- Email: [email protected]
epoxy unidirectional material
Accepted Manuscript Influence of environment conditioning on the interlaminar fracture toughness of a graphite/epoxy unidirectional material L. Boni, D. Fanteria, L. Lazzeri, E. Panettieri, U. Mariani, M. Rigamonti PII:
S1359-8368(17)33988-4
DOI:
10.1016/j.compositesb.2018.07.044
Reference:
JCOMB 5811
To appear in:
Composites Part B
Received Date: 16 November 2017 Revised Date:
19 June 2018
Accepted Date: 22 July 2018
Please cite this article as: Boni L, Fanteria D, Lazzeri L, Panettieri E, Mariani U, Rigamonti M, Influence of environment conditioning on the interlaminar fracture toughness of a graphite/epoxy unidirectional material, Composites Part B (2018), doi: 10.1016/j.compositesb.2018.07.044. This is a PDF file of an unedited manuscript that has been accepted for publication. As a service to our customers we are providing this early version of the manuscript. The manuscript will undergo copyediting, typesetting, and review of the resulting proof before it is published in its final form. Please note that during the production process errors may be discovered which could affect the content, and all legal disclaimers that apply to the journal pertain.
ACCEPTED MANUSCRIPT Influence of environment conditioning on the interlaminar fracture toughness of a graphite/epoxy unidirectional material L. Bonia, D. Fanteriab, L. Lazzeric, E. Panettierid, U. Marianie, M. Rigamontif a
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Email: [email protected] – Department of Civil and Industrial Engineering, University of Pisa, Via G. Caruso 8, 56122, Pisa, Italy. b Email: [email protected] – Department of Civil and Industrial Engineering, University of Pisa, Via G. Caruso 8, 56122, Pisa, Italy. c Corresponding author, email: [email protected] – Department of Civil and Industrial Engineering, University of Pisa, Via G. Caruso 8, 56122, Pisa, Italy d Email: [email protected] – Department of Civil and Industrial Engineering, University of Pisa, Via G. Caruso 8, 56122, Pisa, Italy. e Email: [email protected] – Leonardo Helicopter Division, Via Giovanni Agusta 520, 21017, Cascina Costa di Samarate, Varese, Italy. f Email: [email protected] – Leonardo Helicopter Division, Via Giovanni Agusta 520, 21017, Cascina Costa di Samarate, Varese, Italy.
Abstract
An experimental program has been carried out at the Department of Civil and Industrial Engineering of the University of Pisa, in collaboration with Leonardo Helicopter Division (formerly
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AgustaWestland), for the evaluation of the interlaminar fracture resistance of Hexcel 913C-HTA, a graphite/epoxy system commonly used in many aeronautical platforms. The study comprised mode I tests (carried out on DCB specimens), mode II tests (using ENF specimens) and Mixed Mode I +
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II tests (carried out by means of the Mixed Mode Bending procedure). A complete description of
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the fracture toughness was so obtained, at Room Temperature and on dry specimens. This picture was completed by the assessment of temperature and moisture effects, by means of a few additional tests, performed at high temperature, 70 °C, on DCB and ENF saturated coupons, conditioned in an environmental chamber at 70 °C and 85% RH. The results showed an increase in mode I resistance and a decrease in mode II toughness, in agreement with similar results available in the literature. Keywords A: Polymer-matrix composites (PMCs) B: Fracture toughness
B: Delamination
B: Environmental degradation
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ACCEPTED MANUSCRIPT List of symbols and abbreviations CV
Coefficient of Variation
HTW High Temperature Wet
CC
Compliance Calibration
DCB Double Cantilever Beam
W
Specimen Width
ENF
t
Specimen Thickness
MMB Mixed Mode Bending
E1f
Flexural Modulus
MBT Modified Beam Theory
χ
Delamination Length Correction Parameter
C
Compliance
G
Strain Energy Release Rate
a
Delamination Length
ERR
Energy Release Rate
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End Flexure Notch
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RTD Room Temperature Dry
Introduction
Airworthiness regulations require that, for a safe operation, any flight critical element be capable to
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sustain significant load conditions even in the presence of some damage. For components made of composite materials, the demonstration of such damage tolerance capability must be given by means of a full-scale test. Among the forms of damage that are considered more insidious, low
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speed impact damages are particularly important as they can reduce considerably the strength of the component without being clearly evident. Therefore, a flood of literature has been written on the
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delamination behaviour, since delaminations are typically produced by impacts. Aircraft and helicopter manufacturers are commonly engaged in research programs for the assessment of the interlaminar fracture toughness of the composite material systems used in their applications, because high toughness values are appreciated for an easier compliance with regulatory requirements. Moreover, such requirements must be met in all the situations, for the whole operative life. This obvious consideration points out the importance of evaluating the resistance of composite materials to delamination not only in standard laboratory environment, but also in extreme environments, rare but possible in the whole service life of the aircraft or helicopter. -2-
ACCEPTED MANUSCRIPT Typically, hot / wet situations are considered more demanding and so, with these motivations, an experimental program has been defined and carried out in the framework of a collaboration between the Department of Civil and Industrial Engineering of the University of Pisa and Leonardo Helicopter Division, formerly AgustaWestland.
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A relatively reduced number of papers has been published on the interesting problem, fundamental from the designer point of view, of characterizing the interlaminar fracture resistance of composites, keeping also hygrothermal influences into account. Limiting our attention to unidirectional
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carbon/epoxy material systems, it is worth mentioning the papers by Russell and Street, [1], one of the earliest to focus on such a problem for AS1/3501-6 and HMS/3501-6, more than thirty years
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ago. Indeed, test results reported in [1] showed for AS1/3501-6 a mild decrease in mode I nonprecracked toughness with increasing temperature (in those days, standard test methods were not yet defined), while the opposite effect was observed in the delamination growth phase. Béland et al., [2], also found that the critical toughness in AS1/3501-6 was decreasing with temperature, while
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in IM6/5245C the opposite trend was observed. This conflicting situation could be the result of the lack of common test standard procedures in those pioneers’ days; the important role played by the insert (and its thickness) was not yet underlined. Further experimental work, [3-5], carried out about
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ten years later, showed more clearly that an increase in mode I toughness is normally observed when the temperature and the moisture content increase. Asp [6] and Cowley [7], towards the end
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of the nineties, confirmed this tendency with their experimental evidence. A physical explanation of this phenomenon has been attributed to the fact that an increase in temperature and moisture content is associated with a more plastic behaviour of the matrix, and so the increased matrix ductility was considered, also on the basis of fractographic examination of failure surfaces, as the major responsible for the higher mode I resistance; a different failure mechanism should be active for mode II situations, as a typical tendency to a decrease of mode II toughness with temperature was observed (Russell [1], Asp [6], Kim [8] and Johnson [9]). This
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ACCEPTED MANUSCRIPT behaviour has been ascribed to a degradation of the fibre/matrix bond strength (confirmed by the observation at SEM of a higher number of bare fibres after a high temperature test). Davidson (2009, [10]) studied the influence of moisture content and of test temperature on interlaminar fracture toughness of a thermoplastic-particulate toughened carbon/epoxy system
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(T800H/3900-2). The results from Double Cantilever Beam (DCB), End Notched Flexure (ENF) and Single Leg Bending (SLB) specimens are in line with commonly observed trends: mode I toughness increases with moisture content and temperature, while mode II toughness decreases;
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intermediate behaviour was observed for mixed-mode situations.
As far as the influence of the moisture uptake is considered, its effect was found to be slightly
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detrimental, and sometimes even difficult to be appreciated. Hintikka et al. recently (2010, [11]) evaluated at room temperature (RT) specimens characterized by a different moisture content (dry, at normal ambient humidity and moisture saturated) showing that the effect is slightly appreciable in mode I (where an increase in ductility with increasing moisture content is observed), and almost
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negligible in mode II and mixed mode situations.
More recently, Landry et al. [12] has studied the effect of the contact with aerospace fluids, such as water, hydraulic fluid and deicing fluid; his experimental results from End Notch Flexure (ENF)
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specimens show the adverse influence of the exposure to various fluids on the fracture resistance. One of the most recent papers on the subject is by Davidson (2016, [13]), who has studied the
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elevated temperature effect on the mode I, mode II and mixed (I+II) mode toughness of a carbonpolyimide system: the trend is perfectly in line with the general finding of an increase in mode I toughness and a decrease of mode II toughness, with mixed mode situations being intermediate. In conclusion, the trend emerging from the literature is rather consistent (thermoplastic resins have been deliberately excluded from the review, as they are less sensitive to hygro-thermal effects), but nevertheless the aerospace industries have a considerable interest to investigate the influence of moisture and temperature conditions on the interlaminar delamination toughness of commonly used
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ACCEPTED MANUSCRIPT composite materials, for design and certification purposes. An experimental program has been carried out at the University of Pisa in collaboration with Leonardo HD for the quantitative assessment of such environmental conditions, and it will be described in the following paragraph. Experimental activity
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The material selected for this activity is Hexcel 913C-HTA-12K, a low temperature curing (125 °C) graphite/epoxy system, often used in many helicopter applications, such as rotor blades or hubs.
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The experimental activities consisted in mode I DCB tests and mode II ENF tests, carried out in
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Room Temperature Dry (RTD) and High Temperature Wet (HTW) conditions, and Mixed-Mode Bending (MMB) tests, performed only in RTD conditions. Nominal dimensions of the specimens were 25 mm width, 3 mm thickness (24 plies, at 0°) and 150 mm length. All the specimens were manufactured by Leonardo HD following the internal company specifications and procedures. The mode I tests were carried out following the ASTM D-5528 standard [14], applied both for the
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RT tests on dry specimens and for the HT tests on wet specimens. Aluminium hinges were bonded to the DCB and MMB specimens, by means of an epoxy adhesive, curing at RT. After bonding the hinges, one lateral side of the specimen was painted white, with an acrylic spray painting, in order
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to help the identification of the delamination tip position and the measurement of its distance from
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the hinge axis; to make easier this measurement, a set of marks was traced on the white paint, at regular intervals, as shown for instance in Figure 1. The tests were carried out in displacement control (1 mm/min displacement rate) and force and displacement were recorded throughout the test. Moreover, a camera was used to take pictures of the marked specimen side (1 picture every second) in order to identify the tip position during the delamination growth. The Mixed Mode Bending tests were carried out following the ASTM D-6671 procedure, [16]. Two mode partitions were studied, defined by a GIc /GIIc ratio equal to 2:1 and to 1:2 (or, as more often -5-
ACCEPTED MANUSCRIPT indicated in the literature, GII /Gtot = 0.33 and 0.66). The test data acquisition system, used for the DCB specimens, was also used in these cases: the delamination growth was documented by a number of pictures, collected at the end of the test, and by a file with the load and displacement of the actuator head.
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The mode II tests were carried out using the ENF specimen, loaded in three points bending, with a total span between the external rollers of 100 mm. At the time of the test execution, the relevant ASTM standard [15] had not yet been issued and so an internal Leonardo HD standard was
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followed; all the specimens were pre-cracked before being tested, by means of a three-pointbending fixture, with a shorter distance between the external rollers (total span of 50 mm); a 3-4
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mm natural delamination was introduced, from the insert. Then, the tests were carried out on precracked ENF specimens, using a 100 mm total span. For the wet specimens tested in high temperature, the compliance calibration method for data analysis (that has later been recommended in [15]) was applied; also in this case, a pre-cracking operation was previously executed at RT, in
Procedure for the analysis of the test data
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order to obtain a natural delamination front, originated from the implanted kapton insert.
In order to compare the various results obtained in the different types of tests, a consistent method
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for deriving the fracture toughness value is necessary. So, the three possibilities offered by the ASTM D-5528 ([14]), namely Modified Beam Theory, Compliance Calibration and Modified Compliance Calibration, cannot be used, as they are limited to the mode I test and rely on empirical correction parameters derived during the mode I test itself. The ASTM D-7905 ([15]), that rules the mode II testing, suggests the use of the Compliance Calibration method, that for the ENF specimen simply requires to measure the initial stiffness in correspondence with a number of delamination lengths, sliding the specimen with respect to the test fixture. In the case of the DCB specimen this
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ACCEPTED MANUSCRIPT option is obviously much more complicated, and the specimen compliance cannot be measured for various delamination lengths before the test, but only during the performance of the test itself. It has been therefore decided to use the Kinloch-Williams approach [17-18], adopted in the ASTM D-6671 standard ([16]), that introduces a numerically defined correction to the delamination length
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to explain the lower stiffness of the sub-laminate leg with respect to the beam theory predictions, based on the perfect clamp hypothesis. The delamination length correction involves the factor χ computed according to Equation ( 1 ): where Γ = 1.18
E11 E 22
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2 Γ 3 − 2 1 + Γ
G13
,
(1)
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E11 χ= 11G13
where E11 and E22 are the tensile elastic moduli in the longitudinal and transverse direction and G13 is the shear modulus in 1-3 out-of-plane direction.
Indeed, the correction was deduced in [17-18] on the basis of numerical studies carried out on
Results of the experimental program
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unidirectional graphite/epoxy specimens, and so it is well suited for being used in this similar case.
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The results obtained will be presented in the following, divided by type of test, together with comments and comparisons with literature data (if available). The tests on dry specimens will be presented first. 4.1
DCB tests
Fifteen DCB dry specimens (three batches of five samples each) were manufactured and were tested following the procedure described in ASTM D-5528 [14]. A preliminary pre-cracking was carried out, which allowed the measurement of the material’s bending stiffness. This measurement is particularly valuable, because it refers to an “ideal” condition: a straight delamination front (the end
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ACCEPTED MANUSCRIPT of the insert) and a delamination tip zone that is pristine, free from any damage and/or plastic deformations: these two events allow a reliable estimate of the flexural elastic modulus, a parameter useful for subsequent numerical analyses. The bending modulus of the Non-Pre-Cracked (NPC)
[16]): 3
(2)
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E1 f
t 64 m a + χ 2 = 3 Wt
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specimens has been calculated with the Kinloch-Williams correction to the delamination length (see
where m is the slope of the force-displacement plot, W and t are the specimen width and thickness;
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a is obviously the delamination length, while the correction parameter χ has been defined in Eq. (1). The average value of the bending modulus is 125037.1 MPa, with a standard deviation of 5658.8 MPa (so giving a coefficient of variation equal to 4.53 %).
For what concerns the material properties to be used in eq. (1), the following values were provided
1.72.
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by Leonardo HD: E1 = 137 GPa, E2 = 12.1 GPa, G13 = 5.98 GPa; with such inputs, χ results to be
After pre-cracking, the real test took place, starting from a “natural” delamination (PC). For every
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test it was possible to collect about 5-600 pictures, each one with the current value of the load and the displacement written on it. An example is shown in Figure 1. After the end of the test, the
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pictures were examined and, for a selected number of significant pictures, the relevant compliance was computed. It was therefore possible to calculate the various corrections, suggested by the ASTM D-5528, for keeping into consideration the lower stiffness of the two sub-laminate legs.
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Figure 1 – Example of picture of a DCB specimen (DCB-S-04) during the test. The marking lines spacing is 1 mm on the left (beginning of the propagation) and then 5 mm.
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Figure 2 – Example of determination of the delamination correction length in MBT analysis (specimen: DCB-S-04; C units: mm/N)
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Figure 2 shows an example of linear relationship between delamination length and cubic root of compliance (indeed, it is analogous to Fig. 4 of [14]); this plot allows the determination of a parameter, the crack length increment, to be used in the Modified Beam Theory (MBT) methodology for evaluating the GIc values. The ASTM D-5528 standard procedure ([14]) identifies also three different “critical” points for the calculation of GIc, namely the Non Linear (NL) point (end of the linear part of the force-displacement plot), the VIS point (first visible delamination growth) and the 5%/Max point (the point of maximum force or the intersection of the experimental graph with the 5% reduced stiffness line, whatever comes first). The analysis carried out with the -9-
ACCEPTED MANUSCRIPT MBT methodology and referred to the NL point normally provides the most conservative estimate among the nine total options available in the ASTM D-5528: in this case, the average value of the fifteen tests is 229.61 N/m (standard deviation: 15.93 N/m, CV: 6.9%). In [19] a small test activity was carried out on DCB of this same material and an average initiation
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fracture toughness of about 189 N/m was found, with a plateau of about 250 N/m at the end of the R-curve. Such results are the only fully comparable that the authors have found in the open literature and it is interesting to observe a reasonable agreement.
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Since in this paper comparisons are made between critical ERR values in different mode mix situations, it was considered appropriate to follow a common approach/methodology to evaluate G:
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the Kinloch-Williams Gc value has been calculated for all the specimens. Such an estimation requires first the assessment of the bending modulus; as already explained before, it is deduced from the slope m of the Force-Displacement plot by means of Equation ( 2 ). The average value of the bending modulus measured during the test of the fifteen PC specimens is
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112682.4 MPa, a value significantly lower than the corresponding modulus deduced from the precracking phase (NPC coupons), 125037.1 MPa. This difference can be attributable to two causes: (a) the natural delamination front is not straight any more, but slightly curved, with a longer length
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in the middle of the specimen, and so the dimension measured on the lateral side is an approximation by defect; (b) the plastic zone and the damage state at the tip contribute to a reduced
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constraint in the “clamp”, that induces a lower stiffness. Similar findings were already presented and discussed by the authors, [20], for other experimental results. The Strain Energy Release Rate value is calculated, for the NL point, by means of the Kinloch-Williams formula (see [16]): t 96 P a + χ 2 GI = 2 3 W t E1 f
2
2
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(3)
ACCEPTED MANUSCRIPT and the average Kinloch-Williams value for GIc is 243.98 N/m (standard deviation: 15.17 N/m, CV: 6.2%). These results confirm the reasons why, as already mentioned before, the MBT is widely used in deriving values for design applications: it gives conservative values, while the Kinloch-Williams
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methodology provides a 6 % higher value.
Figure 3 shows, for the 3 tested batches, the force vs displacement curves directly obtained from each of the RTD PC DCB tests (left part of Figure 3) and the distribution of the GIc values evaluated
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for each specimen according to the Kinloch-Williams data reduction method (right of Figure 3),
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referred to the NL point.
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Figure 3 – Force vs displacement (left) of the DCB RTD tests of the specimens belonging to the 3 batches; inter-laminar fracture toughness values (right) of the same specimens evaluated with the Kinloch-Williams data reduction method.
ENF tests
In the literature there is wide evidence that tests carried out on ENF specimens without any precracking operation (NPC, i.e. with the delamination starting at the insert) provide much higher fracture toughness values than in the case of tests carried out on pre-cracked (PC) ENF specimens. Just as an example, a reading of O’Brien’s paper [21] and/or [22] is recommended. Therefore, also - 11 -
ACCEPTED MANUSCRIPT in consideration that the aim of Leonardo HD was to evaluate the material performance in view of design applications, it was decided to pre-crack all the ENF specimens. To this end, the three-pointbending test fixture was modified with a reduced span between the external rollers (namely 50 mm total span), and mounting the ENF so to leave a distance of about 4-5 mm between the insert end
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and the central loading nose. The coupon was then loaded in displacement control, until a load drop was observed, typically associated with a sharp noise and a contemporary visual crack growth, up to the arrest position under the loading nose, or a bit farther. The same type of PC operation was
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carried out both on dry and on wet ENF specimens.
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Figure 4 - Force vs displacement (left) of the ENF RTD tests of the 2 batches; inter-laminar fracture toughness values (right) computed with the Kinloch formulation and the 5%/Max criterion.
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Two batches of five ENF dry specimens each were available. For each test, a load-displacement plot was obtained (e.g. see left part of Figure 4 for the dry specimens), with a large linear part (after a short initial non-linear phase, where the contact is established between the specimen and the loading nose), with evidence of non-linearity only in the very last part (just before the load drop). The linear part of the plot was used to deduce the flexural elastic modulus of the material; this operation was made considering the physical delamination length measurable visually on the lateral surface of the ENF. The Kinloch-Williams formulation was used to deduce the GIIc values, which are shown in graphical form in Fig. 4, right. These GIIc values have been calculated using the - 12 -
ACCEPTED MANUSCRIPT maximum force; this has been considered a better option than the NL point, because in the wet specimens a larger non-linear phase (see Figure 5) was observed in the final phases of the test. Therefore, the comparison of HTW and RTD ENF results on the basis of the NL point would be penalizing for the HTW data, because onset of non-linearity in wet specimens is not an indication of
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imminent failure.
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Figure 5 – Force-displacement plot of HTW test on specimen ENF-Sw5.
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The formula used to compute E1f is, according to [16]:
E1 f
3 3 t m 2 L + 3 a + 0.42 χ 2 = W t3
(4)
where m is the stiffness of the specimen, W is the width and t the thickness; GII was calculated by means of the following relationship (see [16]):
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ACCEPTED MANUSCRIPT t 9 P a + 0.42 χ 2 GII = 2 3 2W t E1 f
2
2
(5)
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All the ten dry specimens available for the RT tests were tested setting the delamination tip at a = 25 mm, with a half span equal to 50 mm.
The results show an average bending modulus equal to 118671.5 MPa (standard deviation: 3800.8 MPa, CV: 3.2%), a value 5% higher than 112682.4 MPa, that was the average for the pre-cracked
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DCB specimens, while they should be fully comparable. The two situations differ for the fact that
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the DCB legs deform separating from each other, while in ENF specimens the two sub-laminates are pressed one against the other and, moreover, tend to slide one against the other, with the separation surface composed by two different areas: a natural delamination (about 4-5 mm in length), plus an area covered by the kapton insert (about 20 mm long). So it is expected that the DCB estimated modulus be “true”, while the ENF deduced modulus may be affected by the sliding
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movement that, through friction, is perceived as an additional stiffness. Moreover, another consideration must be made with reference to the delamination length used in
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equations (2) and (4), that was based on the Kinloch-Williams correction applied to the physical length, estimated visually on the lateral surface of the specimen. In DCB specimens, the GI
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distribution in the specimen width is approximately parabolic, with its maximum in the centre, while in ENF specimens GII is distributed in the opposite way, with maximum on the lateral surfaces and the minimum in the centre. The delamination front resembles the G distribution. Therefore, the visual estimation provides an under-estimate of the “effective” length for the DCB, and an over-estimate for the ENF. So the bending stiffness for the ENF, being referred to a slightly longer delamination length than the “effective”, results a bit higher; the opposite occurs for the DCB specimens.
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ACCEPTED MANUSCRIPT Unfortunately, values of E1f measured from ENF specimens in the NPC state are not available. Further analyses have been carried out to evaluate the possible influence of friction in ENF tests.
4.3
MMB tests
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Interlaminar fracture toughness was evaluated only on dry specimens, at RT. Two groups of specimens were tested, at two mode partition values: GII /Gtot = 0.33 and 0.66. Five samples were available for each mode partition, but only four specimens produced valid results, because of a
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hardware failure in one case and of a strange, singular delamination behaviour (bifurcation) in another case. The recommendations of the ASTM D-6671 procedure [16] were followed. Also in
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this case, the delamination growth was monitored by capturing pictures, and so at the end of the test a number of photographs were available, for the study of the propagation phase (R-curve, that anyhow is not presented in the paper). Figure 6 shows displacement vs. force plots, with evidence
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of linearity for a long part of the recorded trace.
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Figure 6 – Force vs displacement (left) of the MMB RTD tests; inter-laminar fracture toughness values (right) computed with the Kinloch data reduction and the VIS criterion. The Strain Energy Release Rate components in mode I and mode II can be computed by means of
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the equations reported in [16]:
9 P (3 c − L ) a + χ GI = 2W 2 t 3 E1 f
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2
2
t 2
2
t 2 9 P 2 (c + L ) a + 0.42 χ 2 GII = 2 3 2W t E1 f
(6) 2
(7)
where E1f can be deduced from the slope of the linear part of the force-displacement plot. Indeed, the ASTM recommended procedure allows the determination of the delamination initiation toughness (two options are available: NL and 5%/Max) or the study of the propagation (to derive an R-curve, when the propagation takes place in a stable way). The first value is a Non-Pre-Crack - 16 -
ACCEPTED MANUSCRIPT (NPC) value, derived from the insert, and so susceptible to criticism, because it is widely recognized that NPC mode II toughness values are significantly (and artificially) higher than those obtained in tests on Pre-Cracked (PC) specimens; a similar consideration applies also to mode I, with minor quantitative effects. Therefore, for the comparison with pure modes data, it was decided
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to use the MMB results from the VIS point, i.e. the point in the test when the delamination growth from the insert has been observed visually. In the tests with prevailing mode I conditions (GII /Gtot = 0.33), the propagation was very stable and the VIS situation was characterized by a delamination
very quick, with a few millimetres of sudden growth.
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growth of a few tenths of mm; in the tests with prevailing mode II loading, the propagation was
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It is important to note that the Gc values, shown as column plot in Fig. 6, have been derived by making use of E1f, estimated in NPC conditions, but applied in a PC condition (VIS). In the elaboration of the DCB and ENF specimens, on the contrary, the bending modulus E1f was estimated in the PC condition and combined with data (force and displacement) from the test, that
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was also carried out in PC condition.
All the results have been plotted in Figure 7, together with a Benzeggagh-Kenane [23] best-fit (8):
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the global appearance is in line with similar results available in the literature.
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Gc = G Ic + (G IIc − G Ic )(G II / GT )η
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(8)
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Figure 7 – Average and standard deviation (error bars) of the values of the inter-laminar fracture toughness obtained for pure and mixed modes; the dashed line represents the B-K curve that fits the experimental data.
High temperature testing of conditioned specimens
Five DCB and five ENF specimens were conditioned by Leonardo HD in a climatic chamber at 70
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°C and 85% RH up to saturation. Then, the specimens were hermetically packed in sealed envelopes for the shipment to Pisa, where they have been maintained in such package until the day
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of the test. Hinges were cold bonded to DCB specimens by means of an epoxy adhesive, curing at RT; the day after, the specimens were ready and their sides were painted white and marked with ticks. The preparation of ENF coupons was simpler, obviously, requiring only the precracking operation, performed in shear, at RT and marking on the edges. The static tests were performed in a thermally insulated bakelite box, where hot air was insufflated by an industrial heater. Each specimen was instrumented with thermocouples (2 for each DCB, 3 for each ENF) and the transient in temperature from RT to 70 °C took about 25 minutes, with an average heating rate of about 2 °C per minute. No humidity control was performed during the test. - 18 -
ACCEPTED MANUSCRIPT The bakelite box has two large quartz glass windows, so that it was possible to continue to use the same test apparatus already used in the RT tests, capturing and storing pictures of the delamination tip position. Therefore, the DCB tests were carried out following the recommendations of [14], with a precracking phase, followed by the actual test: the data collected allowed to estimate the
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parameters required for correcting the Beam Theory and Compliance Calibration approaches. As far as the ENF specimens are concerned, they were all subjected to precracking, in RT conditions (similarly to the dry specimens), and then to a Compliance Calibration campaign, also at RT. The
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compliance C was measured for the following delamination lengths: 18, 20, 23, 25, 27, 29 mm; for
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completeness, also a measure of the bending stiffness in absence of delamination was performed.
Figure 8 – Example of the CC methodology on a HTW specimen (ENF-Sw4).
All these measurements at RT allowed to collect data for tracing a C vs. ‘a cube’ plot for each specimen; a typical example is shown in Fig. 8. The least squares best-fit linear equation provides the derivative of C with respect to ‘a’, and so the following formula could be used for G:
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ACCEPTED MANUSCRIPT GII =
P 2 ∂C 2W ∂a
(9)
and, with the approximation that C = ma 3 + C0 it follows that:
∂C = 3ma2 ∂a
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( 10 )
This is the CC method, recommended in the ASTM standard [15], and its results were compared
6.1
Comments on the results of the Hot-Wet tests
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6
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with those from the Kinloch-Williams formula (eqn. 5).
Mode I tests
Differently from the analogous RT tests on dry specimens, a more pronounced bridging effect is observed, that is reflected in an appreciable G growth with delamination length; this phenomenon
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has obviously consequences on the semi-empirical parameters that are derived from the test to be used in the MBT, CC and MCC approaches. The results, in terms of Modified Beam Theory,
point):
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relevant to the NL point, are reported in Table II below (P and δ are the coordinates of the NL
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Specimen
DCB-Sw1 DCB-Sw2 DCB-Sw3 DCB-Sw4 DCB-Sw5
P (N) 54.94 52.47 52.47 52.22 50.82
δ (mm) 7.01 6.54 8.40 8.09 7.00
∆ (mm) 6.94 3.21 8.06 6.25 5.84
GIc (N/m) 374.67 354.35 384.03 414.12 345.59
Table I – Results of the HTW tests on DCB specimens (MBT methodology, NL point). The average value is 374.5 N/m (standard deviation: 26.9 N/m; CV: 7.2%), significantly higher than the average RTD value, 229.61 N/m; indeed, the behaviour of the wet specimens in the HT tests seemed clearly more ductile, with a pronounced plastic behaviour. The results have been calculated - 20 -
ACCEPTED MANUSCRIPT also by means of the Kinloch-Williams approach (average: 398.8 N/m; standard deviation: 35.2
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N/m; CV: 8.9%), and are shown in Figure 9 (right).
Figure 9 – Force vs displacement (left) of the DCB HW tests; inter-laminar fracture toughness values (right) computed with the Kinloch-Williams data reduction and the NL criterion.
The remarkable differences, obtained between the HT tests and the RT tests in terms of fracture
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toughness growth with crack length (R-curves), are shown in Figure 10. It turns out that not only the average fracture toughness increases significantly (1.5 times greater), but the bridging effect as well. In fact, while the average slope of the R-curves of the RT tests is essentially negligible, for the
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HT tests the growth of GIc with the crack length is appreciable, though affected by scatter.
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Figure 10 – R-curves obtained for the RT tests (specimens 1-5) and for the HT tests (specimens 1, 2, 4 and 5).
The Kinloch-Williams analysis allows also the comparison of the bending modulus estimated in the
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Non-Pre-Cracked (NPC) state, i.e. with the insert, with the same quantity in the pre-cracked condition (PC), i.e. with a natural delamination. The average value in the HTW NPC specimens is 125651.9 MPa, almost the same value observed in the RTD specimens, 125037.1 MPa: this may be
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ascribed to the fact that the modulus is a fibre dominated property. As a matter of fact, the difference between the RTD and HTW conditions is mainly in the dimensional dilatation of the
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matrix, that slightly increases the thickness, and so the bending stiffness; indeed, the difference is negligible, and can be estimated in less than 1% on the modulus. In the PC test, after pre-cracking, the delamination front is no more perfectly straight, and the measurement on the specimen side edge underestimates the effective value; moreover, damage is present close the tip. The bending stiffness deduced from the PC test data decreases with respect to the NPC value: in the HTW tests the average bending modulus is 119802.9 MPa, while in the RTD tests it dropped down to 112682.4 MPa. The difference between the two situations is quite significant; it can be attributed to the more
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ACCEPTED MANUSCRIPT plastic behaviour of the matrix, with reduced brittleness and therefore also a smaller damage in the tip region after pre-cracking.
6.2
Mode II tests
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The ENF wet specimens, before being tested in high temperature, had been previously subjected to a pre-cracking operation at RT, in shear. After pre-cracking and before the test, for each specimen the compliance vs. delamination length relationship was evaluated, in order to apply the
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Compliance Calibration (CC) methodology for estimating the Strain Energy Release Rate. Such compliance measurements were made at RT, and resulted in a quite consistent, low scatter results,
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as already shown in Figure 8. Moreover, also an experimental assessment of the bending modulus was carried out, shifting the specimen on the rollers in such a way that the loaded part was free from the delamination; the experimental value has been compared with C0, the extrapolated value, from the C = ma 3 + C0 relationship: a very good agreement has been obtained for each specimen.
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Then the real test took place, with a delamination length equal to 25 mm (i.e. the tip was in the middle of the span); the data collected allowed to estimate G, but it was also possible to assess the difference in modulus, ascribed only to the temperature change. The HTW interlaminar fracture
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toughness was estimated by means of the Kinloch-Williams approach and by the CC approach, with
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the ∂C/∂a derivative estimated at RT; this may be not correct. The results are summarized in Table II: the Kinloch-Williams approach gives an average value of 708.4 N/m, which is significantly lower than the RTD value, that was 1047.9 N/m. The CC approach gives a GIIc average value of 640.9 N/m, not so different from the K-W estimate (7% lower). It is interesting to observe the bending modulus measurements: the average value, derived from specimens with a delamination length equal to 25 mm, is 122531.8 MPa for the HTW tests, while the RT analogous value from the dry specimens was 118671.5 MPa. This last value is very consistent with the measurements made on the “un-cracked” ENF wet specimens at RT, i.e. the wet specimens placed in the experimental
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ACCEPTED MANUSCRIPT set-up in such a way to exclude the insert from the rollers span, which showed an average value of 118896 MPa, substantially identical to the RTD value. Table II shows also the directly measured quantities, i.e. the slope of the force-displacement plot, with an average increase of the slope of about 1.5% passing from RT to HT conditions. This is a strange result: the specimens show a slight
systematic, worth to be studied in depth.
Slope RT (N/mm) 291.19 328.05 325.69 339.79 316.08
E1f PC-KW (MPa) 120173.9 124851.3 120415.4 126859.3 120359.0
GIIc – KW (N/m) 642.3 734.4 711.5 741.4 712.2
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ENF-Sw1 ENF-Sw2 ENF-Sw3 ENF-Sw4 ENF-Sw5
Slope HT (N/mm) 295.65 331.30 328.82 347.52 321.46
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Specimen
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increase in stiffness while passing from RT to HT conditions. It is not a big difference, but
GIIc – CC (N/m) 657.0 602.0 630.4 625.5 689.4
Table II – Results of the HTW tests on ENF specimens. It should be underlined that the K-W analysis methodology was developed by means of numerical analysis that were fitted to RT tests; it is obviously straightforward to extend its applicability and
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pretend a reasonable accuracy in HT situations, where the failure mechanisms are sensibly different. Figure 11 shows the force-displacement plot of the HTW ENF tests, together with the toughness,
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with the K-W analysis methodology, while Figure 12 collects all the fracture toughness results.
Figure 11 – Force vs displacement (left) of the ENF HW tests; inter-laminar fracture toughness values (right) computed with the Kinloch-Williams data reduction and the NL criterion.
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Figure 12 – Comparison between the values of inter-laminar fracture toughness obtained in RTD conditions and the one obtained with HTW conditions for the pure modes.
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The temperature increase, in the case of the ENF coupons, seems to have some influence on the measured modulus, due to the occurrence of some mechanism not present in the DCB coupons, where the bending modulus has not shown substantial differences between the RTD and the HTW
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results.
RT and HT friction measurement
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To investigate this aspect, a measurement of the friction coefficient between the delaminated parts of a specimen was organized.
A simple inclined plane device has been designed and realized, as shown in Figure 13. The plane is hinged and can rotate, and is initially placed in horizontal position. Two pieces of delaminated coupons are placed on top of the plane, with a small weight on the upper piece. By applying a slow movement to the rotating plate, it has been possible to measure the angle, α, at the beginning of the upper portion shift, that is related to µ, the coefficient of friction, by means of the simple
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ACCEPTED MANUSCRIPT relationship: µ = tan(α ) . It has been considered appropriate to apply a small pressure between the two portions, and so a weight of about 1 N was fixed over the upper part. The dimensions of the specimens’ portions were 25 mm in width (obviously, as they were cut from the wedge opened
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specimens) and 30 mm in length.
Figure 13 – Friction measurement apparatus used in the friction tests.
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Moreover, a number of different situations have been examined, that differed one from the other for the extension of the kapton film layer, placed between the two pieces of delaminated coupons, so referring to different situations (delamination lengths) during the ENF test. In total, four different
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situations have been investigated, namely: a) no kapton film; b) 5 mm kapton film; c) 15 mm kapton film; d) 30 mm kapton film. It is obvious that the b) situation is representative of the test,
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where the precracking operation had introduced about 5 mm of “natural” delamination. For each of the four different situations, two specimens have been prepared and tested for three replicas at RT and other three replicas at HT (70 °C), to obtain information on the friction coefficient variability in function of the temperature, and also an assessment of the scatter. The scatter observed is high, particularly for the tests b) and c). Possible causes may be: a) the presence of a non-straight delamination front, which disturbed the measurement; b) the high roughness of the surface, generated by wedge opening: no flattening action or wear due to fatigue was present. Anyhow, a tendency was observed, namely an increase in the coefficient of friction - 26 -
ACCEPTED MANUSCRIPT with a lower extent of kapton at the interface: at RT, the average friction coefficient passed from 0.273 in the all kapton interface, to 0.306 for the 15mm/15mm interface, to 0.384 for the no kapton condition. The friction coefficient was influenced by the temperature, but in a contradictory manner: it passed from an average value of 0.384 at RT to 0.510 at HT for the all fibre interface and
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from an average value of 0.273 to 0.241 for the full kapton interface. The decrease of the coefficient of friction for the case with kapton interface at HT could seem puzzling, but is consistent with other literature results [24]. Notwithstanding this confused situation, a numerical study on the influence
Numerical evaluations
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of friction coefficient was carried out, taking the move from classic works, like [25] and [26].
A 2D FE model of a 3-mm thick ENF specimen has been developed in the Abaqus software in order to evaluate the friction effects on the flexural modulus computed according to Eqn. ( 4 ). To reproduce the ENF test, the specimen was sustained by means of 2 cylindrical rigid supports,
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with a span of 100 mm, and loaded by a central nose. Boundary conditions were assigned to the 2 supports through their reference points which were constrained to not translate and rotate. The central cylinder was loaded in displacement control, and the horizontal translation was inhibited,
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such as the in-plane rotation.
The specimen was modelled as two half-laminates tied in the un-cracked region while contact
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interactions were present along the delamination length. The total delaminated length, measured from the first support, was 30 mm: 25 mm were covered by the kapton insert while 5 mm was the length of the natural crack. The geometry of the model, as shown in Figure 14, represents the condition of a pre-cracked specimen. The FE model also presents part of the specimen outside of the left support so that the total delamination length, measured from the specimen left end, is 55 mm (25 mm externally to the support span and 30 mm internally).
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ACCEPTED MANUSCRIPT The two half-laminates were meshed with plane stress elements with a maximum mesh size of 0.5 mm along the specimen length (in the contact regions) and a very fine mesh size, 0.04 mm, around
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the delamination tip. In the thickness, 25 elements for each half-laminate were used.
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Figure 14 – Regions interested by the contact interactions (upper part of the figure) and details of the 2D numerical model of the ENF. To evaluate the effects of the friction contributions in the two regions, 2 tangential contact
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properties have been created to characterize the contact interactions assigned to the kapton insert region and to the natural delamination. Indeed, 2 coefficients have been assigned to each of the
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regions: for the tangential contact property in the kapton insert region, values of the friction coefficient, µt, of 0.2 and 0.3 were used, while, in the crack region, values of 0.3 and 0.4 were assigned to the friction coefficient, µc. Then, 2 non-linear static analyses have been performed by using the lowest and the highest friction coefficient pairs for the 2 contact regions, an arbitrary choice for assessing the influence of a high and a low friction value. Figure 15 shows the distribution of the compression stress and of the shear stresses due to friction, that are generated along the contact line of the half-laminates at the end of both simulations. The
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ACCEPTED MANUSCRIPT results show that the 2 half-laminates are pressed one against the other only in the area of the left support and, consequently, friction forces are generated only in such area. This result is consistent with the stress distribution reported in [25], where the effect of friction on the mode II fracture
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toughness values was investigated.
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Figure 15 – Compression stress and shear stresses along the crack length for the simulations performed with different friction coefficients. According to the results obtained with the simulations, the increase of flexural modulus in the HT
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tests is not fully justified only by the friction effects. Anyhow, another contribution can be ascribed to the transverse thermal expansion coefficient, αT, of the composite system. If a coefficient of
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thermal expansion of about 60 x 10-6 (°C)-1 is considered (typical value for an epoxy resin), a 50 °C temperature increase induces a thermal strain of 0.3% and so the bending stiffness, proportional to the third power of the thickness, increases of about 0.9%, spacing apart the stiff carbon fibres. The more “sticky” behaviour of the new, natural delamination length has been studied in classic literature papers (like [25] and [26]) assessing the influence of the friction coefficient: the influence is limited to the areas where significant pressure is applied between the two sub-laminates; in total, only a few percent increase in apparent toughness (around 2%) can be ascribed to this energy absorbing mechanism. In other words, in HTW conditions the true GIIc value reduces slightly - 29 -
ACCEPTED MANUSCRIPT (around 2%) with respect to the value observed experimentally. Anyhow, since the differences are really small, the influence in aviation design procedures is typically widely covered by the safety factors normally adopted.
Conclusions
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During the present investigation, fracture mechanics tests were carried out on dry and wet specimens, made of Hexcel 913C-HTA; a group of tests were performed at RT on dry specimens
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and another group on wet specimens, at HT. It was not possible to point out the separate influence on interlaminar fracture toughness of moisture and high temperature. Anyhow, the literature
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available on the subject (e.g. [10-12]) clearly underlines that the temperature is the most influencing parameter, with deeper effects, while moisture content has a lower impact. With respect to similar activities performed in the past and available in the literature, this activity has the merits of assessing the sensitivity to environmental conditions of the interlaminar fracture
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toughness of a low temperature curing (125 °C) epoxy resin. The HTW conditions of the tests are therefore relatively more severe than for the case of a 180 °C curing traditional epoxy. Anyhow, the variations observed with respect to the RTD performance are not quantitatively much different with
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respect to composites with traditional epoxy matrix. Nevertheless, it is always important to evaluate quantitatively such aspects, by means of dedicated tests: no theory can predict reliably the
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environmental influence on the interlaminar fracture toughness. The experimental activity has left a number of open questions, topics on which it is important to concentrate future efforts. The first one concerns the way of processing the data collected during the tests: are the rules typically developed and set-up for the evaluation of interlaminar fracture toughness at RT still applicable in a HT situation, where constitutive laws slightly change (a wider non-linearity range is expected)? The widely diffused MBT approach for mode I tests analysis requires the determination of a semi-empirical correction to the delamination length: the hypothesis has been adopted here the experimental procedures set-up for performing tests at RT could be - 30 -
ACCEPTED MANUSCRIPT applied also in HT and that the semi-empirical correction parameter keeps its meaning. The average Delta value (see Figure 2) is 6.06 mm for the HTW DCB specimens, a bit higher than the average RTD value of 5.63 mm; this is according to the expectations, because it is a measure of how far the real situation is from the ideal perfect clamp constraint. In HTW conditions, a less rigid constraint is
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expected. Moreover, the modification of the Beam Theory proposed by Williams and Kinloch [1718] is still meaningful? The ideal clamp constraint is certainly modified by the hygro-thermal conditions; further work is planned on this point.
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Anyhow, once that HT toughness values have been estimated, the results obtained in this experimental activity confirm that the mode I toughness increases with temperature, with a more
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plastic and ductile behaviour of the resin system, while a reduction in the mode II toughness is observed. This tendency is confirmed in the literature by fractographic examination of the failure modes, which gives evidence of a more plastic behaviour in mode I and a fibre/resin bonding interface of reduced strength in mode II tests.
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In mode II tests a discrepancy about the compliance of the ENF specimens has been observed between the RT values (during the CC evaluation) and at HT values, during the test performance: the stiffness of the specimens has shown a small increase with temperature. This point stimulated
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therefore a more detailed investigation, numerically and experimentally. A short campaign for the measurements of the friction coefficient between the two sub-laminates has been carried out,
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showing a slightly higher friction coefficient at high temperature, though in a context of measurements affected by large scatter. The numerical model of the experiment did not explain completely, with such input, the experimental observation; the concurrent presence of another factor, associated with the thermal expansion of the resin, can reduce the gap. Anyhow, the basic work carried out in the early years of the definition of the ENF specimens has found a clear confirmation. Further work is planned on this point, as well as on the questions arose before, on the data reduction methodology.
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References
[1] A.J. Russell, K.N. Street, Moisture and temperature effects on the mixed-mode delamination
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fracture of unidirectional graphite/epoxy. In: Johnson W.S., editor. Delamination and debonding of materials, ASTM STP 876. American Society for Testing and Materials, 1985, pp. 349–70.
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[2] S. Beland, J.P. Komorovski, C. Roy, Hygrothermal influence on the interlaminar fracture energy of graphite/bismaleimide modified epoxy composite (IM6/5245C). In Proceedings of the 6th
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Conference on Comp Mat. Comb/2nd European Conf. on Comp. Mat., London, pp. 3.305-3.316, 1987.
[3] A. Garg, O. Ishai, Hygrothermal influence on delamination behavior of graphite/epoxy
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laminates. Eng. Fract. Mech., vol. 22, pp. 413-427, 1985.
[4] R. Seizer, K. Friedrich, Influence of water up-take on interlaminar fracture properties of carbon
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fibre-reinforced polymer composites. J. Mater. Sci., vol. 30, pp. 334-338, 1995.
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[5] F.X. de Charentenay, J.M. Harry, Y.J. Prel, M.L. Benzeggagh, Characterizing the effect of delamination defect by mode I delamination test. In Effects of Defects in Composite Materials,
ASTM STP 836, American Society for Testing and Materials, Philadelphia, PA, pp. 84-103, 1984.
[6] L.E. Asp, The effects of moisture and temperature on the interlaminar delamination toughness of a carbon/epoxy composite. Comp Sci & Tech, vol.58 (6), pp. 967–97, 1998.
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ACCEPTED MANUSCRIPT [7] K.D. Cowley, P.W.R. Beaumont: The interlaminar and intralaminar fracture toughness of carbon-fibre/polymer composite: the effect of temperature, Comp Sci & Tech, vol. 57, pp. 14331444, 1997.
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[8] H.S. Kim, W.X. Wang, Y. Takao, G. Ben, Effects of temperature on mode II fracture toughness of multidirectional CFRP laminates. Int J Mod Phys; 17(8–9), pp.1717–23, 2003.
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[9] W.S. Johnson, M.M. Pavlick, M.S. Oliver, Determination of interlaminar toughness of IM7/9772 composites at temperature extremes and different thicknesses. NASA TR 20050192417, May
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2005.
[10] B.D. Davidson, M. Kumar, M.A. Soffa, Influence of mode ratio and hygrothermal conditions on the delamination toughness of thermoplastic particulate interlayered carbon/epoxy composite.
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Composites Part A; vol. 40, pp. 67–79, 2009.
[11] P. Hintikka, M. Wallin, O. Saarela, The effect of moisture on the interlaminar fracture
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toughness of CFRP laminate. 27th Congress ICAS, 2010, Nice (FR), Paper ICAS2010-8.3ST.
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[12] B. Landry, G. La Plante, L.R. LeBlanc, “Environmental effects on mode II fatigue delamination growth in an aerospace grade carbon/epoxy composite”, Composites Part A, vol. 43, pp. 475–485, 2012.
[13] M.W. Czabaj, B.D. Davidson, Determination of the mode I, mode II, and mixed-mode I–II delamination toughness of a graphite/polyimide composite at room and elevated temperatures. Journal of Composite Materials, vol. 50, pp. 2235-2253, 2016.
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ACCEPTED MANUSCRIPT [14] ASTM D-5528-01, Standard test method for mode I interlaminar fracture toughness of unidirectional fiber-reinforced polymer matrix composites. American Society for Testing and Materials.
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[15] ASTM D-7905-14, Standard test method for mode II interlaminar fracture toughness of unidirectional fiber-reinforced polymer matrix composites. American Society for Testing and
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[16] ASTM D-6671-04, Standard test method for mixed mode I – mode II interlaminar fracture
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toughness of unidirectional fiber-reinforced polymer matrix composites. American Society for Testing and Materials.
[17] S. Hashemi, A.J. Kinloch, J.G. Williams, Corrections needed in Double Cantilever Beam tests
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for assessing the interlaminar failure of fibre-composites, Journal of Materials Science Letters, vol. 8, pp. 125-129, 1989.
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[18] Y. Wang, J.G. Williams, Corrections for Mode II fracture toughness specimens of composite
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materials, Comp Sci & Tech, vol. 43, pp. 251-256, 1992.
[19] R.W. Truss, P.J. Hine, R.A. Duckett, Interlaminar and intralaminar fracture toughness of uniaxial continuous and discontinuous carbon fibre/epoxy composites, Composites Part A, vol. 28A, pp. 627-636, 1997.
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ACCEPTED MANUSCRIPT [20] D. Fanteria, L. Lazzeri, E. Panettieri, U. Mariani, M. Rigamonti, Experimental characterization of the interlaminar fracture toughness of a woven and a unidirectional carbon/epoxy composite, Comp Sci & Tech, vol. 142, pp. 20-29, 2017.
thresholds for composite materials, NASA TM 89157, August 1987.
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[21] K.T. O’Brien, G.B. Murri, S.A. Salpekar, Interlaminar shear fracture toughness and fatigue
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[22] K.T. O’Brien, W.M. Johnston, G.J. Toland, Mode II interlaminar fracture toughness and fatigue characterization of a graphite epoxy composite material, NASA/TM 2010-216838, August
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2010.
[23] M.L. Benzeggagh, M. Kenane, Measurement of Mixed-Mode delamination fracture toughness of unidirectional glass/epoxy composites with Mixed-Mode Bending apparatus. Comp Sci & Tech,
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vol. 56, pp. 439-449, 1996.
August 1972.
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[24] R.L. Fusaro, Friction and wear life properties of polyimide thin films. NASA TN D-6914,
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[25] L.A. Carlsson, J.W. Gillespie jr., R.B. Pipes, On the analysis and design of the End Notched Flexure (ENF) specimen for mode II testing, J. of Composite Materials, vol. 20, pp. 594-604, 1986.
[26] B.D. Davidson, X. Sun, Effects of friction, geometry, and fixture compliance on the perceived toughness from three- and four-point bend end-notched flexure tests, Journal of Reinforced Plastics and Composites, vol. 24, pp. 1611-1628, 2005.
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S1359-8368(17)33988-4
DOI:
10.1016/j.compositesb.2018.07.044
Reference:
JCOMB 5811
To appear in:
Composites Part B
Received Date: 16 November 2017 Revised Date:
19 June 2018
Accepted Date: 22 July 2018
Please cite this article as: Boni L, Fanteria D, Lazzeri L, Panettieri E, Mariani U, Rigamonti M, Influence of environment conditioning on the interlaminar fracture toughness of a graphite/epoxy unidirectional material, Composites Part B (2018), doi: 10.1016/j.compositesb.2018.07.044. This is a PDF file of an unedited manuscript that has been accepted for publication. As a service to our customers we are providing this early version of the manuscript. The manuscript will undergo copyediting, typesetting, and review of the resulting proof before it is published in its final form. Please note that during the production process errors may be discovered which could affect the content, and all legal disclaimers that apply to the journal pertain.
ACCEPTED MANUSCRIPT Influence of environment conditioning on the interlaminar fracture toughness of a graphite/epoxy unidirectional material L. Bonia, D. Fanteriab, L. Lazzeric, E. Panettierid, U. Marianie, M. Rigamontif a
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Email: [email protected] – Department of Civil and Industrial Engineering, University of Pisa, Via G. Caruso 8, 56122, Pisa, Italy. b Email: [email protected] – Department of Civil and Industrial Engineering, University of Pisa, Via G. Caruso 8, 56122, Pisa, Italy. c Corresponding author, email: [email protected] – Department of Civil and Industrial Engineering, University of Pisa, Via G. Caruso 8, 56122, Pisa, Italy d Email: [email protected] – Department of Civil and Industrial Engineering, University of Pisa, Via G. Caruso 8, 56122, Pisa, Italy. e Email: [email protected] – Leonardo Helicopter Division, Via Giovanni Agusta 520, 21017, Cascina Costa di Samarate, Varese, Italy. f Email: [email protected] – Leonardo Helicopter Division, Via Giovanni Agusta 520, 21017, Cascina Costa di Samarate, Varese, Italy.
Abstract
An experimental program has been carried out at the Department of Civil and Industrial Engineering of the University of Pisa, in collaboration with Leonardo Helicopter Division (formerly
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AgustaWestland), for the evaluation of the interlaminar fracture resistance of Hexcel 913C-HTA, a graphite/epoxy system commonly used in many aeronautical platforms. The study comprised mode I tests (carried out on DCB specimens), mode II tests (using ENF specimens) and Mixed Mode I +
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II tests (carried out by means of the Mixed Mode Bending procedure). A complete description of
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the fracture toughness was so obtained, at Room Temperature and on dry specimens. This picture was completed by the assessment of temperature and moisture effects, by means of a few additional tests, performed at high temperature, 70 °C, on DCB and ENF saturated coupons, conditioned in an environmental chamber at 70 °C and 85% RH. The results showed an increase in mode I resistance and a decrease in mode II toughness, in agreement with similar results available in the literature. Keywords A: Polymer-matrix composites (PMCs) B: Fracture toughness
B: Delamination
B: Environmental degradation
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ACCEPTED MANUSCRIPT List of symbols and abbreviations CV
Coefficient of Variation
HTW High Temperature Wet
CC
Compliance Calibration
DCB Double Cantilever Beam
W
Specimen Width
ENF
t
Specimen Thickness
MMB Mixed Mode Bending
E1f
Flexural Modulus
MBT Modified Beam Theory
χ
Delamination Length Correction Parameter
C
Compliance
G
Strain Energy Release Rate
a
Delamination Length
ERR
Energy Release Rate
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End Flexure Notch
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RTD Room Temperature Dry
Introduction
Airworthiness regulations require that, for a safe operation, any flight critical element be capable to
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sustain significant load conditions even in the presence of some damage. For components made of composite materials, the demonstration of such damage tolerance capability must be given by means of a full-scale test. Among the forms of damage that are considered more insidious, low
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speed impact damages are particularly important as they can reduce considerably the strength of the component without being clearly evident. Therefore, a flood of literature has been written on the
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delamination behaviour, since delaminations are typically produced by impacts. Aircraft and helicopter manufacturers are commonly engaged in research programs for the assessment of the interlaminar fracture toughness of the composite material systems used in their applications, because high toughness values are appreciated for an easier compliance with regulatory requirements. Moreover, such requirements must be met in all the situations, for the whole operative life. This obvious consideration points out the importance of evaluating the resistance of composite materials to delamination not only in standard laboratory environment, but also in extreme environments, rare but possible in the whole service life of the aircraft or helicopter. -2-
ACCEPTED MANUSCRIPT Typically, hot / wet situations are considered more demanding and so, with these motivations, an experimental program has been defined and carried out in the framework of a collaboration between the Department of Civil and Industrial Engineering of the University of Pisa and Leonardo Helicopter Division, formerly AgustaWestland.
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A relatively reduced number of papers has been published on the interesting problem, fundamental from the designer point of view, of characterizing the interlaminar fracture resistance of composites, keeping also hygrothermal influences into account. Limiting our attention to unidirectional
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carbon/epoxy material systems, it is worth mentioning the papers by Russell and Street, [1], one of the earliest to focus on such a problem for AS1/3501-6 and HMS/3501-6, more than thirty years
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ago. Indeed, test results reported in [1] showed for AS1/3501-6 a mild decrease in mode I nonprecracked toughness with increasing temperature (in those days, standard test methods were not yet defined), while the opposite effect was observed in the delamination growth phase. Béland et al., [2], also found that the critical toughness in AS1/3501-6 was decreasing with temperature, while
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in IM6/5245C the opposite trend was observed. This conflicting situation could be the result of the lack of common test standard procedures in those pioneers’ days; the important role played by the insert (and its thickness) was not yet underlined. Further experimental work, [3-5], carried out about
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ten years later, showed more clearly that an increase in mode I toughness is normally observed when the temperature and the moisture content increase. Asp [6] and Cowley [7], towards the end
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of the nineties, confirmed this tendency with their experimental evidence. A physical explanation of this phenomenon has been attributed to the fact that an increase in temperature and moisture content is associated with a more plastic behaviour of the matrix, and so the increased matrix ductility was considered, also on the basis of fractographic examination of failure surfaces, as the major responsible for the higher mode I resistance; a different failure mechanism should be active for mode II situations, as a typical tendency to a decrease of mode II toughness with temperature was observed (Russell [1], Asp [6], Kim [8] and Johnson [9]). This
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ACCEPTED MANUSCRIPT behaviour has been ascribed to a degradation of the fibre/matrix bond strength (confirmed by the observation at SEM of a higher number of bare fibres after a high temperature test). Davidson (2009, [10]) studied the influence of moisture content and of test temperature on interlaminar fracture toughness of a thermoplastic-particulate toughened carbon/epoxy system
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(T800H/3900-2). The results from Double Cantilever Beam (DCB), End Notched Flexure (ENF) and Single Leg Bending (SLB) specimens are in line with commonly observed trends: mode I toughness increases with moisture content and temperature, while mode II toughness decreases;
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intermediate behaviour was observed for mixed-mode situations.
As far as the influence of the moisture uptake is considered, its effect was found to be slightly
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detrimental, and sometimes even difficult to be appreciated. Hintikka et al. recently (2010, [11]) evaluated at room temperature (RT) specimens characterized by a different moisture content (dry, at normal ambient humidity and moisture saturated) showing that the effect is slightly appreciable in mode I (where an increase in ductility with increasing moisture content is observed), and almost
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negligible in mode II and mixed mode situations.
More recently, Landry et al. [12] has studied the effect of the contact with aerospace fluids, such as water, hydraulic fluid and deicing fluid; his experimental results from End Notch Flexure (ENF)
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specimens show the adverse influence of the exposure to various fluids on the fracture resistance. One of the most recent papers on the subject is by Davidson (2016, [13]), who has studied the
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elevated temperature effect on the mode I, mode II and mixed (I+II) mode toughness of a carbonpolyimide system: the trend is perfectly in line with the general finding of an increase in mode I toughness and a decrease of mode II toughness, with mixed mode situations being intermediate. In conclusion, the trend emerging from the literature is rather consistent (thermoplastic resins have been deliberately excluded from the review, as they are less sensitive to hygro-thermal effects), but nevertheless the aerospace industries have a considerable interest to investigate the influence of moisture and temperature conditions on the interlaminar delamination toughness of commonly used
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ACCEPTED MANUSCRIPT composite materials, for design and certification purposes. An experimental program has been carried out at the University of Pisa in collaboration with Leonardo HD for the quantitative assessment of such environmental conditions, and it will be described in the following paragraph. Experimental activity
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The material selected for this activity is Hexcel 913C-HTA-12K, a low temperature curing (125 °C) graphite/epoxy system, often used in many helicopter applications, such as rotor blades or hubs.
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The experimental activities consisted in mode I DCB tests and mode II ENF tests, carried out in
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Room Temperature Dry (RTD) and High Temperature Wet (HTW) conditions, and Mixed-Mode Bending (MMB) tests, performed only in RTD conditions. Nominal dimensions of the specimens were 25 mm width, 3 mm thickness (24 plies, at 0°) and 150 mm length. All the specimens were manufactured by Leonardo HD following the internal company specifications and procedures. The mode I tests were carried out following the ASTM D-5528 standard [14], applied both for the
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RT tests on dry specimens and for the HT tests on wet specimens. Aluminium hinges were bonded to the DCB and MMB specimens, by means of an epoxy adhesive, curing at RT. After bonding the hinges, one lateral side of the specimen was painted white, with an acrylic spray painting, in order
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to help the identification of the delamination tip position and the measurement of its distance from
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the hinge axis; to make easier this measurement, a set of marks was traced on the white paint, at regular intervals, as shown for instance in Figure 1. The tests were carried out in displacement control (1 mm/min displacement rate) and force and displacement were recorded throughout the test. Moreover, a camera was used to take pictures of the marked specimen side (1 picture every second) in order to identify the tip position during the delamination growth. The Mixed Mode Bending tests were carried out following the ASTM D-6671 procedure, [16]. Two mode partitions were studied, defined by a GIc /GIIc ratio equal to 2:1 and to 1:2 (or, as more often -5-
ACCEPTED MANUSCRIPT indicated in the literature, GII /Gtot = 0.33 and 0.66). The test data acquisition system, used for the DCB specimens, was also used in these cases: the delamination growth was documented by a number of pictures, collected at the end of the test, and by a file with the load and displacement of the actuator head.
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The mode II tests were carried out using the ENF specimen, loaded in three points bending, with a total span between the external rollers of 100 mm. At the time of the test execution, the relevant ASTM standard [15] had not yet been issued and so an internal Leonardo HD standard was
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followed; all the specimens were pre-cracked before being tested, by means of a three-pointbending fixture, with a shorter distance between the external rollers (total span of 50 mm); a 3-4
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mm natural delamination was introduced, from the insert. Then, the tests were carried out on precracked ENF specimens, using a 100 mm total span. For the wet specimens tested in high temperature, the compliance calibration method for data analysis (that has later been recommended in [15]) was applied; also in this case, a pre-cracking operation was previously executed at RT, in
Procedure for the analysis of the test data
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order to obtain a natural delamination front, originated from the implanted kapton insert.
In order to compare the various results obtained in the different types of tests, a consistent method
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for deriving the fracture toughness value is necessary. So, the three possibilities offered by the ASTM D-5528 ([14]), namely Modified Beam Theory, Compliance Calibration and Modified Compliance Calibration, cannot be used, as they are limited to the mode I test and rely on empirical correction parameters derived during the mode I test itself. The ASTM D-7905 ([15]), that rules the mode II testing, suggests the use of the Compliance Calibration method, that for the ENF specimen simply requires to measure the initial stiffness in correspondence with a number of delamination lengths, sliding the specimen with respect to the test fixture. In the case of the DCB specimen this
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ACCEPTED MANUSCRIPT option is obviously much more complicated, and the specimen compliance cannot be measured for various delamination lengths before the test, but only during the performance of the test itself. It has been therefore decided to use the Kinloch-Williams approach [17-18], adopted in the ASTM D-6671 standard ([16]), that introduces a numerically defined correction to the delamination length
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to explain the lower stiffness of the sub-laminate leg with respect to the beam theory predictions, based on the perfect clamp hypothesis. The delamination length correction involves the factor χ computed according to Equation ( 1 ): where Γ = 1.18
E11 E 22
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2 Γ 3 − 2 1 + Γ
G13
,
(1)
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E11 χ= 11G13
where E11 and E22 are the tensile elastic moduli in the longitudinal and transverse direction and G13 is the shear modulus in 1-3 out-of-plane direction.
Indeed, the correction was deduced in [17-18] on the basis of numerical studies carried out on
Results of the experimental program
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unidirectional graphite/epoxy specimens, and so it is well suited for being used in this similar case.
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The results obtained will be presented in the following, divided by type of test, together with comments and comparisons with literature data (if available). The tests on dry specimens will be presented first. 4.1
DCB tests
Fifteen DCB dry specimens (three batches of five samples each) were manufactured and were tested following the procedure described in ASTM D-5528 [14]. A preliminary pre-cracking was carried out, which allowed the measurement of the material’s bending stiffness. This measurement is particularly valuable, because it refers to an “ideal” condition: a straight delamination front (the end
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ACCEPTED MANUSCRIPT of the insert) and a delamination tip zone that is pristine, free from any damage and/or plastic deformations: these two events allow a reliable estimate of the flexural elastic modulus, a parameter useful for subsequent numerical analyses. The bending modulus of the Non-Pre-Cracked (NPC)
[16]): 3
(2)
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E1 f
t 64 m a + χ 2 = 3 Wt
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specimens has been calculated with the Kinloch-Williams correction to the delamination length (see
where m is the slope of the force-displacement plot, W and t are the specimen width and thickness;
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a is obviously the delamination length, while the correction parameter χ has been defined in Eq. (1). The average value of the bending modulus is 125037.1 MPa, with a standard deviation of 5658.8 MPa (so giving a coefficient of variation equal to 4.53 %).
For what concerns the material properties to be used in eq. (1), the following values were provided
1.72.
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by Leonardo HD: E1 = 137 GPa, E2 = 12.1 GPa, G13 = 5.98 GPa; with such inputs, χ results to be
After pre-cracking, the real test took place, starting from a “natural” delamination (PC). For every
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test it was possible to collect about 5-600 pictures, each one with the current value of the load and the displacement written on it. An example is shown in Figure 1. After the end of the test, the
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pictures were examined and, for a selected number of significant pictures, the relevant compliance was computed. It was therefore possible to calculate the various corrections, suggested by the ASTM D-5528, for keeping into consideration the lower stiffness of the two sub-laminate legs.
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Figure 1 – Example of picture of a DCB specimen (DCB-S-04) during the test. The marking lines spacing is 1 mm on the left (beginning of the propagation) and then 5 mm.
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Figure 2 – Example of determination of the delamination correction length in MBT analysis (specimen: DCB-S-04; C units: mm/N)
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Figure 2 shows an example of linear relationship between delamination length and cubic root of compliance (indeed, it is analogous to Fig. 4 of [14]); this plot allows the determination of a parameter, the crack length increment, to be used in the Modified Beam Theory (MBT) methodology for evaluating the GIc values. The ASTM D-5528 standard procedure ([14]) identifies also three different “critical” points for the calculation of GIc, namely the Non Linear (NL) point (end of the linear part of the force-displacement plot), the VIS point (first visible delamination growth) and the 5%/Max point (the point of maximum force or the intersection of the experimental graph with the 5% reduced stiffness line, whatever comes first). The analysis carried out with the -9-
ACCEPTED MANUSCRIPT MBT methodology and referred to the NL point normally provides the most conservative estimate among the nine total options available in the ASTM D-5528: in this case, the average value of the fifteen tests is 229.61 N/m (standard deviation: 15.93 N/m, CV: 6.9%). In [19] a small test activity was carried out on DCB of this same material and an average initiation
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fracture toughness of about 189 N/m was found, with a plateau of about 250 N/m at the end of the R-curve. Such results are the only fully comparable that the authors have found in the open literature and it is interesting to observe a reasonable agreement.
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Since in this paper comparisons are made between critical ERR values in different mode mix situations, it was considered appropriate to follow a common approach/methodology to evaluate G:
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the Kinloch-Williams Gc value has been calculated for all the specimens. Such an estimation requires first the assessment of the bending modulus; as already explained before, it is deduced from the slope m of the Force-Displacement plot by means of Equation ( 2 ). The average value of the bending modulus measured during the test of the fifteen PC specimens is
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112682.4 MPa, a value significantly lower than the corresponding modulus deduced from the precracking phase (NPC coupons), 125037.1 MPa. This difference can be attributable to two causes: (a) the natural delamination front is not straight any more, but slightly curved, with a longer length
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in the middle of the specimen, and so the dimension measured on the lateral side is an approximation by defect; (b) the plastic zone and the damage state at the tip contribute to a reduced
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constraint in the “clamp”, that induces a lower stiffness. Similar findings were already presented and discussed by the authors, [20], for other experimental results. The Strain Energy Release Rate value is calculated, for the NL point, by means of the Kinloch-Williams formula (see [16]): t 96 P a + χ 2 GI = 2 3 W t E1 f
2
2
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(3)
ACCEPTED MANUSCRIPT and the average Kinloch-Williams value for GIc is 243.98 N/m (standard deviation: 15.17 N/m, CV: 6.2%). These results confirm the reasons why, as already mentioned before, the MBT is widely used in deriving values for design applications: it gives conservative values, while the Kinloch-Williams
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methodology provides a 6 % higher value.
Figure 3 shows, for the 3 tested batches, the force vs displacement curves directly obtained from each of the RTD PC DCB tests (left part of Figure 3) and the distribution of the GIc values evaluated
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for each specimen according to the Kinloch-Williams data reduction method (right of Figure 3),
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referred to the NL point.
4.2
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Figure 3 – Force vs displacement (left) of the DCB RTD tests of the specimens belonging to the 3 batches; inter-laminar fracture toughness values (right) of the same specimens evaluated with the Kinloch-Williams data reduction method.
ENF tests
In the literature there is wide evidence that tests carried out on ENF specimens without any precracking operation (NPC, i.e. with the delamination starting at the insert) provide much higher fracture toughness values than in the case of tests carried out on pre-cracked (PC) ENF specimens. Just as an example, a reading of O’Brien’s paper [21] and/or [22] is recommended. Therefore, also - 11 -
ACCEPTED MANUSCRIPT in consideration that the aim of Leonardo HD was to evaluate the material performance in view of design applications, it was decided to pre-crack all the ENF specimens. To this end, the three-pointbending test fixture was modified with a reduced span between the external rollers (namely 50 mm total span), and mounting the ENF so to leave a distance of about 4-5 mm between the insert end
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and the central loading nose. The coupon was then loaded in displacement control, until a load drop was observed, typically associated with a sharp noise and a contemporary visual crack growth, up to the arrest position under the loading nose, or a bit farther. The same type of PC operation was
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carried out both on dry and on wet ENF specimens.
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Figure 4 - Force vs displacement (left) of the ENF RTD tests of the 2 batches; inter-laminar fracture toughness values (right) computed with the Kinloch formulation and the 5%/Max criterion.
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Two batches of five ENF dry specimens each were available. For each test, a load-displacement plot was obtained (e.g. see left part of Figure 4 for the dry specimens), with a large linear part (after a short initial non-linear phase, where the contact is established between the specimen and the loading nose), with evidence of non-linearity only in the very last part (just before the load drop). The linear part of the plot was used to deduce the flexural elastic modulus of the material; this operation was made considering the physical delamination length measurable visually on the lateral surface of the ENF. The Kinloch-Williams formulation was used to deduce the GIIc values, which are shown in graphical form in Fig. 4, right. These GIIc values have been calculated using the - 12 -
ACCEPTED MANUSCRIPT maximum force; this has been considered a better option than the NL point, because in the wet specimens a larger non-linear phase (see Figure 5) was observed in the final phases of the test. Therefore, the comparison of HTW and RTD ENF results on the basis of the NL point would be penalizing for the HTW data, because onset of non-linearity in wet specimens is not an indication of
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imminent failure.
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Figure 5 – Force-displacement plot of HTW test on specimen ENF-Sw5.
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The formula used to compute E1f is, according to [16]:
E1 f
3 3 t m 2 L + 3 a + 0.42 χ 2 = W t3
(4)
where m is the stiffness of the specimen, W is the width and t the thickness; GII was calculated by means of the following relationship (see [16]):
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ACCEPTED MANUSCRIPT t 9 P a + 0.42 χ 2 GII = 2 3 2W t E1 f
2
2
(5)
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All the ten dry specimens available for the RT tests were tested setting the delamination tip at a = 25 mm, with a half span equal to 50 mm.
The results show an average bending modulus equal to 118671.5 MPa (standard deviation: 3800.8 MPa, CV: 3.2%), a value 5% higher than 112682.4 MPa, that was the average for the pre-cracked
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DCB specimens, while they should be fully comparable. The two situations differ for the fact that
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the DCB legs deform separating from each other, while in ENF specimens the two sub-laminates are pressed one against the other and, moreover, tend to slide one against the other, with the separation surface composed by two different areas: a natural delamination (about 4-5 mm in length), plus an area covered by the kapton insert (about 20 mm long). So it is expected that the DCB estimated modulus be “true”, while the ENF deduced modulus may be affected by the sliding
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movement that, through friction, is perceived as an additional stiffness. Moreover, another consideration must be made with reference to the delamination length used in
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equations (2) and (4), that was based on the Kinloch-Williams correction applied to the physical length, estimated visually on the lateral surface of the specimen. In DCB specimens, the GI
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distribution in the specimen width is approximately parabolic, with its maximum in the centre, while in ENF specimens GII is distributed in the opposite way, with maximum on the lateral surfaces and the minimum in the centre. The delamination front resembles the G distribution. Therefore, the visual estimation provides an under-estimate of the “effective” length for the DCB, and an over-estimate for the ENF. So the bending stiffness for the ENF, being referred to a slightly longer delamination length than the “effective”, results a bit higher; the opposite occurs for the DCB specimens.
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ACCEPTED MANUSCRIPT Unfortunately, values of E1f measured from ENF specimens in the NPC state are not available. Further analyses have been carried out to evaluate the possible influence of friction in ENF tests.
4.3
MMB tests
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Interlaminar fracture toughness was evaluated only on dry specimens, at RT. Two groups of specimens were tested, at two mode partition values: GII /Gtot = 0.33 and 0.66. Five samples were available for each mode partition, but only four specimens produced valid results, because of a
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hardware failure in one case and of a strange, singular delamination behaviour (bifurcation) in another case. The recommendations of the ASTM D-6671 procedure [16] were followed. Also in
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this case, the delamination growth was monitored by capturing pictures, and so at the end of the test a number of photographs were available, for the study of the propagation phase (R-curve, that anyhow is not presented in the paper). Figure 6 shows displacement vs. force plots, with evidence
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of linearity for a long part of the recorded trace.
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Figure 6 – Force vs displacement (left) of the MMB RTD tests; inter-laminar fracture toughness values (right) computed with the Kinloch data reduction and the VIS criterion. The Strain Energy Release Rate components in mode I and mode II can be computed by means of
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the equations reported in [16]:
9 P (3 c − L ) a + χ GI = 2W 2 t 3 E1 f
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2
2
t 2
2
t 2 9 P 2 (c + L ) a + 0.42 χ 2 GII = 2 3 2W t E1 f
(6) 2
(7)
where E1f can be deduced from the slope of the linear part of the force-displacement plot. Indeed, the ASTM recommended procedure allows the determination of the delamination initiation toughness (two options are available: NL and 5%/Max) or the study of the propagation (to derive an R-curve, when the propagation takes place in a stable way). The first value is a Non-Pre-Crack - 16 -
ACCEPTED MANUSCRIPT (NPC) value, derived from the insert, and so susceptible to criticism, because it is widely recognized that NPC mode II toughness values are significantly (and artificially) higher than those obtained in tests on Pre-Cracked (PC) specimens; a similar consideration applies also to mode I, with minor quantitative effects. Therefore, for the comparison with pure modes data, it was decided
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to use the MMB results from the VIS point, i.e. the point in the test when the delamination growth from the insert has been observed visually. In the tests with prevailing mode I conditions (GII /Gtot = 0.33), the propagation was very stable and the VIS situation was characterized by a delamination
very quick, with a few millimetres of sudden growth.
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growth of a few tenths of mm; in the tests with prevailing mode II loading, the propagation was
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It is important to note that the Gc values, shown as column plot in Fig. 6, have been derived by making use of E1f, estimated in NPC conditions, but applied in a PC condition (VIS). In the elaboration of the DCB and ENF specimens, on the contrary, the bending modulus E1f was estimated in the PC condition and combined with data (force and displacement) from the test, that
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was also carried out in PC condition.
All the results have been plotted in Figure 7, together with a Benzeggagh-Kenane [23] best-fit (8):
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the global appearance is in line with similar results available in the literature.
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Gc = G Ic + (G IIc − G Ic )(G II / GT )η
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(8)
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5
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Figure 7 – Average and standard deviation (error bars) of the values of the inter-laminar fracture toughness obtained for pure and mixed modes; the dashed line represents the B-K curve that fits the experimental data.
High temperature testing of conditioned specimens
Five DCB and five ENF specimens were conditioned by Leonardo HD in a climatic chamber at 70
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°C and 85% RH up to saturation. Then, the specimens were hermetically packed in sealed envelopes for the shipment to Pisa, where they have been maintained in such package until the day
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of the test. Hinges were cold bonded to DCB specimens by means of an epoxy adhesive, curing at RT; the day after, the specimens were ready and their sides were painted white and marked with ticks. The preparation of ENF coupons was simpler, obviously, requiring only the precracking operation, performed in shear, at RT and marking on the edges. The static tests were performed in a thermally insulated bakelite box, where hot air was insufflated by an industrial heater. Each specimen was instrumented with thermocouples (2 for each DCB, 3 for each ENF) and the transient in temperature from RT to 70 °C took about 25 minutes, with an average heating rate of about 2 °C per minute. No humidity control was performed during the test. - 18 -
ACCEPTED MANUSCRIPT The bakelite box has two large quartz glass windows, so that it was possible to continue to use the same test apparatus already used in the RT tests, capturing and storing pictures of the delamination tip position. Therefore, the DCB tests were carried out following the recommendations of [14], with a precracking phase, followed by the actual test: the data collected allowed to estimate the
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parameters required for correcting the Beam Theory and Compliance Calibration approaches. As far as the ENF specimens are concerned, they were all subjected to precracking, in RT conditions (similarly to the dry specimens), and then to a Compliance Calibration campaign, also at RT. The
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compliance C was measured for the following delamination lengths: 18, 20, 23, 25, 27, 29 mm; for
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completeness, also a measure of the bending stiffness in absence of delamination was performed.
Figure 8 – Example of the CC methodology on a HTW specimen (ENF-Sw4).
All these measurements at RT allowed to collect data for tracing a C vs. ‘a cube’ plot for each specimen; a typical example is shown in Fig. 8. The least squares best-fit linear equation provides the derivative of C with respect to ‘a’, and so the following formula could be used for G:
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ACCEPTED MANUSCRIPT GII =
P 2 ∂C 2W ∂a
(9)
and, with the approximation that C = ma 3 + C0 it follows that:
∂C = 3ma2 ∂a
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( 10 )
This is the CC method, recommended in the ASTM standard [15], and its results were compared
6.1
Comments on the results of the Hot-Wet tests
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with those from the Kinloch-Williams formula (eqn. 5).
Mode I tests
Differently from the analogous RT tests on dry specimens, a more pronounced bridging effect is observed, that is reflected in an appreciable G growth with delamination length; this phenomenon
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has obviously consequences on the semi-empirical parameters that are derived from the test to be used in the MBT, CC and MCC approaches. The results, in terms of Modified Beam Theory,
point):
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relevant to the NL point, are reported in Table II below (P and δ are the coordinates of the NL
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Specimen
DCB-Sw1 DCB-Sw2 DCB-Sw3 DCB-Sw4 DCB-Sw5
P (N) 54.94 52.47 52.47 52.22 50.82
δ (mm) 7.01 6.54 8.40 8.09 7.00
∆ (mm) 6.94 3.21 8.06 6.25 5.84
GIc (N/m) 374.67 354.35 384.03 414.12 345.59
Table I – Results of the HTW tests on DCB specimens (MBT methodology, NL point). The average value is 374.5 N/m (standard deviation: 26.9 N/m; CV: 7.2%), significantly higher than the average RTD value, 229.61 N/m; indeed, the behaviour of the wet specimens in the HT tests seemed clearly more ductile, with a pronounced plastic behaviour. The results have been calculated - 20 -
ACCEPTED MANUSCRIPT also by means of the Kinloch-Williams approach (average: 398.8 N/m; standard deviation: 35.2
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N/m; CV: 8.9%), and are shown in Figure 9 (right).
Figure 9 – Force vs displacement (left) of the DCB HW tests; inter-laminar fracture toughness values (right) computed with the Kinloch-Williams data reduction and the NL criterion.
The remarkable differences, obtained between the HT tests and the RT tests in terms of fracture
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toughness growth with crack length (R-curves), are shown in Figure 10. It turns out that not only the average fracture toughness increases significantly (1.5 times greater), but the bridging effect as well. In fact, while the average slope of the R-curves of the RT tests is essentially negligible, for the
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HT tests the growth of GIc with the crack length is appreciable, though affected by scatter.
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Figure 10 – R-curves obtained for the RT tests (specimens 1-5) and for the HT tests (specimens 1, 2, 4 and 5).
The Kinloch-Williams analysis allows also the comparison of the bending modulus estimated in the
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Non-Pre-Cracked (NPC) state, i.e. with the insert, with the same quantity in the pre-cracked condition (PC), i.e. with a natural delamination. The average value in the HTW NPC specimens is 125651.9 MPa, almost the same value observed in the RTD specimens, 125037.1 MPa: this may be
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ascribed to the fact that the modulus is a fibre dominated property. As a matter of fact, the difference between the RTD and HTW conditions is mainly in the dimensional dilatation of the
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matrix, that slightly increases the thickness, and so the bending stiffness; indeed, the difference is negligible, and can be estimated in less than 1% on the modulus. In the PC test, after pre-cracking, the delamination front is no more perfectly straight, and the measurement on the specimen side edge underestimates the effective value; moreover, damage is present close the tip. The bending stiffness deduced from the PC test data decreases with respect to the NPC value: in the HTW tests the average bending modulus is 119802.9 MPa, while in the RTD tests it dropped down to 112682.4 MPa. The difference between the two situations is quite significant; it can be attributed to the more
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ACCEPTED MANUSCRIPT plastic behaviour of the matrix, with reduced brittleness and therefore also a smaller damage in the tip region after pre-cracking.
6.2
Mode II tests
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The ENF wet specimens, before being tested in high temperature, had been previously subjected to a pre-cracking operation at RT, in shear. After pre-cracking and before the test, for each specimen the compliance vs. delamination length relationship was evaluated, in order to apply the
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Compliance Calibration (CC) methodology for estimating the Strain Energy Release Rate. Such compliance measurements were made at RT, and resulted in a quite consistent, low scatter results,
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as already shown in Figure 8. Moreover, also an experimental assessment of the bending modulus was carried out, shifting the specimen on the rollers in such a way that the loaded part was free from the delamination; the experimental value has been compared with C0, the extrapolated value, from the C = ma 3 + C0 relationship: a very good agreement has been obtained for each specimen.
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Then the real test took place, with a delamination length equal to 25 mm (i.e. the tip was in the middle of the span); the data collected allowed to estimate G, but it was also possible to assess the difference in modulus, ascribed only to the temperature change. The HTW interlaminar fracture
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toughness was estimated by means of the Kinloch-Williams approach and by the CC approach, with
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the ∂C/∂a derivative estimated at RT; this may be not correct. The results are summarized in Table II: the Kinloch-Williams approach gives an average value of 708.4 N/m, which is significantly lower than the RTD value, that was 1047.9 N/m. The CC approach gives a GIIc average value of 640.9 N/m, not so different from the K-W estimate (7% lower). It is interesting to observe the bending modulus measurements: the average value, derived from specimens with a delamination length equal to 25 mm, is 122531.8 MPa for the HTW tests, while the RT analogous value from the dry specimens was 118671.5 MPa. This last value is very consistent with the measurements made on the “un-cracked” ENF wet specimens at RT, i.e. the wet specimens placed in the experimental
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ACCEPTED MANUSCRIPT set-up in such a way to exclude the insert from the rollers span, which showed an average value of 118896 MPa, substantially identical to the RTD value. Table II shows also the directly measured quantities, i.e. the slope of the force-displacement plot, with an average increase of the slope of about 1.5% passing from RT to HT conditions. This is a strange result: the specimens show a slight
systematic, worth to be studied in depth.
Slope RT (N/mm) 291.19 328.05 325.69 339.79 316.08
E1f PC-KW (MPa) 120173.9 124851.3 120415.4 126859.3 120359.0
GIIc – KW (N/m) 642.3 734.4 711.5 741.4 712.2
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ENF-Sw1 ENF-Sw2 ENF-Sw3 ENF-Sw4 ENF-Sw5
Slope HT (N/mm) 295.65 331.30 328.82 347.52 321.46
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Specimen
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increase in stiffness while passing from RT to HT conditions. It is not a big difference, but
GIIc – CC (N/m) 657.0 602.0 630.4 625.5 689.4
Table II – Results of the HTW tests on ENF specimens. It should be underlined that the K-W analysis methodology was developed by means of numerical analysis that were fitted to RT tests; it is obviously straightforward to extend its applicability and
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pretend a reasonable accuracy in HT situations, where the failure mechanisms are sensibly different. Figure 11 shows the force-displacement plot of the HTW ENF tests, together with the toughness,
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with the K-W analysis methodology, while Figure 12 collects all the fracture toughness results.
Figure 11 – Force vs displacement (left) of the ENF HW tests; inter-laminar fracture toughness values (right) computed with the Kinloch-Williams data reduction and the NL criterion.
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Figure 12 – Comparison between the values of inter-laminar fracture toughness obtained in RTD conditions and the one obtained with HTW conditions for the pure modes.
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The temperature increase, in the case of the ENF coupons, seems to have some influence on the measured modulus, due to the occurrence of some mechanism not present in the DCB coupons, where the bending modulus has not shown substantial differences between the RTD and the HTW
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results.
RT and HT friction measurement
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To investigate this aspect, a measurement of the friction coefficient between the delaminated parts of a specimen was organized.
A simple inclined plane device has been designed and realized, as shown in Figure 13. The plane is hinged and can rotate, and is initially placed in horizontal position. Two pieces of delaminated coupons are placed on top of the plane, with a small weight on the upper piece. By applying a slow movement to the rotating plate, it has been possible to measure the angle, α, at the beginning of the upper portion shift, that is related to µ, the coefficient of friction, by means of the simple
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ACCEPTED MANUSCRIPT relationship: µ = tan(α ) . It has been considered appropriate to apply a small pressure between the two portions, and so a weight of about 1 N was fixed over the upper part. The dimensions of the specimens’ portions were 25 mm in width (obviously, as they were cut from the wedge opened
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specimens) and 30 mm in length.
Figure 13 – Friction measurement apparatus used in the friction tests.
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Moreover, a number of different situations have been examined, that differed one from the other for the extension of the kapton film layer, placed between the two pieces of delaminated coupons, so referring to different situations (delamination lengths) during the ENF test. In total, four different
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situations have been investigated, namely: a) no kapton film; b) 5 mm kapton film; c) 15 mm kapton film; d) 30 mm kapton film. It is obvious that the b) situation is representative of the test,
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where the precracking operation had introduced about 5 mm of “natural” delamination. For each of the four different situations, two specimens have been prepared and tested for three replicas at RT and other three replicas at HT (70 °C), to obtain information on the friction coefficient variability in function of the temperature, and also an assessment of the scatter. The scatter observed is high, particularly for the tests b) and c). Possible causes may be: a) the presence of a non-straight delamination front, which disturbed the measurement; b) the high roughness of the surface, generated by wedge opening: no flattening action or wear due to fatigue was present. Anyhow, a tendency was observed, namely an increase in the coefficient of friction - 26 -
ACCEPTED MANUSCRIPT with a lower extent of kapton at the interface: at RT, the average friction coefficient passed from 0.273 in the all kapton interface, to 0.306 for the 15mm/15mm interface, to 0.384 for the no kapton condition. The friction coefficient was influenced by the temperature, but in a contradictory manner: it passed from an average value of 0.384 at RT to 0.510 at HT for the all fibre interface and
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from an average value of 0.273 to 0.241 for the full kapton interface. The decrease of the coefficient of friction for the case with kapton interface at HT could seem puzzling, but is consistent with other literature results [24]. Notwithstanding this confused situation, a numerical study on the influence
Numerical evaluations
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7
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of friction coefficient was carried out, taking the move from classic works, like [25] and [26].
A 2D FE model of a 3-mm thick ENF specimen has been developed in the Abaqus software in order to evaluate the friction effects on the flexural modulus computed according to Eqn. ( 4 ). To reproduce the ENF test, the specimen was sustained by means of 2 cylindrical rigid supports,
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with a span of 100 mm, and loaded by a central nose. Boundary conditions were assigned to the 2 supports through their reference points which were constrained to not translate and rotate. The central cylinder was loaded in displacement control, and the horizontal translation was inhibited,
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such as the in-plane rotation.
The specimen was modelled as two half-laminates tied in the un-cracked region while contact
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interactions were present along the delamination length. The total delaminated length, measured from the first support, was 30 mm: 25 mm were covered by the kapton insert while 5 mm was the length of the natural crack. The geometry of the model, as shown in Figure 14, represents the condition of a pre-cracked specimen. The FE model also presents part of the specimen outside of the left support so that the total delamination length, measured from the specimen left end, is 55 mm (25 mm externally to the support span and 30 mm internally).
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ACCEPTED MANUSCRIPT The two half-laminates were meshed with plane stress elements with a maximum mesh size of 0.5 mm along the specimen length (in the contact regions) and a very fine mesh size, 0.04 mm, around
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the delamination tip. In the thickness, 25 elements for each half-laminate were used.
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Figure 14 – Regions interested by the contact interactions (upper part of the figure) and details of the 2D numerical model of the ENF. To evaluate the effects of the friction contributions in the two regions, 2 tangential contact
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properties have been created to characterize the contact interactions assigned to the kapton insert region and to the natural delamination. Indeed, 2 coefficients have been assigned to each of the
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regions: for the tangential contact property in the kapton insert region, values of the friction coefficient, µt, of 0.2 and 0.3 were used, while, in the crack region, values of 0.3 and 0.4 were assigned to the friction coefficient, µc. Then, 2 non-linear static analyses have been performed by using the lowest and the highest friction coefficient pairs for the 2 contact regions, an arbitrary choice for assessing the influence of a high and a low friction value. Figure 15 shows the distribution of the compression stress and of the shear stresses due to friction, that are generated along the contact line of the half-laminates at the end of both simulations. The
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ACCEPTED MANUSCRIPT results show that the 2 half-laminates are pressed one against the other only in the area of the left support and, consequently, friction forces are generated only in such area. This result is consistent with the stress distribution reported in [25], where the effect of friction on the mode II fracture
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toughness values was investigated.
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Figure 15 – Compression stress and shear stresses along the crack length for the simulations performed with different friction coefficients. According to the results obtained with the simulations, the increase of flexural modulus in the HT
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tests is not fully justified only by the friction effects. Anyhow, another contribution can be ascribed to the transverse thermal expansion coefficient, αT, of the composite system. If a coefficient of
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thermal expansion of about 60 x 10-6 (°C)-1 is considered (typical value for an epoxy resin), a 50 °C temperature increase induces a thermal strain of 0.3% and so the bending stiffness, proportional to the third power of the thickness, increases of about 0.9%, spacing apart the stiff carbon fibres. The more “sticky” behaviour of the new, natural delamination length has been studied in classic literature papers (like [25] and [26]) assessing the influence of the friction coefficient: the influence is limited to the areas where significant pressure is applied between the two sub-laminates; in total, only a few percent increase in apparent toughness (around 2%) can be ascribed to this energy absorbing mechanism. In other words, in HTW conditions the true GIIc value reduces slightly - 29 -
ACCEPTED MANUSCRIPT (around 2%) with respect to the value observed experimentally. Anyhow, since the differences are really small, the influence in aviation design procedures is typically widely covered by the safety factors normally adopted.
Conclusions
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During the present investigation, fracture mechanics tests were carried out on dry and wet specimens, made of Hexcel 913C-HTA; a group of tests were performed at RT on dry specimens
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and another group on wet specimens, at HT. It was not possible to point out the separate influence on interlaminar fracture toughness of moisture and high temperature. Anyhow, the literature
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available on the subject (e.g. [10-12]) clearly underlines that the temperature is the most influencing parameter, with deeper effects, while moisture content has a lower impact. With respect to similar activities performed in the past and available in the literature, this activity has the merits of assessing the sensitivity to environmental conditions of the interlaminar fracture
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toughness of a low temperature curing (125 °C) epoxy resin. The HTW conditions of the tests are therefore relatively more severe than for the case of a 180 °C curing traditional epoxy. Anyhow, the variations observed with respect to the RTD performance are not quantitatively much different with
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respect to composites with traditional epoxy matrix. Nevertheless, it is always important to evaluate quantitatively such aspects, by means of dedicated tests: no theory can predict reliably the
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environmental influence on the interlaminar fracture toughness. The experimental activity has left a number of open questions, topics on which it is important to concentrate future efforts. The first one concerns the way of processing the data collected during the tests: are the rules typically developed and set-up for the evaluation of interlaminar fracture toughness at RT still applicable in a HT situation, where constitutive laws slightly change (a wider non-linearity range is expected)? The widely diffused MBT approach for mode I tests analysis requires the determination of a semi-empirical correction to the delamination length: the hypothesis has been adopted here the experimental procedures set-up for performing tests at RT could be - 30 -
ACCEPTED MANUSCRIPT applied also in HT and that the semi-empirical correction parameter keeps its meaning. The average Delta value (see Figure 2) is 6.06 mm for the HTW DCB specimens, a bit higher than the average RTD value of 5.63 mm; this is according to the expectations, because it is a measure of how far the real situation is from the ideal perfect clamp constraint. In HTW conditions, a less rigid constraint is
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expected. Moreover, the modification of the Beam Theory proposed by Williams and Kinloch [1718] is still meaningful? The ideal clamp constraint is certainly modified by the hygro-thermal conditions; further work is planned on this point.
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Anyhow, once that HT toughness values have been estimated, the results obtained in this experimental activity confirm that the mode I toughness increases with temperature, with a more
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plastic and ductile behaviour of the resin system, while a reduction in the mode II toughness is observed. This tendency is confirmed in the literature by fractographic examination of the failure modes, which gives evidence of a more plastic behaviour in mode I and a fibre/resin bonding interface of reduced strength in mode II tests.
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In mode II tests a discrepancy about the compliance of the ENF specimens has been observed between the RT values (during the CC evaluation) and at HT values, during the test performance: the stiffness of the specimens has shown a small increase with temperature. This point stimulated
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therefore a more detailed investigation, numerically and experimentally. A short campaign for the measurements of the friction coefficient between the two sub-laminates has been carried out,
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showing a slightly higher friction coefficient at high temperature, though in a context of measurements affected by large scatter. The numerical model of the experiment did not explain completely, with such input, the experimental observation; the concurrent presence of another factor, associated with the thermal expansion of the resin, can reduce the gap. Anyhow, the basic work carried out in the early years of the definition of the ENF specimens has found a clear confirmation. Further work is planned on this point, as well as on the questions arose before, on the data reduction methodology.
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[2] S. Beland, J.P. Komorovski, C. Roy, Hygrothermal influence on the interlaminar fracture energy of graphite/bismaleimide modified epoxy composite (IM6/5245C). In Proceedings of the 6th
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ACCEPTED MANUSCRIPT [14] ASTM D-5528-01, Standard test method for mode I interlaminar fracture toughness of unidirectional fiber-reinforced polymer matrix composites. American Society for Testing and Materials.
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