Experimentally-validated computational model for temperature evolution within laser heated fiber-reinforced polymer matrix composites

Accepted Manuscript Experimentally-Validated Computational Model for Temperature Evolution Within Laser Heated Fiber-Reinforced Polymer Matrix Composites Sangwook Sihn, Jeremey Pitz, Roger H. Gerzeski, Ajit K. Roy, Jonathan P. Vernon PII: DOI: Reference:

S0263-8223(18)32085-3 https://doi.org/10.1016/j.compstruct.2018.09.041 COST 10187

To appear in:

Composite Structures

Please cite this article as: Sihn, S., Pitz, J., Gerzeski, R.H., Roy, A.K., Vernon, J.P., Experimentally-Validated Computational Model for Temperature Evolution Within Laser Heated Fiber-Reinforced Polymer Matrix Composites, Composite Structures (2018), doi: https://doi.org/10.1016/j.compstruct.2018.09.041

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Experimentally-Validated Computational Model for Temperature Evolution Within Laser Heated Fiber-Reinforced Polymer Matrix Composites Sangwook Sihna,b,∗, Jeremey Pitza,c , Roger H. Gerzeskia , Ajit K. Roya , Jonathan P. Vernona,∗∗ a

Air Force Research Laboratory, Materials and Manufacturing Directorate AFRL/RXA, Wright-Patterson AFB, OH, USA b University of Dayton Research Institute, Structural Materials Division, 300 College Park, Dayton, OH 45469-0060, USA c UES Inc., Dayton, OH, 45432, USA

Abstract Comprehensive experimental and computational studies have been conducted to accurately measure and predict the temperature evolution of polymer matrix composite material subjected to laser heating. Plain-woven 16-ply composites with T650 carbon fibers in a bismaleimide matrix were irradiated with a continuous wave fiber laser (1.0692 µm wavelength) with a 3.0 cm diameter flat-top beam profile. The spatial and temporal evolution of temperature for the front (laser-exposed side) and back (unexposed) surfaces of the composite specimens were monitored and recorded with calibrated infrared cameras. Additionally, internal temperature evolution was measured with thermocouples embedded in the midplane of the composite specimens. A nonlinear transient finite element (FE) analysis, based on conductive, convective, and ∗

Corresponding author: Tel: +1-937-255-9067. Corresponding author: Tel: +1-937-255-6636. Email addresses: [email protected] (Sangwook Sihn), [email protected] (Jonathan P. Vernon) ∗∗

Preprint submitted to Composite Structures

August 7, 2018

radiative heat transfer, was conducted using ANSYS to predict the temperature history during heating (laser exposure) and cooling (after exposure). For all investigated laser irradiances (i.e., 0.71 – 1.95 W cm−2 ), the numerical prediction demonstrated excellent agreement with experimental data. Keywords: A. Polymer-matrix composites (PMCs); B. Thermal properties; C. Finite element analysis (FEA); D. Thermal analysis. 1. Introduction Fiber-reinforced composite materials exhibit exceptional specific stiffness and specific strength, driving increased use in modern structural applications such as airplanes, ships, and sporting goods. For example, composite materials are rapidly assimilating into the aerospace industry with the Boeing Dreamliner 787 now composed of ∼50 % composite materials by weight [1]. However, challenges remain in understanding the damage progression, longterm performance, repair, and replacement of such structural materials. For a given material and environmental conditions, it is of interest to understand the effects of heating (and duration at elevated temperatures) on both mechanical properties and chemical stability. Understanding temperature evolution, and predicting the resulting thermophysical and thermochemical effects of thermal histories, will aid in the development of fleet management tools required to evaluate carbon fiber reinforced polymer matrix composite (PMC) materials subjected to laser paint stripping, laser-assisted manufacturing techniques, onboard fires, and engine exhaust impingement. Now, laser sources are widely used in various composite material manufacturing and processing techniques (e.g., hole drilling, cutting, curing, bonding, 2

sintering, and laser paint stripping) [2–10]. In such applications, relatively low irradiance laser exposures (2 – 20 W cm−2 ) may readily bring carbon fiber reinforced PMCs to temperatures above the matrix glass transition temperature. For a given composite material, heating conditions, and environmental conditions, it is of considerable interest to understand the 3-D spatial and temporal evolution of temperature throughout the irradiated material. To avoid any nonlinear effects (e.g., matrix degradation, evaporation, or pyrolysis), in this study we evaluated heating effects from 0.71 – 1.95 W cm−2 . Even in this thermal response regime, the composite material exhibits complex variations of thermal and mechanical properties upon heating, rendering the prediction of spatial and temporal temperature evolution in the composite materials difficult. The development of appropriate and reliable computational simulation methods and tools, once validated experimentally, can provide cost-effective and reliable means for better understanding the underlying mechanisms of laser/material interactions, and thus identifying critical manufacturing and processing parameters. Furthermore, such predictive capability can guide manufacturing and processing design (e.g., selecting type and wavelength of laser, irradiance level, heating rate, cooling mechanisms) [3, 11–16]. Conventional methods of measuring temperature evolution using thermocouples attached to a surface of the composite specimens (i.e., via chemical adhesive or physical pressure) often result in unreliable temperature measurement because of difficulty in maintaining adequate contact between the thermocouple junctions and the specimen surface throughout the heating and cooling processes. Rather than surface mounting and bonding, the thermo-

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couples can be embedded between plies within the composite laminates. Similarly, improved 2-D temperature/damage monitoring, potentially enabling in situ health monitoring, may be provided through the utilization of fiber bragg grating sensors [17]. Another effective approach is to use non-contact methods such as infrared (IR) cameras for spatially calibrated radiometric measurements. Such imaging-based methodologies require considerable calibration protocols, property measurements, and post-processing procedures to generate accurate temperatures, but provide full 2-D temperature data rather than data at discrete thermocouple locations. In the present study, a comprehensive experimental and computational study was conducted to understand the thermal response of woven composites to laser exposure. Non-contact spatially calibrated radiometric measurements and thermocouple measurements were implemented to monitor both surface and internal temperature evolution during laser heating and post-exposure cooling. Computational simulations were conducted using a finite element analysis for conductive, convective, and radiative heat transfer in the composite materials to predict the temperature rise and fall during the respective heating and cooling stages. Intrinsic thermal properties were measured at various temperatures to be used as input parameters for the simulations. Finally, the model predictions were compared and validated with the experimental measurements.

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2. Experimenal Study 2.1. Fabrication of Test Specimens Composite plates were fabricated with 16 plies of 3k plain-weave fabric prepregs composed of T650 carbon fibers and bismaleimide 5250-4 matrix. Any variation in the specimens would influence the experimental results. Among them, variation in fiber volume fraction would affect the material properties (i.e., density, heat capacity, thermal conductivity). Variation in thickness across the specimen would affect the through-thickness thermal transport, and thus the overall temperature distribution including both frontand back-surface temperatures. Considerable care was taken in producing nearly identical specimens with the exception of intentionally varied embedded thermocouple positions. All the specimens were fabricated as flat plate specimens with heat flux melting the resin flow between each woven prepreg ply when it was laid down. Then, for every four plies, vacuum degassing was conducted under a caul plate for 30 minutes to ensure uniform thickness and eliminate trapped gas voids. Additionally, both of the test specimens were simultaneously cured in the same autoclave run.

During laminate layup, OMEGA® type-K precision fine wire thermocou-

ples wrapped with 0.254 mm (0.01 in.)-diameter glass insulation were embedded at various random locations in the midplane between the 8th and 9th plies of the composite layups. The composite layups with embedded thermocouples were simultaneously cured and post-cured in an autoclave according to the manufacturer’s recommended cure cycle [18]. The cured panels were cut on a diamond saw into a square geometry with a side length of 101.6 mm (4 in.). The cut specimens were designed to have two thermocouples in each 5

square specimen. The cutting residue on all resulting specimens was washed off with distilled water. All of the specimens were dried under vacuum at 121◦ C (250◦ F) for 48 hours. After drying, all the specimens were placed and stored under vacuum in a desiccator before the laser exposure and property measurement. Two square plate specimens (specimen IDs: PMC-1 and PMC-2) were used in this study. The thicknesses of the two specimens (10 independent measurements with a micrometer) were (3.109 ± 0.033) mm and (3.145 ± 0.020) mm, respectively. A36 steel plates were precision-machined by wire electrical discharge machining (EDM) to form x-y axis grids. After affixing the grid to an edge of the square specimen, X-ray images were taken to identify the position of the embedded thermocouples. Figure 1(a) and (b) show a photograph and an X-ray image of PMC-1. Also denoted in Figure 1(b) are the locations of the embedded thermocouples with respect to an origin point of the x-y axis grid. The radial locations of two embedded thermocouples from the center of the laser irradiated area are denoted as r in the figure. 2.2. Measurement of Temperature Evolution under Laser Irradiation A continuous wave (CW) laser beam at a wavelength of 1.0692 µm was generated from an IPG YLR-100-WC-Y100 diode-pumped, solid-state single mode fiber laser system. The beam delivery fiber from the laser was terminated with a 100 mm focal length Optoskand collimator to produce a ∼1.03 cm Gaussian beam with M 2 = 1.21. The beam was then conditioned by pi-shaper beam shaping system, and then reimaged at 3x magnification with a convex lens to form a flat-top beam profile. The nominal diameter of the flat-top laser beam was 3.0 cm. The composite specimens were exposed 6

Figure 1: (a) Photograph of composite test specimen PMC-1 with an x-y axis grid. (b) X-ray image with two thermocouples embedded at the midplane of a 16-ply layup. The radial location of the embedded thermocouple junction tip from original of the x-y axis grid is denoted by r. (c) A video frame from unfiltered CCD camera taken during laser exposure. (d) An image of flat-top beam profile using a high-speed high-sensitivity camera with a bandpass filter allowing transmission of the laser wavelength.

to the flat-top CW laser for two durations (60 and 120 s) at four power levels (5, 7, 10 and 13.8 W), which correspond to irradiances of 0.71, 0.99, 1.41 and 1.95 W cm−2 , respectively. The target irradiation point for the center 7

of the 3 cm diameter laser beam spot on the specimen was desired to be at the (40 mm, 40 mm) point of the x-y axis grid, which is designated as “Laser beam center” in Figure 1(a). An in situ video frame of the composite during laser exposure taken with a Samsung SCB-8001N/EX HD camera is shown in Figure 1(c). The front surface of the irradiated specimen was also imaged at 1 Hz with a filtered PCO1200s camera for monitoring in situ beam location, diameter, and profile. A representative frame from such in situ monitoring is shown in Figure 1(d). The profiles shown within Figure 1(d) are a convolution of the surface morphology of the sample and spatial energy distribution of the laser. The high peak in the center of the profile is the result of a silver marker dot placed on the specimen prior to testing. Given the laser heat source and the InSb focal plane arrays used for non-contact radiometric measurements diagnostics, the total hemispherical spectral reflectance (R) of the T650/5250-4 composite panel was measured from 0.25 to 15 µm and is shown in Figure 2. At the laser wavelength of 1.069 µm, R is approximately 0.09. Given the thickness of the specimens, the measured transmittance of composite laminates was zero, thus the spectral absorptance (α) at the laser wavelength was calculated from α = 1 − R, which results in α ≈ 0.91. This value was used to calculate the heat flux (qin ) provided by the laser: qin = α

Pin , A

(1)

where Pin is an input laser power and A is an effective surface area of the incident laser beam. In the case of a flat-top profile of the laser beam, A = πD2 /4, where D is a diameter of the beam. The spatial and temporal evolution of both front (laser irradiated) and 8

Figure 2: Total hemispherical reflectance of T650/5250-4 composites as a function of wavelength.

back surface temperatures were monitored independently with two calibrated IR cameras (i.e., FLIR SC6813 (front) and SC6101 (back) each equipped with high dynamic range 640×512 InSb photovoltaic focal plane array). The IR cameras were operated at 50 Hz with two preset integration times alternating at 25 Hz. The presets each produced entirely separate videos. Each video was calibrated to a reference wide area blackbody at the target plane after performing a two point non-uniformity correction (NUC). The two points of the NUC data were collected at the extreme range of the linear response region of the focal plane arrays. Reference data for offset corrections to the NUC and calibration were collected just prior to each laser exposure. Temperatures were calculated from the two radiometric videos in Matlab using an inverse Planckian equation developed by Sakuma and Hattori, and

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combined using a high dynamic range (HDR) algorithm to produce a single video [19]. Specifically, a linear fit was applied to convert the signal on the detector to radiance, and the inverse Planckian equation of black body radiation was employed to convert radiance to temperature. Each InSb FPA was operated by cycling through two different integration times. Each individual pixel on the FPA was evaluated to determine the linear response regime. If the signal generated by an individual pixel approached saturation on the longer integration time, it was replaced by the temperature information from the shorter integration time to extend the dynamic range and enhance the temperature resolution. The HDR algorithm produces a single movie that consists of pixels collected from the same camera at different integration times. The temperature calibration (i.e., the emittance used to transform from radiance to temperature) assumed a constant emittance of 0.8605 as determined by an average value of the spectral hemispherical reflectance data between 1.5 to 5.0 µm, which is indicated as “Back surface IR camera” spectral range in Figure 2. Similar to the determination of the laser spectral absorptance, there was no measurable transmittance over this spectral regime. Hence, Kirchhoff’s law of thermal radiation was applied, equating the calculated average spectral absorptance to the emittance, i.e., ε = α [20]. The spectral reflectance was measured as a function of temperature, and it was observed that the reflectance did not exceed a 2 % change over the temperature range of this investigation (i.e., 23.7 – 200◦ C). An aluminum grid of square holes backlit with a wide area black body was employed to collect spatial calibration data for the IR cameras. A spatial

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calibration toolbox for Matlab provided intrinsic and extrinsic parameters for the camera and target relationships [21]. Additionally, images of the aluminum grid held at the target plane permitted the co-registration of the front camera image and the left/right flipped back camera image to produce a two dimensional temperature difference map of each specimen as it was exposed to the laser. These spatially registered 2-D images enable determination of the through thickness temperature difference as a function of time. 2.3. Characterization of Material Properties of Composites Material properties of the T650/5250-4 composites were measured for the computational simulations of the temperature evolution under laser irradiation. The material properties include density, heat capacity, thermal conductivities in both planar and through-thickness directions, and spectral surface absorptance. The composites specimen consists of 60 % volume fraction of T650 plain-woven carbon fibers and 40 % volume fraction of 5250-4 resin. The density (ρ) of the composites measured using the Archimedes’ Principle was 1550 kg m−3 . According to measurement using Thermogravimetric analysis (TGA), the density of the T650/5250-4 composite material is nearly constant for the temperature range considered in the present study. Heat capacity of (Cp ) was measured with a differential scanning calorimeter (DSC) Q2000 from TA Instruments over temperatures ranging from room temperature to 250◦ C at a heating rate of 5◦ C/min. Figure 3(a) shows the results of four repeated runs of Cp measurements. The measured data were approximated with the following quadratic formula: Cp = −0.005 T 2 + 3 T + 650, 11

(2)

where the units of Cp and T are in [J kg−1 K−1 ] and [◦ C], respectively. The quadratic formula is plotted with a dotted line in Figure 3(b). The woven fabric composite plate is assumed to be transversely isotropic having different in-plane and through-thickness thermal conductivities. The through-thickness thermal conductivity (kz ) of T650/5250-4 composites was indirectly obtained from the measurement of thermal diffusivity (αz ) with a laser flash apparatus (LFA 457) from Netzsch over temperatures ranging from room temperature to 150◦ C. Figure 3 (b) shows the temperature-dependent αz measured with two composite specimens. The measured data was linearly fit with the following equation: αz = −5.0 × 10−4 T + 0.5,

(3)

where the units of αz and T are in [mm2 s−1 ] and [◦ C], respectively. The linear formula is plotted with a dotted line in Figure 3(b). Through-thickness thermal conductivity of (kz ) was then calculated from thermal diffusivity, i.e., kz = ρCp αz . The solid line in Figure 3(c) shows the calculated kz . The variation of kz over temperature was further simplified with the following linear formula: kz = 5.66 × 10−4 T + 0.557,

(4)

where the units of kz and T are in [W m−1 K−1 ] and [◦ C], respectively. The linear formula is plotted with a dotted line in Figure 3(c). 3. Computational Study The computational model and methodology were developed to simulate the laser/material interactions, especially the temperature evolution, in the 12

Figure 3: (a) Measured heat capacity of four repeated runs, (b) measured throughthickness thermal diffusivity, and (c) calculated through-thickness thermal conductivity of T650/5250-4 composites as a function of temperature. Fits for data and calculated values utilized in computational study are shown as dashed red lines in each graph.

T650/5250-4 plain-woven composite material subject to the various laser exposures. Time- and space-dependent temperature, T (x, y, z, t), is a primary dependent variable, which is to be determined by solving nonlinear transient Fourier equation for conductive heat transfer along with convective and radiative boundary conditions. The governing equation for the heat transfer

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problems without a heat source or heat sink is written as ∂ (ρCp T ) − ∇ (k∇T ) = 0. ∂t

(5)

The boundary condition of the laser irradiation with convective and radiative heat fluxes is
4 n · (k∇T ) = qin + h(T∞ − T ) + εσ T∞ − T 4 = 0,

(6)

where n is a normal vector pointing out of the boundary surface, qin is an inward heat flux, h is a convective heat transfer coefficient, T∞ is an ambient temperature, ε is a radiative emittance, and σ is the Stefan-Boltzmann constant (5.67 × 108 W m−2 K−4 ). In the present study, T∞ was assumed to be the same as the initial room temperature, i.e., approximately 23.7◦ C (75◦ F). The heat flux density, qin , from the laser source was determined by Eq. 1. The radiative emittance was calculated by averaging the spectral emittance weighted by a perfect blackbody at a given temperature. Radiative emittance varied negligibly (i.e., 0.886 – 0.891) over the temperature range of the experiments (i.e., 23.7 – 200◦ C). Figure 4(a) shows a schematic of the current model description, including the geometry and boundary conditions of the composite plate. Although square panels were actually used in the experimental study, because of circular heat influx of the flat-top laser irradiation, we considered a simplified 2-D axisymmetric model with a cylindrical geometry, whose diameter is the same as the side length (l = 101.6 mm) of the square test specimens. Actual thicknesses of the 16-ply composite specimens were used in the simulation. Although not reported here, we conducted separate simulations with a 3-D model using the actual square-shaped geometry, 14

Figure 4: (a) Schematic of laser exposure set-up and model geometry with specified boundary conditions (BCs) utilized for numerical simulations, and (b) a typical radial temperature distribution shown through the thickness of the composites specimen during a heating stage.

and concluded that the 2-D axisymmetric model yielded sufficiently accurate solutions in less computational time compared to the 3-D model for the present study. Figure 4(b) shows an example of the calculated temperature distribution at a certain time during laser heating. Red arrows indicate the laser heat flux. For the convective boundary conditions, all front, back, and side surfaces were assumed to be under a natural cooling condition with a heat transfer coefficient of h = 5 W m−2 K−1 . The radiative boundary condition was also applied to these surfaces for atmospheric radiation. While most of the

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Table 1: Material properties of T650/5250-4 composite laminates.

Density (ρ)

1550 [kg m−3 ]

Heat capacity (Cp )

Eq. 2 [J kg−1 K−1 ]

Planar thermal conductivity (kx )

3.9 [W m−1 K−1 ]

Through-thickness thermal conductivity (kz )

Eq. 4 [W m−1 K−1 ]

material properties (density, heat capacity, and through-thickness thermal conductivity) of the T650/5250-4 laminated woven fabric composites were experimentally measured as detailed earlier, it was difficult to measure the in-plane thermal conductivity (kx ) of the anisotropic composite laminates. In the absence of measured value, a constant value of kx was assumed for the whole temperature range of this study. It was found that kx = 3.9 W m−1 K−1 results in a good agreement with the experimental data under all conditions considered in this study. The material properties of the T650/5250-4 composites, and boundary- and surface-condition model parameters for the numerical analysis are summarized in Tables 1 and 2, respectively. The model and methodology were implemented into a commercial FE software package, ANSYS, for conductive, convective, and radiative heat transfer to predict the temperature evolution during both heating and cooling stages. The predicted temperatures were compared with the experimental data measured with the IR cameras (front and back surfaces) and embedded thermocouples (midplane of the composite laminates).

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Table 2: Boundary- and surface-condition model parameters for numerical analysis.

Heat transfer coeff. (Front surface) (hfront )

5 [W m−2 K−1 ]

Heat transfer coeff. (Back surface) (hback )

5 [W m−2 K−1 ]

Heat transfer coeff. (Side surface) (hside )

5 [W m−2 K−1 ]

Surface absorptance (α)

0.91

Radiative emittance (ε)

0.89

Initial temperature (Tinit )

∼23.7 [◦ C] / 75 [◦ F]

Ambient temperature (Tamb )

∼23.7 [◦ C] / 75 [◦ F]

4. Results and Discussion The experimental data measured by the IR cameras and embedded thermocouples are denoted as IR and ETC, respectively. The computational prediction results, using finite element analysis, are denoted by FE. The temperature evolution of the front surface and back surfaces are shown in Figure 5(a) and (b), respectively. Figure 5(c) shows a temperature difference map between the front and back surfaces, which were enabled by coregistering front and back surface images and calculating the temperature drop from the front to the back surface. For graphical simplicity, the spatial variations for the remainder of the manuscript were reported only on four straight profile lines, as shown in Figure 5(d). One of the profile lines from specimen PMC-1 irradiated at the laser power of 13.8 W for 60 s duration is shown in Figure 5(e) and (f) for front and back surface temperatures, respectively. Axes labeled with “Time Relative to Laser On (s)” and “Distance Across Profile (mm)” indicate the temporal scale during 0 to 300 s

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and spatial scale along the Profile 1 line, respectively. It was noticed from Figure 5(e) and (f) that a shallow and narrow ’valley’ was observed near a point at 50 mm, which is near the center of the laser beam. This valley was an artifact caused by the lower emissivity of a silver marker dot used to designate the center of the specimen, as shown in Figure 1(a). In the temperature difference images in Figure 5(c), the artifact is pronounced with two dots (one blue dot on front side of sample and one red dot on back side of sample). Not plotted here, it was found that nearly identical plots were drawn for the other three profile lines, indicating the axisymmetric nature of the heat transport phenomena subject to the circular flat-top laser beam. In the remainder of the manuscript, 2-D plots from the post-processed IR camera data overlay all 8 resulting radial plots from these four axisymmetric profile lines. Figure 6(a) shows the experimental data measured with the front-surface IR camera at a radial location of r = 5 mm during the rise and fall of the temperature on the front surfaces of the PMC-1 specimen exposed to the four laser powers (5, 7, 10 and 13.8 W) for two durations (60 and 120 s). Dotted and solid lines indicate the 60 and 120 s durations, respectively. At each power level, 120 s-duration curves (solid lines) follow nearly an identical heating path with 60 s-duration curves (dotted lines), indicating the repeatability of the test at a given laser power. Figure 6(b) compares the front-surface temperature evolution measured on PMC-1 (solid lines) and PMC-2 (dotted lines), irradiated at the four laser powers for 120 s duration. Slight differences between the two specimens might be attributed to differences in the specimen thicknesses (3.109 mm vs. 3.145 mm). Given the experimental repeatability

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Figure 5: (a) Front surface temperature, (b) back surface temperature, and (c) temperature difference at three time steps measured by IR cameras on PMC-1 specimen subjected to 13.8 W laser power for 60 s. (d) A schematic of the four defined, radially symmetric profile lines utilized for referencing temporal and spatial evolution of temperature. 3-D plots of (e) front and (f) back surface temperature evolution along a single line, i.e., Profile 1.

between specimens, only specimen PMC-1 was considered in the subsequent comparison study between the numerical prediction and experiment data. Figure 7 shows the comparison of the front and back surface IR measurements (lines) and the FE model prediction (markers) at three radial (r) locations during the heating and cooling stages of the PMC-1 specimen exposed to 13.8 W for 120 s. Again, the 4 profile lines in Figure 5(d) yield a total of eight IR camera radial measurement results. The difference between front- and back-surface temperatures is considerable (40◦ C and 34◦ C) at both r = 5 mm (circular markers) and 12.37 mm (triangular markers), which are inside the beam profile (i.e., beam radius of r = 15 mm). However, at r = 19

Figure 6: Experimental data measured with a front-surface IR camera at 5 mm away from the center of the irradiated area subject to four laser powers. (a) PMC-1 specimen with laser durations of 60 s (dotted lines) and 120 s (solid lines). (b) Comparison of measurement from PMC-1 (solid lines) and PMC-2 (dotted lines) specimens. An experimental data set representing each power and duration is actually an overlap of eight radial profiles.

19.64 mm (square markers), which is outside the beam profile, the difference becomes negligible (1.9◦ C). For the entirety of the laser exposure and postexposure cooling stages, the simulation results demonstrate good agreement with experimental measurements on both surfaces. Figure 8(a) is a montage of FE simulation images of temperature evolution in a radial cross section of the composite specimen during heating and cooling stages. Note that the laser was turned off at 120 s. The top and bottom of these cross-sectional images represent the front and back of the composites, and can be tied to the radial temperature distributions determined experimentally. Figure 8(b–g) show the comparison of the IR measurements (lines) and the FE model prediction (markers) at distinct time steps (10 – 180 s) after the onset of laser irradiation. Figure 8(b,c,d) depict front surface 20

Figure 7: Comparison of predicted (a) front-surface and (b) back-surface temperature history (markers) with experimental data (lines) measured with PMC-1 specimen via IR cameras at 5, 12.37 and 19.64 mm away from the center of the irradiated area subject to a 13.8 W laser power for 120 s. An experimental data set representing each radius is actually an overlap of eight radial profiles. Both figures share the same legend.

temperatures, and Figure 8(e,f,g) show back surface temperatures. The abscissa indicates the radial position along the four profile lines with respect to the center point. The FE predictions and the experimental results were taken from the PMC-1 specimen irradiated at 13.8 W for 120 s. For the entirety of the radial distance, the simulation results were in excellent agreement with experimental measurements on both surfaces for all laser exposure durations for both heating and cooling stages. Temperature dips near the center point (r ≈ 0) are caused by the lower emissivity silver marker dot. The FE simulations were also conducted for all experimental laser powers. Figure 9 compares the IR measurements (lines) and FE model prediction (markers) for experiments subjected to four laser powers (5, 7, 10 and 13.8 W). The comparisons were made for a point-wise temperature rise and

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Figure 8: (a) A montage of FE simulations of cross-sectional temperature evolution at 10 s intervals for PMC-1 irradiated with 13.8 W for 120 s. Comparison of (b,c,d) front surface and (e,f,g) back surface temperature history (markers) with experimental data (lines) measured via IR cameras along radial positions with respect to the center of the beam profile. (b,e) Heating stage during 10 – 60 s; (c,f) heating stage during 70 – 120 s; (d,g) cooling stage during 130 – 180 s. An experimental data set representing each time is actually an overlap of eight radial profiles.

fall at the radial position of r = 12.37 mm on the PMC-1 specimen irradiated for both 60 s (dotted lines) and 120 s (solid lines). For all cases of laser power levels and exposure times, the simulation results were in good agreement 22

with experimental measurements on both front and back surfaces, as shown in Figure 9(a) and (b), respectively. Figure 10(a) compares the ETC measurements (lines) and FE model prediction (markers) at the radial position of r = 12.37 mm, where the closer thermocouple is located in the middle plane of the PMC-1 specimen irradiated at the four laser powers (5, 7, 10 and 13.8 W) for both 60 and 120 s. For all four power and two duration cases, the simulation results exhibit slightly higher temperatures than the ETC measurements. The discrepancy might be caused by a local cooling effect of the metallic thermocouple tip. Other possible reasons might be slight misalignment of the center of the laser beam or distortions in the spatial calibration of the X-ray images, so that the ETC locations identified by the X-ray images might be inaccurate. To assess the impact of uncertainty in the location of the ETC, various radial positions were selected and compared with the ETC data. Figure 10(b) shows the comparison when the radial position was set as r = 14.0 mm. For all laser powers and all durations tested, the simulation results at this selected radial position were in good agreement with ETC measurements.

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Figure 9: Comparison of predicted (a) front-surface and (b) back-surface temperature history (markers) with experimental IR data (lines) measured with specimen PMC-1 at 12.37 mm away from the center of the irradiated area subject to 5, 7, 10 and 13.8 W laser powers for 60 and 120 s. An experimental data set representing each power and duration is actually an overlap of eight radial profiles. Both figures share the same legend.

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Figure 10: Comparison of predicted temperature history (markers) with experimental data (lines) measured with an embedded thermocouple for midplane temperatures in specimen PMC-1 at 12.37 mm away from the center of the irradiated area subject to 5, 7, 10 and 13.8 W laser powers for 60 and 120 s. Prediction results using (a) 12.37 mm (experimentally identified thermocouple location) and (b) 14.0 mm (proposed thermocouple location). Both figures share the same legend.

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5. Summary and Conclusions Identifying the deleterious thresholds for composite materials based upon temperature evolution will aid in optimizing manufacturing processes and developing fleet management tools. A comprehensive computational and experimental study was conducted to characterize the complex heat transport phenomena and the laser/material interactions of T650/5250-4 plain-woven composite specimens subjected to flat-top CW laser irradiation at four powers and two durations of exposure. The exquisite control of beam profile and power offered by laser heating provided an ideal experimental validation tool for computational simulations. The spatial and temporal evolution of temperatures on both front and back surfaces of composite plates were monitored with the two IR cameras, while the internal temperature evolution was monitored with embedded thermocouples. Full 2-D temperature information and the ability to co-register front and back surface images in 3-D space enabled the in situ monitoring of through thickness temperature differences. Uncertainty in embedded thermocouple junction locations and/or interfacial resistance only reinforced the utility of such full field temperature measurements in generating experimentally validated models. The computational model predictions of 3-D temperature evolution using the nonlinear transient FE simulations conducted in ANSYS were in good agreement with both the heating and cooling stages for all input powers and all durations considered.

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6. Acknowledgments This work was financially supported by U.S. Air Force Research Laboratory (AFRL), Materials and Manufacturing Directorate with a contract number FA8650-17-D-5407. The authors thank Sarah Izor and Josh Kennedy for providing experimental thermal property measurements, and the Optical Measurements Facility team for providing spectral reflectance measurements. 7. References [1] Boeing Company. http://www.boeing.com/commercial/787/by-design/#/ advanced-composite-use [2] M. Sparkes, W. M. Steen, “Light” Industry: An Overview of the Impact of Lasers on Manufacturing, in: J. Lawrence, Advances in Laser Materials Processing, Second Edition, Woodhead Publishing, 2018, Ch. 1, pp. 1–22. [3] C. F. Cheng, Y. C. Tsui, T. W. Clyne, Application of a three-dimensional heat flow model to treat laser drilling of carbon fibre composites, Acta Materialia 46 (12) (1998) 4273–4285. [4] P. J¨aschke, V. Wippo, S. Bluemel, R. Staehr, H. Dittmar, Laser Machining of Carbon Fiber-Reinforced Plastic Composites, in: J. Lawrence, Advances in Laser Materials Processing, Second Edition, Woodhead Publishing, 2018, Ch. 6, pp. 121–152.

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