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In situ residual stress analysis in a phenolic resin and copper composite material during curing
Polymer 182 (2019) 121857
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In situ residual stress analysis in a phenolic resin and copper composite material during curing Atsushi Izumi a, *, Takeshi Kakara a, Midori Wakabayashi Otsuki a, Yasuyuki Shudo a, Tomoyuki Koganezawa b, Mitsuhiro Shibayama c, ** a b c
Corporate Engineering Center, Sumitomo Bakelite Co., Ltd, 2100 Takayanagi, Fujieda, Shizuoka, 426-0041, Japan Japan Synchrotron Radiation Research Institute, 1-1-1 Kouto, Sayo, Hyogo, 679-5198, Japan Neutron Science Laboratory, The Institute for Solid State Physics, The University of Tokyo, 5-1-5 Kashiwanoha, Kashiwa, Chiba, 277-8581, Japan
A R T I C L E I N F O
A B S T R A C T
Keywords: Phenolic resins Curing process Residual stress
Herein, in situ analysis of the residual stress in a phenolic resin and copper foil composite material during curing and thermal-cycle testing was performed via time-resolved X-ray diffraction measurements. The semicured resin–copper composite exhibited a large compressive stress in copper before the curing process. This indicates that the adhesive interface between resin and copper was first formed when the resin melted in the molding process and the magnitude of the thermal contraction of the resin was larger than that of copper in the subse quent cooling to room temperature after the molding. The difference in the magnitude was caused by the dif ference in the coefficients of thermal expansion between resin and copper. This compressive stress decreased as the temperature was increased to curing temperature of 180 � C. As the curing proceeded, the compressive stress in the copper again increased because of cure shrinkage of the resin. When the cured sample was reheated to the curing temperature, the compressive stress in the copper at 180 � C was relaxed. This thermal-annealing-induced stress relaxation suggests that cross-linking reactions during the curing process caused structural strains near the interface between resin and copper, where phenolic resins are two-dimensionally constrained by the copper foil, and that the cross-link strains were relaxed via macroscopic deformation induced by thermal contraction and expansion of the resin.
1. Introduction Phenolic resin composites comprising resin and copper are broadly used in electronic devices such as copper-clad laminates and commu tators because of their excellent physical properties, which include electrical insulator behavior, good heat resistance, and good solvent resistance. One of the important requirements for such composite ma terials is reliable adhesion at the resin–copper interface because delamination causes not only the severe deterioration of physical properties but also fatal defects in the composite products [1]. Typical manufacturing process of composite products comprises molding and curing processes. A composite comprising a semicured phenolic resin and copper is first prepared via laminating or insert molding of an uncured resin together with copper in brief heat treat ment. An additional heat treatment is then performed for several hours as the curing process to drive cross-linking reactions of the phenolic
resins [1]. In this process, thermal stress is generated at the interface between resin and copper because of a difference in the magnitude of thermal expansion and contraction of the materials. This difference is caused not only by the difference in the coefficient of thermal expansion (CTE) between resin and copper but also by the cross-linking reactions that accompany the resin’s cure shrinkage. When this thermal stress is not relaxed, it remains in the material as residual stress, which is an internally balanced stress that exists in the absence of an external force. The residual stress at the interface causes undesirable warpage and interfacial peeling of the composite products. Because the generation of residual stress is inevitable in composites comprising thermosetting resins and other materials, material and process designs for reducing the undesirable thermal stress during the manufacturing process are needed to improve the adhesion reliability at the interface between them. Typically, in industry, the adhesion reliability is evaluated as the num ber of thermal-shock cycles at which the interfacial delamination does
* Corresponding author. ** Corresponding author. E-mail addresses: [email protected] (A. Izumi), [email protected] (M. Shibayama). https://doi.org/10.1016/j.polymer.2019.121857 Received 3 July 2019; Received in revised form 24 September 2019; Accepted 1 October 2019 Available online 1 October 2019 0032-3861/© 2019 Elsevier Ltd. All rights reserved.
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not occur. However, this method is just an evaluation of whether or not the delamination proceeds, which does not provide information about mechanisms that caused delamination or increased adhesion reliability. Therefore, understanding the residual stress generation mechanism in the thermal curing and thermal-cycle testing processes is important [2–13]. Numerous residual stress analysis techniques are available, including methods based on X-ray diffraction (XRD), Raman spectroscopy, ultra sonic acoustics, and strain gauges [14–19]. Among them, the XRD method, which can evaluate lattice strain, is particularly effective for analyzing residual stress in crystalline materials because the lattice spacing is affected by the degree of residual stress and the relation be tween stress directions and lattice planes. When tensile stress is applied to the crystal, the lattice spacing with the lattice plane normal and parallel to the stress axis increases and decreases, respectively, under the relation of the Poisson’s ratio. When compressive stress is applied, the lattice spacing exhibits the opposite behavior. XRD methods for residual stress analysis are divided into cos α methods and sin2 Ψ methods, where α represents the central angle of the Debye ring and Ψ represents the angle between the normal direction of the sample surface and the normal direction of the crystal lattice plane. The cos α methods evaluate an entire Debye ring, and the stress can be calculated from the Debye ring distortion as a function of α. The sin2 Ψ methods evaluate a segment of the Debye ring, and the stress can be calculated from the diffraction-peak angle as a function of Ψ . The sin2 Ψ methods are further divided into iso-inclination and side-inclination methods on the basis of the measurement direction of the diffraction angle. Thus, a wide range of diffraction methods can be selected on the basis of the specimen shape, the source of the incident beam, and the instrumental dif fractometer–detector configuration. In this paper, in situ residual stress analysis in a phenolic resin and copper foil composite material during curing and thermal-cycle testing was performed via a time-resolved XRD measurement using the sin2 Ψ method. The residual stress at the resin–copper interface in the com posite material was calculated as stress in copper via evaluation of the lattice strain of Cu(331) plane. We believe that the proposed study represents the first report of in situ residual stress analysis at the inter face between a phenolic resin and copper during the curing and thermalcycle testing processes.
methylol groups. In their curing, the formation of methylene bridges proceeds with emittions of water and formaldehyde. Herein, the HMTA-cured novolac resin was used as a model resin because the curing system has been intensely investigated in our research project of eluci dating cross-link inhomogeneity in phenoic resins; however, resole-curing system is preferred for the electronic devices in industry because HMTA decomposes to ammonia during curing, thereby causing the corrosion of copper in the device [1]. 2.2. Residual stress analysis XRD measurements were performed at the BL19B2 beamline at the SPring-8 synchrotron radiation facility (Hyogo, Japan). The diffraction profile of the Cu(331) plane was measured using a six-axis diffractom eter (Huber Diffraktionstechnik GmbH, Germany) equipped with a DHS1100 domed hotstage (Anton Paar GmbH, Austria) and a Pilatus 300K two-dimensional X-ray detector (DECTRIS Ltd., Switzerland). The sam ple curing was performed in the air inside a Kapton dome. The energy of the incident X-rays was 8.0 keV, and the longitudinal and horizontal slit sizes for the incident beam were 1 and 6 mm, respectively. As a residual stress measurement method, the sin2 Ψ method was used in conjunction with the iso-inclination method, where a set of Ψ was chosen as 0.0, 20.7, 30.0, 37.8, and 45.0� , which correspond to sin2 Ψ of 0.000, 0.125, 0.250, 0.376, and 0.500, respectively. The X-ray exposure time at each Ψ was 10 s. In the curing process, the temperature of the sample stage was increased from 45 � C to 180 � C, maintained at 180 � C for 6 h, and then cooled to 45 � C. Subsequently, two additional heating and cooling pro cesses between 45 � C and 180 � C were performed as thermal-cycle tests. In the heating and cooling processes, the stress analysis was performed at 15 � C intervals. The heating and cooling rates were �10 � C/min and, after the temperature was increased or decreased by 15 � C, the tem perature was kept constant during sample alignment and diffraction measurements, corresponding to an average heating and cooling rate of approximately �2.3 � C/min. During the 180 � C isothermal treatment, stress analysis was performed at 7 min intervals. Here the sample alignment denotes the direct-beam-half-cut alignment, which corrects the vertical position of the resin–copper interface with respect to the incident beam. The alignment was performed before each sin2 Ψ mea surement because the sample thickness changed as a result of both thermal expansion–contraction and cure shrinkage of the resin. The residual stress (σ) in the copper was calculated by # � �" E 1 π ∂ð2θ* Þ � ; (1) σ¼ ⋅ ⋅ ⋅ 2ð1 þ νÞ tan θ*0 180 ∂ sin2 ψ
2. Experimental 2.1. Materials The resin and copper composite material used in the evaluation had a disk-shaped three-layered structure with the phenolic resin sandwiched between copper foils. A semicured specimen with a diameter and thickness of 24 and 1.5 mm, respectively, was prepared via compression molding at 175 � C for 3 min under a pressure of 10 MPa, corresponding to the composite comprising a semicured phenolic resin and copper, as described in Introduction section. For the resin layer, a 1.0/0.12 (wt/wt) mixture of a random-novolac resin with a number-average molecular weight of 910 g/mol (Sumitomo Bakelite Co., Ltd., Japan) and hexa methylenetetramine (HMTA) was used, wherein HMTA is a typical curing agent for novolac resins. As a copper foil, a surface-roughened electrodeposited copper foil with a thickness of 12 μm (Mitsui Mining and Smelting Co., Ltd., Japan) was used. A specimen of the semicured resin without the copper foil was prepared in a similar manner for temperature-modulated differential scanning calorimetry (TMDSC) analysis. Herein, there are mainly two types of phenolic resins, namely novolac and resole resins. The novolac resins are typically cured with HMTA and curing reaction start with formations of benzoxazine and benzylamine structures, followed by formations of methylene, amide, amine, and imine bridges with liberating ammonia [20]. On the con trary, the resole resins can be cured itself because they contain reactive
where E and ν denote Young’s modulus and Poisson’s ratio, respectively; θ0 and 2θ* denote the diffraction angle of the Cu(331) reflection of the specimen in the non-strained state and the peak top angle of the observed diffraction from Cu(331), respectively. Herein, the positive and negative values of σ represent tensile and compressive stresses, respectively. The values E and ν of the copper foil used in this study were 65.0 GPa and 0.343, respectively. The value θ0 at 25 � C for X-ray energy of 8.0 keV was determined to be 138.04� , which corresponds to a lattice spacing of 0.083 nm assuming that the lattice in the copper foil used in this study was in the non-strained state. The value θ0 at different temperatures was corrected using the copper’s CTE of 16.8 ppm/K. The value ∂(2θ*)/ ∂(sin2 Ψ ) was calculated from the slope of the linear regression of the 2θ*–sin2 Ψ diagram, and the standard error of the slope was used to estimate the error of the residual stress. The value 2θ* was calculated via least-squares fitting for the one-dimensional diffraction profile of Cu (311) using eq. (2). Here, the one-dimensional profile was calculated by sector averaging of the obtained two-dimensional diffraction profile in the center-angle range of �3� with respect to the vertical direction. Eq. (2) is a combination of an asymmetric pseudo-Voigt function and a 2
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linear function representing a diffraction-peak profile and a baseline, respectively, as follows: 8 9 � �" � �2 # 1 > > > > 2η 2θ 2θ* > > > > 1þ4⋅ asymm⋅ > > > > < π⋅FWHM = FWHM Ið2θÞ ¼ scale⋅ # " �� � � � � > 0:5 2 > > > > > 2ð1 ηÞ ln 2 2θ 2θ* > > > > exp 4ln2⋅ asymm⋅ > > : þ FWHM ; π FWHM
200
intensity / a.u.
ψ = 0.0°
þðaþb⋅2θÞ; (2) where FWHM denotes the full-width at half-maximum of the peak; η denotes the fraction of the Lorentz function in the pseudo-Voigt func tion; asymm denotes an asymmetry; scale and a denote scaling parame ters for the pseudo-Voigt and linear functions, respectively; and b denotes the slope of the linear function. Here, asymm was defined using a variable A with 1 < A < 1 as asymm ¼ 1 A for 2θ < 2θ* and asymm ¼ 1 þ A for 2θ � 2θ*. Notably, the structure change in 2.5 min was assumed to be negli gible in this study, where 2.5 min was the time required for the diffraction measurements with five Ψ and axial movement of the goni ometer. In addition, the residual stress was assumed to be plane stress and a stress gradient was assumed to be absent in the thickness direction of the copper foil. It has been confirmed that no internal residual stress was generated in the copper foil when the untreated foil was heated alone to the curing temperature.
150
ψ = 20.7°
100
ψ = 30.0°
50
ψ = 37.8° ψ = 45.0°
0 135
136
137
138
139
140
141
2θ / ° Fig. 1. One-dimensional X-ray diffraction profiles of Cu(331) at 45 � C of the semicured resin–copper composite before the curing process. The profiles are vertically shifted for clarity. The solid lines represent the fitting curves.
138.10
2.3. TMDSC analysis
slope = 0.367 ± 0.015
2θ * / °
TMDSC measurements were performed on a DSC25 (TA Instruments, USA). The specimen was 6.5 mg of small pieces thinly sliced from the disk-shaped resin and used without being ground to a powder. The temperature profiles of the curing and thermal-cycle testing processes used in the residual stress analysis were applied, with the exception that the temperature was modulated and the average heating and cooling rates were �3.0 � C/min. The amplitude and period of the temperature modulation were 1.0 � C and 60 s, respectively. A Tzero aluminum pan and lid set (TA Instruments) was used for sample cells.
138.05
138.00
137.95 0.0
3. Results and discussion
0.1
0.2
sin
Fig. 1 shows the one-dimensional diffraction profiles of Cu(331) at 45 � C of the semicured resin–copper composite before the curing pro cess. The peak diffraction angle increased as Ψ increased, which in dicates the presence of residual stress at the interface between resin and copper. The relation between 2θ* and sin2 Ψ is shown in Fig. 2. It is clear that they have linear relation, indicating that there was no stress gradient in copper toward the thickness direction. The residual stress in copper was calculated as 59.6 � 2.5 MPa from the slope of 0.367 � 0.015 of the 2θ*–sin2 Ψ diagram, which indicates the presence of compressive stress in the copper. This result can be explained by considering the molding process of the resin–copper composite material. The adhesive interface between resin and copper was first formed when the resin melted in the mold, where the interface was in the nonstressed state with σ � 0 MPa. Subsequently, the specimen was removed from the mold and cooled to room temperature. During this cooling process, the resin exhibited greater shrinkage than the copper because of the dif ference in their CTEs. In addition, the specimen exhibited a symmetric three-layer structure in the direction normal to the sample surface, which prevented stress relaxation via macroscopic deformation such as warpage. Therefore, residual stress accumulated at the interface and compressive stress was generated in the copper. Fig. 3 shows the change in residual stress in the copper foil as a function of processing time (t) during the curing and thermal-cycle testing. First, the residual stress largely changed toward the tensile
0.3
0.4
0.5
ψ
2
Fig. 2. 2θ*–sin2 ψ diagram of Cu(331) at 45 � C of the semicured resin–copper composite before the curing process. The solid line represents the fitting line.
stress direction as the temperature was increased to 180 � C. This behavior can be explained by the amount of thermal expansion of the resin being larger than that of the copper, which resulted in a relaxation of the compressive stress and an increase in the tensile stress in the copper. The residual stress changed from negative to positive when the temperature was increased from 90 � C to 105 � C, which indicates that the adhesive interface of the resin and copper formed in this temperature range during the molding process, consistent with the molten temper ature of the uncured resin. In the holding process at 180 � C for 6 h, the stress in copper gradually shifted toward the compressive stress direc tion and the value changed from 11.3 � 0.7 to 9.3 � 0.9 MPa. This behavior can be clearly explained by the cure shrinkage of the resin. In the cooling process from 180 � C to 45 � C after the curing, the stress largely shifted toward the compression direction and the value changed from 9.3 � 0.9 to 113.9 � 2.6 MPa. This behavior is also explained by the magnitude of the thermal contraction of the resin being larger than that of copper because typical HMTA-cured phenolic resins containing no fillers exhibit a CTE of several tens ppm/K, i.e., 46 ppm/K in the temperature range of 20–90 � C [21], which is larger than the copper’s CTE of 16.8 ppm/K. The directions of the residual stress change were 3
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30
–30 180
–60
135
–90
–150
90
stress temperature
–120
temperature / °C
σ / MPa
0
45 0
2
4
6
8
10
12
t/h Fig. 3. Change in the residual stress in copper during curing and thermal-testing processes. The positive and negative values of the stress denote tensile and compressive stresses, respectively. The dashed line represents zero stress. The solid line represents temperature that was recorded on the sample stage.
nonreversing heat flow / W g
−1
toward tensile and compression directions in the heating and cooling processes, respectively, indicating the magnitude relation of CTE be tween the resin and copper did not reverse by curing. On the contrary, the amount of the residual stress change during cooling from 180 � C to 45 � C after curing became significantly larger than that during heating from 45 � C to 180 � C before curing, indicating that the modulus of the resin increased in the curing process by the formation of cross-linked network structures in phenolic resins. The increase in the modulus of the resin resulted in an increase in the force to deform the crystal lattice of copper, which caused the larger stress change after curing. In the thermal-cycle testing process in the t range of 8–13 h, the stress in copper changed toward the tensile stress direction as the temperature increased and toward the compressive stress direction as the tempera ture decreased. The explanation of these behaviors is similar to that of the aforementioned difference in the magnitudes of thermal expansion and contraction between cured resin and copper that resulted from the difference in their CTEs. These behaviors were reversible, and the stresses at 45 � C after the first and second cycles were 114.9 � 1.6 and 113.3 � 2.7 MPa, respectively, which are almost the same as the stress at 45 � C after the curing process. These results indicate that the curing of the resin was essentially completed and that no further structural change occurred during the thermal-cycle testing process, with the exception of the thermal expansion and contraction. However, the stresses in copper at 180 � C in the thermal-cycle testing process (at t ¼ 9.5 and 12 h) clearly differed from the stress at 180 � C after curing for 6 h (at t ¼ 7 h). This difference indicates that the compressive stress in copper at 180 � C generated by the cure shrinkage of the resin was relaxed via the cooling and heating after the curing process. This thermal-annealing-induced stress relaxation behavior suggests that structural strain was generated in the resin during the formation of the three-dimensional cross-linked network structure in the curing process and that the cross-link strain was relaxed via the macroscopic structural contraction and expansion during the cooling and heating processes, respectively. Such macroscopic structural change could accompany the collapse of some nanovoids and rearrangement of local hydrogen bonding network structures of phenolic resins, in which the hydrogen bonding is an important feature [22–28]. The formation of nanovoids and microvoids can occur during curing owing to the emission of volatiles, such as ammonia which were generated via decomposition of HMTA [1] and the presence of nano voids in the HMTA-cured phenolic resin was confirmed via our small-angle X-ray and neutron scattering analyses [29]. To investigate the relaxation behavior, TMDSC analysis was per formed under a similar temperature profile as that shown in Fig. 3; the nonreversing heat flow results are shown in Fig. 4. Herein, the non reversing signals near 45 � C and 180 � C both in heating and cooling processes were missing; however, the specimen was surely treated
0.05 1st & 2nd & 3rd cooling 0.04 0.03
2nd & 3rd heating
0.02 0.01
1st heating
0.00 45 60 75 90 105 120 135 150 165 180 temperature / °C
Fig. 4. Nonreversing heat flow in heating and cooling processes during curing and thermal-cycle testing. The exothermic and endothermic events are plotted as positive and negative heat flows, respectively.
between 45 � C and 180 � C. The absence of the signal was due to a deconvolution procedure of the observed modulated signal using a Fourier transform [30,31]. In the first heating process to 180 � C shown as “1st heating,” an endothermic peak at 65 � C and exothermic heat flow beginning at 140 � C were observed because of enthalpy relaxation and progress of the curing reactions, respectively. Here, the disappearance of the enthalpy relaxation peak was confirmed in the second scan of an additional thermal-cycle measurement between 30 � C and 90 � C in the absence of the curing reaction. However, no distinct behavior was observed in the nonreversing heat flow after curing at 180 � C for 6 h shown as “1st cooling.” In addition, the second and third cooling curves coincide with the first cooling curve and the second and third heating curves coincide with each other. These results indicate that, after curing at 180 � C for 6 h, the sample exhibited no distinct nonreversible macroscopic structural change, including structural relaxation, that affected the heat capacity of the cured resin. This result is inconsistent with the annealing-induced relaxation results of the residual stress analysis; however, because this TMDSC result represents the behavior of the resin as a whole, the observed annealing-induced relaxation of the residual stress at 180 � C might reflect the relaxation behavior near the interface between resin and copper, where phenolic resins are two-dimensionally constrained by the copper foil.
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4. Conclusions
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The residual stress in a phenolic resin and copper foil composite material during curing and thermal-cycle testing was analyzed in situ via time-resolved XRD measurements. The residual stress was estimated from the stress in copper, which was calculated via analysis of the XRD profile of the Cu(331) diffraction using the sin2 Ψ method in conjunction with the iso-inclination method. Before curing, the semicured resin –copper composite exhibited a large compressive stress, indicating that the adhesive interface between resin and copper was first formed when the resin melted in the molding process and the magnitude of the thermal contraction of the resin was larger than that of copper in the subsequent cooling to room temperature, which was caused by the dif ference in the CTEs between resin and copper. While curing the com posite at 180 � C, the cure shrinkage of the resin was successfully investigated as the residual stress change in the copper. Additional heating to the curing temperature in the heat-cycle test revealed that the residual stress generated during the curing was relaxed. The TMDSC analysis indicated that no distinct nonreversible structural change accompanied the change in heat capacity of the cured resin during the thermal-cycle test. We speculated that structural strain was generated in the cross-linked network structure during the curing process and that the cross-link strain was relaxed during annealing, in which this behavior occurred especially near the interface between resin and copper where the resin was two-dimensionally constrained by the copper foil. Inter facial structural analysis using X-ray, neutron scattering and reflection techniques, and molecular dynamics simulations will be the focus of future work in our project of the structure–properties investigation of phenolic resins [28,29,32–37]. Acknowledgment We would like to thank Prof. Takashi Nishino of Kobe University for valuable advice and comments on residual stress analysis of thermo setting resins. The X-ray diffraction experiment and its preliminary ex periments were performed at SPring-8 BL19B2 beamline with the approval of the Japan Synchrotron Radiation Research Institute (pro posal numbers 2017A1813, 2017B1893, 2018A1751, and 2018B1575). References [1] A. Gardziella, L.A. Pilato, A. Knop, Phenolic Resins: Chemistry, Applications, Standardization, Safety and Ecology, 2nd completely rev. ed, Springer, Berlin, 1999. [2] K. Nakamae, T. Nishino, X. Airu, T. Matsumoto, T. Matsumoto, Studies on mechanical properties of polymer composites by X-ray diffraction. I. Residual stress in epoxy resin by X-ray diffraction, J. Appl. Polym. Sci. 40 (1990) 2231–2238. [3] T. Nishino, X. Airu, T. Matsumoto, K. Matsumoto, K. Nakamae, Residual stress in particulate epoxy resin by X-ray diffraction, J. Appl. Polym. Sci. 45 (1992) 1239–1244. [4] M.L. Sham, J.K. Kim, Evolution of residual stresses in modified epoxy resins for electronic packaging applications, Composites Part A 35 (2004) 537–546. [5] M.R. Wisnom, M. Gigliotti, N. Ersoy, M. Campbell, K.D. Potter, Mechanisms generating residual stresses and distortion during manufacture of polymer-matrix composite structures, Composites Part A 37 (2006) 522–529. [6] M. Merzlyakov, G.B. McKenna, S.L. Simon, Cure-induced and thermal stresses in a constrained epoxy resin, Composites Part A 37 (2006) 585–591. [7] S.S. Kim, H. Murayama, K. Kageyama, K. Uzawa, M. Kanai, Study on the curing process for carbon/epoxy composites to reduce thermal residual stress, Composites Part A 43 (2012) 1197–1202. [8] K.T. Hsiao, Embedded single carbon fibre to sense the thermomechanical behavior of an epoxy during the cure process, Composites Part A 46 (2013) 117–121. [9] O.G. Kravchenko, C.Y. Li, A. Strachan, S.G. Kravchenko, R.B. Pipes, Prediction of the chemical and thermal shrinkage in a thermoset polymer, Composites Part A 66 (2014) 35–43. [10] J. Middleton, J. Hoffman, B. Burks, P. Predecki, M. Kumosa, Aging of a polymer core composite conductor: mechanical properties and residual stresses, Composites Part A 69 (2015) 159–167.
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In situ residual stress analysis in a phenolic resin and copper composite material during curing Atsushi Izumi a, *, Takeshi Kakara a, Midori Wakabayashi Otsuki a, Yasuyuki Shudo a, Tomoyuki Koganezawa b, Mitsuhiro Shibayama c, ** a b c
Corporate Engineering Center, Sumitomo Bakelite Co., Ltd, 2100 Takayanagi, Fujieda, Shizuoka, 426-0041, Japan Japan Synchrotron Radiation Research Institute, 1-1-1 Kouto, Sayo, Hyogo, 679-5198, Japan Neutron Science Laboratory, The Institute for Solid State Physics, The University of Tokyo, 5-1-5 Kashiwanoha, Kashiwa, Chiba, 277-8581, Japan
A R T I C L E I N F O
A B S T R A C T
Keywords: Phenolic resins Curing process Residual stress
Herein, in situ analysis of the residual stress in a phenolic resin and copper foil composite material during curing and thermal-cycle testing was performed via time-resolved X-ray diffraction measurements. The semicured resin–copper composite exhibited a large compressive stress in copper before the curing process. This indicates that the adhesive interface between resin and copper was first formed when the resin melted in the molding process and the magnitude of the thermal contraction of the resin was larger than that of copper in the subse quent cooling to room temperature after the molding. The difference in the magnitude was caused by the dif ference in the coefficients of thermal expansion between resin and copper. This compressive stress decreased as the temperature was increased to curing temperature of 180 � C. As the curing proceeded, the compressive stress in the copper again increased because of cure shrinkage of the resin. When the cured sample was reheated to the curing temperature, the compressive stress in the copper at 180 � C was relaxed. This thermal-annealing-induced stress relaxation suggests that cross-linking reactions during the curing process caused structural strains near the interface between resin and copper, where phenolic resins are two-dimensionally constrained by the copper foil, and that the cross-link strains were relaxed via macroscopic deformation induced by thermal contraction and expansion of the resin.
1. Introduction Phenolic resin composites comprising resin and copper are broadly used in electronic devices such as copper-clad laminates and commu tators because of their excellent physical properties, which include electrical insulator behavior, good heat resistance, and good solvent resistance. One of the important requirements for such composite ma terials is reliable adhesion at the resin–copper interface because delamination causes not only the severe deterioration of physical properties but also fatal defects in the composite products [1]. Typical manufacturing process of composite products comprises molding and curing processes. A composite comprising a semicured phenolic resin and copper is first prepared via laminating or insert molding of an uncured resin together with copper in brief heat treat ment. An additional heat treatment is then performed for several hours as the curing process to drive cross-linking reactions of the phenolic
resins [1]. In this process, thermal stress is generated at the interface between resin and copper because of a difference in the magnitude of thermal expansion and contraction of the materials. This difference is caused not only by the difference in the coefficient of thermal expansion (CTE) between resin and copper but also by the cross-linking reactions that accompany the resin’s cure shrinkage. When this thermal stress is not relaxed, it remains in the material as residual stress, which is an internally balanced stress that exists in the absence of an external force. The residual stress at the interface causes undesirable warpage and interfacial peeling of the composite products. Because the generation of residual stress is inevitable in composites comprising thermosetting resins and other materials, material and process designs for reducing the undesirable thermal stress during the manufacturing process are needed to improve the adhesion reliability at the interface between them. Typically, in industry, the adhesion reliability is evaluated as the num ber of thermal-shock cycles at which the interfacial delamination does
* Corresponding author. ** Corresponding author. E-mail addresses: [email protected] (A. Izumi), [email protected] (M. Shibayama). https://doi.org/10.1016/j.polymer.2019.121857 Received 3 July 2019; Received in revised form 24 September 2019; Accepted 1 October 2019 Available online 1 October 2019 0032-3861/© 2019 Elsevier Ltd. All rights reserved.
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not occur. However, this method is just an evaluation of whether or not the delamination proceeds, which does not provide information about mechanisms that caused delamination or increased adhesion reliability. Therefore, understanding the residual stress generation mechanism in the thermal curing and thermal-cycle testing processes is important [2–13]. Numerous residual stress analysis techniques are available, including methods based on X-ray diffraction (XRD), Raman spectroscopy, ultra sonic acoustics, and strain gauges [14–19]. Among them, the XRD method, which can evaluate lattice strain, is particularly effective for analyzing residual stress in crystalline materials because the lattice spacing is affected by the degree of residual stress and the relation be tween stress directions and lattice planes. When tensile stress is applied to the crystal, the lattice spacing with the lattice plane normal and parallel to the stress axis increases and decreases, respectively, under the relation of the Poisson’s ratio. When compressive stress is applied, the lattice spacing exhibits the opposite behavior. XRD methods for residual stress analysis are divided into cos α methods and sin2 Ψ methods, where α represents the central angle of the Debye ring and Ψ represents the angle between the normal direction of the sample surface and the normal direction of the crystal lattice plane. The cos α methods evaluate an entire Debye ring, and the stress can be calculated from the Debye ring distortion as a function of α. The sin2 Ψ methods evaluate a segment of the Debye ring, and the stress can be calculated from the diffraction-peak angle as a function of Ψ . The sin2 Ψ methods are further divided into iso-inclination and side-inclination methods on the basis of the measurement direction of the diffraction angle. Thus, a wide range of diffraction methods can be selected on the basis of the specimen shape, the source of the incident beam, and the instrumental dif fractometer–detector configuration. In this paper, in situ residual stress analysis in a phenolic resin and copper foil composite material during curing and thermal-cycle testing was performed via a time-resolved XRD measurement using the sin2 Ψ method. The residual stress at the resin–copper interface in the com posite material was calculated as stress in copper via evaluation of the lattice strain of Cu(331) plane. We believe that the proposed study represents the first report of in situ residual stress analysis at the inter face between a phenolic resin and copper during the curing and thermalcycle testing processes.
methylol groups. In their curing, the formation of methylene bridges proceeds with emittions of water and formaldehyde. Herein, the HMTA-cured novolac resin was used as a model resin because the curing system has been intensely investigated in our research project of eluci dating cross-link inhomogeneity in phenoic resins; however, resole-curing system is preferred for the electronic devices in industry because HMTA decomposes to ammonia during curing, thereby causing the corrosion of copper in the device [1]. 2.2. Residual stress analysis XRD measurements were performed at the BL19B2 beamline at the SPring-8 synchrotron radiation facility (Hyogo, Japan). The diffraction profile of the Cu(331) plane was measured using a six-axis diffractom eter (Huber Diffraktionstechnik GmbH, Germany) equipped with a DHS1100 domed hotstage (Anton Paar GmbH, Austria) and a Pilatus 300K two-dimensional X-ray detector (DECTRIS Ltd., Switzerland). The sam ple curing was performed in the air inside a Kapton dome. The energy of the incident X-rays was 8.0 keV, and the longitudinal and horizontal slit sizes for the incident beam were 1 and 6 mm, respectively. As a residual stress measurement method, the sin2 Ψ method was used in conjunction with the iso-inclination method, where a set of Ψ was chosen as 0.0, 20.7, 30.0, 37.8, and 45.0� , which correspond to sin2 Ψ of 0.000, 0.125, 0.250, 0.376, and 0.500, respectively. The X-ray exposure time at each Ψ was 10 s. In the curing process, the temperature of the sample stage was increased from 45 � C to 180 � C, maintained at 180 � C for 6 h, and then cooled to 45 � C. Subsequently, two additional heating and cooling pro cesses between 45 � C and 180 � C were performed as thermal-cycle tests. In the heating and cooling processes, the stress analysis was performed at 15 � C intervals. The heating and cooling rates were �10 � C/min and, after the temperature was increased or decreased by 15 � C, the tem perature was kept constant during sample alignment and diffraction measurements, corresponding to an average heating and cooling rate of approximately �2.3 � C/min. During the 180 � C isothermal treatment, stress analysis was performed at 7 min intervals. Here the sample alignment denotes the direct-beam-half-cut alignment, which corrects the vertical position of the resin–copper interface with respect to the incident beam. The alignment was performed before each sin2 Ψ mea surement because the sample thickness changed as a result of both thermal expansion–contraction and cure shrinkage of the resin. The residual stress (σ) in the copper was calculated by # � �" E 1 π ∂ð2θ* Þ � ; (1) σ¼ ⋅ ⋅ ⋅ 2ð1 þ νÞ tan θ*0 180 ∂ sin2 ψ
2. Experimental 2.1. Materials The resin and copper composite material used in the evaluation had a disk-shaped three-layered structure with the phenolic resin sandwiched between copper foils. A semicured specimen with a diameter and thickness of 24 and 1.5 mm, respectively, was prepared via compression molding at 175 � C for 3 min under a pressure of 10 MPa, corresponding to the composite comprising a semicured phenolic resin and copper, as described in Introduction section. For the resin layer, a 1.0/0.12 (wt/wt) mixture of a random-novolac resin with a number-average molecular weight of 910 g/mol (Sumitomo Bakelite Co., Ltd., Japan) and hexa methylenetetramine (HMTA) was used, wherein HMTA is a typical curing agent for novolac resins. As a copper foil, a surface-roughened electrodeposited copper foil with a thickness of 12 μm (Mitsui Mining and Smelting Co., Ltd., Japan) was used. A specimen of the semicured resin without the copper foil was prepared in a similar manner for temperature-modulated differential scanning calorimetry (TMDSC) analysis. Herein, there are mainly two types of phenolic resins, namely novolac and resole resins. The novolac resins are typically cured with HMTA and curing reaction start with formations of benzoxazine and benzylamine structures, followed by formations of methylene, amide, amine, and imine bridges with liberating ammonia [20]. On the con trary, the resole resins can be cured itself because they contain reactive
where E and ν denote Young’s modulus and Poisson’s ratio, respectively; θ0 and 2θ* denote the diffraction angle of the Cu(331) reflection of the specimen in the non-strained state and the peak top angle of the observed diffraction from Cu(331), respectively. Herein, the positive and negative values of σ represent tensile and compressive stresses, respectively. The values E and ν of the copper foil used in this study were 65.0 GPa and 0.343, respectively. The value θ0 at 25 � C for X-ray energy of 8.0 keV was determined to be 138.04� , which corresponds to a lattice spacing of 0.083 nm assuming that the lattice in the copper foil used in this study was in the non-strained state. The value θ0 at different temperatures was corrected using the copper’s CTE of 16.8 ppm/K. The value ∂(2θ*)/ ∂(sin2 Ψ ) was calculated from the slope of the linear regression of the 2θ*–sin2 Ψ diagram, and the standard error of the slope was used to estimate the error of the residual stress. The value 2θ* was calculated via least-squares fitting for the one-dimensional diffraction profile of Cu (311) using eq. (2). Here, the one-dimensional profile was calculated by sector averaging of the obtained two-dimensional diffraction profile in the center-angle range of �3� with respect to the vertical direction. Eq. (2) is a combination of an asymmetric pseudo-Voigt function and a 2
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linear function representing a diffraction-peak profile and a baseline, respectively, as follows: 8 9 � �" � �2 # 1 > > > > 2η 2θ 2θ* > > > > 1þ4⋅ asymm⋅ > > > > < π⋅FWHM = FWHM Ið2θÞ ¼ scale⋅ # " �� � � � � > 0:5 2 > > > > > 2ð1 ηÞ ln 2 2θ 2θ* > > > > exp 4ln2⋅ asymm⋅ > > : þ FWHM ; π FWHM
200
intensity / a.u.
ψ = 0.0°
þðaþb⋅2θÞ; (2) where FWHM denotes the full-width at half-maximum of the peak; η denotes the fraction of the Lorentz function in the pseudo-Voigt func tion; asymm denotes an asymmetry; scale and a denote scaling parame ters for the pseudo-Voigt and linear functions, respectively; and b denotes the slope of the linear function. Here, asymm was defined using a variable A with 1 < A < 1 as asymm ¼ 1 A for 2θ < 2θ* and asymm ¼ 1 þ A for 2θ � 2θ*. Notably, the structure change in 2.5 min was assumed to be negli gible in this study, where 2.5 min was the time required for the diffraction measurements with five Ψ and axial movement of the goni ometer. In addition, the residual stress was assumed to be plane stress and a stress gradient was assumed to be absent in the thickness direction of the copper foil. It has been confirmed that no internal residual stress was generated in the copper foil when the untreated foil was heated alone to the curing temperature.
150
ψ = 20.7°
100
ψ = 30.0°
50
ψ = 37.8° ψ = 45.0°
0 135
136
137
138
139
140
141
2θ / ° Fig. 1. One-dimensional X-ray diffraction profiles of Cu(331) at 45 � C of the semicured resin–copper composite before the curing process. The profiles are vertically shifted for clarity. The solid lines represent the fitting curves.
138.10
2.3. TMDSC analysis
slope = 0.367 ± 0.015
2θ * / °
TMDSC measurements were performed on a DSC25 (TA Instruments, USA). The specimen was 6.5 mg of small pieces thinly sliced from the disk-shaped resin and used without being ground to a powder. The temperature profiles of the curing and thermal-cycle testing processes used in the residual stress analysis were applied, with the exception that the temperature was modulated and the average heating and cooling rates were �3.0 � C/min. The amplitude and period of the temperature modulation were 1.0 � C and 60 s, respectively. A Tzero aluminum pan and lid set (TA Instruments) was used for sample cells.
138.05
138.00
137.95 0.0
3. Results and discussion
0.1
0.2
sin
Fig. 1 shows the one-dimensional diffraction profiles of Cu(331) at 45 � C of the semicured resin–copper composite before the curing pro cess. The peak diffraction angle increased as Ψ increased, which in dicates the presence of residual stress at the interface between resin and copper. The relation between 2θ* and sin2 Ψ is shown in Fig. 2. It is clear that they have linear relation, indicating that there was no stress gradient in copper toward the thickness direction. The residual stress in copper was calculated as 59.6 � 2.5 MPa from the slope of 0.367 � 0.015 of the 2θ*–sin2 Ψ diagram, which indicates the presence of compressive stress in the copper. This result can be explained by considering the molding process of the resin–copper composite material. The adhesive interface between resin and copper was first formed when the resin melted in the mold, where the interface was in the nonstressed state with σ � 0 MPa. Subsequently, the specimen was removed from the mold and cooled to room temperature. During this cooling process, the resin exhibited greater shrinkage than the copper because of the dif ference in their CTEs. In addition, the specimen exhibited a symmetric three-layer structure in the direction normal to the sample surface, which prevented stress relaxation via macroscopic deformation such as warpage. Therefore, residual stress accumulated at the interface and compressive stress was generated in the copper. Fig. 3 shows the change in residual stress in the copper foil as a function of processing time (t) during the curing and thermal-cycle testing. First, the residual stress largely changed toward the tensile
0.3
0.4
0.5
ψ
2
Fig. 2. 2θ*–sin2 ψ diagram of Cu(331) at 45 � C of the semicured resin–copper composite before the curing process. The solid line represents the fitting line.
stress direction as the temperature was increased to 180 � C. This behavior can be explained by the amount of thermal expansion of the resin being larger than that of the copper, which resulted in a relaxation of the compressive stress and an increase in the tensile stress in the copper. The residual stress changed from negative to positive when the temperature was increased from 90 � C to 105 � C, which indicates that the adhesive interface of the resin and copper formed in this temperature range during the molding process, consistent with the molten temper ature of the uncured resin. In the holding process at 180 � C for 6 h, the stress in copper gradually shifted toward the compressive stress direc tion and the value changed from 11.3 � 0.7 to 9.3 � 0.9 MPa. This behavior can be clearly explained by the cure shrinkage of the resin. In the cooling process from 180 � C to 45 � C after the curing, the stress largely shifted toward the compression direction and the value changed from 9.3 � 0.9 to 113.9 � 2.6 MPa. This behavior is also explained by the magnitude of the thermal contraction of the resin being larger than that of copper because typical HMTA-cured phenolic resins containing no fillers exhibit a CTE of several tens ppm/K, i.e., 46 ppm/K in the temperature range of 20–90 � C [21], which is larger than the copper’s CTE of 16.8 ppm/K. The directions of the residual stress change were 3
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30
–30 180
–60
135
–90
–150
90
stress temperature
–120
temperature / °C
σ / MPa
0
45 0
2
4
6
8
10
12
t/h Fig. 3. Change in the residual stress in copper during curing and thermal-testing processes. The positive and negative values of the stress denote tensile and compressive stresses, respectively. The dashed line represents zero stress. The solid line represents temperature that was recorded on the sample stage.
nonreversing heat flow / W g
−1
toward tensile and compression directions in the heating and cooling processes, respectively, indicating the magnitude relation of CTE be tween the resin and copper did not reverse by curing. On the contrary, the amount of the residual stress change during cooling from 180 � C to 45 � C after curing became significantly larger than that during heating from 45 � C to 180 � C before curing, indicating that the modulus of the resin increased in the curing process by the formation of cross-linked network structures in phenolic resins. The increase in the modulus of the resin resulted in an increase in the force to deform the crystal lattice of copper, which caused the larger stress change after curing. In the thermal-cycle testing process in the t range of 8–13 h, the stress in copper changed toward the tensile stress direction as the temperature increased and toward the compressive stress direction as the tempera ture decreased. The explanation of these behaviors is similar to that of the aforementioned difference in the magnitudes of thermal expansion and contraction between cured resin and copper that resulted from the difference in their CTEs. These behaviors were reversible, and the stresses at 45 � C after the first and second cycles were 114.9 � 1.6 and 113.3 � 2.7 MPa, respectively, which are almost the same as the stress at 45 � C after the curing process. These results indicate that the curing of the resin was essentially completed and that no further structural change occurred during the thermal-cycle testing process, with the exception of the thermal expansion and contraction. However, the stresses in copper at 180 � C in the thermal-cycle testing process (at t ¼ 9.5 and 12 h) clearly differed from the stress at 180 � C after curing for 6 h (at t ¼ 7 h). This difference indicates that the compressive stress in copper at 180 � C generated by the cure shrinkage of the resin was relaxed via the cooling and heating after the curing process. This thermal-annealing-induced stress relaxation behavior suggests that structural strain was generated in the resin during the formation of the three-dimensional cross-linked network structure in the curing process and that the cross-link strain was relaxed via the macroscopic structural contraction and expansion during the cooling and heating processes, respectively. Such macroscopic structural change could accompany the collapse of some nanovoids and rearrangement of local hydrogen bonding network structures of phenolic resins, in which the hydrogen bonding is an important feature [22–28]. The formation of nanovoids and microvoids can occur during curing owing to the emission of volatiles, such as ammonia which were generated via decomposition of HMTA [1] and the presence of nano voids in the HMTA-cured phenolic resin was confirmed via our small-angle X-ray and neutron scattering analyses [29]. To investigate the relaxation behavior, TMDSC analysis was per formed under a similar temperature profile as that shown in Fig. 3; the nonreversing heat flow results are shown in Fig. 4. Herein, the non reversing signals near 45 � C and 180 � C both in heating and cooling processes were missing; however, the specimen was surely treated
0.05 1st & 2nd & 3rd cooling 0.04 0.03
2nd & 3rd heating
0.02 0.01
1st heating
0.00 45 60 75 90 105 120 135 150 165 180 temperature / °C
Fig. 4. Nonreversing heat flow in heating and cooling processes during curing and thermal-cycle testing. The exothermic and endothermic events are plotted as positive and negative heat flows, respectively.
between 45 � C and 180 � C. The absence of the signal was due to a deconvolution procedure of the observed modulated signal using a Fourier transform [30,31]. In the first heating process to 180 � C shown as “1st heating,” an endothermic peak at 65 � C and exothermic heat flow beginning at 140 � C were observed because of enthalpy relaxation and progress of the curing reactions, respectively. Here, the disappearance of the enthalpy relaxation peak was confirmed in the second scan of an additional thermal-cycle measurement between 30 � C and 90 � C in the absence of the curing reaction. However, no distinct behavior was observed in the nonreversing heat flow after curing at 180 � C for 6 h shown as “1st cooling.” In addition, the second and third cooling curves coincide with the first cooling curve and the second and third heating curves coincide with each other. These results indicate that, after curing at 180 � C for 6 h, the sample exhibited no distinct nonreversible macroscopic structural change, including structural relaxation, that affected the heat capacity of the cured resin. This result is inconsistent with the annealing-induced relaxation results of the residual stress analysis; however, because this TMDSC result represents the behavior of the resin as a whole, the observed annealing-induced relaxation of the residual stress at 180 � C might reflect the relaxation behavior near the interface between resin and copper, where phenolic resins are two-dimensionally constrained by the copper foil.
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4. Conclusions
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The residual stress in a phenolic resin and copper foil composite material during curing and thermal-cycle testing was analyzed in situ via time-resolved XRD measurements. The residual stress was estimated from the stress in copper, which was calculated via analysis of the XRD profile of the Cu(331) diffraction using the sin2 Ψ method in conjunction with the iso-inclination method. Before curing, the semicured resin –copper composite exhibited a large compressive stress, indicating that the adhesive interface between resin and copper was first formed when the resin melted in the molding process and the magnitude of the thermal contraction of the resin was larger than that of copper in the subsequent cooling to room temperature, which was caused by the dif ference in the CTEs between resin and copper. While curing the com posite at 180 � C, the cure shrinkage of the resin was successfully investigated as the residual stress change in the copper. Additional heating to the curing temperature in the heat-cycle test revealed that the residual stress generated during the curing was relaxed. The TMDSC analysis indicated that no distinct nonreversible structural change accompanied the change in heat capacity of the cured resin during the thermal-cycle test. We speculated that structural strain was generated in the cross-linked network structure during the curing process and that the cross-link strain was relaxed during annealing, in which this behavior occurred especially near the interface between resin and copper where the resin was two-dimensionally constrained by the copper foil. Inter facial structural analysis using X-ray, neutron scattering and reflection techniques, and molecular dynamics simulations will be the focus of future work in our project of the structure–properties investigation of phenolic resins [28,29,32–37]. Acknowledgment We would like to thank Prof. Takashi Nishino of Kobe University for valuable advice and comments on residual stress analysis of thermo setting resins. The X-ray diffraction experiment and its preliminary ex periments were performed at SPring-8 BL19B2 beamline with the approval of the Japan Synchrotron Radiation Research Institute (pro posal numbers 2017A1813, 2017B1893, 2018A1751, and 2018B1575). References [1] A. Gardziella, L.A. Pilato, A. Knop, Phenolic Resins: Chemistry, Applications, Standardization, Safety and Ecology, 2nd completely rev. ed, Springer, Berlin, 1999. [2] K. Nakamae, T. Nishino, X. Airu, T. Matsumoto, T. Matsumoto, Studies on mechanical properties of polymer composites by X-ray diffraction. I. Residual stress in epoxy resin by X-ray diffraction, J. Appl. Polym. Sci. 40 (1990) 2231–2238. [3] T. Nishino, X. Airu, T. Matsumoto, K. Matsumoto, K. Nakamae, Residual stress in particulate epoxy resin by X-ray diffraction, J. Appl. Polym. Sci. 45 (1992) 1239–1244. [4] M.L. Sham, J.K. Kim, Evolution of residual stresses in modified epoxy resins for electronic packaging applications, Composites Part A 35 (2004) 537–546. [5] M.R. Wisnom, M. Gigliotti, N. Ersoy, M. Campbell, K.D. Potter, Mechanisms generating residual stresses and distortion during manufacture of polymer-matrix composite structures, Composites Part A 37 (2006) 522–529. [6] M. Merzlyakov, G.B. McKenna, S.L. Simon, Cure-induced and thermal stresses in a constrained epoxy resin, Composites Part A 37 (2006) 585–591. [7] S.S. Kim, H. Murayama, K. Kageyama, K. Uzawa, M. Kanai, Study on the curing process for carbon/epoxy composites to reduce thermal residual stress, Composites Part A 43 (2012) 1197–1202. [8] K.T. Hsiao, Embedded single carbon fibre to sense the thermomechanical behavior of an epoxy during the cure process, Composites Part A 46 (2013) 117–121. [9] O.G. Kravchenko, C.Y. Li, A. Strachan, S.G. Kravchenko, R.B. Pipes, Prediction of the chemical and thermal shrinkage in a thermoset polymer, Composites Part A 66 (2014) 35–43. [10] J. Middleton, J. Hoffman, B. Burks, P. Predecki, M. Kumosa, Aging of a polymer core composite conductor: mechanical properties and residual stresses, Composites Part A 69 (2015) 159–167.
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