Healing of low-velocity impact damage in vascularised composites

Composites: Part A 44 (2013) 78–85

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Healing of low-velocity impact damage in vascularised composites C.J. Norris, I.P. Bond, R.S. Trask ⇑ Advanced Composites Centre for Innovation and Science (ACCIS), Department of Aerospace Engineering, University of Bristol, Queen’s Buildings, University Walk, Bristol BS8 1TR, UK

a r t i c l e

i n f o

Article history: Received 22 March 2012 Received in revised form 18 August 2012 Accepted 25 August 2012 Available online 7 September 2012 Keywords: A. Carbon fibre B. Impact behaviour D. Mechanical testing Self-healing

a b s t r a c t This research has sought to characterise damage formation and self-healing efficiency within vascularised carbon fibre reinforced polymer (CFRP) laminates over a range of low velocity impact energies. Using ultrasonic C-scanning and compression after impact (CAI) analysis, vascularised laminates were shown to conform to the same damage size to residual compression strength relationship established for conventional laminates. The damage tolerance level of the host laminate was carefully determined, an important consideration in selection of the most appropriate vascule spacing for a reliable self-healing system. The healing functionality imparted full recovery of post impact compression strength over the range of impact energies tested (2.5–20 J), via healing of matrix cracking and delamination damage. The successful implementation of this technology could substantially enhance the integrity, reliability and robustness of composite structures, whilst offering benefits through reduced operational costs and extended lifetimes. However, establishing the benefits of such novel systems to existing design criteria is challenging, suggesting that bespoke design tools will be required to fully attain the potential benefits of self-healing technologies. Ó 2012 Elsevier Ltd. All rights reserved.

1. Introduction In addition to their superior specific strength and stiffness, fibre reinforced polymer (FRP) composites have been shown to have excellent resistance to fatigue cracking and immunity to corrosion [1]. Damage resistance is therefore, one of the primary concerns for the aerospace industry, due to the poor out of plane properties of these materials. The mechanisms involved in damage formation, and the resultant damage morphology, alters significantly depending on the velocity of the impactor. During high velocity impacts, characterised in the range of 300–2500 ms1 [2], the event is so short that the structure has no time to respond in global flexural or shear modes and so damage tends to be localised, taking the form of target penetration. During low velocity impacts (generally regarded to be at velocities less than 10 ms1), the impactor contact time is long enough for the entire FRP structure to respond to the impact [3]. The out of plane displacement generates flexural and shear stresses which, in the main, lead to cracking of the matrix material, although fracture of fibres can also occur. The resultant intra-ply microcracking and inter-ply delaminations can be significant without the presence of visual surface evidence of the impact event itself; a situation termed barely visible impact damage (BVID) [2–5]. BVID resulting from low-velocity impact damage is the most insidious and poses a considerable challenge in its ability

⇑ Corresponding author. Tel.: +44 (0) 117 331 5845; fax: +44 (0) 117 954 5666. E-mail address: [email protected] (R.S. Trask). 1359-835X/$ - see front matter Ó 2012 Elsevier Ltd. All rights reserved. http://dx.doi.org/10.1016/j.compositesa.2012.08.022

to significantly reduce mechanical performance without being readily detectable.

2. Strategies to address BVID To account for this shortfall in performance, the aerospace industry relies on damage resistant designs, damage tolerant designs, non-destructive evaluations (NDEs), structural health monitoring (SHM) and repairs to maintain air worthiness. Impact damage resistance addresses the resistance of a structure to damage formation for a given level of impact [6]. A number of approaches have been adopted to achieve this, such as increasing matrix toughness [7], reducing the overall damage footprint by maximising the number of interfaces available for dissipating impact energy [8], Z-pinning [9] and interleaving [10]. Damage tolerance is the ability of a structure to contain representative weakening defects under typical loading and environmental conditions without jeopardizing aircraft safety, for some specified period of service [11]. This approach is based on the philosophy that only detectable or immediately obvious damage lowers a structure’s residual strength below design allowables [12–14]. This philosophy is provided in schematic form in Fig. 1. Here, the size of damage that lowers the residual strength of the structure to the limit stress (LS), the maximum stress expected during service conditions, is termed the critical design threshold (CDT). An ultimate safety factor above the LS is commonly used in aerospace applications, commonly equal to 1.5, so that any structure is designed to

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Fig. 1. Schematic representation of the damage tolerance design philosophy [13].

withstand this higher ultimate stress (US). The ADL (allowable damage limit) equates to the maximum damage size that can be tolerated whilst sustaining the design ultimate stress [12]. A high probability of detecting the ADL size with the selected detection technique is critical. Damage smaller than the ADL is undetectable, whilst damage greater than CDT is considered to be readily detectable or immediately obvious. It is also evident that a further safety margin exists between the value of design ultimate stress and the actual ultimate strength (undamaged) of the material, again contributing to overweight inefficient composite structures [13]. Based on the ‘no growth’ approach applied to composite structures within the aerospace industry, any damage equal to or greater than the allowable damage limit (ADL) must be rectified immediately. Damage is detected during routine NDE or via SHM, enabling decisions to be made regarding ongoing performance, maintenance requirements and safety thresholds [15]. Structures must be repaired to the state that all static loads, including ageing and environmental effects, can be sustained until the aircraft is removed from service. Applying external patches offers a fast and effective repair, which can be mechanically fastened [16] or adhesively bonded [17]. To prevent crack propagation, the damaged zone is generally removed before the patch is fixed in place. Laminates thicker than 2–3 mm generally carry too much load for external patch repairs and so scarf repairs are often utilised [11]. The key benefit to a scarf repair is the relatively uniform adhesive shear stress distribution along the bonded area, providing increased load-bearing capability. The damaged material is removed prior to replacement with virgin material that closely matches the properties of the parent [18]. Tapered angles of 20:1–60:1 are often required; therefore, a large volume of undamaged material must be removed, especially in thick laminates [19]. An alternate repair approach has been reported, whereby matrix cracks and delamination are infused with a resin via holes drilled in the damaged zone [20–22]. Recovery of compression, tensile and flexural strengths were all demonstrated within these studies, along with ultrasonic C-scan images showing successful infusion of matrix damage. Liu et al. stated that although successful infusion of matrix damage was achieved, patches are also required to account for the loss in strength from fractured fibres [21]. The ability to effectively repair delaminations that exceed a critical size, without the requirement to remove large volumes of material, is advantageous and cost effective. Self-healing has emerged as a candidate for addressing the issues of BVID in composite systems. This bioinspired concept involves the inclusion of secondary functional materials capable of counteracting service degradation whilst still achieving the primary, usually structural, requirements of composite components [23]. The self-healing approaches devised for these materials can be summarised into three categories, (i) solid-state or liquid resin

79

delivery via (ii) compartmentalised vessels or (iii) vascular networks. Solid state (also referred to as intrinsic) self-healing systems have been designed whereby the matrix material has been formulated to enable crack healing upon exposure to an external stimulus (e.g. heat). Development has taken the form of the synthesis of new polymeric matrices or the incorporation of a secondary solid polymer phase within an existing thermosetting matrix material [24,25]. The incorporation of discrete liquid resin filled vessels within FRP materials, in the form of microcapsules [26] or hollow glass fibres (HGFs) [27], has also been extensively studied. Upon exposure to a damage event, these vessels fracture and ‘bleed’ a mobile liquid phase which is reliant on capillary forces for transportation to and subsequent autonomous healing of micro-cracking and delaminations. More recently, embedded vascular networks have been successfully utilised to deliver a pressurised healing agent from an external reservoir to regions of internal damage within bulk epoxy matrices [28], the core structures of composite sandwich panels [29] and laminated FRP coupons [30,31]. Literature on the ability of these self-healing systems to heal impact damage in FRPs is limited, with the majority of studies focusing on fracture geometries. Yin et al. [32] and Patel et al. [33] utilised the compression after impact (CAI) testing protocol to evaluate the recovery of post impact compression strength in woven glass/epoxy laminates through healing via microencapsulated healing agents. Both studies highlighted a reduction in healing efficiency with increased impact energy as a result of the limited volume of healing agent available. Williams et al. [34] assessed the ability of liquid resin filled hollow glass fibres (HGFs) to self-repair low-velocity impact damage using a CAI test protocol. HGF’s were embedded within the 0° plies as these coincided with the longest dimension of the test specimen (thereby maximising healing volume) in a [45/90/45/0]2S CFRP laminate. Although 84% recovery of the undamaged strength was reported after a 2.2 J impact, it is suspected that the potential volume of embedded healing agent may be insufficient to provide similar levels of strength restoration at higher impact energies. The ability to connect embedded vasculatures to external reservoirs potentially eliminates the issues of healing agent availability, with high levels of post-impact strength recovery (P94%) and damage infusion demonstrated in CFRP laminates [30,31]. Damage tolerance is a well established and reliable design approach in the aerospace industry, but as a consequence, structures tend to be significantly over designed to carry normal operating loads in the undamaged state. It is the expectation, from a selfhealing perspective, that reliable preservation of the original undamaged strength could lead to lightweight, efficient structural components that challenge the existing design philosophy. For self-healing to be considered as a viable option for structural FRP components, it must be determined at what level of damage the healing functionality is no longer effective. Lower levels of impact induced damage may not produce the connectivity between delaminations and through the thickness microcracks that such healing systems rely on for full damage infusion, whereas healing efficiencies for high energy impacts are likely to be influenced by the levels of fibre fracture induced. This current study, therefore, focuses on the ability of a vascularised laminate to provide a healing functionality over a range of damage severities, thereby providing an effective ‘working range’ for the system.

3. Materials and testing 3.1. Vascularised laminate manufacture An aerospace grade carbon fibre reinforced epoxy (Hexcel Composites IM7/8552) unidirectional pre-impregnated tape, with a

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nominal fibre volume fraction of 57.7% and very low matrix flow during processing (as specified by the manufacturer’s data) was selected for this study. The laminate stacking sequence [45/90/45/ 0]2S is a preferred lay-up for improved damage tolerance, with two 0° plies located at the centreline. All samples were cured to the manufacturer’s recommendations, 120 °C for 1 h, followed by 180 °C for 2 h and a pressure of 700 kPa. Segments of uncured pre-impregnated tape were removed during lay-up to accommodate 0.5 mm diameter PTFE coated steel wire vascule preforms. This methodology required the removal of material from the central 45/0//0/45 section of the laminate (based on the nominal ply thickness of 0.127 mm). The material to be removed was cut parallel to the 0° fibre direction to prevent termination of the primary load bearing fibres. Steel wires were spaced at 10 mm intervals across the width of the laminate and manually removed post-cure to create open channels. The chosen vascule fabrication route, vascule diameter and location were selected on the basis of minimising disruption to the host’s fibre architecture (see Fig. 2) whilst maximising vascule-damage connectivity [35–38]. A summary of the mechanical performance of this vascule configuration shows minimal influence on the undamaged laminate compression strength and improved Mode I and II fracture toughness, Fig. 3. However, damage size is slightly increased with respect to the control, with an accompanying penalty in damaged compression strength. Vascule diameter must be carefully considered as a trade-off exists between mechanical performance and ease of fluid flow, with narrower 0.25 mm diameter vascules having reduced influence on damage formation but requiring 4 the pumping pressure compared to 0.5 mm vascules. 3.2. Damage and healing test protocols An Instron Dynatup 9250HV drop weight tower fitted with a 20 mm radius hemispherical tup was used to generate low velocity impact damage within specimens, clamped using an impact support fixture with a 75  125 mm cut-out as detailed in ASTM D7136 [39]. Vascularised coupons (absent of infused healing resin) were subjected to a range of impact events (2.5, 5, 10, 15 and 20 J) before assessment via NDE ultrasonic C-scan testing (USL SAM 350 fitted with a NDT UPR receiver and a Panametrics V311 10 MHz/ 0.500 transducer). A series of CAI tests were performed to determine the residual compressive strength of laminates containing the various levels of impact damage. Specimens were placed into an antibuckling fixture based on the specifications detailed in the ASTM standard [40], but modified to allow 89  55 mm panels to be tested as outlined by Prichard and Hogg [41]. The compression testing was conducted on an Instron 1342 test frame fitted with a 250 kN load cell. Specimens were loaded under displacement

-45 90 45 0 -45 90 45 0 0 45 90 -45 0 45 90 -45 Fig. 2. Vascule location and morphology. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)

control at a rate of 0.4 mm/min until failure occurred, with the data recorded via an Instron 8800 controller/data-logger. A second series of coupons were subjected to impact events of 5, 10, 15 and 20 J prior to delivery of a low viscosity two-part epoxy healing resin via the embedded vascular network. A 12v D.C. motor driven peristaltic pump (Williamson Pumps Ltd.) delivered the pre-mixed healing agent (ResinTech RT151) from an external reservoir (flow rate of 0.8 ml/min) to a vascule located in the damage zone for a set time period of 5 min. The RT151 epoxy resin was selected on the basis of the following desirable properties:    

Low-viscosity (mixed viscosity of 0.1 Pa s at 23 °C). Compatibility with FRP matrix. One hour working life after mixing. Rapid ambient cure, 16–24 h at 23 °C.

In all cases, the specimens were left at ambient temperature for a minimum of 24 h to allow cure of the healing agent. These specimens were ultrasonically C-scanned both pre- and post-healing for comparative purposes. Healed compression strength, calculated as a percentage of undamaged strength, was used to derive values of healing efficiency. The number of specimens tested for each condition is provided in Table 1.

4. Impact damage assessment Ultrasonic C-scan analysis was conducted to capture the level of damage introduced to the vascularised laminates, with representative time of flight (TOF) images provided in Fig. 4. TOF provides information regarding the through thickness location of matrix damage, with dark blue highlighting delaminations near the front face and red indicating delaminations toward the laminate back face. Over the range of impact energies tested, it can be seen that the extent of damage induced varies significantly, from no detectable damage at 2.5 J to severe matrix damage in the 20 J case. Severe local fibre fracture and surface indentation was observed visually under the point of impact for the 15 and 20 J cases; this damage corresponds to the black areas visible in the C-scan images. These areas appear ‘black’ due to the depth of impactor indentation, which can be seen to almost match the dimensions of the impactor in the 20 J case. Average indentation depths of 0, 0.05, 0.2 and 0.75 mm were recorded for 5, 10, 15 and 20 J impact events, respectively. It is arguable that the extent of damage imparted by the 20 J impact event goes beyond that defined as BVID. The largest individual delamination occurs at the back face in all cases (red), along matrix tensile cracks induced by bending. The separated sub-laminate created is regarded as being too thin to have a significant effect on the post-impact compression strength and buckling induced delamination growth [2]. However, the level of internal damage is critical, with a number of researchers having shown an excellent correlation (when global buckling is suppressed) between damage width and residual compression strength [5,41,42]. This correlation is attributed to the tendency for damage to propagate transverse to the loading direction in the CAI environment. Therefore, Image J (image analysis software) was used to identify the area of the internal and back face delamination footprints separately, Fig. 5. This data shows that the area of the back face delamination grows significantly with increasing impact energy, and to a greater extent in comparison to the internal damage footprint, which appears to increase linearly with impact energy. The relationship between residual compression strength of the vascularised laminate and damage width (defined as the maximum internal damage width excluding the back face delamination) is provided in Fig. 6. This relationship closely follows that previously

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Mode I Fracture Toughness*

Mode II Fracture Toughness*

Damaged (10J) Footprint†

Control Vascularised

CAI (10J) Damaged Strength†

CAI Undamaged Strength†

Fig. 3. Overall mechanical performance envelope for the selected vascule configuration in comparison to an equivalent plain laminate, as relative percentage ( [37]   [38]). (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)

5. Healing assessment

Table 1 Number of specimens tested for each condition. Impact energy Velocity (ms1) CAI (damaged) CAI (healed)

0 0 6 /

2.5 0.84 2 /

5 1.21 3 3

10 1.74 5 5

15 2.13 3 3

20 2.46 4 3

established for conventional non-vascularised laminates [5,41,42]. A situation expected, as the vascule configuration utilised in this study was found to have an undamaged strength within 3% of the corresponding control laminate [38]. All undamaged specimens and those impacted at 2.5 J, failed at the load introduction points by end brooming. This is a well documented failure mode for undamaged specimens tested in the CAI regime [41,43]. Specimens within the 5 J data set failed by either end brooming or central buckling and collapse. A damage width of approximately 14.9 mm causes a shift in failure mode to central buckling and collapse and consequent drop-off in residual strength and can, therefore, be considered as a critical threshold (for this test scenario). Damage below this critical threshold can be tolerated by the laminate, suggesting that a vascule spacing of 14.9 mm would be sufficient to provide a robust healing system (note that the current vascule spacing was set at 10 mm). Here, this equates to a vascule density of approximately 3% of the specimen width. Identifying the minimum vascular network density required to prevent the existence of critical sized delaminations is an important design consideration, as fewer vascules have reduced potential for disrupting the host’s fibre architecture and provides faster, less complicated fabrication. Beyond this critical damage threshold all specimens fail by central buckling and collapse, with excellent correlation between measured damage width and residual compression strength. This supports the use of internal damage width as the key parameter in predicting residual compression strength and that the largest individual back face delamination has no significant influence on this parameter. It also suggests, for this testing regime at least, that fibre fracture has little influence on the residual post-impact compression strength, in agreement with previous studies that have shown failure mechanisms induced during CAI testing are matrix dominated [5,41].

Pumped delivery of the RT151 epoxy resin healing agent to the impact damage, via the vascules, was found to significantly reduce the damage footprints, Fig. 7. The C-scan TOF images of the ‘healed’ specimens are evidence of large-scale damage infusion, resulting from the internal delamination, microcrack and vascule connectivity, as highlighted in previous self-healing evaluations [30,31,38]. Resin bleed from back face tensile cracks was observed when infusing 10, 15 and 20 J impact damage, and from front face perforation in the 20 J case, proving that sufficient damage connectivity exists for resin flow from the centrally located vascules to the laminate surface. The back face delamination, being open at the surface, is the hardest to infuse and so persists to a greater extent in comparison to the internal delaminations. Interestingly, it appears that infusion occurs to a greater extent for 15 J and 20 J impacts, likely to be the result of increased damage connectivity and larger crack opening displacements resulting from these higher energy events. It should be noted that indentation depths were unaffected by the healing cycle. Although some level of damage remained in all specimens, the average healed internal damage width (excluding the back face delamination) was reduced to below 15 mm for all impact energies, Fig. 8. This level of damage is within the tolerance zone of the vascularised laminate, previously identified in Fig. 5, suggesting that these specimens would be stabilized and fail by end brooming during the CAI testing. This was true for the majority of samples, however, there were specimens within the 10, 15 and 20 J sets that failed by central buckling and collapse. Those specimens failing by end brooming were found to sustain a similar ultimate load to the undamaged laminate, Fig. 9. Healed specimens failing by central buckling and collapse failed at significantly higher loads to those tested in the damaged state and, therefore, do not appear to follow the same relationship between damage width and compression strength described in Fig. 6. The following mechanisms for this increased post damage load bearing capability have been identified:  The ability of the healing agent to accumulate at a delamination crack tip, leading to an effective method of crack blunting, has

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1800

mm

Internal + backface breakout

1600

Damage Area, mm2

20J

Internal damage footprint

1400 1200 1000 800 600 400

Front face perforation – severe level of local fibre fracture

200 0 2.5

5

10

15

20

Impact Energy, J

*Internal damage width

15J

Fig. 5. C-scan damage footprint areas with the back face breakout delamination segregated from the internal damage footprint. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)

450

10J

Compression Strength, MPa

Back face breakout

Central failure

350 300 250 200

y = -4.81x + 359.69 2 R = 0.97

150 100

Damage tolerance zone

50

5J

End brooming

y = 1.04x + 369.11 2 R = 0.13

400

Damage requiring attention

0 0

5

10

15

20

25

30

35

40

Damage Width, mm Front face

Back face

µsec Fig. 4. C-scan TOF analysis of damage footprints at varying levels of impact energy.  Maximum internal damage width excluding the back face delamination.

previously been highlighted [30]. The work of fracture of such blunted cracks will be significantly higher compared to those where cracks propagate from much sharper features typically found in delaminations.  Interleaving, i.e. increasing the depth of the resin layer between the plies in a laminate, leads to an increase in Mixed (I/II) Mode [44] and Mode II [45] fracture toughness. Infiltration of 50 lm depth delaminations with an epoxy resin will result in resin layers of the same thickness between the plies and thus impart a further toughening mechanism to the healed composite.  The toughness of the room temperature cured epoxy resin is likely to be lower in comparison to the laminate matrix, cured under optimum conditions (180 °C).  Infusion of the impact damage effectively creates a blister, increasing the thickness and, therefore, the stability of this region under compression loading. This increase (typically 0.1–0.25 mm) was not accounted for in the residual compression strength data as the aim was to evaluate the ability of a healed structure to behave as if undamaged.

Fig. 6. The relationship between residual compression strength and damage width in vascularised laminates. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)

6. Discussion The CAI assessment highlighted that the vascularised laminate has a similar damage size to residual compression strength relationship found for conventional laminates. However, this relationship alters significantly when the laminate is assessed in the healed state, Fig. 10. Infusion of the damage zone, over the range of impact energies tested, restores the compression strength of the laminate to within 97% of the undamaged strength. This is the first proof that a vascularised self-healing laminate can offer a step-change in damage resistance over a wide range of lowvelocity impact events. The ability to maintain the virgin strength of the laminate has a number of implications when considered in conjunction with the damage tolerance design philosophy currently used in safety critical applications. Fig. 11a displays the damaged and healed CAI data superimposed on the schematic of the design tolerance philosophy (see Fig. 1) as a function of impact energy. The correlation between impact energy (over the range tested herein) and residual compression strength follows a very similar trend to that of damage width and impact energy, justifying the change in x-axis parameter. The maximum level of damage inflicted (at 20 J) was easily detectable visually and appears to fall just short of the CDT but

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450

mm

End brooming 400

Compression Strength, MPa

20J

Central failure End brooming 'healed'

350

Central failure healed

300 250

y = -4.81x + 359.69 2 R = 0.97

200 150 100 50 0

15J

0

5

10

15

20

25

30

35

40

Damage Width, mm Fig. 9. The relationship between residual compression strength and damage width in the healed specimens. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)

10J

Compression strength, MPa

450

5J

Back face µsec

Front face

400 350 300 250 200 150 100 Damaged state

50

Healed state

Fig. 7. Typical ultrasonic C-scan TOF response from healed specimens.

0 0

2

4

6

8

10

12

14

16

18

20

Impact Energy, J

45 40

Fig. 10. The relationship between residual compression strength and impact energy for the vascularised laminate in the damaged and healed states. Note: the grey bar is indicative of the standard deviation around the undamaged strength. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)

Damaged Healed

Width, mm

35 30 25 20 15 10

Damage tolerance zone

5 0 2.5

5

10

15

20

Impact Energy, J Fig. 8. Average internal damage width – post impact and post healing cycle. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)

in the region regarded as readily detectable damage by Razi and Ward [12]. Here, the ADL corresponds to 250 MPa in the damaged state, whereas in the healed state the laminate can sustain a load comparable to the undamaged state. This improved damage resistance could be exploited in the design of a thinner and lighter component. Alternatively, this could provide a drastic reduction in the required inspection frequency as the ADL now occurs at a much higher level of impact. The exact damage size of the healed ADL

is yet to be determined as it falls out of the range of damage size achievable in the CAI test utilised. The healed CAI compression strength data, superimposed on the schematic of the design tolerance philosophy, is provided in Fig. 11b. Here, the x-axis is maintained as a measure of damage width. Over the range of impact energies tested, the healed damage width ranges from approximately 6–15 mm in comparison to 14– 36 mm for the damaged state. Therefore, over the range of impacts tested, damage does not exist between 15 and 36 mm (assuming the healing system is 100% reliable). The upper end of this damage resistant region is again unknown, as further testing of larger geometry sub-elements will be required to gather data for damage sizes above those achievable with CAI. This upper limit will correspond to the effective working range of the self-healing system, and is an obvious requirement for the design of components containing such additional healing functionality. Low levels of damage will exist in a vascularised self-healing laminate, resulting from damage events that do not breach a vascule (dictated by spacing) or small levels of matrix damage that persist post healing cycle. These smaller damage zones could be mitigated by the incorporation of a supplementary healing component, such as healing agent filled microcapsules, to provide a more robust self-healing system.

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450 Healed Damaged

Compression strength, MPa

400 350

New CDT

New ADL

300 US

250

New ADL and CDT

200 LS

150 100 50

CDT

ADL

(a)

0 0

5

10

15

20

Impact Energy, J 450

relate this information to residual compression strength. In addition, the ability to deliver a healing agent to regions of impact induced damage via an embedded vasculature, thereby restoring a proportion of the undamaged material properties, was assessed. From this work, the following conclusions can be drawn:  The CAI and NDE evaluations highlight that vascularised laminates follow the same damage size to residual compression strength relationship as found for conventional laminates.  The ability of the laminate to tolerate a certain damage size was clearly demonstrated. A damage size greater than 14.9 mm was found to cause a shift in failure mode from end brooming to central buckling and collapse, with an accompanying significant decrease in residual compression strength. This data is crucial for the design of self-healing laminates/components as it provides a quantifiable metric for vascule spacing.  The healing functionality imparted full recovery of post impact compression strength over the range of impact energies tested, via healing of matrix cracking and delamination damage.

Compression Strength, MPa

400 350 Damage does not occur in this size region

300 250 200 LS

150 100 50

CDT

(b)

0 0

10

20

30

40

Damage Width, mm Fig. 11. Self-healing in relation to the damage tolerant design philosophy; (a) Demonstration of how the ADL and CDT levels can be influenced by reliable strength recovery and (b) showing a zone of damage sizes that do not occur in the healed vascularised laminate. (For interpretation of the references to colour in this figure legend, the reader is referred to the web version of this article.)

Although the schematic representations outlined in Fig. 11 offer some indication of the benefits of self-healing for FRP design, it also highlights the difficulty of using such multifunctional materials within an existing design philosophy. Undoubtedly, the successful implementation of a lightweight, efficient and reliable self-healing component or structure will be dependent on a new set of design tools which account for a self-healing function from the outset. To assist the development of a new design tool, the effectiveness of the self-healing system needs to be evaluated at the sub-element level to explore and quantify an ability to heal larger damage events. In this scenario, large-scale fibre fracture may influence the efficiency of any healing system as the current technology targets healing of matrix damage only. Significant effort is also required in the development of improved self-healing chemistries that provide fast and repeated healing. The ability to mimic the biological process of ‘haemostasis’, so that the healing agent only polymerises in damaged zones and not within the vascules themselves, is highly desirable and is a current focus of research effort. 7. Concluding remarks This research has sought to characterise damage formation over a range of impact energies within vascularised FRP laminates, and

The successful implementation of this technology could substantially enhance the integrity, reliability and robustness of composite structures, whilst offering benefits through reduced operational costs and extended lifetimes. However, establishing the benefits of such novel systems to existing design criteria is challenging, suggesting that bespoke design tools will be required to fully attain the potential benefits of self-healing technologies. In addition, simply adding such self-healing functionality to a conventionally designed laminate that has already been optimised for damage tolerance is unlikely to yield components with any appreciable enhancement in performance or weight saving. As it stands, the above findings do not justify the incorporation of self-healing in real-life structural, safety critical components as they are currently designed. However, the results strongly suggest that fast, reliable, repeated self-healing can offer a step-change in material performance, thereby justifying the investment required to evaluate such systems beyond the coupon level, and to develop tailored healing chemistries. Acknowledgements The authors would like to thank the UK Engineering and Physical Sciences Research Council and UK Ministry of Defence via the Defence Science and Technology Laboratory for funding this work under CRASHCOMPS (EP/G003599), Airbus UK for their additional financial support. References [1] Chen P, Shen Z, Wang J-Y. Damage tolerance analysis of cracked stiffened composite panels. J Compos Mater 2000;35:1815–43. [2] Davies GAO, Olsson R. Impact on composite structures. Aeronaut J 2004;108:541–63. [3] Richardson MOW, Wisheart MJ. Review of low-velocity impact properties of composite materials. Composites A 1996;27:1123–31. [4] Cantwell WJ, Morton J. The impact resistance of composite materials – a review. Composites 1991;22:347–62. [5] Hull D, Shi YB. Damage mechanism characterization in composite damage tolerance investigations. Compo Struct 1993;23:99–120. [6] Olsson R. Analytical prediction of large mass impact damage in composite laminates. Composites A 2001;32:1207–15. [7] Jordan WM, Bradley WL, Moulton RJ. Relating resin mechanical properties to composite delamination fracture toughness. Compos Mater 1989;23:923–43. [8] Gonzalez EV, Maimi P, Camanho PP, Lopes CS, Blanco N. Effects of ply clustering in laminated composite plates under low-velocity impact loading. Compos Sci Technol 2011;71:805–17. [9] Mouritz AP. Review of Z-pinned composite laminates. Composites A 2007;38:2383–97. [10] Greenhalgh E, Hiley M. The assessment of novel materials and processes for the impact tolerant design of stiffened composite aerospace structures. Composites A 2003;34:151–61.

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