## 3 Results  
Using the RF coil circuit co-simulation approach, the optimal end-ring capacitors were calculated such that the reflection coefficients (S\({}_{ii}\)) at each port and the isolation coefficients (S\({}_{ij}\)) between adjacent ports were at least \(-\)15 dB before including the matching circuit. Figure 3 shows B\({}_{1}\)\({}^{+}\)-field distributions (Figure 3a,b, respectively) in a uniform cylindrical phantom and a human model as well as the SAR distribution (Figure 3c) in the coronal (upper row) and the axial (bottom row) planes. Table 2 summarizes uniformity, transmission efficiency, and maximum B\({}_{1}\)\({}^{+}\) at 20 kW in the cylindrical uniform phantom as well as mean 10 g average SAR and peak 10 g average SAR in the human \"Duke\" model, normalized to 1 \(\mu\)T in terms of B\({}_{1}\)\({}^{+}\) magnitude. The B\({}_{1}\)\({}^{+}\)-field uniformities for a circular birdcage (diameter, 58 cm) driven with quadrature driving and the split-type elliptical birdcage (62 cm \(\times\) 58 cm) driven with semi-elliptical driving were 74.29% and 76.25%, respectively. The simulated SAR values obtained for all the birdcage RF coils were within the IEC SAR limits (IEC 60601-2-33). The mean 10 g average SAR for the circular and the split-type birdcages was identical; however, the peak 10 g average SAR of the split-type birdcage was 19.56% higher than that of the circular birdcage. The S-parameter and the Q measurements of the split-type elliptical birdcage performed inside the MRI bore are listed in Table 3. The unloaded-to-loaded Q ratios (Q\({}_{\text{U}}\)/Q\({}_{\text{L}}\)) were 2.2 and 2.6 for light and heavy loader phantoms, respectively, indicating sample dominance. One of the main difficulties encountered in achieving better isolation was the coupling between the coaxial cable traps and the RF shield, which improved significantly when cable traps were shielded compared with when they were unshielded. After the tuning and the matching of each port, the vertical ports (#1 and #3) and the horizontal ports (#2 and #4) were connected to a \(-3\) dB splitter/combiner, and their respective S-parameters were measured. The reflection coefficient (S\({}_{11}\)) and the isolation coefficient (S\({}_{21}\)) at the splitter/combiner ports are listed in Table 3, indicating acceptable isolation between the horizontal and the vertical ports. The constructed prototype included a tuning/detuning circuit, which was tuned at a forward DC bias of \(+30\) V and detuned at a reverse DC bias of \(-30\) V.  
Figure 3: (**a**) Respective B\({}_{1}\)\({}^{+}\)-field distribution in cylindrical uniform phantom (top–coronal; bottom–axial); (**b**) respective B\({}_{1}\)\({}^{+}\)-field distribution in 3D human “Duke” model (top–coronal; bottom–axial); (**c**) respective SAR distribution in human “Duke” model (top–coronal; bottom–axial). Note that input powers for B\({}_{1}\)\({}^{+}\)-field and SAR distributions were normalized to 1 \(\upmu\)T.  
Figure 4 shows the SNR map and the FA maps of the uniform phantom acquired in axial, sagittal, and coronal orientations using the commercial two-port circular birdcage and the proposed split-type elliptical birdcage. For each orientation, the RF pulse power was calibrated during pre-scan, as presented in Table 3. Furthermore, Table 4 shows that the split-type elliptical birdcage required a moderately higher RF power (average, 18.68%) than the circular birdcage, which is consistent with the simulation analysis (18.33%).