THERMOPLASTIC PULTRUSION DEVELOPMENT AND CHARACTERIZATION OF RESIDUAL IN PULTRUDED COMPOSITES WITH MODELING AND EXPERIMENTS KHONGOR JAMIYANAA

Transcription

1 THERMOPLASTIC PULTRUSION DEVELOPMENT AND CHARACTERIZATION OF RESIDUAL IN PULTRUDED COMPOSITES WITH MODELING AND EXPERIMENTS by KHONGOR JAMIYANAA UDAY VAIDYA, COMMITTEE CHAIR SELVUM PILLAY HAIBIN NING A THESIS Submitted to the graduate faculty of The University Of Alabama At Birmingham, in partial fulfillment of the requirements for the degree of Master Of Science BIRMINGHAM, ALABAMA 2014

2 THERMOPLASTIC PULTRUSION DEVELOPMENT AND CHARACTERIZATION OF RESIDUAL IN PULTRUDED COMPOSITES WITH MODELING AND EXPERIMENTS KHONGOR JAMIYANAA MATERIALS SCIENCE AND ENGINEERING ABSTRACT Pultrusion processing is a technique to make highly aligned fiber reinforced polymer composites. Thermoset pultrusion is a mature process and well established, while thermoplastic pultrusion in still in its infancy. Thermoplastic pultrusion has not been well established because thermoplastic resins are difficult to process due to their high viscosity. However, thermoplastic resins offer distinct advantages that make thermoplastic pultrusion worth exploring. The present work centers on developing a method to design and validate a die for a thermoplastic pultrusion system. Analytical models and various software tools were used to design a pultrusion die. Experimental measurements have been made to validate the models. One-dimensional transient heat transfer analysis was used to calculate the time required for pre-impregnated E- Glass/Polypropylene tapes to melt and consolidate into profiled shapes. Creo Element/Pro 1.0 was used to design the die, while ANSYS Work Bench 14.0 was used to conduct heat transfer analysis to understand the temperature profile of the pultrusion apparatus. Additionally Star-CCM+ was used to create a three-dimensional fluid flow model to capture the molten polymer flow inside the pultrusion die. The fluid model was used to understand the temperature of the flow and the force required to pull the material at any given temperature and line speed. A complete pultrusion apparatus including the ii

3 die, heating unit, cooling unit, and the frame has been designed and manufactured as guided by the models, and pultruded profiles have been successfully produced. The results show that the analytical model and the fluid model show excellent correlation. The predicted and measured pulling forces are in agreement and show that the pull force increases as the pull speed increases. Furthermore, process induced residual stress and its influence on dimensional instability, such as bending or bowing, on pultruded composites was analyzed. The study indicated that unbalanced layup can produce asymmetrical residual stress through the thickness and causes the part to bow. Furthermore, the residual stress through the thickness was mapped with excellent accuracy. A design of experiments around the processing parameters indicated that increase in pull speed or decrease in die temperature increased the residual stress within the part. iii

4 DEDICATION I would like to dedicate this work to my mom and dad. iv

5 ACKNOWLEDGEMENTS I would like to express my most sincere gratitude to those who made this research possible. I would like to thank my mentor, advisor, and teacher, Dr. Uday Vaidya, who guided me, pushed me to challenge myself through this journey, and has thought me so much about the composites industry. Your guidance was always appreciated. I would also like to thank Dr. Brian Pillay and Dr. Haibin Ning who was always there whenever I needed assistance or advice, thank you for giving me your time and encouragements. I want to extend my gratitude to Dr. Balaji Thattaiparthasarathy and Andy Grabany who were always willing help me process and always there whenever I ran into trouble. Thank you for being so selfless. Furthermore, I also need to thank Vernon Merchant, who was always pleasure to work with, for helping me wire the electronics. I definitely owe a huge thank you to Dr. Roy Koomullil and his graduate student Yangyang Hu who dedicated a lot of their time helping me with my models. Additionally, I need to thank my former bosses at Owens Corning who were so kind to let me continue my research during my internship. Thank you Marc Borowczak and Dr. Amol Vaidya for giving me so many invaluable experiences. I also need to thank Dr. Kevin Spoo for sharing his sage wisdom and unending knowledge about pultrusion, I learned so much from our time. Thank you to Peter Pfaff and Glasforms for donating their pultrusion line. Lastly, I must thank all the graduate students in my department who were always so supportive and for making this whole experience a joyful one. Thanks you. v

6 TABLE OF CONTENTS ABSTRACT... ii DEDICATION... iv ACKNOWLEDGEMENTS...v LIST OF TABLES... ix LIST OF FIGURES...x 1 - INTRODUCTION AND LITERATURE REVIEW PULTRUSION APPARATUS Supply and Guidance System Pre-Heating Sections Pultrusion Die Pulling Mechanism OBJECTIVES Objective #1: Design and modeling: Objective #2: Manufacturing the die: Objective #3: Measure the pulling force: Objective #4: Characterization of Residual Stress: MATERIALS DESIGN AND MODELING Analytical Model Methodology vi

7 5.1.2 Results and Discussion (Analytical Model) CAD and FEA Model Methodology Results and Discussions (CAD and FEA Model) CFD Model Methodology Results and Discussions (CFD Model) MANUFACTURE THE DIE Machining the Die for Thermoplastic Pultrusion According To the Models Methodology OBJECTIVE #3: PULL FORCE MEASUREMENTS: Pull Force Experimental Validation Methodology Results and Discussion (Pull Force Experimental Validation) Influence of Processing Parameter on Pull Force Methodology Results and Discussions (Influence of Process Parameter on Pull Force) OBJECTIVE #4: PULL FORCE MEASUREMENTS: Materials Unbalanced Layup Methodology Results and Discussions (Unbalanced Layup) Residual Stress through the Thickness Methodology Results and Discussions (Residual Stress through the Thickness) Process Relation of Residual Stress Methodology Results and Discussions (Process Relation of Residual Stress) vii

8 9 - CONCLUSIONS...53 REFERENCES...55 viii

9 LIST OF TABLES Tables Page 1 - Thermal properties of the E-Glass/Polypropylene system (35) Processing conditions for each sample. Die temperature and pull speed were varied to understand the process to fiber strain relation Comparison of maximum bow and maximum residual stress of all the samples ix

10 LIST OF FIGURES Figure Page 1 - Thermoset pultrusion process Thermoplastic Pultrusion Process Thermoplastic pultrusion apparatus. The material gets pulled through a guiding channel where it enters the pre-heating zone before entering the die. Inside the die, the materials melts, consolidates, and solidifies to take form of the die cross-section. The system is pulled by pneumatic clamps before getting cut at the desired length Creel and guidance system. Guidance system is used to align the tapes and prevent them from twisting or tangling The first pre-heating section forward of the heating die. It is used to reduce the processing period by raising the temperature of the material before it enters the heated forming die Three-point perspective of the pre-heater positions. The pre-heaters (shown in red)are placed in the center to heat the material from the inside-out. (Diagram not to scale.) Guidance plates placed at the front, middle, and the end of the preheater. The tapes move closer together as it goes through each guide plate. (Diagram not to scale.) Second pre-heater. Placed just in front of the heated die to preserve the heat as the material enters the die. (Diagram not to scale.) Pultrusion modeled as a transient heat transfer of a plane wall, with thickness L and initial temperature T i, in contact with constant surface temperature T s where the surface temperature is greater than the initial temperature CAD drawing of the pultrusion die. The die was designed to process a 25.4mm by 4.3mm rectangular bars and has a 5 o taper for 95.25mm. It was made of two symmetrical halves 76.2 mm wide and mm thick x

11 11 - Steady state heat transfer analysis. Only ½ scale model was used due to symmetry. The pultrusion die, along with the heating and the cooling unit, rests on two steel frame members. Natural convection was applied to these surfaces, shown in yellow Meshing created using ANSYS WB. The area of interest, the die cavity, was created with a more refined mesh Temperature profile along the length of the die on the inside surface, as predicted by the FEA heat transfer model. Scale drawing of the temperature contour plot of the die is superimposed Heat transfer analysis using ANSYS WB. Three 300 W cartridge heaters positioned 76.2mm away from each to create the required temperature profile to fully melt the material while pultruding at the highest speed Die cavity meshed for CFD model (above). The cavity walls were assigned the same temperature profile created from the heat transfer FEA model (below). Pull direction left to right Flow characteristic for various types of fluids Viscosity map of Polypropylene at various temperatures and shear rates. The expected shear rate for pultrusion processing also shown. The shear rate range was determined by dividing the expected pull speeds by the height of the thickness Phase change process in the center plane where the material melts, consolidates, and cools inside the die to take form of the die profile at lowest die temperature and the highest pull speed. Pull direction left to right Shear stress along the die cavity walls. Pull direction left to right Prediction of pull speed versus pull force using CFD model. The pull force increases from the increased shear rate Pultrusion die being machined in on a CNC mill. The die was made of mild steel for ease in machining Floating pultrusion system. The forming die is fixed to the front frame (in blue) and the front fame is simply supported on the pultrusion xi

12 frame (in gray) which is fixed. Load cells are positioned in the path of the front frame that transfers the force Pull force measurement over time. The force was measured using load cells while all the parameters were held constant Pull force comparison between the predicted model and experimental results. The experiment was conducted by incrementally increasing the pull speed and while holding the die temperature steady at 185 o C and measuring the pull force. The prediction is based on a CFD model Polyester veil tape was introduced to one side to create an asymmetrical, unbalanced layup Comparison between two samples produced at the same processing condition. The top sample had a balanced layup, the bottom sample had unbalanced layup with a polyester veil tape on the top surface while the rest of the sample consisted of glass rovings. The asymmetric layup of the bottom sample produced a bowed part Predicted residual stress field through the thickness (1) Depicts the residual stress through the thickness of the flat pultruded part. (2) Portion of the thickness was machined off with a surface grinder. (3) The resultant force is offset from center horizontal axis and the residual stress is no longer balanced with respect to the horizontal axis. (4) To relieve the unbalanced, the part bows and the deflection of bow was measured to calculate the stress at various locations throughout the thickness Progression of the bow as the sample was relieved of its residual stress Residual stress through the thickness of a 0.125" thick flat pultruded part The top surface machined off at increments using a horizontal surface grinder to relieve the locked strain Experimental set up for deflection measurement. The sample is fixed at one end and the deflection is measured at the other end with a microscopy tool xii

13 33 - Comparison of the residual stress through the thickness of a 0.125" thick flat pultruded part. One pulled at a Low speed of 1ft/min, the other pulled at a High Speed 4ft/min Comparison of amount of bow between two identical samples, one pulled at 4ft/sec (left) and another pulled at 1ft/sec (right) Comparison of the residual stress through the thickness of a 0.125" thick flat pultruded part. One pulled at a High Temp of 1ft/min, the other pulled at a Low Temp xiii

14 1 - INTRODUCTION AND LITERATURE REVIEW Pultrusion is a processing technique to produce fiber reinforced polymer (FRP) composite of a profiled shape by the use of a temperature controlled die. The FRP material is pulled through a heated die where the matrix cures and assumes the shape of the die cavity. The material is continuously pulled to produce shaped profile parts of constant cross-section at almost any desired length. There are wide ranges of structural and cosmetic applications in transportation, construction, and marine sectors that use pultrusion. Pultruded composite possess high strength-to-weight ratio, corrosion resistance, thermal properties, and electrical insulation, while offering cost effective, rapid, and efficient method of processing (1) (2) (3) (4) (5). Both thermoplastics and thermosets can be pultruded, although their processing techniques differ slightly. In thermoset pultrusion, continuous fibers are typically pultruded. Mats can be included in the laminate to provide transverse stiffness or improved surface finish. The fiber rovings are organized in creels and strategically fed through a guidance system to obtain optimum alignment. The rovings are saturated with the resin by traveling through a resin bath. There the rovings run through series of "S" guides that facilitate full wet out by spreading the filaments apart. Resin can be injected into the die in place of a resin bath but full impregnation may be difficult to achieve in large profiles (6). Prior to entering the die, the rovings run through a series of forming guides or bushings that removes excess resin. The guides are also used to control the weight fraction of the composite. Once inside the die, heat is used to initiate the curing of the 1

15 resin which is accompanied by an exothermic reaction. The cured part is of a shape of the cross-section of the die and it is continuously pulled throughout the process. A cut off saw located at the end of the process is used to cut the formed part at the preferred length. The schematic in Figure 1 illustrates a typical thermoset pultrusion process. Guide Device Cut Off Saw Heated Forming Die Pullers Pullers Fiber Reinforcement Supply Resin Bath Pull Direction Figure 1 - Thermoset pultrusion process. Thermoset pultrusion is a mature process and it is available commercially. The advantage of pultruding with a thermoset matrix (epoxy, vinyl ester, polyesters) is that the processing is much easier than a thermoplastic matrix. This is due to the lower viscosity that thermosets possess which makes it easier to fully wet out the reinforcing fiber and does not require excess pulling force. Thermosets also offer higher thermal and chemical stability and makes them suitable for application where thermal and chemical resistance is needed (7) (8). Thermoplastic pultrusion in general is a relatively new process and there is limited work on thermoplastic pultruded composites. The process is complicated by the high resin viscosity of thermoplastics that are usually two to three orders of magnitude higher than thermoset resins (9). This high viscosity makes it difficult to achieve full 2

16 saturation. Thus, in place of a resin bath, a pre-heating section is usually implemented to improve the process by raising temperature of the pre-forms almost to its melting temperature before it enters the die. Preheating the material shortens the time required to melt the material and improves the efficiency of the die in the process. Furthermore, unlike thermoset pultrusion, the thermoplastic matrix is cooled at the end of the die to solidify and maintain the die shape. It is not uncommon to have two or more separate dies where one is used to melt and the other to cool the composite (10). The schematic for a typical thermoplastic pultrusion process is illustrated in Figure 2. Guide Device Forming Die Cut Off Saw Pre-heating zone Heated Portion Cooled Portion Pullers Pullers Fiber Reinforcement Supply Pull Direction Figure 2 - Thermoplastic Pultrusion Process The high resin viscosity, the non-newtonian resin flow, resin solidification and melting, and possibility of more than one die adds to the complexity and make thermoplastics resin more difficult to process. Nevertheless, thermoplastic matrix offer higher facture toughness and damage tolerance than their thermoset counterparts, which make them ideal for some structural applications. Thermoplastics are also recyclable, can be locally welded, and easy to maintain when compared to thermoset resins. Furthermore, from health and safety perspective, thermoplastic resins do not require special handling 3

17 or ventilation system for styrene emission that is produced by polyesters (8) (10) (11) (12) (13) (14) (15). There is a general lack of well established standards or methods of designing pultrusion lines with thermoplastic matrices. This study attempts to develop a method to determine an optimal design for a thermoplastic pultrusion die based on temperature of the die and pull speed. Previous works in the area includes work from Åström et al who laid the groundwork for thermoplastic pultrusion modeling (16). In their study, Åström et al formulated simplified analytical models to predict the temperature and pressure distribution along the length of the die within a thermoplastic composite as it travels through the pultrusion line. A steady state heat transfer model was used to predict the temperature of the material as it traversed through the die. A one-dimensional matrix flow, non-newtonian matrix viscosity and temperature dependent matrix density pressure model was considered. The work was verified with experimental measurements using pre-impregnated tapes with reasonable accuracy. However, the measurements only recorded the data from a single point within the cross-section and did not represent the conditions through the entire cross-section. It was noted that the repeatability of the measurements of internal pressure were poor and required further investigation (17). In a separate work with pre-impregnated tapes, Åström developed a design criteria for thermoplastic pultrusion aided with experimental measurements. He deduced that time at/or above the molding temperature and pressure of the matrix were two quintessential parameters to improve material consolidation. This was accomplished by reducing the pull speed, increasing temperature gradient of the material to induce more 4

18 pressure from thermal expansion, and increasing the polymer viscosity to increase the pressure by lowering the die temperature (18). Many of other previous works on the topic of thermoplastic pultrusion include process design. Devlin et al developed a pultrusion apparatus for thermoplastic matrices with various different methods of pre-heating with a roll-forming die. The authors found that using an oven with infrared heaters with medium wave emitters in combination with hot air oven heated the pultruding material most efficiently. Devlin et al also set a design of experiments on the pull speed and found that slower speed parts resulted in higher mechanical properties. This was attributed to longer time in the oven, which allowed the materials to attain proper temperature before being formed (19). Several material forms exist for thermoplastic pultrusion. One of the most common material forms used is commingled tows, which consists of the reinforcement fiber and the resin in a fiber form mingled together. Their high durability and flexibility makes it ideal for pultruding complex shapes. They are also more cost effective and efficient than pre-impregnated pre-forms. However, since fibers saturation takes place inside the die, the potential for voids to develop exists. Consequently, the mechanical properties would reduce as void content increases. Since impregnation occurs during the process with commingled yarns, the degree of impregnation is one of the main concerns. Kim et al developed an analytical model to predict the amount of impregnation. The model consists of two sub-models, namely, microscopic impregnation and macroscopic resin flow. The microscopic impregnation model utilized Darcy's law for resin flow through porous medium, or the agglomerated fibers. The macroscopic resin flow model was done by modeling the backflow from 5

19 pressure and excess resin. Kim compared the models to experimental work and obtained satisfactory results. Kim concluded that agglomeration size was the most important parameter for achieving proper impregnation. Kim's models assumed that the matrix had a Newtonian resin flow, which made the models accurate only at slower pull speeds. However, the model would lose accuracy as the pull speed increased (20). Another common form of material for thermoplastic pultrusion is preimpregnated tapes. Impregnating the fibers before the process is one way of resolving the partial saturation issue. For the present study, hot melt impregnated E- Glass/Polypropylene (PP) tapes were used as the starting materials for thermoplastic pultrusion. The impregnation process was done by pulling virgin fibers through a die filled with the resin. The die was attached to a single screw extruder constantly injecting molten resin into the die. The fiber ran through a series of pins inside the die to ensure that the yarn were fully spread to help full fiber wet out. Once the fibers exit the die, they were guided through a series of scrapers. The scrapers were used to control the fiber weight fraction by scraping off the excess resin. The finished product, which exits in a shape of a 0.5 inch wide rectangular tape, is rapidly cooled to maintain the shape and collected on a spool. Much like many other processing techniques, pultrusion is not exempt from common problems faced by manufacturers. Parts develop residual stresses that lead to premature failure or lose their dimensional tolerance from the process conditions that are unique to pultrusion (21) (22). The shape of the die, drag forces in the die, pre-former alignment, resin bath design, creel tension, and puller pad alignment are some factors unique to pultrusion that can dictate part quality in addition to resin shrinkage, fiber 6

20 orientation, and uneven heat treatment, which are typically found in composite manufacturing (23) (24). Prior work on the topic focused on using analytical models to predict residual stress from resin shrinkage, cure kinetics, and stress generated from varying coefficient of thermal expansion (CTE) between the fiber and the resin (25). Although these approaches provide insight into residual stresses, they only consider the out of plane stresses and neglect the in-plane residual stress generated from fiber strain. Attempts to measure local fiber strain during pultrusion with Fabry Perot tube optic sensors have been largely unsuccessful (26). The sensors often failed from the harsh environment inside in the die but were able to measure positive strain before the sensors failed. Finite element analysis (FEA) models have also been used to optimize the process in terms of cure kinetics and resin shrinkage but lacked data pertaining to mechanical properties (27) (28) (29) (30). There is a general lack of research regarding residual stress of pultruded parts. The present paper addresses the need to understand how residual stress is formed, how much is formed, how it affects ultimate tensile strength, and how it can induce dimensional instability with accurate experimental measurements of residual stress through the thickness of a pultruded bar. Fiber strain is one of the dominant factors that can induce residual stress and dimensional instability. It occurs when the wet fiber bundle enters the die, the resin and the fiber experience slight expansion from the heat and the excess resin squeezes out. This creates significant amount of hydrodynamic action and resistance that causes drag and ultimately strain the fibers in tension. Additionally, the fibers that are in contact with the die surface experience friction that imposes shear force on the part. When the resin 7

21 cures, the strain caused by drag is captured and locked within the fibers and residual stress within the part is formed (31). These phenomena are magnified for thermoplastic polymers since the viscosity for thermoplastics can be two to three times greater in magnitude than thermosets (16). The higher viscosity increases the drag that facilitates more fiber strain. Additional tension can be formed throughout the pultrusion apparatus if the subsystems are not perfectly aligned. Pre-form misalignment, friction from the guidance plates, and friction from the fiber roving traveling through resin bath can increase the tension before the fibers even enter the die. Creel tension differences can often be observed from one glass fiber sizing to another. After the part exits the die, the puller pads must also be in alignment with the die in order to avoid additional tension. If the puller pads are slightly higher or slightly off to one side, the part will bias towards that particular direction and will create more friction in the die (31). Once the pultruded part moves past the pullers and cools down to room temperature, the tension is relieved and the fibers tend to return to a stress free state. Bowing occurs if the fibers are strained asymmetrically. This can occur in couple of different ways; when the layup is unbalanced, or when the part shape is asymmetric. When the layup is unbalanced, the strain will not be equally distributed through the thickness. For example, if a flat pultruded part has a tape of unidirectional continuous fiber at the bottom and chopped fiber mat with high resin content at the top, the two sides will have two different strain levels. The continuous fiber, which is under constant tension, will resist the drag force and transfers the drag force to the adjacent fibers at the next level. The chopped fiber however is unconstrained and will not be able to transfer 8

22 the drag force as it will move and glide with the flow of the resin. Consequently, as the pultruded part moves past the puller and cools to room temperature, the two sides with different strain levels will not offset each other. The strain at the bottom side of the continuous fiber will dominate the strain on the top side with chopped fiber, and the part with the unidirectional fibers on the concave side (31). Similar phenomena has been observed not only in pultrusion but in unbalanced laminates as well (23), (32). Unbalanced fiber distribution can occur unintentionally in symmetrical all roving shapes, such as rods, when a significant viscosity reduction occurs in the heated section of the die prior to gelation. This is most prevalent in epoxies as the viscosity reduction allows the glass fiber to settle under the force of gravity ever so slightly. The immediate affect is poor surface finish or loss of an edge on the upper side of the profile. However, bow can also be observed in these situations. This is one of the reasons clay filler is widely used in epoxies to circumvent these problems. When the part shape is asymmetric, similar observations can be made. Consider an L-shaped part with a balanced layup that is asymmetric along both horizontally and vertically neutral axis. As the part is pultruded, the fibers will be strained from the process. However, since the part is asymmetric, there will be more strain on one side than the other, with respect to both horizontal and vertical neutral axis. This strain unbalance will result in fiber bowing towards to one side as the part tends to recover to a strain free state. 9

23 2 - PULTRUSION APPARATUS Supply and Guidance System The pultrusion line utilized for the present research was setup to process hot melt impregnated thermoplastic tapes, as shown in Figure 3. It features a series of creels, guidance systems, two pre-heating sections, heating die, cooling zone, and a pneumatic puller. With this setup, the tapes were pulled from the creel into a guidance system, as illustrated in Figure 4. The guidance system helps align the tapes and prevent them from twisting and tangling with one another. If the tapes become twisted and tangled, they have a tendency to increase the pulling force, break inside the die, and result in lower mechanical properties because the fiber alignment will be reduced. Heating and cooling unit Material Supply Guide Pre-heating zone Pultrusion Die Speed Control Puller Figure 3 - Thermoplastic pultrusion apparatus. The material is pulled through a guiding channel where it enters the pre-heating zone before entering the die. The material melts inside the die, consolidates and solidifies to take form of the die cross-section. The material is pulled by pneumatic clamps and cut to the desired length at the end of the line. 10

24 Guide Material Supply Guide Figure 4 - Creel and guidance system. Guidance system aligns the tapes and prevents them from twisting or tangling Pre-Heating Sections The tapes enter the first pre-heating section prior to entering the heating die, shown in Figure 5. The purpose of the pre-heating section was to heat the resin close to its melting point. Doing so allows the material to melt quickly once it enters the heating die, which reduces the overall processing time (19). It was essential however, not to heat the resin too high to cause the resin to melt before entering the die because this will run the risk of resin dripping from the tape and causing thermal degradation. The pre-heater had four 1000W 12 inch long heaters positioned at the center in two rows as shown in Figure 6. The inner most tapes were exposed to the heat first, allowing the pultruded material to heat from the inside out. By heating the center of the material in the pre-heater helped reduce the overall melt period since the center of the material is always last to witness the heat inside the die. The pre-heater has three aluminum guidance plates placed at the beginning, in the middle, and at the end of the pre-heater, as illustrated in Figure 7. The guidance plates 11

25 keep proper tension on the tapes while drawing them closer together progressively at each guidance plate. Pre-heating zone Pre-heating zone Figure 5 The first pre-heating section forward of the heating die. It is used to reduce the processing period by raising the temperature of the material before it enters the heated forming die. Guide (b) Side view Pull Direction (a) Front view Heaters (c) Top Figure 6 Three-point perspective of the pre-heater positions. The pre-heaters (shown in red)are placed in the center to heat the material from the inside-out. (diagram not to scale.) 12

26 Guide Guidance Figure 7 - Guidance plates placed at the front, middle, and the end of the pre-heater. The tapes move closer together progressively through each guide plate. (diagram not to scale.) Pull A second pre-heating section was placed just prior to the entrance of the die to bridge the gap between the first pre-heating section and the heating die. It allows the tapes to transport from the first pre-heating section and enter the die without losing any heat. This second pre-heater has two 1000 W 12in long heaters running along the sides, heating the material evenly from the outside as illustrated in Figure 8. Pre-heating zone (a) Side View Pull Direction Heaters (b) Top View Figure 8 - Second pre-heater. Placed just in front of the heated die to preserve the heat as the material enters the die. (diagram not to scale.) 13

27 2.3 - Pultrusion Die The quintessential piece of pultrusion is the die. The die is used to control the material shape and is the area where the material melts and consolidates. As soon as the material enters the heated die, it begins to melt. The pressure inside the die also increases as a result of decrease in cross-section and also due to thermal expansion of the material. The die had a rectangular cross-section of 25.4mm by 4.3mm and featured a 5 o taper for 89 mm. This taper was used not only to increase the pressure but also to promote resin backflow, both of which facilitate resin consolidation. A chiller was used to cool the die downstream. Once the material reaches the cooled portion, it begins to freeze and take form of the die cross-section. Cooling the material before it leaves the die allows the material to cool evenly and prevent any form of warping Pulling Mechanism The pultrusion apparatus utilizes two reciprocating clamps that continuously pull the material. It is positioned past the cooling section and before the material cut off saw where the material is cut. The two clamps are made of metal and are coated with polyurethane that grips the material when pulling. These clamps moves in a linear motion and take turn pulling the material; one clamp was pulls the material, and the other clamp was moves back. Once the first clamp stops pulling, it retreats and other clamp simultaneously starts pulling forward. The system is powered by 0.75 horsepower electric motor that powers the pneumatic mechanism. The pultrusion line speed can be varied by controlling the motor speed. 14

28 3 - OBJECTIVES The purpose of this research was to establish an innovative technique using computer and analytical modeling to design a die for thermoplastic pultrusion and develop a technique to characterize residual stress in pultruded composites. To achieve this goal the project was separated into four objectives, each with its own sub-objectives. These four objects are: 1 - Design and modeling of the pultrusion die, 2 - Manufacturing the die, 3 - Prediction and measurement of pulling force, and 4 - Characterization of residual stress. The detailed outline for each objective, the research approach, and the results and discussion follow Objective #1: Design and modeling: 1. Develop a one-dimensional analytical transient heat transfer model to design a thermoplastic pultrusion die. 2. Design a die using Computer Assisted Design (CAD) software based on the analytical results and conduct FEA simulation to create a temperature profile for the pultrusion apparatus. 3. Develop a Computational Fluid Dynamic (CFD) simulation to predict the maximum force required to pultrude Objective #2: Manufacturing the die: 1. Machine the die for thermoplastic pultrusion as guided by the modeling results. 15

29 3.3 - Objective #3: Measure the pulling force: 1. Validate CFD model by measuring pulling force experimentally. 2. Set a design of experiments around processing conditions to understand how processing influences the pull force Objective #4: Characterization of Residual Stress: 1. Study the effect of laminate layup on dimensional stability of pultruded parts. 2. Develop a method to measure the residual stress through the thickness. 3. Characterize the effect of process condition on residual stress. 4 - MATERIALS The thermoplastic matrix used for the research was polypropylene with E-Glass (PP-GF) as the fiber reinforcement. The glass fiber was impregnated with the resin prior to being pultruded via hot-melt impregnation method that produced tapes of 60% fiber PP-GF weight fraction. 5 - DESIGN AND MODELING Analytical Model Methodology A thermodynamic approach was adopted to design the pultrusion die. Thermoplastic pultrusion was viewed as a form of energy transfer where heat is put into and taken out of a system, as opposed to chemical reaction (cross linking) that occurs in 16

30 thermosets. This transfer of heat is used to induce phase change in the material. During the process, the matrix changes phase from solid to liquid from the heat. Melting the material allows to consolidate and form into the die cavity. The latter portion of the die is cooled to freeze the materiel to maintain the shape of the die cross-section. With this approach, the die was designed based on the amount of temperature and time required to melt and solidify the material all the way through its thickness depending upon the pulling speeds. To design the die, the pultrusion process was considered as a one dimensional transient conduction heat transfer problem. The thermoplastic material was lumped as a plain wall of thickness L in contact with the constant die surface temperature T s which was held at a higher temperature than the material melt temperature T m. Lumping the pultruding hot-melt impregnated tapes allowed the materials to be treated as a single composite part based on the volume V f or the weight fraction W f. The thermal properties of the part can be calculated using classical laminate theory and using rule of mixtures through Equation 1, Equation 2, and Equation 3 (33) (34). The standard values for conductivity k, density ρ, and the heat capacity c p for E-glass/polypropylene system are compiled in Table 1 (35). The subscript g and pp denote values for E-glass and polypropylene respectively. k 1 W f 1 W f k g k pp Equation 1 f V f r 1 V f Equation 2 17

31 c p c p_g V f c p_pp 1 V f Equation 3 Table 1 Thermal properties of the E-Glass/polypropylene system (35). Material Thermal Conductivity k Density ρ Heat Capacity c p Melt Temperature T m Weight Fraction W f Polypropylene o C E-Glass >800 o C G/PP o C 1 The material entered the die as unconsolidated pre-impregnated tapes, so it was unavoidable to have air gaps between the layers at the opening. The air gaps were considered when calculating thermal properties of the composite since air can act as insulator and resist the heat transfer between two solid interfaces. For this study, air gaps between pre-formed tapes were considered when calculating the thermal properties by modeling the gaps as thermal contact resistance (36). The die profile for the study was simple rectangular 25.4mm wide and 4.3m thick to produce flat symmetrical bars for research purposes. Since the width of the pultrusion profile was nearly six times larger than the thickness, the heat transfer along the width of the profile was ignored. This assumption means that heat from the top or the bottom of the die would transfer to the center of the profile before the sides of the die. The axial 18

32 heat transfer was also neglected, justifying one-dimensional analysis. A study by Pantaleao et al concluded that at speeds around (similar speeds achieved in this research) the anisotropic conduction rates can be neglected. Pantelaeo's model showed very little difference when axial conduction through the fibers was considered versus when axial conduction was not considered (37). Figure 9 illustrates this approach. Figure 9 - Pultrusion modeled as a transient heat transfer of a plane wall, with thickness L and initial temperature T i, in contact with constant surface temperature T s where the surface temperature is greater than the initial temperature. As the material entered the die, the surface that was in contact with the die was considered to be at the same temperature of the die at T s, which is at a higher temperature than the material melting point T m. It was assumed that the pultruded material entered the die at a uniform temperature of T i and since was heated from the top and bottom symmetrically, only half the thickness was considered. The length d of the die was determined by first calculating the amount of time it took the surface temperature T s to transfer through the material thickness and melt the center by using Equation 4 (38): T m T s T i T s 4 n 1 C n exp n 2 Fo cos n L Equation 4 19

33 where T m is the melt temperature at center of the thickness where it is furthers away from the heating source and the coefficient C n is expressed in Equation 5 below, C n 2 n 4sin n sin 2 n Equation 5 the eigenvalues ζ are positive roots of the transcendental equation, shown in Equation 6, n tan n Bi Equation 6 and Biot number Bi, which relates the material s thermal convection h and conduction k terms shown in Equation 7, Bi h L k Equation 7 the Fourier number Fo in Equation 4 was solved using Equation 8, Fo t L 2 Equation 8 where α there is the thermal diffusivity, which was calculated by using the thermal properties of the composite in Equation 9. k c p Equation 9 Once the time t is solved from Equation 4, the length d was be calculated by d t V Equation 10 where V is the operating pull speed of the pultruder. This length is the minimum length required for the die to melt the pultruding material to melt through its entire thickness at the given speed. 20

34 5.1.2 Results and Discussion (Analytical Model) All the characteristic temperature terms on the left side of Equation 4 were known. T s was the temperature of the material at the surface, which was the same temperature of the die. T m term was desired temperature at the specific location. For this study, this temperature was equal to the matrix melt temperature at center of the thickness. Finally T i was the initial temperature of the material when it entered the die. The only term on right side of Equation 4 was the time variable t that described the amount of time it takes the surface temperature T s to heat the center of the material to T m. The eigenvalues were gathered from Incropera (36) and all the other terms were the material thermal properties from Table 1. Solving for t, the time it took to melt composite to melt through its thickness was calculated to be 10 seconds. Knowing the time it takes to completely melt the material, it was found that die length needed to be 14.4 cm with constant surface temperature of 185 o C. This length was considered at the highest pull speed capable by the machine to be If the operating speed were slower, then the required length to melt the material through would likewise decrease since the material will be in the die for longer time. By considering the maximum pull speed available gave certain safety factor since the normal operating speed is not at maximum speed. Once the material has fully melted and consolidated, it is cooled inside the die to maintain the die shape and initiate crystallization. A reverse calculation was made to determine the length of the die for the center of the material to cool from melt temperature to reach the polymer crystallization temperature T c of 100 o C. Using the same maximum speed, the required length was found to be 11.5 cm. 21

35 The final length of the pultrusion die was cm. The overall die length contained an area for securely mounting, area for melting, and cooling. The die was clamped into a 12.7 mm pin that is attached to a steel frame that was allowed to pivot and was simply supported on another steel frame to help support the weight CAD and FEA Model Methodology Following the determination of minimum die length required to process, the die geometry was designed using CAD software Creo Element/Pro, shown in Figure 10. The die had a cross-section of 25.4mm wide and 4.3mm thick and was made of two symmetrical halves. The two halves were bolted together by eighteen 5/16 inch (7.94 mm) socket head cap screw bolts. It featured a 5 o taper for mm into the die to allow sufficient heat transfer. The rationale behind this was to efficiently transfer heat through the material. A larger taper angle would create big air gaps between the tape layers, creating a large resistance to transfer heat through the gaps. Past the tapered section, the die consists of constant cross-section. 22

36 Figure 10 CAD drawing of the pultrusion die. The die was designed to process a 25.4mm by 4.3mm rectangular bars and has a 5 o taper for 95.25mm. It was made of two symmetrical halves 76.2 mm wide and mm thick. To understand the temperature profile of the die when it is heated, Finite Element Analysis software ANSYS Work Bench was used to create a mesh and perform steady state heat transfer model, as shown in Figure 11 and Figure 12. The model simulates the conductive heat transfer from the cartridge heaters and creates a temperature profile for the entire die apparatus. The cartridge heaters and the chiller were assigned constant temperature surfaces. Thermal contact resistance was assigned between mating surface, and natural convection was applied to exposed surfaces. The pultruding composite was modeled by applying a heat load out of the system in the heating zone, while head load was applied into the system in the cooling zone. The heat load was calculated by energy Equation 11, where m rate was the mass flow rate, c p was the specific heat from Equation 3, and ΔT was the total temperature difference between the die surface and the pultruding material. 23

37 Equation 11 The mass flow rate was determined from Equation 12, where ρ was the composite density, V was the operating velocity, A c was the die cross-sectional area. Equation 12 Since the die was symmetrical along the top and the bottom halves, only a half scale model was created to save simulation time. An iterative process was taken to find the appropriate positions for the cartridge heaters to achieve the desired temperature profile along inner walls of the die. The cooling portion of the die, which consists of the same constant cross-section, was cooled by a temperature controlled chiller. The chiller circulated low temperature coolant through an aluminum blocks that was clamped onto the end of the die. Steel frame member Cooling unit Heating unit Pultrusion Die (1/2 scale) Figure 11 Steady state heat transfer analysis. Only ½ scale model was used due to symmetry. The pultrusion die, along with the heating and the cooling unit, rests on two steel frame members. Natural convection was applied to these surfaces, shown in yellow. 24

38 Figure 12 Meshing created using ANSYS WB. The area of interest, the die cavity, was created with a more refined mesh Results and Discussions (CAD and FEA Model) As discussed in the Die Design section, the die surface temperature T s needs to be at 185 o C for 14.4 cm for the material to fully melt through its thickness, and enable pultruding at maximum speed of An iterative design process was conducted to determine the heating power, number of heating units, and the distance between each heating unit. These optimization steps helped achieve the required 14.4 cm length to fully melt and the 11.5 cm length to cool the material, as shown in Figure 13. The heat transfer study showed that in order to achieve this temperature profile, three 12.7mm diameter 300 W cartridge heaters needed to be spaced 76.2 mm apart and heated to 190 o C. Figure 14 shows the heat transfer FEA result with the cartridge heater position within the aluminum heating block and the corresponding heat transfer FEA result. This design also allows the cartridge heaters to be operated independently if different heating zones were needed within the die. The aluminum block was also designed to be separable 25

39 Temperature (C o ) such that the mm block can be disjoined to a mm or a 76.2 mm long block. This can be useful if a more conductive material, such as carbon fiber, is being pultruded, the heating unit can be shortened Die Length (m) Figure 13 - Temperature profile along the length of the die on the inside surface, as predicted by the FEA heat transfer model. Scale drawing of the temperature contour plot of the die is superimposed. 26

40 Figure 14 Heat transfer analysis using ANSYS WB. Three 12.7 diameter 300 W cartridge heaters positioned 76.2mm away from each to create the required temperature profile to fully melt the material while pultruding at the highest speed. Units in of the scale are in degrees Celsius CFD Model Methodology Computational Fluid Dynamics (CFD) is a powerful tool used to simulate fluid flow within the die. It can used to generate velocity vectors, temperature profiles, predict pressure, and capture shear stress fields to name a few. In thermoplastic pultrusion processing, the matrix is heated beyond its melt temperature and converts into molten fluid. This molten fluid is subjected to pulling forces and thus generates fluid dynamic motion inside the die. However, the phase changes in the process, the non-newtonian fluid flow, and the reinforcing fibers make pultrusion a very difficult process to model and few studies have taken advantage of this tool. In this study, a CFD simulation has been developed using Star CCM+ to model the molten fluid flow inside the die. The overall goal of the model was to predict the 27

41 maximum force required to pultrude the impregnated tape. Knowing the maximum force required helps determine the size of the motor to pull the material. The model had three sections: the inlet, the die wall, and the outlet. The inlet was modeled as the zone where the material enters the die. The input parameters had the same flow rate, temperature, thermal properties, and fluid rheology as composite material. The steady state model also considered the effects of enthalpy of melting and solidification, and multiphase mixture fluid behavior. The die wall, which includes the entire surface of the die cavity, was modeled with polyhedral shell elements and was prescribed the same temperature profile found from the heat transfer model from Figure 14. Since the material exiting the die was released into an open environment, the outlet had no back pressure. The modeling parameters can be found in Figure 15. The fluid model only considered the laminar flow regime since the matrix had a high viscosity and the pull speed was relatively low. The flow regime was identified by calculating the Reynolds number of the fluid flowing inside the die with Equation 13 below. Re V D h Equation 13 The D h term known as the hydraulic diameter is a characteristic length used for irregular shapes such as rectangles and discs, and the µ is the dynamic viscosity of the fluid (36). The hydraulic diameter was calculated using Equation 14 D h 4 A c P Equation 14 28

42 where A c is the cross-sectional area, and the term P is the wetted perimeter. From the above equation, the Reynolds number was calculated to be. This was far below the 2300 Re needed to reach turbulent level in a flow through a pipe (36). Outlet Inlet m rate T i Die Wall Figure 15 Die cavity meshed for CFD model (above). The cavity walls were assigned the same temperature profile created from the heat transfer FEA model (below). Pull direction left to right. Modeling the fluid dynamic for a thermoplastic material however is not trivial and a full understanding of the material rheology behavior must be made before modeling. Thermoplastics are different in that they have a non-newtonian flow behavior. Newtonian fluids, such as water, have a linear applied stress and shear strain rate relation, and their viscosity is only dependent temperature and pressure (39) (40). The viscosity relation for Newtonian fluids is defined in Equation 15, 29

43 Shear Stress (Pa) Equation 15 where τ is the shear stress, η is the viscosity, and γ is the shear strain rate. Thermoplastics on the other hand have a shear thinning behavior, where the shear stress and shear strain relation is nonlinear. The shear thinning effect lowers the viscosity when the rate of deformation increases. The ratio between shear stress and shear rate decreases as either the shear stress or the shear rate increases. The flow characteristic of Newtonian and non-newtonian fluids can be seen in Figure 16 (41) (42). This phenomenon occurs because at lower shear rates, the entanglements in the polymer chains are much higher than it is at higher shear rates. At high shear rates, the polymer chains become disentangled and the mobility of the chains increases, and as a result, the bulk melt viscosity is lowered (43). Shear Thinning Newtonian Shear Thickening Shear Rate (s -1 ) Figure 16 - Flow characteristic for various types of fluids. Ostwald and de Waale formed a simple model to represent the shear thinning region in the viscosity versus strain rate curve known as the power-law, as expressed in Equation

44 m( T) n 1 Equation 16 where m is known as the consistency index and n as the power law index. However, at lower shear rates near the Newtonian plateau region, rates typically witnessed in pultrusion processing, the model overshoots and losses its accuracy (43) (44) (45). For this research, the shear rate was calculated by dividing the pull speed by the height of the cross-section. The range was expected to be between 0.42 s -1 to 3.4 s -1. To accurately predict the resin viscosity, a rheometer with parallel-plate oscillating motion was used. The experiment was conducted by holding temperature at constant level and performing a shear rate sweep at different temperature. Figure 17 maps out the resin viscosity at various temperatures and shear rates of interest. A study by Thomasset et al suggests that the shear viscosity of neat PP did not change much when compared to PP with 40% long fiber by volume (46). The values from the graph were inputted into the fluid simulation in an effort to simulate a realistic molten resin flow. The fluid was also assigned the same thermal properties from Table 1 as the composite. 31

45 Dynamic Viscosity (Pa s) Shear rate range for pultrusion 175 C 185 C 195 C 205 C 215 C Shear Rate (s -1 ) Figure 17 - Viscosity map of Polypropylene at various temperatures and shear rates. The expected shear rate for pultrusion processing is also shown. The shear rate range was determined by dividing the expected pull speeds by the height of the thickness. Once the thermal property and the rheology of the fluid were inputted into the software, a study was done to simulate phase change of the matrix, from solid to liquid to solid again as seen in thermoplastic pultrusion process. This was done to predict if the material had fully consolidated through the thickness inside the die. The simulation was done at the most extreme conditions where it is least likely to melt; i.e. at the highest pull speed of and at the lowest operating temperature of 185 o C. The extreme conditions may be explained as follows:- (a) Pulling at the highest speed the material will have the shortest time to melt, and (b) pultruding at the lowest die temperature will mean least amount of heat to melt the material. All other operating conditions will only increase the chance of melting the material more thoroughly. 32

46 Results and Discussions (CFD Model) Consolidation through the thickness. Figure 18 shows the solid to liquid volume fraction at the center plane of the die at half the thickness (2.15mm) where it was furthest away from the heating source. It can be seen that the resin enters the die in a solid state (red) below the melt point of 167 o C, melts towards middle of the die (blue), and solidifies (red) again before exiting. This study corresponded directly with the analytical approach where the material was predicted to have melted through its center and solidified before leaving the die. It also confirmed that the length of the die of 95.25mm was correctly designed based on the set pull speed of and die temperature of 185 o C. Had the pultruding material not reached molten phase at the center plane, it would not have consolidated in its entirety through the 4.3mm thickness. Figure 18 Phase change process in the center plane where the material melts, consolidates, and cools inside the die to take form of the die profile at lowest die temperature and the highest pull speed. Pull direction left to right. 33

47 Effects of processing parameter on die shear stress and pull force. The CFD model was also used to predict the maximum force required to pultrude the material. Considering that the viscosity of thermoplastic matrices are two to three magnitudes higher than thermosets, understanding the amount of power needed to pultrude is essential to building the overall pultrusion line and determining if the die can handle the stress from the process. As shown on the viscosity map in Figure 17, the highest viscosity occurs at the lowest temperature and the lowest shear rate, i.e. slowest pull speed. However, as the pulling speed increases, the force increases in proportion as a result of other phenomena, such as friction force (16). Therefore, pultruding at a lower speed may not produce the highest pulling force. A further study was conducted on the effect of pultruding speed in relation to pull force. This was examined by keeping the die temperature at a constant 185 o C and holding every other parameter constant while incrementally increasing the pull speed. The results are illustrated in Figure 20. The graph indicates that the pull force increases as the pull speed increases. This is a consequence of increased shear stress on the cavity walls. Even though the viscosity decreases slightly from the increased shear rate, the shear rate also causes the shear stress to increase at a high rate. As a result, the force required to pull the material increases when the pull speed increases. 34

48 Pull Force (N) Figure 19 - Shear stress along the die cavity walls. Pull direction left to right Pull Speed (m/s) Figure 20 - Prediction of pull speed versus pull force using CFD model. The pull force increases from the increased shear rate. 6 - MANUFACTURE THE DIE Machining the Die for Thermoplastic Pultrusion According To the Models Methodology After designing and modeling, the die was manufactured from standard mild carbon steel due to its ease of machinability. It was fabricated using a 3-axis CNC-mill as 35

49 shown in Figure 21. The die cavity was sanded from 80 grit sand paper and worked up to 1200 grit (wet). Additionally, the die cavity surface was polished using a polishing grinder to reduce its friction. Finally, 3-4 layers of mold release agent Frekote was applied to the die cavity to help reduce any seizes inside the die when the resin freezes. Typically, pultrusion dies are made from tool steel and have a hardness of R c. They can be chrome plated and diamond polished to improve wear resistance and reduce the die friction. However, problems can arise if the resin seizes within the die, the chrome plating can be stripped off (47). Dies may also be treated with ion nitride to increase their surface hardness (10). As this die was only used for small scale laboratory type of processing, no post manufacturing treatment was done to improve its longevity. A graphite gasket, capable of withstanding high temperatures up 480 o C, was placed in between the top and bottom halves of the die. This was placed to combat any resin leakage that might occur during processing from any machining imperfections that stemmed from sanding by hand. The two halves are bolted together with 18 5/16 inch socket head cap screws, torqued to 45 ft-lb. 36

50 CNC Lathe CNC Lathe Pultrusion Die Pultrusion Die Figure 21 Pultrusion die being machined in on a CNC mill. The die was made of mild steel for ease in machining. 7 - OBJECTIVE #3: PULL FORCE MEASUREMENTS: Pull Force Experimental Validation Methodology The die was assembled onto the pultrusion line such that the pulling force could be recorded while processing. The objective was to measure the maximum required force to pultrude and to validate the CFD model. This was achieved by having a "floating system" where the die mounted to metal bar that was allowed to slide only along the pulling direction. The bar was stopped by load cells which were mounted to frame directly in the path of the sliding bar. The force generated from pultruding was directly transferred onto the load cells that were recorded via Personal Data Acquisition System (Omega OMB-DAQ-56). The illustration for the floating system and the arrangement is shown in Figure

51 The experiment was conducted at 185 o C which was the lowest temperature needed to fully melt the material all the way through its thickness (see Analytical Model in Section 6.1), and it was also the condition when the matrix had highest viscosity (see Figure 17). This processing condition gave the worst possible condition for pull force requirement. Two samples of 3 meters were pultruded to receive an average force value. The measurements were taken at a constant pull speed and temperature. Top View Floating Frame Fixed Frame Pultruded Bar Load Cell Fixed Frame Side View Pultruded Bar Pull Direction Floating Frame Fixed Frame Load Cell Figure 22 - Floating pultrusion system. The forming die is fixed to the front frame (in blue) and the front fame is simply supported on the pultrusion frame (in gray) which is fixed. Load cells are positioned in the path of the front frame that transfers the force Results and Discussion (Pull Force Experimental Validation) The graph has an oscillating trend, which may be due to having dual reciprocating puller pads that take turn pulling the material, as explained in the 3.4 Pulling Mechanism section. Before the two pullers pads switch turn pulling, both are momentarily engaged. 38

52 Pull Force (N) During this period, both pads are temporarily pulling at the same and therefore require less force to pultrude. When one of them disengages and retreats to the starting position, more work is required by the single puller, and the force required to pultrude consequently increases. Had the pultrusion line was equipped with a continuous style belt pullers, the pull force reaction would expect to be relatively flat over time. Figure 23 shows a typical result force measurement. For this particular case, the average force need to pull was 81N while the maximum force required to pull was 103N. The speed at which it was pulled was while the die temperature was held constant 185 o C Time (sec) Figure 23 - Pull force measurement over time. The force was measured using load cells while all the parameters were held constant. 39

53 7.2 - Influence of Processing Parameter on Pull Force Methodology Like the CFD model, design of experiments was set around the pulling speeds to study its influence on the pulling force. The pulling force measurements were conducted under the same condition as the CFD model at the lowest die temperature of 185 o C when the matrix was most viscous. The pull speed was incrementally increased from to A careful approach was taken to insure that the die temperature had reached steady state and had not changed from the increased speed Results and Discussions (Influence of Process Parameter on Pull Force) The experimental result had a very similar trend to the CFD modeling prediction, as shown in Figure 24. Both curves show an increase in the required pulling force as the pull speed increases. This conflicted with the expectation that the pulling force might decrease since the matrix was a shear thinning fluid, the pulling force should decrease since the viscosity was lowering. But since the force increased in both cases, it suggests that the matrix viscosity was not the biggest contributor to the pulling resistance. The results and the model suggest that the drag from friction from the die walls, compaction, and fiber creels have a dominant influence on the process. 40

54 Pull Force (N) Measured Pull Speed (m/s) Figure 24 - Pull force comparison between the predicted model and experimental results. The experiment was conducted by incrementally increasing the pull speed and while holding the die temperature steady at 185 o C and measuring the pull force. The prediction is based on a CFD model. The difference between the model and the measurement might be attributed to several factors. The model does not take into consideration the resistance generated outside of the die. This means that the total friction in the pre-forms and the total weight of the material were not captured by the model, it only predicted the resistance generated inside the die. Furthermore, the model assumed that the material enters as a single unit, when in reality the material was in tape form and has spaces between the layers near the entrance of the die. For the experimental measurements, the temperature of the die was difficult to control and hold steady. There were two thermocouples positioned along the length of the 41

55 die, one towards the front and the second towards back of the die. But the two thermocouples showed that part of the die changed temperature at different rates and the temperature readings were often in slight disagreement. This was especially true when the pull speed increased. The difficulty in temperature control meant that viscosity of the matrix would not be consistent, which in turn would affect the pull force. 8 - OBJECTIVE #4: PULL FORCE MEASUREMENTS: Materials The residual stress experiments were conducted with thermoset matrix. The samples were prepared at a separate facility that did not have a pultrusion line adapted for thermoplastics. Process induced residual stress occurs in the same manner regardless of the matrix type and the experimental method is applicable to both thermosets and thermoplastics pultruded composites. E-CR glass and polyester resin with 70% fiber weight fraction system was the material systems used. The samples were 2 feet long and had a simple rectangular cross-section of 1 x Unbalanced Layup Methodology A simple qualitative experiment was conducted to observe how a flat asymmetrical layup can introduce bowing. A thin tape of polyester veil was introduced on one side to create an asymmetric layup and the finished part was compared to an identical part with a balanced layup without the veil tape. 42

56 Results and Discussions (Unbalanced Layup) To test the effect of the fiber orientation (layup) of the laminate on dimensional stability in pultrusion, a balanced layup of continuous glass fiber roving was pultruded. As expected it produced an ideal sample in terms of dimensional stability, i.e. no bow was formed. But as the process continued, a thin tape of polyester veil was introduced on one side while all other parameters were kept the same, as shown in Figure 25. This created an unbalanced layup where the veil side with the higher resin content transferred less of the drag force into the fibers. The experiment resulted in a part containing a slight curve, the veil on the convex side and the unidirectional fiber on the concave side. This supported the theory that an unbalanced layup will induce a bow; the difference can be seen in Figure 26. If resin shrinkage was the influencing factor for all dimensional changes in pultruded parts, the veil side of the part with much higher resin content might have been expected to shrink more and forcing the veil side of the part to be the concave side. That did not happen. Veil Side Glass Roving Side Figure 25 - Polyester veil tape was introduced to one side to create an asymmetrical, unbalanced layup. 43

57 Thickness Balanced Sample Veil Side Glass Roving Side Unbalanced Sample Figure 26 - Comparison between two samples produced at the same processing condition. The top sample had a balanced layup, the bottom sample had unbalanced layup with a polyester veil tape on the top surface while the rest of the sample consisted of glass rovings. The asymmetric layup of the bottom sample produced a bowed part Residual Stress through the Thickness Methodology This experiment attempted to investigate the extent to residual stress is in a part. The underlying presumption was that the part had a stress gradient through the thickness analogous to that of a velocity gradient of a laminar fluid flowing through a pipe, as illustrated in Figure 27. The fibers in the center of the part presumably have less stress than fibers on the die boundaries. The outer fibers, which are in contact with the die walls, experience more friction than the fibers in the center. The symmetrical nature of the flat die design creates balanced strain field with respect to the neutral horizontal axis. Residual Stress/Strain Figure 27- Predicted residual stress field through the thickness. 44

58 Thicknes Thicknes Thicknes The experiment took advantage of the symmetric stress gradient through the thickness by machining off a certain layer entirely from one side of the surface with a horizontal surface grinder. By machining off a small layer, it created an uneven stress field and the pultruded part had an imbalanced residual stress. To counter the stress imbalance, the part deflected, or bowed, towards the side with the greater stress. The amount of deflection was measured to back calculate the force needed to create the bow and the residual stress in that specific layer. The process was repeated at various locations throughout the thickness to create a one-dimensional map of the residual stress within the part. Figure 28 illustrates this approach and Equation 17 and Equation 18 describes the calculation of the local residual stress through the thickness. Machined Surface Residual Stress Field Residual Stress Field Resultant Part Bows from Stress Offset (1) (2) Force (3) (4) Figure 28 - (1) Depicts the residual stress through the thickness of the flat pultruded part. (2) Portion of the thickness was machined off with a surface grinder. (3) The resultant force is offset from center horizontal axis and the residual stress is no longer balanced with respect to the horizontal axis. (4) To relieve the unbalanced, the part bows and the deflection of bow was measured to calculate the stress at various locations throughout the thickness. Classical cantilever beam theory was used to calculate the internal forces in the part. Cantilever beam theory is used to calculate deflection of a beam when it is loaded at one end and fixed at the other. Similarly for the pultruded sample, using Equation 17 and Equation 18, the deflection of the bow was measured to back calculate the amount of 45

59 force required to deflect the part. The term M represents the moment that is required to cause a beam to deflect δ or to bow. The equation also considers the material's modulus of elasticity E, the moment of inertia I, and the length L of the part. The elastic modulus was calculated by classical laminate theory for this experiment. Equation 18 simply converts the moment into force F by dividing the moment by the material thickness t. Finally, using Equation 19, the calculated force is applied over the area A that was machined off to find the residual stress P at that particular location. Equation 17 Equation 18 Equation Results and Discussions (Residual Stress through the Thickness) Flat 0.125" by 1" rectangular bars were pultruded to quantify the residual stresses produced from pultrusion process. The samples were relieved of their residual stress to induce bow by machining off a section from the top surface. The deflection of the bow was measured to back calculate the residual stress required to create the bow. The residual stress through thickness was mapped by machining off 0.01", 0.03", and 0.06" from the top surface. Figure 29 illustrates the progression of the bow as the residual stress relieved of a part that was pultruded at 4ft/min and a die temperature of 340 o F. Residual stress through the thickness is represented in Figure 30. It can be seen that the center of the flat pultruded part had almost no residual stress and the residual 46

60 Thickness (in) stress increased non-linearly as it moved away from the center. The maximum stress was at the outside surfaces were it was in contact with the die walls. If the residual stress field been even through the thickness the part would not bow when the stress is relieved since the residual stress field would still be symmetrical and uniform regardless of the thickness. Sample o F, 4ft/min 0.01" Machined Off 0.03" Machined Off 0.06" Machined Off Figure 29 - Progression of the bow as the sample was relieved of its residual stress Sample 1-340F, 4ft/min Residual Stress (psi) Figure 30 - Residual stress through the thickness of a 0.125" thick flat pultruded part. 47

61 8.4 - Process Relation of Residual Stress Methodology In addition to quantifying the residual stress, the study also focused on the influence of process parameters on residual stress. This was examined by setting a design of experiments (DOE) around two processing parameters: die temperature and pull speed. The residual stresses were measured at the different set points to study their effects. The temperature range examined were 280 o F and 340 o F, which were two extreme temperature settings of the pultruder. The pull speed ranged from 1ft/min to 4ft/min, which was the slowest and the fastest pull speed. The entire DOE was established using only three sets of samples as summarized in Table 2. Table 2 Processing conditions for each sample. Die temperature and pull speed were varied to understand the process to fiber strain relation. Die Temperature (F) Pull Speed (ft/min) Sample 1 High Temp, High Speed Sample 2 High Temp, Low Speed Sample 3 Low Temp, Low speed The stress in the samples was relieved using a horizontal surface grinder that ground thin layers from the top surface, as illustrated in Figure 31. This produced an immediate bow. First, 0.01" was machined off from one surface to produce bow. As shown in Figure 32, the bow in the sample was measured by fixing one end against a 48

62 straight flat aluminum bar and the deflection at the opposite end was measured using digital microscopy instrument. Microscopy allowed for a nonintrusive way to accurately measure the deflection of the bow. Next, another 0.02" was machined and the bow was measured again. Finally, 0.03" was machined to reach half the thickness of the original dimension or 0.06" and the bow was measured again. Machined surface: 0.01 off top surface, then 0.03, and finally 0.06 Pultruded Part Total Thickness: Figure 31 - The top surface machined off at increments using a horizontal surface grinder to relieve the locked strain. Microscopy Instrument Fixed End Sample Figure 32 - Experimental set up for deflection measurement. The sample is fixed at one end and the deflection is measured at the other end with a microscopy tool Results and Discussions (Process Relation of Residual Stress) Influence of pull speed. To study the effects of process parameters on residual stress, the samples were pultruded at various process conditions and compared. First, the pull speed was examined while all other parameters were kept constant. The two pull speeds were 4ft/min and 49

63 Thickness (in) 1ft/min. The results showed that the residual stress was greatly influenced by the speed. The sample pultruded at 4ft/min had a total residual stress of 314 psi while the sample pultruded at 1ft/min had a residual stress of 24.4 psi. Figure 33 compares the residual stress through the thickness between the two samples. The results indicated that lowering the speed nearly eliminated the bow in the part when the stress was relieved. The part pultruded at 4ft/min had a bow of 0.80" while the part pulled at 1ft/min had a bow of 0.11", as shown in Figure Sample 1-340F, 4ft/min Sample 2-340F, 1ft/min Residual Stress (psi) Figure 33 - Comparison of the residual stress through the thickness of a 0.125" thick flat pultruded part. One pulled at a Low speed of 1ft/min, the other pulled at a High Speed 4ft/min. 50

64 Sample 1-340F, 4ft/min Sample 2-340F, 1ft/min 0.06" Machined Off 0.06" Machined Off Figure 34 - Comparison of amount of bow between two identical samples, one pulled at 4ft/sec (left) and another pulled at 1ft/sec (right) Influence of die temperature. The influence of die temperature on part residual stress was observed by lowering the die temperature from 340 o F to 280 o F and every other parameter were kept constant including the pull speed, which was held at 1ft/min. The residual stress increased from 24.4 psi to 135 psi when the die temperature was lowered from 340 o F to 280 o F. The increased stress is attributed to the longer resin cure time and the extended gelation zone where the extended time as a viscous gel contributes significantly to the fiber tension. The effect of the temperature change is illustrated in Figure 35 where it shows that the sample pultruded at 280 o F had a higher residual stress near the outer surfaces. Table 3 compiles the results from all three scenarios. 51