A significant opportunity for reducing the weight of composite aircraft is through the development of an economically efficient method to detect barely visible or invisible impact damage sustained in service. In this paper, a structurally integrated, inert, wireless system for rapid, large-area impact damage detection in composite is demonstrated. Large-area inspection from single sensors using ultrasonic-guided waves is achieved with a baseline-subtraction technique. The wireless interface uses electromagnetic coupling between coils in the embedded sensor and inspection wand. Compact encapsulated sensor units are built and successfully embedded into composite panels at manufacture. Chirp-based excitation is used to enable single-shot measurements with high signal-to-random-noise ratio to be obtained. Signal processing to compensate for variability in inspection wand alignment is developed and shown to be necessary to obtain adequate baseline subtraction performance for damage detection. Results from sensors embedded in both glass fibre and carbon fibre-reinforced composite panels are presented. Successful detection of a 10 J impact damage in the former is demonstrated at a range of 125 mm. Quantitative extrapolation of this result suggests that the same level of impact damage would be detectable at a range of up to 1000 mm with an inspection wand alignment tolerance of 4 mm.
Composite materials are susceptible to barely visible impact damage (BVID) caused by low-velocity impacts (energy level of 10–30 J) such as tool drops, bird strikes and hailstones [1,2]. By definition, BVID is hard to detect through routine visual inspection. Although ultrasonic bulk wave and thermal imaging non-destructive testing (NDT) techniques are able to detect BVID, neither are ideal for performing rapid in-service inspections. For example, ultrasonic inspection requires water coupling and the ultrasonic transducer must be scanned over the complete surface to be inspected. Thermography is an inherently large-area technique, but due to the low thermal diffusivity of composite, thermographic methods are primarily limited to near-surface damage .
The current approach adopted by most of the aerospace industry is to over-design composite structures so that they can tolerate the existence of BVID  (more serious impact damage is assumed to be detectable during routine visual inspections between flights). Consequently, there is a real opportunity for reducing the weight of composite aircraft by developing an efficient method of detecting BVID between flights. The potential weight reductions would also help to meet the new environmental targets set for the aerospace industry . Conceptually, a fully integrated structural health monitoring (SHM) system using permanently installed sensors is the ideal solution to provide rapid BVID detection. Such a system could either detect impacts directly (e.g. via acoustic emission) or the damage resulting from an impact [6–9]. Such SHM systems have been widely demonstrated in the laboratory and in some cases trialled on actual aircraft. However, there are two major challenges preventing the widespread deployment of such SHM systems in the aerospace industry.
The first challenge is the fundamental problem of making the necessary physical measurements with suitable levels of reliability, sensitivity and selectivity. For large-area monitoring with a limited number of sensors, ultrasonic-guided waves provide arguably the only generally applicable physical measurement mechanism. In a structure with any degree of complexity, some form of baseline signal comparison approach becomes necessary when guided waves are used in order to discriminate between benign structural features and damage. Significant progress continues to be made in this area [10–12], with increasingly robust detection sensitivity being demonstrated even in the presence of environmental variability.
The second challenge is the more practical problem of data and power connectivity to a distributed sensor network in a large aerospace structure. The easiest connectivity solution is to use wires and a number of strategies for ultrasonic composite SHM based on wired transducers have been investigated [13,14]. The main limitation of a wired system is that a large number of conductive routes and also transducer control units such as multiplexers need to be integrated into the structure. For a large structure, the amount of wiring and electronics not only increases weight, manufacturing and installation costs , but also potentially decreases the reliability of the system, owing to the possibility of failures in the wiring. A widely researched alternative is to use wireless RF protocols (e.g. Wifi, ZigBee) for communication [16,17], but the major challenge for wireless systems is power consumption. If the power source is a battery or other energy storage device then when the power source is exhausted, replacement presents a significant problem if the sensors are embedded or integrated into a structure such as a composite aircraft . An alternative solution is to use power-harvesting technology such as photovoltaic, thermoelectric, piezoelectric devices and rectenna arrays for the wireless unit. However, in order to provide enough power for sensing electronics and data communication, current power-harvesting requires complex, bulky devices .
In this paper, a rapid, impact-damage detection solution is proposed that falls between conventional NDT and fully integrated SHM. Rather than attempting to embed sensors, power sources and data communication into the structure, only the sensing elements are embedded and measurements are made using an external device. The advantage of this approach over a fully integrated SHM system is that the sensors are inert when not in use and no energy storage or long-range data transfer need be considered. The potential advantages over conventional NDT (where the sensors as well as instrumentation are separate to the structure) is (i) the ease and speed with which measurements can be made and (ii) the fact that sensors remain at precisely fixed positions in the structure, enabling baseline signal comparison techniques to be employed if required. The concept of the proposed system for impact detection in a composite structure is shown in figure 1a. A distributed array of inert sensors is permanently embedded in the composite structure and they are probed using an external device termed the inspection wand. The inspection wand could be deployed either manually, as shown in the figure, or using a platform such as a unmanned aerial vehicle. When the inspection wand is in close physical proximity to a sensor, the sensor is activated and a measurement is made. A key attraction of the system is that it enables quantitative measurements to be rapidly obtained from sensors at precisely defined locations using an inspection wand that does not have to be precisely positioned or even physically coupled to the structure.
An inductively coupled transducer system (ICTS) was initially proposed by Greve et al. [19,20], and inductive coupling is employed in the current paper to realize the concept shown in figure 1a. The same technique has also been used by Wu et al. [21,22] for thorough-thickness inspection with flexible transducers. The proposed ICTS is shown in figure 1b and consists of three coils. One coil (termed the transducer coil) is physically connected to a piezoelectric transducer and forms a sensor unit, which is embedded in the structure. The other two coils (termed the transmitting and receiving coils) are in the separate inspection wand, where they are connected to the output and input channels of ultrasonic instrumentation. Importantly, a completely embedded sensor benefits from the physical protection provided by the host structure and also, if manufactured correctly, exhibits superior ultrasonic coupling to the structure compared with a surface-mounted sensor.
Previously, a model of an ICTS parametrized by physical constants, such as the number of turns and diameter of each coil, was developed and validated . In the current paper, this electromagnetic coupling model was used to develop the optimized design for guided wave detection of impact damage described in §2b. A process to practically embed an optimized sensor into composite is demonstrated in §2c. In order to enable single-shot measurements with high signal-to-random-noise ratio and to compensate for the effect of wand misalignment, the linear shift between the axes of the coils, several signal processing techniques are described in §2d. The consistency of embedded sensors, the performance of misalignment compensation methods and finally the use of the system to detect impact damage are reported in §3.
2. System design
(a) Methodology and objectives
The inspection target for the system described in this paper is to detect the damage from a 10 J low-velocity impact in a composite panel at distance of up to 1000 mm from a sensor location [1,2]. Because the intended use of the proposed system is for rapid routine manual inspection, it was decided that the inspection target should be achievable in a single-shot measurement with a realistic degree of inspection wand misalignment relative to the sensor.
Guided elastic waves (Lamb waves) have been widely acknowledged as one of the most promising tools for long range detection of various types of defect in different materials and structures , and deployable-guided wave systems have been successfully used for industrial NDT inspections of simple structures such as pipes [25–27] and rails [28,29]. Because a guided wave sensor can potentially detect damage over a significant surrounding area, a permanent sparse array of guided wave sensors in pitch-catch mode is an ideal candidate for large-area SHM. In this paper, a sparse array of guided wave sensors, each operating independently in pulse-echo mode, is used as the basis of the composite impact damage detection system. The damage in the structure is detected and localized by identifying the appearance of new reflected signals in the pulse-echo mode instead of new diffracted signals in pitch-catch mode. The pulse-echo configuration is more practical for use in this particular application, requiring only a single inspection wand to be aligned with a single sensor to perform a measurement. To achieve a reasonable inspection range and sensitivity to damage, the fundamental extensional (S0) mode in the 0.35–0.5 MHz mm frequency thickness product range is used. The motivation for this is that at higher frequency thickness products, higher order modes exist making signals harder to interpret, while at lower frequency thickness products, the sensitivity to localized defects is diminished. The attenuation of the S0 mode due to material damping is also significantly less than the fundamental flexural (A0) mode  and the S0 mode is also less susceptible to attenuation due to energy leakage if the structure has liquid-loading (as is the case in, e.g. wet-wing designs). Owing to the complexity of aerospace composite structure, it is assumed that the system will operate in a baseline subtraction mode to suppress reflections from benign structural features, such edges and stiffeners [31,32].
The embedded sensor unit should be of minimal size, especially thickness, to reduce its impact on structural integrity. For the reasons outlined above, it is necessary to use the S0 Lamb wave mode, which requires a transducer that is sensitive to in-plane motion. As complete structural coverage in terms of defect detection is desired, the transducers should have omni-directional sensitivity to this mode. Therefore, thin, disc-shaped piezoelectric transducers operating primarily in an axisymmetric radial mode are used (specifically 16 mm diameter, 0.3 mm thick devices manufactured from NCE51 piezoelectric material supplied by Noliac Group, Kvistgaard, Denmark).
It is worth noting that the electromagnetic coupling and ultrasonic measurement are considered as independent challenges in this paper. This allows well-understood electromagnetic modelling  and ultrasonic-guided wave SHM experience  to be adopted separately in the design process.
(b) Electromagnetic design
Figure 2a shows a typical ICTS pulse-echo time-domain response . Cross-talk owing to direct coupling between the transmitting and receiving coils gives rise to a dead-zone in the coverage provided by a sensor unit, as the cross-talk obscures signals from nearby reflectors, as indicated in figure 2b. Inductively coupled systems are susceptible to incoherent electromagnetic interference (EMI) from other circuits causing incoherent or random noise. The signal prior to the cross-talk in figure 2a gives an indication of the level of such noise. Obtaining an adequate signal-to-random-noise ratio, SR, is an important consideration for the design of the ICTS. However, simply designing to maximize the amplitude of the received signal will result in a resonant circuit, which is how most inductive power transfer systems are designed [36,37]. In fact, signal fidelity is of equal if not greater concern in the design of ICTS than SR. Resonance is manifested as long signal ring-down which has several undesirable effects, one of which is poor temporal resolution. More importantly, resonance increases the duration of cross-talk and, consequently, the size of the dead-zone around each sensor. To quantify the signal fidelity of an ICTS response from a sensor on a particular structure, the signal-to-cross-talk ratio, SX, of a received signal, v(t), is defined as 2.1 where indicates a Hilbert transform, t1 is the arrival time of a reference reflection signal and δ is the duration of the input signal used to drive the ICTS whose peak is assumed to start at t=0. The numerator in this expression is signal amplitude and the denominator is the mean cross-talk signal amplitude in the period, yellow region in figure 2, after the input signal has finished. The latter would be theoretically zero if the ICTS had no resonance and is consequently a measure of signal fidelity.
In practice, it has been found that an ICTS designed to maximize SX also achieves an acceptable SR (i.e. one that is not a limit on system performance). Therefore, SR is not considered in the electromagnetic design process, but will be discussed further in the context of signal processing.
As noted, the electromagnetic design and ultrasonic measurements are considered as independent challenges in this paper. The objective of the electromagnetic design here is therefore to optimize the electromagnetic coupling aspect of an ICTS for a specified piezoelectric transducer and ultrasonic instrument. In other words, the electromagnetic optimization is performed on the basis that the specified piezoelectric transducer would yield a suitable ultrasonic signal if it were physically wired to an appropriate ultrasonic pulser–receiver. Here, a combined digital-oscilloscope signal-generator unit (HS3, TiePie Engineering, Sneek, The Netherlands) is used as the ultrasonic instrumentation. An optimization procedure is applied to electromagnetic design to maximize SX for the ICTS for a specified input signal (a 5-cycle Gaussian windowed toneburst at a centre frequency of 160 kHz). The predefined constants of the system such as output impedance of signal generator, impedance of digital oscilloscope and manufacturing limitations are listed in table 1a along with system information such as thickness of printed circuit board (PCB) substrate and input signal.
In order to reduce the physical size of the embedded sensor, the secondary coil in the sensor is a double-sided planar coil fabricated on a PCB that surrounds the piezoelectric disc. Therefore, the inner diameter of this coil is constrained to the diameter of the piezoelectric disc plus 4 mm radial clearance for connection soldering, thus giving a minimum inner coil diameter of 24 mm. Also, in order to decrease the size of all coils and the number of design parameters, the PCB manufacturing resolution, k, is used to define both the PCB track width as well as the distance between the two neighbouring tracks. Because the transmitting and receiving coils are built into opposite side of a single PCB, these coils are single sided.
The specifications of the optimized three-coil design based on the constraints given before are listed in table 1b. Figure 3 shows SR, SX of the optimized system at different reading distances. It can be seen that although SX does decrease with reading distance, the decrease is fairly modest in comparison with SR. SR drops exponentially and approaches zero as the reading distance increases. For this optimized system, the maximum reading distance is 40 mm, where SR is equal to 2 (6 dB). It worth noting that the experimentally measured noise level was used to calculate SR.
(c) Embedded sensor design
Figure 4a shows the process developed to manufacture an encapsulated sensor unit for embedding in a composite structure. First, the piezoelectric disc is placed in a circular cut-out in the centre of a double-sided PCB containing the secondary coil with dimensions determined from the optimization procedure (24 mm inner diameter, 38.4 mm outer diameter, double sided with 19 turns on each side). Then, the assembled unit is cured with 0.075 mm thick insulation layers (Pyralux LF; Dupont, Wilmington, DE, USA) on both sides. The insulation layers consist of polyimide film coated on both sides with acrylic adhesive. The result is an encapsulated sensor unit in the form of a 0.45 mm thick, 55 mm diameter disc, which is ready for embedding into a composite material.
Sensor units were embedded into a glass fibre composite panel (Hexcel 950; Hexcel, Stamford, CT, USA) cured at 125°C and a carbon fibre composite panel (Hexcel M21; Hexcel, Stamford, CT, USA) cured at 180°C with 7 bar pressure applied. The glass fibre and carbon fibre composite panel lay-up sequences are shown in figure 4b,c, respectively, and the finished panels are shown in figure 5. The glass fibre panel is used for impact damage detection tests and the carbon fibre panel for sensor performance studies. The embedded ICTS can be visually identified in the glass fibre panel, but their presence in the carbon fibre material is only detectable via a slight thickening on the non-tool side. In the carbon fibre plate there are three ICTSs and one wired unit for reference. The effect of embedding on the strength of composite structure is open research topics, but has been studied by Yang et al. . It has been found that embedding a sensor with polyimide bond does not decrease the strength of the composite structure.
(d) Signal processing
In this section, the signal processing associated with the embedded ICTS is described. The signal processing is required for two purposes: to allow single-shot measurements with high signal-to-random-noise ratio and to allow successful baseline signal subtraction despite signal variability owing to inspection wand misalignment.
(i) Excitation and pre-processing
Let u(t) be the desired input signal to the ICTS, where t is time. This is typically a short, limited-bandwidth toneburst of duration T. In order to enable single-shot measurements with high signal-to-random-noise ratio, the system response, v(t), to u(t) is recovered from the system response, vc(t), to a narrowband chirped input signal, uc(t), of duration NcT. In the following, and are the forward and inverse Fourier transform operators and ω is angular frequency. The necessary chirp excitation is given by 2.2 where is the frequency spectrum of the desired input signal and is the frequency spectrum of a linear chirp signal, c(t), defined over the interval 0≤t≤NcT given by 2.3 where ω1 and ω2 are the start and end frequencies of the chirp. The bandwidth of c(t) (i.e. from ω1 to ω2) must span the bandwidth of u(t). The stretch factor, Nc, specifies the duration of the chirp signal relative to the duration of the desired input signal. The increase in SR associated with a chirp is proportional to . The frequency domain response of the system, V (ω), to the desired input signal is extracted from the system response to the chirp excitation, vc(t), through the following frequency domain deconvolution augmented with a bandpass filter, G(ω), to remove out-of-band noise 2.4
Thus, the final time-domain response of the system to the desired input signal is 2.5
(ii) Effect of coil misalignment
In an infinitely large, featureless structure, impact damage could be directly detected by the appearance of a reflection in v(t). However, in a real structure, large signals are present owing to reflections from benign structural features and these must be suppressed in order to be able to detect much smaller signals owing to impact damage. The detection of new impact damage therefore requires comparison of the current signal with an earlier baseline signal. Here, baseline signal subtraction is employed as this has the advantage of being a linear operation with predictable performance . It has been found that the amplitude of a received signal from an ICTS varies dramatically with changes in the relative position of coils [19,23], which here is due to imprecise positioning of the inspection wand between repeated measurements at the same sensor. This presents a major challenge for baseline signal subtraction that must be overcome in order to achieve reliable damage detection.
Figure 6a shows waveforms prior to baseline subtraction from the ICTS embedded in glass fibre plate for coil separations of 12 and 17 mm. The coil separation is defined as the distance between the plane of the sensor and the plane of inspection wand under the assumption of perfect coaxial and angular alignment. The effective input signal was a Gaussian-windowed, 5-cycle (defined by the −40 dB points of the Gaussian window) toneburst with centre frequency 160 kHz and the results were obtained using chirp excitation with Nc=1300. The change in amplitude of the first echo signal for a range of coil separations is shown in figure 6b. As the coil separation increases, the signal amplitude drops and decreases by about 65% of its original amplitude when the coil separation is increased by 8 mm.
However, amplitude variation is not the only effect resulting from coil misalignment. The corresponding spectra for the first echo signals shown in figure 6a are shown in figure 6c. There is a small change in the frequency of the spectral peak for the two coil separations and this is plotted for a range of separation distances in figure 6d. Similar results are found for lateral misalignment.
(iii) Received signal compensation
Pure amplitude variations can be effectively compensated by normalization. In this approach to compensation prior to baseline subtraction, both the received signal, v(t), and baseline, v0(t), are normalized to the maximum amplitude of the envelope of the first echo. Therefore, the residual signal using amplitude normalization is 2.6
Since spectral shape as well as signal amplitude are changed by the coil misalignment, amplitude normalization is generally inadequate. Here, an inverse filter  is employed to achieve better compensation for coil misalignment effects. This is calculated from the measurement and baseline spectra of the first echo in each signal 2.7
Consequently, the residual signal obtained using the inverse filter method is 2.8 where and .
(iv) Beam spread compensation
In order to quantify the damage detection capability of the system, it is desirable for the response to a given level of damage to be independent of its distance from a sensor. In a complex structure, this would most likely require some form of experimental calibration procedure. However, in a simple plate, the reduction in signal amplitude with distance is predictable and due to a combination of dissipative and geometric attenuation. As the attenuation of the S0 mode in composites, particularly along the fibre direction, is low , geometric attenuation due to the divergence of both transmitted and scattered waves is the dominant effect. Both result in an amplitude reduction proportional to the square root of propagation distance. Hence, the amplitude of the response from localized damage for a single sensor operating in pulse-echo mode is expected to be proportional to the reciprocal of the distance of damage from the sensor. Therefore, it is instructive to scale the amplitude of residual signals with distance from sensor, d, and to only consider the envelope of the signal. Consequently, the quantity plotted in subsequent residual signal graphs is 2.9 where vg is the group velocity of the S0 mode at the centre frequency of the input signal, 160 kHz.
3. System testing
In this section, the embedded ICTS and signal processing techniques described previously are investigated for consistency, performance in the presence of inspection wand misalignment and ultimately for their ability to detect impact damage.
(a) Sensor consistency
The performance consistency of ICTS is initially tested through the investigation of the frequency response of the devices embedded in carbon fibre plate described in §2c. The three embedded ICTS and a surface-bonded ICTS are wirelessly probed using the inspection wand at a coil separation distance of 12 mm. The effective input signals are 5-cycle, Gaussian-windowed tonebursts with various centre frequencies. They are applied using chirp excitation signals with Nc=1300 and 24 V peak–peak voltage. Meanwhile, a laser vibrometer (OFV-505; Polytec, Harpenden, Hertfordshire, UK) is used to measure the in-plane guided wave displacement generated by each ICTS on the edge of the plate nearest to each sensor. In-plane displacement measurements are necessary as the S0 mode has very low out-of-plane displacement in this frequency range. The sensor positions and laser measuring points are shown in figure 5. A typical signal measured by the laser interferometer is shown in figure 7a. The set-up for the laser vibrometer is fixed for all the measurements and the edge of the carbon fibre plate is polished at the measurement points.
Figure 7b shows the frequency response of both embedded and surface-bonded ICTS. It can be seen from this figure that consistent performance in terms of both signal amplitude and peak response frequency is obtained with the three embedded ICTS. Furthermore, the signal amplitude obtained by embedding the ICTS into composite is almost twice that of a surface-bonded ICTS. This is thought to be because a better bond layer between the unit and structure is achieved through the curing process and because both sides of the sensor are bonded to the structure. The frequency shift of the peak response between the embedded ICTS and directly bonded ICTS is due to the fact that the capacitance between the neighbouring tracks in the transducer coil is increased due to the increased permittivity contributed by the carbon fibre. It is worth noting that the consistency of embedded ICTS in absolute terms is potentially very attractive for an SHM system as it makes system calibration to an absolute level of impact damage much more straightforward.
(b) Baseline signal subtraction
Enveloped subtraction results obtained using both simple normalization, equation (2.6), and the inverse filter technique, equation (2.8), are shown in figure 8 for various degrees of coil misalignment. These results were obtained from the ICTS in the damage-free, glass fibre composite plate. Again, the input signal was a 5-cycle, Gaussian-windowed toneburst with centre frequency 160 kHz and the results were obtained using chirp excitation with Nc=1300. For the baseline signal, the inspection wand was placed 12 mm above the embedded transducer and aligned laterally to maximize the signal amplitude. It is worth noting that the residual signals are normalized to the maximum amplitude of the impact damage echo for impact damage study in the next section.
From figure 8, it can be seen that the residual signal level achieved by using the inverse filter technique is much better than the results obtained using the amplitude normalization technique, with more than three times (10 dB) better suppression of reflections from benign features achieved for all cases. However, the question remains as to whether the residual signal is low enough to allow impact damage to be detected. Also, it is possible that the inverse filter could suppress the signal introduced by the impact damage. These issues are discussed in the next section.
(c) Impact damage detection
The glass fibre panel was subjected to a 10 J impact at a distance of 125 mm from the ICTS, resulting in a delamination around 20 mm diameter, as shown in the inset in figure 5a. Figure 9 shows the residual signals after baseline subtraction using the inverse filter method, when the coils are perfectly aligned with 12 mm separation. The residual signals are normalized to the maximum amplitude of the impact damage echo.
From figure 9b, it can be seen that the reflected signal from a 10 J impact can be clearly identified at the expected distance away from the transducer of 125 mm. This qualitatively shows that the embedded ICTS is capable of detecting impact damage in a composite plate and that the inverse filter did not suppress the signal. Note that the increased residual signals after 125 mm are due to forward scattering from the impact interfering with subsequent edge reflections. Although not directly interpretable, the appearance of these signals provides another opportunity for detecting damage .
Figure 9a shows that the normalized residual signal is still significantly smaller (6 dB) than the 10 J threshold (the dashed horizontal lines) even at a distance of 1000 mm away from the sensor. The amplitude of the threshold is defined as the maximum value of the 10 J impact damage echo. This suggests that the inspection range of the embedded ICTS is around 1000 mm, when the inspection wand is perfectly aligned. The 10 J impact damage threshold is also plotted in figure 8 (the dashed horizontal lines). From this it can be seen that as the misalignment increases, the performance of the inverse filter decreases and the coherent noise level increases due to worsening suppression of reflections from structural features. However, when the coil misalignment is smaller than 5 mm the residual signal is still about 4.5 dB lower than the threshold at a distance of 1000 mm away from the sensor. To maintain a signal-to-coherent-noise ratio of 6 dB for 10 J impact damage at a distance of 1000 mm, more detailed experiments reveal that the maximum horizontal and vertical coil misalignments are 10 and 4 mm.
This paper has demonstrated a structurally integrated sensing system for large-area impact-damage detection in composite panels. The system uses an inert, guided wave sensor embedded in the composite panel, which is externally activated from an inductively coupled inspection wand. The design of the inductive coupling link has been optimized for the application using a previously developed model. To enable highly consistent, single-shot measurements to be made rapidly and without precise alignment of the inspection wand several signal processing steps have been employed: chirp excitation, deconvolution and band-pass inverse filtering. The last step of inverse filtering enables baseline signal subtraction techniques to be employed to detect the occurrence of relatively weak scattered signals from impact damage in the presence of large reflections from benign structural features. The system has been experimentally demonstrated on two composite panels and 10 J impact damage successfully detected. It is estimated that the system could detect this level of impact damage at a range of up to 1000 mm.
It is worth noting that, although the inverse filter technique is specifically developed here for compensation of ICTS coil misalignment, the principles are general and could be applied to compensation of signal changes due to environmental variation, such as temperature. These are major issues when using the baseline signal subtraction for general guided wave SHM applications.
Here, the construction of the inverse filter is based on the assumption that the echo signals used to generate it only change due to coil misalignment effects. This raises a potential concern that damage signals may be masked if they interact with this echo. This has not caused a problem in the impact detection example presented here even though forward scattered signals from the damage must have interacted with the edge echo. Although the presence of these forward scattered signals has been suppressed (as evidenced by the low residual signal level at the position of the first edge echo in figure 9b at 225 mm), the presence of the new direct echo from the impact at 125 mm is separate and is not suppressed. However, there would be a problem if the direct echo from the impact damage itself had overlapped with the first edge echo. In a real system, it would therefore be desirable for all points in a structure to be within the inspection range of more than one sensor.
- Received August 21, 2014.
- Accepted October 28, 2014.
- © 2014 The Author(s) Published by the Royal Society. All rights reserved.