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Abstract
A miniaturized fiber optic hydrophone (FOH) based on a composite metal diaphragm with an air back cavity and a high finesse extrinsic Fabry-Perot interferometric (EFPI) scheme for low-frequency underwater acoustic sensing is proposed and experimentally demonstrated in this paper. A composite metal diaphragm is used to improve the stability of the hydrophone. A balance channel is used to equilibrate the hydrostatic pressure and maintain an air cavity, which improves the mechanical sensitivity. In addition, a white light interferometry (WLI) phase demodulation is used to demodulate the high finesse interferometer consisted of the fiber collimator end face and the diaphragm, which improves the phase sensitivity. Experimental results show that the enhanced phase sensitivity of the hydrophone is about −122.5 dB re 1 rad/µPa @ 200 Hz and the sensitivity fluctuation is below 2.5 dB between 3 Hz and 400 Hz, while the minimal detectable pressure (MDP) is 63.7 µPa/Hz1/2 @ 400 Hz. Due to its miniaturized structure and high sensitivity, the FOH may have an enormous potential in underwater target detection.
© 2022 Optica Publishing Group under the terms of the Optica Open Access Publishing Agreement
1. Introduction
Fiber optic hydrophone (FOH) is a specific application of fiber optic sensor. Compared with the traditional piezoelectric hydrophone, FOH has the advantages of compact size, high sensitivity, anti-electromagnetic interference, and the ability of multiplexing, etc., which make it more suitable for extremely harsh environment. Because of its unique advantages, FOH is extremely popular in military and civilian fields, including underwater target detection [1], biomedical monitoring [2], and oil exploration [3]. Various kinds of FOH structures have been reported such as fiber laser hydrophones [4,5], fiber grating hydrophones [6,7] and fiber interferometer hydrophones. The interferometric structure has higher detection sensitivity and is currently a research hotspot in underwater sound field detection. The mandrel FOHs based on Michelson interferometer (MI) and Mach-Zehnder interferometer (MZI) are mature at present [8–11]. By winding the sensing fiber on the elastic cylinder, extremely high sensitivity can be obtained, but this limits the dimensions of the sensor in a certain extent. Therefore, the extrinsic Fabry-Perot interferometer (EFPI), which can balance sensitivity and miniaturization at the same time, is very valuable for research.
The FOH based on EFPI, known as micro-cavity FOH, utilizes Fabry-Perot interference to demodulate the displacement of a sensing element to achieve underwater sound detection. The performance of the sensor is related to the sensing element and the structure of the micro-cavity. The sensing elements of the FOH include the cantilever [12] and the diaphragm. The diaphragm has better sound pressure response underwater, and researchers have conducted a lot of research on the diaphragm of various materials, including silicon [13,14], photonic crystal [15,16], polymer [17–19], and graphene [20]. The metal diaphragm has many applications in air acoustic sensors [21–25], because of its excellent mechanical properties. However, there is no report on high performance metal diaphragm EFPI hydrophone yet. Choi et al. demonstrated that the air cavity structure can improve the mechanical sensitivity of the diaphragm type hydrophone [26] and added a hydrostatic pressure balancer to increase resistance to ambient pressure of the piezoelectric hydrophone [27]. However, the piezoelectric hydrophones have inherent problems of being sensitive to the electromagnetic interference. Liu et al. applied the balance structure to the fiber optic acoustic sensor [28], but the fiber optic hydrophone based on the similar structure did not obtain a good acoustic response [29]. Low-frequency acoustic signals are less attenuated in water and have a longer propagation distance. The detection of low-frequency underwater acoustic signals is the focus of underwater target detection. Therefore, improving the low-frequency performance of the FOH is an urgent problem to be solved.
In our work, a low-frequency EFPI hydrophone based on a composite metal diaphragm is proposed and demonstrated. The sensitivity of the hydrophone is greatly improved through mechanical and optical enhanced sensing methods, while ensuring the miniaturization. A composite structure improves the strength of the diaphragm, which improves the stability and dynamic range of the hydrophone while ensuring sensitivity. An air back cavity improves the mechanical sensitivity of the hydrophone and ensures the consistency of the spectrum during the hydrophone entering the water. A high finesse EFPI scheme based on fiber collimator improves the phase sensitivity of the hydrophone. The frequency response of the sensitivity enhanced FOH within 3 ∼ 1000 Hz is measured. The phase sensitivity of the sensitivity enhanced hydrophone is about −122.5 dB re 1 rad/µPa @ 200 Hz, which is improved by about 6 dB compared with common interference of the same structure. An average sensitivity of the proposed hydrophone is −124 dB re 1 rad/µPa with a sensitivity fluctuation below 2.5 dB in the frequency range of 3 ∼ 400 Hz, and the resonance frequency of the FOH is 901 Hz. Experimental results show that minimal detectable pressure (MDP) is about 63.7 µPa/Hz1/2 @ 400 Hz, which is lower than the sea background noise [30]. An excellent linearity and a high dynamic range in the excess of 107.6 dB is demonstrated at 250 Hz. The proposed FOH has excellent low-frequency underwater acoustic sensing performance and shows great potential in underwater target detection.
2. Fiber optic hydrophone sensing mechanisms
The vibration of the diaphragm determines the performance of the EFPI to a large extent, and it is related to the performance of the diaphragm and the structure of the hydrophone. The gold diaphragm has an excellent response in the air, due to its good chemical stability and high reflectivity [24]. However, the nanoscale pure gold diaphragm is fragile, which brings difficulties to the fabrication of the sensor and the stability in a complex underwater environment. To address these issues, a two-layer composite diaphragm (chromium thickness hCr = 30 nm and gold thickness hAu = 300 nm) is proposed. The chromium is used as the adhesion layer during the diaphragm growth process to improve the quality of diaphragm formation. Furthermore, the chromium has a body-centered-cubic structure which makes it have higher strength, and the interface between the chromium and the gold can hinder the dislocation motion [31], thereby improving the strength of the diaphragm. For the composite diaphragm, two materials have the same diameter, so the parameters of the diaphragm are related to the thickness of two materials. While density, Young's modulus, and Poisson's ratio of chromium and gold are respectively ρCr, ρAu, ECr, EAu, υCr and υAu, the corresponding parameters of the composite diaphragm with the thickness hdia can be obtained by Eq. (1–3) [32]:
The structure of the proposed FOH is shown in Fig. 1. Corresponding geometrical dimensions and material parameters are shown in Table 1. Stainless steel is chosen as the main body material to fix the diaphragm rim and the fiber collimator. The sensitivity of the micro-cavity hydrophone with the air back cavity is superior to the same structure with the water back cavity [26,33]. However, sealing the back cavity simply will make the hydrophone sensitive to hydrostatic pressure, which will reduce the sensitivity of the hydrophone underwater. To ensure that the hydrophone can still have a high sensitivity underwater, a microchannel for pressure balance is processed in the stainless steel encapsulation. The design of the balance channel geometry can ensure that the water will not flow into the back cavity when working above a certain depth underwater [27].
Table 1. Geometrical Dimensions and Material Parameters
Depending on the force against deflection, the vibration model of the diaphragm can be divided into the plate model and the membrane model. When the tension coefficient of the diaphragm is greater than 20, it can be explained by the membrane model, and the tension coefficient here can be expressed as Eq. (4) [34]:
The tension coefficient of the diaphragm in this paper can be calculated as 531, which means the force against deflection is dominated by the internal stress. For the membrane model, the first order resonant frequency and the deflection u(r) of the diaphragm when the operating frequency is much lower than the resonance frequency can be written as Eq. (5) and Eq. (6), respectively [35]:
Equivalent-circuits model is commonly used to predict and design acoustic-vibration systems. The proposed FOH dimensions are much smaller than the acoustic wavelengths of interest, so the acoustic response of the sensor can be analyzed according to a lumped elements model [16]. An equivalent circuit of the proposed FOH is given in Fig. 1. The circuit elements with the membrane model are defined in Table 2, where the mechanical parameters are transformed into acoustical elements using an effective area [36,37]. In the equivalent model, the voltage drop across the diaphragm compliance corresponds to the pressure difference on the diaphragm. The acoustic performance of the sensor structure can be described by the transfer function, which reflects the relationship between the acoustic pressure in the environment and the pressure applied on the diaphragm. The transfer function is calculated by the equivalent circuit and can be expressed as Eq. (7):
The transfer function of the proposed sensor is simulated in Fig. 2, while the effect of high-pass filtering under the condition of the water back cavity is ignored. As shown in the simulation results, when the balance channel is designed to ensure that water will not completely flow into the back cavity, the amplitude of the transfer function in the flat area is greater, which means the sensitivity of the proposed hydrophone is improved.
Table 2. Circuit elements definitions
As for the Fabry-Perot interferometer shown in Fig. 3, k represents the number of round trips in the cavity, which is an integer. Therefore, the reflected light fields at different k can be expressed as Eq. (8):
Considering the interference of M-th and N-th reflected light, the interference light intensity IMN is the sum of the M-th and N-th light field amplitude multiplied by its complex conjugate, which can be expressed as Eq. (9):
A white light interferometry (WLI) phase demodulation algorithm [38] is used to calculate the phase change of the FP interference. Performing fast Fourier transform (FFT) on the collected spectrum at each moment, a series of characteristic frequency peaks can be obtained in the spatial frequency spectrum. The order of the characteristic peak on the spatial frequency spectrum is D, which is a positive integer. When the length change of FP cavity with time is ΔL(t), the characteristic spatial frequency of the corresponding FFT can be written as Eq. (10):
It shows that when the cavity length and the spectral sampling range are determined, the characteristic spatial frequency of the Fourier spectrum is only related to the difference in the order of the two reflected lights involved in the interference. The D-th characteristic peak contains the information of all the two-beam interference combinations with reflected light orders satisfying M – N = D. When FPI is placed in the sound field, the sensitivity of cavity length change is S0, which is related to the mechanical sensitivity of the diaphragm and the transfer function. If the white-light interference signal is demodulated at the D-th order characteristic peak of the spatial frequency spectrum, the phase sensitivity of FPI can be expressed as Eq. (11):
Therefore, it can be concluded that when performing WLI demodulation at the D-th order peak of the spatial frequency spectrum, the phase sensitivity will be increased by multiples, and its magnification is D. Considering the reduction of light intensity after multiple reflections and the limitation of demodulation equipment, the third and higher order characteristic peaks are difficult to demodulate, so the second order characteristic peak is utilized in this paper.
The divergence of the light emitted by the single-mode fiber is very large [39]. To reduce the space loss of the beam and obtain a high finesse spectrum with more obvious high order peaks, a fiber collimator based on surface-coated Gradient-index lens (G-lens) and single-mode fiber is used. According to the optical thin-film theory [40], the reflectivity of the composite diaphragm is 0.98. The relationship between the reflectivity of G-Lens end face R and the normalized light intensity of second order characteristic peak is simulated and the result is shown in Fig. 4(a). Different space losses are simulated here. It can be found that when the end face reflectivity of the G-Lens is about 0.6, the normalized amplitude value of the second order peak is relatively larger. Therefore, a light splitting film is coated on the G-Lens end face to guarantee the reflectivity is 0.6, which can make the sensor obtain more obvious second order characteristic frequency peaks. The final design of the fiber collimator based on G-Lens is shown in Fig. 4(b). The end face of the optical fiber and the inner surface of the lens are coupled at an angle of 0° and coated with 99.5% anti-reflection coating. The light is transmitted in the optical fiber, and the Gaussian beam passes through the G-Lens to become parallel light.
Fig. 4. Analysis and design of the fiber collimator. (a) Simulated relationship between reflectivity and normalized light intensity. (b) Schematic diagram of the fiber collimator.
3. Hydrophone fabrication and spectral analysis
The hydrophone is mainly composed of the sensing diaphragm, the stainless steel encapsulation and the fiber collimator, where the encapsulation is divided into three parts, namely the adjustor, the main body, and the bottom. The sensing diaphragm is prepared by electron beam evaporation process and transferred to the top of the adjustor with inner diameter of 4.2 mm using the epoxy resin adhesive (353ND), the specific method refers to [41]. The fiber collimator is fixed to the stainless steel bottom by a plastic tail sleeve. The main function of the bottom is to seal the balance channel processed on the main body with a gasket. To ensure that the spectrum obtained by the experimental equipment has enough interferential cycles, the distance between the end face of the fiber collimator and the diaphragm is about 375 µm. For the FP cavity composed of the fiber collimator, the space loss mainly comes from the angle tilt. To adjust the parallelism between the diaphragm and the end face of the fiber collimator during the hydrophone assembly, four adjustment screws and an adjustment gasket are used. While adjusting the parallelism with the screws, the interference spectrum is observed in real time by optical spectrum analyzer (OSA, Yokogawa, AQ6370C-20), and the adjustment is completed when the spectrum has obvious multi-beam interference characteristics. After the spectrum is stabilized, the adjustor is fixed by glue, and the hydrophone assembling process is completed. The assembly schematic diagram and the photo of the hydrophone are shown in Fig. 5(a) and (b), respectively.
Fig. 5. Fabrication and structure of the FOH (a) The assembly schematic diagram of the FOH. (b) The photo of the FOH.
The reflection spectrum removed the light source background intensity of the sensor in the air and underwater can be obtained by the OSA as shown in Fig. 6 (inset). The red curve is the spectrum of the sensor in the air after assembly and the bule curve is the spectrum of the sensor placed underwater. It is clearly that the spectrum of the sensor before and after entering the water is almost the same and the free spectral range (FSR) of them both are about 3.2 nm. The underwater spectrum is slightly deformed compared to the spectrum in the air because the sensor underwater is more sensitive to the external vibration noise, and the sampling rate of the spectrometer is limited, so the states of different cavity lengths are recorded in one frame of the spectrum. Ignoring the subtle differences in the spectrum related to the environment, the identical spectrum shows that there is still air in the back cavity, which proves the effectiveness of the balance channel. The spatial frequency spectrum of the sensor in both environments calculated from the spectrum data is shown in Fig. 6. The result with multiple peaks indicated the effectiveness of the high finesse EFPI scheme.
4. Experimental results and analysis
When underwater acoustic frequency is lower than 1 kHz, it is difficult to obtain a free sound field. To improve accuracy of underwater acoustic calibration, a vibrating liquid column calibration system is used to calibrate the sensitivity of the optical fiber hydrophone by comparison. Measurement setup is shown in Fig. 7. Sinusoidal signal of a specific frequency is output from a data acquisition hardware (B&K LAN-XI 3160) and amplified by a power amplifier to drive a shaker. A designed water tank moves synchronously with the shaker, which can ensure the same sound pressure at the same depth in liquid column. The FOH and a reference hydrophone (B&K Type 8104) are placed in the tank at the same depth of 5 cm. The reference hydrophone is connected to the data acquisition hardware through a converter (B&K Type 2647-B), and its sensitivity can be calculated as 908µV/Pa (0.1Hz ∼ 20kHz). The WLI phase demodulation algorithm is used to calculate the signal of the FOH. A broadband amplified spontaneous emission (ASE) source is adopted to illuminate the FOH, and the reflected light from the sensor is delivered into an interrogation monitor (Ibsen I-MON 512 HS), which is applied to collect spectrum data. Unlike the OSA used to observe the spectrum of the sensor previously, the interrogation monitor is designed for high speed and real-time analysis. Therefore, the phase of the FOH can be demodulated from spectrum by software and output to a computer in real time. Compared with voltage output signal from conventional intensity demodulation system, the phase signal of the WLI demodulation system can directly reflect phase sensitivity of the sensor, and it avoids a complex spectrum stabilization system. Combined with the high speed interrogation monitor, the demodulation system can cover the bandwidth required for calibration. The sampling rate of the interrogation monitor is 1 kHz when the test frequency is lower than 10 Hz, and the sampling rate is 10 kHz in the other experiments.
Different from the OSA, the I-MON 512 HS can only acquire 512 points in the entire spectrum. Due to the limitation of the operating wavelength range of the ASE, the interference spectrum can only be observed in the range of 1526 nm to 1563 nm. Figure 8(a) shows a frame of the spectrum acquired by the interrogation monitor during the experiment. To control the minimum frequency interval in the spatial frequency spectrum to make the characteristic peaks more obvious, the blue dashed part of the interference spectrum is intercepted. Performing FFT on this part, the corresponding spatial frequency spectrum is shown in Fig. 8(b). It can be seen that there is an obvious second order peak in the spatial frequency spectrum, but the third order peak is buried in the background, which is consistent with the previous explanation.
Fig. 8. Spectrum and characteristic peaks of the FOH acquired by the interrogation monitor. (a) Reflection spectrum of the FOH. (b) Spatial frequency spectrum of the FOH.
Sinusoidal signals are used to test time domain response of the proposed FOH. Driven by an input signal, there is a sinusoidal sound pressure signal at the same depth underwater, while amplitude and frequency information of the sound pressure signal can be obtained by the reference hydrophone. At the same time, phase change of the FOH is demodulated through the WLI phase demodulation algorithm mentioned above. Figure 9 demonstrates the time domain signal and corresponding FFT spectrum of the FOH and the reference hydrophone, when the acoustic signal frequency is 200 Hz. The black curve and red curve in Fig. 9(a), (c) are the demodulated results at the first and second order peaks, respectively. It is clear that the demodulated phase amplitude at the second order peak is almost double that of the first order, which is consistent with the results of the previous analysis, and it proves that high finesse EFPI scheme can improve phase sensitivity of the sensor. According to the result in Fig. 9(c), the amplitude of the phase is increased by 6.6 dB. This is because the spectrum of an integer period cannot be obtained during the experiment, and there is a spectrum leakage, which brings the error between the theory and the actual experiments. At the same time, since the diaphragm and the end face of the collimator cannot be perfectly parallel, the phase amplitude is increased by more than 6 dB. Considering the second order peak demodulation, phase amplitude of the sensor is 10.1 dB re 1 rad, while sound pressure level (SPL) is 132.6 dB re 1 µPa. Therefore, the sensitivity of the proposed FOH is calculated to be −122.5 dB re 1 rad/µPa @ 200 Hz. Similarly, the responses of the FOH and the reference hydrophone at 3 Hz are shown in Fig. 10(a), (b), and the sensitivity of the FOH can be calculated in the same way as −122.5 dB re 1 rad/µPa @ 3 Hz.
Fig. 9. Time domain response and sensitivity analysis of the FOH at 200 Hz. (a) Time domain signal of the FOH. (b) Time domain signal of the reference hydrophone. (c) FFT spectrum of the FOH. (d) FFT spectrum of the reference hydrophone.
Fig. 10. Time domain signal and FFT spectrum of the hydrophone at 3 Hz. (a) The FOH. (b) The reference hydrophone.
The frequency response is an important parameter of the hydrophone. Sinusoidal signals with various frequencies are used to test the frequency response curve of the proposed FOH with the range of 3∼1000 Hz. The measured results and the simulated frequency response curves are shown in Fig. 11, while the blue points in the figure are second order peak demodulated results and the red curve is the simulation curve based on the equivalent model. The first order resonant frequency of the diaphragm can be calculated as 11 kHz according to Eq. (5), so the mechanical sensitivity of the diaphragm can be considered as 50.6 nm/Pa according to Eq. (6). As we show in Code 1 (Ref. [42]), the simulation curve is obtained by multiplying the transfer function by the mechanical sensitivity of the diaphragm and the optical phase conversion factor. It can be seen that the measured results are consistent with the simulation results. From the simulation results, it can be considered that the sensitivity of the FOH is averagely −124 dB re 1 rad/µPa between 3 Hz and 400 Hz, and the resonant frequency is about 901 Hz. The sensitivity fluctuation of the FOH between 3 Hz and 400 Hz is less than 2.5 dB, which is not only related to the resonance effect of the cavity, but also to the uncertainty of the measurement setup.
Phase noise of the FOH shown in Fig. 12(a) is characterized in a quiet environment, while the MDP shown in Fig. 12(b) is obtained by dividing the noise-equivalent phase by the frequency response. According to Fig. 12(a), the noise of the FOH at 3 Hz and 400 Hz can be calculated as 5.73 mrad/ Hz1/2 and 44.8 µrad/Hz1/2, while the sensitivity are 0.7494 rad/Pa and 0.7032 rad/Pa, respectively. Therefore, the MDP of the FOH with 7.65 mPa/Hz1/2 @ 3 Hz and 63.7 µPa/Hz1/2 @ 400 Hz can be achieved. Linearity of the FOH is experimentally determined at a frequency of 250 Hz and the result is performed in Fig. 12(c). An excellent linearity with R2 value of 0.99995 is shown in the measured pressure range, and a power spectrum of the FOH in the sound field with a sound pressure amplitude of 32.97 Pa (150.4 dB) is shown in Fig. 12(d). It can be seen from the power spectrum that the higher harmonics appear at this time, and a total harmonic distortion (THD) of the FOH is −30.6 dB, while part of the harmonic energy is caused by acoustic source itself. The MDP of the FOH in a 1-Hz bandwidth at 250 Hz is 138.03 µPa (42.8 dB). Therefore, a dynamic range of the FOH is in the excess of 107.6 dB with THD below −30 dB. Overall, the proposed hydrophone has an excellent low-frequency underwater acoustic signal detection ability.
Fig. 12. Noise property and dynamic range of the FOH. (a) Phase noise of the FOH. (b) MDP of the FOH. (c)Relationship between sound pressure and phase amplitude of the FOH at 250 Hz. (d)Power spectrum of the FOH in the sound field at 250 Hz with a sound pressure amplitude of 32.97 Pa.
5. Conclusions
In summary, a miniaturized FOH with an air back cavity and high finesse EFPI scheme for low-frequency underwater acoustic sensing is proposed. The experimental results prove that the designed structure can realize an air back cavity, and the high finesse scheme can improve the phase sensitivity of EFPI sensors. The proposed FOH has an extremely high sensitivity with −122.5 dB re 1 rad/µPa @ 200 Hz and a flat response range between 3 Hz and 400 Hz with the sensitivity fluctuation below 2.5 dB, which is consistent with the equivalent-circuits model. The MDP of the FOH with 63.7 µPa/Hz1/2 @ 400 Hz can be achieved, which are lower than sea background noise. Overall, the proposed FOH has advantages of high sensitivity and miniaturization, and it has an excellent performance for low-frequency acoustic sensing underwater. By choosing different cavity parameters, the working frequency band and depth of the hydrophone can be changed to adapt to different applications.
Funding
National Natural Science Foundation of China (61775070); NSFC-RS Exchange Programme (62111530153); Science, Technology and Innovation Commission of Shenzhen Municipality (2021Szvup089); Science Fund for Creative Research Groups of the Nature Science Foundation of Hubei (2021CFA033).
Disclosures
The authors declare no conflicts of interest.
Data availability
Data underlying the results presented in this paper are not publicly available at this time but may be obtained from the authors upon reasonable request.
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