**1. Introduction**

Power system demands are increasing due to the increase in economic activities and national developments. With the increase in demand, power systems operators have upgraded their transmission systems to higher voltage levels in order to achieve economical and reliable power transfer between the generation side and the demand side [1]. Therefore, high voltage and extra-high-voltage substations are widely deployed in modern systems, and their number will be greatly increased in the near future [2]. Thus, gas insulated substations (GISs) have been widely used over the last three decades in power systems and other fields such as intelligent transportation systems, high-speed trains, and underground, because of their high reliability, easy maintenance, and small ground space requirements. The higher personal and operational safety level and easy installation and commissioning make high-voltage GISs in significant demand, particularly in heavily industrial areas, in comparison with conventional air insulated substations (AIS) [3–7]. Although GISs have been involved in power systems for a long time, with the development of GISs and the voltage level of transmission lines, many new problems are occurring, such as very fast transient overvoltage (VFTO) caused by SF6 discharge during switching operations in GISs. VFTO could cause damage to power devices. A reduction of the insulating capability of the dielectric gas in GISs is caused mainly by the peak magnitude and high-frequency oscillations of VFTO. The internal VFTO causes stress on the main insulation in the GIS, while the external VFTO poses a threat mainly to the main transformer and the secondary equipment within the substation [8]. These transients have extremely short rise time, in the nanoseconds range.

Several methods have been considered to mitigate such overvoltages [9,10]. VFT suppression has been proposed by many researchers using a disconnector equipped with a damping resistor [11], a disconnector with reduced voltage during the opening operation (the so-called trapped charge voltage (TCV)) [12], surge arresters [13,14], high-frequency resonators [15,16], magnetic rings of di fferent types [17–20], and VFTO mitigation by controlling the voltage conditions preceding voltage breakdown in SF6 gas [21,22]. It could be well observed from the literature that an accurate design using the specific parameters of the GIS DS (Disconnector) contact system is required for VFTO mitigation by controlling the voltage conditions preceding voltage breakdowns in the disconnector contact system. As a consequence, additional costs will increase for new combinations.

Most of the suppression methods may require a change in GIS design and construction, which makes them complicated and expensive with low reliability. Disconnecting switches fitted with damping resistors and grounding switches were developed by Yamagata, Y. et al. [23]. The amplitude of the VFTO was reduced by up to 25% after applying the damping resistor. However, the damping resistor method showed limitations because any increment of the resistance leads the VFTO to be decreased and the dimension of the disconnector to be increased. In addition, the dissipated power requirements for the resistor are increasingly high. Furthermore, the slow operating speed of the GIS DS (2–3 m/s) leads to nonability for arc extinction. As a result, the lifetime of the contactor is reduced because of these nanosecond arcs.

TCV has been noticed on the load side of the disconnector after the occurrence of the last re-strike when the disconnector opening operation is completed. Charge leakage across the insulators leads to decay of the trapped charge, which is a prolonged operation, taking hours or days. During the next closing operation of the disconnector, due to the slow contact speed, the first pre-strike occurs when the source side and the load side have the same voltage with di fferent polarity, where the load-side voltage is TCV resulting from the previous operation. The main challenge in the TCV approach is the optimum design of the disconnector which should be considered to achieve significant VFTO reduction with an acceptable sparking time, as explained clearly by Chen, W. et al. [24].

Moreover, surge arresters can suppress the amplitude of the VFTO, but they have no e ffect on the steepness. The di fficulty in using surge arresters is implementing the optimum number of arrester discs for a noticeable damping e ffect and at a convenient location. The appropriate location of surge arresters in order to eliminate VFTO was investigated by Yadav, D.N. [25].

It could be well seen from the literature [15,16] that the main disadvantage of the cavity resonator method is the absorption of VFTO energy in only a narrow band of its broad frequency spectrum. Consequently, there is no observed damping e ffect when the resonant frequency of the resonator does not fit the dominant harmonic component of the VFTO. However, compared with the ferrite ring method, the magnetic rings become saturated in high current under high frequency and lose their suppressing e ffect, as confirmed by Rama Rao, J. V. G. et al. [26]. Furthermore, ferrite magnetic rings are still in the experimental stage, and they may have no obvious suppressing e ffect on VFTO.

As a consequence, in order to ensure the reliability of substations, it is essential to carry out research on suppressing VFTO. This paper presents a design of a damping busbar to suppress VFTO. We hollow out the conventional busbar to a spiral tube, and then the busbar conductor is changed into a series circuit with a multiturn hollow inductance coil and multiturn gap; based on this, paralleling the damping resistance with the newly designed spiral tube inductance circuit, the spiral tube damping busbar is formed. The mechanism of VFTO suppression by the proposed damping busbar is analyzed, and the distributed equivalent circuit is established. Furthermore, an illustrative structure of the damping busbar is introduced. VFTO with the proposed damping busbar is simulated, and the suppression e ffects before and after installing the damping busbar are compared. An improved design of the damping busbar is proposed, and a higher damping e ffect is verified.

The remainder of this paper is structured as follows. In Section 2, the suppressing mechanism of the new method is analyzed. The simulation results and the damping e ffect of the damping busbar are discussed in Section 3. In Section 4, the influence of the inductance value of the damping busbar on suppressing VFTO is investigated. An improved design of the damping busbar in order to increase the suppression e ffect is proposed in Section 5. This section presents an investigation on a spiral coil made

of litz wire, used to damp VFTO. Such winding meets the requirement of su fficient resistance to damp VFTO. Section 6 summarizes the damping e ffects of the damping busbar and the improved design. Finally, conclusions are drawn in Section 7.

#### **2. The Suppression Mechanism of the New Method**

Several studies have shown that VFTO is generated as a superposition of multiple di fferent reflected electromagnetic waves with complex nonharmonic time dependence and covers a wide frequency range from 100 kHz to 100 MHz [27]. The wavefront steepness, amplitude, and high-frequency components of VFTO can be suppressed by inductance and resistance. Based on this, this research proposes a VFTO suppression method—the spiral tube damping busbar—located before the disconnector switch (DS), as shown in Figure 1b. At the rated frequency, the damping busbar transmits current and voltage waves like a normal busbar, but when VFTO passes through the damping busbar, it will be suppressed. Also, the energy of the travelling waves is consumed by the resistance. When the resistance and inductance of the damping busbar match each other, the biggest wave energy consumption can be obtained and the best VFT damping can be achieved.

**Figure 1.** The new damping busbar installed in a gas insulated substations (GIS): (**a**) general view of the damping busbar with its components; (**b**) the placement of the damping busbar.

#### *2.1. The Structure of the Damping Busbar*

The conventional GIS busbar was hollowed out into a spiral tube, and the busbar conductor was changed into a series circuit with a multiturn hollow inductance coil and a multiturn gap; then, we paralleled the damping resistance with the spiral tube inductance circuit. This method increases the active losses of the busbar by changing the wave impedance of the busbar. As a result, the transient energy is consumed, and the amplitude of the VFTO is reduced. Thus, the main components of the spiral tube damping busbar are (a) a spiral tube damping busbar made by screwing out the conventional busbar to a spiral slotted solenoid busbar; (b) solid metal oxide resistors with an antipulse voltage function as the noninductive damping resistors, as shown in Figure 2; and (c) electrical connection tools.

**Figure 2.** The structure of the damping busbar.

#### *2.2. The Equivalent Circuit of the Damping Busbar*

The equivalent circuit of the damping busbar is resistance and inductance connected in series with the GIS bus, as shown in Figure 3. *Li*(*i* = 0,1,2, ... ) represents the inductance of each unit coil of the metal spiral tube conductor. *Ri* and *LRi*(*i* = 0,1,2, ... ) represent the noninductive resistance and its residual inductance of the parallel connection of each turn, respectively, and the value of the resistance can be adjusted to achieve the best damping effect. *gi*(*i* = 0,1,2, ... ) is the hollowing gap of the damping busbar and *ri*(*i* = 0,1,2, ... ) is the arcing resistance of the gap, which represents the losses formed by the discharge channel.

**Figure 3.** The distributed parameters of the equivalent circuit of the damping busbar.

Inductance Calculation of the Damping Busbar at High Frequency-Damping Busbar Parameters

The purpose of this section is to calculate a fairly precise value of the inductance. Determining the inductance was carried out by studying the effect of high-frequency fields with all the geometrical parameters of the busbar [28]. To this end, a calculation method for the self-inductance of the damping busbar at high frequency was developed and analyzed by the finite element method (FEM). In order to design the damping busbar, we needed to use a simulation method to obtain the VFTO waveform and then calculate the VFTO distributions. Thus, the mathematical expression of the VFTO waveform was theoretically calculated using the curve fitting method to ge<sup>t</sup> the fitting data of the VFTO waveform [29]. Then, the curve fitting method was applied to ge<sup>t</sup> the fitting data of the associated very fast transient current (VFTC) waveform for a single SF6 gas discharge. Due to VFTC being assigned as current excitation for the simulation procedures, the VFTC equation was obtained for Fourier 8, with goodness of fit parameters R-square = 0.9768 and adjusted R-square = 0.9759.

*i*

*ai*

*bi*

ω

$$\text{VFTC}(t) = a\_0 + \sum\_{i=1}^{8} a\_i \cdot \cos(i\omega t) + \sum\_{i=1}^{8} b\_i \cdot \sin(i\omega t) \tag{1}$$

Table 1 illustrates the coe fficients of the VFTC equation.

> fficients of the

> > −7.241

 2540

 403.3

 642.6

**Table 1.** Coe

1025

0

1.372 × 10<sup>7</sup>

veryfast equation.**Values 012345678**

transient current (VFTC)

−28.8

−808.4  46.59

161.9  23.15

−859.5 −1149

1063  1881

−1489

809.5

−3334


**Figure 4.** VFTC for a single breakdown at 0.0195 s.

The simulation results show a decrease in the inductance value of the damping busbar taking place for di fferent frequencies, as illustrated in Figure 5. Notably, however, Figure 5 shows that the inductance value of the damping busbar is 0.33656 mH at 2 MHz. As a consequence, the damping busbar parameters are illustrated in Table 2.

**Figure 5.** Inductance calculation of the damping busbar at high frequency.



#### **3. Simulation Results**

#### *3.1. Modelling of a 1000 kV GIS*

The equivalent circuit diagram of a 1000 kV GIS was simulated by using EMTP (Electromagnetic Transients Program) simulation in order to study the VFTO with and without installing the damping busbar [30]. However, the high-frequency characteristics of VFTO led to simulating GIS components as capacitances dominating the other parameters. In order to model the GIS bus duct, distributed parameters and lumped elements can be utilized. Surge impedance and wave velocity could be calculated for a GIS section of any length by using the physical dimensions of the bus duct [31,32]. In the following equations, C and L are self-capacitance and inductance, Z is surge impedance, and V is wave velocity in the GIS [33,34]:

$$\mathcal{C} = \frac{2\pi\varepsilon\_0\varepsilon\_r}{\ln R/r} \tag{2}$$

$$L = \frac{\mu\_0 \mu\_r}{2\pi} \ln \frac{2R}{r} \,\mathrm{}^{\prime}\tag{3}$$

$$Z = \sqrt{\frac{L}{C}} = \frac{1}{2\pi} \sqrt{\frac{\mu\_0 \mu\_r}{\varepsilon\_0 \varepsilon\_r}} \ln \frac{R}{r} \,\tag{4}$$

$$V = \frac{1}{\sqrt{LC}}.\tag{5}$$

Figure 6 shows the equivalent circuit of a 1000 kV GIS using EMTP simulation after installing the equivalent circuit of the damping busbar, as illustrated in Table 2.

Let r be the inner radius of the GIS shell and R be the inner radius of the GIS bus. In addition, ε0 = 8.85418782 × 10−<sup>12</sup> F/m, εr = 1.0024 and μ0 = 4π × 10−<sup>7</sup> H/m. According to the dimensions of the GIS busbar, the calculations showed that the propagation velocity is 295 m/μs and the surge impedance is about 93.382 ohm/m.

**Figure 6.** The equivalent circuit of a 1000 kV GIS after installing the damping busbar.
