 Original research
 Open access
 Published:
Typology and working mechanism of a hybrid power router based on powerfrequency transformer electromagnetic coupling with converters
Protection and Control of Modern Power Systems volume 8, Article number: 42 (2023)
Abstract
The power router (PR) is a promising piece of equipment for realizing multivoltage level interconnection and flexible power control in the future distribution power grid. In this paper, a hybrid PR (HPR) topology based on powerfrequency transformer electromagnetic coupling with converters is proposed for the medium distribution power grid. The powerfrequency transformer is used to undertake power transmission, voltage conversion, and other main tasks, while the power electronic converters are combined to achieve active control. Equivalent magnetic and electrical circuit models are established to help discuss the operating principle of the proposed HPR. Additionally, the power flow and control principle of the HPR in different operating conditions are analyzed, with the control system design scheme presented. The theoretical analysis results are verified by MATLAB/Simulink + Plecs simulation and a controller hardwareintheloop study, as well as a downscale experimental test, indicating that the proposed HPR is flexible in active voltage support and current control.
1 Introduction
Owing to uncontrollability, lack of fault isolation and fault ridethrough capabilities, as well as other shortcomings, it is challenging for conventional transformers to actively control and optimize the operation of distribution power grids, where nonintelligent techniques, such as bus tie switch and transformer tapping, are mainly adopted to control node voltages and power flows [1]. Additionally, complex grid faults can cause transient disturbances [2], such as unbalanced voltage dip or swell. These can critically affect the safe connection and operation of distributed renewable energy, microgrid, and load. Numerous offgrid events of distributed generation (DG) equipment have been reported. A power router (PR) is significant for renewable DG connection, multivoltage flexible interconnection, and multiport power control, and can notably boost the renewable DG accommodation capacity and continuous power supply capacity of distribution grids, to meet the urgent demand for flexible control and reliable power supply of power distribution grids caused by high penetration of renewable energy sources and power electronic equipment [3].
The PRs currently used for mediumvoltage and mediumpower conversion in power distribution grids use cascaded Hbridge (CHB) typology [4] or modular multilevel converter (MMC) typology [5]. A CHBbased PR topology is proposed by the Future Renewable Electric Energy Delivery and Management System of North Carolina State University [4]. Its main circuit uses an inputseries outputparallel (ISOP) topology, which comprises a mediumvoltage stage CHB, an isolation stage with ISOP dual active bridge (ISOPDAB), and a lowvoltage stage DC/AC inverter. This design meets the fundamental function requirements for PRs. An autobalancing electronic power transformer (EPT) topology, which decouples the mutual influence of the threephase imbalance between the high and lowvoltage sides of the EPT by an interleaving design, is proposed in [6]. The aforementioned topologies [4, 6] use the CHB on the mediumvoltage side. Reference [7] proposes an MMCbased PR topology, which comprises a mediumvoltage stage MMC, an isolation stage with ISOPDAB, and a lowvoltage stage DC/AC inverter. These can realize the interconnection of mediumvoltage AC (MVAC) and DC distribution networks. Several studies have shown the enhancement of the technoeconomic aspects of the full power electronic converters (PEC) modulebased PR through structural optimization [8] or control improvement [9] to achieve certain beneficial effects.
CHB and MMCbased PRs integrate functions, such as voltage conversion, power control, fault isolation, and power quality control, and have become paramount equipment for integrating and optimizing various PEC equipment in power grids, and improving the intelligent power grid management level. Nevertheless, these topologies require numerous PEC modules, and there are still several challenges in replacing powerfrequency (PF) transformers with PRs in AC power distribution grids, particularly the high cost, high loss, and complex typology of CHB and MMCbased PRs (e.g., 1 MVA 10 kV / 400 V PR). They have more than five times the cost and three times the loss of a PF transformer [10]. This poor technoeconomic performance inhibits the widespread application of PRs to this day.
Compared with the CHB and MMCbased PRs scheme, the hybrid scheme of the PF transformer and PEC is a potential direction for PR development. An integrated transformer design scheme based on the inductive filtering principle, which is designed with an extended delta winding on the secondary side for the parallel connection of the converter and realizes current control using the magnetic flux balance principle, is proposed in [11, 12]. Nevertheless, this scheme is used only for power quality control. An integrated transformer design scheme based on a multiwinding tap connection is proposed in [13, 14], in which the center tap is drawn from the mediumvoltage side winding of the transformer for the connection of the converter. This scheme is also used for power quality control. However, it is difficult for these integrated schemes of a PF transformer in parallel with a converter [11,12,13,14] to realize active voltage regulation, and they fail to meet the flexible control requirement for PRs.
Thus, a next generation distribution transformer, which comprises two windings on the secondary side with one connected to an AC/AC converter in series, is proposed in [15]. This structure is further optimized in [16] and extended to a threephase structure [17]. A transformerintegrated dynamic voltage regulator, which realizes the dynamic voltage amplitude and phase angle adjustment of the secondary side by controlling the amplitude and phase angle of the output voltage of the AC/DC/AC converter, is proposed in [18]. However, the voltage amplitude and phase angle adjustment ranges of the transformer are narrow, and all transmitted power flows through the connected AC/DC/AC converter, thereby increasing the operating power loss and reducing operating efficiency and reliability. A new type of PR with seriesshunt architecture is constructed in [19], in which the series PEC is connected to the AC grid in series through the transformer and the shunt PEC is connected to an AC bus in parallel. A transformerintegrated backtoback converter scheme, which realizes the simultaneous correction of voltage and current power quality, is proposed in [20]. A multiwinding integrated threephase hybrid transformer with series–shunt threephase twolevel converters connected by the secondary winding on the lowvoltage side is proposed in [21, 22]. In addition, references [23, 24] propose an improved hybrid distribution transformer, though this structure requires an additional series transformer on the mediumvoltage side, thereby increasing complexity and cost.
In this paper, a hybrid PR (HPR) based on PF transformer electromagnetic coupling with PECs is proposed, one which combines the benefits of the PF transformer and PEC. The core idea of the proposed HPR is that it uses the PF transformer to undertake the voltage conversion, power transmission, and the lowvoltage PEC to realize the flexible active control function. The equivalent magnetic and electrical circuit models are established to expound the HPR’s operating principle. In addition, with the control scheme presented, the power flow and control strategy of the HPR in different operating conditions are analyzed. Finally, the feasibility of the proposed HPR is verified by Simulink + Plecs simulation, controller hardwareintheloop (CHIL) and downscale experimental test results.
2 HPR topology and its wiring scheme
The proposed HPR topology is shown in Fig. 1, where the PF transformers are used for the main power transmission and the converters are used for the active voltage and current adjustment. The proposed HPR topology comprises a grouptype threelimb fourwinding PF transformer, a series converter, and a shunt converter, as shown in Fig. 1a. The wiring scheme of the grouptype threelimb fourwinding PF transformer is shown in Fig. 1b. The threelimb fourwinding PF transformer comprises a threelimb transformer core, mediumvoltage, series, lowvoltage, and parallel windings. The mediumvoltage winding of the PF transformer is deltaconnected to a 10 kV distribution power grid, the series winding is starconnected to the AC side of the series converter, the lowvoltage winding is starconnected to form a 400 V lowvoltage AC (LVAC) grid, and the parallel winding is starconnected to the AC side of the shunt PEC. The DC sides of the shunt PEC and the series converter are connected together to form a lowvoltage DC (LVDC) port, as shown in Fig. 1c. Compared with the conventional PR topology based on multiple PEC modules, the proposed hybrid scheme has the following advantages:

(1)
Simple topology and control complexity. In the mediumvoltage distribution grid scenario, only part of required power needs to be converted by the series and shunt PECs of the HPR, and generally, no multilevel or multistage cascaded typologies are required, leading to a simpler topology and control complexity.

(2)
Low investment cost. With the PF transformer, which is low in price and mature in technology, only a few PEC modules are needed to realize active control. This significantly reduces the investment cost.

(3)
Low operating loss. With the PF transformer, which has an operating efficiency of greater than 99%, the HPR has the advantage of high operating efficiency.

(4)
High reliability. The reliability of the electromagnetic transformer is much higher than that of the PEC device. When a PEC module fails, only the winding bypass or open circuit scheme is required, without disconnecting the transformer’s power supply.
3 Operating principle of HPR
3.1 Equivalent model of the PF transformer
The proposed HPR is modeled to reflect the coupling relationship between the PF transformer and the series and shunt PEC modules. Given the relatively mature modeling methods for series and shunt converters, which can be equivalent to controlled voltage and current sources, it mainly describes the modeling of the new PF transformer and its coupling relationship with the converters in this section. Considering the grouptype transformer as an example, the proposed threelimb fourwinding transformer is formed by adding the left limb, and series and parallel windings, as shown in Fig. 2. The voltage and current are denoted as v_{1} and i_{1} for the primary winding of the transformer yphase (y = a, b, c), v_{2} and i_{2} for the secondary winding, v_{3} and i_{3} for the series winding, and v_{4} and i_{4} for the parallel winding, respectively. Φ_{1δ}, Φ_{2δ}, Φ_{3δ}, and Φ_{4δ} denote the leakage fluxes of the primary, secondary, series, and parallel windings, respectively. Φ_{2y} and Φ_{3y} denote the magnetic fluxes of the two yokes. For simplicity, the mutual leakage fluxes among different windings are ignored. The reference directions of quantities, such as voltage, current, and magnetic flux, are shown in Fig. 2.
From the magnetic flux distribution shown in Fig. 2, the equivalent magnetic circuit (magnetic flux–magnetomotive force) model of the threelimb fourwinding transformer can be obtained as shown in Fig. 3. S_{1L}, S_{2L}, and S_{3L} denote the magnetic resistances of the center, right, and left limb cores, respectively. S_{2y} and S_{3y} denote the magnetic resistances of the yokes, while S_{1δ}–S_{4δ} denote the magnetic resistances of the leakage fluxes of the windings. F_{i} denotes the magnetomotive force of each winding. The total magnetic flux of each winding is equal to the sum of the main magnetic flux and leakage flux, as:
The magnetomotive force of each winding is given by:
where the subscript i (i = 1, 2, 3, 4) represents the number of each winding, while N_{1}, N_{2}, N_{3}, and N_{4} are the turns of the primary, secondary, series, and parallel windings, respectively.
According to Fig. 3 and based on Ohm’s law of the magnetic circuit and the principle of node magnetomotive force balance, the magnetomotive force–magnetic resistance–magnetic flux relationship of the threelimb fourwinding transformer can be expressed in matrix form as shown in (3).
Equation (3) can also be expressed in simple terms as:
To deduce the equivalent circuit of the PF transformer, Eq. (3) can be further transformed with the magnetomotive force of each winding expressed in matrix form as:
where I' denotes the winding current matrix when the turns N_{1}, N_{2}, N_{3}, and N_{4} are normalized to the number of turn N.
According to Faraday’s law of electromagnetic induction, the relationship between the fourwinding transformer port voltage V' and the magnetic flux Φ is:
Substituting (5) and (6) into (4) yields:
Equation (7) can be further transformed into:
Equation (8) is the nodal voltage equation of the threelimb fourwinding transformer. The inductance of each magnetic circuit is defined as:
From (8) and (9), the equivalent circuit model of the threelimb fourwinding transformer can be obtained as shown in (10), where Z = jωL. From (10) and Kirchhoff’s voltage and current laws, the equivalent circuit of the threelimb fourwinding transformer can be obtained as shown in Fig. 4a. With consideration of the normalization turn N, the equivalent circuit can be further transferred to Fig. 4b. It is noted that the equivalent circuit model of the threelimb fourwinding transformer is a series–parallel combination relationship in which the electrical circuit and the magnetic circuit are coupled. Given the small magnetic resistance of the core limb and yoke (i.e., the large excitation impedance corresponding to the main electromotive force of each winding) [24], if the influence of the excitation branches are ignored, i.e., branches Z_{1L}, Z_{2L}, Z_{3L}, Z_{3y}, and Z_{2y} in Fig. 4 are disconnected as shown in Fig. 4c, the winding voltages v_{1}, v_{3} and v_{2} (v_{4}) are in series, while v_{2} and v_{4} are in parallel. Therefore, the following conclusions can be drawn from the equivalent magnetic circuit model in Fig. 3 and the equivalent electrical circuit model in Fig. 4:

(1)
The primary winding N_{1} and series winding N_{3} of the new PF transformer are in series, and thus the amplitude and phase of the lowvoltage side voltage v_{2} can be adjusted when the amplitude and phase of the series winding voltage v_{3} are regulated. For a voltage dip or distortion because of an MVAC distribution grid fault, the compensation by the series winding voltage v_{3} can ensure that the lowvoltage side voltage v_{2} is unaffected, thereby improving power supply reliability.

(2)
The lowvoltage winding N_{2} and parallel winding N_{4} of the new PF transformer are in parallel, and thus the amplitude and phase of the lowvoltage side current i_{2} can be adjusted when the amplitude and phase of the parallel winding current i_{4} are regulated to realize the active power control of the LVAC and LVDC ports. In addition, in the case of reactive power, harmonic and unbalanced power quality problems in the LVAC power grid, the compensation by the parallel winding current i_{4} can ensure that the current and voltage of the MVAC port are in phase and the waveforms are sinusoidal, thereby improving power quality.
3.2 Control principle of HPR
From the modeling analysis in Sect. 3.1, the active voltage and current control of the HPR can be realized by adjusting the voltage and current of the series/parallel windings of the new PF transformer. Therefore, the basic control ideas of the HPR are as follows:

(1)
In steadystate conditions, the active power exchanged between LVAC and LVDC grids, and LVAC grid current quality, are controlled by the shunt converter, whereas the amplitude and phase of the LVAC grid voltage are controlled by the series converter.

(2)
In fault conditions of the MVAC distribution grid, the series converter outputs the compensation voltage to ensure that the LVAC grid voltage is unaffected.
In grid steadystate conditions, the power exchange among the MVAC grid and the LVAC and LVDC grids is controlled by the new PF transformer and the series/shunt PECs. When the capacity is sufficient, there is no power transmitted by the series converter, and the normalized MVAC grid voltage (\(v_{1}^{\prime}\)) is equal to the LVAC grid voltage (\(v_{2}^{\prime}\)) in terms of amplitude and phase.
Figure 5a depicts the voltage–current phasor diagram of the HPR, and Fig. 6 depicts the power flow between the HPR ports. The relationships of the active and reactive power at each port are given as:
where P_{MVac} and Q_{MVac} denote the active and reactive power involved in the interaction between the MVAC port of the HPR and the MVAC grid, respectively, with P_{MVac} > 0 indicating power being absorbed from the grid and P_{MVac} < 0 indicating power being fed back to the grid. P_{sh} and Q_{sh} denote the active and reactive power involved in the LVAC and LVDC grids interaction of the shunt converter, respectively, with P_{sh} > 0 indicating active power being transmitted from the LVDC grid to the LVAC grid and vice versa for P_{sh} < 0. P_{LVac} and Q_{LVac} denote the active and reactive power consumed by the LVAC load, respectively, with Q_{LVac} > 0 indicating inductive reactive power and Q_{LVac} < 0 indicating capacitive reactive power. At this time, the output voltage of the series converter and the output current of the shunt converter satisfy the following relationship:
In addition to compensating for the reactive power of the load with the shunt converter, the series converter can be used to regulate the phase shift between the LVAC and MVAC grid voltages to realize the reactive power regulation function of the series converter, as shown in Fig. 5b, and the power flows between the HPR ports are shown in Fig. 7. In this case, the relationships of the active and reactive power at each port are given as:
where P_{LVdc} = P_{LVdc,DG} + P_{LVdc,ESS}−P_{LVdc,Ld} denotes the active power transmitted by the LVDC port, P_{LVdc,DG} denotes the power supplied by the renewable energy, P_{LVdc,ESS} denotes the power supplied or absorbed by the storage energy, and P_{LVdc,Ld} denotes the power consumed by the LVDC load. P_{se} and Q_{se} denote the active and reactive power controlled by the series converter, respectively.
From Fig. 5b and (13), the reactive power of the LVAC load can be redistributed between the series and shunt converters in any proportion. The series converter can absorb active power while supplying reactive power, while at this time, the output voltage of the series converter and the output current of the shunt converter satisfy the following relationship:
During grid faults, the LVAC grid voltage can be kept in the rated steadystate value by adjusting the voltage injected by the series winding (\(v_{3}^{\prime}\)). The power flows among HPR ports are shown in Fig. 7, and the voltage–current phasor diagram of the HPR is presented in Fig. 8. At this time, the output voltage of the series converter satisfies the following relationship:
where \(V_{1}^{^{\prime\prime}}\) denotes the magnitude of the primary winding voltage after the sudden change of the MVAC grid voltage, while δ denotes the phase jump angle.
4 Control strategy of HPR
As the power supply equipment of a regional distribution grid, a PR should ensure the basic control of the voltage/current conversion, and power transfer of the PEC modules. In addition, a PR should coordinate the operational control of each port by comprehensively considering the characteristics of the sourcegridloadstorage device connected at each port. Also, a PR should flexibly change the operational mode to adapt to different operating environments when the external power grid fails, or the power grid is under maintenance.
In this paper, a hierarchical control strategy including the systemlevel and converterlevel is established to control the power/voltage of each port of the HPR in real time, as shown in Fig. 9. In Fig. 9, v_{MVac,abc} and i_{MVac,abc} are the voltage and current at the MVAC port, v_{LVac,abc} and i_{LVac,abc} are the voltage and current at the LVAC port, while v_{se,abc}, i_{se,abc} and v_{sh,abc}, i_{sh,abc} are the voltage and current at the AC side of the series PEC and shunt PEC, respectively. v_{LVdc} is the DC bus voltage of the LVDC grid. In the sequence decomposition part, x denotes the threephase voltage or current, and q is a phaseshift operator in the timedomain. This obtains the quadraturephase waveform (90degrees lag) of the original inphase waveform. In the PLL part, θ^{+} is the positive phase angle of the MVAC grid voltage. θ_{se} and θ_{sh} are the respective phase angles for the series PEC and shunt PEC. The superscripts ‘ + ’ and ‘−’ indicate positive and negative variables of voltage and current in the dqframe, respectively.
For systemlevel control, the power/voltage command of each port is obtained through the interaction with the power distribution grid control center to realize the coordinated control of the operational mode, fault isolation, and fault ridethrough. When the systemlevel control layer receives a scheduling instruction from the distribution grid control center, the working condition instructions are issued to the HPR controller according to the collected port information and HPR status. Even if there is no scheduling instruction, the appropriate working conditions can be selected based on the current working state of the HPR controller. The systemlevel coordinated control strategies in steadystate and grid fault conditions are shown in Figs. 10 and 11, respectively.
In steadystate conditions, the systemlevel control layer receives a scheduling instruction P_{MVac,ref}, Q_{MVac,ref} and v_{se,abcref} from the distribution grid control center, and collects each port status of the HPR, including P_{LVac}, Q_{LVac}, P_{LVdc} and the stateof charge (SOC) of the ESS. If the SOC of the ESS is within the normal range [SOC_{min}, SOC_{max}], the ESS maintains the DC bus voltage of the LVDC grid, and the shunt PEC provides the active and reactive power (P_{sh,ref}, Q_{sh,ref}) supplement or absorption when a power compensation is needed from the distribution grid control center, shown as:
It should be noted that there is no need to output the compensated voltage v_{se,abc} for the series PEC in steadystate conditions, i.e., v_{se,abcref} = 0. If the SOC of the ESS is out of the normal range, the shunt PEC has to revise the command value to maintain the DC bus voltage. If the SOC of the ESS is larger than SOC_{max}, it means too much power for the LVDC grid. The shunt PEC will transfer surplus power ∆P_{sh1} from the LVDC to the AC grid. Then, the updated active and reactive power command values can be expressed as:
If the SOC of the ESS is smaller than SOC_{min}, the shunt PEC has to absorb insufficient power ∆P_{sh2} from AC grid for maintaining the DC bus voltage. Then, the active power command values can be revised as:
In grid fault conditions, the systemlevel control layer needs to prioritize the voltage compensation function. In order to avoid the phase angle jump at the LVAC grid during grid faults, the series PEC has to compensate the sag/swell voltage so that the LVAC voltage is well compensated with respect to the presag/swell voltage. Because of the Dyn11 connection of the 10 kV / 400 V PF transformer, there exists a π/6degree shift between the MVAC and LVAC sides. Then, the desired LVAC grid voltage v_{LVac,abcref} can be expressed as:
where V_{LVac,N} is the nominal amplitude of the LVAC grid voltage.
Then, to maintain the LVAC voltage at a nominal value and achieve phase angle correction, the series PEC voltage references v_{se,abcref} can be generated as:
where N_{1}, N_{2}, N_{3}, and N_{4} are the turns of the primary, secondary, series, and parallel windings, respectively. v_{MVac,ab}, v_{MVac,bc} and v_{MVac,ca} are the line voltages of the MVAC grid. After the sequence decomposition and abc/dq transformation, the daxis and qaxis components can be respectively derived as \(v\,_{{{\text{se}},{\text{dref}}}}^{ + } /v\,_{{{\text{se}},{\text{qref}}}}^{ + }\), and \(v\,_{{{\text{se}},{\text{dref}}}}^{  } /v\,_{{{\text{se}},{\text{qref}}}}^{  }\).
To avoid overcurrent at the MVAC port during a grid fault, the transferred power P_{MVac,ref}, Q_{MVac,ref} will be revised with respect to the severity of the grid fault. We define D = V_{MVac} /V_{MVac,N} ∈ [0, 1] as a voltage dip factor representing the severity of the grid fault, where V_{MVac} and V_{MVac,N} are the actual and nominal phase voltage amplitudes of the MVAC grid, respectively. Then, the active and reactive power (P_{sh,ref}, Q_{sh,ref}) supplement or absorption by the shunt PEC in (16)–(19) are updated as:
where P_{se} = (1−D)P_{MVac,ref} and Q_{se} = (1−D)Q_{MVac,ref}. In addition, in the case of severe short circuit faults or prolonged voltage sags which exceed the system capacity, the HPR needs to cooperate with the relay protection of the MVAC grid to convert to islanded mode. In this case, loadshedding may have to be considered to maintain healthy operation of the ESS.
For the converterlevel control, the dual closedloop control method under positive and negative dq0frame is adopted [25]. The outer loop of the series converter calculates the current reference value by tracking the required injected series voltage, and the outer loop of the shunt converter calculates the current reference value based on the active and reactive power tracking. The inner loops of the series and shunt converters adopt feedforward decoupling current control. It is noted that either a proportional integral (PI) or proportional integral resonance (PIR) regulator can be used for the outer loop regulator (OLR) and the inter loop regulator (ILR) to precisely track the reference values, whereas the PI is used for the OLR and ILR in this paper. The second order generalized integrator (SOGI) method for decomposing the sequences of voltage and current signals is implemented here [26], as shown in Fig. 9. Also, the SOGIbased phaselocked loop (SOGIPLL) [26] is used to provide an effective solution for grid synchronization in grid asymmetric faulty conditions. Finally, sinusoidal pulse width modulation (SPWM) technology is used to generate the switching signals for the switches of the series and shunt converters.
The function of the ESS with a DC/DC converter in the proposed HPR is to maintain the DC bus voltage of the LVDC grid [25, 27, 28], and to provide the active power supplement or absorption when a power compensation is needed for the power fluctuation in the proposed HPR. Since the LVDC DC bus voltage is controlled by the ESS with the bidirectional DC/DC converter, the shunt converter can operate in active and reactive power control mode and the series converter in voltage control mode. The ESS control scheme with DC/DC converter is also shown in Fig. 9. The voltage loop is formed by comparing the realtime LVDC voltage v_{LVdc} with the reference LVDC voltage v_{LVdc,ref} before feeding into the PI regulator, whose output is then sent to the pulse width modulation (PWM) generator to generate the switching signals.
5 Simulation results
The distribution grid/microgrid shown in Fig. 1 is built on the MATLAB/Simulink + Plecs platform to verify the feasibility of the proposed HPR. The main parameters of the HPR are shown in Table 1. Among them, the LVAC active and reactive load is simulated with a resistor and an inductor, the storage energy is connected to the LVDC bus through the bidirectional Boost converter, and the DC load is resistive.
In the MVAC grid steadystate conditions, the simulation settings are as follows:

From 1.0 to 1.1 s, the active and reactive power of the LVAC grid are 400 kW + j300 kVar. At this time, the shunt PEC only compensates for the reactive power of the LVAC grid and the transmitted active power is 0. As the MVAC grid voltage is normal, the output voltage of the series PEC is also 0.

From 1.1 to 1.2 s, the active and reactive power of the LVAC grid suddenly increase by 400 kW + j300 kVar. At this time, the shunt PEC still only compensates for the reactive power of the LVAC grid, and the transmitted active power by the shunt PEC is 0.

From 1.2 to 1.3 s, the LVDC grid transmits 300 kW active power to the LVAC grid through the shunt PEC.
Figures 12 and 13 show the dynamic simulation results of the power flows among the HPR ports. As shown in Fig. 12c, d, the LVAC grid voltage is always at the rated value under the sudden change of LVAC load. After the reactive power compensation by the shunt PEC, as shown in Fig. 12h, the voltage and current of the MVAC port are in phase, indicating only active power is provided by MVAC grid, as shown in Fig. 12b. Because no compensation voltage is required from the series PEC in steadystate, the active and reactive power transmitted by the series PEC remain at 0, as shown in Fig. 13. Therefore, the active power transmitted by the shunt PEC can be regulated in steadystate to adjust the power control between the LVAC and LVDC grids, and the power quality of the LVAC grid can be compensated by the HPR, ensuring that the voltage and current of the MVAC port are in phase with sinusoidal waveforms.
In the MVAC grid fault conditions, the simulation settings are as follows: from 1.0 to 1.1 s, the threephase voltage of the MVAC grid drops symmetrically by 50%; from 1.2 to 1.3 s, the phase A voltage of the MVAC grid drops by 50%; and at 1.25 s, the active and reactive powers of the LVAC grid increase by 400 kW + j300 kVar.
Figure 14 shows the dynamic simulation results of a conventional transformer in grid fault conditions. Because the conventional transformer has no series voltage compensation correction function, a symmetrical or asymmetrical voltage dip occurs in the LVAC grid during a symmetrical or asymmetrical MVAC grid fault, as shown in Fig. 14a, c. For the asymmetric LVAC voltage, the output threephase current of the shunt PEC without the fault ridethrough function will be severely distorted (see the current waveforms during 1.2–1.3 s in Fig. 14f), and the current at the MVAC port will also be distorted. Therefore, the fault isolation and voltage support capability cannot be achieved using the conventional transformer during the MVAC grid fault.
Figure 15 shows the dynamic simulation results of the HPR in MVAC grid fault conditions. During a symmetrical or asymmetrical voltage dip on the MVAC grid, as shown in Fig. 15a, the series PEC can be controlled to inject a support voltage to ensure that the LVAC grid voltage is always at the rated value and unaffected by the MVAC grid voltage disturbance, as shown in Fig. 15c, e, thereby ensuring highquality power supply for LVAC and LVDC users and renewable energy power generation equipment. During a symmetric or asymmetric voltage dip of the MVAC grid, the power at each port of HPR satisfies P_{MVac} + P_{sh} + P_{se} = P_{LVac}, Q_{sh} + Q_{se} = Q_{LVac}, and Q_{MVac} = 0, as shown in Fig. 16. In conclusion, the proposed HPR can realize fault isolation between MVAC and LVAC grids with the active voltage support of the series winding, thereby improving the power supply reliability.
6 CHIL and downscale experimental test results
6.1 CHIL test results
An RTLABbased CHIL is built to further validate the correctness of the proposed HPR typology and theoretical analysis. As shown in Fig. 17, the HPR and grids are modeled in the OP5600 realtime simulator. The control system of the HPR is implemented in a DSP + FPGAbased control box, with the DSP (TMS320F28335) used for the control algorithm implementation, and the FPGA (EP3C25E144I7N) for the SPWM and optical fiber communication. The control box is connected to RTLAB through the signal conversion module for optical fiber communication. The CHIL parameters are consistent with the simulation parameters.
Figure 18 shows the waveforms of the phase A voltage and current at each HPR port in grid steadystate conditions. The active and reactive power of the LVAC grid are 400 kW + j300 kVar. At this time, the LVAC current lags behind the voltage, as shown in Fig. 18a. The shunt PEC only compensates for the reactive power of the LVAC grid, so the transmitted active power is 0, as shown in Fig. 18b, while the MVAC voltage and current are in phase. Figure 19 shows the dynamic waveforms of power change of the HPR ports in steadystate conditions. The shunt PEC outputs reactive power and linearly increases the output active power by 200 kW within 0.1 s. From Fig. 19, it can be concluded that the power transmitted by the HPR ports can be regulated by the shunt PEC module.
Figure 20 shows the CHIL results of the conventional transformer in MVAC grid asymmetrical fault conditions. The active and reactive powers of the LVAC grid maintain at 400 kW + j300 kVar, and the reactive power is compensated by the shunt PEC. Because the conventional transformer has no series voltage compensation correction function, an asymmetrical voltage dip occurs in the LVAC grid in the case of a MVAC grid asymmetrical fault, as shown in Fig. 20c. In addition, if the conventional PEC is not configured with the negativesequence control loop and fault ridethrough function, the current at the MVAC port is severely distorted because of the asymmetrical fault, as shown in Fig. 20b. Therefore, the voltage support capability cannot be achieved using a conventional transformer.
Figures 21 and 22 show the dynamic results of the HPR under MVAC grid symmetrical and asymmetrical fault conditions, respectively. When the voltage of the MVAC grid drops by 50%, the series PEC can be controlled to inject a support voltage to ensure that the LVAC grid voltage remains at the rated value and is unaffected by the MVAC grid disturbance, as shown in Figs. 21 and 22. Therefore, the fault isolation and voltage support capability can be well realized using the proposed HPR during the symmetrical and asymmetrical grid faults.
To further verify the disturbance/fault isolation performance of the HPR, the MVAC grid voltage is set to contain 0.1 pu 5th and 0.05 pu 7th harmonics. As shown in Fig. 23, the MVAC grid voltage is severely distorted. After the voltage correction by the series PEC, the LVAC voltage maintains a sinusoidal waveform, as shown in Fig. 23b. The CHIL results of these two abnormal conditions agree with the theoretical analysis and simulation results, so the active voltage correction and fault isolation capabilities of the HPR are verified, ensuring the normal operation of nonfaulty ports.
6.2 Downscale experimental test results
To further prove the effectiveness of the proposed HPR, a downscale experimental prototype was built, as shown in Fig. 24. The grid fault is emulated by the programmable AC source, and the ESS is emulated by the DC source. The controller is implemented with DSP TMS320F28377. The experimental parameters are as follows: the root mean square (RMS) value of the phase voltage at the MVAC port is 60 V; the RMS value of the phase voltage at the LVAC port is 60 V; the DC bus voltage is 200 V; the turn ratios of N_{1}, N_{2}, N_{3}, and N_{4} are N_{1}: N_{2}: N_{3}: N_{4} = 1.73: 1: 1: 1; the inductor and capacitor of the LC filter are 5 mH and 22 μF, respectively.
In the downscale experimental test part, we focus on the fault isolation and voltage support performance of the HPR when the MVAC grid faults occur, thereby ensuring the healthy operation of the LVAC grid. The performance in an MVAC grid fault with 100% voltage drop of singlephase and twophase, and 80% voltage drop of threephase is investigated. Figures 25, 26 and 27 show the experimental results of the HPR in MVAC grid asymmetrical and symmetrical fault conditions with phase A 100% voltage drop, twophase 100% voltage drop, and threephase 80% voltage drop, respectively. When the voltage of the MVAC grid drops, the series PEC can be controlled to inject a support voltage (as shown in Figs. 25, 26 and 27b) to block the grid fault, ensuring that the LVAC grid operates in its prefault state and is unaffected by the MVAC grid disturbance, as shown inFigs. 25, 26 and 27c. Therefore, the fault isolation and voltage support capability can be well realized by controlling the series PEC of the proposed HPR during symmetrical and asymmetrical grid faults.
7 Conclusions
In this paper, an HPR scheme based on PF transformer electromagnetic coupling with PECs is proposed. The equivalent magnetic and electrical circuits of the novel PF transformer are established, and its working principle and port power flow are analyzed, with the control scheme presented. Based on theoretical analysis and simulation and prototype experimental tests results, the following conclusions are drawn.

(1)
The series PEC of the HPR can actively regulate the voltage, and the shunt PEC can actively regulate the current. When combined with the active controls, the converters can realize flexible control of voltage and current at each HPR port to ensure that the voltages of the LVAC and LVDC grids are unaffected by the MVAC grid fault disturbance.

(2)
With the PF transformer in the proposed HPR, only part of the power needs to be converted by the PECs of the HPR. Generally, no multilevel or multistage cascaded PEC structures are required, leading to simple topology and control complexity of the proposed HPR scheme.
Availability of data and materials
Not applicable.
Abbreviations
 MVAC:

Mediumvoltage AC
 LVAC:

Lowvoltage AC
 LVDC:

Lowvoltage DC
 HPR:

Hybrid power router
 PF:

Powerfrequency
 PEC:

Power electronic converters
 SOGI:

Second order generalized integrator
 ESS:

Energy storage system
 SOC:

Stateof charge
 CHIL:

Controller hardwareintheloop
References
Ghosh, B., Chakraborty, A. K., & Bhowmik, A. R. (2023). Reliability and efficiency enhancement of a radial distribution system through valuebased autorecloser placement and network remodelling. Protection and Control of Modern Power Systems, 8, 1. https://doi.org/10.1186/s41601022002747
Yu, K., Li, J., Zeng, X. J., et al. (2022). Insulation parameter online measurement technology based on extended zerosequence circuits for resonant grounding distribution networks. IEEE Transactions on Power Delivery, 37(6), 4679–4689.
Huang, K., Li, Y., Zhang, X., et al. (2021). Research on power control strategy of householdlevel electric power router based on hybrid energy storage droop control. Protection and Control of Modern Power Systems, 6, 13.
Huang, A. Q., Crow, M. L., Heydt, G. T., et al. (2011). The future renewable electric energy delivery and management (FREEDM) system: The energy internet. Proceedings of the IEEE, 99(1), 133–148.
Ali, S., Ling, Z., Tian, K., & Huang, Z. (2021). Recent advancements in submodule topologies and applications of MMC. IEEE Journal of Emerging and Selected Topics in Power Electronics, 9(3), 3407–3435.
Wang, D., Tian, J., Mao, C. X., et al. (2016). A 10kV/400V 500kVA electronic power transformer. IEEE Transactions on Industrial Electronics, 63(11), 6653–6663.
Li, Y., Han, J., Cao, Y., Li, Y., Xiong, J., Sidorov, D., et al. (2017). A modular multilevel converter type solid state transformer with internal model control method. International Journal of Electrical Power & Energy Systems, 85, 153–163.
Bu, Z. M., Teng, J. X., Sun, X. F., et al. (2022). Low frequency voltage ripples decoupling with switchedcapacitor conversion for an MMCbased SST. IEEE Transactions on Industrial Electronics, 69(11), 11293–11303.
Zhou, J., Zhang, J., Wang, J., Zang, J., & Cai, X. (2021). Design and control of power fluctuation delivery for cell capacitance optimization in multiport modular solidstate transformers. IEEE Transactions on Power Electronics, 36(2), 1412–1427.
Huber, J. E., & Kolar, J. W. (2014). Volume/weight/cost comparison of a 1MVA 10 kV/400 V solidstate against a conventional lowfrequency distribution transformer. In 2014 IEEE energy conversion congress and exposition (ECCE), Pittsburgh (pp. 4545–4552).
Liu, Q., & Li, Y. (2020). An inductive filteringbased parallel operating transformer with shared filter for power quality improvement of wind farm. IEEE Transactions on Power Electronics, 35(9), 9281–9290.
Li, Y., Peng, Y., Liu, F., Sidorov, D., Panasetsky, D., Liang, C., et al. (2017). A controllably inductive filtering method with transformerintegrated linear reactor for power quality improvement of shipboard power system. IEEE Transactions on Power Delivery, 32(4), 1817–1827.
Lai, J., Yin, X., Zhang, Z., Wang, Z., & Yin, X. (2019). System modeling and cascaded passivity based control for distribution transformer integrated with static synchronous compensator. International Journal of Electrical Power & Energy Systems, 113, 1035–1046.
Lei, E., Yin, X., Zhang, Z., & Chen, Y. (2017). An improved transformer winding tap injection dstatcom topology for mediumvoltage reactive power compensation. IEEE Transactions on Power Electronics, 33(3), 2113–2126.
Aeloiza, E. C., Enjeti, P. N., Moran, L. A., & Pitel, I. (2003). Next generation distribution transformer: to address power quality for critical loads. In IEEE 34th annual conference on power electronics specialist, Acapulco, Mexico (pp. 1266–1271).
Cai, G., Wang, Y., Liu, C., et al. (2020). Direct AC/AC conversion based hybrid distribution transformer with twophase supported regulation. Electric Power Automation Equipment, 40(12), 8–17.
Lee, H. J., Yoon, S. W., & Yoon, Y. D. (2022). Hybrid distribution transformer based on an existing distribution transformer and a seriesconnected power converter. IEEE Transactions on Power Delivery, 37(5), 4202–4211.
Zheng, L. R., Marellapudi, A., Chowdhury, V. R., Bilakanti, N., Kandula, R. P., et al. (2022). Solidstate transformer and hybrid transformer with integrated energy storage in active distribution grids: Technical and economic comparison, dispatch, and control. IEEE Journal of Emerging and Selected Topics in Power Electronics, 10(4), 3771–3787.
Zhao, X. J., Liu, Y. R., Chai, X. H., et al. (2023). Multimode operation mechanism analysis and power flow flexible control of a new type of electric energy router for lowvoltage distribution network. IEEE Transactions on Smart Grid, 13(5), 3594–3606.
Elsaharty, M. A., Candela, J. I., & Rodriguez, P. (2018). Power system compensation using a power electronics integrated transformer. IEEE Transactions on Power Delivery, 33(4), 1744–1754.
Chen, Z., Li, H., Liu, L., Xiang, L., & Bai, B. (2019). Dc bias treatment of hybrid type transformer based on magnetic flux modulation mechanism. IEEE Transactions on Magnetics, 55(6), 1–4.
Yang, B., Zhao, J. F., Ji, Z. D., et al. (2020). Overview of hybrid transformer technologies. Electric Power Automation Equipment, 40(02), 205–213.
Liu, Y., Liang, D., Wang, Y., Kou, P., & Yang, C. (2021). Power flow analysis and dclink voltage control of hybrid distribution transformer. IEEE Transactions on Power Electronics, 36(11), 12579–12595.
Liu, Y. B., Zhang, L. S., Liang, D. L., et al. (2022). Quasi proportionalresonant control for the hybrid distribution transformer with LCLtype converters. IEEE Transactions on Industry Applications, 58(5), 6368–6385.
Yang, R. H., & Jin, J. X. (2021). Unified power quality conditioner with advanced dual control for performance improvement of dfigbased wind farm. IEEE Transactions on Sustainable Energy, 12(1), 116–126.
Rodriguez, P., Teodorescu, R., Candela I., et al. (2006). New positivesequence voltage detector for grid synchronization of power converters under faulty grid conditions. In 37th IEEE power electronics specialists conference (pp. 1–7).
Li, Z., Sheng, W. X., Duan, Q., et al. (2019). Coordinated control strategy of AC/DC hybrid power router based on voltage stabilization by energy storage. Automation of Electric Power Systems, 43(2), 121–129.
Chen, X., Yan, L. F., Zhou, X., et al. (2018). A novel DVRESSembedded windenergy conversion system. IEEE Transactions on Sustainable Energy, 9(3), 1265–1274.
Acknowledgements
Not applicable.
Funding
This work was supported in part by the National Natural Science Foundation of China (Grant No. 52007010) and in part by State Key Laboratory of Advanced Electromagnetic Engineering and Technology (Grant No. AEET 2022KF003).
Author information
Authors and Affiliations
Contributions
JL: Conceptualization, Methodology, Software, Experiments, Investigation, Writing Original Draft, Supervision. XY: Conceptualization, Validation, Formal analysis, Visualization, Writing—Original Draft. XY: Supervision, Writing: Review. JH: Experiments, Writing: Review & Editing. FX: Writing: Review & Editing.
Corresponding author
Ethics declarations
Competing interests
The authors declare that they have no known competing financial interests or personal relationships that could have appeared to influence the work reported in this paper.
Rights and permissions
Open Access This article is licensed under a Creative Commons Attribution 4.0 International License, which permits use, sharing, adaptation, distribution and reproduction in any medium or format, as long as you give appropriate credit to the original author(s) and the source, provide a link to the Creative Commons licence, and indicate if changes were made. The images or other third party material in this article are included in the article's Creative Commons licence, unless indicated otherwise in a credit line to the material. If material is not included in the article's Creative Commons licence and your intended use is not permitted by statutory regulation or exceeds the permitted use, you will need to obtain permission directly from the copyright holder. To view a copy of this licence, visit http://creativecommons.org/licenses/by/4.0/.
About this article
Cite this article
Lai, J., Yin, X., Yin, X. et al. Typology and working mechanism of a hybrid power router based on powerfrequency transformer electromagnetic coupling with converters. Prot Control Mod Power Syst 8, 42 (2023). https://doi.org/10.1186/s41601023003168
Received:
Accepted:
Published:
DOI: https://doi.org/10.1186/s41601023003168