A national laboratory of the U.S. Department of Energy

Office of Energy Efficiency & Renewable Energy

National Renewable Energy Laboratory 

Innovation for Our Energy Future 

Structural Testing of 9 m 

Carbon Fiber Wind Turbine 

Research Blades 

Preprint 

J. Paquette 

Sandia National Laboratories 

J. van Dam and S. Hughes 

National Renewable Energy Laboratory

To be presented at the AIAA 2007 Wind Energy Symposium 

Reno, Nevada 

January 8–11, 2007 

Conference Paper 

NREL/CP-500-40985 

January 2007 

NREL is operated by Midwest Research Institute ● Battelle Contract No. DE-AC36-99-GO10337 
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Structural Testing of 9 m Carbon Fiber Wind Turbine 

Research Blades*

Joshua Paquette†

Sandia National Laboratories‡, Albuquerque, NM, 87111 

Jeroen van Dam§

 and Scott Hughes§

National Renewable Energy Laboratory**, Golden, CO, 80401 

Three 9 m carbon fiber wind turbine blades have been designed through a research 

program initiated by Sandia National Laboratories. The individual designs feature such 

innovations as carbon spar caps, material-induced twist-bend coupling, and flatback airfoils, 

among others. All blades were constructed with conventional dry lay-up and VARTM 

infusion processes. Static tests of these blades were conducted at the National Wind 

Technology Center. The blades were subjected to flapwise loading to simulate the extreme 

wind loads expected for each design in a Class 2b wind site. The blades were loaded with a 

three-point whiffle-tree arrangement. Upon obtaining the predetermined test load, the 

blades were subsequently loaded to failure. Load, deflection, strain, and acoustic emissions 

were monitored throughout the experiments. All blades survived the specified test loads, 

with two designs exceeding it significantly. In addition, carbon strains of over 0.8% in both 

tension and compression were recorded in one of the tests. Finally, acoustic microphones 

were able to detect areas where damage was occurring, and indicated the beginnings of 

failure. This paper outlines the results of the structural tests that were conducted. 

Nomenclature 

VARTM = Vacuum Assisted Resin Transfer Method 

SNL = Sandia National Laboratories 

NWTC = National Wind Technology Center 

HP = High Pressure 

LP = Low Pressure 

I. Introduction 

N 2002, Sandia National Laboratories (SNL) initiated a research program1 to demonstrate the use of carbon fiber 

in subscale blades2

 and to investigate advanced structural concepts through the Blade System Design Study 

(BSDS)3, , 4 5

. From this effort, three 9 m designs were created by SNL with assistance from Global Energy Systems 

Consulting (GEC)††, Dynamic Design Engineering‡‡, and MDZ Consulting§§; and seven blades from each design 

were manufactured by TPI Composites***. All blades were designed for a 100 kW stall controlled turbine. The first 

blade set was called CX-100 (Carbon Experimental), and contained a full-length carbon spar cap, a relatively new 

concept at the time. The geometry of the CX-100 was based on the design of the ERS-1006

 blade at outboard span 

stations, and the Northern Power Systems NW-100 blade in the root area. The second blade design, the TX-100 

(Twist-Bend Experimental), had the same geometry as the CX-100, but featured a significantly different laminate 

I 

 

* This paper is declared work of the U.S. Government and is not subject to copyright protection in the United States.

† Wind Energy Technology Department, MS 1124 

‡ Sandia is a multiprogram laboratory operated by Sandia Corporation, a Lockheed Martin company, for the U.S. Department of Energy under 

contract DE-AC04-94AL85000 

§ National Wind Technology Center, MS3911 

** NREL is operated by Midwest Research Institute, Battelle, Contract No. DE-AC36-99-GO10337 

†† Global Energy Concepts, LLC, 5729 Lakeview Drive NE, Suite 100, Kirkland, WA 98033-7340, (425) 822-9008 

‡‡ Dynamic Design Engineering, Inc., 123 C St., Davis CA, (530) 753-7961 

§§ MDZ Consulting, 601 Clear Lake Road, Clear Lake Shores, Texas 77565, (281) 334-5681 

*** TPI Composites, Inc., 373 Market Street, P.O. Box 328, Warren, RI 02885, (425) 247-1050

1
design. The blade was designed to have passive aerodynamic load reduction by orienting unidirectional carbon 20° 

off of the pitch axis in the skins from approximately 3.50 m outward. Studies7, , , , , 8 9 10 11 12 had indicated the possibility 

of such a method of passive aerodynamic load alleviation through material induced twist-bend coupling, with 20° 

being determined as the optimum angle. Other methods of twist-bend coupling have also been suggested13, , , 14 15 16. 

The TX-100 also contained a fiberglass spar cap which terminated at the mid-span of the blade. The final blade 

design was named the BSDS, owing its name to the research program under which it was created. The BSDS had a 

length of 8.325 m rather than the 9.000 m length of the CX-100 and TX-100. This blade design exhibited a highly 

efficient structure which included such features as a thin, large-diameter root; flatback airfoils; integrated root studs; 

and a full-length, constant-thickness, carbon spar cap. A drawing of the plan form of these blades is shown in Figure 

1, with the carbon-containing areas shown in blue. Note the carbon spar caps of the CX-100 and the BSDS blades, 

and the carbon outboard skins of the TX-100. The unidirectional fiberglass spar cap of the TX-100 is shown in red 

in the figure and extends only to the mid-span of the blade. It was determined that the large amount of carbon 

contained in the skin was adequate to carry loads outboard in this design, making a full length spar cap unnecessary. 

The narrow, constant thickness spar cap of the BSDS blade alludes to the inherent structural stiffness of this design. 

In addition, the BSDS design features a gradual transition between the root and max-chord areas as compared to the 

CX and TX plan form. Figure 2 provides a view of the two different blade geometries as seen from the trailing edge. 

Note the flat back on the inboard portion of the BSDS blade, as well as the larger root and smoother chord thickness 

transition. 

Figure 1: 9 m blade plan forms and major material regions. 

Figure 2: 9 m blade geometries as viewed from the trailing edge. 

Figure 3 shows the airfoil geometries and relative sizes at the root, max-chord, and tip of the CX-100, TX-100, 

and BSDS blades. Note that in the figure the airfoils are shown without pre-twist. The BSDS blade can be seen to 

have a larger root diameter while also having a shorter max-chord length. The tip airfoil of the BSDS blade is 

representative of the outboard airfoils in the blade which are thin and aerodynamically high performing. 

2
A suite of laboratory and field tests were proposed to verify 

that the manufactured blades met their design goals. In the 

laboratory, the blades were to undergo modal, static, and 

fatigue testing. To date, specimens from each design have 

undergone modal and static testing. This paper covers the 

results of the static testing. 

II. Test Setup 

One blade from each design set was tested to failure at the 

National Wind Technology Center (NWTC) near Boulder, CO. 

In addition, a second TX-100 blade was tested to ascertain the 

effect that a manufacturing anomaly had on the blade. This 

anomaly was in the form of extra leading edge reinforcement 

material that was mistakenly added to the first four blades that 

were manufactured. This extra material was thought to have the 

potential to diminish the twist-bend coupling of the blades. TX100 #004 contained the extra leading edge reinforcement and 

was tested to failure while the design specification TX-100 #007 was loaded to a representative proof load of 68% of 

the extreme load. 

The blades were mounted to a 1360 kN-m test stand and subjected to a flapwise bending load case to 

approximate the extreme loading events for the wind class that each blade was designed to17. In the case of the CX100 blade, the extreme load case resulted in a root moment of 86.4 kN-m. For the TX-100 and BSDS blades, the 

extreme load case resulted in a root moment of 53.8 kN-m. All blades were loaded by use of a three-point whiffle 

tree and saddle arrangement connected to an overhead bridge crane which is shown schematically in Figure 4. 

 For the TX-100 blade test, a rope and pulley arrangement was attached to the saddles in an attempt to allow 

twisting without altering the loading point. When optimized correctly, the system can allow for loading through a 

stationary point on the chord line within a given rotation range. A schematic of the setup is shown in Figure 5. 

Figure 5: Rope and pulley 

arrangement for TX-100 

blade test.

A three-saddle whiffle-tree was also used for the BSDS blade test, but with different saddle stations. The saddle 

stations and applied loads at the 100% test load level are given for each blade in Table 1. 

The blades were loaded and unloaded in increasing 25% increments up to the 100% test load as shown 

schematically in Figure 6. At each load step, the load was held for approximately 60 s. The 100% test load was 

calculated by multiplying the design load 

distribution for each blade by 1.10 to 

account for test-to-test variability. The 

resulting distribution was then 

approximated with a piece-wise linear fit 

achieved by point loads applied at the three 

saddle locations. The desired and applied 

test moment distributions for the CX-100, 

TX-100 and BSDS blade tests are shown in 

Figure 7, Figure 8, and Figure 9 

Figure 4: Three point whiffle-tree used in 9 m static blade tests. 

Figure 3: 9 m airfoil geometries at root, 

max-chord, and tip. 

Table 1: Saddle positions and loads* for 9 m static tests. 

CX-100 TX-100 BSDS Saddle

# Position

(m) 

Load 

(kN) 

Position

(m) 

Load 

(kN) 

Position

(m) 

Load

(kN) 

1 3.00 16.91 3.00 9.30 3.00 9.79 

2 5.81 5.47 5.81 3.01 4.80 3.96 

3 7.26 5.59 7.26 3.07 6.60 3.65 

*Loads are for 100% test load

3
respectively. After reaching the 100% test load, loads were increased in 10% increments and held for 60 s until 

failure occurred (see Figure 6). 

 

 

0

20

40

60

80

100

120

140

02468

Blade Station (m)

Moment (kN-m)

Desired Test Load

Applied Test Load

Figure 7: CX-100 test loading 

distribution.

0

10

20

30

40

50

60

70

80

02468

Blade Station (m)

Moment (kN-m)

Desired Test Load

Applied Test Load

Figure 8: TX-100 test loading 

distribution. 

0

10

20

30

40

50

60

70

80

02468

Blade Station (m)

Moment (kN-m)

Desired Test Load

Applied Test Load

Figure 9: BSDS test loading distribution. 

Time

Load/Test Load (%)

50%

75%

100%

110%

120%

130%

Tare

Figure 6: Loading sequence for 9 m blade 

tests. 

An array of sensors was used in the tests to monitor strain, deflection, load, and acoustic emissions. Load was 

monitored with a 100 kN load cell mounted between the whiffle-tree and the overhead crane (see Figure 4). 

Deflection was monitored by string potentiometers (pots) located near the saddle positions. For the CX-100 and 

BSDS blades, a string pot was located near the trailing edge of the blade near each saddle location. For the TX-100 

test, additional string pots were mounted at the leading edge of the blade near the two inner most saddles to measure 

twist. Inclinometers were also mounted at 4.300 m and 6.250 m on the TX-100 blade to measure twist at those 

stations. The test blades were outfitted with a suite of 30-40 strain gages to measure strains along the blade 

centerline, in large panel regions near max-chord, and at other material and structural points of interest. Finally, the 

blades were instrumented with a grid of surface-mounted microphones to detect acoustic events that can point to 

regions of the blade incurring damage and thus, may indicate incipient failure regions. 

III. Test Results and Analysis 

All three blade designs successfully withstood the prescribed test loads. Figure 10 shows the maximum 

deflections that were measured in the three blade tests along with the corresponding root moments. The deflections 

are the result of a polynomial fit of the string pot results. The maximum tip deflections for the CX-100, TX-100, and 

BSDS blades were 1.05 m, 1.80 m, and 2.79 m respectively. Figure 11 compares the deflections at a root moment of 

100 kN-m. The results show that the CX-100 blade was significantly stiffer than the TX-100 and BSDS blades. 

4
0.0

0.5

1.0

1.5

2.0

2.5

3.0

02468

Blade Station (m)

Flap Deflection (m)

CX-100 at 128.6 kN-m

Root Moment

TX-100 at 121.4 kN-m

Root Moment

BSDS at 203.9 kN-m

Root Moment

Figure 10: Maximum deflections for 9 m 

blades. 

0.0

0.2

0.4

0.6

0.8

1.0

1.2

1.4

1.6

02468

Blade Station (m)

Flap Deflection (m)

CX-100

TX-100

BSDS

Deflections Measured at

100 kN-m Root Moment

Figure 11: 9m blade deflections at 100 

kN-m root moment.

The rotation angles about the spanwise axis at different stations of the TX-100 blade were measured during the 

test. The measurements were taken both from leading edge and trailing edge mounted string pots, and inclinometers 

mounted along the blade centerline. A positive angle corresponds to an increasing angle of attack for the airfoil at a 

given station. Figure 12 shows the twist angles that were measured at a 44.0 kN-m root moment in the test of blade 

#004 (blade with added material) and blade #007 (blade without added material) of the TX-100 blade set. Blade 

#004 was did not twist as much as blade #007 at the same load. This provides evidence of the effect that the extra 

leading material had on blade #004. Figure 13 shows the rotations measured along blade #004 at the 100% test load 

and at the failure load. The TX-100 blade showed a maximum twist of about -5° at the 6.25 m station for a root 

moment of 116.9 kN-m. 

 

-2.5

-2.0

-1.5

-1.0

-0.5

0.0

0.5

02468

Axial Position (m)

Rotation (°)

Blade #007

Blade #004

Rotations Measured at

44.0 kN-m Root Moment

Figure 12: TX-100 #004 and #007 twist 

angles at 44 kN-m root moment. 

-6.0

-5.0

-4.0

-3.0

-2.0

-1.0

0.0

1.0

02468

Axial Position (m)

67.7 kN-M

Root Moment

116.9 kN-m

Root Moment

TX-100 Blade #004

Rotation (°)

Figure 13: TX-100 #004 twist angle at test 

load and near failure.

Figure 14 shows the strains that were measured along the high-pressure and low-pressure spar caps for the three 

blades that were tested. Note that for the TX-100 blade, the spar cap terminates at the mid-span of the blade and thus 

strains outboard of that point are along the blade centerline. The carbon spar cap of the CX-100 blade experienced 

strains of around 3000 με before failure. The TX-100 blade exhibited maximum strains of approximately 5500 με

near the termination point of the spar cap and strains elsewhere in the blade were generally lower. The high strain at 

the spar cap termination was likely the result of a stress concentration emanating from the tapered end of the spar 

cap (see Figure 1). The carbon spar cap of the BSDS blade experienced maximum strains of over 8000 με on both 

the high-pressure and low-pressure surfaces. 

5
Figure 15, Figure 16, and Figure 17 show the strains 

measured in the panel regions aft of the spar cap for the 

CX-100, TX-100, and BSDS blades respectively. These 

are regions that are susceptible for buckling in extreme 

loading situations. The CX-100 and TX-100 blades 

showed indications of buckling in the 1.800 m region 

near a root moment of about 117 kN-m. These blades 

were of identical geometry and had similar construction 

in this area. Thus, this result is not unexpected. The 

BSDS blade showed slight signs of buckling at the 

1.575 m station at around a 150 kN-m root moment. 

However, it is important to note that the blade 

continued to withstand load well beyond this point. 

 

0

20

40

60

80

100

120

-2000 -1000 0 1000

0%

50%

100%

150%

Strain (με)

Root Moment (kN-m)1.800 m, 0°

% Test Load

1.800 m, 90°

TX-100 #004

Figure 16: TX-100 aft panel strains parallel 

and perpendicular to blade axis. 

-8000

-6000

-4000

-2000

0

2000

4000

6000

8000

02468

Blade Station (m)

BSDS HP

TX-100 HP

CX-100 HP

BSDS LP

TX-100 LP

CX-100 LP

Strain (με)

Figure 14: 9 m spar cap strains near failure. 

0

40

80

120

160

200

-8000 -6000 -4000 -2000 0 2000 4000 6000

0%

50%

100%

150%

200%

250%

Strain (με)

Root Moment (kN-m)

1.050 m, 0°

% Test Load

1.575 m, 0°

1.050 m, 90°

1.575 m, 90°

BSDS

Figure 17: BSDS aft panel strains parallel and 

perpendicular to blade axis. 

0

20

40

60

80

100

120

-1500 -1000 -500 0 500 1000

0%

20%

40%

60%

80%

100%

Strain (με)

Root Moment (kN-m)

1.800 m, 0°

% Test Load

2.550 m, 0°

1.800 m, 90°

2.550 m, 90°

CX-100

Figure 15: CX-100 aft panel strains parallel 

and perpendicular to blade axis. 

For the BSDS test, strain gages were placed on the 

flat back, perpendicular to the blade axis at 0.580 m and 

at 1.575 m. The 0.580 m station contained one gage in 

the middle of the flatback (labeled 50%). At the 1.575-

m station, gages were placed on the high pressure 

surface near the edge of the flat back (labeled 0%), on 

the low pressure surface near the flat back (labeled 

100%), and at two intermediate positions, 25% and 

75% of the distance between the high-pressure and lowpressure surfaces. The gages were intended to indicate 

if the flat back was collapsing or rotating relative to the 

high-pressure and low-pressure surfaces during the test. 

Figure 19 shows the strains measured at the 

aforementioned locations. Strain was linear with respect 

to load at the 0.580 m location, suggesting that the flat 

back was not buckling inward or outward at this 

location. All gages at the 1.575 m location show 

evidence of changes in the load path, especially beyond 

6
a root moment of 150 kN-m. The 0% and 25% gages 

start to change from compression to tension while the 

75% and 100% gages exhibit an inflection in strain 

direction. This may suggest that the flat back was 

beginning to collapse by the low-pressure edge in a 

concave direction. The strains changing in opposite 

directions at the opposite ends of the flat back suggest 

that the flat back was not experiencing concave or 

convex panel buckling at the 1.575 m location. 

0

40

80

120

160

200

-1500 -1000 -500 0 500 1000 1500

0%

50%

100%

150%

200%

250%

Strain (με)

Root Moment (kN-m)

0.580 m, 50%

% Test Load

1.575 m, 100%

1.575 m, 25%

BSDS

1.575 m, 75%

1.575 m, 0%

 Also of interest for the BSDS blade was the 

behavior of the root mounting studs. A novel root 

mounting stud design was implemented into the BSDS 

blade. The installation of the root mounting fasteners 

was integrated into the laminate infusion. Gages were 

placed near the tip of the most highly loaded root studs 

on the upper (low pressure) and lower (high pressure) 

surfaces. Gages were also placed at the same spanwise 

stations but between the root studs in the 

circumferential direction on the blade surface. Figure 

18 shows the strain at these locations with respect to 

applied load. The resulting strains are relatively similar 

in magnitude and linear throughout the loading, 

indicating that the implementation of the root studs did 

not cause a significant stress concentration, with a 

relatively smooth load path between laminate and root 

stud. The root area was not observed to exhibit any 

signs of damage throughout testing. 

The acoustic emissions sensed by microphones 

placed on the blade surface can be used to indicate the 

location of damage that the blade is incurring. This is 

accomplished by setting a time window in which 

individual microphones can detect the same event. The 

difference in time at which the event is detected by each 

microphone is used to triangulate the position on the 

surface that the microphones are mounted, assuming the velocity profile of the substrate is known. Unfortunately, 

the velocity profile is complicated in a composite structure which contains materials with widely varying acoustic 

transmission properties. The end result is a structure which has a velocity profile which is spatially and directionally 

dependent. For the results from these tests, various fits of the acoustic velocity field were used and were of the form 

( )

n

+= bav α cos (1) 

Figure 19: BSDS flat back strains. 

0

40

80

120

160

200

-3000 -2000 -1000 0 1000 2000 3000

0%

50%

100%

150%

200%

250%

Strain (με)

Root Moment (kN-m)

% Test Load

HP Between Studs

BSDS

HP Stud Tip

LP Between Studs

LP Stud Tip

Figure 18: BSDS threaded rod strains. 

-0.8

-0.6

-0.4

-0.2

0.0

0.2

0.4

0 0.5 1 1.5 2 2.5 3

x (m)

y (m)

CX-100

Spar Cap

*Note: Energy defined as area under V-t curve.

Energy*

<100

100-1000

>1000

where a, b, and n are fitting constants and α

is the angle between the direction vector 

from the sensor to the event and the blade 

pitch axis. The resulting located events for 

the CX-100, TX-100, and BSDS blade tests 

are shown in Figure 20, Figure 21, and 

Figure 22 respectively. Each event is color 

coded with a respective energy range. It 

should be noted that this energy is not the 

true energy, as it represents the integral of 

the voltage vs. time curve. The true energy 

is defined by the integral of the voltage 

squared vs. time curve. These voltage vs. 

time results are useful for comparative 

Figure 20: Acoustic event locations and energies for CX-100 

blade test. 

7
purposes however. The events are overlaid 

on outlines of each blade. Since the 

locations of the acoustic events are twodimensional, the outlines are of the 

flattened low-pressure skins of the blades. 

In addition to the outlines of the blades, 

important regions such as spar caps and 

material boundaries are shown. 

During the CX-100 test, several highenergy events were located between 1.20 m 

and 1.30 m along the spar cap. The CX-100 

blade was observed during failure to 

experience a catastrophic buckling of the 

low-pressure skin near the 1.2-m station. A 

post mortem inspection of the blade in this region showed a large crack in the bond joint between the low pressure 

skin and the shear web. A photograph of the crack is shown in Figure 23. 

-0.8

-0.6

-0.4

-0.2

0.0

0.2

0.4

0 0.5 1 1.5 2 2.5 3

x (m)

Energy*

<100

100-1000

>1000

y (m)

TX-100

Spar Cap

*Note: Energy defined as area under V-t curve.

Glass-Carbon Transition

Figure 21: Acoustic event locations and energies for TX-100 

blade test.

 

Figure 23: Crack in shear web to LP 

skin bond joint at 1.20-1.30 m on the 

CX-100 blade. 

The TX-100 blade test produced high-energy events just beyond the 1.00 m station. The blade eventually failed 

at the 1.0-m station, with a failure described as a buckling of the compressive skin, initiated towards the leading 

edge. A photograph of the fracture is shown in Figure 24. 

During the BSDS test, the only high-energy events were located around 0.350 m, which the station where the 

shear web terminates on the root end. A closer examination of this area after the test showed that a large crack had 

developed between the low-pressure skin and the shear web in the bonding joint. A photograph of the crack is 

shown in Figure 25. 

Figure 24: Crack near blade centerline on 

TX-100 blade. 

 

Figure 25: Crack at shear web 

termination on the LP surface of the 

BSDS blade.

-1.0

-0.8

-0.6

-0.4

-0.2

0.0

0.2

0.4

0 0.5 1 1.5 2 2.5 3

x (m)

y (m)

BSDS

Spar Cap

*Note: Energy defined as area under V-t curve.

Energy*

<100

100-1000

>1000

Flat Back

Figure 22: Acoustic event locations and energies for BSDS 

blade test. 

8
During the initial static test of the BSDS blade, the blade catastrophically failed near the 5.00 m station, with a 

root moment of 203.9 kN-m. The complete catastrophic failure at this station prohibited the assessment of the mode 

of failure. This is of interest as ostensibly panel buckling in the max airfoil stations of the blade was not the limiting 

factor in the static strength of the blade, as observed for other blades of this size. A second static strength test was 

performed on the BSDS in an attempt to obtain a failure in the inboard stations of the blade. A two-point whiffle tree 

was implemented to apply this load. During this second test, the blade failed near the 2.00 m station at a root 

bending moment of 220.2 kN-m. 

In addition to indicating the location of damage occurring within a blade, acoustic emission monitoring is also a 

valuable tool for assessing when failure is occurring locally or globally. Figure 26, Figure 27, Figure 28, show the 

sum of the acoustic emission energy detected by all of the sensors in the root region of the CX-100, TX-100, and 

BSDS blades respectively. In each plot a trend line is drawn through the initial linear portions of the response. At 

some load, the response begins to become non-linear, indicating an acceleration of the acoustic emissions and the 

greater potential of failure. For the CX-100 and TX-100 blades, the acoustic emissions accelerate rapidly beyond 

root moments of 125 kN-m and 115 kN-m respectively. The response from the BSDS blade remained mostly linear 

until a root moment of 140 kN-m was reached. The accumulation of acoustic energy in the BSDS blade did grow 

non-linearly after 140 kN-m, although not as rapidly as for the other two blades. 

 

0

10

20

30

40

50

60

70

80 90 100 110 120

Root Moment (kN-m)

Energy (kV-s)

TX-100

Figure 27: TX-100 acoustic energy 

accumulation. 

0

5

10

15

20

90 100 110 120 130

Root Moment (kN-m)

Energy (kV-s)

CX-100

Figure 26: CX-100 acoustic energy 

accumulation. 

0

10

20

30

40

50

60 80 100 120 140 160 180 200

Root Moment (kN-m)

Energy (kV-s)

BSDS

Figure 28: BSDS acoustic energy 

accumulation. 

A summary of salient results of the static tests of all three 

blade designs is contained in Table 2. The relatively high strength 

of the BSDS blade is evident, with the measured carbon strains 

approaching values seen in coupon testing of pre-preg specimens. 

This high strength combined with the significantly lower weight 

points to the structural advantages of this design in comparison to 

the CX-100 and TX-100. 

9
IV. Conclusion 

Table 2: Summary of results of 9 m blade tests. 

Property CX-100 TX-100 BSDS 

Weight (lb) 383 361 289 

% of Design Load at Failure 115% 197% 310% 

Root Failure Moment (kN-m) 128.6 121.4 203.9 

Max. Carbon Tensile Strain at Failure (%) 0.31% 0.59% 0.81% 

Max. Carbon Compressive Strain at Failure (%) 0.30% 0.73% 0.87% 

Maximum Tip Displacement (m) 1.05 1.80 2.79 

Specimens from three 9 m carbon fiber blade designs underwent static structural testing. All three test blades 

survived factored design test loading. The CX-100 blade displayed exceptional stiffness, deflecting only 1.05 m at a 

root moment of 128.6 kN-m. The blade failed due to panel buckling near max-chord which was likely initiated by a 

separation between the shear web and the low-pressure skin in that region. The TX-100 blade successfully 

demonstrated twist-bend coupling caused by 20° off-axis carbon in the outboard skins. The TX-100 blade failed at a 

slightly lower load than the CX-100 blade, but in a similar location. The BSDS blade displayed exceptional strength 

in comparison to the CX-100 and TX-100 designs, surviving to almost three times the target test load. The flat back 

airfoil feature performed well and did not display non-linear behavior until well after the target test load was 

reached. In addition, the root mounting studs of the BSDS blade were observed to have good static loading 

properties. Finally, the acoustic emission monitoring system detected not only the locations where damage was 

occurring, but also incipient global blade failure. 

Acknowledgments 

The authors would like to acknowledge Mark Rumsey, Perry Jones, and Wesley Johnson of SNL; Mike Jenks, 

Laura Rip, Dave Simms, and Walt Musial of NREL; Derek Berry of TPI Composites, Inc., Alan Beattie of PAC; 

Dayton Griffin of GEC; Kevin Jackson of Dynamic Design Engineering; and Mike Zuteck of MDZ Consulting for 

their assistance in the preparation and performance of these tests. 

References 

1 Ashwill, T, and Laird, D, “Concepts to Facilitate Very Large Blades,” Proceedings, ASME/AIAA Wind Energy 

Symposium, Reno, NV, 2007. 

2 C.-H. Ong and S. W. Tsai, “The Use of Carbon Fibers in Wind Turbine Blade Design: A SERI-8 Blade Example,” 

SAND2000-0478, Sandia National Laboratories Contractor Report, March 2000. 

3 Berry, Derek and Lockard, S., “Innovative Design Approaches for Large Wind Turbine Blades,” SAND2003-0723, Sandia 

National Laboratories, Albuquerque, NM, March 2003 

4 Berry, Derek and Lockard, S., “Innovative Design Approaches for Large Wind Turbine Blades Final Report,” SAND2004-

0074,Sandia National Laboratories, Albuquerque, NM, May 2004 

5 Griffin, Dayton A., “Blade System Design Studies Volume II: Preliminary Blade Designs and Recommended Test Matrix,” 

SAND2004-0073, Sandia National Laboratories, Albuquerque, NM, June 2004 

6 TPI Composites, Inc. Final Project Report, “Blade Manufacturing Improvements Development of the ERS-100 Blade,” 

SAND2001-1381, May 2001. 

7 P. S. Veers, G. Bir and D. W. Lobitz, "Aeroelastic Tailoring in Wind-Turbine Blade Applications," Proceedings, 

Windpower '98 Meeting, pp. 291-304. 

8 Lobitz, D. and Veers, P., “Aeroelastic Behavior of Twist-coupled HAWT Blades,” ASME/AIAA Wind Energy 

Symposium, Reno, NV, 1998, pp. 75-83. 

9 Lobitz, D. and Laino, D., “Load Mitigation with Twist-coupled HAWT Blades,” ASME/AIAA Wind Energy Symposium, 

Reno, NV, 1999, pp. 124-134. 

10 D. Lobitz, P. S. Veers and D. J. Laino, “Performance of Twist-Coupled Blades on Variable Speed Rotors,” Proceedings, 

ASME/AIAA Wind Energy Symposium, Reno, NV, 2000, pp. 404-412. 

11 D. W. Lobitz et al., “The Use of Twist-Coupled Blades to Enhance the Performance of Horizontal Axis Wind Turbines,” 

SAND2001-1303, May 2001. 

12 D. Griffin, “Evaluation of Design Concepts for Adaptive Wind Turbine Blades,” SAND2002-2424, Sandia National 

Laboratories Contractor Report, August 2002. 

10
13 C-H Ong and S. W. Tsai, "Design, Manufacture and Testing of a Bend-Twist D-spar," Sandia National Laboratories 

Contractor Report, SAND99-1324, June 1999. 

14 Wetzel, K., and Locke, J., “Uncoupled and Twist-bend Coupled Carbon-glass Blades for the LIST Turbine,” Proceedings, 

ASME/AIAA Wind Energy Symposium, Reno, NV, 2004, pp. 13-23. 

15 Locke, James and Valencia, J., “Design Studies for Twist-Coupled Wind Turbine Blades,” SAND2004-0522, Sandia 

National Laboratories, Albuquerque, NM, June 2004 

16 Wetzel, K., “Utility Scale Twist-Flap Coupled Blade Design,” Proceedings, ASME/AIAA Wind Energy Symposium, 

Reno, NV, 2005, pp. 382-394. 

17 D. Berry, “Design of 9-meter Carbon-Fiberglass Prototype Blades Final Project Report,” SAND Report in preparation, 

Sandia National Laboratories, Albuquerque, NM, 2007 

11
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Structural Testing of 9 m Carbon Fiber Wind Turbine Research 

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This paper outlines the results of tests conducted on three 9-m carbon fiber wind turbine blades designed through a 

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