This paper gives an overview of current wind turbine rotor blade composite materials research carried out in the material’s laboratory of WMC within the framework of European projects. It outlines the collection and analysis of reference data, research into the effect of temperature and frequency on fatigue life, and experiments to describe the microscopic and subcomponent behaviour.
Nomenclature αT, αC = parameters in regression model, describing curvature of constant life lines a = regression parameter of S-N curve C, D, E = parameters of regression model, describing the constant life diagram boundaries F = maximum load h = web height N = cycles to failure w = width of bondline R = ratio of cyclic minimum and cyclic maximum S = generic term for cyclic stress σa, σm = stress amplitude/mean
I. Introduction Wind turbine rotor blade materials and structures increasingly require optimal and reliable design. In recent
years, the knowledge on these materials has expanded, allowing more accurate strength and life calculations. The experimental and theoretical work is continuing in various research projects, both in the US and in Europe.
The main objective of this paper is to report and analyse recent materials research results on wind turbine
laminates from current European research projects. The work described in this paper is part of the UPWIND and INNWIND projects [1][2]. In these projects, several
disciplines are combined to develop wind turbine technology for the current and next generations of wind turbines. In terms of materials research, the parts of the UPWIND and INNWIND projects described in this paper can be seen, in technical content, as successors of the OPTIMAT programme, which finished in 2006 [3].
Furthermore, this paper focusses on work carried out in the laboratory of WMC, one of the participants in the abovementioned projects.
II. Strength and fatigue of the reference laminate In order to obtain material characteristics that are suitable for comparison under various loading conditions, use
was made of a reference specimen. This is a test specimen with a fixed planform, made of a reference laminate, which is chosen at the start of the project to reflect current material and manufacturing method.
As is shown below, static mechanical characteristics can depend considerably on the specimen geometry and
load introduction – there is a best specimen and test configuration for every mechanical characteristic. When using different specimens and test methods or e.g. tension and compression tests, this will result in difficulties when
48th AIAA Aerospace Sciences Meeting Including the New Horizons Forum and Aerospace Exposition4 - 7 January 2010, Orlando, Florida
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comparing e.g. the influence of environmental conditions on tensile strength with the influence on compression strength. Moreover, in fatigue, the standardisation of test specimens and methods is relatively limited.
Because the current research is aimed at isolating influences rather than finding the highest measurable values for the mechanical properties, a reference specimen is used quite extensively. Incidentally, for the particular laminate, the performance of this specimen in fatigue is satisfactory in most loading types.
This approach, however, does entail the need for testing the same laminate using the relevant standards. Below, this is documented in comparisons of test methods.
The reference specimen is a specimen with planform ‘R08’. This specimen’s nominal dimensions are l x w x t = 130 x 20 x 3 mm, with 55 mm long, 1 mm thick tabs, 20 mm gauge length. The reference laminate is a 4-layer UD glass/epoxy laminate. In previous publications, manufacturing method, static and fatigue strengths of the reference material, and of constituents, under various loading conditions were described (e.g. [4], [5]).
Mechanical characteristics of the laminate are tabulated in Table 1. Table 2 also includes mechanical characteristics of bonding paste.
Figure 1: Reference specimen (planform R08)
Table 1: Static characteristics of reference material, bonding paste
Nominal dimensions: R08=length x gaugelength x width x thickness=155 x 20 x 20 x 3
I03=136 x 10 x 10 x 3; I04=Iosipescu=76x12x3 **strain, stiffness, poisson ratio average from sub-set (not all specimens were equipped with strain gauges) Average FVF and FWF were 48.3% and 67.8%, respectively
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Table 2: Static characteristics of resin and bonding paste
D01 = dogbone specimen, length x gaugelength x width x thickness x radius =100 R08=length x gaugelength x width x thickness=155 x 20 x 20 x 3; I03=136 x 10 x 10 x 3; I04=Iosipescu=76x12x3; S81=double lap shear; S72=tube *bondline thickness < 4 mm only
A. Compression testing In Table 1, the number of tested specimens in compression using the combined loading compression (CLC)
fixture (ASTM D6441) is relatively high. This is part of an ongoing test programme where the influence of small variations in test set-up is investigated. The compression characteristics of a composite are quite relevant for design, as compression strength is lower than tensile strength. Compressive strength test results, in turn, are sensitive to load introduction, and it is well-known, that dedicated test methods such as ASTM 6441 [6] give relatively high results in static tests. This method uses a combined loading fixture, introducing the in-plane compression load relatively benignly, both via shear and end-loading.
The combined loading fixture is shown in Figure 2, together with fatigue data obtained in compression-
compression fatigue using either the reference specimen as the standardised compression specimen and fixture (although it should be added, that the combined loading fixture is not originally intended for use in fatigue tests).
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As far as the extensive scatter –which is very common in compression fatigue data– allows, linear regression of the data shows a tendency for the combined loading fixture to outperform the reference specimen, especially at higher loads. Regardless of whether this were to be verified by more extensive data, this figure illustrates how the conditions under which test data were obtained and (failure to) treat them statistically might influence design, as the line for the CLC-fixture gives far more conservative life estimates at design strain ranges (usually far below the test strain ranges).
B. Shear strength of bonding paste A comparison of shear strength methods of the reference laminate was reported in [5]. A method for
determination of shear strength of adhesives [4] was implemented at the laboratory of WMC, and compared to more conventional test methods, such as lap shear[7] and Iosipescu [8]. Tubular specimens were manufactured using essentially the same method as used for the reference material. Four layers of dry fabric were wrapped around a cylindrical core. The core was inserted into a mould, and then this double-sided mould was used to infuse the specimens using vacuum. After successful infusion, post-curing, and demoulding, the tubes were cut in half transversely and bonded together with controlled bondline thicknesses between 0.5 and 4 mm.
This method shows higher shear strengths when compared to lap-shear tests or Iosipescu tests on bonding paste
specimens. Furthermore, up to 3 mm, the bondline shear strength seems not to be affected,see Figure 4. Beyond that, further testing is required to evaluate any thickness-dependent behaviour.
Figure 3: Test method for shear strength of adhesives
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0
10
20
30
40
50
60
0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5 4.0 4.5
bondline thickness [mm]
Shea
r str
engt
h [M
Pa]
Figure 4: Test results for shear strength of adhesives
1. Fatigue life
Extensive fatigue data form the basis for the derivation and validation of integral fatigue models. The ultimate goal in fatigue research is to generate models that enable optimal design, with as limited experimental work required on beforehand as possible.
0
100
200
300
400
500
600
700
-800 -600 -400 -200 0 200 400 600 800 1000
Mean cyclic stress [MPa]
Am
plitu
de o
f cyc
lic s
tres
s [M
Pa] 100
100010000100000100000010000000
Figure 5: Fatigue behaviour of the reference specimen in piecewise linear constant life diagram
For the reference laminate, testing was done to determine the constant life diagram (CLD). Figure 5 shows this
diagram, which essentially gives a summary of the fatigue behaviour for 9 R-values, where the R-value is the ratio of minimum to maximum cyclic stress. Lines of constant fatigue life are plotted, that connect, for selected nominal lifetimes, the corresponding mean-amplitude combinations for each R-value. These constant life lines are taken from
R=-0.4
R=0.1
R=0.5
R=0.9
R=-1
R=-2.5
R=10
R=2
R=1.1
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interpolation or extrapolation of the available S-N data. The outer lines correspond to relatively shorter lifetimes, the inner lines to longer lives. It should be added, that the outermost line (100 cycles to failure) should probably be deleted from the figure (but was left in for illustrative purposes). This line in particular is a result of extrapolation to lifetimes below which tests were actually carried out for most S-N curves. As a result, for e.g. R=0.9, this line exceeds static strength considerably and in general does not give a very accurate description of fatigue behaviour.
To date, fatigue tests were done at R-values of 0.9, 0.5, 0.1, -0.4, -1, -2.5, and 10, 2, and 1.1.
2. Constant life diagram and statistical fatigue modelling Some statistical analysis of the data from these projects was discussed in [9]. In the below paragraphs, the results
of a regression analysis on the fatigue data is discussed briefly. The CLD of Figure 5 was constructed by plotting projections of the S-N curve at constant R-values onto the
cyclic mean-amplitude plane and connecting points of equal fatigue life with a straight line; a piecewise linear constant life diagram.
Figure 6: Linear Goodman diagram
The simplest form of a piecewise linear constant life diagram is the Linear Goodman diagram, as illustrated in
Figure 6. The figure shows the constant life line for a single load cycle (N=cycles to failure=1). Cyclic stress amplitude is plotted on the ordinate, mean stress is on the abscissa. Hence, the ordinate coincides with the R=-1 (zero mean stress fatigue). This formulation of the CLD is created from R=-1 fatigue data (zero mean stress), and tensile and compressive ultimate strengths (parametrised here as σm/D and σm /E, respectively.
σa
Sa=C
Normalized σm D
mσE
mσ 0
R = -1
N = 1
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Figure 7: Modified Goodman diagram
An alternative formulation is shown in Figure 7. Here, the straight lines between the R=-1 data and the tensile
and compressive strengths are replaced with curved lines. The curvature is determined by αT and αC:
For σm ≥ 0; ⎟⎟
⎠
⎞
⎜⎜
⎝
⎛⎟⎠
⎞⎜⎝
⎛−=
T
DS m
aa
ασσ 1
For σm ≤ 0; ⎟⎟
⎠
⎞
⎜⎜
⎝
⎛⎟⎠
⎞⎜⎝
⎛−=
C
ES m
aa
ασσ 1
Both of the above formulations can be used to reverse engineer the S-N curves at any R-value, which results in
the following sets of equations: For the linear Goodman diagram: for σm ≥ 0 or -1 ≤R< 1:
( ) ⎟⎠⎞
⎜⎝⎛ −−= ma D
CCaaN σσ log.log.)log(
for σm ≤ 0 or R ≤ -1 and R > 1:
( ) ⎟⎠⎞
⎜⎝⎛ −−=
ECCaaN ma σσ log.log.)log(
For the modified Goodman diagram:
σa
Sa=C
Normalized σm D
mσE
mσ 0
R = -1
N = 1
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for σm ≥ 0 or -1 ≤R< 1:
( )C
EaCaaN m
a
ασσ ⎟
⎠
⎞⎜⎝
⎛−−−= 1loglog.log.)log(
for σm ≤ 0 or R ≤ -1 and R > 1:
( )C
EaCaaN m
a
ασσ ⎟
⎠
⎞⎜⎝
⎛−−−= 1loglog.log.)log(
By assuming values for the parameters a, C-E and α, a set of S-N curves can be obtained from the constant life diagram. This set can be compared to the actual measured S-N curves, and the differences can be expressed in a custom criterion. For example, the more these two sets resemble each other, the lower is e.g. the sum of squares of difference in log-life over the entire S-N curves, or the p-value of the paired t-test is higher than the significance level (in this case 0.05).
A next step can be, to select a sub-set of the lifetime parameters and find the optimum values. An example of a summary of such an exercise is shown in Table 3.
Here, for the linear and modified Goodman diagrams, a selection of models was run with less or more free
parameters. For example, in model 1, parameters C-E were fixed (common formulation of the linear Goodman diagram), no optimization was performed. This results in a relatively poor fit of derived and measured S-N curves.
On the contrary, for model 3, where all parameters were freed, the constant life diagram results in derived S-N
curves that fit the measured S-N curves better, from the viewpoint of total sum of squares of log life differences. Also, adding α parameters leads to better overall fits of measured and derived S-N curves.
Other models will be examined and this facilitates the selection or derivation of an optimal fatigue model, which
combines a low number of parameters and required experiments, with a reliable description of the fatigue behaviour of a laminate.
-400 -200 0 200 400 600
020
040
060
080
0
Mean Stress MPa
Stre
ss A
mpl
itude
MP
a
10 2̂10 3̂10 4̂10 5̂10 6̂
Figure 8:CLD modelling, orange = piecewise continuous CLD; black = model according to equations above,
blue dots = fatigue data
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Table 3: Estimated parameters, sum of squares, and p-value of Linear Goodman Diagram and Modified Constant Life Diagram
C. Strength and fatigue at off-reference conditions Based on the reference material fatigue behaviour, various other research topics have been identified that will
contribute to an integral model of material fatigue behaviour, including influence of: • complex stress state and thickness • environmental conditions • strength and fatigue life of constituent fibres and resins • loading rate and cyclic frequency • material creep
Tests on thick laminates under reversed loading conditions were performed in the 1 MN test frame of WMC see
Figure 9. The dimensions of the large specimens were 5 times the dimensions of the small geometry in all directions. These confirmed the suspicion of thickness effects that was also indicated in the tension tests [10]. Furthermore, the low frequencies at which these tests were run, a few of which with forced cooling to enable an increase of testing frequency, emphasised the requirement for further investigation of temperature-frequency effect.
Figure 9: Fatigue performance of thick laminates compared to thin laminates
The results of different investigations considering the effect of frequency on the fatigue (tension-tension and
tension-compression) behaviour were summarized by [11] as follows: “at low frequency ranges where there is negligible heat dissipation, as the load frequency increases, cycles to failure increase also. As higher frequency ranges are considered this increase is at a slower rate. When there is excessive heat dissipation however, a reverse trend can be observed”.
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To investigate the relevance of these observations for wind turbine laminates, and to quantify the effects of temperature and frequency, a dedicated test programme was defined and is being implemented. The results of this research will be useful for design, but also for defining test standards for fatigue of wind turbine laminates, where the choice of test frequency in fatigue can be highly relevant for the reliability of the results.
In a climate facility, fatigue tests in tension-tension were performed on reference specimens at laboratory, high,
and low temperatures, see Figure 10. In addition, tests at high frequencies and at a combination of high frequencies and low temperatures were performed (a selection of the data is presented in Figure 11 - Figure 13).
Figure 10: Comparison of reference laminate fatigue behaviour at reference, hot, and cold conditions
In Figure 10, nominally identical reference specimens were tested at room temperature, -40ºC, and 60ºC. With
the available data and the scatter in lifetime of the results, no statistically significant difference is observed for the low temperature fatigue data. On the other hand, testing the specimens at a temperature of 60ºC, results in approximately a factor 10 of reduction in life. The specimens were predominantly tested at load levels of 32 and 24 kN, so that the majority of the data points is near two distinct stress levels.
The frequency for the high load level was 2 Hz, whereas the low load levels were tested at 6 Hz. Lowering the
frequency for higher load levels is done to limit the surface temperature rise while also limiting the test time for long-life tests.
Another set of tests was done at ambient conditions, but at a higher frequency for part of the set of tests. In
Figure 11, the blue markers indicate tests done at 24 Hz, i.e. with a 12-fold frequency increase at the high load level, and a 4-fold increase at the low load level. The resulting S-N curve has shifted slightly less than a decade to the left, indicating considerable influence of test frequency. It could be considered slightly alarming, that the associated temperature increase is limited to less than 10ºC (which is generally considered as an acceptable temperature increase), see Figure 12. In addition, Figure 12 indicates, that the highest temperature increase is associated with the shortest fatigue life. It should be noted that the temperature is measured by a contact sensor on the surface of the specimen and that the surface is cooled by a fan. Therefore the internal temperature of the specimen may be significantly higher.
When these low versus high frequency tests are repeated in a cooled space at -40ºC, the influence of frequency
on life seems to be eliminated (Figure 13), although the measured surface heating of the specimen is very similar, see Figure 14. This does suggest that the inferior fatigue lives at high frequencies were caused by overheating of the specimen itself, and that there might be a threshold temperature above which the fatigue life is negatively influenced.
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Testing at high temperatures has a comparable effect on life as increasing the test frequency. 800700
600
500
400
300
200108107106105104103102
N
S
S=σmax [MPa] (R = 0.1)AmbientAmbient, high frequencyregr. Ambient (log N=-10.40·log S+32.2)regr. Ambient, high frequency (log N=-10.21·log S+31.0)
Figure 11: Comparison of reference laminate fatigue behaviour at reference, and high frequency
Figure 12: dT during RT tests at 32 kN and an elevated frequency of 24 Hz
600
500
400
300
200108107106105104103102
N
S
S=σmax [MPa] (R = 0.1)-40ºC-40ºC, high frequencyregr. -40ºC (log N=-8.53·log S+27.1)regr. -40ºC, high frequency (log N=-10.60·log S+32.5)
Figure 13: Comparison of reference laminate fatigue behaviour at reference, and high frequency in -40º C
environment
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Figure 14: dT during T = -40°C tests at 32 kN and an elevated frequency of 24 Hz
I. From micromechanics to subcomponents The scope of materials research is expanding to both smaller and larger size-scales compared to previous
coupon-based research. In order to better understand the failure mechanisms of wind turbine materials, micromechanical modelling is required. On the other hand, the connection between material characteristics and behaviour of constructions does not always seem to be obvious, requiring subcomponent research which is aimed at testing and modelling of relatively simple representative constructions.
D. Micromechanical One of the potential benefits of micromechanical modelling, is that coupon-scale experiments can be avoided
and/or better predicted when modelling the actual crack growth and interaction in a laminate. Although a large number of models has already become available, many of these are limited to single-crack growth modes such as interply delamination, static load cases, and relatively ‘neat’ -aerospace- layer structures. There is a need to expand both the theoretical modelling and the experimental database for fatigue of these laminates.
Extensive theoretical modelling is being done, where single fibres or groups of single fibres, and their interaction
with matrix material when loaded are modelled. Next to analytical modelling of crack growth (e.g. [12]), modelling of a ‘representative volume element’ is investigated, as it has the potential of being included in finite element models. Examples of this approach can be found in e.g. [13].
As an input to these models, basic data on the constituents of wind turbine composites need to be collected.
Some fatigue results are summarised in Figure 15. This figure contains tension fatigue data on the reference coupon, but also on pure resin dogbone shaped specimens (for representative test set-up see Figure 16).
Furthermore, data are shown for tensile fatigue of strands (strand data in the figure are taken from the MSU
database [14], [15]) and of single fibres. Single fibre tests were performed at WMC, using a dedicated test set-up with a small 1 N maximum capacity load cell that was mounted in a 25 kN maximum capacity test frame. Currently, the relation between single fibre static strength, fatigue life, and the comparison with the readily available strand data are under investigation, e.g. [16].
Figure 15 illustrates, that the fatigue performance of a composite can be thought of as ‘somewhere in between
fatigue behaviour of fibres and resin’, but how the performance depends on microstructure, fibre-matrix interface, etc. is not thoroughly understood for most composites or wind turbine laminates yet.
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Figure 15: Comparison of fatigue performance of resin, single fibres, impregnated strands, and reference
glass/epoxy composite
Figure 16: Epoxy resin dogbone specimen
E. Subcomponents On the other side of the scale-spectrum, work is done on subcomponents. Testing subcomponents as a material
test was described for wind turbine applications in [6]-[15]. Further use and potential of the subcomponents concept was explained e.g. in [4].
Within the UPWIND project, the first interesting issue in subcomponent testing and analysis that was tackled is
bondline integrity in load-carrying spar structures. In both beams, the shear web and flanges were bonded together with bonding paste adhesive. Two different cross-sectional designs were tested in three-point bending, see Figure 17, with and without reinforcements to mitigate load introduction damage and obtain bondline failure.
As Figure 18 illustrates, bondline failure was not the only failure mode observed in a static test. Also web-
buckling and delamination of the web sandwich were observed.
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The cross-section types differed in the configuration of the shear web, see Figure 19. In beams of type S79, the shear web had an asymmetrical cross-section. Beams of type S78 had an almost symmetrical shear web cross section (bi-axial fabric was wrapped around a prismatic foam core – hence some overlap was present on one side of the web). As a result of the different shear web configurations, the width of the bondline between web and flanges was larger for beams of type S79, resulting in larger failure loads in static tests, and longer fatigue lives at the same load. See Figure 20, where a rather limited number of fatigue tests illustrates this. The parameter plotted on the vertical axis is the maximum (absolute value of the) load exerted on the beam measured at the central load introduction point. The load on the beam was compressive during the complete fatigue cycle, thus the R-value is designated as R=10 (compression-compression fatigue), although the bottom flange will have been in tension, so that R=0.1 is probably more appropriate to describe the fatigue cycles in the bottom flange. More results will be published, among others, in e.g. [17].
As was mentioned, the beam was dimensioned so that the bondine was likely to be critically loaded in shear
during static testing. To support analysis of the bondline in fatigue, shear specimens according to ASTM 7078 [18] were tested in R=0.1 fatigue. The test results to date are reproduced in Figure 21.
A comparison between the shear strength achieved in the beam tests to the shear strength according ASTM 7078
can be made. In it’s simplest form the shear stress in the bonding paste of the beam can be derived if the longitudinal stresses in the shear web are neglected and uniform over the thickness of the flanges. The shear stress in the bondline can then be calculated as:
wh
Fbond ⋅
=21τ
Where F is the applied load, h the shear web height (100 mm) and w the bondline width (taken as ca. 19 for S79 and 12.5 for S78). The data points of Figure 20 are added in Figure 21. The results indicate that tests on bonding paste using ASTM 7078 are relevant for the behaviour of bonding paste in bondlines as used in WT blades.
Figure 17: Beam test set-up
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Figure 18: Beam after static loading. Wooden blocks are load reinforcement blocks
Figure 19: Two different subcomponent beam cross sections
Figure 21: Shear fatigue of bonding paste per ASTM 7078. Open squares indicate shear stress in beam
bonding paste (estimated according to equation above)
III. Concluding remarks In this paper, work carried out in current research projects shows, how various materials research topics are
approached in an integrated manner; • A reference material is defined and extensively described in terms of strength and fatigue. • Extensive off-reference tests are done to investigate various effects separately. • Experimental data are obtained at different scale levels.
The potential of using statistical methods for obtaining the model parameters is explored in the current projects,
as well as the expansion from meso-mechanical (coupon-based) experimental research, to micromechanical (single fibres) up to subcomponent level (beam-type structures).
IV. Acknowledgments The work described is part of the UPWIND and INNWIND projects.
V. References [1] UPWIND Web site, www.upwind.eu [2] INNWIND Web site, www.innwind.nl [3] OPTIMAT Web site, http://www.wmc.eu/optimatblades.php [4] Nijssen, R.P.L., Westphal, T., Stammes, E., Lekou, D., Brøndsted, P., ‘Rotor structures and materials – strength and fatigue
experiments and phenomenological modelling’, proc. European Wind Energy Conference 2008, Brussels Expo, March 31st – April 3rd, 2008)
[5] van Leeuwen, D.A., Nijssen, R.P.L., Westphal, T., Stammes, E., ‘Comparison of static shear test methodologies; test results and analysis’, proc. Global Wind Energy Conference, subject no. 2.11, China International Exhibition Centre (new venue), Beijing, China, October 29th – 31st, 2008
[6] ASTM 6641, ‘Standard Test Method for determining the compressive properties of polymer matrix composite laminates using a combined loading compression (CLC) test fixture’, 2001
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[7] ASTM D3528 – 96, ‘Standard test method for strength properties of double lap shear adhesive joints by tension loading’, 1996 (reapproved 2002)
[8] ASTM 5379, ‘Standard test method for shear properties of composite materials by the V-notched beam method’, 1998 [9] Sari, J., Nijssen, R., Westphal, T., Stammes, E., ‘Statistical analysis of static and fatigue strength characteristics of wind
turbine blade materials’, European Wind Energy Conference (EWEC), 2009 [10] van Wingerde, A., ‘Evaluation on the effect of thick laminates’, OPTIMAT report OB_TG4_R013, May 2005 [11] Hahn, H.T. and Turkgenc, O., The effect of loading parameters on fatigue of composite laminates: part IV information
systems, 2000, contract report DOT/FAA/AR-00/48 [12] Varna, J., ‘Stiffness degradation models’, UPWIND WP 3.2 report, deliverable D.3.2.1.b [13] Mishnaevsky, L. and Brøndsted, P., ‘Micromechanical modeling of strength and damage of fiber reinforced composites’,
annual report on EU FP6 Project UpWind Integrated Wind Turbine Design (WP 3.2), Risø-R-1601, March 2007 [14] Mandell, J.F., Samborsky, D.D., Cairns, D.S., ‘Fatigue of composite materials and substructures for wind turbine blades’,
Sandia National Laboratory report: SAND2002-0771, March 2002 [15] Mandell, J.F., Samborsky, D.D., ‘DOE/MSU composite material fatigue database: test methods, materials, and analysis’,
Sandia National Laboratory report: SAND97-3002, 1997 [16] Qian, C., Nijssen, R., Jansen, E., van Delft, D., Gürdal, Z., Zhang, G., ‘Micro-mechanical modelling of wind turbine rotor
blade composites – Static propertes of fibre bundles’, European Wind Energy Conference (EWEC), 2009 [17] Nijssen, R., Westphal, T., ‘Beam test report’, WMC-2009-32, under construction [18] ASTM 7078, ‘Standard test method for shear properties of composite materials by V-notched Rail shear method’, 2005
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