Fatigue Failure of a Composite Wind Turbine Blade at Its Root End

8/18/2019 Fatigue Failure of a Composite Wind Turbine Blade at Its Root End

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Fatigue failure of a composite wind turbine blade at its root end

Hak Gu Lee ⇑, Min Gyu Kang, Jisang Park

Wind Turbine Technology Research Center, Korea Institute of Materials Science, 797 Changwondaero, Changwon, Gyeongnam 641-831, Republic of Korea

a r t i c l e i n f o

Article history:

Available online 7 August 2015

Keywords:

Fatigue

Failure

Delamination

Wind turbine blade

Bumping motion

a b s t r a c t

As blade failures at wind farms have increased, the structural safety of composite wind turbine blades is

ever more important. The recent implementation of considerably larger blades has made the problem

even more crucial. One of the critical failure modes is the blade root failure, which can result in the bladebeing pulled out from its wind turbine during operation. In this study, we experienced delamination fail-

ure at the blade root during fatigue testing of a 3 MW full-scale wind turbine blade according to interna-

tional standard IEC 61400-23: full-scale structural testing of rotor blades. Comparing the measured data

with the FE analysis results, we simulated the situations the blade had experienced, and then found what

caused the delamination failure as well as the problem of the conventional design approach. The bump-

ing motions of the blade shell caused by geometric complexities between the maximum chord and the

root alter significantly the load distribution at the end of the blade root. Therefore, to enhance the struc-

tural safety of a large composite wind turbine blade, a more detailed FE analysis on the blade root in the

design stage is needed.

2015 Elsevier Ltd. All rights reserved.

1. Introduction

With the recent trend toward large slender wind turbine blades,

questions are being raised regarding their reliability. In order to

evaluate the static strengths and fatigue lives of these larger

blades, static and fatigue tests of full-scale prototype blades should

be conducted according to international standards or equivalent

guidelines [1–4]. Testing methodologies developed so far are well

described in the two review papers of Malhotra et al. [5] and

Yang et al. [6].

Previous studies pertaining to static strength of a full-scale

wind turbine blade are as follows. Jensen et al. [7] tested a 34 m

composite wind turbine blade until its structural collapse.

Debonding of the outer skin was the initial failure mechanism, fol-

lowed by delamination buckling which led to the blade’s collapse.

Jensen et al. believed the main root cause was the Braizer effect of

the shell structure due to bending. Overgaard et al. [8–9] carried

out a static flapwise bending test of a 25 m wind turbine blade

to collapse. The Brazier effect had a large influence on the local

out-of-plane deflection, but its influence on the longitudinal strain

level in the primary load-carrying laminate was insignificant.

Overgaard et al. assert that the structural stability of the generic

wind turbine blade has been governed by buckling and the

delamination phenomena. Yang et al. [10] conducted a static test

of a 40 m wind turbine blade under flapwise loading to collapse.

Yang et al. concluded the Brazier effect was not the dominant

failure mechanism, but debonding between the pressure-side and

the suction-side aerodynamic shells was the initial failure mecha-

nism followed by its instable propagation which leads to collapse.

Previous studies pertaining to fatigue of wind turbine blades are

divided into two categories: a material fatigue behavior and a

structural fatigue behavior. The material fatigue behavior has

extensively studied with uni-axial, in-plane loading of balanced

and symmetrical, relatively thin laminates [11,12], but they are

only remotely representative for blade structures [13]. Tests with

a full-scale wind turbine blade to study the structural fatigue

behavior are so expensive that few studies have been conducted

to date. Leeuwen et al. [14] had carried out fatigue tests of 37 wind

turbine blades 3.4 m in length as well as 35 coupons to compare

fatigue strength from full-scale blade tests with coupon-based

predictions. Flapwise failures occurred at the tensile side, but edge-

wise failures were the result of crack initiation starting in the

bonding at the trailing edge followed by further crack propagation

in the laminate. Blade fatigue data compared with coupon data

fitted reasonably with flapwise tests, but they did not compare

well for edgewise tests. Marín et al. [15,16] inspected fatigue

damage of a 300 kW wind turbine blade, and then performed a

FE analysis to reveal the root cause of the fatigue damage. The

crack initiated at the abrupt geometric-transient region between

the root zone and the aerodynamic zone had been propagated into

the laminate. It should be noted that the aforementioned studies

http://dx.doi.org/10.1016/j.compstruct.2015.08.010

0263-8223/ 2015 Elsevier Ltd. All rights reserved.

⇑ Corresponding author. Tel.: +82 55 280 3261; fax: +82 55 280 3498.

E-mail address: [email protected] (H.G. Lee).

Composite Structures 133 (2015) 878–885

Contents lists available at ScienceDirect

Composite Structures

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dealt with small wind turbine blades that have relatively higher

fatigue margins than the large slender wind turbine blades

presently in development.

Another approach to enhance the reliability of a wind turbine

blade is an FE fatigue simulation for the blade. Kong et al. [17,18]

designed a 750 kW wind turbine blade, factoring in its fatigue life

of 20 years based on the well-known S–N linear damage equation,

the load spectrum, and Spera’s empirical equations. Shokrieh et al.[19] performed a case study with a 23 m wind turbine blade. Using

its FE shell model with a stochastic approach on fatigue loads, the

fatigue life was bounded between 18.66 years and 24 years as

lower and upper limits. Toft et al. [20] estimated the reliability of

a wind turbine blade for a single failure mode, considering statisti-

cal uncertainties.

Despite such contributions by many researchers, about 30 blade

failures are occurring per year throughout the world, and the num-

ber of blade failure are increasing over time [21]. There have been

blade failures not observed in previous studies which have

occurred in the field at Eclipse wind farm and Ocotillo wind farm

in 2013. In these instances, the wind turbine blades were pulled

out from the wind turbines due to delamination at the root. The

authors of this paper have also experienced a similar phenomenon

during fatigue testing of a 3 MW full-scale wind turbine blade. To

find a root cause of the phenomenon, loading conditions calculated

by its FE shell model were applied to the more detailed FE solid

model of the root subcomponent. Comparing the analysis results

with measured strain data of the T-bolt, we adjusted loading con-

ditions of the FE solid model to simulate deformations and stress

distributions of the blade root. Based on the simulation results, this

study has found one of the plausible root causes able to incur

delamination at the root of a wind turbine blade.

2. Test blade and its failure during fatigue testing

The test blade is a 3 MW glass/epoxy composite blade as shown

in Fig. 1. Blade length and weight are 56 m and 14.5 ton, respec-

tively. It has been developed as a result of a R&D project fundedby Korean government. After mounting the test blade on a stand

fixture like a horizontal cantilever beam, the fatigue test setups

including an aerodynamic fairing, two additional masses, a flap-

wise exciter, and an edgewise exciter were attached on the test

blade. Then we carried out a dual-axis resonance fatigue test of

the blade using two different resonance frequencies according to

international standard IEC 61400-23 [1]. The tip motion of the test

blade during dual-axis resonance fatigue testing is shown in Fig. 2.

After flapwise 510,000 cycles under the equivalent amplitude of

5352 kNm and the mean of 5970 kNm and edgewise 780,000

cycles under the equivalent amplitude of 4454 kNm and the mean

of 0 kNm at the end of the blade root, fatigue failure was found asshown in Fig. 3.

Strain values of the T-bolt, located as shown in Fig. 4, were mea-

sured during the fatigue testing. The shape and specification of the

T-bolt are represented in Fig. 5 and Table 1. The graphs in Fig. 6 are

the measured strain values of the two different strain gages

attached at the same cross section of the T-bolt. In each graph

the amplitude of tensile strains larger than that of compressive

strains means separation of the T-bolt joint has occurred.

Fig. 1. 3 MW test blade 56 m in length.

Fig. 2. Tip motion during dual-axis resonance fatigue testing.

Fig. 3. Failure at the end of the blade root.

H.G. Lee et al. / Composite Structures 133 (2015) 878–885 879

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Furthermore, the tensile or compressive amplitude of strain gage 2

being larger than that of strain gage 1 means bending of the T-bolt.

Separation and bending of the T-bolt joint are unexpected

phenomena that must be avoided in the design stage. Before going

into the detailed root cause analysis, the conventional approach

for a blade root design will be explained in the next section to

understand why the approach failed to predict the separation

and bending observed.

3. Conventional blade root design

Fig. 7 represents a schematic diagram of a blade root part

including a pitch bearing assembled by T-bolts and nuts. The blade

root is a very thick composite laminate able to enclose T-bolts and

cross nuts, its thickness being about 100 mm. Thus, blade designers

have believed that, compared with blade shell sandwich structures

whose thickness are less than about 30 mm, a blade root is so stiff

that the stress distribution in it is similar to that of a hollow circu-

lar cylinder structure when subjected to bending. Based on this

presumption, the distributions of local moments and axial stresses

or forces have been calculated as shown in Fig. 7. This conventional

approach for the test blade gives 14 kNm for the maximum local

moment and 267 kN for the maximum axial force at each T-bolt

joint. The axial force is much smaller than the pretension 340 kN

for the T-bolts in Table 1. Thus the separation and bending of the

T-bolt cannot be observed during the fatigue testing based on this

calculation. The presumption regarding the blade root had worked

well before wind turbine blades became larger and more slender.

Fig. 4. Location of the T-bolt where bolt strains were measured.

Fig. 5. Schematic diagram of the M36 T-bolt used in this study.

Table 1

Specifications of the M36 T-bolt used in this study.

Grade Min. diameter

[mm]

Pretension

[kN]

Prestress

[MPa]

Min. yield strength

[MPa]

10.9 28 340 552 940

Fig. 6. Strain values measured by two different strain gages attached at the same cross section of the T-bolt: (a) strain gage 1 and (b) strain gage 2.

880 H.G. Lee et al. / Composite Structures 133 (2015) 878–885


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However, the current trend is requiring a more detailed analysis on

the blade root part.

4. FE simulation for root cause analysis

A flowchart of the root cause analysis carried out in this study is

represented in Fig. 8. Using the test setup and the loading condi-

tions aforementioned in Section 2, a static analysis with the FEshell model for the test blade was conducted to calculate the

plausible amplitudes of an axial force and a local moment at each

T-bolt. The calculated values do not represent the real situation

during the fatigue testing because the clamped boundary

conditions used in the FE shell model does not match with the

separation of the T-bolt joint observed. The axial force and the local

moment were applied to the subcomponent FE solid model for the

test blade root, and then we modified them, comparing the calcu-

lated T-bolt strains with the measured strains during the fatigue

testing. After several modifications, stress distributions at the

end of the blade root were able to explain the observed delamina-

tion that occurred.

The information on the FE shell model for the test blade is in

Fig. 9 and Table 2. The shell model reflects the shape of the test

blade and laminating sequences of composites. We used a com-

mercial FE solver, ABAQUS 6.13, and its 4 node shell element, S4R.

The number of the elements is 57,969. The boundary conditionwas the clamped condition at the end of the blade root, and the

loading conditions were the flapwise and the edgewise test bend-

ing moment distributions along the positive y- and x-directions,

where the positive y means a chord direction toward the trailing

edge in the pitch zero section and the positive x means the cross

product of the positive y with the pitch axis.

Properties of unidirectional glass NCF/epoxy composites used in

this study are in Table 4. Four properties were measured from

coupon tests: E 1 of 40.14 GPa, E 2 of 12.30 GPa, v 12 of 0.26, and

Fig. 7. Schematic diagram of a blade root part and a pitch bearing.

Fig. 8. Flowchart of the root cause analysis taking the fatigue testing conditions into account.


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G12 of 3.40 GPa, where the subscripts 1, 2, 3 mean the fiber

direction, the transverse direction, and the thickness direction,

respectively. A transversely isotropic material needs 5 independent

material properties, so we assumed v 23 of glass NCF/epoxy

composites as 0.38 compared with S-2 glass/epoxy composites in

Ref. [22], the ratio of 1.48 between two Poisson’s ratios v 12 and

v 23. Then G23 can be calculated from Eq. (1), resulting in

4.44 GPa, and the other properties E 3, v 13, and G13 are the same

as E 2, v 12, and G12, respectively.

G23 ¼ E 2

2ð1 þ v 23Þ ð1Þ

The information on the subcomponent FE solid model for the

test blade root is in Fig. 10 and Table 3. ABAQUS 6.13 was also used

as the FE solver for the subcomponent model, which was

constructed with the 8 node solid element, C3D8I . The number of

the element was 53,629. The boundary conditions used in the

model were the axially symmetric and symmetric condition along

the hoop direction, the fixed condition along the blade length

(spanwise) direction on the positions of two bearing ball arrays,

and the contact conditions with the friction coefficient of 0.3 on

the several contact surfaces of the T-bolt and the nuts. The loading

conditions were the axial force and the local moment calculated by

the FE shell model, which were applied to the cross section of the

blade root part along the spanwise direction and the hoop direc-

tion, respectively.

Equivalent orthotropic properties of the blade root laminate

whose stacking sequence is [45/0/45]n were generated for the

convenience of FE modeling with solid elements. Classical lami-

nated plate theory (CLPT) cannot calculate whole equivalent ortho-

tropic properties because even interlaminar stresses at the

interface of two laminae are discontinuous [23]. CLPT gives us only

in-plane laminate properties such as E x, E y, v xy, and G xy and

through-thickness Poisson’s ratios such as v xz and v yz , which are

calculated from Eqs. (2)–(4) [22].

v xz ¼ð A

111F 1 þ A

112F 2 þ A

116F 6Þ

2HA111

ð2Þ

v yz ¼ð A

121F 1 þ A

122F 2 þ A

126F 6Þ

2HA122

ð3Þ

F i ¼XN

k¼1

S k13Q k1i þ

S k23Q k2i þ

S k36Q k6i

t k

ði ¼ 1;2;6Þ ð4Þ

where Aij, S kij ,

Q kij, t k, and 2H are the i, j component of A matrix, the i, j

component of the transformed compliance matrix in the kth lamina,

the i, j component of the transformed reduced stiffness matrix in

the kth lamina, the thickness of the kth lamina, and the totalthickness of the laminate, respectively. The six properties were

calculated using Table 4 and the stacking sequence [45/0/45]s, in

which the thicknesses of 45/45 lamina and 0 lamina were

0.15 mm and 0.60 mm, respectively. The large staking number of

the unsymmetric laminate [45/0/45]n makes its properties

converge into those of the symmetric laminate [45/0/45]s, so we

used the symmetric stacking sequence instead of the unsymmetric

one. The remaining three properties, E z, G xz , and G yz were calculated

from a FE cube model whose stacking sequence is [45/0/45]10s. By

applying normal forces or shear forces to the cube surfaces, we

obtained a pertinent deformation value at each case. From the

deformation value and loading conditions, the cube stiffness was

calculated at each case. Table 5 shows the calculated equivalent

orthotropic properties of the blade root laminate.

Fig. 9. FE shell model used in this study.

Table 2

FE shell model for the test blade.

FE solver Element type No. of the

elements

Boundary condition

ABAQUS

6.13

4 node shell element

(S4R)

57,969 Clampe d c ondition at

the root

Fig. 10. FE solid model used in this study.

Table 3

Subcomponent FE solid model for the test blade root.

FE solver Element type No. of the

elements

Boundary conditions

ABAQUS 6.13 8 node solid

element(C3D8I)

53,629 Axially symmetric condition

Symmetric condition

Fixed condition

Contact condition (l = 0.3)

Table 4

Properties of the glass NCF/epoxy unidirectional lamina.

E 1[GPa]

E 2[GPa]

E 3[GPa]

v 12 v 13 v 23 G12[GPa]

G13[GPa]

G23[GPa]

40.14 12.30 12.30 0.26 0 .26 0 .38 3 .40 3.40 4.44

Table 5

Properties of the [45/0/45]n laminate.

E x[GPa]

E y[GPa]

E z [GPa]

v xy v xz v yz G xy[GPa]

G xz [GPa]

G yz [GPa]

30.69 13.44 12.70 0.42 0 .21 0 .34 6 .26 3.55 4.21



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5. Analysis results

The analysis result of the FE shell model shows that the

maximum values of the axial force and the local moment per each

T-bolt are greater than those of the conventional model. As shown

in Fig. 11(a), the blade root end near 110 degrees from the leading

edge receives severe tension, 387 kN, even larger than pretension

of the T-bolt, 340 kN, resulting in the separation of the T-bolt joint.

The local moment and the axial force in the FE shell model greaterthan those in the conventional model in Fig. 11 come from the

inward bumping motion in Fig. 12, and the lower values come from

the outward bumping motion. The strangest location for these

differences is the trailing edge. The location is expected to move

inward because the flapwise and the edgewise bending moment

are applied along the positive y- and the positive x-direction.

However, it moved outward, as shown in Fig. 12, resulting in a

small axial force and a small local moment. This opposite moving

direction at the trailing edge may be caused by the geometric com-

plexities from the maximum chord of the blade to the root. In the

same cross section of the blade, the alleviation of the axial force

and the local moment in some location incurs an increase of the

axial force and the local moment in another location. Thus the

bumping motions of the blade are thought to be the main reasonof the unfavorable load distribution at the end of the blade root.

Fig. 11. Comparison between the results of the FE shell model and those of the conventional approach: (a) axial force distribution and (b) local moment distribution.

Fig. 12. Bumping motion of the blade shell during the fatigue testing.

Fig. 13. Comparison between the measured and the calculated strain ranges.


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Fig. 14. Partial separation between the blade root and the pitch bearing.

Fig. 15. Stress distributions that incur delamination at the end of the blade root.

Fig. 16. Schematic diagram of the blade root failure caused by delamination followed by crack propagation.



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The subcomponent FE solid model simulated well the T-bolt

joint when applying 100% of the axial force and 73% of the local

moment at the location of 90 degrees. The measured and

calculated strain values at the top and bottom positions of the

T-bolt are shown in Fig. 13; the measured strain value at the top

position was extrapolated based on the two measured data of dif-

ferent strain gages. The simulation shows us partial separation

between the blade root and the pitch bearing as shown inFig. 14. Furthermore, the interlaminar shear stress, r23, and the

peel stress, r33, are very severe when the partial separation occurs.

The positions of the two severe stresses are well matched with the

observed delamination positions as shown in Fig. 15. Therefore, we

conclude that the delamination at the end of the blade root due to

partial separation followed by crack propagation into the root

laminate, as shown in Fig. 16, would bring about the failures in

which blades are pulled out from their wind turbines.

6. Conclusion

In this study we have experienced delamination failure at the

end of the blade root during its full-scale fatigue testing. To find

what caused the failure, FE analyses were carried out using a sub-

component FE solid model as well as a full-scale FE shell model.

The analysis results reveal that for a slender and large wind turbine

blade the real load distribution at the root is very different from

that calculated by the conventional approach, which assumes the

blade root has enough stiffness to be modeled as a bending of a

hollow circular cylinder. The bumping motions of the blade shell

alter load distribution at the end of the blade root, resulting in

the alleviation of load in some locations and the increase of load

in other locations. The severe increase of load incurs partial

separation of the T-bolt joints followed by delamination at the

end of the root, which may lead to pulling of the blade out from

its wind turbine during operation. Therefore, detailed analyses on

the blade root should be carried out to enhance its structural safety

especially for a slender and large wind turbine blade.

Acknowledgements

This work was supported by the New & Renewable Energy of

Korea Institute of Energy Technology Evaluation and Planning

(KETEP) grant funded Korea government Ministry of Trade,

Industry and Energy (No. 2012T100201707).

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