3
MATERIALS PROPERTIES AND LIFE PREDICTION

The characteristics that make composites, especially glass fiber-reinforced and wood/epoxy composites, suitable for wind turbine blades are low density, good mechanical properties, excellent corrosion resistance, tailorability of material properties, and versatility of fabrication methods. Although glass/vinylester and glass/polyvinyl composites based on hand lay-up have been the most widely used materials so far, many more types of fibers and resins have become available recently. The new carbon fibers are stronger and stiffer, while the new resins provide higher toughness and shorter process cycle time. A number of handbooks are now available to the designers of composites (Lubin, 1982; Engineered Materials Handbook, 1987; Composites & Laminates, 1987).

FIBERS

The most commonly used and lowest-priced fiber is E-glass fiber. Over the past several years, however, many new fibers have become available. The commercially available fibers and their typical properties are listed in Table 3-1.

While E-glass fiber is most widely used in wind turbine rotor blades mainly because of its low cost, carbon fibers are the fibers of choice in many aerospace applications. Although more expensive, they provide higher specific modulus and specific strength than glass fibers. The advantage of carbon fibers is further enhanced in fatigue. However, carbon fibers are electrical conductors, and their contact with metals may lead to corrosion of the latter. Polymeric fibers such as aramid and high-density polyethylene are the toughest of all the available fibers and hence can be used where high-impact resistance and toughness are required. These polymeric fibers, however, are weak in compression because of the fibrillar nature of their microstructure. Recently, a variety of ceramic fibers such as alumina and silicon carbide have emerged mainly as reinforcements for metal and ceramic matrices. These ceramic fibers have better oxidation resistance in high-temperature applications than carbon fibers. However, they are still more expensive than most carbon fibers. Mechanical properties of epoxy matrix composites made with the four most widely used fibers--aramid, carbon, E-glass and S-glass--are shown in Table 3-2. Tensile fatigue behaviors of the first three composites are compared in Figure 3-1.

Since one type of fiber does not have all the desired properties, different fibers can be mixed to make a hybrid composite. For example, in a glass/carbon hybrid composite the glass fiber can improve the impact resistance while keeping the cost down, and the carbon fiber can provide the required strength and stiffness with less weight. The weight savings resulting from the use of a hybrid composite reduces the load on the blade and hence will lead to a longer lifetime. Furthermore, the material savings realized can also compensate partially for the higher cost of the carbon fiber.



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Assessment of Research Needs for Wind Turbine Rotor Materials Technology 3 MATERIALS PROPERTIES AND LIFE PREDICTION The characteristics that make composites, especially glass fiber-reinforced and wood/epoxy composites, suitable for wind turbine blades are low density, good mechanical properties, excellent corrosion resistance, tailorability of material properties, and versatility of fabrication methods. Although glass/vinylester and glass/polyvinyl composites based on hand lay-up have been the most widely used materials so far, many more types of fibers and resins have become available recently. The new carbon fibers are stronger and stiffer, while the new resins provide higher toughness and shorter process cycle time. A number of handbooks are now available to the designers of composites (Lubin, 1982; Engineered Materials Handbook, 1987; Composites & Laminates, 1987). FIBERS The most commonly used and lowest-priced fiber is E-glass fiber. Over the past several years, however, many new fibers have become available. The commercially available fibers and their typical properties are listed in Table 3-1. While E-glass fiber is most widely used in wind turbine rotor blades mainly because of its low cost, carbon fibers are the fibers of choice in many aerospace applications. Although more expensive, they provide higher specific modulus and specific strength than glass fibers. The advantage of carbon fibers is further enhanced in fatigue. However, carbon fibers are electrical conductors, and their contact with metals may lead to corrosion of the latter. Polymeric fibers such as aramid and high-density polyethylene are the toughest of all the available fibers and hence can be used where high-impact resistance and toughness are required. These polymeric fibers, however, are weak in compression because of the fibrillar nature of their microstructure. Recently, a variety of ceramic fibers such as alumina and silicon carbide have emerged mainly as reinforcements for metal and ceramic matrices. These ceramic fibers have better oxidation resistance in high-temperature applications than carbon fibers. However, they are still more expensive than most carbon fibers. Mechanical properties of epoxy matrix composites made with the four most widely used fibers--aramid, carbon, E-glass and S-glass--are shown in Table 3-2. Tensile fatigue behaviors of the first three composites are compared in Figure 3-1. Since one type of fiber does not have all the desired properties, different fibers can be mixed to make a hybrid composite. For example, in a glass/carbon hybrid composite the glass fiber can improve the impact resistance while keeping the cost down, and the carbon fiber can provide the required strength and stiffness with less weight. The weight savings resulting from the use of a hybrid composite reduces the load on the blade and hence will lead to a longer lifetime. Furthermore, the material savings realized can also compensate partially for the higher cost of the carbon fiber.

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Assessment of Research Needs for Wind Turbine Rotor Materials Technology TABLE 3-1 Typical Properties of Fibers Fiber Diameter, µm Specific Gravity Modulus, GPa* Tensile Strength, GPa Failure Strain, % CTE, 10-6/°C E-glass 3-20 2.4 72.4 3.45 S2-glass 10-20 2.4 86.9 4.59 Boron 142 2.5 400 2.80 0.7 Carbon, PAN based High strength 1.8 230 2.5 1.1 -0.4 High modulus 1.9 370 1.0 0.5 -0.5 Ultra-high modulus 2.0 520 1.0 0.2 -1.1 High strain 1.8 230-260 3.0-4.8 1.6-2.0 Intermediate modulus 1.8 290 5.2 1.8 Carbon, pitch based P25 10 1.8 140 1.4 1.0 -0.9 P55 10 2.02 380 1.4 0.4 -1.3 P75 10 2.00 520 1.4 0.3 -1.6 P100 10 2.15 700 1.4 0.2 -1.6 P120     820 1.4 0.2 -1.6 P140     960 1.4 0.2 -1.8 Silicon carbide SCS 140 3.0 430 3.4 0.8 Nicalon 15 2.6 180-210 2.5-3.3 1.5 3.1 AVCO 6-10   300 2.8 MPDZ 10-15 2.3 180-210 1.8-2.1 HPZ 10 2.35 140-175 2.1-2.5 MPS 10-15 2.6-2.7 175-210 1.1-1.4 Alumina FP 20 3.9 345-380 1.4 0.36-0.4 Sumitomo 9-17 3.2 210-260 1.8-2.6 0.86 Saffil 3 3.3 300 2.0 0.67 PRD-166   4.2 385 2.1-2.5 Alumina-silica Fiberfrax   2.7 100 1.0 Fibermax   3.0 150 0.8 0.5 Alumina-boria-silica Nextel 312 11 2.7 150 1.4-1.7 Mextel 440 10-12 3.1 200-240 1.4-2.0 0.9 Nextel 480 10-12 3.1 224 2.3 0.9 Si, Ti, C, O Tyranno 8-10 2.3-2.5 200 3.0 1.5 Aramid Kevlar 49 12 1.44 124-131 3.6 2.8 -2.0 Kevlar 29 12 1.44 83 3.6 4.0 Kevlar 149 12 1.47 186 3.4 2.0 HM-50 12 1.3 81 3.1 4.4 Polyethylene Spectra 900 38 0.97 117 2.6 3.5 Spectra 1000 27 0.97 172 3.1 0.7 Tungsten 13 19.4 410 4.0 * 1 GPa = 1.45 × 105 psi.

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Assessment of Research Needs for Wind Turbine Rotor Materials Technology TABLE 3-2 Typical Properties of Unidirectional Composites   E-Glass/Resin KEVLAR Resin Graphite/Resin S-Glass/Resin Fiber Direction Modulus, GPa 44.8 75.8 145 56 Strength tension, MPa 1124 1241 1517 1980 Strength compression, MPa 896 276 2068 626 Coefficient of thermal expansion, microstrain/K 8.8 -4.0 -0.45 5.5 Transverse Direction Modulus, GPa 11.0 5.52 10.3 11.4 Strength tension, MPa 31 14 48 31 Strength compression, MPa 138 55 138 138 Coefficient of thermal expansion, microstrain/K 22.1 57.6 25.2 23.3 Shear (Inplane) Modulus, GPa 4.14 2.41 5.52 4.48 Strength, MPa 71.7 34.5 82.7 71.7 Poisson Ratio Axial-Transverse 0.27 0.34 0.30 0.27 Thickness Direction Modulus, GPa 11.0 5.5 10.3 11.4 Poisson ratio 0.44 0.38 0.55 0.44 To further illustrate the cost differences based on fibers, consider a hybrid composite with a carbon/glass volume ratio vc/g. The ratio of the cost of the hybrid composite to an all-glass composite to provide the same structural stiffness is given by The parameters in Equation (1) are defined as follows: $, price per unit mass; s, specific gravity; E, Young's modulus; g, glass; and c, carbon. On the other hand, the weight ratio, defined as the weight of the hybrid composite divided by that of the all-glass composite, is given by E-glass fiber costs around $2/lb; a high-strength carbon fiber costs about $30/lb. Thus, an all-carbon fiber composite costs 3.5 times more than an all-glass fiber composite to provide the same structural stiffness at a weight savings of 76 percent. When a 50/50 glass/carbon hybrid composite is used, however, the calculated cost ratio is reduced to 2.9 with a weight savings of 58 percent.

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Assessment of Research Needs for Wind Turbine Rotor Materials Technology Figure 3-1 Trends of longitudinal tensile fatigue S-N data for unidirectional composites with various fibers. Source: Mandell (1990). As can be inferred from Table 3-2, the in situ tensile failure strain of E-glass fiber is as high as 2.5 percent, whereas it is only 1 percent for carbon fiber. However, E-glass fiber has a much lower fatigue ratio than carbon fiber, that is, 0.3 versus 0.75 at 10 million cycles (Figure 3-1). Therefore, both fibers have the same fatigue strain of 0.75% at 10 million cycles. Beyond 10 million cycles, however, the carbon fiber is expected to outlast the glass fiber. The cost ratio to obtain the same long-term fatigue strength is then at least the same as that needed to obtain the same stiffness. An additional benefit is that a hybrid composite blade will have a longer lifetime because of the reduced fatigue load resulting from the weight savings. Thus, a careful study is needed to explore the full benefits of a hybrid composite blade from a life-cycle point of view. Fibers are used in various forms. The commonly used fiber preforms include unidirectional tow, woven cloth, knitted fabric, continuous strand mat, chopped strand mat, and braid as well as chopped fibers in sheet and bulk molding compounds. Depending on the application and manufacturing process used, one fiber preform may be preferred to others. The various fiber preforms are shown schematically in Figure 3-2. Where high strength is required, unidirectional bundles of fibers known as tows should be used. Woven cloths and knitted fabrics are easier to use, especially over complex contours. Woven cloths have disadvantages in that

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Assessment of Research Needs for Wind Turbine Rotor Materials Technology Figure 3-2 Schematics of various fiber preforms. Source: Chou et al. (1986).

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Assessment of Research Needs for Wind Turbine Rotor Materials Technology high fiber volume content cannot be attained and the inherent fiber cross-overs are susceptible to premature fatigue damage. The fiber cross-overs also result in low compressive strength. While continuous strand mats provide good strength, chopped strand mats are weaker because of the fiber discontinuity. Where delamination is a prime concern and strengthening is required in the thickness direction, braids can work well. MATRIX MATERIALS The matrix resins commonly used in wind turbine blades are divided into three major classes of thermosetting polymers: unsaturated polyesters, epoxies, and vinyl esters (Lee and Neville, 1967; Launikitis, 1978; Strong, 1989). The unsaturated polyester resins, typically based on the orthophthalic or isophthalic acid, are most widely used because of short cure time and low cost. Despite the long cure time required and higher cost, epoxy resins are gaining greater acceptance due to their superior chemical resistance, good adhesion, low cure shrinkage, good electrical properties, and high mechanical strength. A reasonably good compromise between cost, cure time, and the above-mentioned physical properties is achieved in epoxy-based vinyl ester resins, which have especially shown very rapid growth over the past few years by filling the gap between polyesters and epoxies. The use of all these resins as composite matrices is generally subjected to a temperature limit around 150° to 200°C. The upper temperature limit is lower than the glass transition temperature Tg because above Tg the resin becomes soft and rubbery, thus losing most of the load-carrying capability. Polyesters, as their name implies, are resins in which recurring ester linkages are an integral part of polymer chain backbone. The most commonly used system--isophthalic polyester resins--is more expensive but is chemically more stable and a little less brittle than orthophthalic resins. The term unsaturated comes from the presence of carbon-carbon double bonds in the polyester chain backbone that provide the location for cross-linking, thereby eventually leading to a tight network. The most common type of cross-linking agent is styrene. The styrene also lowers the initial viscosity of the polyester resin to improve processing of the composite by facilitating impregnation and wetting of fibers. The cross-linking reaction, which is addition polymerization, is triggered by an initiator (sometimes erroneously called a catalyst), such as organic peroxide, that produces free radicals. Many polyester resins are cured at room temperature with a rise in temperature due to the exothermic nature of the reaction. In this case, the typical cure time is several hours or overnight. However, with a proper selection of initiator, the cure reaction can be carried out at an elevated temperature in an extremely short time. Experience in automotive composites has shown that the curing of an unsaturated polyester can be completed in 2 to 3 minutes at 150°C with the use of a tert-butyl perbenzoate initiator. Epoxy resins derive their name from the epoxide ring structure that serves as the principal cross-linking site. Although various types of epoxy resins are produced commercially, so-called diglycidyl ether of bisphenol-A (DGEBPA)-type resins (e.g., Shell EPON828), have achieved the widest market acceptance and demonstrated versatility. For applications requiring special properties such as high modulus or higher use temperature, other types of epoxy resins with higher functionality are used in place of difunctional DGEBPA-type resins. For instance, tetrafunctional epoxy resins of tetraglycidyl methylene dianiline (TGMDA) (e.g., Ciba-Geigy MY720) exhibit improved properties at elevated temperature. Greater strength and improved properties at elevated temperature are also achieved in the resins of epoxidized phenolic novolacs (e.g., Dow DEN438) or tetraglycidyl ether of tetrakis hydroxyphenyl ether (e.g., Shell EPON103). All these resins tend to have lower fracture toughness than DGEBPA-type resins. For polyfunctional epoxy resins, cross-linking or cure is effected by use of polyfunctional curing agents or hardeners. The use of a hardener with higher functionality (e.g., tertiary amine versus primary amine) allows a greater cross-link density of cured resins, which generally improves their physical properties. Lewis acids, such as BF3, are catalytic curing agents

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Assessment of Research Needs for Wind Turbine Rotor Materials Technology that promote the cure reaction but do not themselves serve as direct cross-linking agents. These catalytic curing agents have the advantage of very long shelf life and require significant heat to initiate the reaction. The resultant structure of Lewis acid-cured resin is very tightly cross-linked. Curing of most epoxy resins and composites is done at 120° or 177°C with a typical cycle time of 2 to 3 hours under the molding pressure of up to 1.4 MPa. The term vinyl ester can be applied to any number of chemical compounds comprising an ester linkage and terminal unsaturation. There are several types of vinyl esters based on epoxy resins as well as nonepoxy resins. However, to the composites industry, the vinyl ester resins usually mean methacrylate esters of epoxy resins (Launkitis, 1978). Unlike polyesters, vinyl esters do not possess internal unsaturation. However, vinyl esters and polyesters are similar in that they both utilize a coreactant or cross-linking agent, such as styrene, and free radical-producing initiators, such as peroxides, to effect cure. As a result, vinyl ester resins can be cured in a very short time like polyesters, but their static strength and modulus properties are similar or comparable to those of epoxy resins. Typical mechanical properties of polyesters, epoxies, and vinyl esters are shown in Table 3-3. While epoxy resins are at the top of the scale as far as mechanical properties are concerned, they require the longest cure time and are the most expensive. For example, epoxy resins cost about $1.80/lb, whereas unsaturated polyesters cost only $1.00/lb. Vinyl ester resins fall in between these two resins, costing about $1.60/lb. Thus, the final selection of a resin should consider the material cost as well. The fatigue crack propagation rates of these resins vary with the stress intensity factor range ΔK in the case of other plastics for structural use (Hertzberg and Manson, 1980). In amine-cured DGEBA epoxy resins, the values for exponent m in the equation for the crack growth per fatigue cycle, da/dN = const.(ΔK)m, range from 7.7 to 20, which are higher than those for other plastics. In general, a lower fatigue crack growth rate and a higher fracture toughness are observed with increasing molecular weight between cross-links. As discussed previously, all of these resins suffer the problem of brittleness. A brittle resin results in premature matrix cracking in the composite, which in turn facilitates moisture ingress. Although fracture toughness of the resin systems can be raised by increasing Mc or adding diluents, these approaches result in lowering of modulus and temperature resistance (Hertzberg and Manson, 1980; Owen, 1974; Christensen and Rinde, 1979). At present, many commercially available resins utilize toughening agents in the form of discrete particles of elastomers or ductile TABLE 3-3 Properties of Cast Resins   Polyester Vinyl Ester Epoxy Specific gravity 1.10-1.46 1.1-1.2 1.2-1.3 Flexural strength, MPa 60-160 120-140 110-215 Tensile strength, MPa 40-90 70-90 50-130 Compressive strength, MPa 90-200 - 110-210 Tensile elongation, % <5 <6 <9 Modulus, GPa 2-4 3-4 3-4.5   Source: Lee and Neville (1967), Bucknall (1977), Lubin (1982).

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Assessment of Research Needs for Wind Turbine Rotor Materials Technology thermoplastics. In this case the increase of fracture toughness is achieved with a relatively small change of modulus and temperature resistance. Toughening of epoxy resins by the inclusion of elastomer particles is well established. Although several different types of liquid or solid elastomers or their hybrids can be used, the most effective way of increasing resin toughness is by using a liquid rubber with specific terminal functional groups (Lee, Riew, and Moulton, 1980; Bucknall, 1977; Riew, Rowe, and Siebert, 1976; Scott and Phillips, 1975; McGarry and Sultan, 1969; Bascom, Bitner, Moulton, and Siebert, 1980; Moulton and Ting, 1981; Lee, 1986). Examples are butadiene acrylonitrile copolymers with carboxyl, amine, or vinyl groups at both ends of the chain which are selected depending on the type of curing agent and curing mechanisms. When carboxyl-terminated butadiene acrylonitrile copolymer (CTBN) is mixed in liquid state with an amine-cured DGEBPA epoxy resin, their initial compatibility in liquid state followed by phase separation in the solidification process leads to the formation of submicron-sized elastomer particles that are well bonded to the surrounding resin. As a result, adding as small as 5 wt% of CTBN to Epon 828 resin increases the fracture toughness G1c from 350 to 5260 J/m2. Yet the accompanying decrease of tensile modulus is only 2.5 to 2.8 GPa (Riew, Rowe, and Siebert, 1976). A similar type of elastomer toughening in extremely brittle TGMDA resins (G1c = 80 J/m2) is less effective: a maximum twofold increase with considerable reduction of modulus (Lee, Riew, and Moulton, 1980). The use of elastomer-toughened epoxy resins as matrix materials has been shown to improve the interlaminar fracture toughness of fiber-reinforced composites, particularly those with woven fabric reinforcement with resin-rich areas between the plies (Bascom, Bitner, Moulton, and Siebert, 1980). Despite a high level of constraint on thin films of resin matrix by densely packed surrounding fibers (Scott and Phillips, 1975), toughening of epoxy matrices also increases the resistance of composites against local damage initiation and accumulation (McGarry and Sultan, 1969; Bascom, Bitner, Moulton, and Siebert, 1980; Moulton and Ting, 1981; Lee, 1986). However, their effects on fatigue lifetime or fatigue endurance limit of composites have not been fully confirmed. In the case of toughening of unsaturated polyester resins, both liquid elastomers with terminal functional groups and ductile thermoplastics are utilized as a secondary phase (McGarry, Rowe, and Riew, 1978; Lee, Howard, and Rowe, 1983). Examples are vinyl- or epoxy-terminated butadience acrylonitrile elastomers, hydroxyl- or vinyl-terminated epichlorohydrin elastomers, and polyvinylacetate. Compared with elastomer-toughened DGEBPA epoxy resins, the size of elastomer particles formed in situ during cure is much bigger (micron level) because of lower compatibility of the reactive mixture. As a result, the effectiveness of elastomer toughening of polyester resins is much lower than that of DGEBPA epoxy resins. For instance, by adding 8 wt% of vinyl-terminated polyepichlorhydrin elastomer, the fracture toughness G1c of isophthalic polyester resin is increased from 60 to 110 J/m2 with a 23 percent reduction of tensile modulus. However, the use of elastomer-toughened unsaturated polyester resins in short fiber-reinforced composites increases the local damage resistance under impact and, in certain cases, that of tensile strength as well (McGarry, Rowe, and Riew, 1978; Lee, Howard, and Rowe, 1983). In addition to thermosetting resins, many engineering thermoplastics are available now as matrix resins. Table 3-4 lists those new high-performance thermoplastic resins developed mainly for aerospace applications (Witzler, 1988). Compared with thermosetting resins, the thermoplastic resins offer advantages in prepreg (pre-impregnated laminate) stability and short processing cycle. However, they tend to be weak in solvent resistance and fiber impregnation. Trade-offs between thermosets and thermoplastics as matrices are shown in Table 3-5. Unidirectional prepregs also are available with commodity thermoplastics such as nylon and PET. Although they are cheaper than the high-performance thermoplastic prepregs, their full potential has not been realized because of the lack of low-cost manufacturing techniques.

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Assessment of Research Needs for Wind Turbine Rotor Materials Technology TABLE 3-4 High-Performance Thermoplastics Used as Matrix Resins Polymer Type Trade Name Manufacturer Tga, °F Processing Temperature, °F Polyetheretherketone (PEEK) Semicrystalline APC-2b ICI 290 650 Polyphenylene sulfide (PPS) Semicrystalline Ryton Phillips 185 630 Polyarylene ketone Semicrystalline HTX ICI 400 700-789 Polyarylene sulfide Amorphous PAS-II Phillips 410 625-650 Polyetherimide (PEI) Amorphous Ultem (resin) GE Varies Varies     Cypacb American Cyanamid         Cypac 7000/7005   450 575-650     Cypac 7156 545 650-700 Polyarylether Amorphous Radel C Amoco 476 650 Polyethersulfone Amorphous HTA ICI 510 575 Polyamide-imide Amorphous Torlon Amoco 470 650 Polyimide Pseudothermoplastic Avimid Du Pont K-III 480 680     LaRC-TPI Mitsui Toatsu, Rogers Corp. 482 660     2080 Lenzing 536 660     9725 Ciba-Geigy 536 660 a Tg is the glass transition temperature, the temperature at which a polymer changes from a rigid glassy solid to a soft rubbery solid. b Prepreg. Source: Johnston and Hergenrother (1987).

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Assessment of Research Needs for Wind Turbine Rotor Materials Technology TABLE 3-5 Trade-offs Between Thermosets and Thermoplastics as Matrices Property Thermosets Thermoplastics Formulations Complex Simple Melt viscosity Very low High Fiber impregnation Easy Difficult Prepreg tack Good None Prepreg drape Good None to fair Prepreg stability Poor Excellent Processing cycle Long Short to long Processing temperature/pressure Low to moderate High Fabrication cost High Low (potentially) Mechanical properties Fair to good Fair to good Environmental durability Good Unknown Solvent resistance Excellent Poor to good Damage tolerance Poor to excellent Fair to good Data base Very large Small The difficulty associated with fiber impregnation has led to the development of commingled yarns and fabrics as well as powder-coated yarns and fabrics. A commingled yarn consists of both reinforcing fibers and thermoplastic fibers. During processing the thermoplastic fibers melt and impregnate the reinforcing fibers (Lynch, 1989). Thermoplastic powder is smaller in diameter than the corresponding thermoplastic fiber and hence is more evenly distributed in the yarn, which facilitates a more uniform fiber impregnation (Hartnes, 1988). E-GLASS/PLASTIC COMPOSITES E-glass/plastic composites, commonly called glass-reinforced plastics (GRPs), have been widely used in the manufacture of blades of various sizes. Typical resins used in GRPs are polyester, vinyl ester, and epoxy. Although polyester resins can be cured in the shortest time, they show large shrinkage and hence may not be appropriate for use in a hot, dry environment such as California, where moisture-induced swelling would be minimal (Windpower Monthly, 1987). Vinyl ester resins have good environmental stability and are widely used in marine applications. Epoxy resins have good mechanical properties and dimensional stability. Their drawback is longer cure time and higher cost; however, new epoxy resins are now available for pultrusion and resin transfer molding, which require fast curing. GRP is the main structural material for a number of blades of large machines and is also used as cladding over steel load-bearing frames in others (Phillips et al., 1987). GRP is the most popular blade material used in medium-size machines in Denmark and The Netherlands. As of January 1987, approximately 81 percent of 15,059 wind turbines in the California wind farms had fiberglass rotor blades (Stoddard, 1989; Modern Power Systems, 1986). Typical properties of a unidirectional E-glass/epoxy composite are compared with those of other unidirectional composites in Table 3-2. For all the composites in the table, the stress-strain behavior is usually quite linear except in transverse compression and in-plane shear. Another exception is thearamid/epoxy composite, which is quite nonlinear in longitudinal compression also. Elastic properties of multidirectional laminates can be calculated from unidirectional properties using laminated plate theory (Tsai, 1989; Jones, 1975). Tensile stress-strain relationships of several representative E-glass/epoxy laminates are shown in Figure 3-3. When a multidirectional

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Assessment of Research Needs for Wind Turbine Rotor Materials Technology laminate, such as the quasi-isotropic laminate in the figure, is subjected to tension, cracks appear first in the matrix and fiber/matrix interfaces of off-axis plies (Tsai and Hahn, 1975; Reifsnider, 1980). Most of these cracks, called ply cracks, are entirely through the thickness of individual off-axis plies. The crack density increases with increasing load until some of the cracks grow as delamination between plies with different fiber orientations. Further increase of load breaks the fibers in the on-axis plies, leading to failure of the laminate as a whole. When the load is along the 0° direction, 90° plies may fail at a strain as low as 0.3 percent, although the laminate does not fail until 2.3 percent strain is reached. Thus, failure of a multidirectional laminate is preceded by matrix/interface cracking in the off-axis plies and also by delamination between plies with different fiber orientations. The ply cracking in general does not lead to immediate failure of the laminate. However, it facilitates moisture diffusion through the cracks, thereby reducing the durability of the laminate. Even in the absence of cracks, epoxy resins can absorb as much as 7 percent moisture by weight. The absorbed moisture plasticizes the resin, thereby reducing its glass transition temperature. When fully saturated, the glass transition temperature of epoxy resins can be reduced by as much as 100°C (Vinson, 1977). At room temperature, absorbed moisture has a rather minimal effect on mechanical properties. When combined with elevated temperature, however, it can seriously degrade mechanical properties through interfacial debonding and matrix plasticization. The worst combination is thermal spiking to a temperature above the glass transition temperature in the presence of moisture. The thermally spiked specimens show increased moisture absorption and lower strengths (Springer, 1981). Figure 3-3 Stress-strain relationships of glass/epoxy laminates under uniaxial tension (50 vol%). (LONG: strain parallel to loading direction; and TRANS: strain [magnitude] normal to loading direction)

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Assessment of Research Needs for Wind Turbine Rotor Materials Technology Figure 3-8 Effect of moisture content on laminate mechanical properties. Figure 3-9 Typical tensile fatigue strength of wood/epoxy laminate.

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Assessment of Research Needs for Wind Turbine Rotor Materials Technology Figure 3-10 Typical compression fatigue strength of wood/epoxy laminate. Figure 3-11 Typical reversed stress fatigue strength of wood/epoxy laminate.

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Assessment of Research Needs for Wind Turbine Rotor Materials Technology blade, while loads that are nearly fully reversing will be found near the blade's leading and trailing edges due to the deadweight gravity bending moment. Testing (both with and without joints in the laminate) shows the lowest strength levels and most rapid loss of strength with increasing cycles for the reversed load R = -1 case. Strength levels for R = 0.1 tension fatigue can be 50 to 100 percent higher than for reverse axial, but the rate of strength loss with cycles is similar. Depending on size, compression strength will generally be somewhat lower than tension at low cycles, but its strength loss rate is only about two-thirds of that for tension, so it can cross over and exceed tensile strength at higher cycles, particularly for large blades. These strength loss rates tend to be s for each type of loading over a wide range of cycles. Consequently, the data plot as a straight line on a log-log stress versus number of cycles (S-N) plot. This linear trend certainly holds valid up to 107 cycles; there are not enough data beyond that to determine where or if a fatigue endurance limit exists. That knowledge could be valuable for better assessment of very high cycle strength and life, but the appropriate data are expensive and time consuming to obtain. Current life and strength assessments are forced to assume that the loss of strength shown at lower-cycle levels continues at the same rate to arbitrarily high cycles. The effect of size on strength varies dramatically with the type of loading. While size effect for static tensile strength along the grain shows strength loss that can be a factor of 2 between laboratory samples and large blades, it becomes worse for tension fatigue where it can reach roughly a factor of 3. For fully reversed R = -1 fatigue, the size effect drops back about to the 2:1 level of static tension. Size effect data for compression fatigue and static compression strength are minimal, but size effect appears to drop further to perhaps 1.5:1 for R = 0.1 compression fatigue and to some lesser but ill-defined value for static compression (Spera et al., 1990). Recent research work in size effect in the secondary material properties of cross-grain tension and rolling shear was not able to cover the range of R ratios needed to draw conclusions similar to the above, but it revealed a new phenomenon due to carefully controlled sample design and test structure, namely that size effect is more severe for higher-cycle levels. While rolling shear showed very little loss of strength with size in static tests, that was no longer true at 106 cycles; and while cross-grain tension did show substantial size effect in static tests, it again became considerably greater at 10 6 cycles. An associated effect is that the fatigue curve slope, the rate of strength loss with cycles, is larger for the larger specimens. So not only do larger specimens start out with a lower initial static strength, that strength is lost faster on a fractional basis in fatigue than it is for smaller samples (Bertelsen and Zuteck, 1991). These lessons in size effect are important and will be increasingly so as wind turbines grow larger for economic reasons. There is good reason to suspect that size effect will occur in all materials currently contemplated for future wind turbine rotors, including GRP and metals, particularly for high-cycle fatigue. We now have experimental proof that low levels of size effect in static tests are no assurance of its absence in high-cycle fatigue, and designs undertaken in ignorance of these effects could lead to expensive or even dangerous failures in future large-scale machines. The interaction of long-term environmental effects with the fatigue process has been partially addressed in wood/epoxy design by running fatigue tests at a range of temperature and moisture levels. However, it was found that the cycle rates typical for long-duration tests drove substantial moisture from the test specimens, thereby elevating their strength and fatigue performance. This may not occur in the same way at the lower-cycle rates and intermittent use that occurs in the field, so caution must be exercised lest the elevated performance of laboratory samples lead to excessive high-cycle fatigue expectations in the field. The effect of veneer splice joints on strength is much more pronounced in fatigue than in static strength. Early work quickly highlighted this and showed that 10 to 20 percent strength loss could easily occur owing to these tiny joints distributed throughout the laminate. The effects were particularly apparent in tension and in high-cycle fatigue. The first

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Assessment of Research Needs for Wind Turbine Rotor Materials Technology attempts to combat this by using beveled scarf joints between the veneers met with only limited and frustratingly intermittent success. In the large-scale fatigue specimens, performance was decidedly poor, and zipper-like failures through large numbers of joints often occurred. This puzzling state of affairs has been investigated at length in the recent Small Business Innovative Research (SBIR) research work and was traced to expanding gas driving resin from the joints during the vacuum molding process (Bertelsen and Zuteck, 1991). The path to higher-performance joints through mitigating this effect now seems clear in principle, but a practical manufacturing method has yet to be proven. The important general lesson is that defect-driven failure modes that are hard to spot in small specimens tested in static or low-cycle loading may emerge as major issues in bulk specimens tested to moderately high cycles. It is difficult to see how the high-cycle fatigue performance of large structures can be assured without at least a few tests of large-volume specimens to elevated cycles. Fortunately, considerable data of this sort now exist to support the most critical design areas of large wood/epoxy wind turbine blades. However, those data are by no means complete, and comparable data for design in GRP are unavailable. Practical limits to the amount of materials data available will always exist, and perfect knowledge is not necessary to execute a successful design. The benefit of improved materials knowledge is that a more appropriate design that sidesteps pitfalls without undue cost can be provided and can be put into service with a much lower uncertainty as to its operating envelope and expected life. The record of wood/epoxy blades in the field has been excellent despite early materials data limitations due to confining the design window to stay within available knowledge. A wider design window is now available due to knowledge accumulated over the past decade. This will be needed in addressing the larger, more cost-effective designs that are needed in the future. A still broader design window can be opened by addressing the known shortfalls in the existing data and any additional ones that may be revealed as further advanced blade design occurs. RECOMMENDATIONS When designed and manufactured properly, glass/polymer and wood/epoxy rotor blades can provide tens of thousands of hours of operating time. The recent advances in composites technology, however, may provide an excellent opportunity to further improve the blade cost/performance. The knowledge base gained over the past two decades in the aerospace industry can now be used to assess these more efficient blade designs using advanced composites. What is needed is a focused research to refine the design along with improved materials knowledge in certain key areas. Specific recommendations for further research are described below, in the order of priority: Long-term fatigue data should be generated to at least 100 million cycles for the most useful composite laminates and critical elements containing manufacturing defects, under appropriate environmental conditions. Separate tension and compression fatigue data should be generated so that independent lifetime projections for upwind and downwind portions of blade skin can be made. Fully reversed fatigue data should also be obtained for the projection of life against edgewise gravity loads. Moderate-cycle R ratio data may be needed to fully define material fatigue response. The data should include documentation on the modes and growth of damage as well as the effect of damage on properties. At least a few large specimens of the favored composite laminates should be tested in moderate- to high-cycle fatigue to quantify the fatigue size effect allowance that must be made for large rotors in the primary loading regimes. All loading cases are needed for GRP. For wood/epoxy, size effect in static compression and compression fatigue is not yet quantified. Sui tests comparing small and large specimens of matched parent material are needed to fill this gap in the data. Spectrum loading data to assess the available cumulative damage models are not yet available but are needed to fully quantify high-cycle life predictions. A life prediction methodology should be developed to predict

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Assessment of Research Needs for Wind Turbine Rotor Materials Technology blade lifetimes based on constant-amplitude fatigue data. It should also be able to predict the modes and growth of damage. A database should be established for wind turbine blade materials. It should include mechanical properties on fibers, matrix resins, and composites, including wood/epoxy. An extensive search of all fatigue data on both GRP and wood/epoxy should be conducted and the results included in the database. Benefits of hybrid composites should be explored through design and limited testing. Design studies should investigate savings on blade weight and life-cycle cost. The best hybrid reinforcement appears to be carbon for either GRP or wood/epoxy. Duration of load and creep effects are incompletely understood for wood/epoxy laminate. It appears that epoxy stabilization of moisture levels reduces creep rate and magnitude, but data to verify and quantify this for design are needed. Furthermore, moisture correction for laminate in the 5 to 12 percent operating range is still based primarily on old literature data. A better database would improve life predictions. Promising fiber preforms should be identified and examined under fatigue. Low-cost processability should be taken into consideration. Environmental effects, including ultraviolet exposure, moisture absorption, and temperature fluctuations, should be delineated for the candidate composites and surface coatings. Tough new resins should be examined for their applicability to rotor blades.

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Assessment of Research Needs for Wind Turbine Rotor Materials Technology REFERENCES AND BIBLIOGRAPHY Ansell, M. P. 1987. Layman's Guide to Fatigue. The Geoff Pontin Memorial Lecture, sponsored by British Tradewinds. Wind Energy Conversion, ed. by J. M. Galt, Mechanical Engineering Publications Ltd., London. Ansell, M. P., P. W. Bonfield, and K. T. Tsia. Fatigue Testing of Laminated Wood for Generator Blades. School of Materials Science, University of Bath, England. Bach, P. W. 1988. High Cycle Fatigue of Glassfiber Reinforced Polyester. Presented at the IEA Workshop on Fatigue in Wind Turbines in Harwell, United Kingdom. Barrett, J. D. 1974. Effect of Size on Tension Perpendicular to Grain Strength of Douglas-Fir. Wood and Fiber. Barrett, J. D., and R. O. Foschi. 1978. Duration of Load and Probability of Failure in Wood . Part II. Constant, Ramp, and Cyclic Loadings, Canadian Journal of Civil Engineering, Vol. 5. Barrett, J. D., R. O. Foschi, and S. P. Fox. 1975. Perpendicular to Grain Strength of Douglas-Fir. Canadian Journal of Civil Engineering, Vol. 2. Bascom, W. D., J. L. Bitner, R. J. Moulton, and A. R. Siebert. 1980. The Interlaminar Fracture of Organic Matrix, Woven Reinforcement Composites. Bertelsen, W. D., and M. D. Zuteck. 1991. Investigation of Fatigue Failure Initiation and Propagation in Wind-Turbine-Grade Wood/Epoxy Laminate Containing Several Veneer Joint Styles. U.S. DOE Phase II Report, in preparation. Bohannan, B. 1966. Effect of Size on Bending Strength of Wood Members. U.S. Forest Service Research Paper, FLP 56, May. Bonfield, P. W., and M. P. Ansell. Fatigue Testing of Wood Composites for Aerogenerator Rotor Blades. School of Materials Science, Bath University, England. Bowen, D. H., C. W. A. Maskell, D. C. Phillips, T. W. Thorpe, G. M. Wells, and N. J. M. Wilkins. 1984. Materials Aspects of Large Aerogenerator Blades. Proceedings of the European Wind Energy Conference, Hamburg, West Germany, October 22-26, p. 281. Bucknail, C. B. 1977. Toughened Plastics. Applied Science Publishers. Chou, T.-W., R. L. McCullough. and R. B. Pipes. 1986. Composites. Scientific American, Vol. 254, October, p. 193. Christensen, R. M., and J. A. Rinde. 1979. Transverse Tensile Characteristics of Fiber Composites. Polymer Engineering Science, Vol. 19, 506. Clark, R. N., F. C. Vosper, R. G. Davis, and W. E. Pinker. 1985. Operational Data From Bushland Wind Turbines. Final Report. U.S. Department of Agriculture (USDA), Agricultural Research Service, USDA Conservation and Production Research Laboratory, Bushland, Texas. Prepared for Electric Power Research Institute, Palo Alto, California. Composites & Laminates. 1987. Edition 1. D.A.T.A., Inc. Downing, S. D., and D. F. Socie. 1982. Simple Rainflow Counting Algorithms. International Journal of Fatigue, Vol. 4, p. 31.

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Assessment of Research Needs for Wind Turbine Rotor Materials Technology Engineered Materials Handbook. 1987. Vol. 1. Composites. ASM International. Faddoul, J. R. 1981. Test Evaluation of a Laminated Wood Wind Turbine Blade Concept. DOE/NASA/20320-30, NASA TM 81719, May. Faddoul, J. R. 1983. Examination, Evaluation, and Repair of Laminated Wood Blades After Service on the Mod-0A Wind Turbine. DOE/NASA/20320-53, NASA TM-83483, October. Garrad, A. D., and U. Hassan. 1986. Dynamic Analysis of Wind Turbines for Fatigue Life Prediction. Presented at the 8th BWEA Meeting in Cambridge, United Kingdom. Gerhards, C. C. 1977. Effect of Duration and Rate of Loading on Strength of Wood and Wood-Based Materials. USDA Forest Service Research Paper, FLP 283. Gougeon Brothers, Inc. 1985. Engineered Laminates. Technical Bulletin #1, Bay City, Michigan, April. Hahn, H. T. 1979. Fatigue Behavior and Life Prediction of Composite Laminates. Composite Materials: Testing and Design (5th Conf.), ASTM STP 674, p. 383. Han, Y. M., and H. T. Hahn. 1989. Design of Composite Laminates with Ply Failure. Proceedings of the 34th International SAMPE Symposium, p. 529. Hartness, T. 1988. Thermoplastic Powder Technology for Advanced Composite Systems. Proceedings of the 33rd International SAMPE Symposium, p. 1458. Hertzberg, R. W., and J. A. Manson. 1980. Fatigue of Engineering Plastics . Academic Press, New York, p. 124. Hunston, D. 1987. ASTM STP 937. Hunston, D. 1990. Advanced Materials and Material Forms. Presented at the NRC Workshop on Assessment of Research Needs for Wind Turbine Rotor Materials Technology, Washington, D.C., January 22-23. Hwang, W. B., and K. S. Han. 1989. Fatigue of Composite Materials - Damage Model and Life Prediction. Composite Materials: Fatigue and Fracture, 2nd Vol., ASTM STP 1012, p. 87. Johnson, W. S. 1985. Delamination and Debonding of Materials. ASTM STP 876. Johnston, N. J., and P. M. Hergenrother. 1987. High-Performance Thermoplastics, Proceedings of the 32nd International SAMPE Symposium, SAMPE, p. 1412. Jones, R. M. 1975. Mechanics of Composite Materials. Scripta Book Co. Kinloch, A. J., and R. J. Young. 1983. Fracture Behavior of Polymers. Applied Science Publishers. Kommers, W. J. 1943a. Fatigue Behavior of Wood and Plywood Subjected to Repeated and Reversed Bending Stresses. U.S. Forest Products Laboratory Report No. 1327. Kommers, W. J. 1943b. The Fatigue Behavior of Douglas-Fir and Sitka Spruce Subjected to Reversed Stresses Superimposed on Steady Stresses. U.S. Forest Products Laboratory Report No. 1327-A. Lark, R. F., M. Gougeon, G. Thomas, and M. D. Zuteck. 1983. Fabrication of Low-Cost Mod-0A Wood Composite Wind Turbine Blades. DOE/NASA/20320-45, NASA TM-83323, February.

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Assessment of Research Needs for Wind Turbine Rotor Materials Technology Launikitis, M. B. 1978. Chemically Resistant FRP Resins. Technical Bulletin of Shell Chemical Company. Lee H., and K. Neville. 1967. Handbook of Epoxy Resins. McGraw-Hill, New York. Lee, B. L., C. K. Riew, and R. J. Moulton. 1980. Rubber Toughening of Tetrafunctional Epoxy Resin. 12th National SAMPE Technical Conference. Lee, B. L., F. H. Howard, and E. H. Rowe. 1983. Effect of Matrix Toughening on the Crack Resistance of SMC Under Static Loading. 38th Annual Conference of SPI, Session 9-A. Lee, S. M. 1986. Compression After Impact of Composites with Toughened Matrices. SAMPE Journal, March/April. Lubin, G., Ed. 1982. Handbook of Composites. Van Nostrand Reinhold, New York. Lynch, T. 1989. Thermoplastic/Graphite Fiber Hybrid Fabrics. SAMPE Journal, Vol. 25, p. 17. Mandell, J. F. 1990. Fatigue Behavior of Glass Fiber Composites. Presented at the NRC Workshop on Assessment of Research Needs for Wind Turbine Rotor Materials Technology, Washington, D.C., January 22-23. McGarry, F. J., and J. N. Sultan. 1969. Crack Phenomena in Cross-Linked Glassy Polymers. ASTM STP 460, ASTM. McGarry, F. J., E. H. Rowe, and C. K. Riew. 1978. Improving the crack resistance of BMC and SMC. Polymer Engineering Science, Vol. 18, p. 78. Modern Power Systems, July, 1986, p. 19. Morgan, C. A., A. D. Garrad, and U. Hassan. 1989. Measured and Predicted Wind Turbine Loading and Fatigue. Presented at the EWEC Conference in Glasgow, 1989. Moulton, R. J., and R. Y. Ting. 1981. Effects of Elastomeric Additives on the Mechanical Properties of Epoxy Resins and Composite Systems. International Conference on Composite Structure. Murtha-Smith, S. 1985. Loads and Fatigue Evaluation of Hamilton Standard WTS-4 Wind Turbine. Proceedings Wind Power '85, San Francisco, August 27-30, p. 136. Owen, M. J. 1974. Fatigue Damage in Glass Fiber-Reinforced Plastics. In Composite Materials, Vol. 5, ed. by L. J. Broutman, Academic Press, New York. Phillips, D. C., J. McCarthy, A. G. Davis, and J. E. P. Stott. 1987. The Use of Glass Reinforced Plastics and Advanced Composites for Wind Turbine Blades. A Review of Experience in Their Use and of Relevant Design Procedures and Codes. Department of Energy, United Kingdom, June. Properties of Epoxy Resins, Hardeners, and Modifiers. 1982. Adhesive Age, April. Reifsnider, K. L., Ed. 1980. Damage in Composite Materials. ASTM STP 775. Reifsnider, K. L., E. G. Henneke, W. W. Stinchcomb, and J. C. Duke. 1983. Damage Mechanics and NDE of Composite Laminates. Mechanics of Composite Materials. Recent Advances. Pergamon, New York, p. 399.

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Assessment of Research Needs for Wind Turbine Rotor Materials Technology Reifsnider, K. L., and W. W. Stinchcomb. 1986. A Critical Model of the Residual Strength and Life of Fatigue-Loaded Composite Coupons. Composite Materials: Fatigue and Fracture, ASTM STP 907, ed. by H. T. Hahn, p. 298. Riew, C. K., E. H. Rowe, and A. R. Siebert. 1976. Rubber Toughened Thermosets. In Toughness and Brittleness of Plastics, ed. by R. D. Deanin and A. Crugnola, American Chemical Society, Washington D.C. Rosato, D. V., and C. S. Grove, Jr. 1964. Filament Winding. John Wiley & Sons, New York. Scott, J. W. and D. C. Phillips. 1975. Carbon Fiber Composites with Rubber Toughened Matrices. Journal of Materials Science. Vol. 10, p. 551. Sendeckyj, G. P. 1981. Fitting Models to Composite Materials Fatigue Data. Test Methods and Design Allowables for Fibrous Composites, ASTM STP 734, ed. by C. C. Chamis, p. 245. Spera, D. A., J. B. Esgar, M. Gougeon, and M. D. Zuteck. 1990. Structural Properties of Laminated Douglas Fir/Epoxy Composite Material. NASA Reference Publication 1235, Report No. DOE/NASA/20320-76, May. Springer, G. S., Ed. 1981. Environmental Effects on Composite Materials. Technomic Publishing Co. Stoddard, F. S. 1989. Field Problems with Wind Turbine Rotors. Presented at the First Meeting of the NRC Committee on Assessment of Research Needs for Wind Turbine Rotor Materials Technology, Washington, D.C., November. Strong, A. B. 1989. Fundamentals of Composites Manufacturing. Society of Manufacturing Engineers. Sutherland, H. J. 1989. Analytical Framework for the LIFE2 Computer Code. Sandia Report SAND 89-1937 UC-905. Tsai, S. W. 1989. Composites Design, Think Composites. Tsai, S. W., and H. T. Hahn. 1975. Failure Analysis of Composite Materials. Inelastic Behavior of Composite Materials, ed. by C. T. Herakovich, AMD-Vol. 13, ASME, p. 73. Van Delft, D. R. V., F. Hagg, and P. A. Joosse. 1987. The Influence of Fatigue Design Line Criteria on the Rotor Blade Design. Wind Energy Conversion 1987, ed. by J. M. Galt, Mechanical Engineering Publications Ltd., London. Vinson, J. R., Ed. 1977. Advanced Composite Materials--Environmental Effects. ASTM STP 658. Weibull, W. 1939. A Statistical Theory of the Strength of Material, Royal Swedish Academy of Engineering Sciences, Proc. 151. Whitehead, R. S. 1990. Damage Tolerance of Composites. Presented at the NRC Workshop on Assessment of Research Needs for Wind Turbine Rotor Materials Technology, Washington, D.C., January 22-23. Wilkins, N. J. M. 1984. Materials Aspects of Large Aerogenerator Blades . Proceedings of the European Wind Energy Conference. Hamburg, West Germany, October 22-26, p. 281. Williams, J. G., and M. D. Rhodes. 1981. The Effect of Resin on the Impact Damage Tolerance of Graphite/Epoxy Laminates. NASA TM 83213, October. Windpower Monthly. 1987. June, p. 10.

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Assessment of Research Needs for Wind Turbine Rotor Materials Technology Witzler, S. 1988. High-Temperature Thermoplastics: A Progress Report. Advanced Composites, March/April, p. 56. Wood Handbook. 1974. Wood as an Engineering Material. U.S. Forest Products Laboratory, Agriculture Handbook No. 72. Yang, J. N., and D. Shanyi. 1983. An Exploratory Study into the Fatigue of Composites Under Spectrum Loading. Journal of Composite Materials, Vol. 17, p. 511. Zweben, C., H. T. Hahn, and T.-W. Chou. 1989. Mechanical Behavior and Properties of Composite Materials, Vol. 1. Technolmic Publishing Co.

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