At present, electrostatic dust removal equipment is an internationally recognized high-efficiency dust removal device. The high-voltage power supply system is the leading electrical part of the device. Its output voltage directly affects the working effect of the electrostatic precipitator. Generally, high-voltage power supplies for electrostatic precipitators can be divided into three categories: power frequency power supplies, high-frequency power supplies, and pulsed high-voltage power supplies. Due to its low frequency, the power frequency electrostatic precipitator power supply has low power density, significant voltage fluctuation, and low dust removal efficiency. It cannot adapt to the working conditions of high concentration dust and high specific resistance dust. High-frequency power supplies have been improved for this problem [1]. The method adopts high-frequency control to achieve the goals of high power density, small voltage ripple, and high dust removal efficiency. However, it has not been widely used due to its stability and other reasons. Both the power frequency power supply and the high-frequency power supply belong to the high voltage DC power supply.

From the principle of electrostatic precipitator, the pulse power supply is more suitable. It has the following advantages: the pulse voltage is short and less prone to flashover [2]. The high pulse voltage peak value increases the dust charge and improves dust removal efficiency. The pulse power supply is more beneficial for collecting high specific resistance dust. It can suppress the back corona phenomenon and improve the dust removal efficiency.

The development of electrostatic precipitator pulse power supply has formed two topologies of high-voltage side pulse and low-voltage side pulse. The load of the electrostatic precipitator varies greatly, and the pulse power supply has pulse tailing on the pack. This is a complex problem in the design of pulsed power systems. In this paper, the modal analysis is carried out for the high voltage topology of the pulsed power supply, and the differential equations in different modalities are obtained [3]. This paper conveys the effects of parameters such as the coupling inductance, coupling coefficient, and parasitic resistance of the dust chamber on the load oscillation by deriving the transfer function and three-dimensional drawing. At the same time, it provides an essential basis for designing a pulse power supply. Finally, this paper proves the correctness and feasibility of theoretical analysis and design through simulation and experiment.

The circuit topology of the high-voltage side of the electrostatic precipitator pulse power supply is shown in Figure 1. The circuit is divided into three parts: pulse power supply module, resonant cavity, and DC power supply module.

The pulse power supply module consists of a DC power supply _{PS}_{PS}_{f}_{f}_{DC}_{DC}_{PS}_{DC}_{σ}_{1} and DC-side leakage inductance _{σ}_{2}. The quantitative relationship is shown in formulas (1) to (2):

Figure 2 shows the equivalent decoupling circuit. After a steady-state, the course can be comparable to three meshes. Where _{1}, _{2}, _{3} is the current of the three meshes. After the thyristor is turned on, the inductor L and the load capacitor _{f}_{1} first flows through the thyristor to zero and then through the diode to turn off the thyristor. The diode freewheels to zero, and the resonance process ends. During the resonance process, the current _{1} changes in a sine wave. The voltage on the load capacitor is a DC superimposed cosine wave voltage [6]. The mesh current _{2}, _{3} also changes during the resonance process, but due to the sizeable mutual inductance M, the difference between _{2}, _{3} and _{1} is two orders of magnitude. Therefore, its influence is ignored in the resonance process.

Figure 3 shows the oscillation mode after the end of the resonance. After the resonance is over, the thyristor has been turned off in the freewheeling diode stage. Therefore, the branch where the resonant inductor is located is disconnected. The five elements of mutual inductance M, leakage inductance _{σ}_{1}, _{σ}_{2}, _{f}

The load voltage waveform with reasonable system parameters is shown in Figure 4. The load voltage overshoot in the oscillation phase is slight, and the regulation time is short. Figure 5 shows the load voltage waveform with unreasonable parameter design. The significant fluctuation of the DC voltage causes the pulse voltage not to match the expected value. This waveform is not easy to control, seriously damaging the system [7]. Therefore, we must carry out mathematical modeling and detailed parameter design of the system.

Mode 1 is the resonance state. We found five state variables from Fig. 2 to construct the differential equation (4)

In formula (4), _{Cf}_{C}_{1} is the resonant inductor resistance. The initial value of the resonant state circuit is equation (5). The system of differential equations (4) is a higher-order differential equation containing five unknowns. In this paper, we need to use the Laplace transform and bring in the initial value (5) to obtain the numerical solutions of _{1} (_{2} (_{3} (_{Cf}_{C}

Mode 2 is the oscillation state: the initial value of the oscillation state circuit is the final value of the resonant state circuit. We can calculate the resonance period _{1} from equations (7) to (9). We bring _{1} into _{2} (_{3} (_{Cf}_{C}_{2} (_{1}), _{3} (_{1}), _{Cf}_{1}) and.

Figure 3 shows that four state variables can be established in the differential equation system (6). Oscillation state differential equation system contains four unknowns high-order differential equation system. Similarly, we need to obtain numerical solutions to use the Laplace transform and inverse transform.

The load current and voltage in the resonant state are mainly related to _{f}_{CP}_{1} of the resonant inductance, and the initial value of Sinian. Since the oscillation state contains five energy storage elements, the numerical analysis method analyzes the system parameters [8]. The specific parameters are shown in Table 1.

System Parameters

Parameter name | Numerical value |
---|---|

Pulse side voltage VPS/V | 380 |

DC side voltage VDC/V | 790 |

Load resistance Rf/Ω | 950 |

Load capacitance Cf/nF | 100 |

Resonant inductance L/μH | 101.3 |

Coupling capacitor C/μF | 1 |

Resonance period T1/μs | 20 |

Resonant inductor resistance R1/Ω | 0.6 |

_{f}_{f}

_{1}

From the analysis of mode 1, it can be known that the more significant _{1} is, the more resonance energy loss is. At this time, the difference between the final value of the resonance voltage of _{Cf}_{C}

_{CP}

The design of the coupled inductor _{CP}

It can be seen from Fig. 7 that the oscillation state can be divided into oscillation 1 and oscillation 2. Oscillation 1 determines the overshoot of the load voltage, which is determined by the load voltage excitation _{CfV}_{MV}_{σ}_{2}_{V}_{MV}

We extract the parts of _{CfV}_{MV}_{σ}_{2}_{V}

The experimental parameters are shown in Table 1. The observed waveform is shown in Figure 8. _{CP}_{PS}_{2} represents the DC side pulse supply current. The resonance phase _{2} is increased by 0.2A. This is more than two orders of magnitude different from the 15A of the resonant current. This also shows that the system parameter design is more reasonable.

In this paper, the modal analysis of the circuit on the high-voltage side of the electrostatic precipitator power supply is carried out, and its mathematical model is constructed. We obtain the final resonance value of each part according to the Laplace transform circuit. The excitation of coupled inductor voltage and pulse-side leakage inductor current affects the load voltage regulation time.

#### System Parameters

Parameter name | Numerical value |
---|---|

Pulse side voltage VPS/V | 380 |

DC side voltage VDC/V | 790 |

Load resistance Rf/Ω | 950 |

Load capacitance Cf/nF | 100 |

Resonant inductance L/μH | 101.3 |

Coupling capacitor C/μF | 1 |

Resonance period T1/μs | 20 |

Resonant inductor resistance R1/Ω | 0.6 |

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