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Abstract
Metal coatings can protect the fragile optical fiber sensors and extend their life in harsh environments. However, simultaneous high-temperature strain sensing in a metal-coated optical fiber remains relatively unexplored. In this study, a nickel-coated fiber Bragg grating (FBG) cascaded with an air bubble cavity Fabry-Perot interferometer (FPI) fiber optic sensor was developed for simultaneous high temperature and strain sensing. The sensor was successfully tested at 545 °C for 0-1000 µɛ, and the characteristic matrix was used to decouple temperature and strain. The metal layer allows easy attachment to metal surfaces that operate at high temperatures, enabling sensor-object integration. As a result, the metal-coated cascaded optical fiber sensor has the potential to be used in real-world structural health monitoring.
© 2023 Optica Publishing Group under the terms of the Optica Open Access Publishing Agreement
1. Introduction
In a variety of industrial applications, the health status of large equipment in harsh environments such as oil and gas wells, aircraft engines, and nuclear reactors needs to be monitored. Fiber Bragg gratings (FBGs) as optical sensors have demonstrated advantages over traditional electrical sensors in extreme environments, including compact size, chemical corrosion resistance, cost effectiveness, immunity to electromagnetic interference (EMI), and ease of distribution.
FBG sensors have been used for high-temperature sensing in practical industrial applications [1–3]. Conventional Type I FBGs suffer from reflectivity decay when the temperature is over 200 $^{\circ }$C [4], which largely influences their sensing performance at high temperatures. High temperature-resistant FBGs like femtosecond laser-inscribed FBGs (fs-FBGs) have been reported to survive at temperature over 1000 $^{\circ }$C [5,6]. Besides, decoupling temperature and strain in a FBG is another challenge due to the intrinsic cross sensitivity between each other, which limits their ability to provide accurate results when both temperature and strain vary simultaneously.
Researchers have proposed various approaches to achieve simultaneous temperature and strain sensing using FBG sensors. Rao $et$ $al.$ [7] used a hybrid fiber optic sensor consisting of a long-period fiber grating (LPFG) inscribed using CO$_2$ laser pulses and a micro extrinsic Fabry-Perot interferometer (MEFPI) fabricated by applying excimer laser pulses. This approach exploited the temperature insensitivity of the MEFPI and the strain insensitivity of the LPFG to simultaneously record the reflection spectrum of the MEFPI and the transmission spectrum of the hybrid sensor, allowing strain measurements in the range of 0 to 500 $\mu \varepsilon$ at 500 $^{\circ }$C. Similarly, Huang $et$ $al.$ [8] achieved simultaneous temperature and strain measurement at 600$^{\circ }$C in the range of 0 to 4000 $\mu \varepsilon$ by exploiting the features of LP06 and LP07 modes excited in an LPFG. Tian $et$ $al.$ [9] adopted a hybrid approach, using a hollow-core silica tube-based Fabry-Perot interferometer (FPI) sandwiched between two single-mode fibers (SMFs), cascaded with a regenerated FBG (RFBG) inscribed in one arm, and encapsulated the RFBG with an alumina tube to isolate the effects of external stress. This approach enabled strain sensing up to 450 $\mu \varepsilon$ at 800 $^{\circ }$C. Overall, the cascading of various passive optical devices, including LPFG, FBG, and FPI, has proven to be an effective method of simultaneously measuring temperature and strain parameters.
The glass transition temperature ($T_g$) of silica optical fibers has been reported to be around 1200 $^{\circ }$C [10]. Although bare silica optical fibers can withstand high temperatures, direct use of bare fiber is prone to break due to its fragility. A protective coating cover the fiber is needed to improve mechanical strength as a fiber sensor. Polymer coatings such as acrylate [11], polyimide [12], and PDMS [13] are commonly used to improve the mechanical strength of optical fibers at room temperature. However, such polymer coatings tend to burn and carbonize when exposed to high temperatures, with polyimide being the most heat resistant among them, but unable to withstand temperatures higher than 300 $^{\circ }$C for extended periods of time [14].
Compared to polymer coatings, metal coatings generally obtain a higher temperature resistance and a larger mechanical strength. A variety of techniques have been used to metalize the fiber surface, including ultrasonic additive manufacturing (UAM) [15], ultrasonic welding (UW) [16], selective laser melting (SLM) [17], casting [18], magnetron sputtering [19], flash evaporation [20], chemical plating [21], and electroplating [22]. Petrie $et$ $al.$ [15] used UAM to embed copper-coated fibers in a copper matrix for temperature measurement at 500 $^{\circ }$C, while Li $et$ $al.$ [16] used UW to embed nickel-coated fibers in a pure aluminum foil, resulting in a sensor that passed a 40 N load test. Havermann $et$ $al.$ [17] used SLM to embed nickel-coated fibers with FBGs in 316 stainless steel, and Bian $et$ $al.$ [23] achieved simultaneous temperature and strain measurement by embedding cascaded RFBGs in an aluminum test specimen by casting. However, the fragility of optical fibers presents a challenge to the direct use of UAM, UW, and SLM, and the bare fiber must be covered with a metal coating or protective sleeve prior to these techniques. Techniques such as magnetron sputtering, flash evaporation, and electroless plating can deposit a thin layer of metal on the fiber surface with good adhesion, but the coating thickness, which ranges from a few hundred nanometers to a few microns, is not sufficient to provide adequate protection for the optical fiber. For a strain sensor, such a thin coating may not effectively transfer the strain on the metal to the optical sensor.
Electroplating is a widely used surface treatment technology with advantages of low cost, simplicity and ease. The process can be performed at low temperatures to avoid thermal residual stresses. To facilitate electroplating, a conductive metal layer must be deposited on the bare fiber, which is inherently non-conductive [24]. To accomplish this, Sandlin $et$ $al.$ [25] used electroless plating by mixing Torrent reagent and glucose solution and soaking the optical fibers to deposit a silver coating ($\sim$1 $\mu$m) followed by nickel plating ($\sim$25 $\mu$m). On the other hand, Tu $et$ $al.$ [26,27] fabricated Ti-Ag-Ni coated FBG sensors on SMFs and B-Ge co-doped fibers by combining magnetron sputtering and electroplating. Lupi $et$ $al.$ [28] magnetron sputtered a thin gold layer onto the fiber and then electroplated a Cu–Ni double metal layer onto it. Perry $et$ $al.$ [29] deposited a $\sim$1 $\mu$m chromium-gold bilayer on the fibers using evaporative deposition (ED) prior to electroplating a $\sim$100 $\mu$m nickel layer. Although methods like magnetron sputtering and evaporative deposition exist for depositing a metallic conductive layer, electroless plating has unique advantages of its low cost and ease of operation. In addition, it should be noted that previously reported metallized optical fiber sensors have primarily been used for single-parameter sensing, with limited reports of simultaneous temperature and strain measurement.
In this study, we present a novel Ag-Ni coated fiber sensor that combines an femtosecond FBG cascaded with an air bubble cavity FPI for simultaneous high temperature and strain sensing. Metallization is achieved by a combination of electroless plating and electroplating to deposit a conductive silver layer and a protective nickel layer, respectively. The surface metallization improved the temperature sensitivity of the FBG and FPI compared to the bare fiber, while the strain sensitivity remains largely unchanged. The proposed sensor demonstrated simultaneous temperature and strain measurements ranging from 0 to 1000 $\mu \varepsilon$ at 545 $^{\circ }$C, and decoupled the dual parameters of temperature and strain with low error.
2. Principle
2.1 Decoupling of temperature and strain
A schematic diagram of the proposed sensor is shown in Fig. 1. The sensor incorporates a short segment of periodically modulated refractive index, known as an fs-FBG, cascaded with an air bubble FPI fabricated by use of a commercial fusion splicer [30–32]. The exterior of the cascaded sensor is coated with a metal layer primarily for protection. The FBG and FPI obtain distinct differences in temperature and strain sensitivity, which form a characteristic matrix. By using matrix operation, temperature and strain could be measured simultaneously. The Bragg wavelength can be expressed mathematically as follows [33]
where $\lambda _B$ is the Bragg wavelength, $n_{eff}$ is the effective refractive index, and $\Lambda$ is the grating period. When the fiber is disturbed by temperature and strain, the Bragg wavelength $\lambda _B$ would shift with it. The Bragg wavelength shift $\Delta \lambda _B$ is given by [26,34]The sensing principle of the air bubble cavity FPI can be approximated as an ordinary in-line FPI. The air bubble cavity consists of the air and the two ends of the cavity are regarded as two FPI mirrors. The phase difference of two interference reflected light beam $\Delta \phi$ can be expressed as [35]
where $n_{air}$ $\approx$ 1.00 is the refractive index of the air bubble cavity, $L_{fp}$ is the cavity length of the air bubble cavity, $\lambda$ is the free-space wavelength of the reflected light. When the phase difference equals 2$m\pi$ (where $m$ is an integer), a certain dip in the spectrum can be observed. The corresponding $\lambda _m$ can be expressed as [35]The dip wavelength of FPI also shifts with the variation of external temperature and strain. The change in $\lambda _m$ can be expressed as
where ${K_{T}^{fp}}$ is the temperature sensitivity of the FPI, ${K_{\varepsilon }^{fp}}$ is the strain sensitivity of the FPI. In addition, the free spectral range (FSR) of FPI can be expressed as [36]FSR is the difference between two adjacent dip wavelengths, and ${\overline \lambda }$ is the mean value of adjacent dip wavelengths. The cavity length of the air-bubble cavity FPI ${L_{fp}}$ can be approximated from Eq. (7).
When the external temperature and strain change simultaneously, the reflection spectra of the air bubble cavity FPI and FBG will shift at the same time. Using the characteristic matrix, the change of temperature and strain can be expressed as [37]
The sensor was calibrated after the fabrication, so the characteristic matrix could be obtained. The change of temperature and strain can be obtained by observing the shift of $\lambda _B$ and $\lambda _m$ in the reflection spectrum.
2.2 Interactions between the fiber and the metal layer
Figure 2(a) shows a metal-coated monolayer fiber sensor structure. After thermal expansion, the overall fiber sensor is elongated. In reality, the elongation of the metal layer fiber subjected to the same temperature is given by [38]
Fig. 2. (a) Schematic diagram of axial strain of metal coated fiber; (b) cross section of metal coated fiber.
Equation (9) means there is no relative displacement between fiber and metal layer in reality. Where $L$ is the initial length of the fiber optic sensor, $\alpha _f$, $\alpha _m$ is the coefficient of thermal expansion of the silica and metal respectively, and $\Delta L_f$, $\Delta L_m$ is the elongation of the fiber and metal. $\varepsilon _{fzz}$ is the axial strain of the fiber stretched by the metal layer, because the free elongation of the fiber and metal exists differences due to different coefficients of thermal expansion. $\varepsilon _{mzz}$ is the axial strain of the metal. $\sigma _f$, $\sigma _m$ is the axial stress of the fiber and metal. $E _f$, $E _m$ is the Young’s module of the fiber and metal. The equilibrium force between the fiber and metal can be expressed
where ${A_f}$, $A_m$ is the cross-section area of the fiber and metal. From Eq. (9), (10), Eq. (11) could be obtained byIn cylindrical coordinates, according to the Lamé solutions, the the radial stress $\sigma _{mr}$, radial displacement $u _{mr}$, and the radial strain $\varepsilon _{mr}$ of the metal layer can be expressed as [38]
The continuous condition on the fiber-metal boundary can be expressed as
Thus, the internal pressure could be expressed as
From Eq. (13) and (15) the radial strain $\varepsilon _{frr}$, the transversal effect of radial strain $\varepsilon _{frz}$ and pressure temperature sensitivity coefficient of the fiber ${K _2}$ can be expressed as
Above all, the total Bragg wavelength shift $\Delta {{\lambda }_{B}}$ caused by temperature and thermal strain of the metal-coated fiber model could be expressed as
From Eq. (17), the total temperature sensitivity coefficient ${K}_{T}$, the total axial strain of the fiber $\varepsilon _{z}$, the total radial strain of the fiber $\varepsilon _{r}$, the temperature sensitivity coefficient of the bare fiber ${S}_{T}$, the axial strain sensitivity coefficient ${S}_{z}$ and the radial strain sensitivity coefficient ${S}_{r}$ could be obtained. The Bragg wavelength drift of a metal-coated FBG sensor exhibits an almost linear relationship with external temperature. The temperature sensitivity of the sensor is affected by several factors, including the coefficient of thermal expansion of the metal coating, Young’s modulus, Poisson’s ratio, fiber diameter, and thickness. Since thickness is the most easily varied parameter during fabrication, this study focuses on the effect of varying metal coating thickness on temperature sensitivity. A relationship between Ni metal coating thickness and temperature sensitivity of FBG was established using Matlab simulations, as shown in Fig. 3. The results show that the temperature sensitivity increases rapidly at low thicknesses, but tends to reach saturation when the thickness exceeds approximately 50 $\mu$m. From the above, the calculated temperature sensitivity and strain sensitivity of the bare FBG at 1550 nm is 10.46 pm/$^{\circ }$C and 1.20 pm/$^{\circ }$C, respectively.
3. Design of the metal-coating cascaded air bubble FPI-FBG
3.1 Fabrication of the cascaded air bubble FPI-FBG
The fabrication of air bubble cavity FPIs has been proposed by different methods [30–32]. In this work, we have adopted a practical and simple approach using a commercial fusion splicer. The process started with the preparation of an SMF by cleaving the end facet flat using an optical fiber cleaver (VT-20, Vintools). The prepared SMF was inserted into the fusion splicer (S179A, Furukawa Electric) and subjected to arc discharges, which transformed the end face into a hemispherical shape, effectively increasing the surface area of the fiber end. This pre-treatment was similarly applied to another SMF.
In subsequent steps, the hemispherical end facets of two SMFs were immersed in a refractive index matching fluid of low volatility. These immersed fibers were introduced into a commercial fusion splicer and subjected to a controlled arc discharge process. The process resulted in the evaporation of the fluid and rapid solidification of the silica material, ultimately leading to the formation of an air bubble cavity inside the optical fiber after fusion splicing.
To evaluate the sensing performance of the air bubble cavity FPI fabricated by this method, three samples with similar characteristics were fabricated. These samples were analyzed for morphological characteristics using a metallographic microscope (BX51M, Olympus). A circulator (FBY-1x2, Go-Fibereasy) was used to connect the spectrometer (OSA, AQ6370D, Yokogawa), the air bubble cavity FPI samples, and the light source (ASE Broadband Light Source, ASE-CL-100-T-B) for observing the reflection spectra of the samples.
The three fabricated air bubble cavity FPIs were designated as FP01, FP02, and FP03, respectively. Figure 4(a), (b) and (c) show the micrographs of the samples obtained by metallographic microscopy. The lengths of the air bubbles within FP01, FP02, and FP03 were measured to be approximately 56 $\mu$m, 85 $\mu$m, and 108 $\mu$m, respectively. Figure 4(d), (e) and (f) show the reflection spectra of the three air bubble cavity FPIs, respectively. The dip wavelengths marked in red, dip01, dip02, and dip03, were used as the characteristic wavelengths for sensing purposes. The FSRs were calculated to be 19.9 nm, 13.4 nm, and 10.3 nm, respectively, and the estimated cavity lengths were $\sim$60.9 $\mu$m, $\sim$89.9 $\mu$m, and $\sim$114.0 $\mu$m, which agreed with the cavity length measurements obtained through the microscope. It is possible that there is some discrepancy between the microscope measurements and the actual values. This discrepancy might result by measure error.
Fig. 4. (a), (b), (c) Microscope images of the air bubble cavity length of $\sim$56, $\sim$85, and $\sim$108$\mu$m, respectively; (d), (e), (f) the corresponding reflection spectra of the FPI.
To evaluate the changes in temperature and strain sensitivity of the air bubble cavity FPI before and after nickel plating, temperature and strain experiments were performed on each FPIs. The schematic diagrams of the test systems are shown in Fig. 5. The reflection spectrum of the sensors was continuously monitored by the spectrometer during the tests. For the strain experiments, two fiber clamps (OMHF03, RED STAR YANG TECHNOLOGY) were attached at 1 m intervals to each end of the fiber containing the air bubble cavity FPI, with one of the fiber clamps attached to a motorized displacement stage (EPSB100, RED STAR YANG TECHNOLOGY) capable of 40 $\mu$m resolution. The movement of the stage induced tensile strain on the fiber, and the characteristic dip wavelength shift was recorded as the tensile strain was incrementally increased from 0 to 1000 $\mu \varepsilon$ in 200 $\mu \varepsilon$ steps. The results showed that the dip wavelength of each bubble cavity FPIs shifted linearly toward longer wavelengths with increasing tensile strain, with strain sensitivities of 1.81 pm/$\mu \varepsilon$, 2.26 pm/$\mu \varepsilon$ and 3.32 pm/$\mu \varepsilon$ for FP01, FP02 and FP03, respectively. The three bubble cavity FPIs exhibit different strain sensitivities of the same order of magnitude. The difference strain sensitivity are related to the different cavity length [32]. The FP03 sample was selected for subsequent electroplating due to its highest strain sensitivity.
Fig. 5. (a) Schemetic diagram of the air bubble FPI strain test system; (b) schemetic diagram of the air bubble FPI temperature test system.
Temperature experiments were performed on FP01 and FP02 by exposing them to a muffle furnace (Thermconcept, ROHT 40/200/18) at setting temperatures ranging from $\sim$100 to $\sim$500 $^{\circ }$C with steps of $\sim$100 $^{\circ }$C, each step maintaining for 30 minutes to ensure uniform heating. To account for the inaccuracies of the furnace, the actual temperature inside was calibrated using an external thermocouple. Due to the fact that high temperature treatment can reduce the strength of the fiber and affect the subsequent plating process, FP03 was not selected for the high temperature testing. The temperature sensitivities of FP01 and FP02 were determined to be 1.06 pm/$^{\circ }$C and 1.10 pm/$^{\circ }$C, respectively. We believed that FP03 has a similar and low temperature sensitivity like these two bubble cavity FPIs. Similarly, temperature and strain sensitivity testing were performed on a bare femtosecond FBG (XUESEN Technology) with a center wavelength of $\sim$1550 nm to be cascaded with the FP03. The temperature sensitivity of the FBG was 10.60 pm/$^{\circ }$C and strain sensitivity was 0.96 pm/$\mu \varepsilon$, which were comparable to the temperature sensitivity of 9.02 pm/$^{\circ }$C and strain sensitivity of 1.01 pm/$\mu \varepsilon$ of the FBG sensor reported in [39]. The test results of the fiber sensor before electroplating were summarized in Fig. 6.
Fig. 6. Test results of the bare FPI and FBG before eletroplating: (a) fitted result of the wavelength of FP03 at different strains; (b) fitted result of the wavelength of FP01 at different temperature; (c) fitted result of the wavelength of FBG at different strain; (d) fitted result of the wavelength of FBG at different temperature.
3.2 Surface metalization of the fiber
We used a fiber fusion splicer to fuse the FP03 sample to a bare FBG with a distance of 1 cm to form the cascaded sensor. The metallization process of the fiber surface was divided into two main steps: 1) chemical silver plating and 2) nickel plating. Prior to chemical silver plating, some pre-treatments were required for the cascaded sensor. First, the cascaded sensor was immersed in acetone for 15 minutes to remove the acrylate coating. Second, it was washed in deionized water, followed by soaking in a 10 % NaOH solution to remove oil, and washed in deionized water again. Third, the fiber was immersed in a SnCl$_2$ solution to sensitized it for reacting with silver ions in the next step. After pre-treatments, the cascaded sensor was immersed in the silver plating reagent (at 20 $^{\circ }$C for 30 minutes), the chemical composition of which is shown in Table 1. This resulted in the deposition of a thin Ag layer on the fiber surface.
Table 1. Composition of Chemical Silver Plating Reagent
After chemical silver plating, the silver-coated fiber was washed with deionized water and inserted in a copper tube with an inner diameter of 500 $\mu$m and an outer diameter of 1 mm, for better connecting to the power supply. The positive terminal of a power supply (UTP1306S, UNI-T) was connected to the nickel plate by a wire, while the negative terminal was connected to the copper tube. The nickel plate and silver-coated fiber were then placed in an electrolytic cell filled with the nickel plating solution, the composition of which is given in Table 2. The current was adjusted to control the deposition rate over the course of 1 hour.
Table 2. Composition of Nickle Plating Solution
Figure 7(a) shows a microscopic observation of the electroplated FP03, with a thickness of the silver-nickel coated optical fiber of $\sim$131 $\mu$m, which means the coating is $\sim$3 $\mu$m thick (the diameter of the bare fiber is 125 $\mu$m). The shape of the FP03 air bubble cavity after electroplating remains similar to a bare one, as shown in Fig. 4(c). The electroplated area spans approximately 5 cm length. The electroplated sensor were evaluated by using X-ray fluorescence (XRF) spectrometry (ScopeX 980, LANScientific), the result is shown in Fig. 7(b), indicating that the metal layer has high purity of nickle.
Fig. 7. (a) Microscope image of metal-coated FP03 after plating; (b) the XRF spectrum of the metal layer surface.
After connecting the electroplated sensor to the spectrometer, the reflectance spectrum was analyzed as shown in Fig. 8. A comparison of the reflectance spectra before and after electroplating at $\sim$20 $^{\circ }$C reveals a blue shift, with the center wavelength of the FBG drifting from 1550.00 nm to 1549.62 nm, a decrease of approximately 0.38 nm. Further examination of the spectrum shows that each dip of FP03 has a different degree of blue shift. After electroplating, the wavelength of dip03 was found to be 1590.61 nm, exhibiting a blue shift of approximately 0.71 nm compared to its pre-electroplating wavelength of 1591.32 nm. The observed blue shift of the entire spectrum can be attributed to the axial compression caused by the nickel layer during the plating process, as previously reported in [16] and [19]. The relatively small blue shift observed in this study is likely due to the thin plating thickness. In summary, the overall spectrum of the cascaded sensor did not show dramatic deformation, indicating that the residual stress generated during the plating process is evenly distributed along the optical fiber within the plating area, resulting in a shaped spectrum that is suitable enough for high temperature strain measurements.
4. Experiments and results
4.1 Temperature testing
The temperature test of the cascaded sensor used the same experimental equipment and procedure as the temperature test of the FP03 and FBG. The sensor was heated incrementally from 99 to 552 $^{\circ }$C with steps of $\sim$100 $^{\circ }$C using the setup as mentioned in Section 3.1. As shown in Fig. 9(a), at the same temperature change, the red shift of the dip03 wavelength is smaller than that of the FBG center wavelength. This different wavelength shift can be attributed to the difference in temperature sensitivity between FBG and FP03, which was previously demonstrated in the temperature experiments of bare fibers.
Fig. 9. (a) Reflection spectra of the cascaded sensor at 99$^{\circ }$C and 552 $^{\circ }$C; (b) fitted result of the wavelength at different temperature of the metal-coated FBG and FP03 during heating and cooling.
The results of the temperature test on the cascaded sensor are shown in Fig. 9(b). It was observed that the temperature sensitivity of the Ni-coated FBG was 14.92 pm/$^{\circ }$C, which was higher than the bare fiber FBG of 10.60 pm/$^{\circ }$C. As the theory analysis mentioned above, the temperature sensitivity of FBG increases with the thickness of the Ni coating. The calculated temperature sensitivity of FBG of 12.00 pm/$^{\circ }$C with a 3 $\mu$m Ni coating. The discrepancy might be due to the observation error of thickness. A small observation error in thickness can obviously effect the calculated temperature sensitivity at low thicknesses as mentioned in Fig. 3. In addition, Fig. 9(b) showed the temperature response during the heating and cooling process with little difference in temperature sensitivity. The temperature sensitivity of the FP03 after nickel coating was 2.67 pm/$^{\circ }$C. Compared to 1.06 pm/$^{\circ }$C for the bare FP01 (we believed FP01, FP02 and FP03 have similar temperature sensitivities), the temperature sensitivity was improved by approximately 2.4 times after nickel plating. This improvement was attributed to the change in the coefficient of thermal expansion of the sensor after surface metalized, which increased the temperature sensitivity of both the FBG and FP03 in the cascaded sensor.
4.2 Strain testing under high temperature
The high-temperature strain test was performed on the cascaded sensor by securing its two ends with clamps and glue, with the sensing part placed in the heating zone of a furnace. The schematic diagram of the test system is shown in Fig. 10. The spacing between two clamps was 1 m, and the test apparatus were the same as mentioned in Section 3.1. The Young’s modulus of the metal-coated fiber increased slightly after nickel plating. However, due to the thin coating with only $\sim$3 $\mu$m, the increase of the Young’s modulus could be neglected. We assumed that the strain applied to the entire fiber is uniform during the stretching process. The temperature inside the furnace was gradually increased from 98 to 545 $^{\circ }$C in increments of $\sim$100 $^{\circ }$C and held at each temperature for 30 minutes. During the period when the temperature inside the furnace was stabilized, the electrical displacement stage was adjusted to apply a strain from 0 to 1000 $\mu \varepsilon$ with a step of 200 $\mu \varepsilon$. The strain response of the sensor at 545 $^{\circ }$C is shown in Fig. 11. The result shows that with increasing strain, the center wavelength of the FBG and the dip wavelength of the FP03 experience a red shift. The wavelength redshift of the dip03 is observed to be larger than that of the FBG center wavelength due to the higher strain sensitivity of the FP03. At 545 $^{\circ }$C, the strain sensitivity of FP03 and FBG is 3.42 pm/$\mu \varepsilon$ and 0.93 pm/$\mu \varepsilon$, respectively.
Fig. 11. Reflection spectra of the cascaded sensor in a 0-1000 $\mu \varepsilon$ strain testing at 545 $^{\circ }$C.
Figure 12 shows the strain sensitivities of the FBG after nickel plating. The sensitivity at 98 $^{\circ }$C, 218 $^{\circ }$C, 337 $^{\circ }$C, 446 $^{\circ }$C, and 545 $^{\circ }$C was measured to be 0.80 pm/$\mu \varepsilon$, 0.88 pm/$\mu \varepsilon$, 1.11 pm/$\mu \varepsilon$, 1.07 pm/$\mu \varepsilon$, and 0.93 pm/$\mu \varepsilon$, respectively. The results show a small fluctuation range between 0.80 pm/$\mu \varepsilon$ and 1.11 pm/$\mu \varepsilon$ at various temperature. Compared to the bare FBG before nickel plating with a sensitivity of 0.96 pm/$\mu \varepsilon$ mentioned in Fig. 6, the strain sensitivities under high temperature did not change significantly and remained within the fluctuation range.
Similarly, Fig. 13 shows the strain sensitivities of the FP03 after nickel plating. The sensitivity was measured to be 2.93 pm/$\mu \varepsilon$, 2.98 pm/$\mu \varepsilon$, 3.33 pm/$\mu \varepsilon$, 3.57 pm/$\mu \varepsilon$, and 3.42 pm/$\mu \varepsilon$ at 98 $^{\circ }$C, 218 $^{\circ }$C, 337 $^{\circ }$C, 446 $^{\circ }$C, and 545 $^{\circ }$C, respectively. These results show a small range of variation between 2.93 pm/$\mu \varepsilon$ and 3.57 pm/$\mu \varepsilon$. The strain sensitivity of the FP03 before nickel plating is 3.32 pm/$\mu \varepsilon$, which is within this range. It can be concluded that there is no significant difference in the strain sensitivity of FBG and FP03 before and after nickel plating, which may be due to the thinness of the nickel coating.
4.3 Dual parameters decoupling testing
We evaluated the performance and accuracy of the metal-coated cascaded sensor under high-temperature strain conditions. We performed strain tests with 200 $\mu \varepsilon$ increments from 0 to 1000 $\mu \varepsilon$ at three different temperatures (220 $^{\circ }$C, 314 $^{\circ }$C, and 429 $^{\circ }$C) using the same equipment and methods as described above. The temperature and strain values could be obtained by matrix operations. Due to different strain sensitivities of the FBG and FP03 corresponding to different temperatures, we divided the components into five different characteristic matrices, each corresponding to a different temperature range (0-100 $^{\circ }$C, 100-200 $^{\circ }$C, 200-300 $^{\circ }$C, 300-400 $^{\circ }$C, 400-500 $^{\circ }$C).
Three different characteristic matrices were used to calculate the temperature and strain values at 220 $^{\circ }$C, 314 $^{\circ }$C, and 429 $^{\circ }$C. The parameters ${K_{T}^B}$ and ${K_{T}^{fp}}$ remained constant within these matrices, with only ${K_{\varepsilon }^B}$ and ${K_{\varepsilon }^{fp}}$ varying. The comparison between the measured and reference values is shown in Fig. 14, where the reference values are shown as dash lines and the measured values are shown as squares. The dual parameter decoupling tests showed a strain error of approximately $\pm$ 50 $\mu \varepsilon$ and a temperature error of approximately $\pm$ 20 $^{\circ }$C. The error may be due to variations in the strain response of the metal-coated fiber at different temperatures, resulting in errors in the measured values, or to the increased temperature sensitivity of the metal-coated FPI compared to the bare one, which may have increased the cross-sensitivity between temperature and strain.
Fig. 14. Reference values (dash lines) and measured values (squares) obtained from the cascaded sensor at different temperatures and strains.
5. Conclusion
A novel high-temperature strain fiber sensor consisting of a silver-nickel metal-coated FBG-FPI cascaded structure has been developed and fabricated. The metal coating provides improved durability and longevity under harsh high temperature conditions compared to bare fiber sensors. The cascaded structure of FBG and FPI plays a key role in the dual parameter decoupling. The temperature sensitivity of both the FBG and air bubble cavity FPI increased to some extent after electroplating, while the strain sensitivity showed some fluctuation over different temperatures but remained largely unchanged. The effectiveness of decoupling high temperature and strain was demonstrated by subsequent dual parameter tests.
The use of a surface metalized fiber facilitates the integration of the fiber optic sensor with other critical metal components, thereby expanding the range of practical applications for fiber optic sensors in in-situ health monitoring. In addition, the metal-coated cascaded sensor holds promise for other applications, including pressure and vibration sensing in harsh high-temperature environments, and quasi-distributed high-temperature strain sensing using multiple metal-coated fibers.
Funding
National Natural Science Foundation of China (62275269); National Key Research and Development Program of China (2022YFF0706005); China Guangdong Guangxi Joint Science Key Foundation (2021GXNSFDA076001); Guangxi Major Projects of Science and Technology (Grant No.2020AA21077007); Interdisciplinary Scientific Research Foundation of Guangxi University (Grant No.2022JCC014).
Acknowledgments
The authors would like to thank Wenjie Xu, Qiang Bian, Shumao Zhang and for their support during the experiments.
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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