Integrated optic current transducers incorporating photonic crystal fiber for reduced temperature dependence

Woo-Sung Chu; Sung-Moon Kim; Min-Cheol Oh

doi:10.1364/OE.23.022816

1. Introduction

Optical current transducers (OCT) have attracted significant attention from the power industry because of their outstanding features relative to their electrical counterpart [1–3]. Owing to the significant investigation of optical devices for optical communication technology, the potential for various optical sensors is indeed higher than ever and is near commercial penetration [4, 5].

An optical current sensor based on a unique integrated optic chip made of a polymer waveguide was recently demonstrated and consists of various integrated optical devices such as a polarizer, power divider, phase modulator, polarization convertor, and 3-dB coupler [6]. Integrated optics could replace their discrete fiber-optic counterparts owing to the merits of a small footprint and potential for mass production. Furthermore, enhanced polarization controllability, reliability, and reproducibility are the major advantages of integrated optic sensors. Polymer waveguide devices would play a crucial role for the optical sensors because of their structural flexibility and facile fabrication steps. Unique flexible substrate devices enabled widely tunable lasers [7] and an optical waveguide touch panel [8]. For controlling the polarization of the guided mode in the polymeric integrated optic device, a polarization convertor [9], polarization controller [10], and polarization beam splitter [11] have been demonstrated. Otherwise, based on the structural flexibility, a large core waveguide providing a long collimation length [12] and an array device with extremely low crosstalk [13] were also proposed.

Although the integrated optic device provided excellent capability for the manipulation of light signals, the OCT had a temperature dependence problem because of the birefringence of the sensor head. In this work, to reduce the temperature dependence, we incorporated a photonic crystal fiber (PCF), which has small temperature-dependent birefringence [14], as well as a highly birefringent (Hi-Bi) spun fiber [15]. The proposed sensor configuration exhibited significant improvement of temperature dependence compared to previous research.

2. Device configuration and operating principle

Optical current sensors based on polarization rotated reflection interferometry have excellent stability against external perturbations through the compensation of initial retardation between the two orthogonal polarizations [16, 17]. From the basic interferometry, to be suitable for polymeric waveguide devices, we have modified the device configuration by separating the reflected signal into another path as shown in Fig. 1 [6]. The two polarized waves are then interfered to produce an intensity-modulated signal proportional to the applied current signal. Although the reflected returned light does not propagate through the same path as the forward signal, the residual retardation due to the different path length could be negligible because the polymer waveguide has much less birefringence than the PM fiber. Moreover, the lengths of the waveguide for the forward and backward paths could be precisely adjustable in the integrated optic device.

Fig. 1 Schematic configuration of integrated optic current transducers consisting of the sensor head and the integrated optic device for optical signal processing.

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The integrated optic device receives a 45° inclined linear polarization from the SLED, then it excites fast and slow axis components on the PM fiber. In front of the sensing fiber coil, there is a quarter-wave plate (QWP) made of a photonic crystal fiber (or a PM fiber) whose optic axis is aligned to be 45° from the optic axis of the PM fiber. For the sensing fiber coil, instead of the annealed fiber, we used a spun highly birefringent PM fiber, which facilitates fiber coil preparation. In the sensing fiber coil, the two circular polarizations are the eigenmodes, and they accumulate phase retardation due to the magneto-optic effect. After reflecting in the mirror attached at the end of the fiber coil, the two circular polarizations propagate backward to pass through the QWP and convert back to the linear polarizations. The returned light with the retarded phase information is propagating toward output PDs after passing through the phase modulator, half-wave plate, polarizers, and directional coupler as shown in Fig. 1. A detailed explanation of device operation can be found in author’s previous publication [6, 18].

During the sensor operation, although the integrated optic device could stay in a controlled environment, the sensor head is exposed to an outdoor environment; thus, the temperature dependence of the sensor head becomes important. The most sensitive device to temperature is the QWP. A Jones matrix calculation is useful to derive the effect of QWP temperature sensitivity on the OCT response. The polarization of the reflected signal is represented as

{(\begin{matrix} E_{x} \\ E_{y} \end{matrix})}_{out} = M_{QWP}^{- 1} \cdot M_{FR}^{- 1} \cdot M_{R} \cdot M_{FR} \cdot M_{QWP} \cdot {(\begin{matrix} E_{x} \\ E_{y} \end{matrix})}_{in},

M_{QWP} = (\begin{matrix} \cos (\frac{π}{4} + \frac{δ}{2}) & isin (\frac{π}{4} + \frac{δ}{2}) \\ isin (\frac{π}{4} + \frac{δ}{2}) & \cos (\frac{π}{4} + \frac{δ}{2}) \end{matrix}),

M_{FR} = (\begin{matrix} \cos θ_{F} & \sin θ_{F} \\ - \sin θ_{F} & \cos θ_{F} \end{matrix}) .

Where

M_{Q W P}, M_{F R},

and

M_{R}

correspond to the Jones matrices for the QWP, Faraday rotator, and reflector, respectively.

M_{Q W P}^{- 1}

and

M_{F R}^{- 1}

are the inverse matrices of

M_{Q W P}^{}

and

M_{F R}

and represent the Jones matrices of the respective components for the backward propagating wave.

θ_{F}

is the phase retardation caused by the current-induced Faraday effect. is the temperature-induced retardation error on the QWP. For an input of a 45° inclined linear polarization, the output polarization will have a phase difference of

Δ ϕ_{R}

between the slow and fast axes [16].

Δ ϕ_{R} = arc \tan [2 \cos δ \sin 4 θ_{f} / ((1 + \cos^{2} δ) \cos 4 θ_{f} - \sin^{2} δ)]

≅ 4 θ_{f} / \cos δ, f o r θ_{f} ≪ 1.

When

δ

= 0,

Δ ϕ_{R}

= 4

θ_{F}

, and when

δ

≠ 0,

Δ ϕ_{R}

increases as shown in Fig. 2(a). Hence,

Δ ϕ_{R}

is affected by

δ

. When a panda-type PM fiber was used for the QWP, the

δ

could be 9.83° for

Δ T

= 100 °C, which is calculated from the temperature-dependent birefringence

d Δ n / d T = - 4.66 \times 10^{- 7} /

°C. Whereas, PCF has

d Δ n / d T = - 1.24 \times 10^{- 8} /

°C [19], then it will have

δ

= 0.28°. Hence, the PCF-QWP will have a temperature dependence an order of magnitude lower then PMF-QWP.

Fig. 2 The effect of QWP retardation error (a) on the amount of phase retardation of the reflected signal though the QWP and (b) on the sensor output signal detected by PD.

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Prior to the current-sensing experiment, to confirm the temperature dependence of the QWP alone, a fiber-optic mirror was attached to the end of the QWP without connecting the sensing fiber coil. The integrated optical device was then connected to the QWP. In this configuration, by operating the phase modulator of the integrated optic device, one could obtain an interference signal affected by the $δ$ of the QWP. The Jones matrix to obtain the output interference signal of the integrated optic device is as follows:

{(\begin{matrix} E_{x} \\ E_{y} \end{matrix})}_{out} = \frac{1}{2} M_{Pol} \cdot [M_{PM} + {iM}_{HWP}] \cdot M_{QWP}^{- 1} \cdot M_{R} \cdot M_{QWP} \cdot {(\begin{matrix} E_{x} \\ E_{y} \end{matrix})}_{in},

M_{PM} = e^{{iθ}_{PM}} \cdot (\begin{matrix} 1 & 0 \\ 0 & 1 \end{matrix}),

M_{HWP} = (\begin{matrix} 0 & i \\ i & 0 \end{matrix}),

M_{Pol} = (\begin{matrix} 1 & 0 \\ 0 & 0 \end{matrix}) .

M_{P o l}

,

M_{P M}

, and

M_{H W P}

are the Jones matrices of the polarizer, phase modulator (PM), and half-wave plate, respectively.

θ_{P M}

is the phase change induced by the thermo-optic phase modulator. For an input signal with

E_{x} = E_{y} = \frac{1}{\sqrt{2}}

, the final output signal of the sensor becomes

P_{o u t} = {| E_{o u t} |}^{2} = \frac{1}{2} \cos^{2} δ \cos^{2} (\frac{θ_{P M}}{2}) + \frac{1}{4} \sin^{2} δ .

In this result, the terms that do not contribute to the interference due to the limited coherence length of the SLED are omitted. For the phase modulation of a sinusoidal signal with a quadrature point bias,

θ_{P M} = a \cos (w t + π / 2)

. The amplitude of the output signal then becomes dependent on

δ

as shown in Fig. 2(b). Hence, the temperature-dependent retardation error

δ

could be examined from the output signal amplitude variation.

Another source of temperature dependence arises from the bending-induced birefringence of the fiber coil between the refractive indices along the radial and axial directions [20]. When the birefringence is imposed by the fiber bending, it varies depending on the temperature. This effect could be analyzed as a case of QWP retardation error and affect the scale factor of the sensor. To reduce the linear birefringence in previous research, the bare fiber wound on a ceramic jig was annealed at 850 °C to remove the residual stress [21]. However, this high-temperature annealing of the bare fiber was a time-consuming and low-yield process. Another method to remove the linear birefringence of the bent fiber is provided by spinning the PM fiber with initial linear birefringence. The spun PM fiber exhibited negligible linear birefringence even after the coiling of the fiber [22], and it will have highly reduced temperature dependence.

3. Fabrication of integrated optic device and sensing fiber coil

For the comparison of temperature dependence, two kinds of QWPs were prepared by using a panda PM fiber and a PCF, respectively. The QWPs were to be attached at a 45° angle with the optic axis of the PM fiber as outlined in Fig. 3. With a certain length of PM fiber left for quarter-wave phase retardation, the PMF (or PCF) was cleaved. A short length of single-mode fiber was then attached to produce the same effect as the single-mode fiber coil attachment. The lengths of QWP calculated from the birefringence were 0.92 mm and 1.00 mm for the PMF-QWP and PCF-QWP, respectively. The QWP performance was evaluated by measuring the polarization extinction ratio of the output light for the linear input polarization. Due to the stress relaxation by the splicing arc, the effective length of the QWP was shorter than the design. Among many samples with different QWP lengths, the PMF-QWP of 1.25 mm and the PCF-QWP of 0.90 mm exhibited the best output polarization extinction ratios. The retardation error was only 2.0° to 2.6° for the two cases. The PCF-QWP had a length close to the design value because it was spliced with much lower arc power than that of the PMF-QWP to prevent collapse of the air hole. Figure 4 shows the fabricated PCF-QWP spliced between PM fibers.

Fig. 3 The procedures to fabricate the fiber-optic QWP: (a) Panda PM fiber (PMF) or PCF spliced to another PM fiber with their optic axes aligned to be 45°, (b) cleaved to leave the length of QWP, (c) single mode fiber spliced, (d) cleaved for the measurement of QWP characterization, and (e) the completed fiber-optic QWP.

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Fig. 4 Photographs of the QWP device consisting of (a) the splice between the PMF and PCF and (b) the splice between the PCF and spun PM fiber with a smaller cladding diameter. The cut views of (c) PM fiber, (d) PCF, and (e) spun PM fiber are also shown.

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The integrated optic device used for producing an interference signal proportional to the Faraday effect was fabricated based on a well-established polymer waveguide technology. Fluorinated low-loss polymers from ChemOptics Co. were used to form the core and claddinglayers, whose refractive indices were 1.440 and 1.430, respectively. The waveguide had an inverted rib structure with a core dimension of 6.0 × 5.8 μm². For the insertion of the TE-pass polarizer to absorb TM polarization through the surface plasmon absorption, over the first upper cladding layer with a thickness of 1.8 μm, a 9 mm-long Cr-Au metal pattern was fabricated. After the second upper cladding formation, a heating electrode was fabricated for the thermo-optic phase modulator. A groove line was then formed by using a dicing saw for the insertion of a 20 μm-thick polyimide half-wave plate, which was fixed with a UV curable epoxy [23]. The schematic fabrication procedure is outlined in Fig. 5.

Fig. 5 Schematic fabrication procedures of an integrated optic current transducer chip made of polymer waveguide.

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4. Temperature dependence of fiber-optic QWPs and characterization of optical current sensors

To evaluate the temperature dependence of QWP before it is inserted to the complete sensor, a fiber mirror is attached on the right side of the fiber-optic QWP and the other side is connected to the proposed integrated optic device. As explained in section 2, by applying a modulation signal on the phase modulator, one can evaluate the phase retardation error of the QWP. The mirror-attached QWP was placed in a temperature-controlled oven, and the output signal amplitude was then monitored as the temperature was varied from 25 to 80 °C as shown in Fig. 6. The PMF-QWP exhibited a 1.8% change of output signal amplitude, which corresponds to a temperature-dependent retardation of 0.085°/°C, which is close to the design result of 0.098°/°C. In the PCF-QWP case, the output signal amplitude variation stayed withina range of ± 0.3% owing to the low temperature dependent retardation of 0.003°/°C. The birefringence of PMF-QWP is imposed by the stress members, while that of PCF-QWP is introduced by the asymmetric air holes surrounding the fiber core. Hence, the PCF-QWP has lower temperature dependence than the PMF-QWP.

Fig. 6 Temperature-dependent QWP retardation, δ effect on the output signal amplitude. The straight lines indicate the phase retardation error calculated from the temperature-dependent birefringence of the QWPs.

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For complete sensor head assembly, the two QWPs were spliced with each fiber sensing coil with a reflector spliced at the other end. The two types of sensors, one with PCF-QWP spliced to a Hi-Bi spun fiber (PCF-spun fiber sensor) and the other with PMF-QWP spliced to an annealed fiber coil (PMF-annealed fiber sensor), were prepared for the temperature dependence experiment. The Hi-Bi spun fiber had a circular beat length of 72 mm (@ 1550 nm) and a spin pitch of 4.8 mm.

The sensing coil used for temperature dependence measurement was 10.5 turn with a diameter of 90 mm. The sensor head was placed in a temperature-controlled oven with a conducting wire that sent 480 A(rms) through the sensing coil. For temperature variations greater than 70 °C, the sensor signal amplitude was monitored as shown in Fig. 7. The PMF-annealed fiber sensor exhibited output signal amplitude variation greater than 7% of which the effect of the Verdet constant variation accounted for 0.35%, which is calculated from temperature dependent coefficient of the Verdet constant [ $(1 / V) d V / d T \approx 0.7 \times 10^{- 4}$ °C] [17], and the remaining dependence resulted from the retardation error of the PCF-QWP. The PCF-spun fiber sensor exhibited signal variation within ± 1% when the temperature increased to 78 °C; this is only 1/7 of the PMF-annealed fiber sensor.

Fig. 7 Temperature dependence of the IOCT consisting of (a) a PMF-QWP and an annealed fiber, and (b) a PCF-QWP and a spun PM fiber exhibiting temperature dependence less than ± 1%.

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For a special installation of the current sensor enclosing a conductor with a large diameter, we prepared the PCF-spun fiber sensor with a sensing coil of 5.5 turns and a diameter of 1160 mm. On a test bed, a toroid-type current loop was used to amplify the magnetic field. By applying 0.5 to 12 kA(rms), we measured the sensor signal as shown in Fig. 8. The two points with black and empty dots were measured during the increase and decrease of the applied current. The relative error between the applied current and the sensor output was within ± 0.5%, and the result satisfied the standard of 0.5 accuracy class current sensors (IEC 60044-8).

Fig. 8 Output response of the IOCT sensor consisting of a PCF-QWP and a spun PM fiber in which the sensing error is within ± 0.5% and satisfies the standard of 0.5 accuracy class (IEC 60044-8).

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5. Conclusion

Fiber-optic current sensors with low temperature dependence were demonstrated by incorporating a PCF and a spun PM fiber. The scale factor change due to the temperature dependent retarder was analyzed by using a Jones matrix calculation, and experimental results were compared to the design. The PCF-QWP spliced spun fiber sensor head was connected to the integrated-optic phase analyzer device for detecting the current-induced phase change. Compared to a reference sensor consisting of the PMF-QWP and the annealed fiber, the PCF-incorporated sensor exhibited significantly reduced temperature dependence. The PCF-incorporated sensor satisfied the current transducer standard of 0.5 accuracy class (IEC 60044-8).

Acknowledgments

This work was supported by the National Research Foundation of Korea (NRF) grant funded by the Korea government (MSIP) (2014R1A2A1A10051994).

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Integrated optic current transducers incorporating photonic crystal fiber for reduced temperature dependence

Abstract

1. Introduction

2. Device configuration and operating principle

3. Fabrication of integrated optic device and sensing fiber coil

4. Temperature dependence of fiber-optic QWPs and characterization of optical current sensors

5. Conclusion

Acknowledgments

References and links

Cited By

Figures (8)

Equations (10)

Optics Express