Structural investigation of composite wind turbine blade considering structural collapse in full-scale static tests

Composite Structures 97 (2013) 15–29

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Structural investigation of composite wind turbine blade considering structural collapse in full-scale static tests Jinshui Yang a,⇑, Chaoyi Peng a, Jiayu Xiao a,⇑, Jingcheng Zeng a, Suli Xing a, Jiaotong Jin b, Hang Deng b a b

College of Aerospace and Materials Engineering, National University of Defense Technology, Changsha 410073, China Zhuzhou Times New Material Technology Co., Ltd., Zhuzhou 412007, China

a r t i c l e

i n f o

Article history: Available online 23 November 2012 Keywords: Polymer–matrix composites Structural investigation Wind turbine blade Videometrics

a b s t r a c t This study is concerned with an actual collapse testing under the flap-wise loading for a large full-scale composite wind turbine blade, and a discussion is conducted to assess and evaluate the structural response of the blade during loading and after collapse by correlating experimental findings with numerical model predictions. A videometrics technique is adopted to measure the integral deformation and the local deformation of the wind turbine blade under the flap-wise loading. The measured results show that the displacement of the blade tip is up to 11 m at the ultimate load which is 160% of the extreme design load for the tested blade. A simple method is proposed to identify the exact failure location of the blade based on the deformation data. The thorough analysis results indicate that the aerodynamic shells debonding from the adhesive joints is the initial failure mechanism causing a progressive collapse of the blade structure. Ó 2012 Elsevier Ltd. All rights reserved.

1. Introduction Wind energy is a renewable clean energy and inexhaustible. It is an environmental friendly energy and can help in reducing the dependency on fossil fuels. Wind energy is expected to play an increasingly important role in the future energy scene [1,2]. The wind turbine blades that play a key role to collect wind energy are generally regarded as one of the most critical component of the wind turbine system [3]. The wind turbine system converts the kinetic energy of the wind to electrical energy by rotating the blades. The wind turbine blade is a load-carrying aerodynamic structure, which consists of the suction and pressure aerodynamic shells together with shear webs as shown in Fig. 1. As shown in Fig. 1, the upper shell is the aerodynamic suction side and the lower shell is the aerodynamic pressure side. The aerodynamic shells and the shear webs are separately manufactured and then assembled to the final blade by using adhesive joints. In addition to making up the outer geometry and defining the aerodynamic efficiency, the aerodynamic shells are also the load-carrying structure of the blade. The spar caps in the aerodynamic shells (see Fig. 1) are the main load-carrying structure, which can be pre-fabricated as one part of the shells. The function of the shear webs is limited to supporting the aerodynamic shells and transferring shear forces.

⇑ Corresponding authors. Tel.: +86 0731 84576502; fax: +86 0731 84576578. E-mail addresses: (J. Xiao).

[email protected] (J. Yang),

[email protected]

0263-8223/$ - see front matter Ó 2012 Elsevier Ltd. All rights reserved. http://dx.doi.org/10.1016/j.compstruct.2012.10.055

The modern large-scale wind turbine blades are generally made of polymer–matrix composites material system. As shown in Fig. 1, there are sandwich structures and thick laminates in a blade, the involved materials can be divided into the following three groups: resin, reinforcement material and core material for sandwich structures. The resin used in the blade is currently thermosetting resin system such as polyester and epoxy resin. Although glass fiber materials is the primary reinforcement material in the composite blade industry, carbon fiber materials is also considered as a candidate of reinforcement to manufacture large-scale composite blades. A variety of synthetic polymer foams and different wood sorts are used as core materials in the sandwich structures. Although different manufacturing philosophies are widely used in the wind turbine industry, the primary manufacturing technique for large-scale composite wind turbine blades is Vacuum Infusion Molding Process (VIMP) that also known as Vacuum Assisted Resin Transfer Molding (VARTM) or vacuum bag molding. The aerodynamic shells and the shear webs can be fabricated in one process by using the VIMP technique, respectively. The technique is characterized by the fact that only one side of the mold is solid contrary to Resin Transfer Molding (RTM) where a double solid sided mold is used. Furthermore, an applied vacuum is used as the driving force for transferring the resin into the reinforcement. While designed appropriately, the wind turbine blade should be able to withstand the extreme load, the fatigue and other hostile environment over the whole lifetime. Therefore, designers must carefully consider the structural behavior of the blades in their structural design and must test the full-size structure. Structural

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Fig. 1. Illustration of the aerodynamic and structural definitions of the large-scale wind turbine blade under consideration.

design requirements, such as minimum blade tip clearance limit, strain limits along the fiber direction, surface stress limit, and fatigue lifetime over 20 years, are specified by the design requirements of the IEC 61400-1 international specification and the Germanischer-Lloyd (GL) regulations [4–6]. To achieve the desired wind turbine blades, a better understanding of the structural behavior on a different scale is necessary. Detail structural behaviors beyond the elastic range must be investigated. These may include the identification of failure modes which lead to ultimate collapse, and the full-scale test to verify the design and manufacturing procedure as well as to gain confidence of the blade’s in-service structural integrity. A 15 kW wind turbine blade has been designed and tested by Habali and Saleh [7]. The static test indicated that the blade could withstand loads ten times the normal working thrust. A 750 kW horizontal axis wind turbine blade has typically been optimized by experimental tests and simplified analytical methods in [3]. From the results of full-scale static test, it was found that the designed blade had structural integrity and the measured results agreed well with the analytical results. Torsional performance of wind turbine blades has been investigated through experiments and numerical simulation in [8,9]. The flap-wise and edge-wise bending stiffness and torsional stiffness as well as the bend-twist coupling were measured by an experimental investigation on a blade section in [8]. Then, the experimental results have been successfully compared with numerical finite element models, see [9]. A scaled-down prismatic wind turbine blade section has been tested to collapse in a 4-point flap-wise bending, see [10,11]. An interlaminar shear failure originated from local bending due to an initial geometric imperfection has been argued, which triggered a progressive collapse of the wind turbine blade section [10,11]. In [12] an acoustic emission diagnostic technique has been used to monitor the structural behavior of wind turbine blade during a static test. A distributed and controlled static load was applied to the blade using a four-point Whiffle-tree arrangement. Step load increase and hold sequence was repeated until the wind turbine blade breaks. In [13] a 25-m wind turbine blade manufactured by layered orthotropic and isotropic materials has been tested to collapse in

flap-wise bending. The results show that the ultimate strength of the tested blade is governed by instability phenomena in the form of delamination and buckling. Interaction between both instability phenomena occurs causing a progressive collapse of the blade structure. Then a geometrical nonlinear and interlaminar progressive failure finite element analysis of the tested blade has been done and successfully compared with the experimental findings [14]. It was found that the predictive numerical models shown excellent correlation with the experimental findings and observations in the pre-instability response. Based on [13] and [14], a modeling strategy has been developed for the structural analysis of large three-dimensional laminated composite wind turbine blades under geometric and material induced instability, see [15]. Compression strength of a fiber composite main spar in a wind turbine blade has been investigated by Jensen [16], which has great focus on the so-called ovalization effect demonstrated by the Brazier effect, see [17]. The investigations have shown that the blade spar-cap deflects non-linearly and the load carrying box girder of the blade ovalizes during the flap-wise loading. Then a 34-m composite wind turbine blade has been tested to collapse under flap-wise loading, see [18]. Prior to this collapse test, the blade had passed all static and dynamic tests required by the classification authorities. Based on experimental results and finite element calculations, the non-linear Brazier effect has been characterized by a crushing pressure which causes the ovalization in [18]. Local displacement measurements has been conducted and compared with finite element simulations to identify the failure mechanism of the wind turbine blade. Basically, it has been argued that the Brazier effect is the main reason for the structural collapse of the blade. Although some authors [3,7–18] have made some effort in identifying the reason for the structural collapse of the wind turbine blades, there are still not much applied results with practical significance that can be used to optimize the design of large-scale wind turbine blade. Another design parameter (i.e. bending–torsion coupling [19]) induced into the design of large-scale wind turbine blade may have great influence on the structural failure model of the blades. In addition, the size of wind turbine rotors has

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increased greatly in the past decade (i.e. from 40 m to more than 120 m diameter), which means more structural instability phenomena and more complicated location failure. Thus, more experience with practical significance is necessary. In this study, in order to gain some understanding and insight to the chain of events leading to the collapse of large-scale composite wind turbine blades, a 40 m wind turbine blade manufactured by VIMP with E-glass/epoxy composite materials has been tested to collapse under flap-wise loading in a full-scale static test. Based on experimental measurements during testing and post-mortem observations after testing, the structural behavior and failure mode of large-scale wind turbine blade was analyzed, and the exact failure location of the blade was identified by a proposed method. This work will be useful for improving the design of large-scale wind turbine blades. Note that all experimental and numerical simulation values are given in a non-dimensional form considering a respect for the blade manufacturing. 2. Full-scale static test of wind turbine blade 2.1. Materials and lay-ups Three kinds of glass fabrics, used as reinforcement materials in the test wind turbine blade, are illuminated in Table 1. The glass fabrics were provided by Chongqing Polycomp International Corp. in China. Photographs of fabrics are shown in Fig. 2. An epoxy resin system, 760E/766H supplied by Dow Chemical Company is used as matrix material which is a low-viscosity resin system for vacuum infusion molding process in the wind turbine blades. 760E is resin and 766H is additive. Density, viscosity and gelation time of 760E/766H at room temperature (25 °C) are tabulated in Table 2. Adhesive of adhesive joints in the wind turbine blade is 770E/ 778H adhesive system supplied by Dow Chemical Company. 770E is an epoxy resin and 778H is an epoxy hardener. Density, viscosity and gelation time of 770E/778H at room temperature (25 °C) are also tabulated in Table 2. PVC foam and Balsa wood provided by MID (China) Composite Company Ltd. are used as core materials in the sandwich structure. Density of PVC and Balsa is 63 ± 6 kg/m3 and 150 ± 20 kg/m3, respectively. As shown in Fig. 1, the blade lay-ups could be divided into three parts: thick laminates in the spar cap and in the blade edge, sandwich structure in the blade shells and sandwich structure in the shear webs. Specification, materials and schematic of lay-ups for the test blade are illuminated in Table 3. Fiber volume fraction of fiber laminates in the blade structure is approximately 60%. 2.2. Full-scale collapse test A 40 m full-scale wind turbine blade made of E-glass/epoxy for a class horizontal axis wind turbine system, manufactured by VIMP, was tested to failure under flap-wise loading as shown in Fig. 3. Prior to the collapse test, the blade had passed all static tests Table 1 Properties of fabrics. Fabric

Fiber

Specification

Unidirectional Biaxial

Glass Glass

Triaxial

Glass

EKU1150(0)/50E-120 EKB800(45,-45)E1270 EKT1200(0,±45)E1270

Superficial density/g m2 0°

+45°

45°

Total

1250 0

0 404

0 404

1250 808

715

250

250

1215

17

required by the classification bodies. The full-scale test has been conducted at the Blade Test Centre of Zhuzhou Times New Materials Technology Co., Ltd. A flap-wise load is applied corresponding to the aerodynamic lift resulting in compressive and tensile strains in the aerodynamic suction and pressure sides of the blade, respectively. In order to describe clearly the test procedure and results, the blade under consideration was divided into three test regions as shown in Fig. 4. The blade is fixed as a cantilever beam by connecting the root to the test stand. The flap-wise load is obtained by applying shear force at Z = R0.329, R0.462, R0.655 and R0.809 in the local Z-direction giving a flap-wise moment Mx, where R represents the normalized rotor radii. The load is applied to the corresponding radii by means of load clamps, via longitudinal load fixture beams and two or three cranes with ample capacity. The ultimate load is applied stepwise and the load increments are shown in Table 4. The blade was tested in six sequential test steps with increasing load level until the extreme design load (ext), and the final failure event determining the ultimate structural strength of the blade (ult) found in the 5th and 6th test step, respectively, see Table 4. Additional load means the load, additionally applied, after neutralization of the bending moment caused by the own weight of the blade and the weight of the load clamps. 2.3. Deformation measurement In the present study, a videometrics technique is adopted to monitor the large-scale integral deformation of the wind turbine blade and small-scale deformation on the blade aerodynamic surface. The videometric technique is based on the principle of convergent photogrammetry and intended to restore the three dimensions of an object from photographs, which could be captured by common consumer cameras. This method employs at least two photographs from different angles of the three-dimensional specimens, which have been prepared by applying a grid on the surface. In the photographs, a reference item (the target) is included, e.g., a cube with a grid of known dimensions. During analysis, the target is used to define a coordinate system and to calibrate the photographs for the dimensions. Then, the grid points on the sample are identified and computer analysis of the photographs results in the coordinates of the grid points, from which the deformation can be determined. More detail of our implementation of the videometrics technique is available in [20,21]. 2.3.1. Large-scale integral deformation The amplitude of the blade integral deformation under the extreme design load will achieve approximately 10 m, thus a large view field is required and the precision required is relatively low. Taking measuring precision and view field into account, a large view field and a remote (about 30 m) intersection by three still cameras were adopted in case of large-amplitude deformation measurement. Three still cameras (a Nikon D300S camera with a high resolution of 4288  2848 pixel) can keep the blade tip in the view field when it reached the maximum displacement. Fig. 5 shows the schematic of the experimental arrangement for the three still cameras in the blade integral deformation measurements. The relationship among the still cameras to measure the integral deformation is of absolute orientation. Note that the spar cap is the main loading structure of the blade range from the root to the tip, so the integral deformation of the blade can be obtained by monitoring the displacements of the middle axis of the spar cap. More detail of our implementation is available in [21]. 2.3.2. Small-scale deformation on aerodynamic surface Small-scale deformation monitoring is concerned with the minimal deformation of the aerodynamic surface (e.g. warpage and

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(a) Biaxial

(b) Triaxial

(c) Unidirectional

Fig. 2. Photographs of various fabrics.

which connected with a personal computer (PC). The computers are used to control the camera’s operation and store the captured image data. More detail of our implementation is available in [21].

Table 2 Infusion resin and adhesive properties at room temperature (25 °C). Resin

Infusion resin Adhesive

Specification

Mass ratio

760E/766H

100:30 ± 2

770E/778H

100:47 + 5

Density (g cm3)

Viscosity (mPa s)

Gelation time (min)

1.08– 1.12 1.38– 1.39

200–300

350

Thixotropic

200

out-of-plane deflection) and thus relatively high precision is required. The camera adopted for the small-scale deformation measurement was a MVC1SAM-GE14 camera with resolution of 2592  1944 pixels. Fig. 6 shows the schematic of the experimental arrangement for small-scale deformation monitoring in the blade full-scale static tests. The monitoring area (about 3 m length) in the test region 2 is monitored by a group of stereo cameras

2.4. Strain gauge recording Strain gauges are used to monitor the local strain of the wind turbine blade structure under flap-wise loading. The size of wind turbine blade is so large that strain monitoring cannot cover all areas of the blade and only some key points can be monitored. In this study, it focuses on the strain response of some key points at shear webs and spar caps. Strain gauges are located at the leading edge (LE) shear web along longitudinal direction as shown in Fig. 7a, Row 1 is close to the suction shell (SS) and Row 2 is close to the pressure shell (PS). Strain gauges are also located at the trailing edge (TE) shear web along the longitudinal direction as shown in Fig. 7b, Row 3 is close to the suction shell (SS) and Row 4 is close to the pressure shell (PS). In addition, strain gauges are also instrumented on aerodynamic surface at the suction and pressure side

Table 3 Lay-ups of the test wind turbine blade. Type

Specification

Materials

Thick laminate

Thick laminates in the spar cap and in the blade edges

Unidirectional (UD), biaxial and triaxial glass fabrics

Sandwich structure I

Sandwich structure in the blade shells

Biaxial and triaxial glass fabrics; PVC foam and Balsa wood

Sandwich structure II

Sandwich structure in the shear webs

Biaxial glass fabrics and PVC foam

Schematic of lay-ups

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Table 4 represents the normalized flap-wise moment w.r.t. ultimate Load increments M max x flap-wise moment. Load increment

Additional load

M max x (%)

1. Step

Relieving the component from its dead weight and setting all measurement points back to zero 40% extreme design load 60% extreme design load 80% extreme design load 100% extreme design load

0

160% extreme design load

100

2. 3. 4. 5.

Step Step Step Step (ext) 6. Step (ult)

25.0 37.5 50.0 62.5

Fig. 3. A full-scale wind turbine blade under flap-wise loading static test.

along the middle axes of spar caps (rows 5 and 6) as shown in Fig. 7c. A strain gauge, BX120-15AA, provided by Testing Instrument Factory Huangyan Zhejiang in China was used in the test. Gauge resistance and factor is 120.2 ± 0.1 X and 2.08 ± 1%, respectively. The data was continuously collected with a rate of 0.1 Hz and by a trigger file. 2.5. Numerical model The finite element (FE) method has traditionally been used in the development of wind turbine blades mainly to investigate the structural behavior of blade. The purpose of using FE modeling in this work is to correlate the observed structural response from the full-scale test with the results of the predictive models in order to validate the main reason of the blade structural collapse. It is important to correlate the experimentally obtained data with the numerical predictions to gain a further understanding of the sequence of events leading to structural blade collapse. Non-linear characteristics of the wind turbine blade must be considered in the structure numerical calculation due to the large deformation and geometric nonlinearity of the blade structure under the wind loading. ABAQUS has been chosen for this analysis because it has a great capability to make structural analysis considering many nonlinear conditions at the same time, such as material nonlinearity, contact and geometrical nonlinearity. 2.5.1. Finite element and mesh size Wind turbine blades are long, flexible and slender composite structures with cross-sectional dimensions much smaller than the length. The elements named S4R in ABAQUS element library is chose to model the wind turbine blade in the present work. S4R is a 4-node, quadrilateral, first-order and stress/displacement shell element with reduced integration and a large-strain formulation as shown in Fig. 8. Number 1–4 refers to the node and ‘‘1’’ refers to the integration point in Fig. 8. So S4R is commonly recom-

mended when large strains or very high strain gradients are expected and can be used for problems prone to bending-mode hour glassing, in areas where greater solution accuracy is required or for problems where in-plane bending is expected. Fig. 9 shows the numerical model of the wind turbine blade. Typical meshes in the numerical model of the wind turbine blade include mesh1 in the spar cap and in the blade shell edges, mesh2 in the sandwich structure of the blade shells and mesh3 in the sandwich structure of the shear webs as shown in Fig. 10. The in-plane dimension of mesh1, mesh2 and mesh3 is 100  100 mm, 100  100 mm and 100  60 mm, respectively. Mesh1’s thickness varies with the thickness of thick laminates in the spar cap and in the blade shell edges, while the thickness of mesh2 and mesh3 varies with the thickness of sandwich structure in the blade shells and in the shear webs as shown in Table 5. 2.5.2. Non-linearities analysis The non-linear behavior of structures originates from four sources, which can be identified in terms of material, geometric, load and displacement boundary conditions [15]. In this work, only geometric sources of non-linear behavior are taken into account. The wind-turbine blade structure has been considered as an idealized model without any manufacturing imperfections for finite element analysis. In addition, numerical simulation is focused on the correlation with respect to local out-of-plane displacements, strains and the possible existence of limit points originating from geometric sources of non-linear behavior. 3. Results and discussion 3.1. Integral deformation The integral deformation of the full-scale wind turbine blade during the flap-wise loading is displayed in Fig. 11. The data as shown in Fig. 11 was obtained by the videometrics technique, which is a non-contact measurement technique without constraints of conventional measurement techniques. It is impossible

Fig. 4. Definition of test regions of the wind turbine blade in its global rotor XYZ-coordinate system.

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Fig. 5. Schematic of the experimental arrangement for the three still cameras in the blade integral deformation measurement.

(a) Instrumentation of the strain gauges at leading edge shear web

Fig. 6. Schematic of the experimental arrangement for the small-scale deformation measurement in the wind turbine blade full-scale static test.

to record the integral deformation of the blade at the collapse moment by using conventional measurement techniques because the sudden collapse results in no sufficient response time for the techniques. Fortunately, the integral deformation of the full-scale blade at the collapse moment can be recorded by the videometrics technique. As shown in Fig. 11, the ultimate load (the collapse load) of the wind turbine blade is 160% of the extreme design load in our test. The displacement of the blade tip at the extreme design load is up to 7 m, and it is up to 11 m at the ultimate load. Note that the blade is a bending–torsion coupling structure, which could be inferred from the results as shown in Fig. 11. The post-mortem observation after the final failure sometimes can only identify an approximate failure region of the blade (see Fig. 12), and cannot identify the exact failure location of the blade. To find out the main reason for the structural collapse of the blade, it is necessary to identify the exact failure location of the collapsed blade. Based on the integral deformation of the blade, a simple method is proposed to identify the exact failure location of the blade in this paper. Fig. 12 shows that the collapse region of the blade is at R0.3654–R0.5385. The integral deformation of the blade under the extreme design load and the ultimate load (flap-wise load) is shown in Fig. 13. Detail of the collapse region (R0.3654–R0.5385) is shown as Area A in Fig. 13. As shown in Area A (see Fig. 13), the collapse region can be divided into three ranges: range I, range II and range III. There is a turning point at the junction of two ranges. Point A is the turning point at the junction of range I and range II, and point B is the turning point at the junction of range II and range III. If the wind turbine blade is designed appropriately, the strain levels should be constant through the length of the wind turbine blade when subjected to equivalent aerodynamic operational loading. The turning point means that some strain fluctuations occurred along the length of the blade. The probability of collapse at the turning point is greater than that at other positions. Therefore, it can be assumed that the blade failure will occur at point A or point B or a point between A and B. From the post-mortem observation, it is quite evident that the failure location of the blade is neither at point A, nor at point B.

(b) Instrumentation of the strain gauges at trailing edge shear web

(c) Instrumentation of the strain gauges along the spar caps Fig. 7. Schematic illustration of the instrumentation of the strain gauges in the wind turbine blade full-scale static test.

Fig. 8. 4-Node reduced integration element.

It should be at a point between point A and point B. Extended the line in range I and the line in range III, two extended lines intersect at one

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Fig. 9. Finite element model of the wind turbine blade.

Fig. 10. Typical meshes in the numerical model of the wind turbine blade.

point in range II. As shown in Fig. 13, the intersection point is the turning point of the whole blade. The blade section near the turning point will be subjected to large localized non-linear bending and stress concentration increasing with additional load. The ultimate strength of the blade is governed by geometric and material induced instability phenomena, i.e. local buckling and delamination and debonding. Increasing stress concentration near the turning point results in interaction among instability phenomena causing a progressive collapse of the structure. It indicates that the intersection point should be the failure location of the blade. Based on the above-mentioned analysis, it can basically determine that the exact failure location of the collapsed blade is near the region of R0.52. The exact failure location of the blade is very important for identifying the main reason of the blade structural collapse or identifying the instability phenomena governing the ultimate strength of a wind turbine blade. 3.2. Small-scale deformation Fig. 14 shows the extraction results and images of small-scale deformation on the aerodynamic surface (pressure side) in the blade full-scale static tests. The monitoring area is concerned with region R0.335—R0.409 in test region 2, which includes the structural region that governs the ultimate strength of the entire blade when considering the extreme design load of a 50-year recurrence storm. Grid landmarks (see Fig. 14) as the characteristic labels for small-scale deformation measurement were drawn in the monitoring area of the blade aerodynamic surface.

Fig. 11. Integral deformation of the full-scale wind turbine blade under a flap-wise static test.

Fig. 12. The collapse region.

As shown in Fig. 14, there are 340 monitoring points in the monitoring area. Fig. 15 shows the small-scale deformation of the blade aerodynamic surface in monitoring region R0.335—R0.409 under flap-wise loading. Fig. 16 shows the Y-direction displacement of the monitored points on the blade aerodynamic surface under flap-wise loading. As shown in Fig. 16, the maximum experimental displacement of the monitored points is up to 120 mm under the extreme design flap-wise loading. And there is a snap of the displacement under additional load from 40% extreme design load to 60% extreme design load, while the differences between the displacements under 60%, 80% and 100% extreme design load are

Table 5 The typical mesh size in the numerical model of the wind turbine blade. Mesh

Specification

In-plane dimensions/mm

Thickness/mm

Mesh1 Mesh2 Mesh3

Thick laminates in the spar cap and in the blade edges Sandwich structure in the blade shells Sandwich structure in the shear webs

100  100 100  100 100  60

Mesh thickness varies with the thick-laminates’ thickness Mesh thickness varies with the sandwich-structure’s thickness Mesh thickness varies with the sandwich-structure’s thickness

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The blade sections subjected to bending moments can suffer a non-linearity in their bending response and may undergo crushing forces from the rotated compressive or tensile forces due to the curvature that arises from bending. The moment of the airfoil cross-section, and the structure’s ability to resist bending loads, lead to a non-linear bending response. This phenomena is called the Brazier effect, which is best illustrated in the classical example of flattening a circular cross-section into a quasi-oval form which eventually will crush the structure, since it no longer is able to sustain the applied bending moment [17]. Points 1, 18, 35, 52, 69, 86, 103, 120, 137, 154, 171, 188, 206, 222, 239, 256, 273, 290, 307, 324 will be connected to form a curve (see Fig. 14), which is the semi-contour of airfoil section at R0.337. Fig. 18 shows the displacement variation of the monitored points on the curve under flap-wise loading. The variation implies that the aerodynamic surface of pressure side deform gradually into a convex mode shape at the given rotor radii. It indicates that the aerodynamic shell deform in a convex and concave mode shape manner. The convex and concave mode shape suggests that there is the influence of the Brazier effect on the local out-of-plane deflection of the aerodynamic airfoil section. However, there is also no clear evidence to indicate that the Brazier effect is the dominant failure mode for the tested blade. 3.3. Strain gauge recording

Fig. 13. Schematic of the simple method for identifying the exact failure location of the blade based on the integral deformation of the blade.

small. Fig. 17 displays the non-linear local deflection on the aerodynamic surface of the blade. This non-linear bending response suggests that the blade has an initial geometrical imperfection, damage or both, which enforces a non-uniform deflection response of the aerodynamic surface. Unfortunately, the small-scale deformation of the monitored aerodynamic surface at the collapse moment (under the ultimate flap-wise loading) was not recorded in the present work. Because of the monitoring area moving with additional load, it is not in the visual field of the cameras located on the ground at the collapse moment.

The spar caps in the aerodynamic shells (see Fig. 1) are the main load-carrying structure of the wind turbine blade. The strain gauges row 5 and row 6 are located at both spar caps in the suction shell (SS) and the pressure shell (PS) along the middle axes of the spar caps to record the spar-cap strain response during the flapwise loading. The longitudinal strain response of the spar cap recorded during the flap-wise loading is displayed in Fig. 19, where Fig. 19a shows the strain response for gauge row 5 at the middle axes of the SS-spar cap and Fig. 19b shows the strain response for gauge row 6 at the middle axes of the PS-spar cap. As shown in Fig. 19, the location with the highest strain level is near the region of R0.476–R0.548. It indicates that the collapse location of the blade is near the region of R0.476–R0.548, which is consistent with the analyzed result that the failure location of the blade is near the region of R0.52 as mentioned in Section 3.1.

Fig. 14. Images and extraction results of small-scale deformation on the blade aerodynamic surface under flap-wise loading. The monitoring area (R0.335–R0.409) in test region 2 is about 3 m length. Cameras are located on the ground.

J. Yang et al. / Composite Structures 97 (2013) 15–29

Fig. 15. Small-scale deformation on the blade aerodynamic surface under flap-wise loading. The monitoring area (R0.335—R0.409) in test region 2 is about 3 m length.

Compared Fig. 19a with Fig. 19b, it is obvious that the maximum strain recorded during the flap-wise loading is compressive strain at suction side of the blade. It indicates that the blade collapse should be caused initially by the compressive failure of one structure at suction side. There are spar cap, shell sandwich structure,

23

edge thick laminate structure, adhesive joint and shear web with sandwich structure in the blade structure, see Fig. 1. The experimental results does not reveal anything about what happens prior to the collapse point that could determine the initial failure part of the blade and explain the reason for the large difference in strain levels at R0.476–R0.548 compared to the strain level in the remaining blade. Consequently, in order to identify the main reason of the blade structural collapse, it is necessary to determine the initial failure part of the blade structure under the flap-wise loading. The spar cap is an unidirectional E-glass fiber thick laminates structure, its ultimate failure was caused usually by a compressive fiber failure mode of the individual delaminated unidirectional laminate. If the failure occurs initially from the spar cap, the blade structure suffers from intralaminar crushing damage at a main fracture line, which is usually perpendicular to the middle axes of the spar cap. Fig. 20 displays the compressive failure mode of the spar cap for a full-scale wind turbine blade made of E-glass/ polyester. There is no evidence that the spar cap of the present blade failed with the same compressive failure mode as the E-glass/polyester blade. As shown in Fig. 12, although the large deformation occurred in the spar cap, it is still not collapsed with any mode. In addition, the maximum strain recorded during test does not closely correlate with the compressive fiber failure strain of the unidirectional E-glass/epoxy laminate as used in the blade spar cap. Consequently, it suggests that the spar cap do not define the limit point of the present blade and the failure should occur initially from other structures. Fig. 21a displays the strain response for gauges row 1, row 2 located at the leading edge (LE) shear web along longitudinal direction, and Fig. 21b displays the strain response for gauges row 3,

Fig. 16. Y-direction displacement of the monitored points on the blade aerodynamic surface under flap-wise loading.

Fig. 17. Y-direction displacement of some monitored points as a function of the additional load.

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Fig. 18. Displacement variation of the monitored points on the curve under flap-wise loading.

(a) Strain response for gauge row 5 at middle axes of SS-spar cap

(b) Strain response for gauge row 6 at middle axes of PS-spar cap Fig. 19. Normalized longitudinal strain eult y w.r.t. ultimate longitudinal strain as function of the normalized rotor radii R for strain gauges row 5 and row 6 along middle axes of spar caps, where M ult x represents normalized load levels w.r.t the ultimate flap-wise moment.

row 4 located at the trailing edge (TE) shear web along longitudinal direction. Row 1, row 3 are close to the suction shell (SS), row 2, row 4 are close to the pressure shell (PS). The strain response for gauges row 1, row 2, row 3 and row 4 is all the longitudinal strain response, which are near the adhesive joints between the aerodynamic shells and the shear webs. The main stress of the shear webs is shear stress under the flap-wise loading, and the maximum shear strain usually occurs at the middle of the shear web. Therefore, the longitudinal strain responses of row 1, row 2, row 3 and row 4 actually reflect the strain variation of the adhesive joints under flap-wise loading. The strain variation of gauges (row 1–row 4)

is consistent with that of the spar caps (gauges row 5 and row 6) at a level of 62.5% for the ultimate moment ðM ult x ¼ 62:5%Þ as shown in Fig. 21. It indicates that the aerodynamic shells and shear webs are connected into an integral structure by the adhesive joints. The stress transferring results in 20–40% decrease of strain value from the spar cap to the shear web, see Fig. 21. According to this conclusion, it is reasonable to infer that the maximum stress of the adhesive joints suffered under the ultimate loading will be up to 80% of the maximum stress of the spar caps. This maximum value does correlate closely with the failure strain of the adhesive based on the data provided by the supplier. Although the adhesive joints

J. Yang et al. / Composite Structures 97 (2013) 15–29

Fig. 20. Compressive failure mode of the spar cap for a full-scale wind turbine blade made of E-glass/polyester.

debonding is expected to have a major influence on the blade final collapse, unfortunately no data was recorded under the ultimate flap-wise loading for gauges row 1–row 4. Therefore, there is no enough evidence to support that the adhesive joints is an initial failure part of the blade structure. To verify this, it is need more data which can supply some effective evidences. 3.4. Numerical analysis Fig. 22 shows the finite element model results of the longitudinal stress response under the extreme design flap-wise loading for the wind-turbine blade structure. Fig. 22a relates to the compressive

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stress at the blade suction side and Fig. 22b relates to the tensile stress at the blade pressure side. The results indicate that the spar caps are the main load-carrying structure with the maximum stress–strain response during the flap-wise loading. The highest stress–strain region is located near R0.471 clearly indicated by the stress–strain response in the idealized finite element model as shown in Fig. 23. It is not the case as illustrated in Fig. 19 and as mentioned in Section 3.1, which reveals that the region R0.52 is the critical damage region of the tested blade. The experimental response suggests a geometrical imperfection in the damaged region of R0.52. The difference between the experimental response and the numerical model results indicates that the actual damage region can be readily mistaken with a small and uninteresting geometric imperfection in the finite element model. Comparing experimentally measured strain response of the gauges row 1 and row 3 with numerical modeling predictions, an agreement is found to be acceptable, see Fig. 24. Due to measuring difficulties, it is impossible to measure the shear strain of the shear webs during our test. Therefore, numerical modeling is chose to predict the shear strain of the shear webs in the present work. The detailed model with shell elements in the shells and shear webs is as shown in Fig. 25. Fig. 26 shows the numerical-modeling results for the shear strain at the middle axes of shear webs. As shown in Fig. 26, there are five sections at R0.231, R0.327, R0.461, R0.654 and R0.808, respectively, with larger strain fluctuation along the rotor radii. Section at R0.231 is in the test region 1 as shown in Fig. 3, and the strain direction at this section is different from that at the other four sections. It is not difficult to find that sections at R0.327, R0.461, R0.654 and R0.808 are the loading positions of the full-scale blade test. And

(a) Strain response for gauges row 1and row 2 at LE-Shear web

(b) Strain response for gauges row 3 and row 4 at TE-Shear web Fig. 21. Normalized longitudinal strain eult y w.r.t. ultimate longitudinal strain as function of the normalized rotor radii R for strain gauges row 1, row 2, row 3 and row 4 at two shear webs for recording at ðMult x ¼ 62:5%Þ.

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J. Yang et al. / Composite Structures 97 (2013) 15–29

(a) Compressive stress at the suction side

(b) Tensile stress at the pressure side Fig. 22. The finite element model results of the longitudinal stress response under the extreme design flap-wise loading for the wind-turbine blade structure.

indicate that the Brazier flattening effect has a great influence on the shear-deformation of the shear webs at these sections. If the Brazier effect is the dominant failure mode of the blade, the shear-web failure and the blade collapse location should be at the above-mentioned five sections. In fact, the exact failure location of the blade is near the region of R0.52, and the shear web at those sections is still no damage after the blade collapse. It indicates that the Brazier effect is not the dominant failure mode for the tested blade, and the shear web is not the initial failure part of the blade structure in the present work. In [13] a finite element model also has been used to simulate the local out-of-plane deflection of the blade cross-section in relation to the local bending, which is influenced by the Brazier effect. The model results show that the ovalization or flattening of the cross-section reduces the cross-sectional bending stiffness and also flatten the edge-wise curvature of the cross-section. However, it is still no practical influence can be verified that the Brazier effect is the dominant failure mechanism of the wind turbine blade. Fig. 23. The maximum stress–strain region in the finite element model under the extreme design flap-wise loading for the wind-turbine blade structure.

the maximum shear-strain appears at section R0.808, which is in the test region 3 as shown in Fig. 3. These fluctuations observed

3.5. Post-mortem observations The main function of sandwich structure in the blade shells is to provide aerodynamic shape for the blade and to form an integral blade structure. Although some deformations can be observed

J. Yang et al. / Composite Structures 97 (2013) 15–29

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Fig. 24. Comparison of experimental test and numerical modeling results for the strain response of gauges row 1 and row 3 at two shear webs.

Fig. 25. Detailed model of the blade structure.

Fig. 26. Numerical-modeling results for normalized shear strain at middle axes of shear webs.

following the deformation of the spar cap, there is no obvious damage in the shell sandwich structure and it still maintains the structural integrity, see Figs. 12 and 27. It suggests that the failure be impossible to occur initially from this sandwich structure. As shown in Fig. 27, the suction shell debonding from adhesive joints forms a large convex deformation in the collapsed region. Debonding trace can be observed at the adhesive-joint location inside the suction shell of the blade. These post-mortem observations during and after testing support that the adhesive joints between the aerodynamic shells and the shear webs is the initial

failure part of the blade structure. Based on the previous analysis, it can be confirmed that the aerodynamic shells debonding from the adhesive joints is the initial failure mechanism followed by its instable propagation which can lead to collapse. When the debonding reaches a certain size the interaction among instability phenomena becomes critical and causes a progressive collapse of the structure by multiple local buckling-driven delamination propagation processes in the primary load-carrying cap material. The interlaminar fracture processes primarily develops and propagates under almost a pure model II energy dissipation [14] and with a

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J. Yang et al. / Composite Structures 97 (2013) 15–29

Fig. 27. Post-mortem observation of collapsed region.

local buckling deformation of the spar cap (see Fig. 27) without any local buckling of the delaminated regions. This debonding results in an enforcing chain of events, which produces a structure buckling and subsequently a sudden structural collapse following compressive fiber failure in the delaminated cap material. The large released energy of the sudden structural collapse sends large shock waves through the blade structure, which results in a wide range of structure fracture, shell peeling, buckling deformation, fiber fracture and failure, see Fig. 27. 3.6. Final failure–buckling mechanism As a structure, the wind turbine blade is impossible to be a perfect body without any structural defects original from manufacturing defects and design defects. The existence of the structure defect would result in stress concentration inside the blade structure under the ultimate loading. The exact failure location of the blade is determined based on this stress concentration assumption in the abovementioned Section 3.1. The determined result is also verified by strain recording and post-mortem observations. Once the stressconcentration region of the blade originated from the blade structure defects under the flap-wise loading, fracture or damage points will occur with increasing stress in this region. Then the fracture size begins to expand from a progressive damage evolution state to a sudden and complete unrecoverable structural collapse. In addition, the aerodynamic shells debonding from the adhesive joints is the initial failure mechanism in the fracture propagation process. The structure stress-concentration results in an initial debonding in the adhesive joints between the aerodynamic shells and the shear webs, then a progressive debonding evolution starts with increasing loading. Then, instability phenomena (i.e. local buckling and delamination) and its interaction start and propagate with the progressive debonding evolution. When the combined effect of debonding and local buckling and delamination reaches a certain extent, a sudden collapse of the blade occurs. Consequently, the major failure mechanisms that caused the collapse of the blade are the structure stress-concentration, and subsequently shell debonding from the adhesive joints amplified by the local buckling and delamination.

4. Conclusions The present work has primarily been concerned with an actual collapse testing under the flap-wise loading for a large full-scale composite wind turbine blade, and a discussion is conducted to

assess and evaluate the structural response of wind turbine blade during loading and after collapse by correlating experimental findings with numerical model predictions. Thereby taking a discussion to identify the main failure mechanism in the wind turbine blade. A videometrics technique is adopted to measure the integral deformation and the small-scale deformation of the wind turbine blade under the flap-wise loading. The measured integral-deformation results show that the displacement of the blade tip is up to 11 m at the ultimate load which is 160% of the extreme design load for the tested blade. Based on the integral deformation results and the stress concentration assumption, a simple method is proposed to identify the exact failure location of the blade, which is near the region of R0.52. This conclusion is also verified by the strain recording and the post-mortem observations. The measured small-scale deformation suggests that the aerodynamic shell deform in a convex and concave mode shape manner, which reflect the influence of the Brazier effect on the local out-of-plane deflection of the aerodynamic airfoil section. Combined with numerical predictions, it is concluded that the Brazier effect is not the dominant failure mode for the tested blade. The possible initial-failure part explaining the plausible sources and multiple phenomena leading to the structural collapse of the blade has been documented by a thorough analysis. The thorough analysis results indicate that the aerodynamic shells debonding from the adhesive joints is the initial failure mechanism followed by its instable propagation which can lead to collapse. Once the stress-concentration region originated from the blade structure defects under the flap-wise loading, initial debonding will occur with increasing stress in this region. Then the debonding size begins to expand from a progressive damage evolution state to a sudden and complete unrecoverable structural collapse. Acknowledgements This work is fund by the national natural science foundation of China (51103177) and the national ‘‘863’’ plan project of China (2012AA03A205) and the science & technology project of Hunan Province (2011FJ1001). The authors would like to acknowledge the support of the Cooperative Zhuzhou Times New Material Technology Co., Ltd. for this work. References [1] Ezio S, Claudio C. Exploitation of wind as an energy source to meet the world’s electricity demand. Wind Eng 1998;74–76:375–87.

J. Yang et al. / Composite Structures 97 (2013) 15–29 [2] Joselin Herbert GM, Iniyan S, Sreevalsan E, Rajapandian S. A review of wind energy technologies. Renew Sust Energy Rev 2007;11:1117–45. [3] Kong C, Bang J, Sugiyama Y. Structural investigation of composite wind turbine blade considering various load cases and fatigue life. Energy 2005;30:2101– 2114. [4] IEC International Standard. Wind turbine generator system—Part I: Safety requirements; 1994. [5] Technical Note: IEC 1400-1 GL Test Regulation; 2000. [6] Lloyd Germanischer. Regulations for the certification of wind energy conversion system. Germany: Germanischer Lloyd; 1999. [7] Habali SM, Saleh IA. Local design, testing and manufacturing of small mixed airfoil wind turbine blades of glass fiber reinforced plastics Part I: Design of the blade and root. Energy Conv 2000;41:249–80. [8] Berring P, Branner K, Berggreen C, Knudsen HW. Torsional performance of wind turbine blades—Part I: Experimental investigation. In: ICCM16-16th international congress on composite materials, Kyoto, Japan, CD-ROM; 8–13 July, 2007. p. 10. [9] Branner K, Berring P, Berggreen C, Knudsen HW. Torsional performance of wind turbine blades—Part II: Numerical validation. In: ICCM16-16th international congress on composite materials, Kyoto, Japan, CD-ROM; 8–13 July, 2007. p. 10. [10] Kühlmeier L. Buckling of wind turbine rotor blades. Analysis, design and experimental validation. PhD thesis, Aalborg University, Denmark; 2007, ISBN: 87-91464-00-5. [11] Kühlmeier L, Thomsen OT, Lund E. Large scale buckling experiment and validation of predictive capabilities. In ICCM15-15th international conference on composite materials, Durban, South Africa, CD-ROM; 27 June-1 July, 2005. p. 10.

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[12] Rumsey MA, Musial W. Application of acoustic emission nondestructive testing during wind turbine blade tests. J Sol Energy Eng 2001;123:270. [13] Overgaard LCT, Lund E, Thomsen OT. Structural collapse of a wind turbine blade. Part A: Static test and equivalent single layered models. Compos Part A: Appl Sci Manu 2010;41:257–70. [14] Overgaard LCT, Lund E, Thomsen OT. Structural collapse of a wind turbine blade. Part B: Progressive interlaminar failure models. Compos Part A: Appl Sci Manu 2010;41(2):271–83. [15] Overgaard LCT, Lund E, Camanho PP. A methodology for the structural analysis of composite wind turbine blades under geometric and material induced instabilities. Compos Struct 2010;88:1092–109. [16] Jensen FM. Compression strength of a fibre composite main spar in a wind turbine blade. Ris-r-1393, RisNational Laboratory; 2003, ISBN: 87-550-3184-6, ISBN: 87-550-3184-4, ISSN: 0105-2840. [17] Brazier LG. On the flexure of thin cylindrical shells and other thin sections. Proc R soc Lond Ser A 1927;11(6):104–14. [18] Jensen FM, Falzon BG, Ankersen J, Stang H. Structural testing and numerical simulation of a 34 m composite wind turbine blade. Compos Struct 2006;76:52–61. [19] De Goeij WC, Van Tooren MJL, Beukers A. Implementation of bending–torsion coupling in the design of a wind-turbine rotor-blade. Appl Energy 1999;63:191–207. [20] Legac A. Videogrammetry or digital photogrammetry: general remarks, methodology, applications. Proc SPIE 1994;2350:16–21. [21] Yang JS, Peng CY, Xiao JY, Zeng JC, Yuan Y. Application of videometric technique to deformation measurement for large-scale composite wind turbine blade. Appl Energy 2012:22. http://dx.doi.org/10.1016/j.apenergy. 2012.03.040.