1. INTRODUCTION

Integrated electro-optic modulators (EOMs) have recently experienced great interest being critical components in modern optical telecommunication networks [1], microwave photonic systems [2], and emerging quantum-photonic applications [3]. High-performance EOMs that can operate at high data rates while featuring low power consumption and satisfactory signal-to-noise ratios are essential for such applications. Several efforts have been devoted to realizing these target modulation specifications on-chip using various substrates [4–8]. Among them, lithium niobate (LN) has been the preferred candidate due to its superior electro-optic (EO) properties, high refractive index, and wide transparency range [8]. The platform has recently found its way to the integrated photonics area by introducing thin-film lithium niobate (TFLN) technology, capitalizing on advances in nanofabrication techniques [9]. Thanks to the compact and low-loss waveguides achievable on TFLN, the technology can combine the excellent material properties of bulk LN along with high integration capabilities.

Both non-resonant and resonant optical modulators have been successfully demonstrated on TFLN substrate [10–15]. Conventional Mach–Zehnder modulators (MZMs) could feature wide EO bandwidths, up to 100 GHz, and low on-chip optical loss, typically ${\lt}{0.3}\;{\rm dB/cm}$, while operating at CMOS-compatible voltages [10,11]. However, they require long phase shifters (${\sim}1\;{\rm cm} $) combined with a co-traveling modulation design to achieve the required phase modulation. On the other hand, resonant modulators based on microring resonators [12], photonic crystals [13], or phase-shifted Bragg grating resonators (BGRs) [14,15] could achieve a more compact footprint and lower power consumption. Nevertheless, resonator-based modulators have a low modulation bandwidth (${\lt}{30}\;{\rm GHz}$), limited by the cavity photon lifetime, and are very sensitive to fabrication and temperature variations due to their narrow optical bandwidth [9]. To address these issues, uniform waveguide Bragg grating (WBG) modulators have been proposed on TFLN, based on linear EO tuning of the WBG transmission spectrum [16,17]. A wide modulation bandwidth that is not limited by cavity photon lifetime can be achieved; moreover, the modulator footprint can be smaller than EO-MZMs. The WBG modulators were implemented on X-cut LN, where a straight WBG is modulated using two side electrodes to harness the EO effect toward the Z-axis of LN (${r_{33}} \approx 31\;{\rm pm}/{\rm V}$). Such configuration is most commonly used for wideband EO modulators depending on traveling wave design; however, it results in long WBG modulators, which are still unsuitable for highly integrated realizations of optical modulators.

Recently, spiral-shaped WBGs have been demonstrated on silicon [18,19] and silicon nitride [20] substrates in order to shrink the device footprint of long gratings, which are useful for highly selective optical filters and dispersion control applications. Such compact devices could not be realized on LN in the common X-cut configuration to benefit from its efficient EO tuning, achievable using LN Pockels effect, because the use of X-cut constrains the electrodes’ orientation. Hence, the tunability of spiral WBGs on other substrates was limited to other modulation techniques such as thermo-optic tuning [21] or free carrier modulation [22], which have much lower bandwidths and higher optical loss than EO modulation.

In this work, we utilize the Z-cut LN platform and its isotropic in-plane refractive index to achieve an efficient and ultra-compact uniform WBG modulator based on the spiral configuration with top and bottom electrodes [23,24]. The fabricated modulators could fit extremely long gratings into a highly dense area. Different grating designs were fabricated and optically characterized with regard to grating periodicity and width corrugation to assess the effect of the spiral configuration on WBG optical characteristics. Moreover, the central wavelength of the WBG could be efficiently tuned using the linear EO effect of LN with a tuning coefficient of 8.36 pm/V, while achieving a 3 dB EO bandwidth of up to 25 GHz.

2. DEVICE DESIGN AND FABRICATION

Figure 1(a) shows a three-dimensional (3D) schematic representation of the proposed spiral uniform WBG modulator. Width corrugations are added to two Archimedean spirals connected with an S-shape to have an input and transmission output as shown in Fig. 1(b) [18,23]. We designed the S-shape with an inner radius of ${R_o} = 40\;{\unicode{x00B5}{\rm m}}$ to provide a good balance between the WBG footprint area versus curvature-related perturbations in refractive index and optical insertion loss. According to Lumerical MODE simulations, for ${R_o} = 40\;{\unicode{x00B5}{\rm m}}$, the bending loss is 0.01 dB/cm, and the effective index variation $\delta {n_{{\rm eff}}}$ due to curvature is $1.2 \times {10^{- 3}}$. Such a small index variation can be ignored in our design as it causes a minimal effect on the grating response. Moreover, the required change of the grating period in order to compensate for this variation is less than 0.25 nm, which is below the fabrication resolution [19]. The bending loss and $\delta {n_{{\rm eff}}}$ increase to 0.31 dB/cm and $4.8 \times {10^{- 3}}$, respectively, at ${R_o} = 20\;{\unicode{x00B5}{\rm m}}$; then, both dramatically increase for ${R_o} \le 20\;{\unicode{x00B5}{\rm m}}$. While a radius of 40 µm was chosen for initial implementation, more compact spiral designs could be achieved in the future with ${R_o}$ down to 20 µm, while accounting for the extra loss and $\delta {n_{{\rm eff}}}$. The gap between spiral turns was chosen to be $g = 2\;{\unicode{x00B5}{\rm m}}$ to minimize the evanescent coupling of optical modes between different turns. The grating structure is mapped to the whole curve, wrapping a 2.2 mm long WBG into a circular spiral with a maximum radius of 60 µm. A nominal waveguide width of ${w_g} = 1\;{ \unicode{x00B5}{\rm m}}$ is chosen to minimize the propagation loss of the transverse magnetic (TM) mode, which was chosen to harness the EO effect by maintaining the optical mode polarization parallel to the Z axis of LN [23,25].

 

Fig. 1. (a) 3D schematic representation of the spiral WBG modulator on Z-cut LN. (b). Top view of the spiral WBG showing its relevant geometrical parameters. (c) Cross-sectional view of the implemented device showing different layers in the proposed stack.

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Using top and bottom electrodes, an RF signal is applied to the grating causing a shift in the central Bragg wavelength (${\lambda _B}$) of the WBG transmission stop band due to change in the effective index of the optical mode. As a result, by operating the WBG modulator at the bandgap edge, the input optical signal is intensity modulated [16,17]. Long grating lengths, achievable using the spiral configuration, would be highly desirable to increase the filter extinction ratio (ER) and roll-off slope at edges of the bandgap, hence enhancing the effect of wavelength shift on the amplitude of the modulated optical signal and reducing the required voltage for intensity modulation.

The optical filter characteristics are defined by the grating geometrical parameters shown in the inset in Fig. 1(b), namely the grating period ${\Lambda}$, the width corrugation depth $\Delta w = ({{w_2} - {w_1}})/2$, where ${w_1}$ and ${w_2}$ are the grating widths of narrow and wide grating segments, and the total spiral grating length $L$. The eigenmode expansion (EME) method in Lumerical simulation tools was used to simulate the performance of the optical band structure at different grating designs, while accounting for the rib-etched LN waveguides and sidewall angles. The period ${\Lambda}$ is chosen to adjust the center Bragg wavelength of the WBG, calculated as ${\lambda _B} = 2{n_{{\rm eff}}}{\Lambda}/m$, where ${n_{\rm{eff}}}$ is the average effective index of the grating, and $m$ is the Bragg order. The bandwidth of the stop band ($\Delta \lambda$) is directly proportional to the grating coupling coefficient ($\kappa \cong 2\Delta n / \lambda_B$) for long gratings in the strong coupling regime $({\kappa \gg \pi /L})$, where is the index perturbation between narrow and wide grating parts [26]. Higher ER is achieved through the increase of peak power reflectivity, ${R_P} = {\tanh ^2}({\kappa L})$. A deep width modulation $\Delta w$ results in high $\Delta n$ and, hence, a stronger coupling coefficient $\kappa$. That leads to a wider $\Delta \lambda$ and a higher ER, but at the expense of higher optical insertion loss due to extra roughness loss at the modulated grating sidewalls [27]. At high Bragg orders, the grating has larger period ${\Lambda}$, reducing the device fabrication complexity; however, the propagation loss increases because a fraction of optical power is diffracted toward other directions outside the WBG with certain angles depending on the diffraction order, and the bandwidth $\Delta \lambda$ decreases because the grating has a weaker coupling coefficient [26,28]. To test the quality of the spiral WBG on TFLN platform, we fabricated and measured both first-order $({{\Lambda} = 400\;{\rm nm}})$ and third-order $({{\Lambda} = 1200\;{\rm nm}})$ grating designs with two different values of $\Delta w$, as shown in Fig. 2(b). All grating designs have a total length of $L = 2.2\;{\rm mm}$.

 

Fig. 2. (a) Optical microscope image of the fabricated modulator after patterning the top cladding and electrode layers. (b) Top view SEM image of the spiral WBG before adding SU-8 and Al layers. Zoomed-in SEM images of the four fabricated grating designs are displayed, highlighting the period ${\Lambda}$ and corrugation depth $\Delta w$ of each design.

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The modulator was fabricated on a Z-cut TFLN wafer with the cross section shown in Fig. 1(c). The initial stack is composed of a silicon handle substrate, 140 nm thick bottom metallization layer made of gold (Au) with chromium (Cr) adhesion layers, 1 µm buried silicon dioxide (${\rm SiO}_2$) layer, and 800 nm Z-cut LN film. The optical waveguiding structures were patterned using electron beam lithography and then were formed by direct etching of 600 nm of the LN layer resulting in a sidewall angle of ${\sim}70^\circ$. The etching depth is chosen to minimize the overlap between optical mode and the rough sidewalls, while supporting the TM mode. Then, a top dielectric layer made of SU-8 polymer was used to cover the structure before adding the top electrode above the modulated spiral WBGs. We chose SU-8 polymer for top cladding as it provides a highly uniform surface with controllable thickness using spin coating [23]. Moreover, it exhibits excellent optical and dielectric properties (${n_r} \approx 1.57$ at 1550 nm, ${\varepsilon _r} \approx 3$) [29]. The thickness of the top cladding layer identifies the gap between the top and bottom electrodes ($d$) and should be precisely engineered for optimum RF path design. A thickness of 1 µm was selected to reduce the optical loss and device capacitance, while maintaining an appropriate electric field across the optical mode for efficient EO modulation. Aluminum (Al) was chosen for top electrodes as it results in lower propagation loss compared to other metals at 1550 nm. Although both Al and Au electrodes gave almost the same optical loss in TM mode simulations, Au requires a Cr adhesion layer, which increases the optical loss to 1 dB/cm at 1550 nm. More details on the fabrication process can be found in our previous work [23,24].

An optical microscope image of the fabricated device after top cladding and electrode deposition is displayed in Fig. 2(a). The top electrode is patterned to cover the S-curve without filling the inner circular area of the spiral to minimize the device capacitance for higher EO bandwidth. Figure 2(b) shows a top view scanning electron microscope (SEM) image of the spiral WBG before adding SU-8 and Al electrode. The inset shows zoomed-in images of the different grating designs. As displayed in the images, the sidewall angles appear along the fabricated gratings after the dry etching of LN, which would affect the grating coupling strength and optical filter response [23,30].

3. DEVICE CHARACTERIZATION

First, the optical transmission of the fabricated spiral WBG modulators was measured, so that the optical filter characteristics of different grating designs could be analyzed and compared. As shown in the measurement setup in Fig. 3(a), a tunable laser source (Santec 710 TSL) with input power of 10 dBm and a fiber polarization controller (FPC) are used to adjust the input light’s wavelength and polarization to the optical TM mode. Grating couplers [31] are used to couple the input into the modulator chip and then couple the modulated output to a high-speed InGaAs photodiode (Thorlabs DET08CFC). A high gain transimpedance amplifier (Thorlabs AMP100) amplifies the generated photocurrents, and output voltages are read out by a laptop-controlled data acquisition card (DAQ). The measurement shows a high optical insertion loss of the device, which is mainly due to the grating couplers’ loss (${\gt}{10}\;{\rm dB}$ for each side). The propagation loss was measured for the fabricated waveguides on Z-cut LN to be ${\lt}{1.3}\;{\rm dB/cm}$, using a ring resonator structure with 40 µm radius [23]. Although the simulated bending and electrode losses are 0.21 dB/cm for the 1 µm waveguides, the excess loss is attributed to the fabrication-induced scattering loss, which may be improved by optimizing the fabrication process to minimize surface roughness at the etched LN sidewalls [32].

 

Fig. 3. (a) Experimental setup to measure the optical response of the WBG modulator. FPC, fiber polarization controller; DUT, device under test; TIA. Transimpedance amplifier; DAQ, data acquisition card. (b) to (e) Normalized transmission spectra of the spiral WBG modulator for different grating designs; (b) ${\Lambda} = 400{\rm \;nm}$, $\Delta w = 100\;{\rm nm};$ (c) ${\Lambda} = 400{\rm \;nm}$, $\Delta w = 200\;{\rm nm}$; (d) ${\rm \;\Lambda} = 1200\;{\rm nm}$, $\Delta w = 100\;{\rm nm}$,; (e) ${\Lambda} = 1200\;{\rm nm}$, $\Delta w = 200\;{\rm nm}$.

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The measured normalized optical transmission spectra of the different grating designs are displayed along with simulated responses in Figs. 3(b)–3(e). As depicted, the experimental results are in good agreement with EME simulations. The small mismatch in the filter response and bandwidth ($\Delta \lambda$) between measurement and simulation can be attributed to imperfections in the fabricated grating shape due to the rounded grating edges at LN sidewall angles and perturbations in the spiral shape mapped period ($\Lambda$). These effects are not accounted for in simulations, where straight gratings are defined with sharp periodic transitions. The small shift in ${\lambda _B}$ between measurement and simulation is attributed to the fabrication tolerances in the grating period and the etch depth of TFLN. An optical ER of ${\sim}35\;{\rm dB} $ or higher is measured for first-order designs, which was limited by the noise floor in our experimental setup. The third-order designs exhibit a lower noise-limited ER of ${\sim}27\;{\rm dB} $ because they have ${\sim}8\;{\rm dB} $ excess propagation loss. Higher ER could be achieved by reducing the fiber-to-chip coupling loss, e.g., using an edge coupling method. The bandwidth ${\Delta}\lambda$ increases for high $\Delta w$ for both grating orders. The first-order grating design has a good filter response at small $\Delta w$ and skewed filter characteristics for the deeply modulated grating. This non-ideal filter response is believed to be due to the aforementioned fabrication tolerances, which have more impact at a small grating period. The third-order grating’s response is less affected at high $\Delta w$; hence, it would be appropriate for spiral grating designs with large width corrugations. As expected, the third-order design has a narrower bandwidth $\Delta \lambda$. It is worth noting that, using the spiral configuration, the high-order gratings could be implemented without significant increase in the device area.

Subsequently, the EO modulation response of the fabricated devices was characterized by recording the transmission spectrum at different applied voltages. Unfortunately, the measurement could not be done at DC due to the inadequate DC response observed for the proposed Z-cut LN modulator. That response is attributed to the flow and redistribution of the surface charges among the device structure under the application of DC voltage [9,33]. Such effects cause DC bias drift in the state-of-the-art X-cut MZMs and could be much more severe in Z-cut LN [23]. This problem can be mitigated using active thermo-optic phase shifters that could be added to the modulator design to bias the WBG modulator at the target wavelength [34]. While the weak DC response does not affect the usage of the proposed Z-cut spiral configuration as a modulator, this drawback hinders its possible implementation in the future as a phase-shifted BGR for tunable filter applications. Nevertheless, the spiral phase-shifted BGRs can still be employed in future designs to achieve compact resonantly enhanced modulators on LN, where coupled BGRs are combined with MZM to achieve phase modulation enhancement in Mach–Zehnder arms [35].

The effective EO AC tunability was measured by applying a 500 Hz square wave signal with alternating polarities to the modulator using a function generator followed by a ${50} \times$ piezo amplifier, while recording the output using the same setup in Fig. 3(a). Figure 4(a) shows the recorded optical spectra of the first grating design (${\Lambda} = 400\;{\rm nm}$, $\Delta w = 100\;{\rm nm}$) at the positive values of different pulse amplitudes. The shift in the Bragg wavelength is plotted as a function of the applied AC voltage amplitude in Fig. 4(b), demonstrating an EO tunability coefficient of 8.36 pm/V up to 120 V. A similar result was obtained for other grating designs. The measured result is in good agreement with the predicted sensitivity (8.43 pm/V) from calculating the effective index variation with voltage $V$ as given by [9]

 

Fig. 4. (a) Normalized transmission spectra at different AC voltages. The green point shows the selected operating wavelength for the modulation ER and EO frequency response measurements. (b) Bragg wavelength shift as a function of the AC voltage amplitude, demonstrating an EO tunability of 8.36 pm/V. (c) Measured modulation ER at different peak-to-peak voltages, while operating at $\lambda = 1547.65\;{\rm nm}$.

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Next, the modulation ER, defined as the output intensity ratio between the high and low levels of the output modulated signal (${P_{{\rm high}}}/{P_{{\rm low}}}$), was measured at different applied voltage levels to characterize the modulator electrical power consumption. The voltage signal was applied using a function generator, and the modulated RF signal was measured from the modulator output using an ultra-fast photoreceiver (Thorlabs RXM40AF) that was connected to an oscilloscope. The input light wavelength was tuned to the WBG bandgap edge at $\lambda = 1547.65\;{\rm nm}$ [shown in Fig. 4(a)] using the tunable laser source. The modulation ER was calculated from the modulated optical power (${P_{{\rm mod}}}$) and the carrier power (${P_o}$) at the operating wavelength, and it was plotted versus the applied peak-to-peak voltage (${V_{\textit{pp}}}$) in Fig. 4(c). The ER depends on the tuning sensitivity as well as the roll-off slope at the bandgap edge. Having an average roll-off slope of 0.21 dB/pm in the WBG transmission spectrum, the measured results are in good agreement with prediction. An optimum voltage would be selected to provide a good balance between the modulator power consumption and the required output ER that determines the bit error ratio (BER) in the communication link.

 

Fig. 5. (a) Experimental setup for measuring EO modulation frequency response. VNA, vector network analyzer. (b) Measured small signal EO response, showing 3 dB bandwidth of the modulator.

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The EO modulation frequency response of the proposed modulator was measured from 100 MHz to 30 GHz using a two-port vector network analyzer (VNA) (Agilent N5230A) according to the setup in Fig. 5(a). Port 1 of the VNA is used to apply a sinusoidal microwave signal to RF electrodes; port 2 is connected to the output of the photoreceiver to measure EO ${{\rm S}_{21}}$. The performed measurement using the first grating design is shown in Fig. 5(b), where the input light wavelength was tuned to the WBG bandgap edge. As depicted in the output results, a 3 dB bandwidth of 25 GHz was achieved. The bandwidth is mainly restricted by the resistance-capacitance (R-C) electrical limit because the RF electrodes are considered lumped in the proposed Z-cut configuration [23,24]. The four grating designs have similar RF bandwidth because they have the same length and electrode area. In such lumped configuration, the grating length can be used to optimize the trade-off between drive voltage for a given ER and EO bandwidth. This voltage-bandwidth trade-off exists because longer gratings can be driven with a lower voltage due to the steeper roll-off but have smaller EO bandwidth due to the increased capacitance. Moreover, another trade-off arises between EO bandwidth and optical loss because decreasing the grating length for wider bandwidth would require higher $\Delta w$, in order to increase coupling ($\kappa$) to have the same peak reflectivity, at the expense of excess roughness loss. Wider modulation bandwidth could be achieved by reducing the driving RF resistance (R) through custom designs of the electrical driving circuit. Moreover, thicker top and bottom cladding layers could be used to decrease the device capacitance (C); however, that would reduce the EO tuning sensitivity due to the increased electrode gap (d).

4. CONCLUSION

A highly compact spiral WBG modulator has been demonstrated on a Z-cut LN platform. The modulator utilizes the strong EO effect of LN to efficiently tune the transmission stop band of spiral uniform WBGs using top and bottom electrodes. The fabricated WBGs were optically characterized with regard to grating periodicity and corrugation depth, demonstrating an ER of over 35 dB. EO modulation experiments show an AC tunability factor of 8.36 pm/V for the central Bragg wavelength. Moreover, a 3 dB EO bandwidth of 25 GHz was measured, mainly limited by the electrical R-C time constant. The spiral-shaped configuration fit a 2.2 mm WBG into a total footprint area of ${120} \times {120}\;{{\unicode{x00B5}{\rm m}}^2}$, allowing for extremely long WBGs that are required for high ER drop response with steep roll-off. The presented implementation provides an ultra-compact, compared to conventional MZMs, and wideband, compared to microresonators with finite cavity photon lifetimes, integrated photonic modulator.