(s=\) True
    Else if \(\mathbb{B}_{i}=0 \wedge s=\) True
        \(\mathbb{P}_{\text {OFF }} \leftarrow \mathbb{P}_{\text {OFF }} \cup\{i\}\)
        \(s=\) False
Output: \(\mathbb{P}_{\text {ON }} \cup \mathbb{P}_{\text {OFF }}\)

The resulting pulse detection is shown in Fig. 5. This diagram shows the voltage and the current waveform excerpt with the generated signal for pulse detection. The excerpt clearly shows that the pulse detection system has accurately detected the start and the end of each discharge. This enables the characteristics of a discharge to be determined for each detected signal. Subsequently, each detected discharge was analysed for its characteristics, such as open-circuit voltage 𝕊i= and average discharge voltage (𝔽i−L…𝔽i) as shown in Fig. 1. The analysed characteristics of the discharges, as well as the applied process parameters, name of the measurement file, and pulse position were stored in a database. A total of 290,774 discharges from all experiments were stored in this database.

[picture]

Fig. 5. Excerpt from the measurement of the top: voltage waveform and bottom: current waveform and the resulting pulse detection, highlighted in purple color

Fig. 5. Excerpt from the measurement of the top: voltage waveform and bottom: current waveform and the resulting pulse detection, highlighted in purple color

In addition, the MRR and TWR were determined. For this purpose, the electrode and the workpiece were measured on a microbalance XA 210.4Y PLUS from RADWAG Waagen GmbH before and after the experiments.

Classification. An unsupervised classification was selected in order to obtain correlations from the characteristics of the discharges. This means that the pulses were not manually labelled to predict a certain result. Instead, the aim was to cluster identical pulses and identify correlations with the performance of the machining process.

The specific characteristics of a discharge were utilised as features as shown in Table 3. To improve the model, the data was pre-processed. Therefore, a standardisation of the features was applied. Standardisation is a transformation that removes the mean value of the data for each feature. The data was then scaled by dividing the features by their standard deviation. After standardising the data, the mean equals to zero and the standard deviation to one [12].

Table 3. Features for the classification for each discharge

Discharge | ûi | ūe | īe | îe | ti | td | te
D₁ | ûi₁ | ūe₁ | īe₁ | îe₁ | ti₁ | td₁ | te₁
D₂ | ûi₂ | ūe₂ | īe₂ | îe₂ | ti₂ | td₂ | te₂
D₃ | ûi₃ | ūe₃ | īe₃ | îe₃ | ti₃ | td₃ | te₃
... | ... | ... | ... | ... | ... | ... | ...
Dn | ûin | ūen | īen | îen | tin | tdn | ten

For a better overview and visualisation of the different clusters, principal components analysis (PCA) was utilised to summarise the correlated variables that explain most of the variability in the original data [13]. With the implementation of PCA, the dimension was reduced from 7 features to 3 features. In this case, Kernel PCA with a gaussian radial basis function was applied to maximize the variance in the dataset.

K-means clustering was applied to cluster the data. The K-Means algorithm clusters the data by separating the data points into 𝔹U,𝕄,𝕊 groups with equal variance. For this purpose, centroids were selected for the clusters with the aim of minimising the within-cluster sum-of-squares (WCSS) criterion [12]. The Python library scikit-learn was applied to calculate K-Means. In order to determine the optimal number of clusters, the elbow method and the silhouette score IP was utilised. Therefore, WCSS and the silhouette score were determined for a number of clusters ranging from 2 to 10 . According to the elbow method, the number of clusters is optimal when a clear bend, the so-called elbow, is visible [10]. Accordingly, the 3 calculated principal components, derived from the first parameter set, were clustered into groups of 2 to 10 clusters, as seen in the diagram in Fig. 6. A significant elbow can be noticed at a number of clusters of 4 . To verify the elbow method, the silhouette score for each group of clusters was calculated. The silhouette score is calculated based on the mean intra-cluster distance and the mean distance to the nearest cluster for each data point [14]. The silhouette score ranges from -1 to 1 , with values close to 1 indicating a well clustered group and values close to 0 indicate overlapping clusters. The diagram in Fig. 6 shows that the peak of the silhouette score is reached at 0.74 with 4 clusters. Thus, the discharges were grouped into 4 clusters.

[picture]

Fig. 6. WCSS and silhouette score for the number of clusters from 2 to 10

Fig. 6. WCSS and silhouette score for the number of clusters from 2 to 10

The discrete discharges from the parameter set 02 with their principal components PC1, PC2 and PC3 are shown in Fig. 7. The specific clusters are highlighted in colour and marked with their centroid. A total of 38,167 discharges were recorded from the parameter set 02 . The diagram shows a clear separation of the individual clusters. To determine the discharge type corresponding to the individual clusters, the 100 discharges with the smallest euclidian distance to the corresponding cluster centroid are shown in Fig. 8.

[picture]

Fig. 7. Discrete discharges of parameter set 02 with their derived principal components

Fig. 7. Discrete discharges of parameter set 02 with their derived principal components

[picture]

Fig. 8. Cumulative data points from 100 discharges for the specific cluster a) C1 b) C2
c) C3 and d) C4

Fig. 8. Cumulative data points from 100 discharges for the specific cluster a) C1 b) C2 c) C3 and d) C4

Fig. 8a) shows discharges from cluster C4, which have a recognisably high ignition delay time of approximately 1.43 ms . This cluster is therefore referred to as a "Delayed Ignition". Fig. 8b) shows discharges from cluster C2. Here, the ignition delay time varies around approximately 0.1 ms . The open-circuit voltage is approximately -155 V and the discharge current is approximately -17.2 A . These values are within the range of a normal discharge and therefore the cluster is referred to as a "Normal Discharge". Fig. 8c) shows discharges for cluster C3. Here, the sudden extinction of the discharge is noticeable at an ignition delay time of approximately 0.007 ms . This discharge is known as a transient arc [15]. Hence, this cluster is referred to "Arc". The discharges in the last cluster C4 are shown in Fig. 8d). It is noticeable that the discharge voltage is nearly 0 and that no open-circuit voltage occurred. Based on this, the cluster is classified as "Short".

Fig. 9 illustrates the count of classified pulses per measurement over the process time. The diagram shows that short circuits rarely occur up to a time of 50 s and that there is a consistent distribution of pulse clusters with normal, arcing and delayed discharges. The number of short circuits and transient arcs increases after 70 s and remains at a high percentage after 150 s . In contrast, the percentage of normal discharges decreases. This is partly due to the deterioration in flushing conditions, as flushing and readjustment of the electrode were deliberately omitted. Furthermore, the fuzzy logic was deactivated. The short circuits could also be caused by an excessively high feed rate. The delayed ignitions remain at a constant percentage throughout the process time. The resulting peak values in the count of discharges could be due to the system performing a dummy movement of the spindle in Z -direction at regular intervals to lubricate the ball screw.

[picture]

Fig. 9. Number of classified pulses as function of the process time DC

Fig. 9. Number of classified pulses as function of the process time D C

The results of all parameter sets derived from a total of 290,774 discharges are summarised in a correlation matrix, as seen in Fig. 10. In the correlation matrix, the blue cells indicate an indirect proportionality and the red cells indicate a direct proportionality. The correlation coefficient indicates the degree of correlation between the values between -1 and 1 , where the value 1 indicates a strong direct proportional correlation and the value -1 indicates a strong indirect proportional correlation. At the left-hand matrix, the impact of the discharge types to the MRR and TWR are illustrated. The arcing discharge exhibits the strongest direct proportional influence on the MRR, as indicated by a correlation coefficient of 0.56 . This implies that an increase in the number of arcing discharges leads to an increasing MRR. At an arcing discharge proportion of 5.1%, an MRR of 𝔹U and a TWR of 𝔹I were determined. In comparison, with a proportion of ℙON arcing discharges, an MRR of ℙOFF and a TWR of 𝔹 were determined.

[picture]

Fig. 10. left: Impact of the discharge types to the MRR and TWR, right: Influence of input current 𝔹U and pulse on time s to the discharge types

Fig. 10. left: Impact of the discharge types to the MRR and TWR, right: Influence of input current 𝔹 U and pulse on time s to the discharge types

However, this is at the cost of an increase in TWR. In contrast, the MRR decreases with an increasing number of short discharges. This correlates with the assumption that short circuits do not lead to material removal. The right-hand matrix displays the influence of the selected process parameters input current m and pulse on time 𝔹U. From this, it can be deduced that n has a significant influence on the occurrence of short discharges and arcing. As 𝔹I increased, fewer short circuits occurred, but more arcing. In comparison, the input current influences short circuits, normal discharges and delayed discharges. As m increased, more short circuits and fewer normal discharges and delayed discharges occurred.

Conclusion

The original version of this paper is available on https://www.scientific.net/SSP.391.43.pdf

This paper aims to extend the evaluation of the process of electric discharge machining. Therefore, this paper presents a detection method for detecting discharges and subsequently classifying the discharge type. Furthermore, experiments were conducted by varying the input current and the pulse on time. Solid carbide WC-6Co was chosen as the workpiece material and copper was chosen as the electrode material. Throughout the experiments, the current and voltage signals were measured via an oscilloscope, and the material removal rate (MRR) and tool wear rate (TWR) were determined. Discharges were detected from the current and voltage measurements using a statistical dispersion method. Subsequently, the discharges were grouped into 4 clusters applying the unsupervised classification method K-means.

Based on the results, it was concluded that a higher proportion of arcs results in a higher MRR, but with increased TWR. Thus, with a proportion of n arcing discharges, an MRR of m and a TWR of n were determined. In comparison, with a proportion of 𝔹i arcing discharges, an MRR of 𝔹i and a TWR of 𝔹 were determined. Moreover, an analysis of all discharges indicated that as the pulse on time increases, the proportion of arcing discharges increases and short circuits decrease, which can lead to an increase in the MRR.

Based on the results, further investigations can be carried out. For instance, the influence of flushing conditions and electrode readjustment on the discharge types can be investigated. As well as repeating the experiments with fuzzy logic enabled. In addition, a time-dependent adjustment of the pulse on time and input current can be carried out.

Acknowledgments

The original version of this paper is available on https://www.scientific.net/SSP.391.43.pdf

This project is funded by the State of Saxony-Anhalt with funds from the European Union under the European Regional Development Fund under the project number ZS/2023/12/182059.

The machine tool Mitsubishi SG12R was funded by the German Research Foundation (DFG) with project number 509924008.

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The original version of this paper is available on https://www.scientific.net/SSP.391.43.pdf

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Solid State Phenomena, ISSN: 1662-9779, Vol. 391, pp 55-63

doi: 10.4028/p-fqy4QR

© 2026 The Author(s). Published by Trans Tech Publications Ltd, Switzerland.

Submitted:
2026-01-16

Revised:
2026-02-20

Accepted:
2026-02-23

Online:
2026-04-20

Concurrent Effect of Laser Texturing and Resin Pre-Coating on the Performance of CFRP Single Lap Joints

The original version of this paper is available on https://www.scientific.net/SSP.391.55.pdf

Gianluca Parodo 1,a*, Luca Sorrentino 2,b, Sandro Turchetta 2,b, Giuseppe Moffa 2, b

¹Department of Engineering and Sciences, Faculty of Technological and Innovation Sciences, Universitas Mercatorum, Piazza Mattei 10, Rome, 00186, Italy

2 Department of Civil and Mechanical Engineering, University of Cassino and Southern Lazio, Cassino,04023, Italy

a  gianluca.parodo@unimercatorum.it, b  luca.sorrentino@unicas.it, c  sandro.turchetta@unicas.it, c  giuseppe.moffa@unicas.it

Keywords: Adhesive bonding, Single lap joints (SLJ), Surface treatments.

Abstract

Adhesive bonding is increasingly used in lightweight structural applications, though its effectiveness for carbon fiber reinforced polymers is often limited by low surface energy. This study investigates the concurrent effect of infrared laser surface texturing and resin pre-coating on the mechanical performance of carbon fiber-reinforced polymer single-lap joints. Using a CO2 laser system, micro-dimples were generated on the substrate to promote mechanical interlocking, while various concentrations of an epoxy resin ( 5wt%,10wt%, and 20wt% ) were applied as precoating to enhance surface wetting and interfacial continuity. To validate the experimental findings, a numerical investigation was performed using a Cohesive Zone Model implemented within a finite element framework. This model utilized a bilinear traction-separation law and a quadratic delamination criterion to predict the progressive debonding and failure of the interface. Experimental results indicate that the 10wt% resin pre-coating concentration provides the highest lap shear strength.

Introduction

The original version of this paper is available on https://www.scientific.net/SSP.391.55.pdf

Adhesive bonding is increasingly adopted in lightweight structural applications as an alternative to traditional mechanical fastening, particularly in the field of fiber-reinforced polymer composites. Compared to bolted or riveted joints, adhesively bonded joints enable a more uniform stress distribution, avoid stress concentrations associated with holes, and preserve the integrity of loadbearing fibers [1]. Additional advantages include weight reduction, improved fatigue performance, sealing capability, and enhanced design flexibility, making adhesive bonding particularly attractive for aerospace, automotive, and high-performance structural components [2,3].

For polymer matrix composites, however, joint performance is not governed solely by the bulk properties of the adhesive or adherends, but rather by the quality of the interfacial region. In this context, synergistic strategies combining surface modification and interfacial engineering have gained increasing attention. Rather than relying on a single surface treatment, the concurrent action of topographical modification, chemical activation, and improved wetting can significantly enhance adhesion mechanisms such as mechanical interlocking, physico-chemical bonding, and interphase continuity. This synergistic approach is especially relevant for carbon fiber reinforced polymers (CFRP), where the intrinsic heterogeneity of the material adds further complexity to the bonding process [4,5].

A major challenge in bonding polymer-based substrates lies in their intrinsically low surface energy, which limits adhesive wetting and interfacial adhesion. Thermoset matrices, despite offering good mechanical performance, often exhibit chemically inert surfaces after curing, while thermoplastic matrices present additional difficulties related to chain mobility and surface contamination. As a consequence, adhesively bonded joints on untreated polymer composite surfaces frequently show premature interfacial failure and poor reproducibility [6,7].

To overcome these limitations, a wide range of surface pre-treatments has been developed, including mechanical abrasion, grit blasting, peel ply techniques, chemical etching, and plasma treatments. While these methods can effectively increase surface roughness or surface energy, they also present significant drawbacks from an industrial perspective. Traditional mechanical treatments are often non-selective, generate large amounts of dust and debris, and strongly depend on operator skill, leading to variability in joint performance. Chemical and plasma-based treatments, on the other hand, may involve hazardous substances, limited process stability, or high operational costs, which restrict their scalability and industrial adoption.

In recent years, laser surface texturing has emerged as a promising alternative for the pre-treatment of composite adherends prior to adhesive bonding. Laser processing offers several advantages, including high repeatability, precise spatial control, and easy automation. By tailoring laser parameters, it is possible to selectively modify surface morphology, increasing roughness and creating micro-scale features that promote mechanical interlocking with the adhesive. As a noncontact and highly controllable process, laser texturing is particularly appealing for advanced composite structures and complex geometries [ 8,9 ].

Despite these advantages, laser surface treatments (especially those based on infrared (IR) wavelengths) are not free from limitations. The interaction between laser radiation and CFRP may induce localized thermal effects, potentially affecting the fiber-matrix interface or causing matrix degradation and resin recession [10]. Such damage, if not properly controlled, can compromise load transfer mechanisms at the microscale and negatively impact joint durability [11]. Moreover, highly textured surfaces may hinder effective adhesive penetration, particularly when high-viscosity structural adhesives are employed. In this context, resin pre-coating (RPC) has recently been proposed as an effective strategy to further enhance adhesion on pre-treated surfaces [12]. The application of a diluted resin layer prior to bonding can improve surface wetting, promote better penetration into laserinduced micro-features, and enhance interfacial continuity between the adherend and the adhesive layer [13]. When combined with laser texturing, RPC has the potential to mitigate some of the limitations associated with adhesive viscosity and thermal surface modification, leading to a more effective and robust interphase [8,14]. However, despite its promising potential, the combined application of infrared laser texturing and resin pre-coating on CFRP surfaces remains insufficiently explored, particularly in joint configurations representative of industrial practice [14].

The goal of the present work is to investigate the concurrent effect of infrared laser surface texturing and resin pre-coating (RPC) on the mechanical performance of adhesively bonded carbon fiber reinforced polymer (CFRP) single lap joints. The study aims to assess whether the combination of a controlled laser-induced surface morphology and a thin pre-applied resin layer can act synergistically to enhance interfacial adhesion, overcome limitations related to adhesive viscosity, and improve joint strength and damage tolerance.

Materials and Methods

The original version of this paper is available on https://www.scientific.net/SSP.391.55.pdf

The material investigated in this study was a unidirectional carbon fiber reinforced polymer (CFRP) prepreg known as Cytec T152/x751/135 and supplied by Cytec Engineered Materials, (USA), consisting of carbon fibers embedded in an epoxy resin matrix. Laminates were manufactured using a vacuum bagging technique and cured in a convection oven. The technological layout is illustrated in Figure 1.

[picture]

Fig. 1. Technological layout for the adopted vacuum bagging process.

Fig. 1. Technological layout for the adopted vacuum bagging process.

The curing of laminates and surface treatments was carried out in a Nabertherm High Temperature Oven (model NA 120/45), which provides a controlled temperature environment up to 450∘C with forced air circulation. The curing cycle consisted of a heating ramp of 2∘C/min up to 125∘C, followed by an isothermal dwell of 2 h at the same temperature. Vacuum was applied from the beginning of the cycle. After curing, the laminates were allowed to cool down naturally to room temperature inside the oven.

After curing, the laminates were demolded and visually inspected to ensure the absence of macroscopic defects. Prior to specimen preparation, all laminate surfaces were degreased using acetone to remove contaminants and release agent residues. Adherends for single lap joint manufacturing were then cut from the cured laminates according to the geometrical requirements specified by ASTM D5868.

Laser surface pre-treatment was performed using a CO2 laser system (GIOTTO CO2, nominal power 25 W ). The laser texturing strategy was based on the generation of a regular array of microdimples, specifically designed to promote mechanical interlocking while minimizing heat-affected zone (HAZ) formation and potential damage to the fiber-matrix interface. This surface patterning approach was selected based on previous optimization studies by the authors, which demonstrated its effectiveness in enhancing adhesive bonding performance while preserving the integrity of CFRP substrates [15,16].

The laser-induced dimples exhibited an average depth of approximately 10μ m, comparable to the thickness of the superficial resin-rich layer of the CFRP laminate. Laser parameters were selected from previous studies and reported in Table 1 [8,15].

Table 1. Adopted laser process parameters.

Parameters | Value | Units
Power | 10 | W
Wavelength | 10.6 | μm
Fluence | 1.44 | J/cm²
Frequency | 25 | kHz
Writing speed | 5950 | mm/s
Pulse duration | 18 | μs
Focal distance | 200 | mm
Laser spot diameter | 200 | μm
Mode | TEM₀₀ | -

The resulting textured surfaces exhibited localized topographical modification without evidence of extensive matrix degradation or fiber exposure, making them suitable for subsequent resin precoating and adhesive bonding operations. A sketch of the texture and a SEM image of the CFRP surface subjected to laser texturing are reported in Figure 2.

[picture]

Fig. 2. Laser texturing strategy: a) adopted scheme (dimples in square grid); b) SEM observation for laser textured specimen

Fig. 2. Laser texturing strategy: a) adopted scheme (dimples in square grid); b) SEM observation for laser textured specimen

Resin pre-coating (RPC) was performed using the structural epoxy adhesive EA9309.3NA. The adhesive was diluted to obtain three different resin concentrations, namely 5wt%,10wt%, and 20wt%, in order to investigate the effect of resin content on the surface condition of laser-textured CFRP adherends. The resin concentrations ( 5wt%,10wt%, and 20wt% ) were selected based on literature indications and on previous ENF investigations performed by the authors [8], where the effect of resin pre-coating on interfacial fracture behaviour was analysed. The diluted resin solutions were applied by brushing onto the laser-textured surfaces, aiming to promote resin penetration into the laser-induced micro-dimples and to improve interfacial continuity prior to bonding. Following application, solvent evaporation was carried out in a convection oven at 60∘C for 30 min . After drying, the pre-coated adherends were immediately used for joint assembly.

Single-lap joints were manufactured in accordance with ASTM D5868 using the epoxy adhesive EA9309.3NA. After assembly, the specimens were positioned within a dedicated bonding fixture, previously developed and validated by the authors [17], designed to constrain the adherends and to maintain a controlled bondline release agent throughout polymerization. This setup ensured geometrical repeatability of the single-lap joint configuration and minimized variability associated with adhesive flow. The joints were then cured at room temperature for 5 days prior to mechanical testing. This curing protocol was selected to ensure complete adhesive crosslinking.

Lap shear tests were performed in accordance with ASTM D5868 using a universal testing machine with a maximum load capacity of 100 kN . The tests were conducted under displacement control at a crosshead speed of 1.3 mm/min, as specified by the standard and consistently adopted in previous investigations by the authors [16,17], ensuring methodological continuity and enabling comparison of results. During testing, load-displacement curves were recorded for all specimens to evaluate the mechanical response and ultimate lap shear strength. All tests were carried out at room temperature. For completeness, the experimental plan summarizing the investigated surface conditions, resin pre-coating concentrations, and testing activities is reported in Table 2.

Table 2. Experimental plan

Factors | N° levels | Levels

Surface pretreatment
5
Degreasing
Laser texturing
Laser Texturing + RPC 5wt%
Laser Texturing + RPC 10wt%
Laser Texturing + RPC 20wt%

Replication | 5 |
Total | 15 |

To model the progressive debonding of the SLJ, a Cohesive Zone Model (CZM) was implemented within a finite element method. The method used is the same proposed by the authors in [18]. This approach characterizes the Fracture Process Zone (FPZ) through two superimposed surfaces governed by a bilinear traction-separation law, which describes the gradual degradation of the interface as separation increases [19]

GI,c=12·tn0·snf=12·Kn·sn0·snfGII,c=12·tt0·stf=12·Kt·st0·stf

(1)(2)

The maximum traction values in the normal ( tn0 ) and tangential ( tt0 ) orientations were established based on the documented interfacial strengths of the adhesive substance. Subsequently, the related final separations ( snf,stf ) were determined utilizing the bilinear cohesive law assumption, as delineated in Eq. (1) and Eq. (2). The initial stiffness coefficients ( Kn,Kt ) were calculated in accordance with Eq. (3):

K=t0s0(3)

The onset of damage is predicted using a quadratic delamination criterion, which assumes initiation occurs when the quadratic sum of the stress ratios across the three primary modes reaches a critical threshold:

(TI Tm,I)2+(TII Tm,II)2+(TIII Tm,III)2=1(4)

For subsequent crack growth, a power law criterion is adopted. The numerical implementation utilizes zero-thickness interface elements placed specifically at the adherend-adhesive boundaries to focus the failure simulation on the adhesive region. The values implemented in the model are taken from [8]:

(GIGI,c)+(Gshear GS,c)=1(5)

Finally, to ensure numerical stability and accuracy, the mesh was optimized following the Turon analysis, ensuring that at least five elements span the calculated cohesive zone length [20,21].

Results and Discussions

The original version of this paper is available on https://www.scientific.net/SSP.391.55.pdf

The lap shear test results highlight a significant influence of surface preparation on the mechanical performance of CFRP joints, as shown in Figure 3. Joints manufactured on degreased surfaces exhibit lower lap shear strength and an early, abrupt failure at relatively small displacement levels. This behavior may be associated with limited interfacial adhesion and a predominantly adhesive failure mode, which is consistent with the low surface energy typically exhibited by untreated polymeric matrices.

Laser surface texturing leads to an increase in lap shear strength compared to the degreased condition, suggesting that the induced surface topography could promote mechanical interlocking at the adhesive-adherend interface. However, the overall displacement at failure remains comparable, indicating that laser texturing alone might not be sufficient to ensure complete adhesive penetration into the laser-induced micro-features, particularly when high-viscosity structural adhesives are employed.

The introduction of resin pre-coating (RPC) in combination with laser texturing results in a further enhancement of joint performance, indicating a possible synergistic effect between the two surface treatments. In particular, the RPC concentration of 10wt% is associated with the highest lap shear strength and a larger displacement at failure compared to the other investigated conditions. This behavior could be attributed to improved resin infiltration within the laser-generated micro-cavities and to enhanced interfacial continuity between the adherend and the adhesive layer.

Conversely, joints produced with RPC concentrations of 5wt% and 20wt% exhibit lower performance than the optimal condition. At lower resin concentrations, the amount of pre-coating material might be insufficient to fully saturate the textured surface, whereas at higher concentrations the formation of a resin-rich layer could partially mask the laser-induced topography, thereby reducing the effectiveness of mechanical interlocking. The non-monotonic trend observed as a function of RPC concentration suggests the existence of an optimal resin content that balances infiltration capability and preservation of surface morphology.

Overall, the analysis of the load-displacement curves suggests that the combined application of laser texturing and resin pre-coating may contribute to an increase in the load-bearing capacity of the joints and to a modification of the failure mechanisms, promoting a transition from predominantly adhesive failure towards more complex and progressive damage modes.

[picture]

Fig. 3. Average lap shear strength results

Fig. 3. Average lap shear strength results

The comparison between the numerical predictions and the experimental findings for the various SLJ configurations is presented in Figure 4, while an example of numerical result is reported in Figure 5. The developed Cohesive Zone Model (CZM) demonstrates high reliability in characterizing the mechanical response of the joints across all surface conditions. As illustrated in the comparison, the numerical failure loads closely align with the experimental averages for the different configurations investigated.

Table 3. Comparison between experimental and numerical results

Surface pretreatment | Experimental results [MPa] | Numerical results [MPa] | Error
| average | dev.st. | c.v. |  |
Degreasing | 5.04 | 0.51 | 10.2% | 5.21 | 3.44%
Laser texturing | 9.40 | 0.80 | 8.5% | 9.38 | -0.25%
Laser texturing + RPC 5wt% | 10.13 | 0.72 | 7.1% | 10.62 | 4.88%
Laser texturing + RPC 10wt% | 11.60 | 0.95 | 8.2% | 12.19 | 5.11%
Laser texturing + RPC 20wt% | 10.60 | 0.86 | 8.1% | 11.06 | 4.37%

[picture]

Fig. 4. Comparison of numerical predictions and experimental results for various single-lap joint (SLJ) configurations

Fig. 4. Comparison of numerical predictions and experimental results for various single-lap joint (SLJ) configurations

[picture]

Fig. 5. Finite element model of the single-lap joint (SLJ) specimen showing the von Mises stress distribution (MPa)

Fig. 5. Finite element model of the single-lap joint (SLJ) specimen showing the von Mises stress distribution (MPa)

Conclusions

The original version of this paper is available on https://www.scientific.net/SSP.391.55.pdf

This work investigated the concurrent effect of infrared laser surface texturing and resin precoating (RPC) on the mechanical performance of adhesively bonded CFRP single lap joints. Compared to degreased surfaces, laser texturing resulted in an increase in lap shear strength, confirming the beneficial role of controlled surface topography in promoting mechanical interlocking.

The combination of laser texturing and RPC led to a further improvement in joint performance, highlighting a synergistic interaction between the two surface treatments. Among the investigated conditions, the RPC concentration of 10wt% provided the most effective balance between resin infiltration and preservation of the laser-induced surface morphology, resulting in the highest lap shear strength and a more progressive failure response.

The non-monotonic trend observed as a function of RPC concentration suggests the existence of an optimal process window, beyond which excessive resin content may partially mask the laserinduced micro-features and reduce the effectiveness of mechanical interlocking. This finding underlines the importance of controlling both surface morphology and pre-coating formulation when combining laser-based treatments with resin pre-application.

Moreover, a numerical Cohesive Zone Model (CZM) to characterize the Fracture Process Zone (FPZ) was used. The model, governed by a bilinear traction-separation law, successfully captured the failure loads for each investigated surface morphology. By employing a quadratic delamination criterion for damage initiation and a power law for crack growth, the numerical predictions remained consistent with experimental dispersion.

Overall, the results demonstrate that the combined application of infrared laser texturing and resin pre-coating represents a promising and scalable surface-treatment strategy for improving the performance of adhesively bonded CFRP joints in configurations representative of industrial practice. Future work will focus on a more detailed microstructural characterization and on extending the approach to different joint geometries and loading conditions.

References

The original version of this paper is available on https://www.scientific.net/SSP.391.55.pdf

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This ePub file was created using machine-generated methods and automated software to convert the source text into an ePub format and may contain formatting inconsistencies from the original version. Formatting may vary depending on the device used.

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Solid State Phenomena, ISSN: 1662-9779, Vol. 391, pp 65-77

doi: 10.4028/p-vSvpd4

© 2026 The Author(s). Published by Trans Tech Publications Ltd, Switzerland.

Submitted:
2026-01-16

Revised:
2026-02-16

Accepted:
2026-02-18

Online:
2026-04-20

Enabling the Reuse of Single-Lap Adhesive-Bonded Joints through Laser Cleaning and Texturing

The original version of this paper is available on https://www.scientific.net/SSP.391.65.pdf

Vincenzina Siciliani 1,a*, Matteo Benvenuto 2, b, Chiara Mandolfino 2,c, Barbara Reggiani 1,3,d and Leonardo Orazi 1,4,e

¹DISMI - University of Modena and Reggio Emilia, via Amendola 2, 42122, Reggio Emilia, Italy

2 DIME - University of Genoa, Via Opera Pia 15, 16145, Genoa, Italy

3 INTERMECH - University of Modena and Reggio Emilia, Piazzale Europa 1, 42124, Reggio Emilia, Italy

4 EN&TECH - University of Modena and Reggio Emilia, Piazzale Europa 1, 42124, Reggio Emilia, Italy

a*  vincenzina.siciliani@unimore.it, b  matteo.benvenuto@edu.unige.it, c  chiara.mandolfino@unige.it, bbarbara.reggiani@unimore.it, eleonardo.orazi@unimore.it

*corresponding author

Keywords: laser ablation, bonding joint, rebonding.

Abstract

Adhesive bonding is widely employed for joining complex-shaped components due to its ability to distribute stresses uniformly, preserve material integrity, and eliminate mechanical fasteners. However, the permanent nature of conventional adhesives and the difficulty of clean disassembly hinder the reuse of bonded joints, limiting recyclability and sustainability. This study investigates laser cleaning as an enabling technology for re-bonding single-lap joints. An ultrashort pulsed laser was applied to remove residual epoxy adhesive from AA6061 aluminum alloy substrates after debonding. Surface characterization revealed that laser ablation produces micrometric roughness, and laser processing can be further tailored to generate laser-induced periodic surface structures (LIPSS), influencing morphology, chemistry, and wettability. Mechanical testing demonstrated that laser cleaning also improves the tensile lap-shear strength of re-bonded untreated joints. These findings confirm that laser cleaning effectively restores substrate surfaces and, when combined with controlled texturing, can enhance bonding performance. The proposed approach supports repair and recycling strategies, contributing to extended component lifecycles and circular economy objectives.

Introduction

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Adhesive bonding has become a key joining technology in modern engineering, particularly in sectors such as aerospace, automotive, and transportation, where lightweight structures and complex geometries are essential. Unlike traditional mechanical fastening or welding, adhesive bonding offers several advantages: it enables uniform stress distribution, reduces stress concentrations, preserves the integrity of the base material, and eliminates the need for additional components such as rivets or mechanical fasteners [1-4]. Furthermore, adhesive joints contribute to weight reduction, vibration damping, and corrosion resistance, making them highly attractive for structural applications [5]. Among the various materials used in these industries, aluminum alloys stand out due to their high strength-to-weight ratio, corrosion resistance, and excellent machinability. Alloys such as AA6xxx and AA7xxx series are widely employed in automotive body structures, aircraft components, and marine applications [6,7]. However, the performance of adhesive joints strongly depends on the surface condition of the adherends, which governs adhesion mechanisms at the interface [8].

Surface preparation is therefore a critical step in adhesive bonding. Conventional methods include mechanical abrasion (e.g., grit blasting, sanding) and chemical treatments (e.g., etching, anodizing), which aim to increase surface roughness, remove contaminants, and modify surface chemistry to enhance wettability [9-11]. While effective, these methods present drawbacks such as limited
accuracy, environmental concerns, and difficulties in automating the process [12]. In this context, laser-based surface treatments have emerged as a promising alternative, offering high precision, flexibility, and environmental sustainability [13,14]. Laser texturing has been extensively investigated for its ability to create controlled micro and nano-scale features that increase the effective bonding area and promote mechanical interlocking [15-19]. Studies have shown that laser-induced grooves, dimples, and periodic structures can significantly improve lap-shear strength and shift failure modes from adhesive to cohesive [20-22]. Moreover, laser processing can simultaneously modify surface chemistry and wettability, further enhancing adhesion [23].

Despite these advances, the reuse of adhesive-bonded joints remains a major challenge for sustainable manufacturing. Conventional adhesives are generally considered permanent, and disassembly typically requires destructive methods such as mechanical cutting, thermal degradation, or chemical dissolution, which often damage the substrates and hinder their reuse [24]. This limitation conflicts with the principles of the circular economy, which aims to minimize waste and maximize resource efficiency by extending product lifecycles and enabling component recovery [25]. In highvalue sectors such as aerospace and automotive, where aluminum components are costly and energyintensive to produce, the inability to reuse bonded joints represents a significant environmental and economic burden. Therefore, developing technologies that allow clean disassembly and re-bonding is essential to reduce raw material consumption, lower greenhouse gas emissions, and comply with increasingly stringent sustainability regulations [26].

To address this issue, laser cleaning represents an innovative solution for post-debonding surface preparation. Laser cleaning employs a continuous or pulsed laser to ablate residual adhesive and contaminants: the pulsed laser has the major advantage of not damaging the underlying substrate [27,28]. Compared to abrasive or chemical cleaning, laser cleaning offers several advantages: it is a non-contact process, leaves no chemical residues, minimizes substrate alteration, and can be easily automated [29]. Furthermore, by adjusting process parameters, laser cleaning can generate microtextures that improve wettability and create favorable conditions for re-bonding [30]. Recent studies have shown that the chemical effect of laser cleaning on aluminum alloys leads to increased mechanical performance compared to untreated surfaces [7,8]. For instance, Morello et al. [8] reported that laser cleaning alone increased the apparent shear strength of re-bonded AA6082 joints compared to untreated samples, while a combined approach of laser texturing followed by cleaning achieved even higher strength values. Similarly, Min et al. [7] highlighted that laser-induced modifications in morphology, chemistry, and wettability are key factors influencing adhesion performance.

These findings suggest that integrating laser cleaning, in the sense of ablating adhesive residues together with the chemical treatment produced by the laser, into the lifecycle of adhesive-bonded assemblies can significantly enhance repairability and recyclability, reducing waste and supporting sustainable manufacturing strategies. By enabling the reuse of bonded joints, laser cleaning contributes to resource efficiency and aligns with global efforts to reduce environmental impact in manufacturing. In particular, the use of lasers could perform the dual function of removing adhesive residues and pre-treating the surface for subsequent re-bonding.

The present work investigates the application of ultrashort pulse laser texturing and cleaning for enabling the reuse and improving the performance of single-lap adhesive-bonded joints made of Aluminum AA6061. The study compares two parallel methodologies. In the first case, three different pre-treatments are carried out on the substrates to increase the adhesive strength, followed by the adhesive bonding and the evaluation of the shear strength. The second methodology involves the pretreatment of the surfaces, bonding the joints by adding a component to enable the controlled detachment, detaching the joints, and removing the residual epoxy adhesive after detachment. Finally, the joints are re-bonded, and the strength of the rebonding is evaluated. The resulting surface morphology is characterized to relate it to mechanical performances.

The goal is to assess the feasibility of laser texturing and cleaning as a robust and environmentally friendly method for extending the service life of adhesive-bonded structures and advancing circular economy principles in structural joining technologies.

Materials and Methods

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Substrates and adhesive used.

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The single-lap joint specimens were prepared in accordance with ASTM D1002 [31], which defines the geometry and the procedures for preparing specimens for shear strength tests of adhesive joints. In particular, the standard specifies the use of substrates with dimensions of 100×25×1.6 mm, overlapped with a length of 12.5 mm . The substrates were made of 6061-T6 aluminum alloy, an Al-Mg−Si alloy widely used in structural applications due to its favorable combination of mechanical strength, light weight, corrosion resistance, and workability. The substrates were bonded using a thixotropic two-component structural epoxy adhesive ( 3MTM EPX DP490), characterized by high mechanical strength, good toughness, and the ability to maintain a controlled adhesive layer thickness during assembly. Thanks to its formulation, the adhesive is suitable for bonding metals and for structural applications subjected to both static and dynamic loads. The main technical specifications of the adhesive and substrates used are reported in Table 1.

Table 1. Mechanical properties of 6061-T6 aluminum substrates and main characteristics of the 3M TM  EPX DP490 structural adhesive used for single-lap joints.

Substrate | AA6061-T6 | Adhesive | 3M™ DP490
Density [g/cm³] | 2.70 | Chemical composition | Epoxy
Tensile strength [MPa] | 327.1±0.48 | Mix ratio | 2:1
Yield strength [MPa] | 273.3±4.21 | Work life @ 23°C | 1.5 h
Elastic modulus [GPa] | 36.09±1.95 | Handling time @ 23°C | 4÷6 h
Elongation at break [%] | 18.6±0.3 | Full strength time @ 23°C | 7 days

Surface preparation methods.

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Substrates underwent different surface preparation protocols for comparison. The first consists of degreasing (DG) with acetone-soaked absorbent paper for the reference condition samples. The second is mechanical abrasion (AB) using a Scotch-Brite TM 7447 PRO pad, which is the standard industrial pretreatment. The third is an alternative laser texturing treatment (TXT).
Particular attention is given to the role of laser-induced surface features (LIPSS) in promoting adhesion. The laser employed is a picosecond Ekspla Atlantic 50,≈ with a pulse duration of about 10 ps. The wavelength used is 1064 nm . The repetition rate is set at 333 kHz as a compromise between speed and pulse overlapping. The system includes a Raylase Superscan V galvanometric scanning head to ensure precise beam trajectory control, while an F-theta lens provides a focused spot size of roughly 10μ m at the 1/e2 intensity. The chosen scanning strategy is parallel lines, 4.5μ m spaced, scanned at 1200 mm/s to have a homogeneous treated area, having the same energy deposition between and along the scanlines. The power was set at 0.3 W to have regular and clean LIPSS. This strategy was chosen for two main reasons: to increase the hydrophilicity of the base material, primarily through chemical oxidation [32], and to do this with a sufficient areal throughput. The process is carried out in an area of 16x28 mm, to ensure that the gluing area is covered, which is then suitably masked.
Twenty aluminum sheets were textured, for a total of five pairs per two conditions to be tested (bonding with and without the debonding method). For clarity, the entire methodology is summarized in Fig. 1, which explains the bonding and debonding/rebonding phases and where laser texturing and cleaning come into play. In this preliminary study, after cleaning, surface preparation was performed, followed by bonding without debonding methods. Bonding can then be performed using debonding methods to evaluate recyclability over multiple cycles.
Before bonding, all adherends were degreased to remove treatment residues and dirt, taking care that after laser treatment, the samples were only dried with hot air to avoid damaging the microstructure with mechanical actions.

[picture]

Fig. 1. Methodology of the work: a) laser texturing for the bonding process, b) laser texturing and cleaning for the rebonding process.

Fig. 1. Methodology of the work: a) laser texturing for the bonding process, b) laser texturing and cleaning for the rebonding process.

Single-lap joint manufacturing and debonding method used.

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The bonding area was defined by using masking tape to ensure consistent overlap geometry. The adhesive was dispensed using a manual applicator equipped with a 2:1 ratio plunger and static mixer to ensure proper resin-hardener mixing. After assembly, the joints were subjected to uniform pressure to eliminate excess adhesive and to shape the fillet, minimizing stress concentrations at the edges. The specimens were then cured at room temperature for 24 hours, followed by a post-curing cycle at 80∘C for 1 hour, as recommended by the adhesive manufacturer.
An electrothermal debonding technique was developed and evaluated using a metal mesh (MM) embedded within the adhesive layer of single-lap joints. The mesh served as a conductive element to locally heat the joint via the Joule effect, facilitating adhesive degradation and enabling manual disassembly.
The mesh is a woven stainless steel 304 fabric with a wire diameter of 0.05 mm and an aperture of 0.104 mm . It was cut to match the overlapping area and positioned between the aluminum alloy substrates during adhesive application.
Debonding was performed by connecting the mesh to a DC power supply (Kert Cosmo 3000), capable of delivering adjustable voltage and current. Preliminary trials were conducted to determine optimal electrical parameters. The final debonding protocol involved applying a voltage of 3.5−4.0 V and a current of 10 A for 3 minutes, resulting in a surface temperature of approximately 130∘C, measured by a thermocouple. This condition was sufficient to soften the adhesive without exceeding the thermal limits of the substrate. Following treatment, the joints could be manually disassembled without tools, but the mechanical strength was reassessed via lap shear testing.

Laser Cleaning and surface preparation after debonding.

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The methodology comprises the ablation of the thermosetting epoxy resin, while leaving a surface treatment on the aluminum.
The adopted strategy derives from preliminary investigations aimed at identifying the most effective approach for removing the adhesive without damaging the aluminum substrate. Two methods were initially considered: direct and selective ablation of the resin, and detachment by targeting the Al−Al− adhesive interface. Tests performed at 355 nm and 1064 nm revealed that UV radiation rapidly attacked aluminum, while near-IR provided better control over resin removal, despite its limited transparency through the adhesive. At high energy, both wavelengths ablated the resin. Considering material response and process efficiency, the definitive choice was 1064 nm , which offered higher power and energy per pulse, enabling faster and more productive cleaning.

The scanning strategy is a homogeneous treatment, with 400 kHz of repetition rate, scanning 4μ m spaced parallel lines at 2000 mm/s. Power is set at 1.42 W , and the number of passes changes depending on the adhesive thickness, from 20 to 30 . The cleaning is carried out on an area of 16×28 mm . The process is carried out by keeping two dust-suction systems active to aspirate the particulate and fumes generated. Thirty aluminum sheets were cleaned, for a total of fifteen pairs, for three tested conditions (degreased, abraded, and textured after debonding).

Surface morphology and wettability characterization.

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A VHX Keyence digital microscope equipped with two optical objectives (100-500x and 500-2500x) was used to obtain the morphology of the treated samples compared to the untreated and standard ones. The instrument with the supplied optics has a resolution of 0.1μ m, which covers the dimension expected from the LIPSS of the I harmonic.
Sample measurements of roughness were also taken using the Talyscan 150 profilometer to get a quick idea of the topography of the untreated material, treated with different surface preparation methods, then with laser cleaning and subsequent surface preparation methods. The instrument operates using triangulation laser technology. The area to be analysed was scanned with a spacing of 5μ m, and the measured data were processed using the dedicated TalyMap software. In the case of the first surface pretreatment, the area analyzed was 5×5 mm as it was visibly uniform. In the case of laser cleaning and subsequent preparation methods, the area analyzed was extended to 12.5×25 mm to account for all irregularities. The measure is based on an average of three different samples for each condition.
To validate the choice of texture evaluated in this study, wettability was measured using inks with known surface tension. The procedure involved applying the solution with the greatest surface tension onto the sample's surface. If the liquid spreads as a continuous film, the sample surface tension is at least equal to that of the applied solution. Conversely, if the liquid contracts and forms droplets, the sample has a lower surface tension, and a subsequent test using a solution with reduced surface tension is required. The solutions used for surface tension assessment were supplied by Plasmatreat GmbH in accordance with DIN 53914/ISO 304:1985 standards. Ethanol-based test inks (C Series) were employed, covering a surface tension range of 28−72mN/m (surface tension of water) in 2mN/m increments. Since the untreated material appears anisotropic due to the rolling process from which it is obtained, the wettability was evaluated both parallel and perpendicular respect the rolling direction.

Shear strength and bonding performance validation.

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Lap shear tests were performed using an Instron 8802 universal testing machine, and the testing protocol was defined within the Bluehill software. The specimens were loaded at a constant displacement rate of 1.3 mm/min. To ensure pure shear loading, spacers matching the specimen's thickness were placed in the grips to prevent bending in the overlap region.
Lap shear tests were performed on pretreated and untreated samples, on samples prepared for detachment without applying voltage to degrade the joint, and on samples prepared for detachment but after electrothermal degradation, to quantify the effects on mechanical strength.

Results and Discussion

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Surface preparation morphological results.

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The morphology of the degreased (a), abraded (b), and textured (c) samples is shown Fig. 2, in which the signs of lamination of the as-is sample, the random scratches of abrasion, and the uniformity of the laser treatment are clearly visible all over the surface. The sample roughness values for the different treatments are shown in Table 2. As might be expected, mechanical abrasion slightly increases roughness values, while laser texturing, which leaves a micrometric texture, reduces them.

Table 2. Surface roughness values for the different surface pretreatments: Arithmetic mean height (Sa); Root mean square height (Sq); Maximum height (Sz); Skewness (Ssk); Kurtosis (Sku).

Surface pretreatment | Sa [μm] | Sq [μm] | Sz [μm] | Ssk [-] | Sku [-]
Degreased | 3.88±0.30 | 5.06±0.42 | 53.3±5.39 | -0.26±0.35 | 4.26±0.31
Abraded | 5.08±0.33 | 6.45±0.47 | 52.4±5.19 | 0.21±0.19 | 3.28±0.19
Laser textured | 2.11±0.38 | 2.64±0.48 | 20.1±2.48 | 0.05±0.09 | 2.96±0.01

[picture]

Fig. 2. Digital microscope images of the degreased (a), abraded (b), and textured (c) samples.

Fig. 2. Digital microscope images of the degreased (a), abraded (b), and textured (c) samples.

Fig. 3-b shows the periodic surface structures induced by the laser on the Aluminum substrate, at high magnification, in which the regularity of the structures and the size, equal to approximately 1μ m, can be appreciated. In the image, the structures appear almost vertical, perpendicular to the direction of pull, to offer maximum resistance to the shear force. At the macroscopic level, texturing overlaps with the structure of the untreated material, which is visible in Fig. 3-a.

[picture]

Fig. 3. Digital image of untreated (a) and LIPSS textured (b) Aluminum substrate at high magnification.

Fig. 3. Digital image of untreated (a) and LIPSS textured (b) Aluminum substrate at high magnification.

Wettability results.

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A solution with a surface tension of the water was applied to both treated and untreated specimens. Measurements were repeated at least four times for each condition. The surface tension assessment was performed immediately after the laser treatment to capture its initial effect on wettability. After 24 and 48 hours, the results remained consistent. On all treated samples, the liquid spread uniformly, forming a continuous film, which indicates hydrophilic behavior, see Fig. 4-b.

In contrast, on untreated samples, the liquid retracted into droplets. Additional tests using solutions with progressively lower surface tension values were carried out on the untreated substrates, revealing that their surface tension was below 46mN/m in the rolling direction and below 44mN/m in the perpendicular direction. An example of this can be seen in Fig. 4-a. Indeed, the choice of texture for the proposed objective is validated.

[picture]

Fig. 4. Surface tension evaluation of the a) untreated and b) laser-treated sample.

Fig. 4. Surface tension evaluation of the a) untreated and b) laser-treated sample.

Bonding mechanical tensile test results and fracture surface analysis.

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The results of tensile tests on bonded joints are visible in
Fig. 5. Laser treatment can increase the mechanical strength to 157% respect to the degreased bonded joint, while also increasing the displacement value at break. This leads to an increase in absorbed energy of more than three times compared to degreased material and more than double compared to abraded material. The results also demonstrate that the insertion of the metal mesh as a debonding method does not significantly affect the strength of the joint, which becomes around 93% for the textured sample, 95% for the abraded, and 84% for the degreased.
Furthermore, the fracture surfaces analysis in Fig. 6 shows that the de-adhesive failure occurred on the degreased samples, becomes mixed for the abraded sample, and is completely cohesive for the textured sample. This demonstrates greater adhesion between the substrate and adhesive surfaces. Laser texturing treatment uniformly modifies both the chemistry and the morphology of the surface, improving mechanical interlocking.

[picture]

Fig. 5. Tensile test results on bonded joints without and with the metal mesh (MM).

Fig. 5. Tensile test results on bonded joints without and with the metal mesh (MM).

[picture]

Fig. 6. Fracture surfaces analysis of the a) degreased, b) abraded, c) textured bonded joints.

Fig. 6. Fracture surfaces analysis of the a) degreased, b) abraded, c) textured bonded joints.

Cleaning morphological and wettability results.

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The joints subjected to debonding were laser cleaned for the subsequent rebonding. An example of the cleaning process effect is shown in Fig. 7, which depicts the two textured adherends separated after the mechanical test, with the adhesive on both surfaces ( a,c ) and after cleaning ( b,d ). The sample roughness values for the different treatments after laser cleaning are shown in Table 3. In the case of post-laser cleaning treatments, roughness increases due to some localized adhesive residues remaining on the surface, visible as small dots highlighted by the white arrows in Fig. 7−b and d .

[picture]

Fig. 7. Cleaning surfaces on debonded samples: example of textured: the two adherends before (a, c) and after laser cleaning (b, d): white arrows indicate localized adhesive residues.

Fig. 7. Cleaning surfaces on debonded samples: example of textured: the two adherends before (a, c) and after laser cleaning (b, d): white arrows indicate localized adhesive residues.

Table 3. Surface roughness values for the different surface pretreatments after laser cleaning.

Surface pretreatment | Sa [μm] | Sq [μm] | Sz [μm] | Ssk [-] | Sku [-]
Degreased | 4.96±1.39 | 9.66±2.15 | 199.9±86.5 | 6.90±4.00 | 107.8±89.8
Abraded | 8.80±4.45 | 15.5±7.83 | 182.8±68.1 | 1.49±1.97 | 24.9±8.23
Laser textured | 14.2±9.93 | 34.3±22.2 | 552.6±319.2 | 7.64±2.04 | 96.1±48.9

Laser-cleaned surfaces were then re-prepared according to the original three procedures before the rebonding process: degreasing, abrasion, and texturing. It is worth analyzing in detail how the texturing appears on the cleaned material: cleaning creates an uneven surface, visible in Fig. 8-a, with micrometric roughness, in some places there are LIPSS, but they are quite irregular. The subsequent LIPSS texturing adapts to the rough substrate and becomes more regular, see Fig. 8-b. Furthermore, wettability confirms that surface energy is constantly above 72mN/m for post-cleaning texture.

[picture]

Fig. 8. Digital image of laser cleaned (a) and LIPSS texturing after cleaning (b) at high magnification.

Fig. 8. Digital image of laser cleaned (a) and LIPSS texturing after cleaning (b) at high magnification.

Rebonding mechanical tensile test results and fracture surface analysis.

The original version of this paper is available on https://www.scientific.net/SSP.391.65.pdf

The mechanical results of the bonded, debonded, and rebonded joints after cleaning are compared in Fig. 9. Shear tests on the cleaned and rebonded joints show that the surface condition left by laser cleaning increases the mechanical strength of the degreased rebonded joint up to 178% and to 103% compared to the previously textured one. Therefore, surface improvement over the original degreased surfaces can be achieved with the cleaning treatment itself. Only the rebonding of the cleaned and abraded surfaces shows, on average, a slight decrease to 97%, a value that remain inside the standard deviation. Subsequent abrasion does not further improve resistance but restores the original surface. On the other hand, combining these methods would not be advantageous from an industrial point of view, because once the surface has been cleaned with the laser, better results can be achieved (see rebonding results on the degreased surfaces). It is presented for comparison purposes.

In addition, the fracture surfaces in Fig. 10 show that laser cleaning produces a cohesive fracture for both degreased and textured samples, while abrasion again shows a mixed fracture.
When testing the mechanical strength of joints subjected to debonding, shear strength is reduced by 7.20% for textured, 5.42% for abraded, and 4.58% for degreased, respect to the bonded one. The lowest resistance of the degreased allowed separation even manually.

[picture]

Fig. 9. Comparison between tensile test results: bonded, debonded, and rebonded joints.

Fig. 9. Comparison between tensile test results: bonded, debonded, and rebonded joints.

[picture]

Fig. 10. Fracture surfaces analysis of the a) degreased, b) abraded, c) textured rebonded joints.

Fig. 10. Fracture surfaces analysis of the a) degreased, b) abraded, c) textured rebonded joints.

Conclusion

The original version of this paper is available on https://www.scientific.net/SSP.391.65.pdf

The growing demand for sustainable manufacturing and circular economy practices calls for innovative solutions to overcome the limitations of permanent adhesive joints, which hinder component reuse and increase waste. This work explored laser cleaning and texturing as strategies to enable the reuse of single-lap adhesive-bonded joints, addressing the challenge of recyclability in structural bonding. The study showed that controlled debonding using a metal mesh embedded in the adhesive promotes the separation of the joint once it has been disused, preserving the strength of the original joint during operation and preventing post-debonding deformation. Subsequent laser cleaning effectively removes the residual adhesive and not only restores the substrate, but also introduces micrometric roughness, enhancing adhesion performance.

Initial bonding tests confirmed that laser texturing significantly improves the shear strength (up to 1.5 times) and energy absorption (up to 3 times) compared to degreasing and abraded conventional surface preparation ( 1.36 times and 2 times, respectively).

After debonding, laser cleaning alone proved sufficient to recover surface properties and even increase the strength of rebonded joints beyond their original degreased condition. While additional abrasion after cleaning offered no further benefit, the results highlight the potential of laser-based treatments to simplify and optimize the rebonding process.

These findings demonstrate that laser cleaning can be effectively integrated into circular manufacturing strategies, as the process not only enables surface preparation for rebonding but also enhances joint strength. This study primarily aimed to validate the overall methodology, acknowledging that cleaning parameters can be further optimized using different laser sources.

The concept was initially validated through preliminary tests on 6016 aluminum alloy; however, due to its inherently low baseline strength, the research ultimately focused on 6061 aluminum alloy. Given that the method relies on controlled surface-morphological modification, it is reasonable to expect that it can be extended to other structural materials and a wide range of adhesive systems. Laser ablation is already known for its versatility in processing complex materials, with no significant inherent limitations.

Future research should investigate the feasibility of multiple bonding cycles without compromising the substrate's mechanical integrity. Since laser treatment typically affects only a few micrometers of the surface, we are confident that several successive bonding cycles are achievable. This approach also shows promise for industrial-scale adoption, provided that process parameters are optimized for
the different phases, and that economic assessments could confirm the advantages of closed-loop recycling systems aimed at reducing waste and extending component lifetimes.

From an industrial perspective, laser cleaning is both economically and environmentally sustainable over the long term, despite higher initial equipment costs compared to conventional techniques. Its advantages stem from the absence of consumables (e.g., abrasives, chemical agents), low energy demand, and reduced waste-management requirements. When combined with laser texturing, which further enhances material performance while retaining similar sustainability benefits, the approach becomes even more effective. Successful integration requires precise control of laser power and parameters such as pulse duration and frequency. Modern pulsed fiber lasers allow seamless switching between cleaning and texturing modes by adjusting the software settings alone.

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The original version of this paper is available on https://www.scientific.net/SSP.391.65.pdf

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Solid State Phenomena, ISSN: 1662-9779, Vol. 391, pp 79-88

doi: 10.4028/p-sb8yXx

© 2026 The Author(s). Published by Trans Tech Publications Ltd, Switzerland.

Submitted:
2026-01-17

Revised:
2026-02-19

Accepted:
2026-02-20

Online:
2026-04-20

Unconventional Recycling of Aeronautical Ground Carbon Fiber Composite Wastes

The original version of this paper is available on https://www.scientific.net/SSP.391.79.pdf

Pietro Russo 1,a*, Francesco Napolitano 2, b, Fabrizio Quadrini 3,c, Denise Bellisario 4,c

1Institute for Polymers, Composites and Biomaterials, National Research Council, Via Campi Flegrei 34, 80078 Pozzuoli (NA), Italy

2 Institute for Polymers, Composites and Biomaterials, National Research Council, p.le E. Fermi 1, 80055 Portici (NA), Italy

3 Department of Industrial Engineering, University of Rome Tor Vergata, Via del Politecnico 1, 00133 Rome, Italy

a  pietro.russo@cnr.it, b  francesconapolitano@cnr.it, c  fabrizio.quadrini@uniroma2.it, d  denise.bellisario@uniroma2.it

Keywords: Ground CRFP waste, Direct Molding, Thermo-mechanical properties, Recycling

Abstract

Carbon fiber-reinforced polymer (CFRP) composites are extensively used in aerospace applications; however, their end-of-life management remains a critical challenge. This study investigates an unconventional recycling route based on the direct hot compression molding of CFRP waste powders, aiming to valorize industrial composite scraps without the addition of virgin polymers or binding agents. The material investigated corresponds to the finest fraction ( ~300μ m ) obtained from a sieving process applied to industrial CFRP scrap powders derived from trimming residues and partially cured aeronautical prepregs. The use of this fine powder fraction promotes effective particle aggregation and consolidation during molding, preventing powder loss during demolding and enabling the fabrication of relatively thick panels despite the absence of additional bonding agents. Compression molding was carried out at 250∘C and 1.5 bar for 20 min . Two material configurations were analyzed: uncoated compression-molded panels and panels coated with a thin polyester layer. The recycled materials were characterized through morphological, thermomechanical, and mechanical analyses. The results indicate that the polyester-coated panels exhibit improved mechanical performance compared to the uncoated configuration. In comparison with previous studies focused on coarser powder fractions ( ≤1 mm ), the present work highlights the potential of the finest powder fraction for effective consolidation, demonstrating the strong influence of particle size on the processability and properties of compression-molded recycled CFRP. These findings confirm the viability of direct compression molding as a sustainable and scalable recycling strategy for tailoring CFRP waste reuse as a function of powder size.

Introduction

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Industrial manufacturing processes generate substantial quantities of powder-based waste, particularly in sectors such as metallurgy, ceramics, additive manufacturing, and mineral processing. These waste powders often consist of high-value raw materials, including metals, alloys, ceramics, and composite residues, but are frequently discarded due to contamination, irregular particle size distributions, partial degradation, or the absence of economically viable recycling routes [ 1−3 ]. Therefore, large volumes of potentially recoverable material are rerouted to landfilling or incineration, practices that not only impose significant environmental burdens but also result in the irreversible loss of resources and energy [4,5].
In recent years, increasing regulatory pressure and the global push toward decarbonization have driven industries to adopt circular economy principles, emphasizing material recovery, waste minimization, and resource efficiency [2,6]. Within this framework, the valorization of industrial powder waste has emerged as a critical challenge and opportunity. Powder-based residues are particularly attractive candidates for recycling due to their intrinsic form, which is already suitable
for shaping and consolidation processes, potentially bypassing several intermediate manufacturing steps [7].
Conventional recycling approaches for industrial powders typically rely on energy-intensive reprocessing methods, such as remelting, high-temperature sintering, or chemical treatments aimed at restoring material purity [8,9]. While effective in some cases, these strategies often involve high energy consumption, complex separation procedures, and significant operational costs, which can undermine their overall sustainability and limit large-scale implementation [10]. Moreover, repeated thermal or chemical processing may lead to material degradation, further reducing the performance of recycled products [11].
In contrast, direct molding and consolidation of waste powders represent a promising alternative recycling route. This approach enables the conversion of industrial residues into functional components with minimal preprocessing, relying on powder compaction mechanisms such as hot pressing, compression molding, or hybrid thermo-mechanical techniques [12,13]. By avoiding melting or extensive chemical modification, direct molding can significantly reduce energy demand while preserving the intrinsic properties of the original material. Previous studies have demonstrated the feasibility of producing structural and semi-structural components from recycled powders with competitive mechanical, thermal, and functional properties [14].
Despite these advantages, several challenges should be solved in the direct molding of industrial waste powders. Variability in particle morphology, size distribution, chemical composition, and the presence of residual binders or contaminants can strongly influence powder flowability, compaction behaviour, and interparticle bonding [15,16]. Furthermore, the optimization of processing parameters, such as pressure, temperature, dwell time, and binder or coating content, is essential to ensure adequate mechanical integrity, dimensional stability, and reproducibility of the final components.
In this context, the objective of this study is to assess the direct molding approach as a sustainable and scalable non-conventional process for the valorization of industrial powder waste. Aeronautical powder scraps are transformed into panels through direct molding technology, and the resulting materials are characterized in terms of their thermal, mechanical, and morphological properties. This characterization aims to verify the suitability of the recycled panels for potential industrial applications and to identify key parameters governing their performance. By advancing the understanding of direct powder molding from industrial waste streams, this work seeks to contribute to the development of sustainable manufacturing practices and to support the implementation of circular economy strategies within powder-based industries. Consequently, the present work explores an alternative route for the valorization of carbon fiber-reinforced polymer (CFRP) machining residues, focusing on their application as a functional material for panels. The feedstock is mechanically milled waste generated during aerospace-grade composite machining processes, containing short carbon fibers (CF), fragmented thermoset matrix, and minor residual impurities. The recycled powder was consolidated via direct compression molding, deliberately avoiding the addition of external binders, virgin reinforcements, or supplementary curing agents. The consolidation mechanism is driven by a combination of thermo-mechanical effects activated during grinding and compression molding [17-19]. Mechanical milling generates localized shear forces and frictional heating, which can activate residual chemical functionality within the thermoset matrix, promoting matrix softening and interparticle adhesion during molding at elevated temperature [20-22].
Compared to the coarser powder fractions investigated in earlier studies by the authors for CFRP powders with particle sizes ≤1 mm, the finest powder fraction examined in the present work exhibits enhanced aggregation behaviour, leading to improved packing, reduced powder loss during handling and demolding, and more uniform consolidation [23]. This improved aggregation enables the fabrication of relatively thick panels without the use of additional binders or virgin resin systems, highlighting the strong influence of particle size on processability and structural integrity. These results demonstrate that tailoring the recycling strategy as a function of powder size is a key factor in maximizing the effectiveness of direct compression molding as a sustainable route for CFRP waste valorization.

Materials and Methods

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Materials. A carbon fiber-reinforced sandwich panel was manufactured using exclusively recycled, commercially available materials. The carbon fiber-reinforced polymer (CFRP) scraps were derived from mechanically processed CFRP wastes sourced from an aerospace production facility, where it originated from dismantling operations and end-of-life treatment of composite components. As the waste was generated through machining and grinding of certified aeronautical structures, the resulting material exhibits intrinsic heterogeneity, consisting of a mixture of short carbon fibers, cured epoxy matrix fragments, and minor residual impurities. Given the complex and variable nature of the recovered material, no detailed ultimate or proximate chemical analyses were performed. Nevertheless, the composition is representative of conventional aerospace-grade CFRP systems based on epoxy matrices reinforced with carbon fibers. The waste material was collected immediately following the grinding process, a stage at which both particle size distribution and contamination levels are typically broad and poorly controlled. To obtain a powder suitable for further processing, the recovered CFRP scraps were classified via mechanical sieving using a Retsch AS 200 basic sieve shakes (Retsch GmbH, Haan, Germany). A three-step sieving procedure was employed, consisting of an upper sieve with a 4 mm aperture followed by a lower sieve with a 1 mm mesh size and a final one with 300μ m mesh size. This classification yielded a fine fraction with particle sizes below 300 μm. The pronounced heterogeneity of the resulting powder fraction was qualitatively assessed through stereomicroscopic observations carried out with a Leica S9i stereo microscope (Leica Camera AG, Wetzlar, Germany) (Fig. 1a). In addition, the same instrument was used to evaluate the particle size distribution of the sieved powder.

Fabrication of Recycled CFRP Panels. Recycled CFRP panels were manufactured by direct compression molding of industrial CFRP powder without the addition of virgin polymer or binding agents. A square specimen ( 200×200 mm2 ) was produced in a monolithic recycled CFRP panel from 500 g of powder. Molding was performed using a hydraulic hot press (ATS FAAR; maximum load 264 kN , platen size 300×300 mm2 ). The upper and lower platens were set at 250∘C and 220∘C, respectively, to accommodate mold geometry and promote effective consolidation. The recycled CFRP powder was placed in an aluminum mold with an internal cavity of 200×200 mm2 and compressed at 1.5 bar for 20 min . Fluorinated ethylene propylene (FEP) films were used as release layers on both mold interfaces. After molding, samples were cooled to room temperature inside the press for 120 min prior to demolding. The fabrication procedure is schematically illustrated in Fig. 1a. One half of the resulting panel was subsequently coated with a thin polyester layer applied by spray deposition.
Thermomechanical analysis. Dynamic mechanical analysis (DMA) was conducted using a TA Instruments DMA Q800 (New Castle, DE, USA) operating in dual cantilever bending mode. Measurements were performed over a temperature range from 30∘C to 250∘C at a heating rate of 3∘C/min. Rectangular specimens with nominal dimensions of 10×60×3 mm3 were tested, and five samples were analyzed to ensure repeatability.

Thermal analysis. Thermogravimetric analysis (TGA) was performed on both raw powder samples and compacted specimens using a TA Instruments TGA Q500. The tests were carried out under a nitrogen atmosphere with a constant heating rate of 10∘C/min over a temperature range from 30∘C to 900∘C.

[picture]

Fig. 1. Fabrication process scheme (a), recycled CFRP panel coated and uncoated (b).

Fig. 1. Fabrication process scheme (a), recycled CFRP panel coated and uncoated (b).

Microscopic and morphological observations. A first microscopic observation of the CFRP panels surfaces with and without coating were carried out by a digital 3D microscope (Hirox HRX-01, HIROX EUROPE, JYFEL Corporation, Limonest, France). The 3D microscope also enabled a preliminary qualitative analysis of the three-dimensional surface morphology, with extraction of surface parameters for both coated and uncoated CFRP printed panel samples. A more accurate microstructural characterization was performed with scanning electron microscope (SEM) to compare the morphology of unprocessed powders and consolidated materials. Prior to analysis, the samples were sputter-coated with a thin gold-palladium conductive layer to minimize charging effects. Observations were conducted using a field-emission scanning electron microscope (FE-SEM, FEI Quanta 200 F, Zurich, Switzerland), operating under high-vacuum conditions at an accelerating voltage of 30 kV . Finally, to evaluate the effect of the coating on the surface properties, particularly the wettability, measurements were performed using an optical tensiometer (Attension Theta Lite, by Biolin Scientific). A 3μ L drop was deposited on the surface, and the average contact angle was measured by OneAttension software both at 1 s after deposition and after 30 s .

Mechanical tests. The flexural behavior of the manufactured specimens was evaluated through threepoint bending tests conducted in accordance with ASTM D7264. Experiments were performed on an electromechanical universal testing system (MTS Insight 5, MfigTS Systems S.r.l., Torino, Italy) operating under displacement-controlled conditions at a constant crosshead velocity of 1 mm/min. The span length between supports was fixed at 40 mm , and an initial preload of 1 N was applied to ensure proper specimen seating. Throughout the tests, the applied force and corresponding mid-span displacement were continuously monitored. The acquired data were subsequently analyzed to determine flexural strength, flexural modulus, and deformation at failure using the formulations specified in the ASTM standard.

Results and Discussion

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Recycled CFRP panels. The recycled CFRP panel exhibited an average thickness of 12.7±0.32 mm and a bulk density of approximately 1.18 g/cm3. Compared to the recycled panel produced in earlier work from CFRP powders with particle sizes ≤1 mm (density ≈1.3 g/cm3 ) [23], the panel fabricated in this study exhibits a lower apparent density ( ≈1.2 g/cm3 ), likely associated with the presence of differences in powder packing and consolidation behavior due to the sieved powder and the dimensions. In particular, the use of powders with more uniform and finer particle sizes ( <300μ m ) tends to reduce packing efficiency and increase apparent porosity in the consolidated panel, because
a narrow size distribution and lack of large particles limit the ability of smaller particles to fill interstitial voids effectively. This is a consolidated concept in the literature, with narrower and more uniform size distributions generally leading to lower random packing densities compared to broader, multimodal distributions that better fill void spaces [24]. After consolidation, the recycled panel displayed a largely continuous and homogeneous surface, indicating good cohesion across the panel (Fig. 1b). Only limited surface defects were detected, mainly localized cracks attributed to stresses induced during the demolding stage.
The surface morphology of the panels fabricated via compression molding, both coated and uncoated, was further examined using optical microscopy to corroborate the visual assessments. The microscopic images are shown in Fig. 2. Microscopic observations reveal effective consolidation and agglomeration of the recycled constituents during direct compression molding. At higher magnification, clusters composed of randomly oriented short carbon fibers, fragmented epoxy matrix, and dispersed metallic inclusions, originating from machining operations on composite components, are clearly identifiable. These metallic residues, commonly present in aerospace-grade CFRP waste, contribute to the intrinsic heterogeneity of the recycled material. Importantly, the absence of large voids or unbonded regions indicates a high degree of densification, demonstrating efficient utilization of the recycled powder and supporting the potential of this approach for producing structurally consistent materials from CFRP waste.

[picture]

Fig. 2. Microscopic images of the panel surface for the uncoated one (a,b) and coated one (c,d).

Fig. 2. Microscopic images of the panel surface for the uncoated one (a,b) and coated one (c,d).

An in-depth analysis of the surface morphology was conducted using optical profilometry, and the data for the different profiles examined are reported as mean values with standard deviations in Table 1. The surface parameters highlight a pronounced modification of the surface topography induced by the coating. The areal roughness parameters Sa and Sq are significantly reduced in the coated panel, decreasing from 322μ m to 123μ m and from 413μ m to 149μ m, respectively. This reduction indicates an effective smoothing and leveling of the surface after coating application. Similarly, the peak-tovalley parameter Sz exhibits a substantial decrease, from 2973μ m for the uncoated surface to 878 μm for the coated one, reflecting a marked reduction in overall surface height variations. Consistent with this trend, both the maximum peak height (Sp) and maximum pit depth (Sv) are considerably
lower for the coated surface, confirming that the coating effectively fills surface valleys and attenuates prominent asperities. The skewness parameter (Ssk) shifts from a negative value ( -1 ) in the uncoated surface to approximately zero after coating, suggesting a transition from a valley-dominated topography to a more symmetric height distribution. In addition, the kurtosis (Sku) decreases from 4 to 3, indicating a reduction in sharp, extreme surface features and a surface height distribution closer to a normal profile. Overall, these results demonstrate that the coating produces a more uniform and leveled surface morphology, which is expected to influence functional properties such as wettability and interfacial behavior. Furthermore, wettability measurements indicate that the uncoated surface demonstrates a higher contact angle, and thus lower wettability, compared to the coated surface (Table 2). This behavior is consistent with literature reports showing that surface coatings can modify surface roughness and topography, leading to changes in wettability; smoother and more uniform surfaces produced by coatings typically exhibit altered solid-liquid interactions and lower apparent contact angles due to increased effective solid-liquid contact area and surface energy modifications [25]. Moreover, the decrease in the contact angle for the coated sample occurs more rapidly, highlighting a stronger spreading effect. The influence of surface roughness and coatings on wettability is well documented in surface and coatings research, where the interplay between morphology and contact angle is shown to govern wetting behavior.

Table 1. Surface parameters for the uncoated and coated recycled CFRP panel surfaces

Parameter | Uncoated | Coated
Sa | 322 μm | 123 μm
Sq | 413 μm | 149 μm
Sz | 2973 μm | 878 μm
Ssk | -1 | 0
Sku | 4 | 3
Sp | 997 μm | 364 μm
Sv | 1975 μm | 514 μm

Table 2. Contact angle measurements on uncoated and coated recycled CFRP panel surfaces

| CA (°) after 1 s | CA (°) after 30 s
Uncoated | 146.7 ± 6.7 | 138.9 ± 4.9
Coated | 108.4 ± 5.7 | 86.8 ± 5.5

Dynamic-mechanical analysis. Given that, as expected, preliminary analyses have shown that the presence of the coating does not alter the viscoelastic behaviour of the materials considered, the results of the dynamic-mechanical analysis are summarised in Fig. 3 in the form of representative curves. In particular, the dynamic-mechanical analysis of 5 specimens cut from CFRP panels showed the variation of storage modulus ( E ') and loss factor ( Tanδ ) as a function of temperature, and the processing of the curves returned the values of the same parameters collected in Table 3. Clearly, the compression-molded material exhibits a high initial stiffness at room temperature ( E′≈1180−1620 MPa) with a progressive drop in stiffness, up to approximately 120−140MPa, with increasing temperature, particularly marked in the temperature range between 100 and 175∘C (glass transition). The amplitude of this transition phase between the glassy and rubbery states, centered at approximately 167∘C, as evidenced by the Tanδ peak, reflects the heterogeneity of the base waste powders including, among others, short carbon fibers and partially degraded or post-cured thermoset phases.

[picture]

Fig. 3. a) Storage modulus ( E′ ), b) Loss factor ( Tanδ ).

Fig. 3. a) Storage modulus ( E ′ ), b) Loss factor ( Tan δ ).

Table 3. Results of the dynamic mechanical analysis

Sample | E' at 30 °C [MPa] | E' at 175 °C [MPa] | T at Tan δ peak [°C]
1 | 1182.66 | 141.72 | 168.64
2 | 1345.52 | 142.32 | 169.84
3 | 1467.15 | 140.14 | 167.26
4 | 1433.01 | 121.18 | 166.11
5 | 1622.49 | 124.98 | 164.97

Thermogravimetric Analysis (TGA). Fig. 4 shows representative TGA curves for processed and unprocessed CFRP powders. Again, the curve processing provided the values of some thermal degradation parameters specific to the CFRP waste before and after the direct molding process (Table 4). Both powders show a multistep decomposition pattern, likely due to concurrent phenomena such as volatile evaporation, resin breakdown, and the formation of stable char and carbon-fiber residue. The processed powder shows higher decomposition temperatures, indicating greater thermal resistance and a more consolidated, fully cured matrix. The lower final residue for the processed sample suggests different char formation behavior due to thermomechanical history.

[picture]

Fig. 4. TGA graph of processed and unprocessed composite powders.

Fig. 4. TGA graph of processed and unprocessed composite powders.

Table 4 Thermogravimetric analysis parameters.

Sample | T2% weight loss [°C] | Tpeak [°C] | Residue at 900 °C [%]
Preprocess | 188.44 | 325.28 | 61.87
Postprocess | 198.46 | 359.23 | 55.59

Mechanical results. The results of the three-point bending tests are shown in Fig. 5 as loaddisplacement (Fig. 5a) and stress-strain (Fig. 5b) curves for four representative samples of each composite panel type. The coated samples exhibited a maximum load of 146.8±14.7 N, a maximum flexural stress of 6.1±0.4MPa and an average flexural modulus of 401.0±12.0MPa, significantly higher than the 117.4±16.4 N,4.5±0.5MPa and 332.0±13.8MPa measured for the uncoated samples, corresponding to an increase of about 21%. The flexural modulus of the recycled CFRP panel without coating was approximately 330.0 MPa . Although significantly lower than that of continuous-fiber CFRP laminates, this value is consistent with the highly discontinuous and porous nature of the recycled material. The short, randomly oriented fibers and limited consolidation efficiency result in a matrix-dominated bending response and reduced load-transfer efficiency. The use of a three-point bending configuration further emphasizes the influence of surface defects and local heterogeneities, which are intrinsic to compression-molded recycled CFRP panels produced from powdered feedstock. The panels with and without coating behaved in a brittle manner, showing no significant plastic deformation or damage precursors prior to fracture, typical of a discontinuous material like this.

[picture]

Fig. 5. Load vs displacement (a) and stress vs. strain (b) curves for both panels uncoated and coated.

Fig. 5. Load vs displacement (a) and stress vs. strain (b) curves for both panels uncoated and coated.

These findings suggest that, beyond the observed mechanical and surface modifications, the presence of the coating contributes to the development of a more continuous and resistant surface, thereby enhancing the overall performance of the panel. Importantly, the simple application of a thin coating on the top and bottom surfaces of relatively thick recycled panels may significantly expand the potential application range of this class of materials, offering an effective and low-complexity strategy to improve surface-related properties without altering the bulk manufacturing process.

Summary

The original version of this paper is available on https://www.scientific.net/SSP.391.79.pdf

This work demonstrates that the application of a thin surface coating effectively enhances the surface quality and functional performance of recycled CFRP panels produced via a non-conventional compression molding process based on CFRP powder waste. Surface analyses revealed a significant reduction in roughness and height-related parameters, indicating effective surface leveling and improved uniformity in the coated panels. Mechanical testing under three-point bending confirmed that the recycled panels exhibit a matrix-dominated flexural response, consistent with their highly discontinuous and porous microstructure resulting from the nature of the feedstock material. The application of the coating enhances the bulk mechanical properties, as evidenced by slight increase in the flexural modulus from 330 MPa for the uncoated panel to 400 MPa for the coated one. In addition, the coating markedly improves surface integrity and wettability, contributing to more sturdy and functional panel surfaces. Overall, the results highlight that the simple application of thin coatings on the top and bottom surfaces of thick recycled CFRP panels represents an effective, scalable
strategy to extend the applicability of materials produced through non-conventional compression molding, without increasing process complexity or compromising the sustainability advantages of recycled composites.

References

The original version of this paper is available on https://www.scientific.net/SSP.391.79.pdf

B. Reck, T.E. Graedel T.E., Challenges of metal recycling, Science 337(6095), (2012) 690-695.

J. Kirchherr, D. Reike, M. Hekkert, Conceptualizing the circular economy: An analysis of 114 definitions, Resources, Conservation and Recycling 127 (2017) 221-232.

J. R. Duflou, J. W. Sutherland, D. Dornfeld, C. Herrmann, J. Jeswiet, S. Kara, M. Hauschild, K. Kellen, Towards energy and resource efficient manufacturing: A processes and systems approach, CIRP Annals 61(2) (2012) 587-609.

R. Geyer, J. R. Jambeck, K. L., Law Production, use, and fate of all plastics ever made, Science Advances 3(7) (2017) e1700782.

European Commission. Critical Raw Materials and the Circular Economy. COM(2017) 490 final.

M. Geissdoerfer, P. Savaget, N. M. P. Bocken, E.J. Hultink, The Circular Economy - A new sustainability paradigm?, Journal of Cleaner Production 143 (2017) 757-768.

R. M. German, Powder metallurgy and particulate materials processing. Ed. Metal Powder Industries Federation, University of California, 2005.

S. J. Pickering, Recycling technologies for thermoset composite materials-current status, Composites Part A 37(8) (2006) 1206-1215.

G. Oliveux, L. O. Dandy, G. A. Leeke, Current status of recycling of fibre reinforced polymers, Progress in Materials Science 72 (2015) 61-99.

J. M. Allwood, M. F. Adhby, T. G. Gutowski, E. Worrell, Material efficiency: A white paper, Resources, Conservation and Recycling 55(3) (2011) 362-381.

S. Job, Recycling glass fibre reinforced composites - history and progress, Composites Part A 105 (2018) 299-321.

M. Golvaskar, S.A. Ojo, M. Kannan, Recent Advancements in Material Waste Recycling: Conventional, Direct Conversion, and Additive Manufacturing Techniques, Recycling 9(3) (2024), 43.

S.S. Alrwashdeh, A. Al-Kaseasbeh Qusay, M. Al-Qaralleh, Ala'a M. Al-Falahat, Recycling of materials industrial waste into a new composite alloy using sintering technique, Results in Engineering 26 (2025) 104975.

P.S. Stelzer, U. Cakmak, L. Eisner, L. K. Doppelbauer, I. Kállai, G. Schweizer, H. K. Prammer, Z. Major, Experimental feasibility and environmental impacts of compression molded discontinuous carbon fiber composites with opportunities for circular economy, Composites Part B: Eng. 234 (2022), 109638.

A. Santomaso, P. Lazzaro, P. Canu, Powder flowability and density ratios: The impact of granules packing, Chem. Eng. Sci. 58(13) (2003) 2857-2874.

J. K. Prescott, R. A. Barnum, On powder flowability, Pharm. Techn. 24(10) (2000) 60-84.

F. Quadrini, D. Bellisario, L. Iorio, A. Proietti, M. Regi, L. Santo, Recycling of Powders from Cutting of Medium Density Fiberboard. In Proceedings of the 25th International Conference on Material Forming, ESAFORM 2022, Braga, Portugal, 27-29 April 2022; Key Engineering Materials, Volume 926; KEM: pp. 1719-1724, by Trans Tech Publications Ltd, Stafa-Zurich, Switzerland.

D. Bellisario, L. Iorio, A. Proietti, F. Quadrini, L. Santo, Recycling of thermoset fiberglass by direct molding of ground powders, Mater. Res. Proc. 41 (2024) 2410-2418.

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D. Bellisario, L. Iorio, A. Proietti, F. Quadrini, L. Santo, Recycling of thermoset fiberglass by direct molding of ground powders, Mater. Res. Proc. 41 (2024) 2410-2418.

D. Bellisario, F. Quadrini, F. Napolitano, P. Russo, Enhancing Through-Thickness Electrical Conductivity in Recycled Carbon Fiber-Reinforced Polymer Composites Using Machining Waste, J. Compos. Sci. 9 (2025) 451.

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Solid State Phenomena, ISSN: 1662-9779, Vol. 391, pp 89-94

doi: 10.4028/p-1gR22Z

© 2026 The Author(s). Published by Trans Tech Publications Ltd, Switzerland.

Submitted:
2026-01-18

Revised:
2026-02-17

Accepted:
2026-02-18

Online:
2026-04-20

Electrodeposited Metallic Coatings as a Manufacturing Route to Enhance the Durability of NdFeB Magnets

The original version of this paper is available on https://www.scientific.net/SSP.391.89.pdf

Scherillo Fabio 1,a*, Perna Alessia Serena 1, b and Viscusi Antonio 2,c

¹Department of Chemical, Materials and Production Engineering, University of Naples Federico II, Piazzale V. Tecchio 80, 80125 Napoli, Italy

2 Department of Engineering and Sciences, Faculty of Technology and Innovation Sciences, Universitas Mercatorum, Piazza Mattei 10, 00186 Rome, Italy

ª*fabio.scherillo@unina.it, b alessiaserena.perna@unina.it, c antonio.viscusi@unimercatorum.it

*corresponding author

Keywords: electrodeposition, rare-earth magnets, metallic coatings, corrosion protection.

Abstract

NdFeB rare earth magnets are widely used in high performance applications but suffer from poor corrosion resistance due to their multiphase microstructure and the preferential dissolution of rare earth rich intergranular phases. In this study, the corrosion behaviour of NdFeB magnets was investigated and the effectiveness of galvanic nickel coatings as a protective solution was evaluated. Microstructural and compositional analyses were combined with potentiodynamic polarisation tests, comparing bare and coated magnets with carbon steel. Coating adhesion was assessed by pull off testing according to ASTM D4541. The results show that NdFeB magnets are significantly more susceptible to corrosion than steel, while nickel exhibits a passive behaviour in the investigated environment. Galvanic nickel coatings provide effective protection and display adhesion comparable to that measured on steel substrates, demonstrating their suitability for corrosion protection of NdFeB magnets.

Introduction

The original version of this paper is available on https://www.scientific.net/SSP.391.89.pdf

Rare earth based magnets are well known for their high performance and have become essential components in numerous industrial sectors since the 1980s [1]. Within this group, NdFeB magnets are especially valued because they combine high coercivity, strong remanence, elevated saturation magnetization, and the ability to deliver significant magnetic output even when produced in compact geometries [2]. These characteristics have encouraged their widespread incorporation into technologies such as electric motors and renewable energy devices, including wind turbine generators, whose importance continues to increase in response to global efforts toward sustainability and reduced environmental impact [3].
Despite these advantages, their practical use is limited by a pronounced sensitivity to environmental degradation. Exposure to moisture or aggressive atmospheres frequently compromises their stability and may necessitate repeated replacement in operating systems [4]. NdFeB magnets exhibit poor corrosion resistance largely due to their multiphase microstructure, characterized ferrite phase and rare earths enriched intergranular matter. The corrosive attack occurs inside the intergranular matter, coupled with cathodic reactions occurring on the Fe-rich domains, results in the intergranular corrosion commonly described in the literature. Because these magnets are used in both indoor and outdoor environments, it is essential that their magnetic properties remain stable over prolonged service periods. Various approaches have been explored to enhance their corrosion resistance, generally falling into two main categories: modification of the alloy composition and the application of protective coatings. Adjusting the alloy chemistry may influence the microstructure in ways that reduce magnetic performance, making this strategy less attractive for many applications [5]. Consequently, protective coatings such as organic layers, polymeric films, and metallic deposits (typically nickel or zinc) are frequently considered. Each coating system presents specific benefits and limitations [6,7]. Among available solutions, galvanic plating represents an effective option for safeguarding rare earth magnets. This method allows the formation of uniform metallic layers with
thicknesses of only a few tens of micrometers, which helps preserve the magnetic behavior of the substrate while improving its resistance to corrosive attack [8]. Galvanic coatings have been widely researched and implemented across different metals, demonstrating good adhesion on various substrates, favorable wear resistance, possibility to accurately control the coating thickness, and suitability for components with complex shapes without additional processing costs [9]. One of the critical aspects is that these coatings often exhibit poor adhesion, particularly when applied to materials characterized by microstructural heterogeneity.
However, published studies do not yet provide a comprehensive evaluation of their use on rare earth based magnets. It is therefore necessary to assess whether such coatings can achieve adequate adhesion on NdFeB substrates and whether thin deposits are sufficient to enhance their corrosion performance. For these reasons, this study investigates the corrosion behavior of NdFeB rare earth magnets and evaluates their protection through galvanic nickel based coatings, which are among the most widely used materials for this type of surface treatment.

Materials and Methods

The original version of this paper is available on https://www.scientific.net/SSP.391.89.pdf

The NdFeB magnets used in this study were supplied by Magfine Srl (Sendai, Japan) in the form of discs with a diameter of 25 mm and a thickness of 10 mm . Prior to characterization, the magnets were demagnetized and heated to their Curie temperature of 380∘C for 4 hours, followed by furnace cooling for 24 hours [10]. The specimens were then sectioned, mounted, and subjected to water free metallographic preparation.
The prepared samples were examined using a Hitachi TM 3000 scanning electron microscopy, and their elemental composition was assessed through a Oxford Instrument Swift ED 3000 detector for energy dispersive X ray spectroscopy (EDX).

[picture]

Fig. 1. Schematic representation of the experimental setup used for the nickel galvanization of NdFeB magnets.

Fig. 1. Schematic representation of the experimental setup used for the nickel galvanization of NdFeB magnets.

Following the microstructural and compositional analysis, galvanic nickel coatings were produced A schematic representation of the experimental apparatus employed, is shown in Fig.1. Before coating deposition, the magnets were mechanically polished with abrasive paper to enhance coating adhesion. To evaluate adhesion on a reference substrate with established galvanic plating performance, the same coatings were also deposited on AISI 1040 carbon steel samples, for which good adhesion has been consistently reported in the literature.
Galvanic deposition was carried out using nickel electrodes in combination with the corresponding electrolyte solutions. The anodes employed consisted of rolled nickel with a controlled chemical
composition. For the deposition, a commercial nickel electroplating solution designed for bright and smooth coatings was used, applying a current density of 1.8 A/dm2 for 15 min .
To ensure consistency in the coated area across all samples, a masking system was applied to both NdFeB and steel cylinders, resulting in a defined deposition area with a diameter of 10 mm . The applied current density was selected according to values recommended in the literature, while a coating thickness of 5μ m was adopted to maintain a balance between corrosion protection and preservation of geometric characteristics.
After deposition, coating adhesion to the substrate was evaluated through pull off testing using a portable adhesion tester (PosiTest AT M), following the ASTM D4541 standard.
To assess the corrosion resistance provided by the coatings, potentiodynamic polarization tests were performed in a aqueous solution containing 1 g/l of NaCl . In this analysis, coated NdFeB magnets were compared with coated steel cylinders, which served as the benchmark for the electrochemical measurements. All measurements were carried out on three different samples for both magnets and steel.

Results and Discussion

The original version of this paper is available on https://www.scientific.net/SSP.391.89.pdf

In order to evaluate the electrochemical behavior and, consequently, to assess the feasibility of an electrochemical material removal process, microstructural analysis was performed prior to deposition, as shown in Fig. 2.
A detailed understanding of the substrate phase composition and distribution is required in order to properly interpret the effects induced by the galvanic plating process.
The magnet exhibits a biphasic microstructure, characterized by equiaxed grains surrounded by an intergranular phase, which appears brighter under SEM contrast. The overall composition of the magnet, as determined by EDX analysis, is reported in Fig. 2b. The compositions of the two aforementioned phases are also provided. The grey phase corresponds primarily to a ferritic phase, while the intergranular phase is enriched in cesium and praseodymium relative to overall composition. Among the rare earth elements, neodymium is distributed approximately uniformly between the two phases.

[picture]

Fig. 2. Microstructure and elemental composition of the NdFeB magnet prior to coating. a) SEM image showing the biphasic structure with equiaxed grains and brighter intergranular regions. b) EDX analysis reporting the overall composition and the composition of phases.

Fig. 2. Microstructure and elemental composition of the NdFeB magnet prior to coating. a) SEM image showing the biphasic structure with equiaxed grains and brighter intergranular regions. b) EDX analysis reporting the overall composition and the composition of phases.

The presence of the two phases plays a fundamental role in the corrosion mechanism governing magnet degradation, which the galvanic coatings are intended to mitigate. In order to gain deeper insight into the corrosion process, the magnet was exposed to an industrial corrosive environment at Fater SpA, characterized by the presence of chloride ions. Fig. 3 shows a section of an industrial magnet employed in a corrosive environment.

[picture]

Fig. 3. Cross-section of NdFeB magnet after exposure to a chlorine-containing corrosive environment.

Fig. 3. Cross-section of NdFeB magnet after exposure to a chlorine-containing corrosive environment.

The micrograph clearly indicates that corrosion proceeds through an intergranular mechanism. The grey ferritic phase is distinctly visible, whereas the intergranular phase, which is enriched in rare earth elements, is no longer detectable, as it selectively dissolves under the action of the corrosive environment.
Accordingly, galvanic coatings were applied, as described in the Materials and Methods section, on both magnet and steel substrates. Potentiodynamic polarization tests were then carried out on coated and uncoated specimens, as shown in Fig. 4.
Potentiodynamic polarization tests indicate that, compared with a low alloy reference steel, the magnet is overall less noble. It exhibits a lower corrosion potential and, at a given potential, higher dissolution currents, indicating a significantly greater susceptibility to corrosion than carbon steels.

[picture]

Fig. 4. Potentiodynamic polarization curves of uncoated and galvanic nickel coated NdFeB magnets, compared with carbon steel.

Fig. 4. Potentiodynamic polarization curves of uncoated and galvanic nickel coated NdFeB magnets, compared with carbon steel.

In contrast, nickel shows a passive behavior in the investigated corrosive environment, characterized, relative to both steel and the magnet, by higher corrosion potentials and anodic currents that are several orders of magnitude lower. Nickel therefore represents an excellent material for protective applications. It should be noted that this protection mechanism is based on a metal that is more noble to respect the substrate and act as barrier between the corrosive environment and the magnet. The key factor governing the feasibility of applying this coating to magnets lies in adhesion performance, since galvanic coatings must be well adhered and free of delaminated areas through which the corrosive solution may penetrate.
For this reason, adhesion tests were performed by means of pull off testing in accordance with ASTM D4541, and the corresponding results are reported in Fig. 5

[picture]

Fig. 5. Pull off test results for galvanic nickel coatings on NdFeB (left) magnets and carbon steel (right).

Fig. 5. Pull off test results for galvanic nickel coatings on NdFeB (left) magnets and carbon steel (right).

The pull off test shown in Fig. 5 demonstrates that the adhesion of the nickel coating is fully comparable to that measured on carbon steel substrates.

Conclusion

The original version of this paper is available on https://www.scientific.net/SSP.391.89.pdf

This work investigated the corrosion behavior of NdFeB rare earth magnets and evaluated the suitability of galvanic nickel coatings as a protective solution. The main results of the work can be summarized as follows:

NdFeB magnets exhibit a biphasic microstructure in which rare earth rich intergranular phases play a dominant role in corrosion degradation through an intergranular mechanism.

Exposure to aggressive industrial environments leads to selective dissolution of the intergranular phase, confirming the intrinsic susceptibility of these magnets to corrosion.

Potentiodynamic polarisation tests show that NdFeB magnets are less noble than carbon steel, with lower corrosion potentials and higher anodic dissolution currents. Nickel displays a passive electrochemical behaviour in the investigated environment, with significantly higher corrosion potentials and anodic currents several orders of magnitude lower than those of the magnet.

Galvanic nickel coatings provide protection to NdFeB magnets, effectively improving their corrosion resistance.

Pull off adhesion tests demonstrate that the adhesion of galvanic nickel coatings on NdFeB magnets is comparable to that obtained on carbon steel substrates.

The combination of effective corrosion protection and adequate adhesion indicates that galvanic nickel coatings are a technically viable solution for extending the service life of NdFeB magnets in corrosive environments.

References

The original version of this paper is available on https://www.scientific.net/SSP.391.89.pdf

B. D. Cullity, C. D. Graham, Introduction to Magnetic Materials, Wiley, 2009.

J.D. Livingston, Magnetic domains in sintered NdFeB magnets, J. Appl. Phys. 57 (1985) 41374139. .

Y. Matsuura, Recent development of Nd-Fe-B sintered magnets and their applications, J. Magn. Magn. Mater. 303 (2006) 344-347. .

J.-M. Hu, X.-L. Liu, J.-Q. Zhang, C.-N. Cao, Corrosion protection of Nd-Fe-B magnets by silanization, Prog. Org. Coat. 55 (2006) 388-392. .

S. Constantinides, Permanent magnet coatings and testing procedures, in: Modern Permanent Magnets, Elsevier, 2022: pp. 371-402. .

D.J. Blackwood, B. Balakrisnan, Y.Z. Huang, C.K. Tan, Influence of the chemical composition of the plating solution on the ability of nickel coatings to protect Nd 2 Fe 14 B magnets against corrosion, J. Magn. Magn. Mater. 223 (2001) 103-111. .

D.C. Nababan, R. Mukhlis, Y. Durandet, M.I. Pownceby, L. Prentice, M.A. Rhamdhani, Kinetics of high temperature oxidation of end-of-life coated NdFeB rare earth permanent magnets, Corros. Sci. 189 (2021). .

G.O. Mallory, J.B. Hajdu, Electroless plating: fundamentals and applications, Cambridge University Press, 1990.

Y. Liu, X. Zheng, S. Hu, D. Cai, G. Li, Effects of electroless nickel plating process on the microstructure and properties of complex shape ceramic capacitors, Journal of Materials Research and Technology 35 (2025) 6739-6748.

G. Dai, W. Yin, X. Tang, J. Ju, L. Wu, Y. Cui, R. Chen, A. Yan, High-temperature magnetic performance and corrosion resistance of hot-deformed and sintered Nd-Fe-B magnets with similar room-temperature magnetic properties, J. Magn. Magn. Mater. 587 (2023) 171321. .

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The original version of this eBook is available on https://www.scientific.net/book/non-conventional-processes/978-3-0364-1985-5

Solid State Phenomena, ISSN: 1662-9779, Vol. 391, pp 95-106

doi: 10.4028/p-Az4zHs

© 2026 The Author(s). Published by Trans Tech Publications Ltd, Switzerland.

Submitted:
2026-01-19

Revised:
2026-02-20

Accepted:
2026-02-20

Online:
2026-04-20

Study of the Effect of Foam Substrate Material on Hydrophobic Coating Efficiency

The original version of this paper is available on https://www.scientific.net/SSP.391.95.pdf

Baxevani Aikaterini 1,a, Lamprou Eleni 1, b*, Stergioudi Fani 1,2,c, Michailidis Nikolaos 1,2, d

¹Physical Metallurgy Laboratory, School of Mechanical Engineering, Aristotle University of Thessaloniki, Greece

2 Centre for Research & Development of Advanced Materials (CERDAM), Centre for Interdisciplinary Research and Innovation, Balkan Centre, Building B, 10th km Thessaloniki-Thermi road, Thessaloniki 57001, Greece

a ampaxeva@meng.auth.gr, b ealamprou@meng.auth.gr, c fstergio@meng.auth.gr, d  nmichail@meng.auth.gr

Keywords: Metal Foams, Coating, Hydrophobicity, Stability, Oil-separation Efficiency, Corrosion

Abstract

This study investigates the fabrication and performance of hydrophobic copper and nickel foams produced through a two-step immersion coating process using silver nitrate and stearic acid, aiming at oil-water separation and corrosion protection in marine environments. The coating process generated hierarchical surface roughness and reduced surface energy, leading to enhanced hydrophobic behavior. Surface morphology and wettability were systematically evaluated, confirming the formation of stable hydrophobic coatings on both substrates. The modified foams demonstrated high oil-water separation efficiency, good reusability, and strong thermal and chemical stability. Corrosion resistance in saline solutions was also significantly improved after coating. Although the same treatment was applied to both substrates, differences in wettability and long-term stability were attributed to their intrinsic microstructural characteristics. Overall, the results highlight the effectiveness of the developed coating strategy for multifunctional surface protection applications.

Introduction

The original version of this paper is available on https://www.scientific.net/SSP.391.95.pdf

In recent years, oil-contaminated wastewater and accidental oil spills at sea have increased the need for effective and sustainable oil-water separation methods. Industrial activities such as petroleum refining, offshore drilling, metal processing, mining, textile manufacturing and food production produce large amounts of oily wastewater, harming aquatic ecosystems, public health and economic stability [1-4]. Large-scale incidents, including major offshore oil spills, further emphasize the urgent need for advanced separation materials capable of operating effectively under harsh environmental conditions [5,6].

Conventional oil-water separation methods, including centrifugation, gravity settling, chemical dispersants, bioremediation and flotation-based techniques, are widely applied but remain limited by low selectivity, high operational costs, complex system requirements and, in some cases, the risk of secondary pollution [7-11]. As a result, increasing attention has been directed toward surfaceengineered materials that enable selective oil capture while repelling water, offering simpler operation and improved separation efficiency.

Nature-inspired superhydrophobic surfaces are a promising solution, inspired by plants such as lotus leaves that exhibit water repellency and self-cleaning behavior [12, 13]. These properties result from a combination of hierarchical micro-/nanostructures and low surface energy. Applying these principles to materials has led to the development of superhydrophobic coatings for removing pollutants [14-16].

Among the various substrates explored, porous materials have shown particular promise due to their high surface area and fluid permeability. While textiles, sponges and polymeric foams have been investigated, their limited durability and structural stability restrict their long-term applicability. In contrast, metallic foams offer superior mechanical strength, thermal stability and reusability, making
them attractive candidates for oil-water separation in demanding environments [17, 18]. However, achieving durable, cost-effective and scalable superhydrophobic coatings on metal foams remains a challenge, as many reported fabrication techniques rely on complex, multi-step or expensive processes.

Hydrophobicity enables mechanical oil-water separation by preventing water from wetting the surface while promoting preferential oil transport through the porous structure. Surfaces with high water contact angles resist water penetration, allowing oil, which has lower surface tension, to selectively wet and pass through the material. The presence of micro-/nano-scale surface roughness further enhances this selectivity by strengthening water repellency and maintaining stable separation performance during repeated operation. These principles are well established in studies on hydrophobic foams and membranes for oil-water separation, where surface morphology and wettability are directly linked to separation efficiency and durability [19-21].

Various strategies have been explored to create superhydrophobic surfaces on metallic substrates, but many involve elaborate procedures or specialized treatments. For example, Xin et al. [22] combined laser ablation with chemical functionalization using FAS and graphene oxide. Zhang et al. [23] achieved superhydrophobicity by growing HKUST-1 in situ on copper foam followed by surface modification with 1-Hexadecanethiol, while Yong et al. [24] fabricated hydrophobic nickel foams via repeated slurry-based coating and oven drying, involving multiple processing cycles and the use of toxic solvents such as DMSO. Although these approaches demonstrate effective surface modification, their multi-step nature, energy demand and environmental burden limit their practicality and scalability.

In this research work, a robust and efficient hydrophobic coating was developed on copper and nickel foam substrates using a simple and rapid immersion-based method. The proposed approach minimizes processing steps and fabrication time, avoids complex equipment or energy-intensive procedures, and requires only small amounts of chemicals. By applying the same coating strategy to both substrates, the influence of the metallic foam material on wettability, coating adhesion, corrosion resistance and oil-water separation performance is systematically examined. This enables a direct comparison of substrate effects while demonstrating that a fast, cost-effective and scalable method can produce durable hydrophobic metallic foams suitable for environmental and marine applications.

Since the coating process is reproducible, it can also be easily scaled up for larger samples or industrial use. The fabrication process is based on a simple immersion method, with the samples allowed to dry gently in air under ambient conditions, without requiring any specialized equipment. Its straightforward procedure and minimal use of chemicals make it easy to implement for multiple samples simultaneously or for larger foam surfaces. By proportionally increasing solution volumes and container sizes, the coating can be applied uniformly to larger areas without compromising the hierarchical surface structure that is critical for hydrophobicity. These features ensure that the method can be scaled up efficiently, maintaining consistent coating quality, high separation performance, and durability. As a result, the approach is suitable for industrial applications such as large-scale oil-water separation, wastewater treatment, and filtration systems, demonstrating both practical applicability and scalability while remaining cost-effective and environmentally friendly.

Experimental Procedure

The original version of this paper is available on https://www.scientific.net/SSP.391.95.pdf

Materials. Silver nitrate ( AgNO3, purity > 99.8%) was supplied by Panreac (Barcelona, Spain), while stearic acid (C18H36O2, purity >97%) and ethanol (purity ≥99% ) were obtained from Merck (Darmstadt, Germany). Copper and nickel foams were provided as sheets with dimensions of 100 mm×100 mm×2 mm, an average pore size of 1.6 mm and a porosity of 94%. Sodium hydroxide ( NaOH,98% ) and hydrochloric acid solution ( HCl,37% ), used for chemical stability tests, were obtained from Honeywell (Seelze, Germany) and Panreac (Barcelona, Spain), respectively, for pH regulation.

Fabrication of Hydrophobic Metallic Foams. To ensure direct comparability between the two substrates, the same simple and rapid immersion-based coating procedure was applied to both copper
and nickel foams. Circular foam discs ( 15 mm in diameter and 2 mm in thickness) were cut from the sheets and ultrasonically cleaned in alternating acetone and anhydrous ethanol baths for 5 min per cycle, repeated three times, to remove surface contaminants. The coating process consisted of two immersion steps, each carried out in 50 mL of ethanolic solution without stirring. In the first step, the pretreated foams were immersed in a 0.2 M ethanolic solution of silver nitrate at 50∘C for 20 min to induce micro-/nano-scale surface roughness through silver deposition. The samples were rinsed with deionized water. In the second step, the samples were transferred into a 0.15 M ethanolic solution of stearic acid at 50∘C for 50 min , aiming to reduce surface energy (Fig. 1). After coating, the samples were dried in a desiccator containing silica gel under ambient conditions.

[picture]

Fig. 1. Fabrication process of hydrophobic copper and nickel foams.

Fig. 1. Fabrication process of hydrophobic copper and nickel foams.

The proposed method requires minimal processing time and very small quantities of chemicals; for instance, only 0.17 g of silver nitrate and 0.21 g of stearic acid were sufficient to coat four foam samples. These low material requirements highlight the cost-effectiveness, environmental friendliness and scalability of the approach. The successful application of the same coating strategy to both copper and nickel foams confirms its suitability for different metallic foam substrates.

To ensure reproducibility, the experiments for optimizing AgNO3 and stearic acid treatment were repeated five times. Water contact angles are reported as the average of five measurements. Similarly, nitrate ion absorbance was measured at different intervals, with each value representing the average of five independent measurements. These repeated trials confirm that the coating process is reliable and produces consistent results for both surface wettability and chemical deposition.

Characterization and Testing. The fabricated hydrophobic metallic foams were systematically characterized to evaluate the effectiveness, uniformity, and reproducibility of the two-step immersion coating process from a manufacturing perspective.

Surface morphology and coating development were examined using scanning electron microscopy (Phenom ProX desktop SEM, Thermo Fisher Scientific, Eindhoven, The Netherlands), enabling assessment of microstructural features and coating homogeneity induced by the fabrication process. UV-Vis spectroscopy (UV-1800 spectrophotometer, Shimadzu, Tokyo, Japan) was employed to examine the immersion time in silver nitrate (AgNO3) solution for both copper and nickel foam substrates to determine the optimal conditions for silver deposition and to compare the behavior of the two materials. Nitrate ion concentrations were determined using the Beer-Lambert law (Eq. 1):

A=αλ·b·c(1)

where A is the absorbance, αλ the molar absorptivity, b the optical path length, and c the concentration. Samples were collected at regular intervals over 30 min , and concentrations were averaged from five measurements per point.

Thermal stability and mass variation of the coated foams were evaluated by thermogravimetric analysis (TGA) using a Discovery SDT 650 thermal analyzer (TA Instruments) over the temperature range of 25−200∘C, providing insight into coating robustness and durability. Chemical stability was assessed through controlled immersion tests in NaOH,HCl, and saline (NaCl) solutions, followed by
water contact angle (WCA) measurements to quantify the retention of superhydrophobic properties after chemical exposure. The objective was to assess how the nature of the substrate influences the durability of the hydrophobic coatings when exposed to thermal stress, chemically aggressive media, and saline environments. All experiments were performed five times to ensure reproducibility, and the reported results correspond to the average values.

Corrosion performance was investigated in a corrosive electrolyte of 3.5wt%NaCl via electrochemical polarization tests using a three-electrode configuration, with a saturated calomel electrode (SCE) as reference and a platinum rod as counter electrode. Measurements were conducted using a VoltaLab PGZ 402 potentiostat (Radiometer Analytical, Villeurbanne Cedex, France). Prior to polarization, the open circuit potential (OCP) was recorded until stabilization. Polarization scans were performed at 2mVs−1 within a potential window of -900 mV to +900 mV versus SCE, relative to OCP. Due to the intrinsic porosity and geometrical complexity of metallic foams, precise determination of the effective electrochemical surface area is challenging, particularly in the presence of non-uniform surface coatings. Therefore, a nominal geometric area of 220 mm2 was adopted for all calculations to ensure consistency and comparability. All specimens were fabricated with identical dimensions and immersed to half their height during testing, ensuring standardized exposure conditions and enhancing measurement reproducibility.

Oil-water separation experiments were conducted in laboratory-scale separation units operating at 25∘C with an oil-to-water volume ratio of 1:3. The separation system consisted of a twocompartment rectangular tank containing 60 mL of water and 20 mL of oil, allowing gravity-driven permeation through the porous foam structure. Separation performance was evaluated using optimally processed superhydrophobic copper foams. All experiments were performed at least five times to ensure reproducibility, and the reported results correspond to the average values.

To assess the influence of fluid properties on separation efficiency, oils with distinct viscosity ranges were selected, including low-viscosity sunflower oil, medium-viscosity engine oil (10W-40), and high-viscosity engine oil ( 80 W−90 ). The dynamic viscosities of the oils were determined using a TA Discovery Hybrid Rheometer HR30 (TA Instruments) prior to testing, ensuring accurate correlation between rheological properties and separation performance.

Results and Discussion

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Effect of Immersion Time. The influence of immersion time in silver nitrate ( AgNO3 ) solution was examined for both copper and nickel foam substrates to determine the optimal conditions for silver deposition and to compare the behavior of the two materials. For copper foams, the absorption spectrum (Fig. 2a) shows a peak at 274 nm corresponds to Ag2, while a peak at 308 nm corresponds to nitrate ions (NO3−). The highest peak arises from solvent effects (ethanol) and does not provide reliable information about changes in silver concentration. The peak at 325 nm corresponds to charged silver aggregates ( Ag2+), formed through interaction with electrons in the solution environment and containing reduced silver atoms (Ag), while the absorption at 340 nm is attributed to Ag4.Ag2 and Ag4 are neutral silver particles (or molecules) formed through the aggregation of smaller silver atoms. These neutral intermediates represent temporary stages in the growth of larger silver particle [25, 26]. For copper foams, the molar absorptivity was calculated from the first sample ( 15.65M−1 mm−1 ), where both absorbance and nitrate concentration were known.

Over the first 20 min of immersion, nitrate concentration increased almost linearly for both foams due to the gradual coating of silver on the foam surface, while nitrate ions remained constant because they do not participate in the reaction. For copper foam at 21 min , a sharp decrease in nitrate concentration was observed (red arrows in Figs. 2c and 2b), indicating partial detachment of silver and suggesting the formation of multilayer coatings. Based on these results, 20 min was determined as the optimal silver deposition time, corresponding to the highest nitrate concentration and absorbance, ensuring uniform coating. For nickel foams, the optimal immersion time was similarly investigated using UV-Vis spectroscopy. The molecular absorption coefficient ( 5.7M−1 mm−1 ) was determined from the first sample, where both nitrate concentration and absorbance were known. A representative absorption spectrum from sixteen measurements (Fig. 2a) shows a peak around 304 nm
corresponding to nitrate anions. The absorption intensity is influenced by ethanol, making it difficult to directly determine the Ag+concentration. Nitrate concentration increased up to 20 min , which was identified as the optimal silver deposition time, corresponding to the maximum absorbance and nitrate concentration, ensuring effective and uniform silver coating on the nickel foam (Figs. 2c and 2d). Immersion time at each stage of the process is a critical parameter for the morphological modification of metallic foams and the development of hydrophobic coatings, as it directly affects surface microroughness and surface energy reduction.

[picture]

Fig.2. (a) Indicative UV-vis absorption spectra for copper and nickel foams, (b) WCA measurements for different time of immersion in AgNO3 and in stearic acid solutions, (c) nitrate ion absorbance versus time for copper and nickel foams, and (d) nitrate ion concentration versus time for copper and nickel foams.

Fig.2. (a) Indicative UV-vis absorption spectra for copper and nickel foams, (b) WCA measurements for different time of immersion in A g N O 3 and in stearic acid solutions, (c) nitrate ion absorbance versus time for copper and nickel foams, and (d) nitrate ion concentration versus time for copper and nickel foams.

To further confirm the optimal immersion time in the AgNO3 solution identified by UV-Vis analysis, water contact angle (WCA) measurements were performed for both copper and nickel foams, as shown in Fig. 2b. The maximum WCA values were obtained at an immersion time of 20 min, verifying this duration as optimal for silver deposition. Subsequently, the effect of the stearic acid treatment was also evaluated through WCA measurements (Fig. 2b), since UV-Vis spectroscopy is not suitable for monitoring stearic acid in the UV-visible range. For a fixed AgNO3 immersion time of 20 min , the WCA reaches its maximum at 50 min of immersion in the stearic acid solution for both substrates. Under these optimal conditions, copper foams exhibit a superhydrophobic behavior with a WCA of 180∘, while nickel foams reach a maximum WCA of approximately 147∘. The highest standard deviations measured in this study were ±1∘ for WCA on copper and ±2∘ on nickel substrates. For nitrate ion absorbance, the maximum standard deviation was ±0.015 for the copper substrate and ±0.034 for the nickel substrate. Regarding nitrate ion concentration, the corresponding standard deviations were ±0.0018M for copper foams and ±0.0026M for nickel foams. These low standard deviation values indicate good repeatability and high sample homogeneity across the performed measurements.

Characterization of Hydrophobic Metallic Foams. SEM images of copper foams reveal silver moss-like structures along with stearate (copper or silver) nanowires (Figs. 3a and 3b). In the case of nickel foams, SEM images show uniformly distributed silver particles, likely formed by in situ reduction of silver ions on the foam substrate followed by nanoparticle agglomeration, and larger flake-like surface features, likely arising from the formation of a silver stearate phase during the stearic acid treatment (Figs. 3c and 3d).

[picture]

Fig. 3. SEM images of the morphologies developed on ( a,b ) copper and ( c,d ) nickel foams.

Fig. 3. SEM images of the morphologies developed on ( a , b ) copper and ( c , d ) nickel foams.

Thermal and Chemical Stability of Hydrophobic Copper and Nickel Foams

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The stability of the hydrophobic copper and nickel foams was investigated under various environmental conditions, with particular emphasis on comparing the behavior of the two metallic substrates.

Thermogravimetric analysis (TGA) was performed for both copper and nickel foams over a temperature range of 25−200∘C to evaluate their thermal stability. Regarding the thermal stability of the foams, no measurable weight loss was observed for either foam throughout the entire temperature range, indicating the absence of thermal degradation under the applied conditions (Fig. 4a). Under thermal exposure at −15∘C for residence times of up to 48 hours, clear differences between copper and nickel foams were observed. Copper foams exhibited superior thermal stability, maintaining high WCA values throughout the exposure period with only minor fluctuations (Fig. 4b). In contrast, nickel foams showed a more pronounced, time-dependent decrease in WCA. Although the nickel surfaces remained hydrophobic, their contact angles decreased from approximately 146∘ to about 135∘ (Fig. 4c).

[picture]

Fig. 4. (a) Thermogravimetric (TGA) curve of the samples. Contact angle measurements after thermal treatment at −15∘C, (b) of the hydrophobic copper foams and (c) of the hydrophobic nickel foams. (d) Contact angle measurements of the samples after immersion in sodium chloride solution. Contact angle measurements after chemical treatment at different pH values, (e) of the hydrophobic copper foams and (f) of hydrophobic nickel foams.

Fig. 4. (a) Thermogravimetric (TGA) curve of the samples. Contact angle measurements after thermal treatment at − 15 ∘ C , (b) of the hydrophobic copper foams and (c) of the hydrophobic nickel foams. (d) Contact angle measurements of the samples after immersion in sodium chloride solution. Contact angle measurements after chemical treatment at different pH values, (e) of the hydrophobic copper foams and (f) of hydrophobic nickel foams.

A decrease in the water contact angle is observed after immersion in a 3.5wt%NaCl solution, indicating a reduction in surface hydrophobicity (Fig. 4d). Despite this decrease, hydrophobic behavior is maintained. Copper foams exhibit contact angle values above 140∘ after 48 h of exposure, while nickel foams retain contact angles above 130∘, confirming the stability of hydrophobic properties under saline conditions (Fig. 4d).

Similar substrate-dependent behavior was observed during chemical stability tests (Figs. 4e and 4f). Copper foams retained their hydrophobic properties in both acidic ( pH 2 ) and alkaline ( pH 13 ) solutions, with minimal variation in WCA even after extended immersion. Nickel foams, on the other hand, exhibited an initial reduction in WCA during the first hours of exposure, followed by stabilization at approximately 135∘. While this response confirms that the hydrophobic coating on
nickel remains functional, it also highlights a comparatively lower resistance to chemical exposure relative to copper.

Figs. 4e and 4f also show the evolution of the water contact angle WCA of the coated copper and nickel foams after immersion in deionized water ( pH 7 ) for different residence times. For copper foams, the initial superhydrophobic state (WCA =180∘ ) gradually decreases with increasing immersion time. Nevertheless, even after 48 h of exposure, the WCA remains above 140∘, indicating good retention of hydrophobicity. In the case of nickel foams, the initial WCA ( 147∘ ) also shows a gradual decline with residence time. The decrease is moderate, suggesting enhanced stability of the coating on nickel substrates. After 48 h, WCA values remain around 135∘, confirming the durability of the hydrophobic behavior under prolonged aqueous exposure. Regarding the WCA measurements in chemical stability tests, the maximum deviations observed were ±3∘.

Overall, the comparative analysis demonstrates that copper foams offer higher stability and better retention of hydrophobic performance across all tested environments, whereas nickel foams, although slightly less stable in terms of WCA retention, maintain their hydrophobic character and structural integrity.

Oil-water Separation

The original version of this paper is available on https://www.scientific.net/SSP.391.95.pdf

Figs. 5a and 5b show the experimental setup for the oil-water separation experiments and the viscosities of the oils used, respectively. The results demonstrated that separation efficiency is strongly influenced by oil viscosity. Low-viscosity oils achieved separation efficiencies of nearly 99%, whereas higher-viscosity oils exhibited slightly reduced performance. The thickest oil ( 80 W− 90) showed a minor decrease in efficiency; however, in all cases, it remained above 95%. Oils with intermediate viscosity, such as 10 W−40 engine oil, performed between the low- and high-viscosity oils. For sunflower oil, copper foams achieved a slightly higher efficiency of 96%, while nickel foams maintained 95% (Figs. 5c and 5d). The maximum deviations regarding the separation efficiencies were ±0.5% for both substrates.

These findings suggest that increased viscosity hinders oil permeation through the porous superhydrophobic structure, causing small reductions in efficiency without compromising overall performance. Both copper and nickel foams maintained stable separation efficiencies over 15 consecutive filtration cycles, demonstrating their robustness and reusability. Overall, the results confirm that hydrophobic metallic foams provide efficient and reliable oil-water separation across a wide range of oil viscosities, with optimal performance observed for low-viscosity oils and consistently high efficiency maintained even for highly viscous engine oils.

Separation experiments were also conducted under forced filtration conditions. Fig. 5e illustrates a representative copper foam and the corresponding forced filtration setup. In all cases examined, separation efficiencies remained above 95%, and the same viscosity-dependent trend observed under gravity-driven filtration was preserved.

[picture]

Fig. 5. (a) Tank for oil-water separation, (b) oil viscosities used, separation efficiency of coated (c) copper and (d) nickel foams, for different types of oil after 15 filtration cycles. (e) Schematic illustration of the foam placed in a beaker with water and oil and the foam immersed in oil during filtration.

Fig. 5. (a) Tank for oil-water separation, (b) oil viscosities used, separation efficiency of coated (c) copper and (d) nickel foams, for different types of oil after 15 filtration cycles. (e) Schematic illustration of the foam placed in a beaker with water and oil and the foam immersed in oil during filtration.

Potentiodynamic Polarization Measurements

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The polarization curves reveal the potentiodynamic electrochemical behavior of hydrophobiccoated copper and nickel foams compared with their uncoated counterparts. As shown in Fig. 6a, the uncoated copper foam exhibits a corrosion potential of approximately -220 mV vs. SCE and relatively high anodic current densities at low overpotentials, indicating active corrosion in the chloride-containing medium. In contrast, the hydrophobic-coated copper foam shows a positive shift in corrosion potential to around -200 mV vs. SCE, together with a significant suppression of current density over a wide potential range. This behavior indicates improved corrosion resistance due to reduced anodic dissolution and inhibited cathodic reactions. The presence of two current peaks in the anodic region of the uncoated copper foam, observed at less active dissolution potentials (around +100 mV vs. SCE), is attributed to the sequential formation of Cu(I) and Cu(II) corrosion products, likely associated with chloride-induced complexation, in agreement with literature [27,28].

Similarly, Fig. 6b shows that the uncoated nickel foam has a corrosion potential near -400 mV vs. SCE and exhibits a steeper rise in current density, reflecting more pronounced anodic dissolution. Upon hydrophobic modification, the nickel foam displays a marked noble shift in corrosion potential to approximately -100 mV vs. SCE, accompanied by substantially lower current densities across the investigated potential range. The break in the anodic branch of the uncoated nickel foam observed near 0 mV vs. SCE is attributed to the formation of nickel oxide/hydroxide species, indicating the onset of passivation.

Overall, hydrophobic surface modification significantly improves the corrosion resistance of both metallic foams by limiting electrolyte penetration and reducing chloride ion transport within the porous structures. The enhancement is particularly pronounced for nickel foam, where the corrosion current is reduced by nearly two orders of magnitude, whereas copper foam exhibits a more moderate yet still significant reduction. These results suggest that the hydrophobic treatments induce a superhydrophobic surface state through increased micro-roughness of the coating, promoting air entrapment within the foam pores (Cassie-Baxter regime). This composite solid-air interface effectively repels water and delays corrosion initiation. While nickel foam benefits from its inherently higher baseline nobility, copper foam shows a stronger relative dependence on hydrophobic modification due to its poorer corrosion resistance in the uncoated state. The multiple anodic features observed for the coated copper foam may indicate intermittent electrolyte penetration through trapped air pockets, leading to localized oxide formation, whereas the coated nickel foam exhibits more stable anodic behavior even at low overpotentials.

[picture]

Fig. 6. Potentiodynamic polarization curves for the (a) uncoated and superhydrophobic copper and (b) uncoated plus hydrophobic nickel foams, after immersion in sodium chloride ( NaCl ) solution.

Fig. 6. Potentiodynamic polarization curves for the (a) uncoated and superhydrophobic copper and (b) uncoated plus hydrophobic nickel foams, after immersion in sodium chloride ( NaCl ) solution.

SEM images of coated copper and nickel foams after 48 h of immersion in NaCl solution are presented in Figs. 7a and 7b. The nanowire and dendritic structures on copper foams show some reduction and partial degradation. For nickel foams, slight alterations of the silver nanoparticles are observed, with some particles appearing partially modified. Despite these morphological changes, contact angle measurements indicate that the foams retain their hydrophobic properties as already shown in Fig. 4d.

[picture]

Fig. 7. SEM images of coated (a) copper and (b) nickel foams, after immersion in sodium chloride (NaCl) solution for 48 h .

Fig. 7. SEM images of coated (a) copper and (b) nickel foams, after immersion in sodium chloride ( N a C l ) solution for 48 h .

Summary

The original version of this paper is available on https://www.scientific.net/SSP.391.95.pdf

Superhydrophobic copper and hydrophobic nickel foams were prepared via a two-step chemical solution immersion process. Copper foams exhibited dual microstructures from silver precipitation and stearic acid deposition, achieving WCA up to 180∘. They demonstrated high thermal and chemical stability plus oil-water separation efficiencies above 96%. Nickel foams showed dense silver nanoparticle coverage with flake-like structures, reaching WCA of 147∘ and maintaining above 135∘ after 48 h under harsh conditions. They also exhibited separation efficiencies over 95% across multiple cycles and enhanced corrosion resistance. Both foams exhibit strong hydrophobicity, thermal plus chemical stability, and excellent separation performance, offering superior durability, reusability, and corrosion protection. Both materials are highly effective for oil-water separation and environmental rehabilitation. Potentiodynamic polarization results show that the hydrophobic coating decreases corrosion current for both copper and nickel foams. Overall, these results underline the influence of the metallic substrate on coating durability and support the selection of copper foams for applications requiring enhanced long-term stability, while nickel foams remain a reliable alternative for demanding environments.

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The original version of this paper is available on https://www.scientific.net/SSP.391.95.pdf

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Solid State Phenomena, ISSN: 1662-9779, Vol. 391, pp 107-114

doi: 10.4028/p-7Bow8J

© 2026 The Author(s). Published by Trans Tech Publications Ltd, Switzerland.

Submitted:
2026-01-19

Revised:
2026-02-25

Accepted:
2026-02-25

Online:
2026-04-20

Origami-Inspired Design for Manufacture of Composite Parts

The original version of this paper is available on https://www.scientific.net/SSP.391.107.pdf

and James Kratz d 

Bristol Composites Institute, University of Bristol, United Kingdom

a  sutharsanan.navaratnarajah@bristol.ac.uk, b  vincent.maes@bristol.ac.uk, c  m.schenk@bristol.ac.uk, djames.kratz@bristol.ac.uk (*corresponding author)

Keywords: forming, origami, composite structures, non-crimp fabrics

Abstract

During dry fibre processing, manufacturing geometrically complex composite parts often produces wrinkles when fabrics deform beyond their shear limits. This work proposes a design-formanufacture approach based on origami principles, which modifies component geometry so that deformation remains within allowable deformation limits. A baseline aircraft spar geometry is considered; an origami-inspired version, along with several intermediate designs between these two extremes, are generated. Preliminary forming trials with unidirectional non-crimp fabrics show that the origami-based geometry is inherently manufacturable without defects, and that a selected intermediate design also form successfully, confirming a larger manufacturable design space than classical origami permits. Results further show that wrinkle severity increases with increasing angular defect. This provides a foundation for linking geometric measures to draping mechanics to guide the design of wrinkle-free composite components without requiring computationally expensive simulations.

Introduction

The original version of this paper is available on https://www.scientific.net/SSP.391.107.pdf

To satisfy market demands, the automotive and aerospace industries aim to produce lightweight and high-performance composite components at high production rates. Dry fibre processing such as Liquid Composite Moulding (LCM) enables high production rates at relatively low cost across various stages of the manufacturing and assembly process [1]. Non-crimp fabrics (NCFs) are preferred for dry fibre processing due to their higher deposition rates and superior mechanical properties (strength and stiffness) resulting from the absence of tow undulations, making them more advantageous than woven fabrics [2,3]. However, NCFs introduce new challenges when creating composite parts with more complex geometries.

In order to conform to doubly curved three-dimensional surfaces, the flat composite sheets must undergo in-plane shear deformations; when the shear deformation exceeds the fabric's shear limit, out-of-plane deformations such as wrinkles occur during the forming process, which negatively impact the mechanical performance of the final part. NCFs have reduced formability compared to woven fabrics due to the lack of 'cross over points' between tows [1,2], which may therefore result in additional defects (thereby offsetting the superior mechanical properties of the NCF). Wrinkles can be managed by either compressing them, cutting and redistributing excess material and adding extra material to reinforce the part. However, these approaches require intensive manual labour and introduce inherent variability. Therefore, the industry aims to minimize wrinkles, focusing on improving process repeatability with reduced labour demands.

Addressing this challenge requires controlling or modifying the key influencing parameters [4]: the process [5-8], the geometry [9, 10], and the material [11, 12]. While process and material effects are well studied, the role of geometry offers a more fundamental route to reducing wrinkling. However, geometry-based approaches are still underexplored because structural design and manufacturing have historically been treated as separate stages, and the effect of geometry on defect formation is not yet fully understood. To help address this gap, the present work adopts a design-formanufacture strategy, using origami-inspired folding techniques to create geometries which are inherently manufacturable without defects.

Engineering origami is a rapidly evolving field that draws inspiration from the art of origami to address engineering problems. Its applications range from deployable space structures [13] to miniaturized medical devices [14] and mechanical metamaterials [15]. Crucially, origami can generate complex 3D geometries from a flat sheet without any material stretching or shearing, by bending the material and folding along defined creases. Hence, origami-based designs can minimize stretching and shearing during the forming process, thereby reducing the occurrence of wrinkling defects. Our previous studies on origami geometries with curved folds have shown that they are inherently manufacturable without defects [16]. As NCFs can absorb some shearing without wrinkling [17] the geometric design space of origami-inspired geometries can be expanded, enabling the design of inherently manufacturable components that continue to meet structural and other design requirements.

To demonstrate the application of these design principles we consider a case study geometry as a baseline, inspired by Hallander et al. [19] - a ramped aircraft spar with recessing features in the flange and web as shown in Fig. 1. These recessing features on all faces make the fabric prone to defects during forming [19]. The aim of this study is to systematically modify the baseline spar geometry to achieve manufacturability without wrinkles, while continuing to satisfy the design constraints. First an origami-based version of the ramped spar is created, which exhibits zero defects; by allowing for the shearability of the composite fabric, the design is gradually adjusted to more closely match the baseline geometry. Initial forming trials with NCF fabrics show a transition between geometries that are manufacturable without defects, to those which exhibit significant wrinkling.

[picture]

Fig.1. Baseline ramped aircraft spar-inspired geometry, symmetric about the xz-plane, with a 1:5 ramp on both the flanges and the web.

Fig.1. Baseline ramped aircraft spar-inspired geometry, symmetric about the xz-plane, with a 1 : 5 ramp on both the flanges and the web.

Case Study Geometries

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The baseline geometry considered here is inspired by Hallander's C-spar [18], which contains features representative of a wing spar. The recessing feature in the flanges ensures proper clearance and mating surfaces, allowing the wing ribs and the upper and lower wing covers to fit properly against the spar without geometric interference. Hence, the design requirement is to preserve the flange surfaces in their original form. In this study, instead of recesses only in the flanges, we introduce a recess in the web as well, with similar ramping as in [19]. The addition of the web recess increases the risk of wrinkling due to the added geometric complexity. We present a geometry with more aggressive ramping (1:5) in both directions compared to [19]. This provides more space for additional components to pass through the flange recesses and serves as a proof of concept for how extreme the ramping can be while still producing a double-curved surface that does not exhibit
wrinkles. In this section, we present a simple method to quantify geometrical features linked to wrinkle formation.

Gaussian curvature ( K ) is an intrinsic geometric property of points on a surface; the Gaussian curvature remains invariant under bending deformations, including folding along sharp creases, but changes when the surface undergoes in-plane deformations [20,21]. We here differentiate between two types of surfaces: developable ( K=0 ) and non-developable ( K≠0 ) surfaces. When forming a flat sheet of fabric - which has zero Gaussian curvature - over a developable surface, the material undergoes only isometric deformations (no shearing or stretching). In contrast, when conforming over a non-developable surface, the material experiences stretching and shearing. Hence, Gaussian curvature can be linked to the tendency for wrinkles to form.

The baseline geometry consists of discrete flat surfaces joined at vertices, as shown in Fig. 2a. This forms a polygonal surface, for which the Gaussian curvature can be expressed as [21]:

K≈βA= angular defect of vertex  area associated with vertex (1)

The angular defect ( β ) at a flattened vertex is defined as:

β=2π− sum of interior angles (2)

When β>0, it is referred to as an angular defect, and when β<0, it is referred to as an angular excess. As shown in Fig. 2b, unfolding a small area around the vertex results in a gap for an angular defect and an overlap for an angular excess. Consequently, forming a flat continuous fabric without any cuts over an angular defect leads to excess material, whereas forming over an angular excess result in insufficient material.

[picture]

Fig. 2. (a) Vector notation and geometric parameters used to derive vertex angles at a discrete vertex of the baseline geometry, including the position vector r→AB, ramp angles θf and θw, and local coordinate axes. (b) Schematic illustration of angular defect and angular excess at vertices A and B, showing the relationship between vertex angles α and γ and the resulting angular defect (βA) and angular excess (βB).

Fig. 2. (a) Vector notation and geometric parameters used to derive vertex angles at a discrete vertex of the baseline geometry, including the position vector r → A B , ramp angles θ f and θ w , and local coordinate axes. (b) Schematic illustration of angular defect and angular excess at vertices A and B, showing the relationship between vertex angles α and γ and the resulting angular defect ( β A ) and angular excess ( β B ) .

Since the baseline geometry is symmetric about the xz-plane and the mid yz-plane, we only need to consider the angular defect at vertex A and angular excess at vertex B. Further, the magnitude of the angular defect is found to be equal to that of the angular excess, as shown in Fig. 2b. Accordingly,

β=βA=|βB|=π−(α+γ)(3)

Hence, for this geometry, a single angular defect term ( β ) is sufficient to describe the nondevelopability of the geometry. It should be noted that βA is of opposite sign to βB. Although the total Gaussian curvature over the surface satisfies ∑K=0, this does not imply that the geometry is developable or formable. The angles α and γ are derived via vector algebra; the position vector r→AB can be expressed as:

r→AB=Lı→+Ltanθfȷ→+Ltanθwk→(4)

where L is the longitudinal length of the ramping section along the x -axis, and θf and θw are the ramp angles of the flange and web, respectively. The vertex angles α and γ can be calculated as:

cosα=r→‖r→‖·J→=tanθftanθf2+tanθw2+1cosγ=r→‖r→‖·k→=tanθwtanθf2+tanθw2+1

(5)(6)

Accordingly, the angular defect (and equivalently the magnitude of the angular excess) at the vertex can be expressed as:

β=βA′=π−(cos−1(tanθftanθf2+tanθw2+1)+cos−1(tanθwtanθf2+tanθw2+1))(7)

As mentioned above, the design constraints require that the flanges remain fixed, so θf is constant, while the web angle θw is allowed to vary. To make the baseline geometry developable, the angular defects at all vertices must be zero. Thus,

β=π−(α+γ)=0(8)

Solving Eqs. 7 and 8 yields:

cosα=−cosγ⇒θw=−θf(9)

This implies that the web must be ramped upward with the same slope as the flange to make the surface developable; see Fig. 3(d). Accordingly, the origami-based geometry features web ramping at 1:5. However, composite fabrics can accommodate some shearing, so the spar geometries do not need to be strictly developable; this allows us to design intermediate geometries with allowable angular defects. The intermediate geometries include one with no web ramping and other with an upward web ramping of 1:10. As illustrated in Fig 3, β decreases from left to right: the baseline exhibits the highest angular defect, while the origami-inspired geometry achieves zero angular defect.

[picture]

Fig. 3. Case-study geometries with decreasing angular defect β : (a) baseline geometry with high angular defect, (b-c) intermediate geometries with progressively reduced angular defect, and (d) origami-inspired geometry with zero angular defect ( β=0 ).

Fig. 3. Case-study geometries with decreasing angular defect β : (a) baseline geometry with high angular defect, (b-c) intermediate geometries with progressively reduced angular defect, and (d) origami-inspired geometry with zero angular defect ( β = 0 ).

Initial Fabric Forming Trials and Discussion

The original version of this paper is available on https://www.scientific.net/SSP.391.107.pdf

This section presents preliminary experimental forming results and discusses their implications. A fillet with 5 mm radius was introduced along all fold lines to accommodate the bending of the fabric. Note that the fillet introduces a small amount of double curvature, and thus non-developability, at the vertices, even in surfaces that are developable; this is not yet accounted for in the calculation of the angular defect or excess of each vertex. All geometries were rapidly prototyped using thermoplastic (PLA) via Fused Deposition Modelling (FDM), allowing iterative, geometry-driven forming studies. Tool tack was applied to ensure a smooth forming surface.

Preliminary trials were carried out using a hand layup procedure. A unidirectional single-ply NCF (SGL, 300 gsm ), with fibres aligned in the spanwise direction, was placed and formed in a consistent sequence, as indicated by the arrows in Fig. 4a. Visual inspection shows that forming over origamibased geometries with slight non-developability produces no visible defects. As discussed earlier, forming over origami-based geometries leads primarily to isometric deformation, i.e., without stretching or shearing. However, the addition of a fillet introduces local non-developability, and forming paper over such geometries would typically lead to defects [16]. In contrast, fabrics, unlike paper, can accommodate a certain amount of shear; therefore, the presence of fillet edges does not result in visible defects.

Similarly, the geometry with web ramping at 1:10(β=0.1, Fig. 4b) exhibits no defects. This further demonstrates that the NCF can absorb a limited amount of shear without wrinkling. In contrast, the other two geometries display visible defects, and the baseline geometry shows significant wrinkling due to its large angular defect (Fig. 4d).

As expected, increasing angular defect correlates with increased severity of wrinkling. However, angular defect is a highly localised measure. For this geometry, this metric can be linked to kinematic draping models [22], which account for the continuous nature of fabrics. This enables the design of geometries that remain manufacturable while being closer to the baseline.

This origami-based design-for-manufacture method enables the process to begin with an origamibased version of a baseline geometry that would otherwise exhibit wrinkling, and then progressively modify this origami version toward the baseline design while accounting for the shear capacity of the fabric.

[picture]

Fig. 4. Preliminary hand layup results for the case-study geometries. (a) Origami-inspired geometry ( β=0 ) with non-developable fillet edges formable without wrinkles; the arrow indicates the tape layup sequence. (b) Intermediate geometry with reduced web ramping ( β=0.1 ), also formable without defects. (c) Geometry with a flat web ( β=0.2 ) shows noticeable wrinkles and some difficulty during forming. (d) Baseline geometry (β=0.4) exhibiting severe wrinkling and poor formability.

Fig. 4. Preliminary hand layup results for the case-study geometries. (a) Origami-inspired geometry ( β = 0 ) with non-developable fillet edges formable without wrinkles; the arrow indicates the tape layup sequence. (b) Intermediate geometry with reduced web ramping ( β = 0.1 ), also formable without defects. (c) Geometry with a flat web ( β = 0.2 ) shows noticeable wrinkles and some difficulty during forming. (d) Baseline geometry ( β = 0.4 ) exhibiting severe wrinkling and poor formability.

Conclusion and Future Work

The original version of this paper is available on https://www.scientific.net/SSP.391.107.pdf

Complex geometries with double curvature often lead to defects during the composite forming process because fabrics undergo in-plane deformations that exceed their allowable limits. Our origami-inspired design-for-manufacture approach addresses this challenge by identifying shapes in which the material deforms within its shear capacity. Starting from a baseline geometry that exhibits significant wrinkling, we generated an origami-derived design and subsequently introduced two intermediate geometries between the origami and baseline configuration.

The results show that the origami-based geometry, even with the introduction of non-developable filleted edges, forms without any defects. Although origami imposes a strict geometric constraint, where ramping in one direction requires an equal ramp in the opposite direction, the preliminary results demonstrate that an intermediate geometry with reduced web ramping ( 1:10 ) also forms successfully. This confirms that the manufacturable design space extends beyond strict origami constraints when fabric shearability is taken into account. Moreover, the results suggest that wrinkle severity increases with angular defect, highlighting its importance as a geometric indicator for manufacturability.

Unlike the conventional process of forcing fabrics to conform to complex designs, this origamibased, non-conventional approach allows design modifications that satisfy functional requirements
while keeping material deformations within allowable limits. In doing so, it ensures manufacturability. This case study demonstrates the potential of the approach, and ongoing work is exploring additional aspects, as discussed in the following section.

These preliminary findings open several avenues for future work. First, the methodology is currently being used with different fabric types and fibre orientations to better understand material dependent limits. Second, the effect of fillet radii will be incorporated into the angular defect formulation. Third, the localised geometric measure (angular defect) will be linked with continuum kinematic draping models to predict maximum shear angles. This will enable correlation between angular defect, material behaviour, and a limiting value of θw for different fabrics beyond which wrinkles begin to form. Establishing this relationship will allow us to identify manufacturable geometries that remain as close as possible to the original design while avoiding defects.

References

The original version of this paper is available on https://www.scientific.net/SSP.391.107.pdf

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Solid State Phenomena, ISSN: 1662-9779, Vol. 391, pp 115-122

doi: 10.4028/p-eSLz2C

© 2026 The Author(s). Published by Trans Tech Publications Ltd, Switzerland.

Submitted:
2026-01-19

Revised:
2026-02-18

Accepted:
2026-02-18

Online:
2026-04-20

Surface Roughness Control in Incoloy 800HT via Laser Engraving Parameters

The original version of this paper is available on https://www.scientific.net/SSP.391.115.pdf

Evangelos Nikolidakis 1,a, Panagiotis Karmiris-Obratański 2, b, Emmanouil L. Papazoglou 3,c, Lisa Fatyga 2, d, Angelos P. Markopoulos 3,e*

1 School of Production Engineering & Management, Micromachining & Manufacturing Modeling Lab, University Campus Kounoupidiana, Technical University of Crete, 73100, Chania, Greece

2 Advanced Manufacturing Laboratory, Department of Manufacturing Systems, Faculty of Mechanical Engineering and Robotics, AGH University of Krakow, Krakow, Poland

3 School of Mechanical Engineering - Laboratory of Manufacturing Technology, National Technical University of Athens, Athens, Greece

a  krhthkournas@gmail.com, b  karmiris@agh.edu.pl, c  mlpapazoglou@gmail.com, d  dekster@agh.edu.pl, e*amark@mail.ntua.gr

Keywords: laser engraving, surface roughness, taguchi DoE, nickel.

Abstract

Although Laser engraving (LE) is increasingly adopted for precision surface texturing, the resulting surface is highly sensitive to the coupled thermal and hydrodynamic mechanisms governing laser-material interaction. In this experimental work, LE of Incoloy 800 HT is systematically investigated using an L9 Taguchi design of experiments considering Laser Power (LP), Laser Scanning Speed (LSS), and Laser Pulse Frequency (LPF) as control parameters. Surface roughness is quantified using the arithmetical mean height (Ra), maximum profile height (Rz), skewness (Rsk), and the height at material ratio Rmc=20%, enabling both amplitude- and function-oriented assessment of the engraved textures. The contribution of each parameter is evaluated through Analysis of Variance and response ranking, and regression-based correlations are established to support predictive selection of processing conditions. The results show that LP is the dominant factor for Ra, Rz, and Rmc( 20% ), while LSS primarily governs Rsk, reflecting the role of scanning speed in controlling melt redistribution and peak-valley balance. High cumulative energy conditions promote thermal accumulation, melt ejection, spatter redeposition, and recast formation, leading to substantially rougher surfaces, as corroborated by topography and SEM observations.

Introduction

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As laser-based machining technologies continue to mature and to be integrated into modern manufacturing systems, understanding the Surface Roughness (SR) and Surface Quality (SQ) generated after laser processing has become both a topic of strong scientific relevance and an issue of pure industrial significance. Laser-material interaction is intrinsically multi-physics and highly non-linear, which renders laser-based machining and manufacturing processes challenging to predict and control [1]. This challenge is further intensified in material-removal applications, where multiple removal mechanisms may take place simultaneously, including surface vaporization, melt ejection, phase explosion, spallation, Coulomb explosion, critical-point phase separation, fragmentation, and hydrodynamic expansion [2]. Compared with conventional machining material removal processes, laser-based approaches offer the ability to produce small-scale and geometrically complex features with ultra-high precision, while avoiding mechanically induced deformation and residual stresses. Moreover, despite the high local energy densities involved, the Heat Affected Zone (HAZ) and the associated thermally driven residual stresses and distortions are typically limited [3].

Among established laser-based material-removal processes, Laser Engraving (LE) remains one of the most widely implemented. In LE, material is removed in a layer-wise manner through repeated scanning passes of the laser beam over the target surface. In practical implementations, most commonly pulsed laser sources are utilized, and thus, each removed layer results from the spatial overlap of successive ablation craters generated during pulse-by-pulse interaction with the material
[4]. Consequently, the primary controllable parameters in LE include Laser Power (LP), Laser Scanning Speed (LSS), Laser Pulse Frequency (LPF), laser spot size, and pulse duration. These parameters require systematic investigation and optimization, considering both the processed material and the intended functional or geometrical outcome [5]. The available literature includes several works studying how LE process parameters affect the SR and SQ of machined surfaces. Šugár et al. [6] investigated the relationship between LE processing conditions and surface finish during the machining of an aluminum-copper alloy using a pulsed ytterbium fiber laser. Their findings indicated that LP and, more specifically, laser intensity (LI) are the most important factors governing the final machined Surface Texture (ST), as these parameters are directly associated with the dominant material-removal mechanisms and the thermo-hydraulic state of the irradiated material [7]. Chen et al. [8] concentrated on the surface morphology and SR of micro-grooves produced in an aluminum alloy using a nanosecond pulsed laser. Pritam et al. [9] evaluated the SR of Al-6063 considering control process parameters the LE, the LP, the LSS, the LPF and the scanning hatch spacing as control parameters. Notably, their analysis of variance (ANOVA) identified LSS as the most influential factor for average roughness (Ra), followed by hatch spacing and LPF, whereas LP exhibited the lowest statistical significance. In contrast, the work of Pattanayak et al. [10] on micro-engraving of 316L stainless steel for orthopedic applications reported LP as the primary determinant of surface quality, with LPF and LSS exerting secondary effects, thereby further emphasizing the material- and processdependent nature of parameter sensitivity.

From a materials science standpoint, alloys within the 800 series typically incorporate multiple alloying additions, including Ni, Cr, Mo, Al, and Ti. These elements improve high-temperature strength, mechanical performance, and corrosion resistance through a combination of precipitation strengthening, solid-solution effects, and the formation of protective oxide films [11]. In particular, Incoloy 800 HT is an austenitic, heat-resistant Ni−Fe−Cr alloy in which titanium nitrides, titanium carbides, and chromium carbides may be present within the microstructure, depending on composition and thermal history [12]. Owing to its stable performance over a broad temperature range, Incoloy 800 HT is widely applied in industrial sectors where thermal stability and environmental resistance are essential, including chemical processing, power generation, petrochemical systems, heat-treatment equipment, nuclear engineering, aerospace, and other specialized manufacturing environments [11,13].

Based on the above, while substantial work exists on related processes (laser polishing, additive manufacturing, and machining), direct studies specifically addressing engraving-driven roughness control in nickel-based alloys remain limited. In the present study, LE of Incoloy 800 HT is examined using an L9 Taguchi design of experiments. Three processing parameters are considered at three discrete levels: LP ( 8/12/16 W ), LSS ( 300/400/500 mm/s ), and LPF ( 30/40/50kHz ). SR is characterized using the arithmetical mean height (Ra), maximum profile height (Rz), skewness (Rsk), and load length ratio (Rmr(C)), whereas the influence of the selected parameters is quantified through Analysis of Variance (ANOVA). Finally, regression-based correlations are developed to support predictive modeling of surface roughness as a function of the applied machining conditions.

Materials and Methods

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For the present work, Incoloy 800 HT plates with a thickness of 3 mm were employed, with flatness and perpendicularity verified prior to machining. All experiments were performed on a LASERTEC 40 laser machining centre equipped with a nanosecond, pulsed, Q-switched fibre laser operating at a wavelength of 1064 nm , a nominal spot diameter of 30μ m, and a pulse duration of 100 ns . The experimental campaign followed an orthogonal L9 Taguchi DOE, considering LP, LSS, and LPF as control factors, each investigated at three levels (Table 1). A unidirectional cross-hatching strategy was adopted, and twenty layers were engraved for each parameter set to ensure operation under steady-state engraving conditions. The hatch spacing was set equal to the distance between successive pulses, thereby maintaining uniform areal coverage over the machined region. Surface roughness was evaluated using the arithmetical mean height (Ra), maximum profile height (Rz), skewness (Rsk), and the profile height at a load length ratio of c=20%(Rmr(c)). To ensure standards-compliant
characterization (ISO 4287, ISO 12085, ISO 13565-2/3, and ISO 25178-2), a randomly selected engraved area of 4×4 mm was evaluated. Measurements were performed using a Formtracer Avant H8 D4000-D. A 4 mm evaluation length was scanned, and 61 parallel profiles were acquired at a traverse speed of 0.5 mm/s to enhance the statistical robustness of the roughness characterization. The influence of the selected process parameters is quantified using ANOVA. Subsequently, regression-based correlations are established to enable predictive modeling of surface roughness as a function of the applied machining conditions.

Results and Discussion

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The machining parameters along with the SR measurements are listed in Table 1. Before presenting and discussing the results, it should be noted that the experimental design follows a Taguchi L9 orthogonal array, which efficiently screens the main effects of the selected factors at three discrete levels with a limited number of trials. However, using only three levels per factor limits the resolution of potential non-linear (higher-order) responses between levels, and the L9 array is primarily suited to estimating main effects rather than fully quantifying factor-factor interactions. Therefore, any interaction trends observed in the response plots are interpreted qualitatively, and the reported parameter effects should be understood within the discrete level space investigated.

Table 1. Machining conditions and SR measurements.

LP (W) | LSS (mm/s) | LPF (kHz) | VED (J/mm³) | Ra (μm) | Rz (μm) | Rsk | Rmc (μm)
8 | 300 | 30 | 37.73 | 1.08 | 8.18 | 0.39 | 4.56
8 | 400 | 40 | 28.29 | 1.13 | 8.32 | 0.29 | 3.75
8 | 500 | 50 | 22.64 | 0.85 | 6.69 | 0.25 | 3.29
12 | 300 | 40 | 56.59 | 2.87 | 18.68 | 0.49 | 9.65
12 | 400 | 50 | 42.44 | 1.86 | 14.01 | 0.32 | 8.56
12 | 500 | 30 | 33.95 | 1.43 | 10.16 | 0.10 | 4.47
16 | 300 | 50 | 75.45 | 8.63 | 50.92 | 0.29 | 23.54
16 | 400 | 30 | 56.59 | 1.91 | 13.61 | 0.26 | 7.10
16 | 500 | 40 | 45.27 | 1.68 | 11.99 | 0.17 | 5.80

Based on the results of Table 1, Fig. 1 presents the main-effects and interaction plots for Ra as a function of the machining parameters. The main-effects trends support physically consistent conclusions, given that Ra is governed primarily by the net energy delivered per unit area. Specifically, increasing LP increases the effective energy input (Volumetric Energy Density - VED ∝ LP/LSS) and, consequently, elevates peak surface temperature and molten material volume, intensifying phenomena like recoil pressure-driven melt expulsion, and spatter redeposition of the material. The rapid re-solidification of ejected droplets and recast material produces pronounced asperities and a thicker re-solidified layer, thereby increasing Ra. The increase in Ra with LPF can be rationalized by the higher pulse Overlap Ratio (OR) and the reduced inter-pulse cooling time, which, within the examined regime, these effects outweigh any reduction in laser peak power per pulse, leading to thermal accumulation, sustained molten films, and repeated crater-rim formation that coalesces into a rougher surface topography. On the other hand, higher LSS reduces the effective dwell time and the VED, limiting heat accumulation and melt-pool growth, favoring a more controlled ablation-limited-melt regime and thus lowering Ra. The interaction plots provide deeper insight into the combined parameter effects and confirm that the general trends remain valid, while also revealing notable non-linearities. The most pronounced increase in Ra occurs when LP rises from 12 to 16 W , particularly when combined with the lowest LSS ( 300 mm/s ) and the highest LPF ( 50 kHz ), i.e., under conditions that maximize cumulative energy deposition and promote unstable melt ejection and re-solidification. Additionally, different parameter combinations yield comparable Ra values (e.g., LP=8 W with LPF=30−40kHz ), highlighting that surface finish is dictated not only by the magnitude of each factor but also by their coupling through overlap, dwell time, and thermal accumulation. Finally, some seemingly counterintuitive behaviors, such as the decrease in Ra at LPF
=40kHz when LP increases from 12 to 16 W , or analogous deviations at LSS =400 mm/s, further evidence the strongly non-linear and condition-dependent nature of LE, where modest shifts in overlap and melt dynamics can transition the process between regimes that significantly alter the roughness of the machined surface.

[picture]

Fig. 1. Main Effect Plot and Interaction Plot of the Ra as function of the machining conditions.

Fig. 1. Main Effect Plot and Interaction Plot of the Ra as function of the machining conditions.

The Rz main-effects and interaction plots that are presented in Fig. 2, exhibit trends fully consistent with Ra which can be interpreted through the same energy- and melt-dynamics framework, with the key distinction that Rz is more sensitive to extreme topographic features (deep valleys and high peaks) generated during laser engraving. In the main-effects plot, Rz increases strongly with LP and LPF, and decreases with increasing LSS, mirroring the Ra response, i.e., higher LP and higher LPF increase cumulative energy deposition (via increased dwell/overlap and reduced inter-pulse cooling), promoting a melt-dominated regime characterized by deeper ablation craters, more pronounced rim formation, and extensive spatter/recast redeposition, which preferentially amplifies peak-to-valley height and thus elevates Rz. Conversely, higher LSS reduces the effective VED, limits melt-pool growth and hydrodynamic instabilities, and suppresses large craters and resolidified protuberances, yielding lower Rz. The interaction plots reinforce this coupling and highlight that the largest Rz excursions occur under "high-energy" combinations (e.g., high LP together with low LSS and/or high LPF), where process stability deteriorates and localized melt ejection and re-solidification generate severe peaks/valleys. Therefore, while Ra captures the overall roughness level, the larger dynamic range of Rz in these plots indicates that the same parameter changes that increase Ra also tend to magnify isolated defects and crater-spatter features, making Rz a complementary metric that accentuates the extremes of the laser-affected surface morphology.

[picture]

Fig. 2. Main Effect Plot and Interaction Plot of the Rz as function of the machining conditions.

Fig. 2. Main Effect Plot and Interaction Plot of the Rz as function of the machining conditions.

Rsk is a surface roughness parameter that quantifies the asymmetry of the surface profile relative to the mean line, indicating how peaks and valleys are distributed across the surface. A positive Rsk value signifies a surface dominated by peaks, while a negative value indicates a surface dominated by valleys. An Rsk value of zero corresponds to a symmetrical height distribution, typically seen in Gaussian profiles [14]. Surface skewness is related with the initiation of erosion damages with surfaces with negative Rsk (valley-dominated) be more prone to certain types of damage [15], with load-bearing and friction characteristics, as for negative Rsk fewer peaks are rubbed off during contact, while the skewness of a surface can also influence bacterial adhesion and biofilm formation, with positive Rsk (more peaks) or negative Rsk (more valleys) affecting the attachment of cells or
bacteria differently. The Rsk plots in Fig. 3 provide a complementary interpretation to Ra and Rz by indicating whether the LEed topography is predominantly governed by peaks or valleys. Across all investigated parameters levels, Rsk remains positive, demonstrating that the generated surfaces are generally peak-dominated, which is consistent with LE mechanisms that promote the formation of recast ridges, crater rims, and spatter redeposition. The interaction plots indicate that the most significant and consistent influence on Rsk is exerted by LSS, as increasing LSS, most clearly from 300 to 500 mm/s, typically reduces Rsk. This behavior is physically reasonable since at low LSS, the longer dwell time and stronger thermal accumulation promote melt-pool perseverance, rim build-up, and droplet-based redeposition, leading to a height distribution skewed toward high asperities. In contrast, although the main-effects plot suggests weak tendencies with respect to LP and LPF, the interaction plots reveal non-linear and non-monotonic responses, and thus, a robust monotonic relationship between LP, LPF and Rsk cannot be established within the examined parameter space. Importantly, parameter combinations that maximize Ra/Rz do not necessarily maximize Rsk, highlighting that Ra and Rz primarily quantify roughness amplitude, whereas Rsk reflects the balance between peak-forming recast/spatter features and crater-induced valley formation, which can shift depending on the local stability of melt ejection and re-solidification during LE.

[picture]

Fig. 3. Main Effect Plot and Interaction Plot of the Rsk as function of the machining conditions.

Fig. 3. Main Effect Plot and Interaction Plot of the Rsk as function of the machining conditions.

Finally, the Relative Material Curve in surface texture analysis, specifically indicating the percentage of the surface that remains at a given height level (c) when cut horizontally, showing how much material is present vs. air at that height. It offers a complete picture of a surface's functional properties, crucial for understanding wear, lubrication, and friction, especially in engineering applications. In practice, Rmc support functional surface design by linking topography to performance since high valley material ratio indicates good lubricant retention, while low peak material ratio reflects smoother, wear-resistant summits. They are especially useful for stratified/coated or textured surfaces, where different layers or features contribute differently to load bearing and fluid storage. The main-effects and interaction plots for the height at Rmc=20% (Fig. 4) follow the same overall trends observed for Ra/Rz, indicating that the parameter sets that increase roughness amplitude also deepen the profile level required to reach a 20% material ratio. In particular, increasing LP and LPF raises the height at Rmc(20%), whereas increasing LSS reduces it. These results are consistent with higher cumulative energy input promoting a melt-dominated regime with stronger crater formation, rim build-up, and redeposited spatter that broadens the height distribution and shifts the lower loadbearing portion of the profile to more extreme values. The interaction plots further confirm that the largest Rmc(20%) heights occur under "high-energy" combinations (notably high LP coupled with low LSS and/or high LPF), where thermal accumulation and unstable melt ejection generate pronounced peak-valley contrasts. Conversely, at higher LSS the reduced dwell time suppresses melt pooling and resolidified ridges, narrowing the distribution and lowering the Rmc(20%) height. Overall, this behavior is in line with the Ra/Rz response and supports the interpretation that LE conditions controlling melt generation, ejection, and recast largely dictate both amplitude roughness and functional load-bearing/lubricant-retention descriptors.

[picture]

Fig. 4. Main Effect Plot and Interaction Plot of Height for Rmc 20% as function of the machining conditions.

Fig. 4. Main Effect Plot and Interaction Plot of Height for Rmc 20 % as function of the machining conditions.

Table 2. ANOVA for the Ra, Rz, Rsk and Rmc(20%) in respect of the machining conditions.

|  | LP | LSS | LPF | LP·LSS | LP·LPF | LSS·LPF | error | R-sq (adj)
Ra | Contr. (%) | 29.89 | 26.45 | 17.10 | 9.08 | 15.82 | 0.28 | 1.37 | 94.51
Rz | Contr. (%) | 31.75 | 26.73 | 17.57 | 8.31 | 14.57 | 0.17 | 0.90 | 96.39
Rsk | Contr. (%) | 6.74 | 67.16 | 2.32 | 2.09 | 16.25 | 0.10 | 5.34 | 78.64
Rmc(20%) | Contr. (%) | 32.71 | 30.99 | 19.64 | 5.95 | 10.57 | 0.09 | 0.05 | 99.80

Table 2 summarizes the ANOVA for Ra, Rz, Rsk, and Rmc( 20% ) with respect to the examined machining parameters and also considering the interactions terms. It confirms that the amplituderelated roughness descriptors (Ra, Rz) and the functional height parameter Rmc(20%) are governed primarily by the process variables controlling net energy input and thermal accumulation, with LP and LSS consistently ranking first and second and together accounting for the majority of variance ( ≈56−64% combined contribution), while LPF provides a smaller but still measurable effect (≈17−20% ). In contrast, Rsk is overwhelmingly dominated by LSS ( ≈67% contribution), indicating that LSS is the key driver of peak-valley asymmetry. Among interaction terms, LP•LPF is the most influential ( ≈11−16% ) across all metrics, highlighting the role of pulse overlap/inter-pulse cooling in modulating melt stability, spatter generation, and re-solidification when power is increased. The LP•LSS term shows a secondary but non-negligible contribution ( ≈6−9% ) for Ra/Rz/Rmc(20%), consistent with the coupling between power and dwell time (line energy), whereas LSS•LPF is negligible ( ~0.1−0.3% ). Finally, the low residual errors for Ra,Rz, and Rmc(20%) and the high adjusted R2 values (≈94−100%) indicate a robust statistical description of these responses within the tested factor space, while the higher error and lower adjusted R2 for Rsk ( ≈79% ) reflect its greater sensitivity to localized, non-linear melt-ejection and redeposition events that are not fully captured by main effects and low-order interactions.

In Fig. 5, the topography and SEM micrographs corresponding to the two extreme cases, i.e., lowest and highest SR, are presented. The two parameter sets yield significantly different surface morphologies, in agreement with the previously identified energy-governed trends in Ra and Rz. For 8 W−500 mm/s−50kHz, the effective energy delivered VED is relatively low due to the short dwell time and reduced thermal accumulation, resulting in a shallow and comparatively uniform topography with a height range of approximately 0−10μ m. The associated SEM image shows predominantly fine resolidified features, consistent with a regime closer to controlled ablation and limited melting, where the melt layer remains thin and melt expulsion/redeposition is less pronounced. Conversely, for 16 W −300 mm/s−50kHz, which maximizes cumulative energy deposition, enhanced heat accumulation promotes a thicker and more persistent molten layer, leading to more intense recoil-pressure-driven melt ejection and hydrodynamic instabilities. This is reflected in the substantially larger height range (up to ~70μ m ) and in the SEM evidence of abundant resolidified spherical droplets (spatter), indicating thick recast material, and porous/ligament-like structures, indicative of vigorous melt splashing, droplet coalescence, and partial remelting.

Overall, the high-energy condition amplifies both peak formation (recast/spatter) and crater-related valley development, increasing Ra/Rz and shifting functional height descriptors (e.g., Rmc-related levels), whereas the low-energy condition suppresses these melt-driven extremes and produces a more stable and moderate surface texture.

[picture]

Fig. 5. Topography and the respective x 5000 SEM images for the two extreme scenarios of Ra for 8 W−500 mm/s−50kHz and 16 W−300 mm/s−50kHz.

Fig. 5. Topography and the respective x 5000 SEM images for the two extreme scenarios of Ra for 8 W − 500 m m / s − 50 k H z and 16 W − 300 m m / s − 50 k H z .

Conclusion

The original version of this paper is available on https://www.scientific.net/SSP.391.115.pdf

In the current work the correlation between the machining parameters during LE of Incoloy 800 HT and the SR is experimentally studied. Specifically, an L9 Taguchi DoE was implemented for laser LP, LSS and LPF, while SR was evaluated using Ra, Rz, Rsk, and Rmc( 20% ), whilst, the parameter effects were quantified via ANOVA ranking. The observed trends were interpreted through energydriven removal physical based underlying mechanisms, linking thermal accumulation, melt generation, recoil-pressure-driven ejection, and recast/spatter deposition to the resulting height distributions and peak-valley balance. The main conclusions that have been deduced are:

LP and LSS are the dominant factors for Ra, Rz, and Rmc(20%), followed by LPF and then LP•LPF interaction term, confirming that roughness amplitude is primarily controlled by the net energy input and the severity of melt/recast formation.

Increasing LP and LPF generally increases Ra/Rz and elevates Rmc(20%) heights, whereas increasing LSS decreases these metrics due to reduced dwell time and lower thermal accumulation.

Rsk remains positive, indicating peak-dominated surfaces consistent with crater rims, recast ridges, and redeposited droplets typical of LE. The ANOVA defines LSS as the primary driver of Rsk, implying that scan speed most strongly controls whether the surface is dominated by redeposited peaks versus crater-driven valleys.

Topography and SEM observations corroborate the statistical trends, namely, low-energy conditions produce shallower, more uniform textures with limited spatter, while high-energy conditions generate thick recast layers, abundant spherical droplets, and pronounced height ranges. These morphological differences directly explain the concurrent increase in Ra/Rz and the shift of Rmc(20%), with clear implications for tailoring surfaces toward wear-resistant plateaus or lubricant-retentive valley structures.

References

The original version of this paper is available on https://www.scientific.net/SSP.391.115.pdf

N.B. Dahotre, S.P. Harimkar, Laser Fabrication and Machining of Materials, Springer US, Boston, MA, 2008.

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The original version of this eBook is available on https://www.scientific.net/book/non-conventional-processes/978-3-0364-1985-5

Solid State Phenomena, ISSN: 1662-9779, Vol. 391, pp 123-131

doi: 10.4028/p-Ju0c5j

© 2026 The Author(s). Published by Trans Tech Publications Ltd, Switzerland.

Submitted:
2026-01-19

Revised:
2026-03-02

Accepted:
2026-03-03

Online:
2026-04-20

Non-Conventional Method of Evaluation of the Abrasion Resistance of a Polymeric Material Using a Hole Enlargement Process by Abrasion

The original version of this paper is available on https://www.scientific.net/SSP.391.123.pdf

Roxana Gabriela HOBJÂLĂ a , Răzvan MITITELU b , Vasile ERMOLAI c  Ștefan JURESCHI d, loan SURUGIU e, Vlada ȚISARI f, Margareta COTEAȚĂ g, Laurențiu SLĂTINEANU h

"Gheorghe Asachi" Technical University of lași, Department of Machine Manufacturing Technology, Blvd. D. Mangeron, 39 A, 700050 Iaşi, Romania

a  roxana.hobjala@student.t. (*corresponding author), b  nicolae-razvan.mititelu@academic.tuiasi.ro, c  vasile.ermolai@academic.tuiasi.ro, d  stefan.jureschi@student.tuiasi.ro, eioan.surugiu @academic.tuiasi.ro, f  vlada.tisari@student.tuiasi.ro, 9  margareta.coteata@academic.tuiasi.ro, h  laurentiu.slatineanu@academic.tuiasi.ro

Keywords: abrasion resistance, polymer, enlarging, experiment, empirical mathematical model.

Abstract

Some parts made of polymeric materials are used in abrasive environments. Any damage to their integrity by abrasive environments could have a negative effect on the operation of those parts. In the paper, it is proposed to use a process and equipment for evaluating the resistance to abrasive erosion based on the application of a process of enlarging a previously made hole in the polymer material specimen using a conical abrasive tool. The feed movement of the conical abrasive tool takes place under the action of a counterweight of known size. The duration of the process of enlarging by abrasion provides information on the resistance to abrasive erosion of the specimen material. An experimental study of the resistance to abrasive erosion of a polymeric material using the enlarging process with a conical abrasive tool was designed and implemented. The experimental results were processed mathematically, and empirical mathematical models were obtained that highlight the influence of some input factors in the abrasive process on the resistance to abrasive erosion evaluated through the duration of the enlarging process with a conical abrasive tool. The possibility of using enlarging by abrasion as a procedure for evaluating the resistance to abrasive erosion was confirmed.

Introduction

The original version of this paper is available on https://www.scientific.net/SSP.391.123.pdf

The concept of wear can be attributed to the meaning of irreversible modification of the surfaces of friction couples during operation [1]; most frequently, this irreversible modification manifests itself through the removal of material from the surface of the part affected by the wear process [2]. The concept of mass wear refers to how a wear process manifests and is evaluated, and it concerns the amount of material removed from a part as a result of the aforementioned process, developed under predetermined conditions.

Abrasive wear takes into account the action of free abrasive particles located in the space between two parts that are in relative motion with respect to each other, or of abrasive particles present in one of the parts involved in the motion. In industrial practice, abrasive wear is one of the important factors capable of affecting the service life of a part.

The expansion of the use of polymeric materials has led, among other things, to the manufacture of parts intended to operate in abrasive environments. In such situations, the availability of information regarding the resistance of polymeric materials to abrasive erosion is of interest. One of the ways to obtain such information has been based on the design of experimental methods that aim to study the influence, over time, of various factors on the amount of material removed from a part as a result of the development of abrasive erosion processes.

Some researchers have directed their activities toward a better understanding of the resistance of polymeric materials to abrasive erosion. Thus, Zhang et al. proposed the use of abrasive erosion resistance testing equipment that simulates the operating conditions of a hydro cyclone in the oil field
[3]. The experimental research carried out facilitated the identification of explanations regarding the processes that occur during abrasive erosion.

Biswas et al. developed a synthesis regarding the effects of several input factors in the abrasive erosion process, such as the nature of the abrasive material, the impact angle, and the impact velocity on the abrasive wear rate [4]. They also analyzed different models proposed by other researchers to explain how the abrasive erosion wear process is initiated and develops.

The presentation of experimental methods that can be used to evaluate the resistance to abrasive erosion of materials used in hydraulic machinery, as well as the results of experimental tests performed on two categories of steels, was carried out by Crîngaşu et al. [5].

Some methods ensure the development of friction processes in the presence of abrasive particles between bodies having contact surfaces of different shapes and between which relative motion exists. In this regard, the pin-on-disk, pin-on-flat, pin-on-ring, pin-on-cylinder, crossed cylinder, and thrust washer methods can be mentioned [6-10].

The pin-on-cylinder method was used in an experimental investigation of the abrasion resistance of two polymeric composite materials [11], using a device adapted to a universal lathe.

The study of the specialized literature has highlighted the fact that for the evaluation of the resistance to abrasive erosion of various materials, methods that involve measuring the amount of material removed from the samples as a result of their pressing and displacement on abrasive surfaces, under specified experimental conditions and on appropriate equipment, are mainly used. It was considered that the experimental tests carried out by the known methods require specialized equipment and long durations of performance.

In the present paper, the presentation of results from research on the resistance to abrasive erosion of polymeric materials is considered under the conditions of the proposal of a non-traditional method of enlarging holes using a conical abrasive tool. The duration of the enlargement of a hole in a polymer material specimen was considered as an indicator for evaluating the abrasion resistance of the specimen material. It was assumed that, in this way, it would be possible to obtain information regarding the resistance to abrasive erosion of polymeric materials. Accordingly, a preliminary analysis of the conditions under which an abrasive erosion process takes place and of the factors capable of influencing the intensity of such a process was performed. An experimental investigation was planned and carried out in order to obtain an empirical mathematical model corresponding to the stated objective.

Experimental Section

The original version of this paper is available on https://www.scientific.net/SSP.391.123.pdf

One of how the wear process manifests itself is the removal of small amounts of material from the part affected by wear in the form of micro-chips. In principle, if, between the surfaces of parts in relative motion with respect to each other and under the action of one part being pressed against the other, there are particles made of sufficiently hard materials, having sharp edges and a certain orientation with respect to the parts, there is a high probability of initiation and development of microcutting processes. Some of the particles may penetrate the surface layer to a small depth and, being displaced as a result of the relative motion between the two parts involved in the process, will contribute to pressing the material of the part until shear occurs, resulting in the generation of a microchip (Fig. 1a).

If a force F acting perpendicular to the surface on which the abrasive grain presses the material of a part, along a direction coinciding with the direction of motion of the grain, is considered, this force can be decomposed into two components: a component Fa, which presses a material layer of thickness ap, and another component, Fp, which exerts pressure on the surface layer generated as a result of the micro-cutting process. The component Fa will cause the detachment of a chip which, according to machining theory, will have a thickness smaller than the thickness ap of the removed material layer. The component Fp will contribute to a change in the shape of the grains in the surface layer, possibly increasing the hardness of the material in this layer and generating the effect known as work hardening.

Assuming that the surface of the abrasive grain involved in the micro-cutting process forms a clearance angle γ, the magnitude of the Fa component will be given by a relationship of the form:

Fa=Fcosγ(1)

This Fa component must generate a pressure pa greater than the compressive strength σc of the material of the tested part, evaluated in terms of resistance to abrasive erosion, for the initiation and development of a shear process that generates a micro-chip to be possible:

pa>σc(2)

The aspects mentioned above make it possible to conclude that the resistance of a material to abrasive erosion can be evaluated, for example, by determining the amount of material removed from the tested part within a given time interval or by determining the length of the time interval required for a certain amount of material to be removed from the tested part by abrasive erosion; the actual implementation of the abrasion process can take various forms, corresponding, for example, to methods of machining a workpiece by abrasive processes such as grinding, polishing, honing, and superfinishing, all of which involve the use of tools having different shapes and in which the abrasive grains are embedded in a binder, while there are also processes that use free abrasive grains, as in the case of lapping.

The values of a material's resistance to abrasive erosion may be influenced by the following groups of factors or factor categories:

the nature and physical-mechanical properties of the material of the tested specimen; for example, abrasive erosion resistance is considered to be correlated with the material's compressive and shear strength, hardness, etc.;

the strength and physical-mechanical properties of the material of the abrasive grains; the abrasive grain material must possess mechanical properties (compressive and shear strength, hardness) superior to those of the material of the tested part;

the size and geometry of the abrasive grains; if the abrasive grains are larger, there is a higher probability that they will contribute to the generation of micro-chips of larger size than in the case of smaller abrasive grains;

the nature and physical-mechanical properties of the material that embeds the abrasive grains, referred to as the binder in the case of abrasive bodies used to materialize cutting processes; if this embedding material has high mechanical properties, there is a higher probability that larger quantities of material will be removed from the part whose resistance to erosion is being determined;

[picture]

Fig. 1. The action exerted by an abrasive grain in motion and pressed against the surface of the part whose material is being tested for abrasive erosion (a), and a detail from the micro-chip formation area (b).

Fig. 1. The action exerted by an abrasive grain in motion and pressed against the surface of the part whose material is being tested for abrasive erosion (a), and a detail from the micro-chip formation area (b).

the presence and nature, in the contact zones between the abrasive grains and the material of the tested part, of liquids capable of generating cooling-lubrication processes; such liquids may facilitate
the penetration of the sharp edges of the abrasive grains into the material of the tested part, which would imply a decrease in the abrasive erosion resistance of the tested material;

the values characterizing the dynamics of the micro-chip generation process by the abrasive grains; therefore, the relative sliding velocity between the grains and the tested material, as well as the magnitude of the normal force exerted by the abrasive grains on the tested material, are important;

the heat generated by micro-cutting processes; in the case of high relative velocities between the abrasive grains and the material of the tested part, larger amounts of heat may be generated, which could lead to softening of the tested material and a reduction in abrasive erosion resistance; in this respect, the thermal properties of the tested material are also important, as a higher heat transfer into the bulk of the material may affect its plasticity and, consequently, the values of abrasive erosion resistance.

Through the experimental tests, the aim was to identify empirical mathematical models that would highlight the intensity of the influence exerted by certain input factors in the process of testing resistance to abrasive erosion on an output parameter characterizing this process.

As the specimen material, a flexible polyurethane polymer was chosen. The main physicalmechanical properties of this material are [12]: tensile strength Rm≥300kgf/mm2, elongation at break δ≥300%, abrasion loss Δm<40 mm3, volume variation in oil ΔV≤2.2% at 72 h/70∘C, and a shrinkage coefficient of 1.2%. From the polyurethane, a two-step specimen was manufactured with thicknesses of 10 mm and 20 mm . In each of the two steps, holes with a diameter of 4.5 mm were drilled using a drill bit.

The equipment used for the experimental tests was a bench drilling machine, previously employed for evaluating the machinability of metallic materials [13,14]. The drilling machine allows the use of multiple spindle speeds by shifting a belt across pulleys of different diameters. For the objectives of the present research, it was particularly relevant that on the shaft normally used for manually advancing the drill into the workpiece, a pulley had been mounted, around which a cable was attached and wound. At the free end of the cable, a pan was attached to hold weights of varying mass (Fig. 2a). Under these conditions, it became possible to perform drilling or hole-enlarging processes with a feed force of known magnitude (Fig. 2b).

During the experimental tests, to enlarge a hole with an initial diameter of 5.5 mm , a conical abrasive tool (multi-purpose diamond drill bit, Stoolxi, China) with a maximum diameter of 6 mm was used, which was expected to produce holes with diameters equal to or greater than 6 mm after

[picture]

Fig. 2. Enlarging a hole with an initial diameter di=5.5 mm using a conical abrasive tool with a maximum diameter of 6 mm , employing a drilling setup with a constant feed force (a), and, respectively, performing the working stroke for enlarging the hole with the conical abrasive tool (b).

Fig. 2. Enlarging a hole with an initial diameter d i = 5.5 m m using a conical abrasive tool with a maximum diameter of 6 mm , employing a drilling setup with a constant feed force (a), and, respectively, performing the working stroke for enlarging the hole with the conical abrasive tool (b).

enlargement. The abrasive properties of the conical tool were achieved by the presence of diamond grains embedded in a metal coating layer on the conical surface of the tool (Fig. 2b). The conical abrasive tool was located and clamped in a chuck typically used for drill-type tools on the drilling machine.

A full factorial L8 experimental design was chosen, with three independent variables, each at two levels of variation. For this purpose, two values of the rotation speed were used ( vmin=400rev/min, vmax=560rev/min ), the drill feed force ( Fmin=18 N,Fmax=24 N ), and the specimen thickness (which actually corresponds to the length of the hole to be enlarged) ( tmin =10 mm,tmax =15 mm ).

The output parameter selected to evaluate the abrasion resistance of the polymeric material used for the specimens was the time τ required to enlarge the existing hole in the specimen. This time was measured with a stopwatch. It was considered that the three input factors in the testing process (abrasive tool rotation speed, feed force, and specimen thickness) can exert an important influence on the duration of the enlarging process using the conical abrasive tool. At the same time, the duration of the enlarging process could provide information on the abrasion resistance of the specimen material on which the enlargement of a hole by abrasion was performed.

Initially, the drill was placed in contact with the edge of the hole to be enlarged, and the time was recorded from the moment the drill began to penetrate the specimen material until the drill exited freely after the hole-enlarging process.

It was hypothesized that the higher the rotation speed and feed force, the shorter the time required for abrasive enlargement, while an increase in the length of the hole to be enlarged would contribute to a longer enlargement time. These hypotheses take into account that, according to chip formation theory, increasing the rotation speed of the conical abrasive tool improves the flow of removed material, thereby reducing both the force required for the enlargement process and the time needed to complete the hole enlargement, compared to results obtained at lower cutting speeds.

The values of the input factors that were varied during the experimental tests are listed in columns 2−8 of Table 1. This table also includes the coded values of the input factors, using the symbol "-1" for the minimum values and " +1 " for the maximum values of the input factors used in the abrasive erosion tests.

Result and Discussion

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The values of the output parameter used (the time τ required to perform the hole enlargement) were recorded in column 9 of Table 1. These experimental results were processed mathematically, with the requirement to determine first-degree and second-degree polynomial functions, including the use of two artificial intelligence programs for their verification [15, 16].

In this way, the following two empirical mathematical models were obtained, corresponding to the two types of polynomials considered:

τ=63.945−0.08834n−0.05833F+0.38300t,(3)

for which the coefficient of determination has the value R2=0.523 and, respectively:

τ=0.0006+0.1438n+0.0152F+0.0262t+0.0004n2−0.0234nF−0.0132nt+0.3704F2−0.352Ft+0.5632t2(4)

in this case, the value of the coefficient of determination is R2=0.990, the adjusted R2=0.9302, the sum of residuals SSR=7.7618,Fstatistic =16.5514 and p-value =0.1860. Due to the high value of the coefficient of determination R2, it can be considered that the proposed model explains the experimental results well, but the F test does not reject the null hypothesis for α=0.05, which could be justified by the small number (8) of experimental trials.

Table 1. Experimental test conditions and obtained results.

Exp. no | Input factors | Output parameter
Rotation speed | Pressing force | Sample thickness
Coded value | Rotation speed, n [rev/min] | Peripheral speed, v [m/min] | Coded value | Real value, F [N] | Coded value | Real value, t [mm] | Time τ [s]
Column no. 1 | 2 | 3 | 4 | 5 | 6 | 7 | 8 | 9
1 | -1 | 400 | 7.53 | -1 | 18 | -1 | 10 | 21,48
2 | -1 | 400 | 7.53 | -1 | 18 | +1 | 15 | 31,98
3 | -1 | 400 | 7.53 | +1 | 24 | -1 | 10 | 35,67
4 | -1 | 400 | 7.53 | +1 | 24 | +1 | 15 | 39,55
5 | +1 | 560 | 10.55 | -1 | 18 | -1 | 10 | 21,88
6 | +1 | 560 | 10.55 | -1 | 18 | +1 | 15 | 25,77
7 | +1 | 560 | 10.55 | +1 | 24 | -1 | 10 | 17,55
8 | +1 | 560 | 10.55 | +1 | 24 | +1 | 15 | 06,94

The first-degree polynomial empirical mathematical model explains only 52.3% of the variation in the output parameter (the time τ required to enlarge the hole using the conical abrasive tool), but it provides general information regarding the influence exerted by the considered input factors.

It can thus be observed that, in accordance with the initial hypotheses, the time τ for abrasive hole enlargement decreases as the rotation speed n of the conical abrasive tool and the feed force F increase, and, respectively, it increases when the specimen thickness t increases, of course, within the predetermined variation ranges of the input factor values. Analysis of this empirical mathematical model shows that, among the three input factors, the greatest influence is exerted by the specimen

[picture]

Fig. 3. Influence of the rotation speed on the abrasive hole-enlarging time τ(F=18 N,t=15 s).

Fig. 3. Influence of the rotation speed on the abrasive hole-enlarging time τ ( F = 18 N , t = 15 s ) .

thickness t, followed by the feed force F, and lastly by the rotation speed n of the conical abrasive tool. This observation is based on the order of the absolute values of the coefficients assigned to each input factor in the testing process (0.38300>0.05833>0.08834).

[picture]

Fig. 4. Influence of the sample thickness on the abrasive hole-enlarging time t (rotation speed n=400 and 480rev/min, feed force F=18 N ).

Fig. 4. Influence of the sample thickness on the abrasive hole-enlarging time t (rotation speed n = 400 and 480 r e v / m i n , feed force F = 18 N ).

A better fit of the empirical mathematical model to the experimental results was achieved by using a second-degree polynomial function, in which case the coefficient of determination was R2=0.990. This mathematical model also takes into account the interactions between the considered input factors, making it possible to observe that the variation of the output parameter with respect to one input factor is influenced by the values of the other two input factors.

Based on the second-degree polynomial empirical model, the graphical representations in Figures 3 and 4 were developed, showing that, in principle, increasing the rotation speed of the conical abrasive tool and the feed force contribute to a reduction in the time required for the complete enlargement of a previously drilled hole in the specimen. However, the second-degree polynomial model also indicates the need to extend the experimental investigations to more thoroughly verify the existence of interactions between the input factors in the abrasive hole-enlarging process.

It was considered that the determined empirical mathematical models are valid only for the variation intervals established for the values of the input factors in the experimental process. The main limitations of these models come from the relatively small number of experimental tests based on which the empirical mathematical models were developed, from the fact that these tests have so far been carried out only on a single polymer material and for small ranges of variation of the input factor values, and from the low precision with which it was possible to determine the duration of the expansion process.

Summary

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The concept of a material's resistance to abrasive erosion refers to the ability of the material to resist the removal of micro-chips caused by the relative motion and pressing of abrasive grains on the surface of a part or specimen made from that material. Knowledge of a material's resistance to abrasive erosion is important because it allows certain aspects of the material's behavior under conditions involving abrasive wear to be anticipated, situations that sometimes occur in the practical use of various mechanical equipment. A review of the specialized literature highlighted the interest of researchers in the abrasive erosion behavior of polymeric materials. Focusing on the investigation of the abrasive erosion behavior of a polyurethane polymer, an experimental testing scheme was designed involving the enlargement of pre-drilled holes in polyurethane specimens under the action of a constant feed force. It was proposed to use an equipment that would allow determining by a nonconventional method the duration of the enlarging of a hole previously made in the specimen with the help of a drill. For the enlargement, a conical abrasive tool was used. It was proposed to use the duration of the enlarging process as an indicator for evaluating the abrasion resistance of the
polymeric material. The determination of an empirical mathematical model was considered to highlight the direction and intensity of the influence exerted by the tool's rotation speed, the feed force, and the specimen thickness on the duration of the abrasive hole-enlarging process. The resulting empirical mathematical model confirms that, under certain testing conditions, the duration of the abrasive hole-enlarging process decreases as the rotation speed and the magnitude of the constant feed force increase. Some limitations of the validity of the empirical mathematical models may derive in principle from the small number of experimental tests performed for their determination, from the low precision of measuring the duration of the enlarging process using the conical abrasive tool. In the future, the experimental investigations are intended to be extended to study the abrasive erosion behavior of different polymeric materials, to provide a clearer understanding of the existence of interactions between some of the input factors in the abrasive holeenlarging process. It is also intended to compare the results obtained using the proposed method with those obtained using other methods for evaluating resistance to abrasive erosion.

References

The original version of this paper is available on https://www.scientific.net/SSP.391.123.pdf

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E.C. Crîngaşu, C.A. Safta, B. Ghiban, Experimental methods for determining the resistance to abrasive erosion of materials used in the manufacture of hydraulic machines (in Romanian), XVIIth International-Multidisciplinary Conference "Professor Dorin Pavel - the founder of Romanian hydropower", Sebes, 2017,

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I. Hutchings, M. Gee, E. Santner, Friction and Wear. In: Czichos, H., Saito, T., Smith, L. (Eds.) Springer Handbook of Materials Measurement Methods. Springer Handbooks, Springer, Berlin, Heidelberg, 2006.

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S. Bheemappa, H. Gurumurthy, V.V. Badami, P.R. Hegde, 13 - Tribological behavior of polymeric systems in lubricated surfaces or conditions, in: S.C. George, J.T. Haponiuk, S. Thomas, R. Reghunath, P.S. Sarath (eds.), Elsevier Series on Tribology and Surface Engineering, Tribology of Polymers, Polymer Composites, and Polymer Nanocomposites, Elsevier, 2023, pp. 357-399.

R. Hobjâlă, V.Ermolai, R. Mititelu1, A. Hrițuc, E. Crăciun, I. Surugiu, L. Slătineanu1, C. Bisog, Abrasion resistance of some composite polymer materials, Material Forming - ESAFORM 2025, Mater. Res. Proc. 54 (2025) 2112-2120.

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The original version of this eBook is available on https://www.scientific.net/book/non-conventional-processes/978-3-0364-1985-5

Solid State Phenomena, ISSN: 1662-9779, Vol. 391, pp 133-141

doi: 10.4028/p-T5RRvs

© 2026 The Author(s). Published by Trans Tech Publications Ltd, Switzerland.

Submitted:
2026-01-23

Revised:
2026-02-23

Accepted:
2026-02-23

Online:
2026-04-20

Non-Conventional Recycling Process of Hemp/Carbon Hybrid Laminates via Thermo-Mechanical Disassembly

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Luca Boccarusso 1,a*, Denise Bellisario 2, b, Alessandro Canneva 1,c, Dario De Fazio 1, d, Massimo Durante 1,e, Fabrizio Quadrini 2,f, Loredana Santo 2, g

1 Department of Chemical, Materials and Production Engineering, University of Naples "Federico II", P.le Tecchio 80, 80125, Naples, Italy

2 Department of Industrial Engineering, University of Rome Tor Vergata, via del Politecnico 1, Rome 00133, Italy

a* luca.boccarusso@unina.it, b denise.bellisario@uniroma2.it, c alessandro.canneva @unina.it, d dario.defazio@unina.it, e mdurante@unina.it, f fabrizio.quadrini@uniroma2.it, g  Ioredana.santo@uniroma2.it

Keywords: hybrid composites, non-conventional recycling, sustainability, CFRP, hemp.

Abstract

The increasing use of fibre-reinforced composites raises critical issues related to sustainability and end-of-life management, particularly for thermoset-based systems. In this work, a non-conventional thermo-mechanical recycling strategy is proposed for hemp/carbon hybrid laminates, aiming at the recovery and reuse of intact reinforcement plies without destructive fibrematrix separation. Full carbon, full hemp, and two hybrid laminate configurations with different stacking sequences were manufactured, recycled through controlled thermo-mechanical disassembly, and reprocessed into new laminates. The flexural and interlaminar shear behaviour of virgin and recycled materials was investigated to assess the influence of the recycling process on mechanical performance.

Introduction

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Over the last decades, fibre-reinforced composite materials have experienced a steadily increasing adoption across several industrial sectors due to their high strength-to-weight ratio, outstanding design flexibility, and excellent structural performance. In particular, polymer matrix composites reinforced with carbon fibres have progressively replaced metallic materials in high-value applications such as aerospace, automotive, and sports equipment, where weight reduction represents a primary design requirement. However, the growing awareness of environmental issues, together with increasingly stringent regulations on sustainability and end-of-life management, has highlighted the limitations of these solutions, fostering strong interest in the development of more sustainable composite materials capable of reducing environmental impact without significantly compromising mechanical performance [1].

In this context, the use of natural fibres as reinforcement in composite materials represents a particularly promising strategy, owing to their renewability, low energy demand during production, and potential contribution to the reduction of CO2 emissions. Nevertheless, fully "green" composites reinforced exclusively with vegetal fibres still exhibit significant application limitations when high structural performance is required. The intrinsic variability of natural fibre properties, their sensitivity to moisture, and their inferior mechanical characteristics compared to synthetic fibres confine their use mainly to non-structural or semi-structural applications [2-4].

To overcome these limitations, a solution widely explored in the literature is the adoption of hybrid composites, in which natural fibres and high-performance fibres are combined within the same polymer matrix [5, 6]. Hybridisation enables the exploitation of the advantages of different reinforcement types while mitigating their respective drawbacks. In particular, carbon/natural fibre hybrid composites allow the high stiffness and strength of carbon fibres to be combined with the enhanced ductility, energy dissipation capability, and potential damping behaviour of vegetal fibres, resulting in an effective compromise between mechanical performance, cost, and sustainability.

Previous studies [7,8] have demonstrated that, through careful design of the stacking sequence, it is possible to tailor the flexural and shear behaviour as well as the damage tolerance of hemp/carbon hybrid laminates, in some cases approaching the performance of full carbon laminates while benefiting from a partially bio-based content.

Beyond the sustainability of raw materials, another critical issue affecting composite materials, especially those based on thermosetting matrices, concerns end-of-life management and the limited recyclability of such systems, which remains a significant barrier to circularity due to the irreversible cross-linked structure of thermosets that prevents remelting or conventional reprocessing (e.g., CFRP and GFRP) [9]. Historically, the end-of-life problem of composites was considered marginal owing to their relatively limited early diffusion; however, the exponential growth in the use of carbon and glass fibre-reinforced polymers is leading to a significant increase in composite waste volumes, while traditional disposal routes such as landfilling and incineration are increasingly unsustainable and restricted by regulation, prompting the need for cost-effective recycling strategies [10]. Current recycling technologies for thermoset composites can be broadly classified into three main categories: mechanical, thermal, and chemical recycling [11, 12]. Mechanical recycling, based on cutting and grinding operations, produces fibre-rich and resin-rich fractions that are mainly reused as fillers in new materials, but it results in a severe reduction of fibre length and integrity. Thermal processes, such as pyrolysis, allow the recovery of relatively clean fibres but require high temperatures, entail high energy consumption, and can cause significant degradation of fibre mechanical properties. Similarly, chemical processes, such as solvolysis, enable the recovery of fibres with properties closer to virgin ones, but they present critical issues related to solvent use, processing time, and by-product management. Overall, these strategies are generally characterised by a "down-cycling" approach, in which the value of the recovered material is lower than that of the original composite. Moreover, in previous studies, the authors introduced and validated a novel, non-conventional thermo-mechanical recycling strategy specifically designed for full carbon fibre-reinforced thermoset composites, demonstrating the feasibility of separating and recovering intact reinforcement plies without resorting to destructive fibre-matrix separation techniques [13].

Consequently, while the effectiveness of non-conventional recycling approaches has been mainly demonstrated for full carbon laminates, their application to hybrid composites containing natural fibres remains a largely unexplored scenario. The presence of vegetal fibres, characterised by a different thermal and mechanical response compared to synthetic fibres, introduces additional challenges related both to the disassembly process and to the preservation of mechanical performance after recycling. At the same time, the possibility of recovering and reusing plies containing natural fibres opens new perspectives for extending the service life of hybrid composites with reduced environmental impact.

Considering these factors, the present study aims to extend the current knowledge on hemp/carbon hybrid composites by introducing a non-conventional recycling methodology based on a thermomechanical disassembly approach as a tool for the recovery and reuse of reinforcements [14]. To this purpose, full hemp, full carbon, and different hybrid laminate configurations were manufactured, and their recyclability and the reuse of the recovered individual plies were investigated. The mechanical properties of recycled laminates were then compared with those of the corresponding virgin materials, with particular focus on flexural and shear behaviour. The objective of this work is to assess the extent to which the thermo-mechanical recycling process affects the key mechanical properties of these systems and to evaluate the potential of recycled hybrid laminates as sustainable structural solutions capable of combining adequate mechanical performance with reduced environmental impact.

Materials and Methods

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Materials and composite manufacturing.

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The composite laminates investigated in this study were manufactured using carbon fibre and hemp fibre fabrics as reinforcements combined with a SX10 epoxy matrix (supplied by Mates Srl). The carbon fibre reinforcement consisted of a woven fabric with an areal density of 200 g/m2 (supplied
by Toray international Srl), while the woven hemp fabric with an areal density of 160 g/m2 (supplied by Maeko Srl). Before the impregnation, the hemp fabrics were dried in a climatic chamber for 12 h at 60∘C and a relative humidity of 20% to reduce moisture content and improve impregnation quality.

The laminates ( 250 mm×250 mm ) were manufactured by combining hand lay-up and vacuumassisted press forming, following a consolidated procedure already adopted in previous studies [6]. In detail, the impregnated fabrics were manually stacked inside the mould according to the desired stacking sequence and subsequently cured under vacuum and pressure.

Four types of virgin laminates were produced and used both as reference materials and as starting structures for the recycling process: full hemp laminates (H), full carbon laminates (C), and two hemp/carbon hybrid configurations, referred to as S and U, characterized by different ply distributions through the laminate thickness. All laminates, shown in Fig.1, were manufactured with the same total number of plies (i.e., 15) in order to allow a meaningful comparison of mechanical properties. All stacking sequences are labelled in Table 1.

Table 1. Main laminate characteristics.

Label | Reinforcement | Stacking sequence | Thickness [mm] | Fibre volume fraction [%]
C | Carbon | Carbon stacking sequence [C15] | 2.87 | 58.1
S | Carbon/Hemp | Carbon/Hemp stacking sequence [C6H3C6] | 3.00 | 55.1
U | Carbon/Hemp | Carbon/Hemp stacking sequence [CHC5HC5HC] | 3.00 | 55.1
H | Hemp | Hemp stacking sequence [H15] | 3.70 | 43.2

[picture]

Fig. 1. Virgin laminates (a) and their cross section (b).

Fig. 1. Virgin laminates (a) and their cross section (b).

Recycling process and Manufacturing of recycled laminates.

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The recycling of the virgin laminates was performed using a thermo-mechanical disassembly approach aimed at recovering individual plies while preserving fibre integrity and maintaining the matrix bonded to the fibres. Unlike conventional recycling technologies, this method does not rely on matrix degradation or destructive fibre separation but exploits the increase in ductility of the thermosetting matrix when heated above its glass transition temperature.

Specifically, the virgin laminates were heated in a muffle furnace to 200∘C for 5 min . The selected temperature was high enough to promote epoxy matrix softening and facilitate thermo-mechanical disassembly, while the limited exposure time prevented the onset of significant thermal degradation. Subsequently, the laminates were subjected to controlled mechanical deformation using a manual roll-bending machine with a roll diameter of 38.5 mm , as shown in Fig.2. The cyclic application of bending deformation promoted progressive interlaminar delamination, enabling the separation of individual plies or small ply stacks without inducing macroscopic fibre damage. When necessary, the heating and bending steps were repeated until complete laminate disassembly was achieved. Particular attention was paid to the adjustment of the roll gap and the number of bending passes in order to maximize delamination while preserving the structural integrity of the recovered plies. At the end of the process, the individual plies appeared mechanically intact, flexible, and suitable for reuse in subsequent restacking and consolidation steps.

To indirectly assess the residual resin content and to evaluate possible matrix redistribution or loss induced by the recycling process, each recovered ply was individually weighed. The mass of recycled plies was then compared with that of the corresponding virgin plies, allowing an estimation of resin retention after thermo-mechanical disassembly.

The recovered plies were reused to manufacture recycled laminates, preserving the same reinforcement configurations as the original virgin laminates (H, C, S, and U) without introducing additional reinforcing or matrix materials.

The production of recycled laminates was performed by following the same procedure adopted for the manufacturing of virgin laminates.

[picture]

Fig. 2. Thermo-mechanical disassembly of a composite laminates: (a) roll-bending system used, (b) the start of delamination and the separation between two hemp plies (c) and among carbon fibre ply and hemp ply.

Fig. 2. Thermo-mechanical disassembly of a composite laminates: (a) roll-bending system used, (b) the start of delamination and the separation between two hemp plies (c) and among carbon fibre ply and hemp ply.

Mechanical characterization.

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The mechanical characterisation of both virgin and recycled laminates was carried out by means of three-point bending tests and interlaminar shear strength (ILSS) tests, in order to evaluate the effect of the recycling process on the main structural properties of the materials.

Three-point bending tests were performed in accordance with ASTM D790 using a universal MTS Exceed E43 universal testing machine equipped with a 50 kN load cell. Specimen dimensions, span length, and crosshead speed were selected as a function of laminate thickness and reinforcement type, with the aim of minimising shear effects and ensuring a reliable evaluation of flexural behaviour. Three specimens for each sample configuration were tested with a span-to-depth ratio of 16:1 was adopted.

The flexural stress σf was then evaluated according to Eq. 1, where P is the load in (N),L is the support span in (mm), b and d are respectively the width and the thickness of the specimen in (mm)

σf=32PLbd2(1)

Interlaminar shear strength tests were conducted according to ASTM D2344 using a short-beam configuration to emphasise shear stresses between adjacent plies. For each configuration, three specimens were tested using a span-to-depth ratio of 5:1. The shear stress τ was evaluated according to Eq. 2:

τ=0.75Pbd(2)

Results and Discussion

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Virgin laminates.

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The flexural behaviour of the virgin laminates is strongly affected by the reinforcement type and stacking sequence. The carbon laminate (C) exhibits the typical response of CFRP systems under three-point bending, characterised by a linear elastic region with high stiffness followed by an abrupt catastrophic failure. From the experimental results, laminate C shows a flexural modulus of approximately 55 GPa and a flexural strength close to 680 MPa , confirming the high stiffness and strength but limited damage tolerance typical of carbon fibre composites. In contrast, the hemp laminate (H) displays a markedly different behaviour. The stress-strain curve is characterised by a much lower slope and a progressive, ductile response without a clear catastrophic failure. The flexural modulus of H is approximately 7 GPa , while the flexural strength reaches about 130 MPa , corresponding to reductions of nearly −80% in stiffness and −90% in strength compared to the carbon laminate. However, the hemp laminate reaches significantly higher strain levels, exceeding 0.03 mm/mm, highlighting its superior deformation capability and ductile nature, in agreement with the behaviour reported by Pinto et al [6].

The symmetric hybrid configuration (S), with hemp plies positioned at the laminate mid-plane, exhibits a flexural response that combines the advantages of both reinforcements. The initial linear portion of the stress-strain curve is characterised by a flexural modulus comparable to that of the pure carbon laminate. This confirms that placing hemp layers close to the neutral axis does not significantly affect the elastic stiffness of the laminate. The flexural strength of configuration S is around 610 MPa , corresponding to a moderate reduction (around −10% ) compared to laminate C . At the same time, the maximum strain increases to about 0.022 mm/mm, indicating a more progressive damage evolution and enhanced deformation capability with respect to the carbon reference. The uniform hybrid configuration (U), characterised by a symmetric distribution of hemp layers throughout the thickness, shows a further modification of the flexural response. The flexural modulus decreases to approximately 43 GPa , reflecting the greater influence of hemp layers located away from the neutral axis. The flexural strength is reduced to about 460 MPa , while the maximum strain at failure remains close to 0.015 mm/mm.

Compared to configuration S, laminate U exhibits lower stiffness and strength but still maintains significantly better mechanical performance than the hemp laminate, confirming the effectiveness of hybridisation in balancing performance and ductility. In Fig. 3, typical stress-strain curves for each sample configuration are plotted.

[picture]

Fig. 3. Typical flexural stress-strain curves for each sample configuration under inspection.

Fig. 3. Typical flexural stress-strain curves for each sample configuration under inspection.

The interlaminar shear behaviour of the virgin laminates further highlights the effect of hybridisation on damage mechanisms, as shown by the stress-displacement curves of Fig.4.

The carbon laminate (C) exhibits the highest ILSS value, reaching approximately 42 MPa , with a sharp peak in the shear stress-displacement curve followed by an abrupt load drop, indicative of a brittle interlaminar failure.

The hemp laminate (H) shows a significantly lower ILSS, with a maximum shear stress of approximately 15 MPa , corresponding to a reduction of about −64% compared to the carbon laminate. The shear response of H is characterised by a smooth and progressive curve, associated with combined shear and flexural deformation rather than a purely interlaminar failure.

Both hybrid configurations (S and U) exhibit an ILSS of approximately 37 MPa , slightly lower than that of the carbon laminate ( −12% ). The shear stress-displacement curve closely resembles that of laminate C , indicating that the shear load is still mainly sustained by the carbon rich regions of the laminate, while the hemp layers do not significantly compromise interlaminar strength.

[picture]

Fig. 4. Typical ILSS stress-displacement curves for each sample configuration under inspection.

Fig. 4. Typical ILSS stress-displacement curves for each sample configuration under inspection.

Recycling process.

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The thermo-mechanical recycling process led to a partial disassembly of the virgin laminates, as the complete separation into individual plies was not achievable for all configurations. In particular, the full hemp laminate (H) was the only configuration that allowed the complete disassembly into all 15 individual plies. This behaviour is directly related to the lower interlaminar shear strength characterising the hemp laminate, which facilitates delamination under thermo-mechanical loading.

Conversely, for carbon and hybrid laminates ( C,S and U ), the disassembly process resulted in a combination of single plies and small blocks consisting of 2−3 plies still bonded together, due to their higher interlaminar cohesion. For the manufacturing of recycled laminates, only fully disassembled single plies were used, while multi-ply blocks were discarded, in order to ensure a controlled and repeatable restacking process.

Each recovered ply retained a certain amount of residual epoxy resin on its surface, quantified in terms of mass fraction. Specifically, for the recycled carbon plies, the mass fraction of residual resin was 33%, whereas for the recycled hemp plies it was 54%. Details of the recycled laminates are reported in Table 2.The recycled laminates were manufactured using the same processing route and curing parameters adopted for the virgin laminates, with additional epoxy resin introduced to ensure proper impregnation and bonding. As a consequence, the recycled laminates exhibited an increased thickness and a corresponding reduction in fibre volume fraction, as reported in Table 2.

Table 2. Main recycled laminate characteristics.

Label
Thickness [mm]

Fibre volume

fraction [%]

rC | 4.0 | 42
rS | 4.4 | 38
rU | 4.3 | 38
rH | 4.9 | 33

Properties of recycled laminates.

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The flexural stress-strain curves of the recycled laminates (Fig. 5) reveal a significant modification of the mechanical response compared to the virgin counterparts, mainly associated with changes in material used and fibre volume fraction.

The recycled carbon laminates ( rC ) exhibits a lower stiffness compared to the virgin carbon laminate, with a flexural modulus of approximately 33 GPa , corresponding to a reduction of about 40%. The flexural strength is also reduced, reaching values close to 275 MPa , compared to approximately 680 MPa for the virgin laminate. Despite this reduction, the stress-strain curve of rC shows a markedly more progressive post-peak behaviour, characterised by multiple load drops and an extended deformation regime, indicating enhanced damage tolerance and energy dissipation capability induced by the recycling process.

The recycled hemp laminate (rH) maintains a very compliant and ductile flexural response. The flexural modulus remains low, around 6 GPa , while the flexural strength reaches approximately 108 MPa . Compared to the virgin hemp laminate, rH exhibits very similar behaviour, with a truly minimal loss of properties; this aspect distinguishes it from all other types of recycled composites. The recycled symmetric hybrid laminate (rS) shows a flexural modulus of approximately 31 GPa , close to that of rC, while the flexural strength reaches about 255 MPa . Although these values are significantly lower than those of the virgin S laminate, the stress-strain curve displays a progressive damage evolution with multiple stress drops and a long post-peak plateau, indicating effective stress redistribution and enhanced deformation capability.

The recycled uniform hybrid laminate (rU) exhibits a flexural modulus of approximately 25 GPa and a flexural strength close to 245 MPa . Compared to rS, rU shows a further reduction in stiffness, consistent with the more uniform distribution of hemp layers through the thickness. Nevertheless, the flexural response remains stable and progressive, with extended strain levels exceeding 0.05 mm/mm,
highlighting the high damage tolerance of the recycled hybrid configuration. It was therefore observed that, unlike the virgin specimens, although a reduction in mechanical properties is evident, the differences between the various hybrid configurations and the full carbon laminate are much less pronounced.

The ILSS results further confirm the effectiveness of the recycling strategy. The recycled carbon (rC), symmetric hybrid (rS), and uniform hybrid (rU) laminates do not show significant variations compared to their virgin counterparts. All these configurations exhibit ILSS values of approximately 30 MPa , indicating only a moderate reduction with respect to the virgin laminates.

This behaviour suggests that the interlaminar adhesion in the recycled laminates is mainly governed by the new epoxy resin introduced during reprocessing, which effectively restores the bonding between adjacent plies despite the presence of residual resin on the recycled fibres.

A different trend is observed for the recycled hemp laminate (rH), which exhibits an ILSS of approximately 17 MPa , slightly higher than that of the virgin hemp laminate. This unexpected improvement can be attributed to the rough surface morphology of the recycled hemp plies, which enhances mechanical interlocking with the newly added epoxy resin. The increased surface roughness promotes a more effective mechanical anchoring mechanism, leading to improved interlaminar shear resistance.

[picture]

Fig. 5. Typical flexural stress-strain curves for each sample recycled sample configuration.

Fig. 5. Typical flexural stress-strain curves for each sample recycled sample configuration.

Conclusion

The original version of this paper is available on https://www.scientific.net/SSP.391.133.pdf

The proposed thermo-mechanical recycling approach represents a significant conceptual shift compared to conventional composite recycling technologies. Rather than relying on destructive fibrematrix separation or matrix degradation, this strategy exploits the increased ductility of thermosetting matrices at temperatures above the glass transition to promote controlled laminate delamination. Through the application of suitably calibrated mechanical deformation, composite laminates can be disassembled into individual or partially bonded plies while preserving fibre integrity and fibrematrix adhesion. The experimental results demonstrate that this approach can be successfully applied not only to full carbon laminates but also to hemp/carbon hybrid composites, and full hemp composites enabling the recovery of reusable reinforcements with limited damage. Although recycled laminates exhibit a reduction in stiffness and strength mainly associated with changes in fibre volume fraction and additional matrix content, they retain stable, progressive mechanical responses and satisfactory interlaminar shear properties. Notably, after recycling, the mechanical differences between hybrid configurations and full carbon laminates become less pronounced, highlighting the effectiveness of hybridisation in combination with the proposed recycling route.

Overall, the thermo-mechanical disassembly strategy allows the recovery of high-added value structural elements while reducing energy consumption and secondary waste generation, opening new perspectives for up-cycling thermoset-based hybrid composites and improving their overall sustainability.

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The original version of this paper is available on https://www.scientific.net/SSP.391.133.pdf

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Keywords Index

A

Abrasion Resistance Non-Conventional Method of Evaluation of the Abrasion Resistance of a Polymeric Material Using a Hole Enlargement Process by Abrasion

Adhesive Bonding Concurrent Effect of Laser Texturing and Resin Pre-Coating on the Performance of CFRP Single Lap Joints

Am-built Holes Preliminary Comparative Assessment of Hole Quality and Geometrical Tolerances in AM-Built and CNC-Drilled PLA Components

B

Bonding Joint Enabling the Reuse of Single-Lap Adhesive-Bonded Joints through Laser Cleaning and Texturing

C

Cemented Carbides Modulated Pulsed Electrochemical Machining of Tungsten Carbide-Cobalt: An Experimental Study According to DIN SPEC 91399

Cfrp Non-Conventional Recycling Process of Hemp/Carbon Hybrid Laminates via Thermo-Mechanical Disassembly

Cnc Drilling Preliminary Comparative Assessment of Hole Quality and Geometrical Tolerances in AM-Built and CNC-Drilled PLA Components

Coating Study of the Effect of Foam Substrate Material on Hydrophobic Coating Efficiency

Composite Structures Origami-Inspired Design for Manufacture of Composite Parts

Corrosion Study of the Effect of Foam Substrate Material on Hydrophobic Coating Efficiency

Corrosion Protection Electrodeposited Metallic Coatings as a Manufacturing Route to Enhance the Durability of NdFeB Magnets

Critical Raw Materials Mining Critical Raw Materials from Waste Printed Circuit Boards by Thermo-Mechanical Disassembly

D

Detection Method Classification and Detection Method for Discharges in Electrical Discharge Machining through Unsupervised Learning

Die Sinking Edm Classification and Detection Method for Discharges in Electrical Discharge Machining through Unsupervised Learning

Direct Molding Unconventional Recycling of Aeronautical Ground Carbon Fiber Composite Wastes

Discharges Classification and Detection Method for Discharges in Electrical Discharge Machining through Unsupervised Learning

E

Electrodeposition Electrodeposited Metallic Coatings as a Manufacturing Route to Enhance the Durability of NdFeB Magnets

Empirical Mathematical Model Non-Conventional Method of Evaluation of the Abrasion Resistance of a Polymeric Material Using a Hole Enlargement Process by Abrasion

Enlarging Non-Conventional Method of Evaluation of the Abrasion Resistance of a Polymeric Material Using a Hole Enlargement Process by Abrasion

Experiment Non-Conventional Method of Evaluation of the Abrasion Resistance of a Polymeric Material Using a Hole Enlargement Process by Abrasion

F

Forming Origami-Inspired Design for Manufacture of Composite Parts

Fused Deposition Modelling (fdm) Preliminary Comparative Assessment of Hole Quality and Geometrical Tolerances in AM-Built and CNC-Drilled PLA Components

G

Ground Crfp Waste Unconventional Recycling of Aeronautical Ground Carbon Fiber Composite Wastes

H

Hemp Non-Conventional Recycling Process of Hemp/Carbon Hybrid Laminates via Thermo-Mechanical Disassembly

Hole Quality Preliminary Comparative Assessment of Hole Quality and Geometrical Tolerances in AM-Built and CNC-Drilled PLA Components

Hybrid Composites Non-Conventional Recycling Process of Hemp/Carbon Hybrid Laminates via Thermo-Mechanical Disassembly

Hydrophobicity Study of the Effect of Foam Substrate Material on Hydrophobic Coating Efficiency

I

Incremental Forming Mechanical Performance of Incrementally Formed Flax Fiber-Reinforced Polypropylene Caps

L

Laser Ablation Enabling the Reuse of Single-Lap Adhesive-Bonded Joints through Laser Cleaning and Texturing

Laser Engraving Surface Roughness Control in Incoloy 800HT via Laser Engraving Parameters

M

Machine Learning Classification and Detection Method for Discharges in Electrical Discharge Machining through Unsupervised Learning

Mechanical Performance Mechanical Performance of Incrementally Formed Flax Fiber-Reinforced Polypropylene Caps

Metal Foams Study of the Effect of Foam Substrate Material on Hydrophobic Coating Efficiency

Metallic Coatings Electrodeposited Metallic Coatings as a Manufacturing Route to Enhance the Durability of NdFeB Magnets

Modulated Pulsed Electrochemical Machining Modulated Pulsed Electrochemical Machining of Tungsten Carbide-Cobalt: An Experimental Study According to DIN SPEC 91399

N

Natural Fiber-reinforced Plastics Mechanical Performance of Incrementally Formed Flax Fiber-Reinforced Polypropylene Caps

Nickel Surface Roughness Control in Incoloy 800HT via Laser Engraving Parameters

Non-conventional Recycling Non-Conventional Recycling Process of Hemp/Carbon Hybrid Laminates via Thermo-Mechanical Disassembly

Non-crimp Fabrics Origami-Inspired Design for Manufacture of Composite Parts

O

Oil-separation Efficiency Study of the Effect of Foam Substrate Material on Hydrophobic Coating Efficiency

Origami Origami-Inspired Design for Manufacture of Composite Parts

P

Polymer Non-Conventional Method of Evaluation of the Abrasion Resistance of a Polymeric Material Using a Hole Enlargement Process by Abrasion

R

Rare-earth Magnets Electrodeposited Metallic Coatings as a Manufacturing Route to Enhance the Durability of NdFeB Magnets

Rebonding Enabling the Reuse of Single-Lap Adhesive-Bonded Joints through Laser Cleaning and Texturing

Recycling Unconventional Recycling of Aeronautical Ground Carbon Fiber Composite Wastes

S

Single Lap Joints (slj) Concurrent Effect of Laser Texturing and Resin Pre-Coating on the Performance of CFRP Single Lap Joints

Stability Study of the Effect of Foam Substrate Material on Hydrophobic Coating Efficiency

Surface Roughness Surface Roughness Control in Incoloy 800HT via Laser Engraving Parameters

Surface Treatments Concurrent Effect of Laser Texturing and Resin Pre-Coating on the Performance of CFRP Single Lap Joints

Sustainability Non-Conventional Recycling Process of Hemp/Carbon Hybrid Laminates via Thermo-Mechanical Disassembly

T

Taguchi Doe Surface Roughness Control in Incoloy 800HT via Laser Engraving Parameters

Thermo-mechanical Disassembly Mining Critical Raw Materials from Waste Printed Circuit Boards by Thermo-Mechanical Disassembly

Thermo-mechanical Properties Unconventional Recycling of Aeronautical Ground Carbon Fiber Composite Wastes

U

Urban Mining Mining Critical Raw Materials from Waste Printed Circuit Boards by Thermo-Mechanical Disassembly

Authors Index

B

Baxevani, A. Study of the Effect of Foam Substrate Material on Hydrophobic Coating Efficiency

Bederitzky, O. Classification and Detection Method for Discharges in Electrical Discharge Machining through Unsupervised Learning

Bella, G. Preliminary Comparative Assessment of Hole Quality and Geometrical Tolerances in AM-Built and CNC-Drilled PLA Components

Bellisario, D. Mining Critical Raw Materials from Waste Printed Circuit Boards by Thermo-Mechanical Disassembly; Unconventional Recycling of Aeronautical Ground Carbon Fiber Composite Wastes; Non-Conventional Recycling Process of Hemp/Carbon Hybrid Laminates via Thermo-Mechanical Disassembly

Benvenuto, M. Enabling the Reuse of Single-Lap Adhesive-Bonded Joints through Laser Cleaning and Texturing

Boccarusso, L. Mechanical Performance of Incrementally Formed Flax Fiber-Reinforced Polypropylene Caps; Preliminary Comparative Assessment of Hole Quality and Geometrical Tolerances in AM-Built and CNC-Drilled PLA Components; Non-Conventional Recycling Process of Hemp/Carbon Hybrid Laminates via Thermo-Mechanical Disassembly

C

Canneva, A. Non-Conventional Recycling Process of Hemp/Carbon Hybrid Laminates via Thermo-Mechanical Disassembly

Chairi, M. Preliminary Comparative Assessment of Hole Quality and Geometrical Tolerances in AM-Built and CNC-Drilled PLA Components

Clauß, P. Modulated Pulsed Electrochemical Machining of Tungsten Carbide-Cobalt: An Experimental Study According to DIN SPEC 91399

Coteata, M. Non-Conventional Method of Evaluation of the Abrasion Resistance of a Polymeric Material Using a Hole Enlargement Process by Abrasion

D

Damm, P. Modulated Pulsed Electrochemical Machining of Tungsten Carbide-Cobalt: An Experimental Study According to DIN SPEC 91399; Classification and Detection Method for Discharges in Electrical Discharge Machining through Unsupervised Learning

Durante, M. Mechanical Performance of Incrementally Formed Flax Fiber-Reinforced Polypropylene Caps; Preliminary Comparative Assessment of Hole Quality and Geometrical Tolerances in AM-Built and CNC-Drilled PLA Components; Non-Conventional Recycling Process of Hemp/Carbon Hybrid Laminates via Thermo-Mechanical Disassembly

E

Ermolai, V. Non-Conventional Method of Evaluation of the Abrasion Resistance of a Polymeric Material Using a Hole Enlargement Process by Abrasion

F

Fatyga, L. Surface Roughness Control in Incoloy 800HT via Laser Engraving Parameters

Fazio, D. Mechanical Performance of Incrementally Formed Flax Fiber-Reinforced Polypropylene Caps; Non-Conventional Recycling Process of Hemp/Carbon Hybrid Laminates via Thermo-Mechanical Disassembly

Formisano, A. Mechanical Performance of Incrementally Formed Flax Fiber-Reinforced Polypropylene Caps

H

Hackert-Oschätzchen, M. Modulated Pulsed Electrochemical Machining of Tungsten Carbide-Cobalt: An Experimental Study According to DIN SPEC 91399; Classification and Detection Method for Discharges in Electrical Discharge Machining through Unsupervised Learning

Hobjâlă, R. Non-Conventional Method of Evaluation of the Abrasion Resistance of a Polymeric Material Using a Hole Enlargement Process by Abrasion

I

Iorio, L. Mining Critical Raw Materials from Waste Printed Circuit Boards by Thermo-Mechanical Disassembly

Irace, G. Mechanical Performance of Incrementally Formed Flax Fiber-Reinforced Polypropylene Caps

J

Jungblut, G. Modulated Pulsed Electrochemical Machining of Tungsten Carbide-Cobalt: An Experimental Study According to DIN SPEC 91399

Jureschi, Ş. Non-Conventional Method of Evaluation of the Abrasion Resistance of a Polymeric Material Using a Hole Enlargement Process by Abrasion

K

Karmiris-Obratański, P. Surface Roughness Control in Incoloy 800HT via Laser Engraving Parameters

Kratz, J. Origami-Inspired Design for Manufacture of Composite Parts

L

Lamprou, E. Study of the Effect of Foam Substrate Material on Hydrophobic Coating Efficiency

Langella, A. Preliminary Comparative Assessment of Hole Quality and Geometrical Tolerances in AM-Built and CNC-Drilled PLA Components

Lerez, C. Classification and Detection Method for Discharges in Electrical Discharge Machining through Unsupervised Learning

M

Maes, V. Origami-Inspired Design for Manufacture of Composite Parts

Mandolfino, C. Enabling the Reuse of Single-Lap Adhesive-Bonded Joints through Laser Cleaning and Texturing

Markopoulos, A. Surface Roughness Control in Incoloy 800HT via Laser Engraving Parameters

Meichsner, G. Modulated Pulsed Electrochemical Machining of Tungsten Carbide-Cobalt: An Experimental Study According to DIN SPEC 91399; Classification and Detection Method for Discharges in Electrical Discharge Machining through Unsupervised Learning

Michailidis, N. Study of the Effect of Foam Substrate Material on Hydrophobic Coating Efficiency

Mititelu, R. Non-Conventional Method of Evaluation of the Abrasion Resistance of a Polymeric Material Using a Hole Enlargement Process by Abrasion

Moffa, G. Concurrent Effect of Laser Texturing and Resin Pre-Coating on the Performance of CFRP Single Lap Joints

N

Napolitano, F. Unconventional Recycling of Aeronautical Ground Carbon Fiber Composite Wastes

Navaratnarajah, S. Origami-Inspired Design for Manufacture of Composite Parts

Nikolidakis, E. Surface Roughness Control in Incoloy 800HT via Laser Engraving Parameters

Noqra, D. Mining Critical Raw Materials from Waste Printed Circuit Boards by Thermo-Mechanical Disassembly

O

Orazi, L. Enabling the Reuse of Single-Lap Adhesive-Bonded Joints through Laser Cleaning and Texturing

P

Panico, M. Preliminary Comparative Assessment of Hole Quality and Geometrical Tolerances in AM-Built and CNC-Drilled PLA Components

Papazoglou, E. Surface Roughness Control in Incoloy 800HT via Laser Engraving Parameters

Parodo, G. Concurrent Effect of Laser Texturing and Resin Pre-Coating on the Performance of CFRP Single Lap Joints

Patrizii, G. Mining Critical Raw Materials from Waste Printed Circuit Boards by Thermo-Mechanical Disassembly

Perna, A. Electrodeposited Metallic Coatings as a Manufacturing Route to Enhance the Durability of NdFeB Magnets

Plänitz, P. Classification and Detection Method for Discharges in Electrical Discharge Machining through Unsupervised Learning

Proietti, A. Mining Critical Raw Materials from Waste Printed Circuit Boards by Thermo-Mechanical Disassembly

Q

Quadrini, F. Mining Critical Raw Materials from Waste Printed Circuit Boards by Thermo-Mechanical Disassembly; Unconventional Recycling of Aeronautical Ground Carbon Fiber Composite Wastes; Non-Conventional Recycling Process of Hemp/Carbon Hybrid Laminates via Thermo-Mechanical Disassembly

R

Reggiani, B. Enabling the Reuse of Single-Lap Adhesive-Bonded Joints through Laser Cleaning and Texturing

Riefer, A. Classification and Detection Method for Discharges in Electrical Discharge Machining through Unsupervised Learning

Russo, P. Unconventional Recycling of Aeronautical Ground Carbon Fiber Composite Wastes

S

Santo, L. Mining Critical Raw Materials from Waste Printed Circuit Boards by Thermo-Mechanical Disassembly; Non-Conventional Recycling Process of Hemp/Carbon Hybrid Laminates via Thermo-Mechanical Disassembly

Schenk, M. Origami-Inspired Design for Manufacture of Composite Parts

Scherillo, F. Electrodeposited Metallic Coatings as a Manufacturing Route to Enhance the Durability of NdFeB Magnets

Schäfer, M. Modulated Pulsed Electrochemical Machining of Tungsten Carbide-Cobalt: An Experimental Study According to DIN SPEC 91399

Siciliani, V. Enabling the Reuse of Single-Lap Adhesive-Bonded Joints through Laser Cleaning and Texturing

Slătineanu, L. Non-Conventional Method of Evaluation of the Abrasion Resistance of a Polymeric Material Using a Hole Enlargement Process by Abrasion

Sorrentino, L. Concurrent Effect of Laser Texturing and Resin Pre-Coating on the Performance of CFRP Single Lap Joints

Stergioudi, F. Study of the Effect of Foam Substrate Material on Hydrophobic Coating Efficiency

Surugiu, I. Non-Conventional Method of Evaluation of the Abrasion Resistance of a Polymeric Material Using a Hole Enlargement Process by Abrasion

T

Thielecke, A. Modulated Pulsed Electrochemical Machining of Tungsten Carbide-Cobalt: An Experimental Study According to DIN SPEC 91399

Turchetta, S. Concurrent Effect of Laser Texturing and Resin Pre-Coating on the Performance of CFRP Single Lap Joints

V

Viscusi, A. Electrodeposited Metallic Coatings as a Manufacturing Route to Enhance the Durability of NdFeB Magnets

Vogel, J. Classification and Detection Method for Discharges in Electrical Discharge Machining through Unsupervised Learning

Ţ

Ţisari, V. Non-Conventional Method of Evaluation of the Abrasion Resistance of a Polymeric Material Using a Hole Enlargement Process by Abrasion