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Abstract
Optical phased array (OPA) is a useful device for achieving the solid-state beam scanner required in compact light detection and ranging. However, conventional OPAs actively control the phase difference between arrays. Therefore, power consumption is extremely high in a high-resolution OPA. Herein, we fabricated a passive OPA with a 128-channel silicon arrayed waveguide and Si-dot grating antennas with large apertures. Moreover, we integrated a hybrid wavelength-tunable laser diode with a passive OPA. The field of view was 43.9° × 10.4°, and the FWHM of the beam width was 0.233° × 0.0495°. The power consumption per antenna was 0.397 mW.
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1. Introduction
Recently, the development of optical sensing technologies for automated driving has increased significantly. To achieve high-level automated driving, accurate detection of distant obstacles is required. Therefore, light detection and ranging (LiDAR) technology has attracted considerable attention. LiDAR detects obstacles by illuminating them and receiving reflected light. In general, two methods are used to irradiate obstacles with a laser light in a LiDAR. The first is the flash method, and the second is the beam scanning method. For two reasons, beam-scanning LiDAR outperforms flash LiDAR in measuring the distance between long-range obstacles. First, the reflected light intensity is high because a focused laser beam is used. Second, the beam-scanning method can be combined with the frequency-modulated continuous-wave (FMCW) method, which enables highly accurate distance measurements without being affected by sunlight. The simplest beam-scanning method uses a mechanically rotated laser light source or a polygon mirror. However, mechanical beam scanners are too large to be used in cars. Therefore, the solid-state beam scanner must be realized [1–7].
The optical phased array (OPA) using optical integrated circuit technology is a promising device for solid-state beam scanner [3,4,8–10]. There are three types of OPA structures. The first is 2D-OPA consisting of active phase shifters and a two-dimensional antenna array [10–14]. The second is 1D-OPA consisting of active phase shifters and a one-dimensional waveguide grating antenna (WGA) array [9,15–24]. The third is a passive OPA consisting of a WGA array and a passive phase shifter with a waveguide delay line [8,25–32]. To reduce the divergence of the emitted beam, a wide aperture area is required. To achieve this goal, the number of antennas must be increased. However, in 2D- and 1D-OPA, large power consumption is a problem because they require an active phase shifter in each antenna [9–24]. Therefore, passive OPA are attracting attention owing to their ability to reduce power consumption. Silicon and silicon nitride are the primarily materials used in passive OPA. With silicon nitride, it is relatively easy to fabricate a passive OPA using several antennas. However, achieving a wide field of view (FOV) is difficult because the antenna pitch cannot be reduced [28–30]. In contrast, silicon can realize a wide FOV using a high-density antenna array because of the strong light confinement in the waveguide [33,34]. In addition, because silicon photonics is based on a CMOS-compatible fabrication process, silicon-passive OPA are suitable for mass production. However, increasing the number of antennas in silicon passive OPA is difficult because the effect of manufacturing errors is large [25,26]. We previously proposed and demonstrated beam scanning using a passive silicon OPA with a low manufacturing error effect [27,31,32]. In a previous study, we realized a 64-channel silicon passive OPA and the FWHM of the beam width was 0.517° × 3.67° [31].
In this study, we fabricated a silicon-based high-resolution passive OPA consisting of 128-channel arrayed waveguide grating (AWG) and WGA with long apertures. We reduced the FWHM of the beam width to 1/2 in the horizontal direction and 1/100 in the vertical direction compared to that in our previous study. Additionally, we integrated a hybrid wavelength-tunable laser diode as the light source for OPA [35–41]. A fully integrated two-dimensional beam scanner with high resolution and low power consumption was realized.
2. Operation principle and design of the device
2.1 Operation principle of passive OPA
Figure 1 shows a schematic of the designed passive OPA.
It consists of a tree branch with 1 × 2 multi-mode interference (MMI) couplers, 128-channel AWG, and WGA with a long aperture length. In the AWG, there is a path length difference, $\Delta L$, between adjacent waveguides. This causes a phase difference based on the wavelength $\lambda $ of the incident light [42]. Therefore, the beam emission angle in the array direction ${\theta _x}$ can be calculated as follows:
where ${n_{eff,AWG}}$ represents the effective refractive index of the waveguide in the AWG, d represents the pitch of the antenna, and m represents the diffraction order in the ${\theta _x}$ direction. Moreover, the pitch in WGA $\mathrm{\Lambda }$ causes a phase difference based on $\lambda $. Therefore, the beam emission angle in the antenna direction ${\theta _y}$ can be calculated as follows: where ${n_{eff,WGA}}$ represents the effective refractive index in WGA and n represents the diffraction order in the ${\theta _y}$ direction. As shown in Eqs. (1) and (2), both ${\theta _x}$ and ${\theta _y}$ depend on $\lambda $. Therefore, two-dimensional beam scanning was performed using a wavelength sweep.In ${\theta _x}$ direction, there were some beams because of adjacent diffraction order ($m - 1$ or $m + 1$). This restricts FOV in ${\theta _x}$ direction ($FO{V_x}$). $FO{V_x}$ can be calculated as follows:
Therefore, a high-density antenna array is required to achieve a wide $FO{V_x}$. In ${\theta _y}$ direction, Eq. (2) usually does not hold by $n - 1$ or $n + 1$. Therefore, FOV in ${\theta _y}$ direction ($FO{V_y}$) is decided by the wavelength sweep range of the light source and the sweep efficiency (°/nm) of WGA.
2.2 Arrayed waveguide grating with low phase error
The number of antennas should be increased to increase the aperture length in the ${\theta _x}$ direction. When designing large-scale AWG with numerous antennas, it is important to consider manufacturing errors. The variation of waveguide width by manufacturing error causes index error $\mathrm{\Delta }{n_{eff,AWG}}$ because ${n_{eff,AWG}}$ depends on waveguide width [43]. This index error causes phase error $\mathrm{\Delta }\phi $ in the light diffracted from WGA and $\mathrm{\Delta }\phi $ disturbs beam formation. $\mathrm{\Delta }\phi $ can be calculated as follows:
In our previous study, we identified that the standard deviation of $\mathrm{\Delta }{n_{eff,AWG}}$ was reduced to 4.49 × 10−5 in a 3.2-µm-wide waveguide [27]. However, $\mathrm{\Delta }{n_{eff,AWG}}$ was not suppressed considerably in waveguides with width of over 3.2 µm because height error was more dominant than width error. In this study, we reduced $\mathrm{\Delta }L$ for further reduction of $\mathrm{\Delta }\phi $. Figure 3 shows the calculation results of beamforming in a 128-channel passive OPA. For $\mathrm{\Delta }L = \; $ 80 µm, the beam intensity decreased to 38.0%. In contrast, 85.3% of the beam intensity was maintained at $\mathrm{\Delta }L = \; $ 20 µm. Reducing $\Delta L$ results in the coarse resolution in ${\theta _y}$ direction because the number of scan line decreases. However, in autonomous driving, it is not important to accurately know the height of an obstacle. Therefore, we prioritized reducing the phase error in the AWG.
2.3 Waveguide grating antenna with long aperture length
Increasing the length of WGA is required for a long aperture length in the ${\theta _y}$ direction. In the design of long WGA, a small grating strength $\alpha $ is required. $\alpha $ is defined as follows [44]:
where z denotes the propagation length in the WGA, ${P_0}$ denotes the intensity of incident light in the WGA, and $P(z )$ denotes the remaining optical power in the waveguide. Figure 4 shows the calculation results for $P(z )$ for the three grating strengths. If the grating strength is too high, $P(z )$ converges to zero at a short propagation length, and the effective aperture length is reduced.
Fig. 4. Calculation results of relationship between propagation length and remaining optical power for each grating strength
Figure 5 shows the relationship between grating strength and the FWHM of beam width in the ${\theta _y}$ direction. Grating strength should be suppressed to less than 3.5 mm-1 to realize a FWHM of approximately 0.1°.
Fig. 5. Calculation result of relationship between grating strength and FWHM in ${\theta _y}$ direction
Figure 6 shows examples of WGAs with low grating strengths [45,46]. In a sharrow etching grating, ${d_e}$ denotes the depth of etching. In the Si-dot grating, g denotes the gap between the waveguide and the silicon dot, and ${w_s}$ denotes the width of the silicon dot. The duty cycle is defined as $a/\mathrm{\Lambda }$ in both gratings.
Fig. 6. Schematic of WGA with low grating strength: (a) 3D structure of sharrow etching grating; (b) sharrow etching grating observed from side; (c) 3D structure of Si-dot grating; (d) Si-dot grating observed from above
Figure 7 shows the calculation results for $P(z )$ for the sharrow and Si-dot gratings. The red data represent the results for the Si-dot grating, and the blue data represent the results for the sharrow etching grating. In sharrow etching grating, $\mathrm{\Lambda }$ is 900 nm. In Si-dot grating, $\mathrm{\Lambda }$ is 780 nm and g is 200 nm. In both gratings, the waveguide width is 440 nm, and the duty cycle is 0.5. For comparison, the calculation result of half-etching grating with 80.4-mm-1 grating strength is also shown in Fig. 7. In sharrow etching grating, grating strength is 0.701 mm-1 for ${d_e} = \; $ 10 nm and 1.60 mm-1 for ${d_e} = \; $ 15 nm. A 5-nm change in ${d_e}$ causes a large change in the grating strength. This result indicates that manufacturing errors significantly reduce the effective aperture length. In Si-dot grating, grating strength is 1.41 mm-1 for ${w_s} = \; $ 200 nm and 1.54 mm-1 for ${w_s} = \; $ 205 nm. Therefore, the change in the effective aperture length owing to manufacturing errors is small in the Si-dot grating.
Fig. 7. Calculation results of relationship between propagation length and remaining optical power in each WGA
Thus, the Si-dot grating realizes both a long aperture length and a high manufacturing tolerance. The combination of a Si-dot grating and an AWG consisting of a multimode waveguide results in a Si passive OPA with narrow beam divergence and high manufacturing tolerance.
2.4 Hybrid wavelength tunable laser diode
Figure 8 shows a schematic of a hybrid wavelength-tunable laser diode [35,41]. This laser diode consists of a semiconductor optical amplifier (SOA) chip and a wavelength-tunable filter using silicon photonics. The facets of the SOA chip connected to the silicon photonic chip have an antireflection coating. The other facet of the SOA chip has a 6% reflection mirror with a dielectric multilayer film. The combination of a waveguide mirror on a silicon photonic chip and an SOA facet mirror constitutes a Fabry–Perot interferometer.
The reflectance of the waveguide mirror can be tuned by adjusting the input power of the heater. As shown in Fig. 9, this mirror can be considered a Mach–Zehnder interferometer (MZI) optical switch. The directional coupler in the MZI has a power-coupling efficiency of 0.5. The incident light is divided into bar and cross ports. The output power in the bar and cross ports can be considered as reflected and transmitted light, respectively. The ratio of the output power in the two ports is based on the phase difference between the propagating light in each arm, which can be controlled using the thermo-optic effect. Therefore, reflectance can be tuned using a mirror heater.
The reflectance of waveguide mirror ${R_{mirror}}$ can be calculated as follows:
A wavelength-tunable filter is a serial connection of two ring resonators with different circumferences [35,38,41]. The free spectral range (FSR) of this filter is wide because of the Vernier effect. The filtering characteristics of the ring resonator can be controlled using heaters placed on the ring resonators. Thus, the designed laser diode can achieve arbitrary single-mode oscillations over a wide wavelength range. The FSR of the ring resonators was 939 GHz for ring1 and 846 GHz for ring2. Therefore, the wavelength-tunable range of the designed laser diode was 68.5 nm. The Q factor of ring resonator was estimated to be 2.72 × 103 for ring1 and 3.11 × 103 for ring2. The linewidth of the laser was estimated to be several tens kHz [38]. The micro heater on the bus waveguide controls the longitudinal mode of the hybrid laser cavity.
2.5 Simulation of beam scanning
Figure 11 shows the calculation results of beam scanning using the designed beam scanner. Here, the number of antennas N was 128, pitch of antenna d was 2.07 µm, path length difference in AWG $\mathrm{\Delta }L$ was 20 µm, pitch in WGA $\mathrm{\Lambda }$ was 780 nm, duty cycle $a/\mathrm{\Lambda }$ was 0.5, width of Si dot ${w_s}$ was 200 nm, gap between waveguide and Si dot g was 200 nm, and length of the WGA was 3 mm. The wavelength of the incident light varied from 1515 to 1584 nm in the 1.5 nm step. FOV was 44.0° × 10.7° and FWHM of the beam width was 0.284° × 0.0414°. The sweep efficiency of the WGA was 0.156 °/nm along ${\theta _y}$ direction.
3. Fabrication
In this study, passive OPA and wavelength-tunable filters were fabricated in a multi-project silicon photonics foundry using a 300-mm silicon on insulator (SOI) wafer. The thickness of the SOI layer was 220 nm, and the buried oxide (BOX) layer was 3 µm. The Si waveguides and dots were patterned using ArF immersion lithography and reactive-ion etching. A TiN micro heater was fabricated on the SiO2 over cladding layer with a thickness of 1.2 µm.
After aligning the silicon photonic and SOA chips, they were edge-bonded using a UV-curing resin. Figure 12 shows a photograph of the fabricated beam scanner. The footprint of the beam scanner was 6 × 2.4 mm2.
4. Results and discussion
In the following experiments, the temperature of the beam scanner chip was set to 20 °C using a thermoelectric temperature controller. Figure 13 shows the measured relationship between the reflectance of the waveguide mirror and the power consumption of the mirror heater. The reflectance was estimated from the change in the threshold current. The phase-shift efficiency of the mirror heater was 21.5 mW/π. The reflectance with no temperature change was significantly different between the calculated and experimental values. This may have been caused by manufacturing errors in waveguide width. In the subsequent experiments, a power of 14.7 mW was applied to the mirror heater to achieve 60% reflectance in the waveguide mirror. The reflectance of SOA facet and waveguide mirror were selected to realize a stable laser operation.
Fig. 13. Measured relationship between power consumption and reflectance in reflectance tunable mirror
Figure 14 shows the measured relationship between SOA injection current and optical power above the beam scanner chip. The threshold current is 27.7 mA. The maximum optical power is 3.88 mW for the SOA injection current of 115 mA. The optimization of output light extraction method could realize higher output power [47,48]. From Fig. 14, the coupling loss between the SOA and silicon photonics chips is estimated to be 1.79 dB. In this estimation, we used the gain of the one-way SOA chip at the threshold current (9.55 dB), mirror loss of the waveguide mirror (2.22 dB), mirror loss of the SOA facet (12.2 dB), and the propagation loss of the one-way silicon photonic chip (0.549 dB). In the subsequent experiments, the SOA injection current was set to 90 mA.
Figure 15 shows the measured relationship between the lasing wavelength and power consumption in each ring heater. When the lasing wavelength returns to its original value, the power consumption is 52.2 mW in the ring1 heater and 49.5 mW in the ring2 heater. This is the required power consumption for a 2π phase shift in a ring resonator. Therefore, the phase shift efficiency is 26.1 mW/π in the ring1 heater and 24.7 mW/π in the ring2 heater. This result shows that the power consumption required for controlling the fabricated beam scanner is only 50.8 mW. Figure 16 shows the measured change in the wavelength spectrum by tuning the ring1 heater. Good single-mode oscillation is observed in the range of 1514.95–1582.25 nm. This range of 67.3 nm matches well with the design values. The smallest side mode suppression ratio (SMSR) is 42.8 dB. Figure 17 shows the measured fine wavelength tuning by controlling both the ring heaters. In the white area, two lasing wavelengths are observed at the same interval as the FSR of the wavelength filter. This result indicates that the fabricated laser diode can select a lasing wavelength almost continuously.
Figure 18 shows the measured far-field pattern (FFP) of the beam emitted from the passive OPA. The emitted beam was Fourier-transformed using an f-θ lens and two lenses (Hamamatsu Photonics A3267-12) and observed by an InGaAs infrared camera (Hamamatsu Photonics C14041-10U) (Figs. 18(a)–(c)). The FOV of this measurement system is approximately 68° × 54°, and the angular resolution is 0.3°. The FOV of the beam scanner is 43.9° × 10.4°, and the sweep efficiency of the WGA is 0.163 °/nm. These results match the design values well. The smallest peak-to-sidelobe ratio was 4.16 dB. The intensity of the side lobe and background may be increased by interference with the light reflected from the silicon substrate after downward diffraction in the WGA. This problem can be solved using the WGA, which is asymmetric in the vertical direction, to increase the amount of upward diffraction [49,50]. The distortion in the scan line have been caused by the distortion of the lenses. As shown in Fig. 18(b) and (c), the FWHM of the beam width appears wider than the designed value owing to the coarse angular resolution of the measurement system. Therefore, as shown in Fig. 18(d), the emitted beam is observed by direct injection into the infrared camera [51]. The angular resolution of the system is approximately 0.02°. The FWHM of beam width is 0.233° × 0.0495°. This result matches the design value well. The resolution of the beam scanning is 188 × 3.
Fig. 18. Measured beam scanning by fabricated beam scanner: (a) 2D profile of beam scanning; (b) 1D profile of beam scanning in ${\theta _x}$ direction; (c) 1D profile of beam scanning in ${\theta _y}$ direction; (d) observation of beam by direct injection into the camera (distance between camera and chip was 100 mm)
Table 1 compares the proposed passive OPA with previous passive OPAs. In previously reported silicon-passive OPAs, the number of antennas remained at approximately 30 because of the effect of manufacturing error. Our study realized a 128-channel silicon AWG and significantly increased the resolution of a passive silicon OPA. In silicon nitride, a 128-channel OPA had been demonstrated because the effect of manufacturing errors was small. However, realizing a high-density antenna array in silicon nitride was difficult because the width of the single-mode waveguide was several micrometers. Therefore, the silicon-passive OPA was superior in expanding the FOV. Thus, our study contributes to realizing a passive OPA with a wide FOV and high resolution.
Table 1. Comparison with Previous Passive OPA
Table 2 compares the proposed beam scanner with previous beam scanners that required active phase shifters for each antenna. Previously reported beam scanners required power consumption of several tens of milliwatts or several hundred milliwatts per antenna. By contrast, our beam scanner required only two ring heaters for beam scanning. Therefore, the power consumption per antenna was less than 0.5 mW.
Table 2. Comparison with OPA Requiring Phase Shifter in Each Antenna
5. Conclusions
In this study, we integrated a 128-channel silicon passive OPA and a hybrid wavelength-tunable laser diode. The footprint of the beam scanner was 6 × 2.4 mm. The wavelength-tunable range of the fabricated laser diode was 67.3 nm, and the SMSR exceeded 40 dB. We demonstrated beam scanning by controlling an integrated laser diode. The FOV was 43.9° × 10.4°, and FWHM was 0.233° × 0.0495°. This passive OPA was realized by combination of Si dot grating and AWG which is reduced its path length. It reduced the beam divergence to 1/200 compared with our previous work. As far as we know, our OPA has the largest number of WGA in Si passive OPA. The power consumption for the beam scanning was 50.8 mW, and the power consumption per antenna was only 0.397 mW. As shown by these data, our study achieved both high resolution and low power consumption. In addition, controlling this beam scanner was easy because only two microheaters were available for beam scanning. Our beam scanner could be fabricated using a CMOS-compatible fabrication process and bonding with a SOA chip. In addition, waveguide structure could be fabricated by only once etching process on the SOI platform. Therefore, our beam scanner is suitable for mass production. This could be a key device for realizing compact and low-power-consumption LiDARs.
Funding
Ministry of Internal Affairs and Communications (SCOPE, JP235003005); Japan Society for the Promotion of Science (23H01472); Adaptable and Seamless Technology Transfer Program through Target-Driven R and D (JPMJTR23RG); National Institute of Information and Communications Technology (JPJ012368C, 1301).
Acknowledgments
This study was partly supported by Japan Society for the Promotion of Science (KAKENHI 23H01472); Adaptable and Seamless Technology Transfer Program through Target-Driven R and D (JPMJTR23RG); National Institute of Information and Communications Technology (JPJ012368C, 1301); Ministry of Internal Affairs and Communications (SCOPE, JP235003005).
Disclosures
The authors declare that there are no conflicts of interest related to this article.
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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