Defect-Tolerant Structural Design of Wind Turbine Blades

Transcription

1 5th AIAA/ASME/ASCE/AHS/ASC Structures, Structural Dynamics, and Materials Conference<br>17th 4-7 May 29, Palm Springs, California AIAA Defect-Tolerant Structural Design of Wind Turbine Blades Kyle K. Wetzel * Wetzel Engineering, Inc., Lawrence, Kansas USA The structural design of a wind turbine rotor blade can influence its ability to withstand defects commonly resulting from industrial-quality composites manufacturing. Analysis of structures with typical manufacturing defects a practice common in other industries is not standard in the wind turbine blade industry and is not required by certification agencies. The present study focuses on the two most common approaches to utility-scale blade construction clamshell structures with the spar caps embedded in outer shells that are bonded with internal shear webs and clamshell structures with cosmetic shells bonded to a separate internal spar and their ability to resist common manufacturing defects in bond lines, including voids and thickness that is out of tolerance. The results demonstrate that these defects cause a reduction in the strength of blades of both construction methods and suggest that certification analyses should probably include typical defects in the modeling. However, the results show significant differences between the two blade designs abilities to withstand the defects. The conclusion is that blades with cosmetic shells bonded to separate internal spars are weakened by bond line defects substantially more than designs with stiffening elements embedded in the shells. This conclusion should discourage the use of this design or encourage levels of manufacturing quality control generally higher than what is frequently observed in the wind blade industry. I. Introduction Several dramatic failures of utility-scale wind turbine rotor blades have appeared in the industry news during the past few years, including the failure in 27 of several Gamesa blades that exhibited complete debonding of the shells at the trailing edge combined with significant debonding of the shells from the internal spars (Figure 1). The failure was ultimately attributed by Gamesa to a defective adhesive applicator that had applied inadequate quantities of adhesive. 1 This episode, while it may be a particularly dramatic example, is instructive regarding the relatively poor quality control present in much of wind blade manufacturing. While the obvious solution would be to improve quality control in the industry, given the current very high demand for products and the very low costs and cycle times expected by the industry for manufacturing blades, it is unlikely that significant improvement can be expected in this arena in the near future. Among the more common manufacturing defects identified by this author in his own work with manufacturers are: Adhesive Bond Defects o Thickness out of Tolerance (Figure 2) o Voids (to the point of missing adhesive), such as in the Gamesa blades. Laminate Defects o Ply Wrinkling & Waviness (Figure 3) o Misplaced Laminates Fiber Fraction Problems o Resin-Rich Regions (Figure 3) o Dry Spots * AIAA Senior Member. KyleK.Wetzel@kwetzel.com. Copyright 29 by Wetzel Engineering, Inc. Published by the American Institute of Aeronautics and Astronautics, Inc., with permission. 1

2 Figure 1. Debonding Failure of Gamesa 43m Blades Nominally 5mm Bond Gap Measured at 15mm Figure 2. Bond Gap Out of Tolerance Figure 3. Laminate Waviness and Resin-Rich Regions 2

3 However, an equally important issue has been generally overlooked some blade designs and manufacturing processes are less tolerant of manufacturing defects than alternatives. The defects in the bond lines in the Gamesa blade may have been particularly poor, but generally poor bond lines are in the personal experience of the author relatively common in the wind blade industry. Thickness of the bond line between the shells and the internal spars or shear webs, which was commonly specified at 3mm or less several years ago, is now frequently specified at 6mm or less, and yet visual inspection of blades frequently reveals gaps much larger than this. The author has observed gaps as large as 12-15mm. Voids in bond lines some up to 5mm long are observed. Nevertheless, most blades appear to be relatively tolerant of such defects. The question then arises of whether there is something inherent in the design of the Gamesa blades that made them particularly susceptible to this defect in bonding. Gamesa, like its former parent Vestas, manufactures the 43m blades in question using an approach that differs from what much of the rest of the industry uses. Where many other manufacturers (e.g., LM Glasfiber, General Electric, Suzlon, Mitsubishi, Clipper, DeWind, etc.) embed the spar caps (generally referred to in Europe as girders) into the outer shells that are then bonded to internal shear webs, Gamesa fabricates an internal box spar that contains all of the structural material. The skins are essentially cosmetic sandwich core material with no significant stiffness that are then bonded to the spars. The focus of the current study is to determine if the latter approach is more susceptible to defects in the bonds. The theory being tested is that the lack of significant stiffening material in the sandwich core shells makes the shells more susceptible to local panel buckling in the presence of defects in the bonds. A defect in the bond will create a concentration of lap shear stress at the site of the defect, which will in turn translate to a concentration of interlaminar shear stress in the underlying substrates. Such interlaminar shear promotes buckling. For blades with stiffening structure embedded in the shells, the author proposes that this structure allows the outer shells to resist such buckling. However, for blades without such material, there is little to prevent buckling of the cosmetic shells in response to these stress concentrations. This could contribute to making designs with internal spars more susceptible to bond line defects and poor manufacturing quality control. Analysis of composite structures with included defects is relatively common in some other industries, including most notably aerospace. The more sophisticated of these analyses include the influence of defects such as micro-cracks and track crack growth and progressive ply failure. Such analyses are often used to determine the minimum time interval between inspections to ensure that cracks cannot initiate and progress to complete failure during the time between inspections. Aerospace industry methods of analysis and testing also include damaged and repaired structures. In contrast, standard practice in the wind blade industry is to model essentially ideal structures, with uniformly nominal laminate thicknesses, bond gaps according to specifications, material properties according to coupon test results (less required safety factors), etc. Figure 4 presents a flowchart of current wind turbine rotor blade design practice. The engineering analysis is supported by material coupon testing, but the quality of the coupons typically reflects laboratory conditions, not fabrication quality. While some effort is made to match process cycles and fiber volume fraction, the coupons as is standard practice do not include any of the defects mentioned above. These coupon tests are used to calculate characteristic values for the stiffness and strength, but again, such calculations reflect only the significance of statistical scatter in the results in determining values of a required confidence level. Post-processing of the results does not reflect defects present in parts as fabricated. Material safety factors are applied to these characteristic values. Table 1 summarizes the material SFs required by Germanischer Lloyd for the certification of wind turbines. 2 The base safety factor, γ m, and the safety factor for process, C 3, may be intended to include discrepancies between coupon data and parts as fabricated. However, this may be dangerous to assume, as these factors account for a variety of influences, not just defects. It is not clear that there is any rigorous engineering to support a claim that these particular values are sufficient to account for the types of manufacturing defects noted above. The relatively frequent occurrence of blade failures in recent years suggests that they are not universally adequate. However, rather than simply increasing the safety factors across the board, it may be more rigorous and accurate to begin modeling some of the common manufacturing defects and design the structures to withstand defects typical of the quality of manufacturing found in the industry. Only analysis including realistic defects and appropriate safety factors can help ensure structural integrity and sufficient life. This should be applied to all typical defects. The only defects specifically addressed by GL in their guidelines are defects in bonds. Section of Reference 2 specifies that Stress concentrations within the bonding surfaces and flaws shall be taken into account. In the experience of the author, this is widely ignored or addressed unrealistically by most manufacturers. 3

4 WHERE ARE MANUFACTURING DEFECTS IN CURRENT DESIGN PRACTICE Lab Quality Confidence Only Maybe? Too Late Material Coupon Testing Testing with Low Loads Safety Factors But Not on the Test Blade Characteristic Stiffness & Strength Manufacturing Of Course Materials Safety Factors Design & Analysis Not As Required Certification Installation & Operation Figure 4. Flowchart of Current Blade Design Process Field Failures Partial Safety Factor Table 1. Material Safety Factors per Germanischer Lloyd x=a Lam Strength x=b Lam Fatigue x=c Static Stability x=d Bond Strength x=e Bond Fatigue γ m Base C 1x 1.35 Aging n/a Type of Structure 1.5 Aging C 2x Temperature C 3x Process Fabric n/a 1.1 Surface Reproducibility C 4x Post Cure II. Objective of the Present Study While there are many defects than could be studied, the author has chosen to focus on what he believes to be the most common manufacturing defect, which is poor bond line quality. The objectives of the present study are Present a method for including manufacturing defects in the Blade Design Process Determine if embedding the spar caps into the outer shells creates a structure that is more resistant to bond line defects than bonding cosmetic outer shells to an internal spar structure. 4

5 III. Method of Analysis Two wind turbine rotor blades with bond line defects have been subjected to structural analyses using finite element methods. All analyses have been conducted in accordance with the requirements of IEC Standard and the requirements of GL for certification of wind turbines. 2,3 A. Blade Designs Both blades exhibit the same aerodynamic contour. A 44m blade designed for a 2.MW variable-speed, pitch-regulated wind turbine has been employed for the present study (Figure 5). This blade features a conventional aerodynamic design using Delft University (DU) airfoils of thickness 25% to 4% through the midspan and NACA636XX airfoils of thickness 18-21% near the tip (Figure 6). The internal laminate schedule for the two blades is essentially the same. The structural model of the two blades differs primarily as shown in Figure 7. For Blade A, the spar cap is embedded in the outer shell, and the shells are bonded to the shear webs. For Blade B, the spar is an integral I beam that is bonded to the outer cosmetic shells. Figure 5. 44m Blade Analyzed in the Present Study Figure 6. Three-View of the 44m Blade Analyzed in the Present Study 5

6 Adhesive Bonds Embedded Spar Caps (a) Blade A Adhesive Bonds Integrally Molded I beam Spar (b) Blade B Figure 7. Two Blade Structural Configurations Analyzed in the Present Study B. Bond Line Defect Modeling The following bond line defects have been modeled for each blade: 1. Excessive bond gaps for the main shell-spar or shell-web bonds (Figure 8). The thickness of the bond gaps, nominally specified as no more than 5mm, were modeled as 5mm, 8mm, 12mm and 15mm. 2. Voids in the bond lines for the main bonds (Figure 9). This analysis was conducted assuming a uniform bond thickness of the nominal 5mm. The length of the region of missing adhesive was varied from 5cm to 25cm. Elements in the bond gap were removed in regions of the blade previously shown to be critical for buckling. C. Finite Element Models A finite element model of the blade was built in ANSYS. 4 The finite element model was constructed using SOLID186 2-node layered brick elements, as shown in Figure 1. The element features a higher-order quadratic displacement behavior and large strain and stress stiffening capabilities. Accurate modeling required approximately 1.16million elements and 17 million degrees of freedom. 6

7 Bond Gap, Nominally 5mm Modeled as 5, 8, 12, and 15mm (a) Blade A Figure 8. Models of Bond Gaps Out of Tolerance (b) Blade B Figure 9. Voids in Bond Line The use of brick elements as opposed to shell elements allows for more accurate determination of interlaminar and lap shear stresses at the bond lines. The use of bricks also enables more accurate modeling of the bonds between the shells and the shear webs or spar. Using shell elements requires all of the nodes to be placed around the outer contour of the shell, which in turn requires that the tops and bottoms of the shear webs extend all the way to the outer shell (Figure 11). This is obviously not accurate and represents a significant deterioration of model fidelity that compromises accurate calculation of the stresses in the bond line. The maximum distributed design loads for the blade were distributed as point loads to the nodes on the spar caps. The nodes at the blade root were constrained against motion with respect to all degrees of freedom. D. Laminate Strength and Ply Failure Analysis The laminate structures of the blades were modeled layer-by-layer (layer material, thickness, and orientation). The blades feature carbon UD prepreg girders and shells constructed of glass double bias and triaxial (/±45) nonwoven fabrics. PVC foam is used forward and aft of the girders for sandwich core reinforcement. Details of the laminate structure are proprietary and cannot be published. They are not particularly pertinent to the analyses or results presented here. Table 2 summarizes the material properties employed in the present study. Note that the base material safety factor was set to 1. because in the present analysis wherein defects are being included, it was desired to determine what the margin of safety was without considering GL s base safety factor that in part accounts for short-comings in manufacturing. Figure 12 summarizes the lap shear strength of the adhesive used in the bonding of the actual 44m blade. This data is the result of proprietary testing. The strength reflects both the influence of bond thickness as well as exposure to hot, humid conditions (5C, 1% relative humidity for 3 days). Because these tests already include the influence of aging and temperature, we did not apply further safety factors to the lap shear values. 7

8 Figure 1. Solid (Brick) Element Geometry of a Wind Turbine Blade Figure 11. Shell Element Geometry of a Wind Turbine Blade Cross Section 8

9 Each of the blade models was subjected to a large displacement static solution (i.e., so-called nonlinear buckling ). The analysis was performed using progressive ply failure. The methods of the Puck criteria for fiber failure and interfiber failure were applied to determine if a layer had failed. 5 For the adhesive bond layer, the failure criteria was whether lap shear stress exceeded the lap shear strength of the bond. In the event of an exceedance of the failure criteria, the failed layer was removed (element by element), the stiffness matrix was recalculated, and the solution continued. This resulted in a redistribution of the stresses. Global failure occurred when the stresses could not be redistributed without resulting in cascading failure of surrounding layers or the same layer in surrounding elements. Table 2. Material Properties Employed in the Present Analysis Carbon Glass Triax Glass DB PVC Foam Adhesive UD Axial Tensile Modulus E11 GPa Normal Tensile Modulus E22 GPa Shear Modulus G12 GPa Density ρ kg/m Axial Tensile Strength R11+ MPa Axial Compressive Strength R11 MPa Normal Tensile Strength R22+ MPa Normal Compressive Strength R22 MPa Shear Strength R12 MPa Lap Shear Strength S 13 [MPa] Non post cured RT Test 5 Test, Hot Humid Accelerated Aging Bond Gap Thickness t [mm] Figure 12. Lap Shear Strength of the Adhesive 9

10 IV. Results & Discussion Figure 13 illustrates the influence of the main bond line thickness on the load factor at failure. At the nominal bond thickness of 5mm, Blades A and B demonstrate load factors at failure of 1.51 and 1.56, respectively. Recall that these results do not include the material safety factors required by GL of γ m and C 3, the product of which is Therefore, the results indicate that Blades A and B are marginally sound for the nominal bond thickness. The first failure mode with the nominal bond thickness is static instability (i.e., buckling), which results in large strain in the carbon girder and subsequent failure of the carbon fibers due to compressive failure. This is illustrated in Figure 14. Increasing the bond line thickness has a very small impact on the load factor at failure out to 12mm. Only for the 15mm bond gap does the load factor drop below the 1.49 that would be required to satisfy GL s safety factors, assuming that the latter do not already intend to partially account for out-of-tolerance bond thickness. Even at 15mm, however, the influence is small, and could rather easily be addressed by the addition of a small amount of material to the girders. Although the lap shear strength of the adhesive drops with increasing bond thickness, the stress in the bond line also drops, and so the influence on the load factor at failure is weak. These results suggest that bond line thickness should not be a source of overwhelming concern. Figure 15 illustrates the influence of voids in the bond line on the load factor at failure. The results show that for both blades, the buckling load factor drops with the presence of voids and continues to drop further as the extent of the voids increases. However, the effect is significantly larger with Blade B, so much so that Blade B no longer exhibits positive margin of safety under these conditions. For blade A, the void creates a concentration of lap shear, and the bond fails before the blade buckles. The failure is cascading to the extent that once the stress concentration at the edge of the void reaches the lap shear strength of the adhesive, the failure propagates rapidly in both directions the bond line essentially unzips. But this occurs as a result of stress in the bond line itself. This causes the sudden drop in minimum load factor that is observed for Blade A moving from the baseline case (no void) to the minimum 5cm long void. As the extent of the void is increased, the concentration of stress at the edge of the void increases, and so the minimum load factor drops. Increasing the void length from 5 to 25 cm (2 to 1 inches) reduces the minimum load factor from 1.32 to That is significant, and suggests that very extensive voids would result in premature failure of the blade Load Factor at First Failure LF Blade A Blade B Bond Line Thickness t [mm] Figure 13. Influence of Bond Line Thickness on Failure Load Factor 1

11 Blade A Blade B Load Factor at Failure Figure 14. Base Buckling Mode 1.6 Load Factor at First Failure LF Blade A Blade B Bond Line Void Length ΔL [cm] Figure 15. Influence of Bond Line Void Length on Failure Load Factor 11

12 But the influence of voids on the structural response of Blade B is quite dramatic. The influence of a small 5cm void is not as severe as for Blade A, but for more extensive voids, the minimum load factor for Blade B drops very rapidly. For the small 5cm void, the first failure mode for Blade B remains buckling of the spar. However, for the 1cm void and all the longer voids, the first failure mode shifts to panel buckling of the cosmetic shell at the location of the void, which induces high stresses in the bond line that result in a cascading failure of the bond line. As with Blade A, the bond unzips. However, the root causes are different for the two blades. For Blade A, the root cause is simply a stress concentration in the bond due to the void. In Blade B, the stress concentration is exacerbated by the panel buckling of the cosmetic shell. This accounts for the very dramatic drop in minimum load factor as the void length is increased. The increasing extent of the void reduces the load factor at which the cosmetic shell buckles by increasing the length of the shell that is unreinforced by the spar. Given the very uncertain nature of such voids, the results obtained here are bothersome, as they suggest that blades of Design B could be susceptible to significantly premature failures during extreme wind events as a consequence of typical and rather unpredictable manufacturing defects. V. Conclusions and Recommendations The conclusions of the present study are: Blades with embedded spar caps might be more tolerant of voids in adhesive bonds. Bond gap thickness is not a strong influence on structural integrity. Standard safety factors may be inadequate to accurately reflect the influence of some bond line defects. The first conclusion should discourage the use of designs with cosmetic shells bonded to internal spars or encourage levels of manufacturing quality control generally higher than what is frequently observed in the wind blade industry. The author recommends that the wind industry generally reassess standard practice for analysis of blade structures. Specifically, he recommends that as standard practice, blade designers should Model composite structures with finite brick elements rather than shell elements Include all realistic defects in the finite element model Adjust the safety factors down to account for the fact that defects are included in the analysis This should result in a more accurate and realistic assessment of structural integrity than the present practice of rather blindly assuming ideal structures and applying safety factors of dubious reliability. VI. References K. Mellott, Gamesa pinpoints turbine-blade flaw, The Tribune-Democrat (Johnstown, PA), May 11, 27. Guidelines for the Certification of Wind Turbines, Edition 23, Germanischer Lloyd, Hamburg, 1 November 23. Wind turbines Part 1: Design requirements, Ed. 3, IEC Standard 614-1, International Electrotechnical Commission, Geneva, August 25. ANSYS, ver. 11., ANSYS, Inc. Canonsburg, Pennsylvania. A. Puck, Festigkeitsanalyse von Faser-Matrix-Laminaten Modelle für die Praxis, Carl Hanser Verlag, Munich,