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International Journal of Material Forming

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Open Access 01-11-2023 | Original Research

Integral forming of continuous CFRP sandwich sheet by additive manufacturing

Authors: Kazusa Nishi, Yuji Sato, Jun Yanagimoto

Published in: International Journal of Material Forming | Issue 6/2023

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Abstract

Sandwich sheets comprising continuous carbon fiber reinforced plastics (CFRP) are applied mainly in the aerospace industry due to their light weight and high rigidity. However, sandwich sheets require separate formation and bonding of the face sheets and core, resulting in high labor costs and early fracture due to delamination of the adhesive layer. The purpose of this study is to overcome these problems by integrating sandwich sheet using additive manufacturing. The mechanical properties of the integrally formed sandwich sheets were compared with those of adhesively formed sandwich sheets using a three-point bending test. The strain distribution was captured by digital image correlation (DIC) during the test. Additionally, the geometric design parameters of a core with superior mechanical properties were investigated. The test results showed that the integrally formed specimens exhibited superior properties compared to those of the adhesively formed specimens. It was also observed that the larger the width angle of the corrugated core, the better the mechanical properties.

Publisher’s Note

Springer Nature remains neutral with regard to jurisdictional claims in published maps and institutional affiliations.

Introduction

In recent years, sandwich sheets have been implemented in different applications to simultaneously reduce weight and increase rigidity [1]. Sandwich sheets are particularly effective in aircraft and satellites, because weight reduction directly affects energy efficiency and payload [2, 3]. Additionally, the material should be selected as light as possible [4]; and thus, continuous carbon fiber reinforced plastics (CFRP), which have excellent specific strength and stiffness, are considered suitable materials for sandwich sheets [5‐11]. In fact, sandwich sheets comprising CFRP face sheets and an aluminum or Nomex honeycomb core are widely used in the aerospace industry [12]. Flying cars are currently being developed as next-generation mobility, and are expected to be used in delivery, transportation, entertainment, and other applications. They can also be an effective means of relieving traffic congestion in urban areas. To popularize flying cars, it is important to manufacture flying car structures that reduce the weight in a cost-effective manner. One promising solution to meet these demands is the use of CFRP sandwich sheets [13].

However, sandwich sheets require two face sheets and a core adhesively bonded separately, which increases production costs. Furthermore, the adhesive layer causes face sheet delamination, which often leads to failure. The strength of the adhesive layer depends on the properties of the adhesive, and face sheet delamination indicates that the mechanical properties of the sandwich sheet have not been fully demonstrated.

Xiong et al. developed a lightweight sandwich sheet by forming a CFRP truss core via hot press forming and bonding with two CFRP face sheets [14]. However, the results of bending tests revealed that the adhesive area between the core and face sheets was insufficient, resulting in delamination [15]. To solve this problem, Wu et al. installed an aluminum alloy (6061) connector between the truss core and face sheets, and improved the mechanical properties by adding an interlocking mechanism between the connector and truss core [16]. Compared with the experimental results obtained by Xiong et al., this structure suppressed the delamination of the adhesive layer. However, the formation of sandwich sheets using connectors is a costly production process involving several steps.

Zhang et al. developed a new sandwich sheet using steel for the face sheet and CFRP for the core, and proposed a truncated dome core considering the bending strength, rigidity, and mass producibility [17‐19]. Sandwich sheets with truncated dome cores have excellent specific stiffness and formability, and can be mass-produced by hot stamping. Although the truncated dome core exhibited outstanding properties, it was fabricated empirically based on a conventional core structure without analytical optimization. Shibuya et al. optimized the core structure and verified the superior specific stiffness through bending tests and finite element analyses [20, 21]. In studies by both Zhang et al. and Shibuya et al., formability evaluations indicated that buckling of the face sheet or delamination of the adhesive layer were the main failure modes. However, neither failure mode fully exploits the properties of the sandwich sheets. From previous studies, it can be concluded that the delamination of the adhesive layer is a crucial problem for sandwich sheets.

The purpose of this study is to improve the mechanical properties of CFRP sandwich sheets by eliminating the adhesive layer via integral formation of sandwich sheets. The elimination of the adhesive interface is expected to prevent early failure due to the delamination of the adhesive layer. A 3D printer manufactured by Markforged, MarkTwo™, which is capable of printing continuous CFRP, was used because additive manufacturing is suitable for integral forming. MarkTwo™ is a novel 3D printer capable of printing continuous CFRP; it has a lower cost and shorter production time than that of conventional CFRP molding methods [22‐24].

One of the advantages of forming sandwich sheets by additive manufacturing is its high design flexibility, where it is expected that the mechanical properties can be improved by optimizing the core shape. However, in this study, we focused on the differences between integral and adhesive forming. Therefore, only a simple corrugated shape is examined as the core structure. Integrally formed sandwich sheets and adhesively formed sandwich sheets were manufactured using additive manufacturing, and subjected to a three-point bending test to compare their mechanical properties. During the test, the strain distribution on the surface of the specimens was measured by digital image correlation (DIC), and the influence of the difference in the forming method was discussed. Moreover, as a step toward optimizing the core shape of the integrally formed sandwich sheets, specimens with different geometrical design parameters were manufactured and the optimal parameters were investigated.

Specimen configuration

The core of the sandwich sheet comprised a repeating arrangement of unit cells with identical geometries, and in this study, the unit cell of the corrugated core was trapezoidal. The geometrical design parameters were set as shown in Fig. 1, where the z axis represents the direction of lamination. The geometrical design parameters of the unit cell were as follows: $l$ is the length of inclined strut, $d$ is the length of horizontal strut, ${t}_{f}$ is the thickness of face sheets, ${t}_{c}$ is the thickness of core, ${h}_{c}$ is the height of core, $\phi$ is the inclined angle of strut, and $\theta$ is the angle of core in the width direction. Here, $\theta$ is set so as to suppress the buckling of the face sheet and improve the mechanical properties by aligning the cores along both the width and longitudinal directions. The specimens were uniformly formed with $d=8\text{m}\text{m}, {t}_{f}={t}_{c}=2.4\text{m}\text{m}$ and $\phi =55^\circ$. The shear stiffness of the corrugated core is maximized at [25]. Then, $l$ and $\theta$ were set to $l=5\text{m}\text{m}$and $3\text{m}\text{m},$ and $\theta =0^\circ , 30^\circ ,$ and $45^\circ$, respectively. Six different specimens were prepared for this study. From Eq. (1), the relative density of the cores was calculated to be $\rho =0.37$ for the $l=5\text{m}\text{m}$ specimen and $\rho =0.49$ for the $l=3\text{m}\text{m}$ specimen.

$$\begin{array}{c}\rho =\frac{{t}_{c}}{{h}_{c}}=\frac{{t}_{c}}{{t}_{c}+l\text{sin}\varphi}\end{array}$$

(1)

Figure 2 shows an integrally formed sandwich sheet. Onyx, comprising short carbon fibers and PA6, was used for the face sheets, whereas a continuous carbon fiber (CCF) filament and Onyx were used for the core. Table 1 lists the material properties of Onyx and CCF. The carbon fiber volume fraction of the CCF was 35% as per manufacturer specifications.

Table 1

Material properties of Onyx and CCF filament [26]

Material property	Onyx	CCF
Tensile modulus [GPa]	2.4	60
Tensile strength [MPa]	37	800
Flexural modulus [GPa]	3.0	51
Flexural strength [MPa]	71	540
Density [g/cm³]	1.2	1.4

The $l=5\text{mm}, \theta =0^\circ$ specimens were manufactured not only by integral forming, but also by adhesive bonding. The other geometric design parameters were the same as those described above. The face sheet and core were printed separately and adhered using MOS-8 (Konishi Co., Ltd.), a two-component silicone polymer and epoxy adhesive with working temperature ranging from $-40^\circ\mathrm C$ to $120^\circ\mathrm C$. Adhesive was cured at room temperature with a curing time of 12 hours.

After specimen preparation, a random pattern of speckles was sprayed onto the surface of the specimens to track the strain distribution using DIC.

Three-point bending test

Figure 3 shows a schematic of the three-point bending test. The jig was made of an aluminum alloy with an indenter radius of 5 mm and distance of 50 mm between the lower indenters. The crosshead speed was set at 10 mm/min. The test was terminated when the stroke reached a depth of 20 mm. A multi stage compression testing machine (Thermecmaster, Fuji Electronic Industrial Co., Ltd.) was used. Six different integrally formed specimens with different geometric design parameters were tested 10 times each, and the adhesively formed $l=5\text{m}\text{m}, \theta =0^\circ$ specimens were tested three times. The load and stroke were measured during the tests, and the maximum bending load and stroke at the maximum bending load point were compared for each type of specimen. To evaluate the effect of the geometric design of the core on the mechanical properties of the integrally formed specimens, t-tests at a $5\%$ significance level were conducted with the null hypothesis that the core geometry does not affect the maximum bending load and stroke at the maximum bending load point.

The strain distribution was captured by DIC during a three-point bending test. The strain distributions of the integrally formed and adhesively formed specimens were compared.

Advantage of integral forming

First, the integrally formed and adhesively formed specimens were compared with $\theta =0^\circ$. Table 2 presents the three-point bending test results of the integrally formed specimens with $l=5\text{m}\text{m}$ and $l=3\text{m}\text{m}$ and the adhesively formed specimens with $l=5\text{m}\text{m}$. Figure 4 shows a typical load-stroke diagram. The test results showed that in the $\theta =0^\circ$ geometry, integrally formed specimen have higher maximum bending load than adhesively formed specimen.

Table 2

Result of three-point bending tests (integrally formed specimen vs. adhesively formed specimen)

	Maximum bending load ${F}_{max}$ [kN]	Stroke at maximum load point ${U}_{max}$ [mm]
$l=3\text{m}\text{m}$	2.36	4.31
$l=5\text{m}\text{m}$	2.28	4.65
Adhesive ($l=5\text{m}\text{m}$)	1.55	6.03

Figure 5 shows the strain distribution obtained using DIC. The integrally formed specimen exhibited a typical strain distribution during bending tests, with compressive deformation on the upper side of the specimen and tensile deformation on the lower side. The adhesively formed specimen exhibited a similar strain distribution at the beginning of the test. However, as the test progressed, a localized deformation appeared in the adhesive layer, as shown in Fig. 5. As the test progressed, the strain in the adhesive layer increased, and delamination occurred at approximately $3\text{m}\text{m}$ of the stroke to induce specimen fracture. This is attributed to the low modulus of ductility of the adhesive layer compared to those of the face sheets and core. The stress singularity, where the adhesive layer becomes an interface, induces discontinuous of stress state. Therefore, the fracture occurred earlier, and a difference in the maximum bending load was observed.

Figure 6 shows the appearance of the specimens after the three-point bending test. The adhesively formed specimen clearly showed fracture owing to delamination of the adhesive layer. In the integrally formed specimens, the face sheets and cores fractured simultaneously. Further investigations, including finite element method (FEM) analyses, are required to determine the initiation point and fracture details.

Effect of core geometries manufactured by integral forming

The effects of the different core geometries on the mechanical properties of the integrally formed specimens were compared. Figure 7 shows a typical load-stroke diagram when the parameters $l$ and $\theta$ were changed. Table 3 shows the maximum bending load for different values of $l$ and $\theta$. Table 4 shows the stroke at the maximum bending load point, and Tables 5 and 6 show the p-values obtained using t-tests to verify the effects of $l$ and $\theta$, respectively. Table 5 shows that for all $\theta$, the p-values are greater than 0.05 when comparing $l=5\text{m}\text{m}$ vs. $l=3\text{m}\text{m}$. Therefore, there was no significant difference in the maximum bending load or stroke at the maximum bending point in a range of $3\text{m}\text{m}$ to $5\text{m}\text{m}$. Further investigations on the effect of the parameter $l$ will be continued through experiments and FEM analysis because it is possible to have some impact on mechanical properties seeing that $l$can determine the thickness of the specimen. Table 6 shows that for both $l=5\text{m}\text{m}$ and $3\text{m}\text{m}$, the p-value is above 0.05 for $\theta =0^\circ$ vs. $\theta =30^\circ$, and below 0.05 for $\theta =30^\circ$ vs. $\theta =45^\circ$ and $\theta =45^\circ$ vs. $\theta =0^\circ$. Therefore, it can be said statistically that the specimens with $\theta =45^\circ$ have superior maximum bending load compared to the specimens with $\theta =0^\circ$ and $\theta =30^\circ$. As mentioned in the section on the specimen configuration, this can be attributed to the fact that the cores are aligned not only in the width direction, but also in the longitudinal direction, which improves the mechanical properties. Since the face sheet thickness was set to 2.4 mm, which is relatively high, buckling of the face sheet did not occur in any of the specimens.

Table 3

Maximum bending load${F}_{max}$[kN]

Angle of core$\theta$[deg]	$l=5\text{m}\text{m}$	$l=3\text{m}\text{m}$
0	2.28 (0.33)	2.36 (0.28)
30	2.49 (0.31)	2.38 (0.23)
45	3.13 (0.57)	2.85 (0.15)

The values in the parenthesis show standard deviation

Table 4

Stroke at maximum bending load point${U}_{max}$[mm]

Angle of core$\theta$[deg]	$l=5\text{m}\text{m}$	$l=3\text{m}\text{m}$
0	4.65 (0.26)	4.31 (0.50)
30	4.58 (0.72)	4.35 (0.44)
45	4.63 (0.65)	4.51 (0.46)

The values in the parenthesis show standard deviation

Table 5

The p-value calculated by the t-test (compared parameter: $l$)

Two samples for t-test ($l=5\text{m}\text{m} \text{v}\text{s}. l=3\text{m}\text{m}$)	Maximum bending load	Stroke at maximum bending load point
0	$5.49\times {10}^{-1}$	$9.04\times {10}^{-2}$
30	$4.87\times {10}^{-1}$	$8.25\times {10}^{-1}$
45	$2.29\times {10}^{-1}$	$7.82\times {10}^{-1}$

Table 6

The p-value calculated by the t-test (compared parameter: $\theta$)

Two samples for t-test	Maximum bending load		Stroke at maximum bending load point
	$l=5\text{m}\text{m}$	$l=3\text{m}\text{m}$	$l=5\text{m}\text{m}$	$l=3\text{m}\text{m}$
0 deg vs. 30 deg	$1.87\times {10}^{-1}$	$9.24\times {10}^{-1}$	$7.52\times {10}^{-1}$	$8.73\times {10}^{-1}$
30 deg vs. 45 deg	$8.60\times {10}^{-3}$	$6.75\times {10}^{-3}$	$6.40\times {10}^{-1}$	$4.72\times {10}^{-1}$
45 deg vs. 0 deg	$1.16\times {10}^{-3}$	$3.63\times {10}^{-3}$	$3.62\times {10}^{-1}$	$4.14\times {10}^{-1}$

In the three- point bending test, the contact between the jig and the specimen affects the results. Furthermore, the specimens are thicker, the effect of shear deformation in the three-point bending test is greater. Therefore, the four-point bending test is considered effective for further investigation in order to rule out those error factors.

Summary

In this study, sandwich sheets were manufactured using additive manufacturing to eliminate delamination of the adhesive layer, which often causes early fracture of the sandwich sheets. The integrally formed sandwich sheet was subjected to three-point bending tests to investigate its mechanical properties. In addition, strain measurements using DIC, and a t-test were conducted. The results were compared with those of an adhesively formed specimen. The following conclusions were drawn.

(1)

The adhesively formed specimens fractured at the adhesive layer by delamination, whereas the integrally formed specimens fractured at the core and face sheets. This is consistent with the results of the strain distribution measurements using DIC.

(2)

Integrally formed specimens have superior mechanical properties compared to adhesively formed specimens, because integrally formed specimens avoid early fracture due to delamination of the adhesive layer.

(3)

In integrally formed specimens, the $\theta =45^\circ$ specimen had the highest maximum bending load, which attributes to the core alignment.

In this study, only a simple corrugated structure was used for the core, and no optimization was performed. However, one of the major advantages of additive manufacturing is its high design flexibility; thus, appropriate optimization is effective. Therefore, optimization of the core structure to maximize the stiffness, strength, and formability is necessary in the future. In addition, the details of the fracture of the integrally formed specimens need to be further investigated, and the stress state and fracture should be analyzed using FEM.

Declarations

Competing interests

The authors declare that we have no conflict of interest.

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Yang DY, Bambach M, Cao J, Yang DY, Bambach M, Cao J, Duflou JR, Groche P, Kuboki T, Sterzing A, Tekkaya AE, Lee CW (2018) Flexibility in metal forming. CIRP Ann 67(2):743–765. https://doi.org/10.1016/j.cirp.2018.05.004CrossRef

Kim BJ, Lee DG (2010) Development of a satellite structure with the sandwich T-joint. Compos Struct 92(2):460–468. https://doi.org/10.1016/j.compstruct.2009.08.030CrossRef

Zenkert D (1997) The handbook of sandwich construction. Engineering Materials Advisory Services, Worcestershire

Zhao T, Yang J, Chen J, Guan S (2022) Review of carbon fiber-reinforced sandwich structures. Polym Polym Compos 30:1–17. https://doi.org/10.1177/09673911221098729CrossRef

Haldar A, Guan ZW, Cantwell WJ, Wang QY (2018) The compressive properties of sandwich structures based on an egg-box core design. Compos Part B Eng 144:143–152. https://doi.org/10.1016/j.compositesb.2018.03.007CrossRef

Fan HL, Meng FH, Yang W (2006) Mechanical behaviors and bending effects of carbon fiber reinforced lattice materials. Arch Appl Mech 75(10–12):635–647. https://doi.org/10.1007/s00419-006-0032-xCrossRefMATH

George T, Deshpande VS, Sharp K, Wadley HNG (2014) Hybrid core carbon fiber composite sandwich panels: fabrication and mechanical response. Compos Struct 108:696–710. https://doi.org/10.1016/j.compstruct.2013.10.002CrossRef

George T, Deshpande VS, Wadley HNG (2013) Mechanical response of carbon fiber composite sandwich panels with pyramidal truss cores. Compos Part A Appl Sci Manuf 47:31–40. https://doi.org/10.1016/j.compositesa.2012.11.011CrossRef

Russell BP, Deshpande VS, Wadley HNG (2008) Quasistatic deformation and failure modes of composite square honeycombs. J Mech Mater Struct 3:1315–1340. https://doi.org/10.2140/jomms.2008.3.1315CrossRef

10.

Hu Y, Li W, An X, Fan H (2016) Fabrication and mechanical behaviors of corrugated lattice truss composite sandwich panels. Compos Sci Technol 125:114–122. https://doi.org/10.1016/j.compscitech.2016.02.003CrossRef

11.

Sebaey TA, Mahdi E (2017) Crushing behavior of a unit cell of CFRP lattice core for sandwich structures’ application. Thin-Walled Struct 116:91–95. https://doi.org/10.1016/j.tws.2017.03.016CrossRef

12.

Vasiliev VV, Barynin VA, Razin AF (2012) Anisogrid composite lattice structures - development and aerospace applications. Compos Struct 94(3):1117–1127CrossRef

13.

Saito Y, Endo M, Matsuda Y, Sugiyama N, Watanabe M, Sugahara N, Yamamoto K (1996) A Conceptual study of an advanced VTOL transport aircraft. Technical Report of the National Aerospace Laboratory 1292:1–11

14.

Xiong J, Ma L, Wu L, Xiong J, Ma Li, Wu L, Wang B, Vaziri A (2010) Fabrication and crushing behavior of low density carbon fiber composite pyramidal truss structures. Compos Struct 92(11):2695–2702. https://doi.org/10.1016/j.compstruct.2010.03.010CrossRef

15.

Xiong J, Ma L, Pan S, Xiong J, Ma L, Pan S, Wu L, Papadopoulos J, Vaziri A (2012) Shear and bending performance of carbon fiber composite sandwich panels with pyramidal truss cores. Acta Mater 60(4):1455–1466. https://doi.org/10.1016/j.actamat.2011.11.028CrossRef

16.

Wu Q, Ma L, Wu L, Xiong J (2016) A novel strengthening method for carbon fiber composite lattice truss structures. Compos Struct 153:585–592. https://doi.org/10.1016/j.compstruct.2016.06.060CrossRef

17.

Zhang J, Taylor T, Kizaki T, Yanagimoto J (2020) Bendable metal-based composite sheets with a truncated dome core made of carbon fibre reinforced thermoplastics. Compos Struct 236:111918. https://doi.org/10.1016/j.compstruct.2020.111918CrossRef

18.

Zhang J, Taylor T, Yanagimoto J (2021) Mechanical properties and cold and warm forming characteristics of sandwich sheets with a three-dimensional CFRTP core. Compos Struct 269:114048. https://doi.org/10.1016/j.compstruct.2021.114048CrossRef

19.

Zhang J, Yanagimoto J (2019) Design and fabrication of formable CFRTP core sandwich sheets. CIRP Ann 68(1):281–284. https://doi.org/10.1016/j.cirp.2019.04.060CrossRef

20.

Shibuya Y, Zhang J, Sato Y, Yanagimoto J (2022) Enhancement of mechanical property and formability of CFRP core sandwich sheets by additive manufacturing process-induced material and structural anisotropies. J Mater Process Technol 310:117778. https://doi.org/10.1016/j.jmatprotec.2022.117778CrossRef

21.

Shibuya Y, Zhang J, Sato Y, Yanagimoto J (2022) Evaluation of the mechanical properties and deformability of metal-based composite sheets made of thin stainless-steel sheets and carbon fiber reinforced plastics. Int J Mater Form 15(4):1–13. https://doi.org/10.1007/s12289-022-01661-zCrossRef

22.

Justo J, Távara L, García-Guzmán L, París F (2018) Characterization of 3D printed long fibre reinforced composites. Compos Struct 185:537–548. https://doi.org/10.1016/j.compstruct.2017.11.052CrossRef

23.

Matsuzaki R, Ueda M, Namiki M, Matsuzaki R, Ueda M, Namiki M, Jeong T-K, Asahara H, Horiguchi K, Nakamura T, Todoroki A, Hirano Y (2016) Three-dimensional printing of continuous-fiber composites by in-nozzle impregnation. Sci Rep 6(1):1–7. https://doi.org/10.1038/srep23058CrossRef

24.

Cheng P, Peng Y, Li S, Cheng P, Peng Y, Li S, Rao Y, Le Duigou A, Wang K, Ahzi S (2023) 3D printed continuous fiber reinforced composite lightweight structures: a review and outlook. Compos Part B Eng 250:110450. https://doi.org/10.1016/j.compositesb.2022.110450CrossRef

25.

Lu TJ, Hutchinson JW, Evans AG (2001) Optimal design of a flexural actuator. J Mech Phys Solids 49(9):2071–2093. https://doi.org/10.1016/S0022-5096(01)00024-2CrossRefMATH

26.

Zhang J, Ding S, Yanagimoto J (2022) Bending properties of sandwich sheets with metallic face sheets and additively manufactured 3D CFRP lattice cores. J Mater Process Technol 300:117437. https://doi.org/10.1016/j.jmatprotec.2021.117437CrossRef

Title: Integral forming of continuous CFRP sandwich sheet by additive manufacturing
Authors: Kazusa Nishi
Yuji Sato
Jun Yanagimoto
Publication date: 01-11-2023
Publisher: Springer Paris
Published in: International Journal of Material Forming / Issue 6/2023
Print ISSN: 1960-6206
Electronic ISSN: 1960-6214
DOI: https://doi.org/10.1007/s12289-023-01788-7

Two samples for t-test (\(l=5\text{m}\text{m} \text{v}\text{s}. l=3\text{m}\text{m}\))	Maximum bending load	Stroke at maximum bending load point
0	\(5.49\times {10}^{-1}\)	\(9.04\times {10}^{-2}\)
30	\(4.87\times {10}^{-1}\)	\(8.25\times {10}^{-1}\)
45	\(2.29\times {10}^{-1}\)	\(7.82\times {10}^{-1}\)

Two samples for t-test	Maximum bending load		Stroke at maximum bending load point
	\(l=5\text{m}\text{m}\)	\(l=3\text{m}\text{m}\)	\(l=5\text{m}\text{m}\)	\(l=3\text{m}\text{m}\)
0 deg vs. 30 deg	\(1.87\times {10}^{-1}\)	\(9.24\times {10}^{-1}\)	\(7.52\times {10}^{-1}\)	\(8.73\times {10}^{-1}\)
30 deg vs. 45 deg	\(8.60\times {10}^{-3}\)	\(6.75\times {10}^{-3}\)	\(6.40\times {10}^{-1}\)	\(4.72\times {10}^{-1}\)
45 deg vs. 0 deg	\(1.16\times {10}^{-3}\)	\(3.63\times {10}^{-3}\)	\(3.62\times {10}^{-1}\)	\(4.14\times {10}^{-1}\)

Springer Professional

Integral forming of continuous CFRP sandwich sheet by additive manufacturing

Abstract

Publisher’s Note

Introduction

Specimen configuration

Three-point bending test

Advantage of integral forming

Effect of core geometries manufactured by integral forming

Summary

Declarations

Competing interests

Publisher’s Note

Premium Partners

Springer Professional

Abstract

Publisher’s Note

Introduction

Specimen configuration

Three-point bending test

Advantage of integral forming

Effect of core geometries manufactured by integral forming

Summary

Declarations

Competing interests

Publisher’s Note

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