laser ultrasonics experiment

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1 Introduction

2 the principles of fmc/tfm, 3 experiment and discussion, 4 discussion, 5 conclusions, conflict of interest, acknowledgments.

• List of tables
• List of figures

Technical & Applied Article

Laser-induced ultrasonic measurements for the detection and reconstruction of surface defects

Feiming Qian 1 , Guangzhen Xing 2 , Ping Yang 2 , Pengcheng Hu 1 , Limin Zou 1 * and Triantafillos Koukoulas 3

1 Institute of Ultra-Precision Optoelectronic Instrument Engineering, Harbin Institute of Technology, Harbin 150080, PR China 2 Division of Mechanics and Acoustics, National Institute of Metrology, Beijing 100029, PR China 3 Metrology Research Centre, Mass and Related Quantities Team, Acoustics, Ultrasound and Vibration Metrology, National Research Council Canada, 1200 Montreal Road, Ottawa, Ontario K1A 0R6, Canada

* Corresponding author: [email protected]

Received: 28 June 2021 Accepted: 13 August 2021

Laser-induced ultrasonic measurement is a non-contact non-destructive technology that can be employed for the testing and assessment of surface defects. In order to improve the correct identification of defects, the full matrix capture (FMC) and total focusing method (TFM) are applied on the imaging process. FMC data includes A-scans resulting from the combination of all measurement axes defined by the sequential generation and detection of utilized laser beams in the system. In this paper, an aluminium block with four holes whose diameters range from 1 mm to 2.5 mm is assessed through B-scans, the synthetic aperture focusing technique (SAFT) and FMC/TFM. The results demonstrate that the FMC/TFM technology can significantly improve the imaging quality and signal-to-noise ratio (SNR). In addition, this method has higher lateral resolution and larger imaging range compared with traditional B-scans.

Key words: Laser-induced ultrasound / Surface defects detection / Full matrix capture / Total focusing method

© F. Qian et al., Published by EDP Sciences, 2021

Surface defects are the most common damage forms in materials, such as fatigue cracks on rail surfaces and micro-cracks on the surfaces of metal containers. These defects are shallow and small in size at the beginning (generally less than 1 mm) and are not easy to detect, but will gradually deepen and become larger with usage, potentially causing serious accidents. Therefore, the detection of such surface defects is of great significance in aviation, aerospace and machinery industries amongst other fields [ 1 , 2 ]. Rayleigh waves propagate along the surface with low attenuation and their energy is manly concentrated within the depth of one wavelength, thus giving the possibility of scanning a large area at a time [ 3 , 4 ]. This technique is suitable for the detection of surface defects and is more accurate than penetrant testing using liquids or dyes and safer compared to radiographic testing. As a result, Rayleigh waves offer the potential for the detection of surface and subsurface defects. When Rayleigh waves interact with defects, part of the propagated waves will be transmitted and reflected, that is subsequently used for the defects visualization and depth analysis [ 5 ].

Traditional surface defect detection techniques mainly use monolithic wedge transducers to induce and receive Rayleigh waves. The transducers have a large beam profile that limits the accuracy for small defect detection and significantly limits the off-axis detection capabilities of defects due to the low signal-to-noise ratio (SNR). In view of these limitations, array transducers have been used with a wedge whose angle is equal to the third critical angle to improve the sensitivity and range of detection, but this method limits the imaging area directly below the wedge and requires a series of time delays to realize the deflection of the acoustic beam [ 6 ]. In recent years, the full matrix capture (FMC) using total focusing method (TFM) has been applied to accurately characterize surface defects as shown in Figure 1 , in addition to imaging outside the wedge area without complex control [ 7 , 8 ]. However, all the methods require simple geometry wedge-coupled samples and have significant limitation in detecting complex structures. Additionally, the attenuation within the wedge and the leaky loss of Rayleigh waves impose further detection limitations especially for high frequency signals [ 9 ].

In order to overcome the problems mentioned above, laser-induced ultrasound can be used for surface defect detection. For applications in non-destructive evaluation, ultrasonic waves are generated by laser irradiation based on the thermo-elastic effect, while laser interferometry is used for detection. Compared to traditional piezoelectric and electrical vibration measurement methods, there is no requirement to use wedges, which in turn facilitates large area scanning and avoids the limitations imposed due to high frequency signals. Moreover, laser-induced ultrasound can detect vibration signals in many harsh environments, such as high temperature, pressure and radiation as well as strong corrosion [ 10 , 11 ]. Nonetheless, the detection efficiency and robustness are still the main obstacles for wide application of laser-induced ultrasonic technology. In view of this, strategies of laser-induced ultrasonic phased array technology have been proposed. Some use multiple optical fibers with variable length to generate phased-array laser beams [ 12 ], others synthesizes laser ultrasonic array to image the internal defects [ 13 , 14 ]. However, most of the research of laser-induced ultrasonic phased array technology has been limited to longitudinal waves, to the authors’ knowledge, this technique has not been used for surface defect detection.

In this paper, we present laser-induced ultrasonic phased array measurements based on FMC/TFM technology for the assessment of surface defects by the usage of Rayleigh waves. The utilization of a laser-induced ultrasonic phased array technology avoided the effect of coupling on detection, a major deficiency of traditional array detection. In addition, the synthetic aperture focusing technique (SAFT) combined with B-scans has also been used to surface flaw imaging. First, the principles of FMC/TFM using a laser-induced ultrasonic array to image surface defects are described. Then, we introduce the experimental design and the results from an aluminium block with drilled holes whose diameters range from 1 mm to 2.5 mm. Finally, the discussion and conclusions are presented. The usage of laser-induced ultrasonic array and FMC/TFM imaging technology offers a good solution for surface defect detection, and may reach a wide application in non-destructive area.

Similar to SAFT, FMC/TFM utilizes synthetic apertures to improve the detection resolution and the system SNR. The key difference is on the data acquisition approach. To be specific, FMC data includes A-scans from all the transmitted and received components obtained by the sequential excitation and response of the array elements. However, SAFT data is acquired by the generation and detection of each element which is equivalent to B-scans. Thus, the FMC/TFM method may be regarded as an improvement of the SAFT approach.

Assuming that the number of array elements is N , the FMC can be described as follows. Full matrix data acquisition includes ultrasonic emissions of N elements, where each is sequentially excited as shown in Figure 2a . When the i th element generates an ultrasonic signal, all the elements receive the A-scan signals of echo data S ij ( i  = 1, 2, …, N ; j  = 1, 2, …, N ) simultaneously, so the full matrix data { S ij } of N  ×  N can be obtained as shown in Figure 2b . The advantages of FMC is that most conventional phased array ultrasonic detections can be achieved by applying specific post-processing algorithms to matrix data.

In order to adapt FMC for laser-induced ultrasonic array, the laser is moved to induce Rayleigh waves in all directions at the array positions in sequence. When a surface defect is encountered along the propagation path, Rayleigh waves will scatter and the detection interferometer is used to scan across array positions along x -axis to collect scattering signals. Following, the full matrix of N  ×  N can be acquired, where N is the number of array positions, as described in Figure 3 , and the data of S jj ( j  = 1, 2, …, N ) can be used for SAFT imaging.

Unlike phased arrays, the delay or focus scheme is not required in FMC, instead the image reconstruction algorithm is used for total focusing. The TFM is considered as the reference standard in phased array imaging [ 15 ]. For full-matrix imaging, the region of interest is identified and then discretized into grid points P i ( x p , y p ) at an appropriate resolution, for instance half of the acoustic wavelength. The amplitude of sound pressure at any point in the region of interest is considered as the contribution of acoustic waves transmitted by different array elements. To be specific, the acoustic amplitude A ( x p ,y p ) at point P i ( x p , y p ) can be expressed by:

It may be seen from Figure 3 that G i ( x i , y ) is the i th emission position, and D j ( x j , y ) is the j th detection position; Equation (1) represents the acoustic pressure at point P i ( x p , y p ). Therefore, the amplitude at each point in the region of interest can be obtained to image any surface defects present.

3.1 Rayleigh wave experiment

The aluminum block used for laser-induced ultrasonic measurements is a thin plate. Rayleigh or Lamb waves can be generated in the plate, while the velocity of these waves being similar. In order to verify what kind of waves are measured, a wedge aluminum block with different thickness was constructed. The schematic diagram of the testing block is shown in Figure 4 .

The utilized laser was an infrared Nd:YAG (INNOLAS SpitLight 500) generating 8 ns duration pulses at 1064 nm with a repetition rate of 20 Hz and an energy of approximately 100 mJ. The spot size of laser is 6 mm and an adjustable aperture was used for adjusting the diameter of the laser beam in order to change the laser energy. After being focused by a convex lens, the laser was directed towards the surface of sample and induced ultrasonic signals along the y direction, which were detected by a two-wave mixing interferometer (Bossa Nova TEMPO-2D) operating at 532 nm. The detecting beam was focused to a 75 μm diameter spot. The interferometer used an array receiver to improve detection sensitivity, while the out-of-plane and in-plane displacements were obtained simultaneously [ 16 ]. It may be noted that high sensitivity can be achieved even on a slightly rough surface because of the photorefractive effect [ 17 ]. A digital oscilloscope (Tektronix DPO 4054) was utilized to capture the output signal of interferometer, which was then averaged 64 times to increase the SNR. To determine what type of wave is measured, the wedge aluminum block was moved along the x direction to change the testing thickness. The Lamb wave (highly dispersive A0 mode) [ 18 ] could be detected when the thickness was 1 mm, as shown in Figure 5a .

The Lamb wave will approach the Rayleigh wave as the thickness of sample increases. As may be seen in Figure 5b , when the thickness (1.5 mm) is bigger than the wavelength of the Rayleigh wave (1.465 mm), Lamb wave begins to convert to Rayleigh wave [ 19 ]. The Rayleigh wave becomes dominant and the Lamb waves can be ignored when the thickness reaches 5 mm.

3.2 Experimental setup

An aluminium block with dimensions 50 × 90 × 10 mm was used for the detection and reconstruction of surface defects as shown in Figure 6a . There were 4 holes (H1, H2, H3, H4) at a surface depth of 3 mm whose diameters were 1 mm, 1.5 mm, 2 mm and 2.5 mm respectively. The experiment arrangement was the same as Figure 3 . The Nd:YAG laser was stationary during the experiment, while the sample and the interferometer were moved sequentially.

Regarding this particular experimental arrangement, the number of array elements was N  = 20 with a space of 1 mm ranging from 10 mm to 30 mm along the x axis, as illustrated in Figure 6b . In addition, the center coordinates of the four holes were (8, 9.5) (16, 9.5) (24, 9.5) and (32, 9.5) respectively. In addition, separation between adjacent features must not less than a wavelength, or strong overlapping may appear in the TFM reconstruction [ 20 ].

3.3 Results

The center frequency of the Rayleigh waves generated by the laser was 2 MHz and the bandwidth of the selected filter was 1–4 MHz. The separation of the four holes is far greater than one wavelength, so defects can be well identified. Figure 7a shows a typical filtered A-scan waveform generated and detected at x  = 16 mm. The velocities of Rayleigh and longitudinal waves in aluminium are 2930 m/s and 6100 m/s respectively. It may be seen in this figure that the longitudinal wave is reflected from the bottom of the aluminium block and the Rayleigh waves are identified as the scattered signal after encountering four holes. The SNR of the reflected waves is significantly low. At the same time, the small size of these defects (i.e. holes), diffraction and scattering effects will occur. Part of the wave will mode-convert into bulk waves (longitudinal or shear waves) [ 21 ]. Figure 7b gives the B-scan result using a part of the out-of-plane signals S jj ( j  = 1, 2 … N ). As expected, it is difficult to distinguish the surface defects from the B-scan image with poor lateral resolution which is mainly determined by the scanning step.

Figures 8a and 8b show the results of SAFT and FMC/TFM reconstruction. These two methods achieve improvements over B-scan technology including higher lateral resolution and larger imaging range. Moreover, SAFT and B-scans use the same data the former can significantly enhance the contrast of the defects boundary and reduce the interference of the longitudinal wave (bottom reflection). SAFT and FMC/TFM images can clearly show the location of four holes, though H1 and H4 are out of the scanning range. However, the SAFT image has relatively high background noise compared to FMC/TFM especially in the range from 0 mm to 5 mm. The reason is the number of signals ( N 2 ) utilized in the FMC/TFM process as opposed to SAFT ( N ). As a result, FMC/TFM improves imaging quality at the cost of measurement time.

In order to avoid the effects of the bottom-reflected longitudinal wave and the ultrasonic wave source on the results, the unwanted signals can be identified and removed based on the propagation time and pulse width [ 13 ]. The resulting SAFT and FMC/TFM images are presented in Figures 9a and 9b . The black dotted line marks the actual position of the holes. It may be seen from Figures 9a and 9b , that both imaging results of the two methods agree well at the actual defects boundaries. However, the SAFT imaging quality is relatively poor, and the defects are over-shadowed by noise. The maximum of the projection of the SAFT and FMC/TFM along the x direction are shown in Figures 9c and 9d , which shows FMC/TFM technology can better suppress the background noise.

In this study, SNR is used as a metric to compare the performance of imaging algorithms, which is defined as:

The SNR values of different methods.

As may be seen in Table 1 , B-scan technology can only image H2 and H3 in the scanning range and it is difficult to locate defects with low SNR, while SAFT and FMC/TFM technology allow for larger area testing with high SNR. Furthermore, FMC/TFM technology can greatly enhance the SNR than the other techniques.

Surface defects like small holes were fabricated and used to verify the FMC/TFM technology. It has been shown that this detecting mechanism can improve the imaging quality compared to SAFT and B-scan technologies. Although this method can significantly improve the SNR and successfully identify the surface defect boundary, FMC/TFM technology needs N  ×  N points data acquisition instead of N points for SAFT and B-scan technologies, which consequently increases the scanning time. To overcome this, a pulsed laser with higher repetition rates as well as a high-speed scanning mechanism is required. In addition, the appropriate array spacing should also be taken into consideration, which determines the directivity of the array.

The directivity pattern is the same as the optical fiber phased array, which is given as [ 22 , 23 ]:

As can be concluded from equation (7) that the lobe width decreases as the Na increases, so the maximum spacing a can not only improve the scanning speed but also strengthen the directivity of the sound beam. However, when a  = 1.465 mm (=  λ ) and θ ranging from 0 to π, the grating lobes will appear and effect the imaging quality as shown in Figure 10 , the red line and black line represent a  = 1.465 mm (= λ), a  = 2 mm (> λ), respectively. In order to optimize the directivity of the array elements, the array spacing with a  =  λ may be the best choice.

Surface defects have various shapes and patterns, such as holes and cracks. Further work will focus on the experimentation of different defects regarding sizes and shapes. In addition, this technology will be experimented on the detection of surface defects for objects in harsh environments especially under high temperature and pressure, such as non-destructive testing of high-temperature pipelines.

In conclusion, this study presented a method to locate and image surface defects using full matrix capture and total focusing method based on laser-induced ultrasonic non-destructive testing. As opposed to traditional array ultrasonic surface defect detection, laser-induced ultrasound is a non-contact technology does not require wedge-coupling that avoids attenuation of high frequency acoustic signals and can detect complex structures in many harsh environments. Additionally, the FMC/TFM technology is applied in characterizing surface cracks without complex delay controls. Similar to SAFT, this technology utilizes synthetic apertures to improve the resolution and SNR of detection. As an example, an aluminium block with 4 holes at a depth of 3 mm with diameters ranging from 1mm to 2.5 mm is used for testing. Compared to B-scan and SAFT, FMC/TFM has lower background noise and higher imaging contrast. In order to reduce the scanning time and optimize the directivity of array elements, the array spacing has been taken into consideration. When an array spacing of a  = λ the main lobe width should be minimized without doing so for the grating lobes. This method offers significant improvement and thus potential over existing methods for non-contact testing.

The authors declare that they do not have any conflict of interest.

The authors would like to acknowledge the funding and support of the National Key R&D Program of China (2017YFF0206302), National Natural Science Foundation of China (NSFC) (11904347, 51805506), China Postdoctoral Science Foundation (2018M630192, 2019T120125) and Natural Science Foundation of Heilongjiang Province of China (LH2021E060).

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Cite this article as : Qian F. Xing G. Yang P. Hu P. Zou L, et al. 2021. Laser-induced ultrasonic measurements for the detection and reconstruction of surface defects. Acta Acustica, 5 , 38.

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• Published: 14 June 2022

Real-time laser ultrasonic monitoring of laser-induced thermal processes

• Rosa E. Morales 1 ,
• Kathryn J. Harke 2 ,
• Joseph W. Tringe 2 ,
• David M. Stobbe 2 &
• Todd W. Murray 1  

Scientific Reports volume  12 , Article number:  9865 ( 2022 ) Cite this article

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• Characterization and analytical techniques
• Mechanical engineering
• Optical techniques

Intra- and inter-layer integrity of components fabricated with advanced manufacturing techniques, such as laser powder bed fusion, is dependent upon rapid heating, melting, and solidification processes. There is a need for new techniques to provide in situ feedback of these processes. Here a laser-based ultrasonic technique to probe thermal effects induced by a high-power continuous wave laser in titanium samples is described. Numerical simulations were performed to show that, for a spatially uniform heating beam, laser-induced surface acoustic waves are strongly influenced by surface heating conditions, are dispersive in the case of rapid heating, and that an abrupt velocity reduction happens upon the onset of surface melting. Furthermore, laser-based ultrasound experimental results which monitor the transient change of surface wave travel time associated with high power laser surface heating are provided. A pulsed laser is used to generate high frequency surface acoustic waves that propagate through the laser-heated region and are detected using a photorefractive crystal-based interferometer. Qualitative agreement is observed between theory and experiment with both showing a rapid reduction in the surface wave velocity at the onset of illumination and further decrease in surface wave velocity associated with melting. It is demonstrated that changes in the surface wave velocity can be used to track local heating and detect the onset of surface melting in real time.

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Introduction.

Lasers have seen widespread use in materials processing and manufacturing applications including laser cutting, welding, ablation, drilling, surface texturing, and advanced manufacturing 1 . In these applications, laser energy is absorbed by a material leading to local heating, melting, and vaporization, and in situ control of these laser-induced processes is critical to ensure the integrity of the final product 2 . For example, in laser powder bed fusion an object is built in consecutive layers by laser melting of powder mechanically distributed over the build surface 3 , 4 . Intra- and inter-layer integrity is dependent upon the rapid heating, melting, and solidification processes during which defects and material discontinuities are likely to form 3 , 4 , 5 , 6 . If the heating laser power is too high for a given scan speed, it can lead to evaporation of the material from a molten pool, a subsequent recoil force, and collapse of the melt pool, resulting in porosity in the component. Alternatively, a low laser heating power can lead to an incomplete melting and solidification process and lack of fusion defects 7 , 8 . The optimal laser parameters, defined as those required to produce a part with the lowest number of processing defects, are difficult to determine a priori due to the complexity of the process and additional variables including powder quality and machine to machine variations in laser characteristics 9 .

The success of laser-based manufacturing techniques, such as powder bed fusion, hinges on the ability of the operator to set and control process variables such as laser power and speed 10 , 11 . Nondestructive evaluation can aid in this process and provide in situ feedback on the build quality such that process variables can be adjusted in real time 12 . Several nondestructive evaluation methods such as thermal imaging, optical imaging, and conventional ultrasonics have been employed to monitor laser powder bed fusion builds 9 , 12 . These techniques offer remote access to the high temperature laser-material interaction site. Thermal imaging can provide information regarding the surface temperature distribution, and optical imaging can be used to ascertain surface morphology changes 9 . Conventional acoustic emission and ultrasound can potentially provide additional data regarding the build process 13 . In the case of acoustic emission, the sound generated during the laser heating process is detected using a microphone or transducer and analyzed to infer information about the process 14 . Using signal processing techniques, including those employing machine learning, it is possible to categorize the laser-material interaction into different regimes 15 .

In addition, the physical, mechanical, optical, and thermal properties of materials are functions of temperature. It follows that the velocity of ultrasonic waves is also temperature-dependent with an increase in temperature generally leading to a softening of the material and lower longitudinal, shear, and surface acoustic wave velocities. Ultrasound thermography is a well-known technique used to determine the temperature of a uniformly heated sample or to map out temperature distributions below the surface 16 . Ultrasound thermography has attracted much attention in the medical ultrasonics community to continuously monitor and provide feedback to thermal treatment processes in biological tissue such as high intensity focused ultrasound therapy 16 , 17 . Some of the advantages of conventional ultrasonic techniques for process monitoring are that they are not limited to surface temperature determination and that the transducers can be placed remote from the process zone. For additive manufacturing, they do require physical access to the build surface, but the contact location can be somewhat removed from the high temperature environment. The dependence of bulk wave propagation on temperature, mechanical properties, and phase state of the material has, for example, been used to infer the temperature in the processing zone, and to predict the melt pool size based on monitoring bulk waves reflected and scattered from the melt pool 18 , 19 .

Laser-based ultrasonic techniques are well suited for the real-time monitoring and nondestructive evaluation of laser-induced thermal processes. The laser detection probe beam and laser source exciting the ultrasonic waves can be separated from the high temperature manufacturing process environment using optically transparent windows or other means. Laser ultrasonic techniques have been used to evaluate advanced manufacturing builds by using surface acoustic waves and bulk waves to detect surface and near subsurface defects ex situ 20 , 21 , 22 , 23 . Surface acoustic waves and bulk waves have also been used to evaluate material microstructure and grain size 24 , to infer the surface temperature in laser-induced thermal processes 22 , 25 , to predict internal temperature distributions based on waves propagating over multiple paths 26 , and to observe melting and solidification during crystal growth 27 . Finally, laser ultrasonic methods have been used for high temperature measurements of materials properties 28 and for phase transformation studies in metals 29 , 30 , 31 .

In this work, laser-based ultrasound is used to monitor laser-induced heating and melting processes. Numerical simulations of a spatially-uniform heating beam show that laser-induced surface acoustic waves are strongly influenced by surface heating conditions. They are dispersive in the case of rapid heating where the thermally-induced mechanical property change is on the same spatial scale as the wavelength of the surface acoustic waves, and they show an abrupt velocity reduction upon the onset of surface melting. Furthermore, using a laser line source for surface acoustic wave excitation and a point interferometric detector, experimental results demonstrate the transient change of surface wave travel time associated with high power laser surface heating with a Gaussian beam placed between the source and receiver positions. In agreement with the numerical simulations, a deviation in the response is observed when the heating laser power is sufficient to cause local surface melting. The proposed technique may find effective application in the mapping of transient laser-induced thermal fields and melt zones, providing critical information for real-time process control.

Background and theory

The thermal and elastic properties of materials are dependent on temperature, with an increase in temperature generally resulting in a decrease in both material density and stiffness. A uniformly heated sample of a titanium alloy (Ti-6Al-4V) that is assumed to be homogeneous and isotropic is considered first. The temperature-dependent density, elastic modulus, and Poisson’s ratio are obtained from the software JMatPro 32 and are used to determine the longitudinal, \({c}_{L}\) , and shear, \({c}_{T}\) , wave speeds as a function of temperature. These values are, in turn, used to calculate the Rayleigh wave speed using the characteristic equation 33 . The result is shown in Fig.  1 where the Rayleigh wave speed is seen to decrease from 3000 m/s at room temperature (293 K) to 1880 m/s at the melting temperature of 1943 K. The decrease is relatively monotonic outside of a small region between 1100 and 1275 K, the temperature range at which Ti-6Al-4V undergoes an hcp (α) → bcc (β) phase transformation 34 . In the case of uniform heating, the thermal field and elastic properties are not depth-dependent and the Rayleigh waves propagate without dispersion.

Rayleigh wave speed of Ti-6Al-4V as a function of temperature ranging from room temperature to the melting temperature of 1943 K.

Rayleigh wave propagation becomes more complex in the case of transient heating, such as that produced by a high power laser source, as the thermal field and temperature-dependent elastic properties are functions of both time and space. Furthermore, the thermal properties are also temperature-dependent, and, at sufficiently high heating powers, the material will undergo a phase transformation resulting in surface melt. Here, the elastic displacement response generated by a nanosecond pulsed laser source incident upon a surface that is being heated by a spatially uniform continuous wave (CW) laser with a step-function time dependence is calculated. First, the one-dimensional temperature field produced by a CW laser is calculated using the implicit finite difference method presented by Singh and Narayan 35 . Temperature-dependent thermal properties and density are included for both the solid and liquid Ti-6Al-4V phases 36 , 37 , 38 , 39 . These properties are summarized in the Appendix. The model is used to determine the temperature as a function of time ( t ) at each depth ( d ) below the surface. Figure  2 a shows the surface temperature (at d  = 0) as a function of time where the heating laser is turned on at t  =  0 and a heating laser power density of 6 kW/cm 2 . The surface temperature rises until it reaches the melting temperature of 1943 K at a time of 465 ms, where it briefly remains until the net heat absorbed exceeds the latent heat of the phase change 35 , 40 . The melt front then begins to propagate into the material and, as shown on the right axis of Fig.  2 a, proceeds rapidly to a depth of over 35 μm. Figure  2 b shows the full extent of the calculated thermal data, with the color bar indicating the temperature rise at each depth and heating time.

( a ) Surface temperature (solid line) and melt size (dotted line) as a function of heating time for a 6 kW/cm 2 laser power density incident on Ti-6Al-4V. ( b ) Calculated temperature field as a function of heating time and depth for the 6 kW/cm 2 power density. ( c , d ) Results for a higher laser power density of 250 kW/cm 2 .

Surface acoustic waves are confined to propagate in the near surface region with a penetration depth on the order of the wavelength. For example, for a frequency of 30 MHz, surface waves in Ti-6Al-4V will be sensitive to mechanical property changes that occur over a depth of approximately 100 μm. In Fig.  2 b, the temperature rise is somewhat uniform over the near surface region, and thus it is expected that the mechanical property changes will also be relatively constant over the penetration depth, resulting in a surface acoustic wave delay that is frequency independent. The surface temperature for a significantly higher heating power density of 250 kW/cm 2 is shown in Fig.  2 c. Here, surface melting occurs at about 300 μs and there are marked thermal gradients in the near surface region within the 500 μs time window as shown in Fig.  2 d. These thermal gradients can cause dispersion of surface waves since the higher frequency waves, with a shorter wavelength, will be more influenced by the near surface region while lower frequency waves will penetrate further into the cooler substrate.

This one-dimensional thermal model allows for the calculation of temperature as a function of depth and melt front position at any time after the heating laser is turned on. Next, the pulsed laser excitation and interferometric detection of the CW laser-heated surface at a given time is modeled. The material near the surface is discretized into 400 layers, with a layer thickness of 0.6 µm. The elastic properties of each layer are calculated from the mean temperature of the layer, and the elastic wave propagation problem is then reduced to an analogous problem of wave propagation in a homogeneous, isotropic layered media 41 . When surface melting occurs, the thickness of the surface layer is set as the thickness of the melt pool and the density 38 (3920 kg/m 3 ) and longitudinal wave velocity 39 (4407 m/s) of liquid Ti-6Al-4V are used. Laser generation of ultrasound in plates 42 , 43 , 44 and in multi-layer plates on a semi-infinite substrate 45 have been previously addressed by others. The approach presented by Cheng et al. is followed in which the excitation laser source is represented as an equivalent elastic boundary source (Gaussian in space with a 10 ns pulse width), and the transfer matrix technique is used to enforce the continuity of stress and displacement across all homogeneous and isotropic layer boundaries 41 . The problem is solved in cylindrical coordinates using the integral transform technique where a Hankel transform of the elastic wave equation is taken with respect to the radial coordinate ( r ) and a Laplace transform is taken with respect to time. The normal surface displacement as a function of time at a given r is found through numerical inversion of the Hankel-Laplace transforms.

The excitation laser spot size was set to 100 μm full width at half maximum (FWHM) with a detection location at r  = 1.0 mm. The normal surface displacement as a function of time is given in Fig.  3 a for a heating power of 6 kW/cm 2 . The top curve shows the room temperature response in the absence of CW laser surface heating. A small amplitude wave arrival corresponding to the surface skimming longitudinal wave is seen at 0.16 μs followed by the larger surface acoustic wave (SAW) amplitude arrival at about 0.30 μs. The other curves show the displacement response at various times after the CW heating laser is turned on. For the signals between t  = 0 and t  = 450 ms, the shape of the surface acoustic wave remains relatively uniform, but the arrival is delayed as heating proceeds. During the last three time steps: t  = 508.5, 517.0, and 525.6 ms, melting has occurred with melt depths of 6.0, 16.8 and 27.0 μm, respectively. More prominent dispersion is seen when the surface waves traverse the molten layer; the higher frequency, short wavelength, components are delayed due to the strong interaction with the melt layer. Note that in this case, the longitudinal wave velocity in the molten liquid is higher than the shear wave velocity in the substrate so the surface waves are not leaky 46 . In general, the velocity of waves propagating on a liquid-covered half space transition from the Rayleigh wave velocity at zero thickness to the Scholte wave velocity when the thickness of the liquid is large with respect to the wavelength 47 , 48 .

( a ) Calculated laser ultrasonic signals with a source to detector distance of 1.0 mm at room temperature and heating times of 150.0, 300.0, 450.0, 508.5, 517.0, and 525.5 ms for a heating laser power density of 6 kW/cm 2 . ( b ) Temporal evolution of the displacement field for the 6 kW/cm 2 power density. ( c ) Calculated laser ultrasonic signals from the 250 kW/cm 2 heating laser power density at room temperature and heating times of 100, 200, 300, 330, 410, and 490 μs. ( d ) Temporal evolution of the displacement field for the 250 kW/cm 2 power density.

Figure  3 b shows the evolution of the displacement field calculated throughout the heating time. Here the abscissa gives the time after the heating laser is turned on while the ordinate gives the time after the excitation laser pulse. The color bar represents the normal displacement of the surface. In this image, the SAW arrival has the largest negative amplitude and is shown in red. The initial pronounced change in the SAW arrival time is associated with the rapid rise in the near surface temperature as seen in Fig.  2 a,b. The SAW arrival time is then relatively constant between 75 and 100 ms during which the α → β phase transformation in Ti-6Al-4V occurs. After this transition region, the arrival of the SAW continues to be delayed with heating time in a monotonic fashion until approximately 508 ms when a sharp break in the curve associated with surface melting is observed. The SAW signal sensitivity to the presence of melt makes it an attractive option for sensing melt depth.

Figure  3 c shows the normal displacement of the surface for a higher heating power density, 250 kW/cm 2 . The excitation source characteristics are the same as those given above. In this case, however, the heating takes place much more rapidly and surface melting starts at about 305 μs. Such rapid heating leads to strong near-surface thermal gradients (see Fig.  2 c,d) which, in turn, lead to sharp changes in the mechanical properties within the wavelength range of the broadband surface acoustic wave. At heating times between t  = 0 and t  = 300 μs, a significant degree of surface acoustic wave dispersion is evident, with the higher frequency components that probe the near-surface temperature delayed with respect to the lower frequencies that penetrate further into the cooler bulk of the material. This effect is more pronounced at later times ( t  > 300 μs) where the higher frequency SAW components are also delayed by the presence of surface melt. Figure  3 d shows the temporal evolution of the displacement field with surface heating. While the dispersion is certainly more pronounced than in Fig.  3 b, the onset of surface melting is not as evident. Note that the dispersion of the SAWs in a multilayer system can be used to back out the depth-dependent mechanical properties using a model-based inversion approach 49 . For the heating case, depth-dependent mechanical properties could ultimately be related to the subsurface temperature profile.

Experimental setup

A laser ultrasonic system was used to study surface acoustic wave propagation through a laser-heated region of Ti-6Al-4V samples. A schematic of this experimental configuration is shown in Fig.  4 . A pulsed Nd:YAG laser operating at the fundamental frequency (λ = 1064 nm) and a repetition rate of 15 Hz is used to generate broadband surface acoustic waves. The generation laser is collimated and focused to a line on the sample surface using a cylindrical lens. At the sample surface, the line source was approximately 15 mm in length and had a Gaussian FWHM of 60 μm. The laser energy at the sample was 1.6 mJ, sufficiently low that generation remained in the thermoelastic regime.

Schematic of the experimental setup depicting the three laser beams on the sample surface. The following abbreviations are used: PBS, polarizing beam splitter; λ/2, half-wave plate; λ/4, quarter-wave plate; L1, L2, L3, focusing lenses.

The displacement normal to the surface associated with the laser-generated acoustic wave was detected using a photorefractive crystal (PRC) based interferometer using a Bismuth Silicon Oxide (BSO) PRC 50 , 51 , 52 . The detection laser was a 300 mW frequency doubled Nd:YAG laser operating at 532 nm. The laser output was directed to a beamsplitter where it was divided into reference and signal beams. The reference beam was sent directly to the PRC, while the signal beam was focused onto the polished specimen surface, and, upon reflection, was sent to the PRC where it interferes with the reference beam at a 5-degree angle and creates a sinusoidal index grating inside the crystal. A portion of the reference beam diffracts from the grating in the two-wave mixing process and interferes with the signal beam at the photodetector. In addition, an AC electric field was applied across the PRC to enhance the two-wave mixing gain. Polarization optics after the PRC were used to ensure that the diffracted reference beam and transmitted signal beam were in quadrature, optimizing the detection sensitivity. The output of the photodetector was sent to a digital oscilloscope with a 200 MHz bandwidth limit and subsequently transferred to a computer and filtered using a 40 MHz second order low pass Butterworth filter. The distance between the SAW excitation line and detection point was set to 4.0 mm.

A fiber-coupled 60 W continuous wave Nd:YAG laser operating at a wavelength of 1064 nm was used to heat the sample surface. The laser output was collimated and sent through a spherical lens to the surface. The Gaussian spot size at the surface was 644 μm and the heating laser was positioned directly in between the SAW excitation laser line and the detection point using translation stages. The samples were polished Ti-6Al-4V disks with a diameter of 25 mm and a height of 13 mm. A LabVIEW code was used to control the heating laser power and to acquire laser ultrasonic signals during the heating process at a data acquisition rate of 15 Hz (corresponding to the excitation laser repetition rate). At a given heating laser power, data acquisition commenced at a heating time t h  = − 3 s, and single shot laser ultrasonic signals were acquired continuously throughout the experiment. At t h  = 0 s the heating laser was switched on and the sample was illuminated with a constant power for 10 s, after which the heating laser was turned off. To monitor sample cooling, data acquisition continued for 5 s after the heating laser was turned off. The sample was then allowed to cool to room temperature and translated to a new position. The experimental procedure was repeated a total of 10 times at each heating power and the laser ultrasonic signals collected at each time, with respect to heating laser turn on at t h  = 0, were averaged to improve the signal-to-noise ratio. After each experiment, the sample surface was inspected using an optical microscope for signs of surface melting and discoloration.

Results and discussion

Figure  5 a shows the ultrasonic signals detected at several heating times, with the prominent feature (negative dip) corresponding to the arrival of the surface acoustic wave. The top waveform shows the response in the absence of surface heating ( t h  < 0) and the negative peak in the surface wave arrival is seen at about 1.33 μs. The next four curves show the displacement responses at various times after the heating laser is turned on, and the last two curves show the response at 1 and 3 s after the heating laser is turned off. The shape of the surface acoustic wave remains relatively constant, but the arrival is delayed as heating proceeds. This lack of dispersion is expected as the heating times are long and the thermal field is thus relatively constant throughout the surface wave penetration depth. Figure  5 b shows similar results for a CW laser heating power of 46 W. The basic features of the waveforms at 30 W and 46 W are similar but the delay in the SAW is more pronounced with the larger heating power due to the increase in temperature in the laser-heated region.

Laser ultrasonic signals in Ti-6Al-4V samples with a source to detector distance of 4.0 mm at room temperature, heating times of 1, 3, 5, and 10 s, and cooling times of 1 and 3 s for CW laser heating powers of ( a ) 30 W and ( b ) 46 W. Temporal evolution of the displacement fields, where experiment time of 0 s indicates the time at which the heating laser is turned on, for CW laser heating powers of ( c ) 30 W and ( d ) 46 W.

Figure  5 c,d show the evolution of the displacement field throughout these experiments for CW laser heating powers of 30 W and 46 W, respectively. The abscissa gives the experiment time, where t h  = 0 represents the time at which the heating laser is turned on and the heating laser is turned off at t h  = 10 s, while the ordinate gives the time after the excitation laser pulse. The color bar represents the out of plane displacement of the surface. In these images, the SAW arrival has the negative-most amplitude shown in red. In both images, the surface wave arrival shows a marked delay associated with the onset of surface heating. This is followed by a more gradual change between t h  = 2 s and t h  = 10 s as heat diffuses through the sample and the sample temperature begins to approach the steady state. Finally, a rapid decrease in the arrival time is observed when the heating laser is turned off. Interestingly, the 46 W heating power not only causes more of a surface wave delay but the shape of the arrival over heating time is quite different than the 30 W heating laser case, and there is not a simple linear scaling between them. Furthermore, optical microscopy revealed no visible markings left on the surface from the 30 W heating power experiments, while there were clear discolorations and surface texture changes, indicative of surface melting, resulting from the 46 W heating power experiments.

The frequency content of the surface acoustic waves generated in these experiments extends to 32 MHz. At this frequency and with the slow heating time of 10 s, the SAW is minimally dispersive with the increasing temperatures as seen in Fig.  5 , while the delay of the SAW is sensitive to the heating power or surface state. The arrival time of the negative peak of the surface wave was determined as a function of heating time for different laser powers. The surface wave delay was then determined by subtracting the room temperature arrival time. Figure  6 a shows the real-time transient surface wave delays for heating powers of 30, 34, and 40 W. Note that at these lower heating powers, all the curves show a similar shape. If these curves are divided by the heating laser power to determine the normalized surface wave delay (in units of ns/W), all the curves collapse to a single curve as shown in Fig.  6 b. This indicates that in this regime, the surface wave delay is a simple linear function of the heating power. Optical microscopy confirmed that there were no visible changes to the surface subsequent to heating in this power range. At higher heating powers the response is quite different, however: experiments performed at 46, 48, and 56 W resulted in visible markings and surface texture changes. The delay curves for these heating powers are plotted in Fig.  6 c together with the results from the 30 W experiment during which no surface changes were observed. The SAW delay is significantly more pronounced and the normalized surface wave delay plots in Fig.  6 d show a distinct change in shape. While all the curves follow the 30 W heating curve for early heating times, deviations begin at later times, with the deviations occurring earlier for higher power. It is hypothesized that the nonlinearity in the curve is associated with phase change and the presence of surface melt.

( a ) Real-time transient surface wave delays and ( b ) normalized surface wave delays, in units of ns/W, for heating powers of 30, 34, and 40 W. ( c ) Real-time transient surface wave delays and ( d ) normalized surface wave delays for heating laser powers of 30, 46, 48, and 56 W.

Incremental heating experiments were performed in which the laser power was fixed at 48 W but the illumination time was varied between 0.5 s and 10.0 s in 0.5 s intervals. The sample surface was optically observed after each interval, and the sample was translated between measurements such that a new region was illuminated. Figure  7 a shows a subset of the surface wave delay data, again demonstrating that surface waves can be very effective for the characterization of the transient thermal field. Figure  7 b shows a zoom-in of the first 5.0 s of normalized surface wave delay for heating laser powers of 30 W and 48 W. The curves are nearly identical up until about 2.0 s of illumination, after which the 48 W curve shows a significantly higher delay. Optical micrographs of the surface are shown in Fig.  7 c. There is no detectable surface discoloration before 2.0 s of heating. However, there is a clear indication of a melt region starting at 2.0 s and this region continues to grow at longer heating times.

( a ) Subset of surface wave delay data from the incremental heating experiment at a fixed laser power of 48 W. ( b ) Zoom-in of the normalized surface wave delay for heating laser powers of 30 W and 48 W. ( c ) Optical micrographs of the sample surface after the indicated times of heating.

Note that in the modeling results, a one-dimensional illumination model was used to elucidate the effects of surface heating and melting on surface wave propagation. Thus a quantitative comparison with experimental results, in which the waves propagate through a region heated with a Gaussian laser source, is not possible. Nevertheless, there is qualitative agreement between the two with a rapid delay in the surface wave velocity at the onset of illumination and further decrease in surface wave velocity associated with melting. The surface acoustic wave delay curves, such as those shown in Figs.  6 and 7 , may prove useful in monitoring laser-induced thermal processes, particularly in more complex cases involving phase transformation. Changes in these curves indicate variations in the heating laser parameters or laser-material interaction. These experimental results are limited to relatively slow heating, where the temperature is uniform over the surface wave penetration depth. A higher repetition rate pulsed laser can be used to probe more rapid thermal processes and potentially obtain quantitative information about the heated zone, including temperature distribution and melt pool depth. It is also important to point out that laser additive manufacturing techniques are conducted in a layer-wise manner, and track formation and overlap add complexity to the heating and melting processes.

Laser-based ultrasonics is a non-contact technique that can be used to monitor transient laser-induced heating and melting processes. Numerical simulations were performed to show that, for a spatially uniform heating beam, laser-induced surface acoustic waves are strongly influenced by surface heating conditions, are dispersive in the case of rapid heating, and that an abrupt velocity reduction happens upon the onset of surface melting. The complementary experimental results monitor the transient change of surface wave travel time associated with high power laser surface heating. Qualitative agreement between theory and experiment is observed: both show a rapid reduction in the surface wave velocity at the onset of illumination and further decrease in surface wave velocity associated with melting. This technique may ultimately find application in the mapping of transient laser-induced thermal fields and melt zones, providing critical information for real-time process control in advanced manufacturing systems including those relying on laser powder bed fusion.

Data availability

The datasets generated during and/or analyzed during the current study are available from the corresponding author on reasonable request.

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Acknowledgements

This work was performed under the auspices of the U.S. Department of Energy by Lawrence Livermore National Laboratory under Contract DE-AC52-07NA27344 and was supported by the LLNL-LDRD Program under Project No. 19-ERD-008. The document release number is LLNL-JRNL-831155. R.E.M. thanks the National GEM Consortium for supporting the preparation of this manuscript. The authors thank Mark McDaniel, lead technical illustrator at Lawrence Livermore National Laboratory, for his assistance in creating Fig. 4 .

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Contributions

T.W.M., D.M.S., and R.EM. suggested the experiments. R.EM. constructed the analytical models, performed the experiments, analyzed the data, and prepared figures. J.W.T., D.M.S., and T.W.M. acquired funding for this work. J.W.T. and K.J.H. provided resources for relevant work. R.E.M and T.W.M. wrote the manuscript. All authors reviewed the manuscript.

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Morales, R.E., Harke, K.J., Tringe, J.W. et al. Real-time laser ultrasonic monitoring of laser-induced thermal processes. Sci Rep 12 , 9865 (2022). https://doi.org/10.1038/s41598-022-13940-5

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Received : 10 February 2022

Accepted : 30 May 2022

Published : 14 June 2022

DOI : https://doi.org/10.1038/s41598-022-13940-5

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