On the Performance of a Miniaturized Reactive Loaded Monopole Antenna for Ex Vivo Catheter Applications

In this paper, we propose a miniaturized series L-C loaded monopole antenna for catheter application in microwave ablation systems. Initially, a quarter wavelength long monopole antenna (Design1), having length 30 mm (<inline-formula> <tex-math notation="LaTeX">$\approx 0.24\lambda _{0}$ </tex-math></inline-formula> where <inline-formula> <tex-math notation="LaTeX">$\lambda _{0}$ </tex-math></inline-formula> represents the free-space wavelength at 2.4 GHz operating frequency) using series LC loading concept is designed. The initial design is further extended to a miniaturized version (Design2) having length 5 mm (<inline-formula> <tex-math notation="LaTeX">$\approx 0.04\lambda _{0}$ </tex-math></inline-formula>), showing nearly 83% size reduction compared to its former.Both the antennas are designed and simulated by immersing inside high permittivity egg white phantom (<inline-formula> <tex-math notation="LaTeX">$\epsilon _{r}$ </tex-math></inline-formula> = 63.84). Antenna Design 1 and 2 exhibit good impedance matching (<inline-formula> <tex-math notation="LaTeX">${S_{11} < -12}$ </tex-math></inline-formula> dB) at and around the operating frequency with uniform monopolar radiation pattern having co-to-cross isolation of nearly 30–35 dB. Further, the in-phantom Specific Absorption Rate (SAR) values for both antennas are evaluated using simulation and verified with the measured SAR values such as 22 W/kg and 19.6 W/kg for Design1 and 2 respectively, using an in-house experimental setup. Moreover, the microwave ablation property of the proposed antennas is studied using simulated transient thermal gradient which translates to ablation zone formation. Here, Design 2 shows the advantage of having a low invasive diameter of 1.5 mm and near unity aspect ratio (AR) for the ablation zone as compared to Design1.

The evolution of better impedance-matched catheter design started with properly designed baluns for good impedance match with improved ablation zones and stable performance with respect to changes in insertion depth [20]. Some of the documented examples depict coaxial baluns implemented by surrounding the coaxial cable with a hollow circular conductor as in cap-choked antenna [13], coaxial slot dipole [14], triaxial choked dipole [15], [16], and expanded tip wire antenna [17] or the floating sleeve dipole VOLUME 11, 2023 This work is licensed under a Creative Commons Attribution 4.0 License. For more information, see https://creativecommons.org/licenses/by/4.0/ (FSD) [20]. Designs improvised without baluns for achieving compactness and invasiveness of the catheter antennas are documented in [26], [28], [30], [31].
In this paper, a compact series L-C loaded 30 mm (≈ 0.24λ 0 ) long monopole type catheter applicator (Design1) and its 5 mm (≈ 0.04λ 0 ) long 83% miniaturized version (Design2) both working at and around ISM bands centered at 2.4 GHz are thoroughly explored. The reactive reactance and the various degrees of freedom obtained by varying the parameters of L-C loading elements can be strategically engineered to obtain a good impedance match as well as control the desired level of miniaturization needed for catheter applications. A concrete design guideline for both designs is laid down in the following subsection. The two designs are characterized inside egg white phantom (ϵ r = 63.84) which replicates high permittivity environment encountered in real time microwave ablation of targeted biological tissue mass. A quantitative comparison in terms of in-phantom impedance matching in the ISM band, transient temperature-rise, aspect ratio of ablation zone formation for the same levels of input power and change in SAR profile with increasing radial distance from the catheter applicators is enunciated in the following subsections. An in-house laboratory set up providing 35 dBm (2.5 W) of input RF-power has been developed for experimental verification of the simulated results obtained using CST Thermal Studio Suite (v.2020). Miniaturized Design 2 not only exhibits very low invasive diameter C D =1.5 mm (with catheter encapsulation) but also near unity Aspect Ratio of simulated ablation zone which translates to spheroid-shaped ablation zone in real time catheter applications. On the contrary, tear-drop-shaped ablation zone caused by shaft heating in many applicators cause unwanted charring of healthy tissues as well, hence, undesired. Design2 portrays a slightly reduced maximum achievable SAR value (22 W/Kg) than that of Design1(19 W/kg) for the same level of input power 2.5 W. However, a considerable trade-off amongst invasive diameter, maximum achievable SAR, and ablation zone shape indicates that Design2 is a more suitable configuration for real-time catheter applicators in microwave ablation of deep-seated solid tumors.

II. ANTENNA DESIGN AND GEOMETRY
A. DESIGN 1 Fig.1(a) shows the schematic of Design1, an initial L-C loaded monopole antenna of length h t = 30 mm (≈ 0.24λ 0 ≈ 1.92λ g where λ 0 the free-space wavelength at 2.4 GHz operating frequency and λ g is the effective wavelength inside egg white phantom having ϵ r = 63.84).Reactive loading of monopole antenna is a standard documented technique for obtaining desired levels of miniaturization at a lower frequency with respect to an equivalent unloaded quarter wavelength monopole of same length designed to work at a higher frequency [37], [38]. In Design1, this reactive loading approach has been strategically exploited for obtaining good impedance matching inside high permittivity non-perfused egg white phantom (ϵ r = 63.84) which replicates high permittivity in-body medium for catheter applications while in Design2 the miniaturization aspect has been explored. Design1 has been conceived initially as in [44] and further improvised in [45]. The stages of topological development are shown in Fig.2. As evident from Fig.2(a), a bare quarterwavelength monopole antenna (h t = 30 mm) fitted to a small ground plane with radius G R = 2.38 mm lacks satisfactory impedance matching and has an operating frequency around 2.8 GHz. A higher G R value provides better impedance matching for a standalone monopole but too high G R is detrimental for catheter antenna design which requires a minimum invasive diameter C D . Therefore, an optimum trade-off in obtaining good impedance matching and uniform monopolar radiation at 2.4 GHz as well as maintaining the relation G R ≤ C D is needed. This can be achieved by primarily fixing the value of G R = 2.38 mm followed by adding in additional reactive L-C loading elements in series. The helical wire with turn radius R h and number of turns N is attached to the ground plane on one end and to the monopole wire via a short conducting pin of length L s (= R h − h d /2) = 1.5 mm, where h d is the monopole wire diameter on the top end. This is further connected in series with a parallel plate metal capacitor introduced in between the monopole wire at a distance fro the ground plane, best optimized as L CG = 8.5 mm. Impedance matching is obtained in here through extensive parametric variation of the degrees of freedom offered by the L-C loading elements such as number of turns of the helical coil, N , radius of the helical turn, R h , the pitch of the helix, P, the position of the metal capacitor from the ground plane L CG , the gap between metal plates of the capacitor C g and radius of the capacitor, R G . The best optimized design topology thus works at 2.4 GHz and provides excellent impedance matching by exploiting the sharp rise on the real part of the impedance curve. A concrete set of design guideline is formulated as follows: • It is ideal to keep: 2 < Cap r /C g < 3.5.
• No of turns, N can be optimized in accordance to the |Z R | and |Z I | plot on the impedance curve to ensure good impedance match. For ease of fabrication, large value of N is not encouraged. In here, the optimized value of N = 5.
• Limitation on capacitor radius and radius of the helical coil R h : • Ratio of the thickness of the monopole (L thck ) and that of the helix filament (H thck ) should be maintained as, B. DESIGN 2: MINIATURIZED Fig. 1(b) portrays the schematic for Design2, which is an 83% miniaturized version of Design 1 having length h t = 5 mm (≈ 0.04λ 0 ≈ 0.32λ g where λ 0 the free-space wavelength at 2.4 GHz operating frequency and λ g is the effective wavelength inside egg white phantom having ϵ r = 63.84). Documented literature depicts miniaturization through series inductive loading or shunt capacitive loading using a top hat structure in monopoles or both. These common techniques have their respective disadvantages [37], [38]. The concept of miniaturization technique using series L-C loading is documented for dipoles and folded dipoles in [37] and [38] by introducing a resonance frequency f R prior to the first antiresonance (f AR ) of the folded dipole so as to exploit the sharp rise on the real part of the impedance curve. This is further improvised herein to obtain a balun-free design suitable for catheter applications. A detailed parametric variation of the L-C loading obtained through simulation is shown in Fig.3. Design2 has an invasive diameter C D = 1.5 mm and a non-radiating insertion length, L = 50 mm made of RG178 coaxial cable having an outer diameter of about 1.75 mm. Abiding by the design guideline laid down in the previous subsection, the consequent values of the number of turns, N = 18 and pitch, P = 0.28 mm, gap between the capacitor plates, C g = 0.50 mm, radius of the helix, R h = 0.65 mm and the radius of the capacitor plates, R G = 0.625 mm yield the best-optimized values for Design2. The ratio of the H thck /L thck is maintained at 0.196 which is less than 1 as specified in the design guideline. A very low pitch value (P = 0.28 mm) and monopole wire diameter h d = 0.3 mm calls in for challenges in fabrication which are discussed in the subsequent subsections. The main difference in configuration of Design1 and 2 lies in the length of the helical wire, L h (= N * P) wound over the quarter wavelength monopole antenna of height h t . The radius of the helical turn, R h , precisely determines the maximum value of C D . The length L s (= R h − h d /2) is around 1.5 mm for Design1 and 0.5 mm long for Design 2. In Design 1, the relation between L h and h t (0.5 h t < L h < h t ) obtained through parametric variation of N and P for the best impedance matching at 2.4 GHz stands as L h = 0.67h t . In Design2, the length of the miniaturized monopole is reduced from 30 mm in Design1 to 5 mm, hence the need to increase the inductive loading is unavoidable to provide impedance matching. Thus, it is not feasible from the fabrication aspect to follow the aforesaid relation between L h the and h t due to the practical limitation of having too low pitch, P, and too high value of number of turns, N for a miniaturized monopole length of h t = 5 mm. Therefore, the best optimized relation stands as L h = h t , so as to obtain good impedance matching at around 2.4 GHz.This directly translates to the fact that in Design1 minimum invasive diameter at the tip is A little improvisation is achieved during fabricating the prototype where the ground plane radius G R is further reduced to 1.5 mm, an optimum value in between C Dmin and C Dmax , thus making it more advantageous in invasive applications.

A. SIMULATION: FREQUENCY DOMAIN
The two versions of the proposed catheter antenna are simulated and investigated thoroughly in the frequency domain. Fig. 4(a)-(b) depict the frequency variation of the S 11 characteristics of the miniaturized and non-miniaturized prototypes for various topologies dipped in egg white phantom: • With the parylene C (ϵ r = 3) external catheter tubing • Without the parylene C (ϵ r = 3) external catheter tubing • With the parylene C (ϵ r = 3) external catheter tubing and an additional non radiating coaxial insertion length L for both Design1 and 2.
Each topology depicts a good impedance match inside the egg phantom at and around 2.4 GHz. The S 11 characteristics of the bare and unloaded monopole of length h t = 30 mm with a ground plane radius, G R = 2.38 mm for Design1 and h t = 5 mm with ground plane radius G R = 0.875 mm for Design2 is also plotted in Fig.4(a) and (b). The bare monopole of length h t = 5 mm has a distinctively higher operating frequency, f opt = 16 GHz as compared to its series L-C loaded versions, Design1 and Design2. Due to the difference in design topology as discussed in the previous subsection, Design 2 is indeed a highly miniaturized (≈ 83%) version of Design1 while maintaining the same operating frequency centered around 2.4 GHz. The relative permittivity, density, and thermal conductivity of the egg white medium are calculated using regression analysis as documented in [39]. The calculated values are used to assign the thermal and electric properties of the medium in the CST thermal studio suite to maintain close corroboration with simulated and measured scenarios. The calculated values for T = 25 0 C are as follows: • Loss tangent δ = tan −1 (ϵ ′′ /ϵ ′ ) = 0.2505 Fig. 5(a) and (b) portray E-plane radiation patterns of Design1 and miniaturized Design2. The simulated pattern for each topology ensures monopolar type radiation and the in-phantom co-to-cross isolation for Design1 is around 30 dB and similarly for the miniaturized Design2 is around 35 dB. The peak gain portrayed for both Design1 and 2 is around −20 to −25 dBi which is quite the case for high permittivity in-phantom applications.
The equivalent circuit for Design1 and Design2 is modelled following the standard five-element approach of unloaded dipole of any arbitrary length as documented in [40], [41], and [42]. The proposed equivalent circuit has additional series inductor L load and capacitor C load which arise due to the series reactive L-C loaded topology in Design1 and Design2. The values of L load and C load are calculated following equations in [43] and the circuit is analyzed at 2.4 GHz operating frequency with help of Advanced Design System (ADS) as shown in Fig.3. The miniaturized Design2 with length h t =5 mm has higher inductive loading for better impedance matching at the desired operating frequency with number turns (N =18, P =0.33 mm). This is reflected in the value of L load (= 0.45µH) being nearly 4 times higher than that of Design1 (L load = 0.11µH) having N =5, P =4 mm. Similarly, since the area of the capacitor plates is smaller in miniaturized Design2, C load (=1.4 pF) for miniaturized Design2 is smaller than that in Design1(C load =15.9 pF). Thus, keeping the other parameters almost unaltered except for a slight change in the value of C ′ 0 for Design1 and Design2 in the five-element design, the optimized values of C load and L load give the clear picture of how the series inductive and capacitive loading changes in Design1 and Design2 in order to achieve miniaturization. The proposed equivalent circuit analyzed at 2.4 GHz depicts the respective S 11 plots as shown in Fig. 6(b:1) and 6(b:2).

B. SIMULATION: TRANSIENT THERMAL ANALYSIS
The proposed prototypes are thoroughly studied in CST Thermal Studio Suite (v.2020) to trace out the rise of temperature inside egg phantom for a specific input power of 35 dBm (2.5 W) at 2.4 GHz. Our prime focus is to exhibit the physical changes that take place inside the phantom which further paves the way for the ultimate cell necrosis in real-time ablation systems. CST Thermal Studio Suite provides co-simulation where electric and magnetic thermal losses, surface current, power loss densities are evaluated at 2.4 GHz and serve as thermal sources for high frequency transient thermal analysis. An input power of 2.5 W(∼ 35 dBm) is used for thermal simulation as that matches with the input power applied for an economic laboratory friendly measurement set up. Fig.7(a) and 8(a) show the temperature versus time plot of Design1 and Design2 for different thermal monitors placed at increasing radial distances from the applicator dipped inside egg phantom. The monitors are placed at a depth D = h t /2, equivalent to the half the length of the catheter applicators. In thermal simulation, no catheter covering made of Parylene C is used for both the designs to maintain strong corroboration in simulation and experimental scenario. Fig.7(b) and 8(b) show and numerically computed SAR value calculated from the slope of the temperature high-frequency for each distance using the simple calorimetric equation documented in [13]: where D t is the transient temperature rise in 0 C, d t is duration in seconds, and c is the specific heat capacity in Kcal/kg/ 0 C. VOLUME 11, 2023 FIGURE 7. (a) Simulated Temperature versus Time curve for Design 1 applicator at thermal monitors being placed over 5 mm to 30 mm with distance varying radially from the catheter applicator, (b) Plot of numerically computed SAR profile with varying distance from the temperature versus time graph in Fig. 7(a)).

FIGURE 8. (a) Simulated Temperature versus Time curve for miniaturized
Design 2 applicator at thermal monitors being placed over 5 mm to 50 mm with distance varying radially from the catheter applicator, (b) Plot of numerically computed SAR profile with varying distance from the temperature versus time graph in Fig. 8(a)).
It can be inferred that heat capacity of a medium M = mc, where m is the mass of the medium in which the applicator is immersed, and c is the specific heat capacity, plays a pivotal role in determining transfer from the applicator to the biological medium and thus cause rise in temperature due to increased kinetic energy of the medium molecules. Mass and specific heat of the medium being constants, the input microwave power is transformed into thermal energy by the catheter applicator and this strongly determines temperature rise in a stipulated transient time duration. Once steady state is reached, irrespective of the change in time duration, temperature gradient will remain unaltered unless there is a boost in the input microwave power. Fig. 7(b) and 8(b) show the SAR plots which depict the physical process of deposition of electromagnetic energy that decreases with an increase in radial distance from the applicators. This translates to ablation zone formation for Design1 and Design2, simulated for 2.5 W input power as shown in Fig. 9. The Aspect Ratio (AR), i.e., the ratio of the radial width of the ablation zone W to the axial length L for 38 0 C temperature contour line value, is evaluated for both the designs. Input power level is kept at 5W for simulation of ablation zone formation to enable temperature to reach beyond 40 0 C where protein denaturation of biological tissues starts. In real-time scenarios, complete cell necrosis happens beyond 50 0 C only which requires greater than 10 W of input power. The calculated value of AR for  Design1 is 15.8/17.08 = 0.89 while that for Design2 is 8.6/8.1 = 1.06. Thus, for Design2, AR ∼ 1 translates to spheroid-shaped ablation zone formation which is highly desired in ablation applications. Moreover, as seen from Fig.9(a)-(b), the simulated temperature profile at and around the tip of the monopole applicators in both designs is much below the maximum achievable temperature as shown in the respective colourmap. This is indeed an advantage for real time ablation scenarios in deep tissue environment as it can prevent any unwanted damage to healthy and vulnerable tissues/organs/nerves located in the immediate vicinity of the catheter applicator tip beyond the target tissue.

C. FABRICATION AND MEASUREMENT
The fabricated prototypes of the proposed Design1 and 2 are shown in Fig.10(a) and (b) respectively. Both Design1 and 2 are manually assembled and have fabrication tolerance of ±5% and ±20% respectively. In Design2, winding the helical wire manually with pitch value P = 0.33 mm and the number of turns, N = 18 around the thin stranded inner conductor of RG178 cable (h d = 0.3 mm) serving as the monopole has been challenging. The copper windings in Design2 wound over the central conductor of the RG178 flexible coaxial cable (Bend diameter = 10.36 mm) tend to compress and expand as they lack tensile strength owing to the extremely small wire thickness of about 0.2 mm. This could be avoided in real-time scenarios which require robust catheter tube encapsulation unlike in laboratory testing where owing to limited input power, experiments are carried on without Parylene C catheter encapsulation to avoid any unwanted power loss incurred in the heating up the dielectric enclosure. Fig.11 compares the measured and simulated S 11 characteristics for the prototypes of Design1 and Design2 without their catheter tubing, dipped inside egg phantom and both portray close corroboration.
An inexpensive low-profile yet elaborate laboratory set-up is shown in Fig 12. The complete laboratory-friendly in-house set up-comprises three main components: • Signal generation using an easily available SDR Tx-Rx Module (ADALM Pluto) capable of transmitting 7 dBm power. The SDR module runs on a code-based program in which, input attenuation and frequency of operation can be manually entered into the laptop/PC. The output Signal power is measured in a Spectrum Analyzer (Tektronix, RSA306B) working up to 6 GHz having a power limit of 20 dBm.
• Two cascaded gain block amplifiers (single-ended RF/IF on-chip) ADL5610-EVALZ having fixed gain 15.4 dB working over 30 MHz to 6 GHz, are connected in series to the output of SDR module as shown in Fig.12(a)-(b). A total of around 35 dBm taking into account the cable wiring loss and the interconnector losses which are estimated and shown in Fig.12.Thus, the maximum obtainable input power is; P in = P 01 (SDR)+P 02 (Gain block) +P 03 (Gain block cascaded)-Cable Loss = 34.45 dBm (∼ 2.5W).
• Phantom preparation from around 18 eggs carefully separated out into a clean 1L beaker, and weighing around 689 gm filling up to the 400 ml mark in a 1L beaker is shown in Fig.12(c:c1). 35 dBm power is delivered via a low-loss cable to the prototype dipped in egg phantom. The bottom surface of the beaker is marked in gradations of 1cm along the diameter of the circular base so as to facilitate the temperature reading when the thermal probe is moved away radially outwards from the catheter applicator as shown in Fig.12(c:c2).  Fig.13 shows some instantaneous temperature readings for the in-phantom miniaturized prototype when the thermal probe is inserted at a depth midway to the applicator length D = h t /2. The obtained set of data for temperature versus time is plotted and variation of SAR with distance is portrayed in Fig.14(a)-(d) for both prototypes. Simulated and measured maximum achievable SAR values slightly vary owing to real-time ambient temperature fluctuations encountered during experimental studies. The maximum temperature rise encountered in the experimental study is not more than 35.5 0 C for the Design1 prototype and around 33 0 C for VOLUME 11, 2023    and rapid and exploit the rapidly rising graph of the transient temperature till before the steady state which is clinically achieved in between 5 − 10 mins where temperature shoots up to > 50 0 C for 30 − 40 W input power.
Since the input power available is low and strictly for laboratory demonstration purposes, both the miniaturized and initial versions are used without the catheter covering. A detailed chronological comparison of different genres of antenna applicators along with the proposed designs for catheter applications is enunciated in Table 1 Table 2.

IV. CONCLUSION
Two different L-C loaded antenna configurations named Design1 and Design2 are developed. Design2 exhibits: • Lowest invasive diameter (C D = 1.5mm) amongst the catheter applicators documented by far. A non-radiating insertion length (L = 50 mm) made of 1.83 mm thick RG178 coaxial cable is incorporated for reaching deep-seated solid tumor targets for catheter ablation application.
• Highly miniaturized (≈ 83%) version of Design1, with length h t = 5 mm, the lowest documented radiating length by far. The level of miniaturization can be easily controlled as per requirement by varying the degrees of freedom associated with L-C reactive loading.
• The transient thermal studies depict the required near spherical simulated ablation zone having unity AR for Design2 which makes it a more suitable contender for catheter applications. Both the designs are fabricated and the respective antenna parameters are measured. An in-house experimental setup is developed and in-phantom Figures of Merit such as Transient Temperature gradient and SAR are measured and compared with the simulated results. The corroborative simulation and experimental results show that Design2 suits the purpose of this genre of reactive monopole type antennas to be used as catheter applicators in microwave ablation of deep seated renal, hepatic, intestinal tumors and abnormal tissue mass prone to being carcinogenic. The proposed designs can be redesigned in multiport configurations using the concept of MIMO antenna systems deployable in the sub-6 GHz range of 5G [46], [47], [48]. These multiport catheter applicators can be future contenders to ablate large hepatocellular carcinoma and other larger (> 3cm) deep-seated solid tumors.