Eccentricity Fault Identification in Round Rotor Synchronous Motors considering Load Variation

Published by Bashir Mahdi EBRAHIMI1, Jawad FAIZ1, Mohammad ETEMADREZAIE2, Mojtaba BABAIE3, Center of Excellence on Applied Electromagnetic Systems, School of Electrical and Computer Engineering, University of Tehran, Tehran, Iran(1), Department of electrical and computer engineering, Delft, Zuid-holland, Netherlands(2), Faculty of Engineering, Science and Research Campus, Islamic Azad University, Tehran, Iran(3)

Abstract. Impacts of load variation have not been so far investigated on the eccentricity fault diagnosis in synchronous motors. In this paper, a synchronous motor under static and dynamic eccentricities with different load levels is modeled using winding function method (WFM) and finite element method (FEM). Self and mutual inductances of the stator windings, stator currents and torque are calculated and analyzed. Spectrum of the stator currents is utilized for eccentricity fault detection, its type recognition and its degree determination. Variation of eccentricity degrees and load levels on the selected indices is scrutinized separately and simultaneously. The accuracy of the obtained simulation results is verified by FEM and experimental results.

Streszczenie. Zbadano wpływ obciążenia na błąd diagnozowania ekscentryczności w silniku synchronicznym. Modelowano silnik używając metody WFM (Winding function method) i metodę elementu skończonego. Analizowano indukcyjności własna i wzajemną stojana oraz moment. Analizę widmową prądu stojana użyto do oceny ekscentryczności. Oceniono dokładność diagnozowania. (Błąd oceny ekscentryczności w silniku synchronicznym o różnym obciążeniu)

Keywords: synchronous motor, WFM, FEM, fault diagnosis.
Słowa kluczowe: silnik synchroniczny, ekscentryczność.

Introduction

Synchronous machines are essential and valuable parts of power systems. These machines are generally well constructed and robust. However, they are subjected to a wide variety of abnormal operations due to stress involved in electromechanical energy conversion process. Operation of synchronous machines under fault conditions disturbs their performances and declares their life spans. Also, persistent faults damages these machines and consequences outage time for repairing is costly. Thus, fault detection and condition monitoring of the synchronous machines allow more flexibility in operation by knowing the performance and extend machine life by adjusting the operation to avoid known operating regimes or ranges and cost effectiveness. Consequence of many electrical and mechanical faults occurring during the operation of electrical machines is the eccentricity between the rotor and stator [1]. Eccentricity is categorized into three general groups: static eccentricity (SE), dynamic eccentricity (DE) and mixed eccentricity (ME).

In [2], two air gap search coils method applied to diagnose the static and dynamic eccentricity faults in the synchronous generators. It has been shown that the odd multiple harmonics of the fundamental frequency present at the EMF of search coils in the presence of the SE fault. It is noticeable that the stator currents and the shaft signals of the synchronous generators are two essential parameters that so far have been used to analyze synchronous machines under eccentricity fault. Unbalanced flux-linkages on the shaft caused by the asymmetrical condition can generate shaft signals. Variations of magnetic properties, tolerances in physical dimensions and any asymmetrical induced currents in the winding or stray paths may cause unbalanced flow of fluxes around the shaft [3, 4]. Also, sectional frames and segmental punching are the major cause of asymmetries of synchronous machine [5]. The rotor eccentricity is one of the prominent causes of magnetic asymmetries and shaft voltage. It has been shown that the shaft signals reduce in tilted rotor condition due to the opposite shaft flux- linkage at both ends. Also, the amplitude of the shaft signal is proportional with the eccentricities. Shaft signals of large turbo generator have been used to detect the faults [6, 7]. The shaft signals and other related methods have been frequently used for the eccentricities diagnosis of asynchronous machine [8, 9].

The measured stator currents are utilized for signature analysis in condition monitoring [10]. So far static and mixed eccentricity faults in synchronous machines have not been diagnosed using the analysis of the stator current signature. However, analysis and diagnosis of dynamic eccentricity faults using stator current signature analysis has been reported in [10, 11]. The amplitude of the 17th and 19th harmonic components as an index was utilized for eccentricity fault recognition in synchronous generators. Stator and rotor windings distribution and air-gap permeance have been taken into account in [12] while this was not considered in [10, 11]. In [13], effects of different degrees of static eccentricity (SE) and dynamic eccentricity (DE) on the synchronous generator inductances have been investigated. However, influence of load variation on the fault detection procedure has not been investigated in [13].

Eccentricity fault diagnosis in synchronous machines is subject to consider faulty generators under different conditions. These conditions are variation of load levels, and variation of static and dynamic eccentricities degrees. Impacts of these states on the nominated indices should be analyzed to evaluate indices efficiency for incisive fault detection. According to [1]-[13], albeit eccentricity fault identification has been illustrated, the eccentricity severity has not been determined in faulty synchronous machines. Furthermore, effects of load variation on the nominated indices (17th and 19th harmonic components) have not been studied.

In this paper, a synchronous motor under different degrees of SE and DE is modeled by winding function method (WFM) and finite element method (FEM). Stator currents are calculated using these approaches for processing and feature extracting. Then, spectrum of the simulated currents is evaluated to estimate the degree of eccentricity. After that, competency of the aforementioned indices is evaluated in different load levels. Finally, simulated results are validated by experimental results.

Modeling of Synchronous Motors using WFM

Winding function theory facilitates the performance analysis of electrical machines under different internal faults such as eccentricity between stator and rotor. The major part of this theory is calculation of machine inductances considering time and spatial harmonics; the first step in the evaluation of the machine inductances is determination of the magnetic permeance distribution of the air gap in the eccentricity fault condition. This distribution has an inverse relationship with the air gap length and a direct relationship with the radius. The air gap length and radius are precisely calculated in this paper. The next step is the use of Ampere’s law on a closed-path, the result defines a turn function of the motor circuits and calculates the field intensity in the air gap. The resulting field intensity depends on the air gap magnetic permeance distribution and magnetic field intensity at the origin. Application of Ampere and Gauss’s laws lead to an expression for the magnetic field intensity of the winding, which is proportional to the magnetic permeance distribution. The proportionality factor is called the winding function of the specific winding. Having the magnetic field intensity from any winding and applying the electromagnetic laws, an integral form equation is obtained for different inductances of the motor. By applying the Ampere’s law and considering the turn function [10, 11], the following equation for winding is obtained as follows:

.

Since the iron permeability, µfe , is considerably larger than 1, the mmf drop within the iron can be ignored, compared to that of the air gap. Also, because the air gap length is small, it is assumed that the magnetic field intensity within angle φ is independent of the radius and its value is equal to that at the middle of the air gap. The inductance between two arbitrary windings is:

.

where p(φ) is the air gap magnetic permeance distribution, nH(φ) is the turn function of winding H, and NK(φ) is the winding function of winding K. The last step entails computation of the motor inductances. The calculated inductance matrices lead to a quick solution of the differential equations. If such analytical expressions are not used, all inductances profiles of the motor must be calculated at several rotor angular positions and stored in the inductances matrices. For solving it, an interpolation is then carried out for all inductances. This results in a longer computation time and lower computations accuracy. The method presented here is based on a manual computation of all inductances and deriving their closed-form analytical equations. Computerizing this method leads to the development of a software function that calculates the inductance matrix of the motor by receiving the mechanical position of the rotor and geometry of the eccentricity condition. After calculation of self- and mutual-inductances of the healthy and faulty motor for several rotor angular positions and storing data in the inductances matrices, computation of stator currents, torque and speed are necessary for processing and feature extracting. Hence the electromagnetic coupling model of the synchronous motor is solved using 4th- and 5th- order Runge-Kutta approach. Fig.1 depicts the per phase self-inductance of the stator winding of a synchronous motor in healthy, 10% DE and 40% DE cases. It is seen the dynamic eccentricity increases the magnitude of the self-inductance of the stator winding. As shown in Fig. 1, the high level DE distorts the self-inductance of the stator winding. The reason is that in the dynamic eccentricity case, the air gap permeance depends on the rotor angular position and this angle varies continuously.

Fig.1. Per phase self- inductance of stator winding for healthy and different faulty cases

Fig.2. Flux distribution in faulty motor cross-section at no load and rated excitation current

Modeling Synchronous Motors by TSFEM

According to [14, 15], the model of the faulty motors is the first stage of any reliable fault recognition algorithm. Thus, practical conditions of the faulty motor should be considered for accurate modeling. The modeling methods which are based on the magnetic field determination and also they consider different aspects of faulty machines can be selected as an efficacious approach to calculate required signals and parameters for processing. In this paper, the healthy and faulty synchronous motors under SE and DE are modeled using TSFEM. In this modeling, geometry of the motor elements including stator, rotor and shaft are taken in to account. Moreover, spatial distribution of the stator windings, non-uniform air-gap, physical conditions of the stator conductors, rotor, shaft and air-gap, and nonlinearity of the core materials are taken into account. Three-phase sinusoidal voltage applied to the terminals of the motor is the input of the simulation procedure. Specifications of the proposed synchronous motor have been given in Table I.

In this simulation, transient analysis of rotating machines is employed for modeling and analyzing the synchronous motor with mechanical coupling. The electrical equations due to the external circuits which exhibit supply and electrical circuits are combined with magnetic field equation in FEM and motion equations of the mechanical coupling.

Fig. 2 illustrates flux distribution in the simulated synchronous motor. As shown in Fig. 2 the eccentricity clearly affects the flux distribution within the round-rotor motor and the magnetic flux distributions are not identical on both sides of the motor in faulty condition. The reason is that the flux path reluctance in this type of motor is mainly determined by the length of the air gap. When the rotor is displaced to the stator, the length of gap between rotor poles and stator core in one ha lf of the rotor surrounding air is reduced, while rising in the other half. This leads to the creation of unbalanced magnetic reluctance paths for the flux and consequence asymmetry in flux distribution within motor. The eccentricity is too capable to more clearly affect to the total reluctance of the motors in low excitation current, in which the motors magnetic circuit are linear and saturation not occurred. Therefore, the full reluctance paths of the motors dominantly are determined by the air gap reluctance.

Fig.3. Per phase mutual- inductance of stator winding for healthy and different faulty cases

Table I. Salient Pole Synchronous Generator Parameters Used in the Simulation

.

Calculation of winding inductances is a prominent approach to analyze the electrical machine under different fault conditions. In order to predict performance of the motors under dynamic eccentricity, inductances of the stator circuits should be estimated precisely. Due to the symmetrical distribution of the windings in the motors slots, the inductance profile of stator windings will have identical variation with a phase shift in the presence of the eccentricity. Consequently, one phase of stator (phase A) is considered in this paper. Self inductance of phase A (Laa) is calculated from the flux-linkage seen by the phase (λaa), when a dc current passes through it, and other windings are open-circuited, as follows:

.

Fig. 3 demonstrates the self-inductance of phase A in the healthy and in case of different DE degrees in round-rotor synchronous motor. They have been obtained for different rotor angular positions from 0° to 360° with step-angle of 3°. According to Fig. 3, it is clear that the minimum and maximum amplitude of the self-inductance of the phase A rises with the increase of DE. It is seen 10% DE and 40% DE increase the self-inductance mean 4% and 12%, respectively.

Modeling Synchronous Motors by TSFEM

Following the calculation of the stator current using WFM and TSFEM, determination of the stator current spectrum is necessary to introduce a criterion for fault detection. The maximum required frequency is around 1000 Hz (21th harmonic component) and according to the Nyquist law, sampling frequency is set equal to 2000 Hz. However, in order to have an acceptable resolution, sampling frequency has been set to 6400 Hz (128 sample per cycle).

Fig.4. Normalized line current spectra of the motor using WFM in (top) healthy, (middle) 30% SE and (bottom) 50% SE

Fig.5. Amplitude variation of 19th harmonic component versus SE and DE degrees obtained by WFM.

Furthermore, the stator current signal is simulated over 4 seconds which allows analyzing the signals with a frequency resolution of 0.25 Hz. As shown in Fig. 4, 30% SE increases the amplitude of the 17th and 19th harmonic components about -2.05dB and -2.21dB, respectively. Furthermore, fault expansion to 50% SE raises the amplitude of the 17th and 19th harmonic components about 6 dB and -6.1dB, respectively. Comparison between amplitude of the 17th and 19th harmonic components in the stator current spectra due to SE reveals that incremental rate of the SE and DE degrees. Fig. 5 shows that the slope of AB is as same as BC. Indeed, incremental rate of the amplitude of the 17th harmonic component due to different SE degrees is as equal as DE cases. Therefore, it may be concluded that influence of the SE and DE fault on the stator currents harmonics is equal. It is related to the unbalanced magnetic pull (UMP) due to eccentricity which is larger for the SE cases. Fig. 6 exhibits the spectrum of the stator current which has been calculated by FEM.

Fig.6. Normalized line current spectra of the motor using FEM in (top) healthy, and (bottom) 50% DE

Fig.7. Normalized line current spectra of the motor using experiment in (top) healthy, and (bottom) 50% DE [13]

Comparison between Fig. 4, Fig. 6, and experimental results in Fig. 7 demonstrates that the obtained results by FEM is more precise than obtained results by WFM. The reason is the utilized approximation in the WFM in which non-linearity characteristics of the ferromagnetic materials, spatial distribution of the stator windings and stator slots effects are ignored.

Impacts of Load Variation on Fault Recognition

A. Stator Currents

Fig. 8 shows the variation of the amplitude of the 19th harmonic components versus different load levels. According to Fig. 8, amplitude of the aforementioned harmonics rises due to the increase of the eccentricity degree and it is fairly constant for load variation between 0% and 120% rated load. It is seen, that the load variation has no influence on the magnitude of the aforementioned harmonics. This may be justified based on the speed profile of the healthy and faulty synchronous motors. Since the synchronous motors rotate with a constant synchronous speed, load variation has no noticeable impacts on the amplitude of harmonic components at the same eccentricity degrees. Indeed, these criteria are robust against the load variation and it is not necessary to specify motor load for accurate fault detection. Therefore, these criteria are much efficient than the global indices which are utilized for the eccentricity fault diagnosis in induction motors. Amplitude of the aforementioned harmonic components has been summarized in Table II.

Fig.8. Amplitude variation of 19th harmonic component versus SE degrees and different load levels obtained by WFM

Fig.9. Torque profile of the full load synchronous motor (top) healthy and (bottom) 50% SE

B. Developed Torque

Fig. 9 depicts the time variation of the developed torque of the healthy and faulty motor for 50% and 100% rated load cases. As shown in Fig. 9, eccentricity raises the variation rate of the motor torque. This is due to the increase of magnitude of harmonic components in the stator currents and distortion of the airgap flux density. In addition, it is seen that load increase causes to decrease of the overshoot of the motor torque.

Modeling Synchronous Motors by TSFEM

In this paper, WFM and FEM were used to simulate the healthy and faulty synchronous motors under different SE and DE degrees. It was illustrated that SE and DE have the identical impacts on the inductances and stator currents spectra of the synchronous motors. In addition, it was demonstrated that load variation has no influence on the 17th and 19th harmonic components of the stator currents spectra. Indeed, these criteria are robust against load variation which is the noticeable feature for accurate fault recognition in different conditions of the motor.

Table II. Amplitude of the 17th Harmonic Components for different SE degrees and different Loads

.

REFERENCES

[1] B. M. Ebrahimi and J. Faiz, “Diagnosis and performance analysis of three-phase permanent magnet synchronous motors with static, dynamic and mixed eccentricity,” IET Electric Power Applications, vol. 4, no. 1, Feb. 2010, pp. 53- 66.
[2] Stoll, R. L., Hennache, A., “Method of detecting and modeling presence of shorted turns in DC field winding of cylindrical rotor synchronous machines using two airgap search coil,” IEE Proceeding Vol. 135, No. 6, pp. 281-294, November 1988.
[3] Hsu, J. S., and Stein, J., “Effect of eccentricities on shaft signals studied through windingless rotors,” IEEE Transaction on Energy Conversion, Vol. 9, No. 3, pp. 564:571, September 1994.
[4] Hsu, J. S., and Stein, J., “Shaft signal of salient-pole synchronous machines for eccentricity and shorted-field-coil detections,” IEEE Transaction on Energy Conversion, Vol. 9, No. 3, pp. 572-578, September 1994.
[5] Alger, P. L., and Samson, H. W., “Shaft currents in electric machines,” AIEE Transaction pp. 235-245, February 1924.
[6] Verma, S. P., and Girgis, R. S., Shaft potentials and currents in large turbo generators. Report for the Canadian Electrical Association, Research & Development, Suite 580, One Westmount Square, Monteral, Quebec, H3Z 2P9, May 1981.
[7] 61. Meyer, A., Joho, R., Posedal, Z., Reichert, K., and Ammann, C., “Shaft voltage in turbosets: Recent development of a new grounding design to improve the reliability of the bearings,” Int. Conference on large high voltage electric systems, paris, August 1988.
[8] Cameron, J. R., Thomson, W. T., and Dow, A. B., “Vibration and current monitoring for detecting airgap eccentricity in large induction motor,” IEEE Proceeding, Vol. 133, No. 3, pp. 155- 163, May 1986.
[9] Pollock, G. B., and Lyles, J. F., “Vertical hydraulic genertors experience with dynamic airgap monitoring,” IEEE Transaction on Energy Conversion, Vol. 7, No. 4, pp. 660-667, December 1992.
[10] Toliyat, H. A., and Al-Nuaim, N. A., “Simulation and detection of dynamic airgap eccentricity in salient-pole synchronous machines,” IEEE Transaction on Industry Application, Vol. 35, No. 1, pp. 86-93, January/February 1999.
[11] Al-Nuaim, N. A., and Toliyat, H. A., “A novel method for modeling dynamic air-gap eccentricity in synchronous machines based on modified winding function theory,” IEEE Transaction on Energy Conversion, Vol. 13, No. 2, pp.156- 162, June 1998.
[12] Tabatabaei, I., Faiz, J., Lesani, H., and Nabavi-Razavi, M. T., “Modeling and simulation of a salient–pole synchronous generator with dynamic eccentricity using modified winding function theory,” IEEE Transaction on Magnetics, Vol. 40, No. 3, pp.1550-1555, May 2004.
[13] J. Faiz, Bashir Mahdi Ebrahimi, M. Valavi and H. A. Toliyat “Mixed eccentricity fault diagnosis in salient-pole synchronous generator using modified winding function method,” Journal of Progress In Electromagnetics Research B, vol. 11, 2009, pp. 155–172.
[14] J. Faiz, Bashir Mahdi Ebrahimi, B. Akin, and B. Asaie, “Criterion function for broken-bar fault diagnosis in induction motor under load variation using wavelet transform,” Journal of Electromagnetics, Taylor & Francis, vol. 29, May 2009, pp. 220-234.
[15] J. Faiz and Bashir Mahdi Ebrahimi, “Determination of number of broken rotor bars and static eccentricity degree in induction motor under mixed fault,” Journal of Electromagnetics, Taylor & Francis, vol. 28, Aug 2008. pp. 433 – 449.

Source & Publisher Item Identifier: PRZEGLĄD ELEKTROTECHNICZNY (Electrical Review), ISSN 0033-2097, R. 87 NR 5/2011

The Optimisation of the Usage of Electricity from a Wind Turbine in a Household

Published by Wiesław MICZULSKI1, Oliwia PERS2, Uniwersytet Zielonogórski, Instytut Metrologii, Elektroniki i Informatyki (1), ETB Solution GmbH (2) ORCID: 1. 0000-0002-8175-0078; 2. 0000-0002-2962-5261

Abstract. The article presents the concept of an algorithm that optimizes energy management in a household. The algorithm optimizes the usage of the energy from a wind turbine and ensures the greatest possible comfort of using energy-efficient smart appliances by the user.

Streszczenie. W artykule przedstawiono koncepcję algorytmu optymalizującego zarządzanie energią w gospodarstwie domowym. Algorytm optymalizuje wykorzystanie energii z turbiny wiatrowej oraz zapewnia jak największy komfort korzystania z energooszczędnych inteligentnych urządzeń przez użytkownika. (Optymalizacja wykorzystania energii elektrycznej z turbiny wiatrowej w gospodarstwie domowym).

Keywords: energy management algorithm, renewable energy sources, smart appliances, wind turbine.
Słowa kluczowe: algorytm zarządzania energią, odnawialne źródła energii, inteligentne urządzenia, turbina wiatrowa

Introduction

The energy systems in many countries are developing rapidly in order to meet the increasing demand for electricity, as well as the expectations of energy consumers. Renewable energy sources (RES) are more and more often used as basic power sources also in households. The increase in renewable energy installations in Poland and Germany has occurred as a result of the implementation of appropriate financing projects for such investments. These solutions make it possible to reduce the high costs of electricity consumption from the grid. The projects promoting the financing of different types of renewable energy are being implemented in Poland and Germany.

The usage of solar and wind energy to generate electricity for households depends heavily on the weather conditions, the season and the time of day. This has a significant impact not only on the proper functioning of the electrical grid (EG) [1] but also the costs associated with energy consumption. It also has an impact on the proper management of the operation of devices in households. The work [2] presents an overview of the latest literature on energy management solutions in households.

The article presents the concept of an energy management algorithm (EMA) in a household. In the optimization process, EMA takes into account the power generated by the wind turbine (WT), which results in the reduction of energy consumption costs from EG. EMA also provides the greatest possible comfort for the user of the energy-efficient smart appliances. The choice of a WT has been determined by increased electricity production and economic considerations compared to a solar power plant [3]. The research results presented in this article are a continuation of previous research papers, the results of which are described in [4, 5].

Energy management algorithm

The goal of the EMA (Fig. 1) is to ensure that the power generated by WT (PWT) is used as much as possible to power smart household appliances. It will also reduce the cost of consumed energy from EG. In the process of optimization, EMA also provides the greatest possible comfort for the user to use smart appliances. For this purpose, it has been assumed that each of the appliances will be assigned a priority (pr) indicating which of them will be able to change power (PA) first, will be turned off or turned on with a specific time shift (ts).

Fig.1. The block diagram of an energy management algorithm (EMA) in a household

After reading the current PWT(i) value and the sum of power of the powered on appliances in a given iteration Ph(i) and the current tariff tz, the condition is checked

.

If condition (1) is not met, iteration (i) is terminated. This means that appliances that are turned on are powered only from the WT. However, failure to meet this condition will result in further EMA operation, i.e. the calculation of ∆Ph(i) according to the formula

.

In the next step of the iteration, the simultaneous fulfillment of the condition is checked

.

where: tz – tariff time zone, ∆Phys – accepted allowable value of changes |∆Ph(i)|, limiting the number of power changes of smart appliances (PA).

If condition (3) is not met, iteration (i) is terminated. However, if this condition is met, the sub-program “Optimizing the power of smart appliances” will be started, preceded by the reading of the current parameters of active appliances. For the optimization process, the GRASP (greedy randomized adaptive search procedure) heuristic algorithm [4, 6] has been used, in which the objective function takes into account: the current PWT values and power consumed from EG (PEG), current power values of individual appliances turned on (PA), remuneration (cWT) for 1 kWh transferred from WT to EG and the cost of 1 kWh from EG (cEG) determined for the relevant tariff, ts and pr values of active smart appliances and the total number of them in the household. Compared to [4], an additional condition has been introduced in the optimization process

.

where: (Ph_EMA) – value of the sum of the power of all appliances after optimization by EMA.

Condition (4) allows for a better adjustment of the power of appliances to the power generated by the WT at the expense of a small amount of energy taken from the EG, resulting from the adopted ∆P value.

As a result of the optimization process, the following decisions can be made in relation to a smart appliances:

– disagreement with its activation,
– its activation or shift of the activation time (ts),
– change of power value,
– its deactivation.

In the sub-program “Confirmation by the user” will be displayed information about the actions taken by the EMA in relation to the smart appliances. The user may knowingly allow or reject the proposed proposal. If the user does not react to the information displayed, the EMA will automatically consent to the implementation of the proposed proposal after the set time.

Characteristics of smart appliances in the household

In order to verify the functioning of the EMA algorithm, it has been assumed that the household would use energy efficient smart appliances, equipped with systems for measuring the power consumed by each appliances (PA) (e.g. [7]), and remotely controlled to change the power consumed, turn them off and time delay (ts) of their activation.

Table 1 presents a list of smart appliances installed in a sample household for which the values of PA, ts and pr have been defined. For each appliances turned off, shown in Table 1, PA = 0 kW has been assumed. The boiler has 3 heaters installed with the power of 1.5, 1.5 and 2 kW, which allow the EMA to set one of the following PA values: 1.5, 2, 3, 3.5 and 5 kW. The air conditioner also has the option to set the PA values: 2.2, 3 and 3.5 kW. Both of these appliances turn off automatically as a result of reaching the predetermined values of the output parameters.

Table 1. Parameters of smart appliances

.

*PA – the power value depends on the individual control sub-program installed in the smart appliances.
**PA – the power value depends on the EMA decision.

It has been also assumed that the EMA can turn off appliances with pr equal to 2, 3 and 5. On the other hand, appliances with pr = 6 will not be turned off by the EMA. Each appliances will be turned on after obtaining approval from the EMA or as a result of an informed decision of the user. For washing machines and dishwashers, pr = 1 has been assumed, as these are appliances whose starting time can be shifted by the EMA to a very large extent. This will have a positive effect on the optimization process. The appliances with pr = 4 will be turned off automatically as a result of completing the program in progress. It has also been assumed that the electric car battery charger will be turned on in the night time zone. The battery charging time from 10% to 100% of its capacity is 9 hours and 40 min. The first phase of battery charging is carried out with a constant current value and lasts about 1.5 hours.

In order to manage the energy in the household, it has been proposed to build an intelligent network Home Area Networks (HAN), the diagram of which is shown in Figure 2. The basic element of this network is the energy management module (EMM) consisting of, among others, from a microcontroller with an installed EMA and a system for measuring the power generated by WT. It has been assumed that smart appliances (Tab. 1) in a household would be powered by energy generated by WT with a power of 15 kW [8]. When determining the power value of the wind turbine, the car battery charger has not been taken into account. This has been based on the assumption that the electric car battery would be charged during the night tariff, which is cheaper than the daily tariff. Also at this time, there is often a small number of household appliances turned on.

In the case of a lower production or a complete deficit of energy from the WT, it will be replenished with EG. EMA, by optimizing the usage of the energy from WT, and thus the energy costs, will accordingly reduce the energy consumed by smart appliances from EG. Smart appliances will be controlled via the HAN network. The amount of energy taken from the EG, as well as the amount of energy transferred to this grid as a result of the excess energy generated by the WT, will be measured by an smart meter (SM). The communication of the SM with EMM will also be carried out by HAN.

Fig.2. The diagram of a smart grid in a household

The results of simulation tests of the energy management algorithm

The conducted simulation tests have been aimed at checking the correctness of the developed EMA concept. In addition to the above assumptions, simulation tests have been conducted under the following additional assumptions:

– the household is located in Jena (Germany), with a strictly defined Time-of-Use (ToU) pricing program with two time zones (tz) [9]:
• tz = 1 from 0:00 to 6:00 and from 22:00 to 24:00 and the price for 1 kWh of energy (cEG_1) is 0.32 €,
• tz = 2 from 6:00 to 22:00 and the price for 1 kWh of energy (cEG_2) is 0.41 €,
– a recompense for 1 kWh of energy (cWT) transferred from WT to EG is equal to 0.092 € (applies to Jena) [10],
– ∆Phys = 0.2 kW,
– ∆P = 0.4 kW.

The simulation tests have been carried out for the period ofThe simulation tests have been carried out for the period of 1 day for a household equipped with appliances presented in Table 1. The time characteristics of these appliances have been determined on the basis of a database [11]. The results assessing the benefits of using EMA and WT for five analyzed cases are presented in Table 2.

In case 2, the EMA performed an energy optimization for tz = 2 by shifting the reported washing machine operation from 7:14 per hour 22:00 (tz = 1). At 9:06 there has been a notification that the dishwasher had started working. EMA proposed to move the operation of this appliances to the zone tz = 1 per hour. 23:00. The user did not agree to the algorithm proposal, as he had another launch in the plan. After this decision, the dishwasher has been turned on and finished its work at 9:37. At 10:09 there has been another application of the dishwasher for operation. In this case, the user agreed to move the dishwasher’s operation to 23:00. In cases 1 and 2, the same amount of energy was taken from the EG by all appliances. However, in the second case, the energy costs (Tab. 2) decreased by 1.16% (0.21 €) as a result of the operation of the EMA.

In case 3, the reduction in energy costs, compared to cases 1 and 2, has been only due to energy generation by the WT. The household profit was 6.60 €.

Table 2. Results of simulation tests for five cases

.

In case 4, the EMA used reduced the amount of energy consumed from the EG for tz = 2 by adjusting the total power of the appliances included (Ph_EMA) so that it is not greater than the PWT power by the adopted value of ∆P = 0.4 kW (Fig. 3).

Included the values of the quantities defined in the objective function, the EMA performed the following sequence of actions on the air conditioner: at 6:27 he was reported to work, but due to the low value of the power generated by the WT, it has not been turned on (Fig. 3). The consent of the EMA to turn on the air conditioner has been made at 6:42, but with PA = 2.2 kW taking condition (4) into account. At 6:47 the air conditioner turned off automatically. Later in the day, the EMA did not take any action on this appliance. In the case of the boiler, EMA performed from 6:45 to 7:32 activities to optimize Ph_EMA, so that according to condition (4), the determined value ∆P with respect to PWT changes is not exceeded. The sequence of EMA activities during this time period has been as follows: at 6:45 a boiler has been reported for operation and it has not been turned on as the air conditioner has been running at that time. After the air conditioner has been turned off automatically at 6:47 EMA turned on the boiler with PA = 2 kW and turned it off at 6:50 after turning on the kettle, which could have been turned on according to condition (4). At 6:53 the EMA turned on the boiler with PA = 3 kW, and at 6:58, according to the condition (4), it reduced its power to 2 kW. Failure to take into account the condition (4) in the algorithm would cause its reaction at 6:53 to not turning on the boiler with PA = 3 kW but with PA = 2 kW. Such action would extend the operation of the boiler. At 7:09 the EMA again increased the boiler capacity to 3 kW. After turning on the washing machine at 7:15 EMA reduced the boiler power to 1.5 kW, by 7:19 increased to 2 kW; and at 7:24 turned off the boiler due to the air conditioner being turned on. After the air conditioner turned off automatically, the boiler has been turned on again at 7:29 with PA = 3 kW and turned off at 7:32 due to the lower PWT power generation with the washing machine still running and the kettle turned on. The next action of the EMA has been at 14:06. When turning on the kettle, EMA reduced the boiler power from 5 kW to 2 kW, and after turning off the kettle at 14:09 the boiler power has been increased to 5 kW. At 14:13 the boiler has been turned off by the EMA because the air conditioner has been turned on and it turned off automatically at 14:20. The EMA then turned on the boiler with PA = 5 kW again, and after 1 minute it turned off automatically. At 16:16, with the air conditioner and the EMA kettle on, he changed the boiler power from 5 kW to 3 kW for a period of 1 minute. As a result of the above EMA actions, in case 4, the final profit for the household (Tab. 2) was higher than in case 3 by 3.93% (0.27 €).

In case 5 (Tab. 2), between 6:00 and 24:00 there has been a large power generation by WT (PWT > Ph) which resulted in no EMA operation for tz = 2. Moreover, a high profit (16.35 €) has been achieved in the analyzed household.

Fig.3. Illustration of EMA operation for case 4

Summary

The obtained results from simulation studies shows that the developed EMA concept makes it possible to optimize the usage of the energy generated by WT, which transferred into the costs of energy used in the household. This effect has been obtained by introducing an additional condition (4) in the optimization process, ensuring the best possible usage of the power generated by WT (PWT) to power smart household appliances and improving the comfort of using these appliances by the user. The introduction of specific priority values for individual appliances in the optimization process also affects the achievement of the greatest possible user comfort in using the appliances.

An additional effect of EMA may be the educational function in terms of shaping the user behavior profile in terms of appropriate activation of household appliances when there is the energy generation from RES.

REFERENCES

[1] Powroźnik P., Szcześniak P., Turchan K., Krysik M., Koropiecki I., Piotrowski K., An Elastic Energy Management Algorithm in a Hierarchical Control System with Distributed Control Devices. Energies, 15 (2022), 4750, https://doi.org/10.3390/en15134750
[2] Shewale A., Mokhade A., Funde N., Bokde N.D., A Survey of Efficient Demand-Side Management Techniques for the Residential Appliance Scheduling Problem in Smart Homes, Energies, 15, (2022), 2863, https://doi.org/10.3390/en15082863
[3] Bątkiewicz-Pantuła M., Comparative analysis of production costs of electricity generated from various renewable energy sources, Przegląd Elektrotechniczny, 95 (2019), nr 2, 71-74
[4] Powroźnik P., Szulim R., Miczulski W., Piotrowski, K., Household Energy Management, Appl. Sci., 11 (2021), 1626. https://doi.org/10.3390/app11041626
[5] Powroźnik P., Szcześniak P., Piotrowski K., Elastic Energy Management Algorithm Using IoT Technology for Devices with Smart Appliance Functionality for Applications in Smart-Grid, Energies, 15 (2022), 109, https://doi.org/10.3390/en15010109
[6] Pitsoulisi L., Resende M.G., Greedy randomized adaptive search procedurę, in Handbook of Applied Optimization, Pardalosi P.M., Resende M.G., Oxford University Press: Oxford, UK, (2002), 168–181
[7] https://www.fibaro.com/en/products/wall-plug/
[8] http://www.s-und-w-energie.de
[9] https://www.stromauskunft.de/de/stadt/stromanbieter-injena.html
[10]Erneuerbare-EnergienGesetz,https://www.erneuerbareenergien.de/EE/Redaktion/DE/Standardartikel/FAQ/faq_eeg_2009.html
[11] http://archive.ics.uci.edu/ml/datasets/ElectricityLoadDiagrams20112014

Authors: D.Sc. Wiesław Miczulski, University of Zielona Góra, Institute of Metrology, Electronics and Computer Science, str. Prof. Szafrana 2, 65-965 Zielona Góra, E-mail: W.Miczulski@imei.uz.zgora.pl; M.Sc. Oliwia Pers, ETB Solution GmbH, Dessauer Straße 8, 06-886 Lutherstadt Wittenberg, E-mail: oliwia.pers@etb-solution.com.

Source & Publisher Item Identifier: PRZEGLĄD ELEKTROTECHNICZNY, ISSN 0033-2097, R. 98 NR 11/2022. doi:10.15199/48.2022.11.28

Improvement of Energy Quality through the Application of BtB STATCOM in a Power System

Published by 1. Kheira BELHACEL1, 2. Mohamed BOUDIAF2, 3. Mohamed MOUDJAHED3, 4. Abderrahmane BERKANI4, L2GEGI Laboratory, Electrical Engineering Department, Ibn Khaldoun University, Tiaret, Algeria (1, 3, 4). LAADI Laboratory, Electrical Engineering Department, ZIANE Achour University, Djelfa, Algeria (2)

Abstract. Currently and in the last few years, the subject of integrating FACTS in a power system will have more importance within industrial scientific research. This is mainly due to the liberalization of the electricity sector and the development of power electronics. Very often, the quality of electrical transmission is restricted by constraints of voltage setting and the maximum transmissible power of the lines. These constraints can be overcome by the creation of new lines. However, creating new lines is not always possible for various reasons. The implementation of FACTS devices and more particularly of the BtB STATCOM system constitute an alternative to the creation of new lines. It can lead to the strengthening of the power system and the improvement of the energy quality. It is this solution that we have examined in this work which shows that the BtB STATCOM device improves, in steady state conditions, the performances of a power system such as the reduction of voltage drops and power losses in the electrical transmission lines.

Streszczenie. Obecnie iw ostatnich latach temat integracji FACTS w systemie elektroenergetycznym będzie miał coraz większe znaczenie w przemysłowych badaniach naukowych. Wynika to głównie z liberalizacji sektora elektroenergetycznego i rozwoju energoelektroniki. Bardzo często jakość przesyłu energii elektrycznej jest ograniczona ograniczeniami nastawy napięcia i maksymalnej dopuszczalnej mocy linii. Te ograniczenia można przezwyciężyć, tworząc nowe linie. Jednak tworzenie nowych linii nie zawsze jest możliwe z różnych powodów. Wdrożenie urządzeń FACTS, aw szczególności systemu BtB STATCOM, stanowi alternatywę dla tworzenia nowych linii. Może prowadzić do wzmocnienia systemu elektroenergetycznego i poprawy jakości energii. Właśnie to rozwiązanie, które zbadaliśmy w niniejszej pracy, pokazuje, że urządzenie BtB STATCOM poprawia w warunkach stanu ustalonego działanie systemu elektroenergetycznego, takie jak redukcja spadków napięć i strat mocy w liniach elektroenergetycznych. (Poprawa jakości energii poprzez zastosowanie BtB STATCOM w systemie elektroenergetycznym)

Keywords: FACTS Devices, BtB STATCOM, Power System, Electrical Network, Energy quality, Electrical Transmission.
Słowa kluczowe: FACTS Urządzenia, BtB STATCOM, System elektroenergetyczny, Sieć elektryczna, Jakość energii, Przesył elektryczny.

Introduction

Independent of the structure of a power system, the power flows throughout the electrical network are largely distributed as a function of transmission line impedance, then a transmission line with low impedance enables larger power flows through it than does a transmission line with high impedance. The problems of network operation are as many as they are varied. Both large and small power disturbances can cause power system instability. Failure to control power flows can lead to high power losses in transmission lines and violation of limit voltage constraints. If no solution is found to these problems, a risk of tripping the cascade network exists and can lead to a black out [1, 2, 3].

Conventional network control means such as adjustable transformers, additional capacitors and inductors, etc. sometimes turn out to be too slow and insufficient to respond effectively to power disturbances. This constraint can be overcome by the use of devices based on power electronics. They are faster regulation devices. FACTS (Flexible Alternative Current Transmission Systems) devices are part of these tools. The FACTS concept brings together all the power electronics-based devices that improve the operation of the electrical network. With their ability to modify the apparent characteristics of the lines, the FACTS are able to increase the capacity of the network as a whole by controlling the transits of both active and reactive power [4, 5].

Depending on the connection of FACTS to the electrical networks, the FACTS can be classified in three types : the series type such as Static Synchronous Series Compensator (SSSC) and Thyristor Controlled Series Capacitor (TCSC), the parallel type such as Static Synchronous Compensator (STATCOM) and Static Var Compensator (SVC), the hybrid type such as Unified Power Flow Controller (UPFC).

We are interested, in this work, in a parallel type of FACTS called Back-to-Back STATCOM (BtB STATCOM) and its action on a power system [1, 6, 7]. The BtB STATCOM is composed of a rectifier station and an inverter station which are joined back-to-back. The converters can use either conventional thyristors or the new generation of semiconductor devices such as gate turn-off thyristors (GTOs) or insulated gate bipolar transistors (IGBTs) [3, 8, 9]. The main aim of this paper is to analyze the impact of a BtB STATCOM on a power system.

Structure and Modeling of BtB STATCOM

The simplest BtB STATCOM consist of two back-toback DC-to-AC converters (the one as rectifier and the other as inverter), which are connected as shown in figure 1.The transmission regions i and j are connected through parallel coupling transformers T and the BtB STATCOM. The dc terminals of the converters are connected together via a common dc link.

With this BtB STATCOM, in addition to providing shunt reactive compensation and active power regulation, any converter can be controlled to supply real power to the common dc link from its own transmission line [1, 10].

Fig.1. Configuration of BtB STATCOM

Fig.2. Equivalent circuit of BtB STATCOM [11, 12]

The equivalent circuit of the BtB STATCOM is shown in figure 2.

The inverter and rectifier converter are represented by voltage sources Vsh1 and Vsh2 respectively. The shunt impedance is modeled by: Zshi=rshi+jLshiω (i=1 or 2).

By performing Park transformation, the AC current transmission can be described by the following equations.

.

(ish = ish1 and V = Vi) for rectifier, (ish = ish2 and V = Vj ) for inverter.

BtB STATCOM regulation

Decoupled control system has been employed to achieve simultaneous control of the inverter bus voltage and the DC link capacitor voltage in the 1st region. The rectifier converter of the BtB STATCOM provides simultaneous control of real power and AC voltage (or reactive power) in the 2nd region.

The power flow control is then achieved by using properly designed controllers to force the line currents to follow their references. It is desired to obtain a fast response with minimal interaction between the real and reactive power together with a strong damping of the resonance frequency. according to equations (1) and (2), the interaction between the current loops is caused by the ωLsh coupling term. Decoupling is achieved by feeding back this term and subtracting [13, 14, 15].

The figure 3 shows the diagram of this decoupling system.

Fig.3. PI decoupling system

The principle of this control strategy is to convert the measured three phase currents and voltages into d-q values and then to calculate the current references and measured voltages. Taking into account equations (4) and (5), we obtain below the equations (6) and (7). The superscript defines the reference quantities. The principle of this control strategy is to convert the measured three phase currents and voltages into d-q values and then to calculate the current references and measured voltages. Taking into account equations (4) and (5), we obtain below the equations (6) and (7). The superscript defines the reference quantities.

.

The network equation is given by:

.

The figure 4 shows the model control.

Fig.4. PI Control for dc voltage

.

The global diagram of control circuits for BtB STATCOM is given in the figure 5.

Fig.5. Block diagram control of BtB STATCOM

Presentation of the power system test

The power system test of the figure 6 is a Single Machine Infinite Bus (SMIB). It is made of a generator connected to an electrical network through a transformer T. The data of the system is given in the Appendix.

Fig.6. Power system test

Insertion of STATCOM on the power system

The BtB STATCOM is installed on line 2-3 (side of bus 2) as shown in the figure 7 to force the power flow in the desired direction.

Fig.7. Configuration of the power system equipped with a BtB STATCOM

We examine the effect of BtB STATCOM on the power system in the steady state condition in order to assess its performance. The power system’s behavior is also studied when the generator is equipped of the conventional regulator. We study the influence of changing BtB STATCOM setpoints on the electrical and dynamic characteristics of the power system stability electrical in small disturbance condition.

The static behavior of the power system is examined in the absence and in the presence of BtB STATCOM. The maximum admissible power transit for each line Pmax is 300 MW. We give in all simulation scenarios:

– For load C1: Active power =1000 MW and reactive power = 0MVAR,

– For load C2: Active power = variable value and reactive power = 0MVAR.

Power system simulation without BtB statcom

The following figure 8 shows the initial state of network voltages and load flows. For this configuration, the active power losses are equal to 2.3MW. We remark a nonuniform distribution of energy on the transmission lines: weak power flow on the lines 2-3 and 2-4, but the line 3-4 is more loaded.

The results shown in the figure 9 below are obtained for load C2=500MW. The figure 9 shows that there is an imbalance in the power flow and an increase in active power losses which reach 6.35MW. The power flow of the line 3-4 reachs 379.2MW which is greater than the maximum admissible power transit 300MW. Then the line 3- 4 is overloaded

Fig.8. Power flow test (Active power of load C2 = 300MW)

Fig.9. Power flow test (Active power of load C2 = 500MW)

Power system simulation with BtB STATCOM

The BtB STATCOM is installed in the power system as shown in the figures 10, 11 and 12. The voltage references and the voltage of dc link are:

V1ref = 1 pu = 230kV,
V2ref = 1pu = 230 kV,
Udc = 1 pu = 82kV.

Figure 10 shows the static behavior of the network under the following conditions: Load C2 = 300MW and PBtBSTATCOM=0MW. Note the active power transit on line 2-3 respects the order of the command P23 = P1ref = 0MW.

Consequently the load flow of line 2-4 decreases from 63,29MW to 3,122MW. Likewise, the two voltages in the two alternating sides of BtB STATCOM respect the voltage commands V1 = Vref1 = 1pu and V2 = Vref2 = 1pu. So, the STATCOM system is a capable controller to modify the transit of electrical power: Charging or discharging power lines.

Figure 11 shows the static behavior of the network under the following conditions: Load C2 = 300MW and PBtBSTATCOM=200MW. We note that the active power transmitted through line 2-3 always follows the setpoint P32=PBtBSTATCOM = 200MW, even for voltage references V1=Vref1=1pu and V2=Vref2=1pu. We also note an acceptable balance in the power flow that explains the beneficial effect of STATCOM in controlling power system. BtB STATCOM injects reactive powers into the two alternating sides to ensure the regulation of voltages V1 and V2.

Fig.10. Power flow test for: Active power of load C2=300MW and PBtBSTATCOM=0MW

Fig.11. Power flow test for: Active power of load C2=300MW and PBtBSTATCOM=200MW

Fig.12. Power flow test for: Active power of load C2=500MW and PBtBSTATCOM=200MW

Figure 12 shows the static behavior of the network under the following conditions : Load C2 = 500MW and PBtBSTATCOM=200MW.

It can be seen that all the characteristics controlled by BtB STATCOM (V1, V2, P23) respect the command (Fig. 12). By comparing figures 12 and 9, we find the major capacity of BtB STATCOM to ensure the protection of power lines against overloads:

– Improvement of power transit in the 3 power transmission lines,

– Make the power P34 to a value (284,5MW) inferior than 300 MW (Pmax), – Voltage regulation by reactive injection to the electricity network.

Power system characteristics with BtB STATCOM

The figures 13 to 17 show the behavior of the characteristics of the generator (rotor speed, load angle, electrical power responses and voltage generator) when the system is equipped with classical regulation and BtB STATCOM.

The parameters conditions are:

.

In the following figures, the red curve is for the reference and the blue curve is for the response. We observe maintaining stability and synchronism of the power system after each variation of the reference Psh1ref with slight transient disturbances.

Fig.13. Speed variation of generator (pu)

Fig.14. Load angle variation of generator (pu)

Fig.15. Active power variation of generator (pu)

Fig.16. Voltage variation of generator (pu)

BtB STATCOM control characteristics

The figures 17, 18 and 19 show the behaviors of active power, the alternating voltage magnitude and the reactive power on the inverter side. The characteristics behaviors attest to the good performance of the used configuration of the inverter. Indeed, each response follows the order with slight disturbances at each variation of the setpoint. The progressive variation of the active power makes it possible to avoid dangerous transient peaks.

Fig.17. Active power variation of inverter (pu)

Fig.18. Voltage variation in the AC bus of inverter (pu)

Fig.19. Reactive power variation of inverter (pu)

The figures 20, 21, 22 and 23 show the behaviors of the dc bus voltage, the AC voltage magnitude, the reactive power and the active power on the rectifier side. We can note also the good performance of the used configuration of the rectifier. Each characteristic response respects the order of the command with slight disturbances at each variation of the setpoint. The variation of the active power of the rectifier is confused with the active power of the inverter.

Fig.20. Voltage variation of dc Bus (pu)

Fig.21. Voltage variation in the AC bus rectifier (pu)

Fig.22. Reactive power variation of rectifier (pu)

Fig.23. Active power variation of rectifier (pu)

Conclusion

This paper has covered the topic of power flow models of FACTS controllers and assessed their role in wide network applications. We examined the behavior of the power system test in static state conditions when the power system operates without and with BtB STATCOM. The simulation results indicate that, by employing BtB STATCOM in the weakest bus of the system, the loading margin can be greatly increased. The results also show the effectiveness of BtB STATCOM in enhancing the load flow of the power system test. The transient mode analysis proves that the impact of setpoint variation at BtB STATCOM on the characteristics of the power system is considerably negligible. Then, a small disturbance is noted. These considerations make that a BtB STATCOM is a very powerful tool by its efficiency in the control of a power system.

APPENDIX

Table 1. System data

.
Bloc diagram for a representation of voltage regulation

Bloc diagram for a representation of speed regulation

Table 2. Nomenclature

.

Table 3. Symbols

.
.

REFERENCES

[1] Enrique Acha, Claudio R. Fuerte-Esquivel, Hugo AmbrizPe´rez, Ce´sar Angeles-Camacho, “FACTS Modelling and Simulation in Power Networks”, BOOK Wiley Editorial Offices : Copyright # 2004 John Wiley & Sons Ltd, The Atrium, Southern Gate, Chichester, West Sussex PO19 8SQ, England.
[2] Ruslan Ufa, Alexandr Gusev, Ahmed A. Zaki Diab, Aleksey Suvorov, Nikolay Ruban, Mikhail Andreev, Alisher Askarov, Vladimir Rudnik, Omer Abdalla, Ziad M. Ali, Ahmed Ibrahim, Raef Aboelsaud, “Analysis of application of back-to-back HVDC system in Tomsk electric power system”, Energy Reports, 6 (2020), 438-444. https://doi.org/10.1016/j.egyr.2020.01.017
[3] Ahmet M. Vural, Kamil C. Bayindir., “Transient stability improvement using quasi-multi pulse BTB-STATCOM”, Advances in Energy Research, 2 (2014), No. 1, 47-59. http://www.techno- press.org/fulltext/j_eri/eri2_1/eri0201005.pdf
[4] Yeshitela Shiferaw Maru, K. Padma, “The Optimal Power Flow Solution by Optimal Location of STATCOM Device using AHP Method”, International Journal of Innovative Technology and Exploring Engineering (IJITEE), 9 (2020), Issue. 4, 2561- 2569. DOI: 10.35940/ijitee.D1944.029420
[5] Dr R. Ilango, T. Santhosh kumar, S. Sathiesh kumar, “Load Flow & Transient Stability Analysis For IEEE 5 Bus System Using UPFC”, INTERNATIONAL JOURNAL OF SCIENTIFIC & TECHNOLOGY RESEARCH Vol. 9, Issue. 4, April 2020, pp. 1349-1352. http://www.ijstr.org/final-print/apr2020/Load-Flow-TransientStability-Analysis-Forieee-5-Bus-System-Using-Upfc.pdf
[6] IMDADULLAH, SYED MUHAMMAD AMRR , M. S. JAMIL ASGHAR, IMTIAZ ASHRAF 1, MOHAMMAD MERAJ, “A Comprehensive Review of Power Flow Controllers in Interconnected Power System Networks”, IEEE Access, 8 (2020), 18036- 18063. DOI: 10.1109/ACCESS.2020.2968461
[7] Aditya Chorghade, Vandana A. Kulkarni (Deodhar), “FACTS Devices for Reactive Power Compensation and Power Flow Control – Recent Trends”, 2020 International Conference on Industry 4.0 Technology (I4Tech) Vishwakarma Institute of Technology, Pune, India. Feb 13-15, 2020, 217-221. https://ieeexplore.ieee.org/document/9102640
[8] Young Ok Lee, Hyun Jae Kang, Youngseong Han, and Chung Choo Chung, “A Nonlinear Control for a BTB STATCOM System with Asymmetrically Structured Converters”, 2011 IEEE TRONDHEIM POWERTECH CONFERENCE, JUNE 19–23, TRONDHEIM, NORWAY. https://ieeexplore.ieee.org/document/6019219
[9] Jelena Stojkovi´c , Aleksandra Leki´c, Predrag Stefanov, “Adaptive Control of HVDC Links for Frequency Stability Enhancement in Low-Inertia Systems”, Energies, 13 (2020), Issue. 23, 1-20. https://doi.org/10.3390/en13236162
[10] Alcalá, J.; Cárdenas, V.; Espinoza, J.; Durán, M., “Investigation on the limitation of the BTB-VSC converter to control the active and reactive power flow”, Electric Power Systems Research,143 (2017), 149–162. https://doi.org/10.1016/j.epsr.2016.09.012
[11] Zeng, L.; Yao,W.; Zeng, Q.; Li, D.; Fang, J.; Ai, X.;Wen, J.; He, H., “Design and real-time implementation of data-driven adaptive wide-area damping controller for back-to-back VSCHVDC”, Int. J. Electr. Power Energy Syst. 109 (2019), 558–574. https://doi.org/10.1016/j.ijepes.2019.02.024
[12] MOHAMADREZA BARADAR, “Modeling of Multi Terminal HVDC Systems in Power Flow and Optimal Power Flow Formulations”, PHD Thesis, KTH School of Electrical Engineering, SE-100 44 Stockholm, SWEDEN 2013. http://www.diva-portal.org/smash/get/diva2:588780/FULLTEXT0.pdf
[13] Yunfeng Li; Guangfu Tang; Ting An; Hui Pang; Puyu Wang; Jie Yang; Yanan Wu; Zhiyuan He, “Power compensation control for interconnection of weak power systems by VSC-HVDC”, IEEE Transactions on Power Delivery,32 (2017), No. 4, 1964–1974. DOI: 10.1109/TPWRD.2016.2602890
[14] Wang Xiang, Ruizhang Yang, Chang Lin, Jiapei Zhou, Jinyu Wen, Weixing Lin, “A Cascaded Converter Interfacing Long Distance HVDC and Back-to-Back HVDC Systems”, IEEE Journal of Emerging and Selected Topics in Power Electronics, 8 (2020), Issue. 4, 4109 – 4121. DOI: 10.1109/JESTPE.2019.2913915
[15] Yinbiao Shu, Senior, Guangfu Tang, Hui Pang, “A Back-to-back VSC-HVDC System of Yu-E Power Transmission Lines to Improve Cross-region Capacity”, CSEE JOURNAL OF POWER AND ENERGY SYSTEMS, 6 (2020), No. 1, 64-71. DOI: 10.17775/CSEEJPES.2018.01280

Authors: Kheira BELHACEL, Laboratory of Energy Engineering and Computer Engineering (L2GEGI), Department of Electrical Engineering, Faculty of Applied Sciences, University of Tiaret, BP 78 Size Zarroura, Tiaret 14000, Algeria, E-mail: belhacelelt@yahoo.fr
Dr Mohamed BOUDIAF, Laboratory of Applied Automation and Industrial Diagnostics (L2GEGI), Department of Electrical Engineering, Faculty of Sciences and Technology, University of Djelfa, BP 3117, Moudjbara Road, Djelfa 17000, Algeria, E-mail: boudhiaf_mohamed@yahoo.fr
Prof Mohamed MOUDJAHED, Laboratory of Energy Engineering and Computer Engineering (L2GEGI), Department of Electrical Engineering, Faculty of Applied Sciences, University of Tiaret, BP 78 Size Zarroura, Tiaret 14000, Algeria, E-mail: moudjahedm@yahoo.fr
Dr Abderrahmane BERKANI, Laboratory of Energy Engineering and Computer Engineering (L2GEGI), Department of Electrical Engineering, Faculty of Applied Sciences, University of Tiaret, BP 78 Size Zarroura, Tiaret 14000, Algeria, E-mail: abderrahmane.berkani@univ-tiaret.dz

Source & Publisher Item Identifier: PRZEGLĄD ELEKTROTECHNICZNY, ISSN 0033-2097, R. 98 NR 7/2022. doi:10.15199/48.2022.07.27

A Simple LED Lamp with a Stable Luminous Flux Supplied from 230V AC and Free rom Inrush Current

Published by Jacek CHĘCIŃSKI, Zdzisław FILUS, Silesian University of Technology, Department of Electronics, Electrical Engineering and Microelectronics

Abstract. Modern replacements for 230V AC lamps with the bases E14 and E27 are mainly based on LEDs. Today the efforts of their designers are concentrated, in particular, on elimination of detrimental effects associated with the flicker characteristic of the LED lamps supplied from the 230 VAC grid. This paper presents a simple supply circuit of a converter-less LED lamp supplied from the 230 VAC grid which features a stable luminous flux and no inrush current at turn-on. The design criteria are based on ensuring the conformance to the legally binding EMC regulations.

Streszczenie. Współczesne zamienniki żarówek z trzonkami E14 i E27 oparte są głównie na diodach LED. Obecnie wysiłki projektantów tych lamp ukierunkowane są, w szczególności, na eliminację negatywnych zjawisk związanych z migotaniem lamp LED zasilanych z sieci 230 VAC. W pracy opisano prosty układ bezprzetwornicowej lampy LED ze stabilnym strumieniem światła zasilanej z sieci 230 VAC bez udaru prądu podczas załączania. Kryteria projektowe zostały oparte na zapewnieniu zgodności z uregulowaniami dotyczącymi kompatybilności elektromagnetycznej. (Prosta lampa LED ze stabilnym strumieniem świetlnym zasilana z sieci 230 VAC bez udaru prądu podczas załączania).

Keywords: SSL, LED lamp, inrush current, stability of the luminous flux
Słowa kluczowe: SSL, lampa LED, udar prądu, stabilność strumienia świetlnego

Introduction

In order to use LED lamps as replacements for ordinary bulbs with the bases E14 and E27 their designers must solve a number of problems. The most important of them concern the capability of fitting the electronic circuit into a relatively small space inside the lamp and ensuring its proper operation at an elevated temperature. Modern LEDs are characterized by a very high luminous efficacy (over 160 lm/W) and a high energetic efficiency (reaching 40%), which is understood as a ratio of the power of the emitted electromagnetic wave (in the visible light range) to the supplied electric power. To ensure the equivalent luminous flux of LED lamps which are considered as replacements for ordinary bulbs whose rated power was usually within the range 10–150 W, it should be expected that the required power of LED lamps will be included, roughly, between 2 W and 20 W. If we consider the efficiency of LEDs solely, we must assume that at least 60% of the electric power drawn by the lamp from the grid will be converted to heat. In addition, the efficiency of the supply system and aging of LEDs will also contribute to the temperature rise in the lamp, so while designing a heat sink for an LED lamp it is reasonable to assume that the whole electric power drawn by the lamp may be converted to heat. In practice, the internal temperature of 10–15 W LED lamps, whose shape is expected to imitate the old incandescent lamps, may often exceed 100°C. Due to that some problems may arise concerning the use of electronic components. Although individual electronic components and modern LEDs may operate above 100°C (with decreased durability, remaining, however, at an acceptable level), such excessive temperatures may make impossible proper operation of advanced integrated circuits (e.g. drivers for switching converters), electrolytic capacitors and cored inductive components.

Each electronic product that is to be offered on the consumer market of the European Union must conform to the respective legally binding regulations. In the case of LED replacements for tungsten lamps the most difficult task of their designers is to comply with the Electromagnetic Compatibility (EMC) Directive. In the standard harmonized with this directive [1] the allowed levels of radioelectric disturbance emissions are specified together with the acceptable limits for harmonic components present in the supply current drawn from the 230 VAC grid. To meet the requirements stated in the standard, less or more complicated EMI (Electro-Magnetic Interference) filters have to be usually used together with the chokes correcting the waveform of the supply current. It should be noticed that the requirements of the standard differ depending on the type of equipment (e.g. lighting equipment) and the active power drawn by the equipment from the 230 VAC grid.

The supply circuits used in LED lamps (replacements for conventional 230 VAC bulbs) can be basically divided into three categories. In the first category the 230 VAC line voltage is rectified and after smoothening the ripple with a capacitor it feeds a DC/DC converter operated at a high frequency [2]. Due to the principle of operation, based on commutation of voltage and current, they must include RFI (Radio Frequency Interference) filters. The second group includes designs without any switching converters in which a few tens of series connected LEDs or the so called COB (Chip On Board) modules are used, whose total operating voltage is usually within a range 150–250 VDC [2]. Such lamps also include electrolytic capacitors to reduce the ripple of the rectified 230 VAC voltage and simple current regulators to stabilize the current flowing through the LEDs. LED lamps from the third group do not include any electrolytic capacitors. They are sometimes referred to as lamps with sequential supply. The idea of their operation is based on using a few strings of LEDs which become gradually connected in series to the previously conducting ones as the rectified line voltage increases and then, as the supply voltage decreases, they are turned off in reverse order [3, 4]. The current supplying the LED strings being connected in sequence is usually shaped as a staircase waveform in order to resemble a sine wave. The most important advantages of this solution are: a close to 1 value of the Power Factor (PF), high luminous efficacy of the lamp, fairly simple design and high durability (the lamp does not include electrolytic capacitors whose lifetime at high temperatures is a few times shorter than the lifetime of LEDs). On the other hand, the essential disadvantage of such lamps, resulting directly from their principle of operation, is high variability of the generated luminous flux, known as flicker. Due to that these lamps are not recommended for general lighting applications but many efforts are made to reduce this flicker [5,6].

Today, the lamps which are the most widely available on the market as replacements for conventional bulbs are the LED lamps from the second group discussed above. Stability of the luminous flux produced by these lamps is dependent on the capacitance of the electrolytic capacitor at the output of the 230 VAC rectifier. To minimize flicker and stroboscopic effects, this capacitance should be possibly large and chosen with a decent surplus in order to counteract its reduction resulting from operation at a high temperature. However, by increasing this capacitance it may become more difficult to satisfy the restrictive requirements of the standards harmonized with the EMC Directive which describe the acceptable waveform of the supply current or the allowed limits for harmonic components. A large value of this capacitance contributes also to excessive current surges at start-up of the lighting installation. This is especially a serious problem when many lamps are being turned on at the same time. Figure 1 shows the waveform of the supply current drawn by a 13 W LED lamp (equivalent to a 100 W incandescent bulb) from the moment of its turn-on.

Fig.1. Waveform of the supply current of a 13 W LED lamp from the moment of its turn-on (as voltage across a series 1Ω resistor)

The initial peak value of the current, which is a few dozen times greater than in steady-state operation, causes problems difficult to solve by designers of lighting installations. The amplitude of the inrush current can greatly change depending on the current value of the line voltage at the very moment of turning the lamp on. Because of that current surge the contacts of switches and contactors become burnt much faster. Therefore, it is necessary to use slow circuit breakers for overcurrent protection of lighting installations or to choose their current rating greater than resulting from the power balance. Another problem appears when LED lamps are used for emergency lighting to illuminate escape routes, where the voltage 230 VAC is obtained from DC/AC converters supplied from batteries. A rectifier with a large filtering electrolytic capacitor is the worst type of load for DC/AC converters. Large amplitudes of the inrush current disturb the start-up of the converter and due to a greatly non-linear load distortions of the sinusoidal voltage appear which are difficult to eliminate [7].

Requirements concerning the waveform of the supply current for devices connected to the 230 VAC grid

The acceptable distortions (allowed limits for harmonic components) of the supply current drawn by a load connected to a single-phase power grid have been specified in regulations concerning electromagnetic compatibility. According to the standard PN-EN IEC 61000- 3-2 [1] LED lamps are classified as loads belonging to the very restrictive Class C. However, in Section 7.4.3 of the standard less restrictive requirements are specified for lighting equipment with the power rating not exceeding 25 W. Lamps with the power between 5 and 25 W must conform to at least one of the three specified sets of requirements. The easiest to comply with and the most often chosen by designers is the second set [1], according to which the third harmonic of the supply current must no exceed 86% of the fundamental current and the fifth is limited to 61%. In addition, the waveform of the input current must reach 5% of its peak value before or at 60° (t ≈ 3.33 ms for a 50 Hz line voltage), reach the peak value before or at 65° (t ≈ 3.61 ms) and not fall below 5% of the peak value before 90° (t = 5 ms). The angle values are referenced to any zero crossing of the fundamental supply voltage. The 5% current threshold is referenced to the highest absolute peak value which appears in the measurement window and phase measurements are made in the cycle which includes this absolute peak value. It should be noted that the criteria described above were specified a few dozen years ago and initially they referred to discharge lighting equipment only (mainly compact fluorescent lamps – CFL).

In turn, the third set of requirements was introduced fairly recently, so it is very likely that the lamps complying with these requirements will be allowed to be commercially offered for the longest time. In this group of requirements the total harmonic distortion THD of the supply current must not exceed 70%, and the allowed limits for the harmonics, expressed as a percentage of the fundamental current, are: the second – 5%, the third – 35%, the fifth – 25%, the seventh – 30%, the ninth and eleventh – 20% [1]. As for today, the standard PN-EN IEC 61000-3-2 does not specify any limits on distortion of the supply current for lamps with a rated power less than 5 W.

A modified method of supplying LED lamps

Figure 2a shows the circuit diagram of the proposed modification of the supply circuit for LED modules.

Fig.2. Simplified circuit diagram of a modified LED lamp (a) and the waveforms of current and voltage explaining its operation (b)

Instead of a choke shaping the waveform of the supply current (according to the second set of requirements specified in Section 7.4.3 of the standard [1]) an IC current regulator is used. In this case the waveform of the supply current becomes rectangular (Fig. 2b) and its amplitude is independent of both the capacitance and the amount of charge in the electrolytic capacitor, which filters the voltage waveform across a string of LEDs connected in series. The proposed solution completely eliminates harmful current surges when the lamp is being turned on. What is more, as it is possible to use electrolytic capacitors with much higher capacitance, the luminous flux generated by the lamp can be much better stabilized.

Although the presented circuit seems to be a very simple and obvious solution, it is not used by manufacturers or, at least, the authors of this paper have not been able to find a similar solution on the consumer market. Manufacturers often refrain from modifying their established products in order not to bear the costs of tests and acquiring the necessary certificates for a new product. In most cases, the input circuits of LED lamps offered on the market have been based on solutions transferred from CFLs. The compliance of these solutions with the standard PN-EN IEC 61000-3-2 has been declared on the basis of satisfying by them the second set of requirements presented in Section 7.4.3. For the proposed circuit it would be more appropriate to consider the conformance with the third set of requirements rather than the second one. The rectangular waveform of the supply current can be very easily decomposed into harmonic components and, then, refer their amplitudes to the maximum levels indicated in the standard. Apart from that, the third set of requirements has been introduced to the standard most recently, so it is justified to expect that it will remain in force with no modifications for the longest time.

The waveform i(t) of the supply current drawn by an LED lamp from the 230 VAC grid can be represented by the Fourier series with A0 = 0:

.

The Fourier coefficients are dependent on the moments tON and tOFF at which the current i(t) reaches its amplitude (±IA) and drops to zero, respectively, in each half-period of the sinusoidal supply voltage 230 VAC. If the capacitance C of the filtering capacitor is sufficiently large to ensure that the peak-to-peak ripple will be much smaller than the average voltage across the capacitor, we may assume that the waveform i(t) is close to symmetry and tOFF ≈ T/2 – tON (where T = 20 ms). The value of tON is dependent on the currents IA and IB adjusted in both regulators and can be found from the charge balance for the capacitor:

.

The coefficients Ak and Bk result from the equations:

.

The coefficients Bk obtained from (5) are zero for even harmonics (k = 2N, where N is a natural number), whereas odd harmonics (k = 2N – 1) are given by the following:

.

for k = 1, 3, 5, 7, 9, 11 …
The coefficient THD can be found from (7):

.

where IRMS is the RMS value of the supply current and I1 is the RMS value of its first harmonic. Figure 3 shows the percentage THD and normalized amplitudes of harmonics |Bk/B1| as functions of tON.

Fig.3. Amplitudes of harmonics of the supply current in the modified version of the LED lamp as a function of tON (a) and the range of tON which guarantees the conformance to the third set of requirements specified in Section 7.4.3 of the standard [1]

Analyzing the above plots we can notice that the third set of requirements from Section 7.4.3 of the standard [1] will be satisfied by the circuit from Fig. 2a, if the moment tON is smaller than 2.53 ms. In [8] it has been shown that in order to obtain high efficiency of an LED lamp with a single string of diodes, which is fed with a rectangular current waveform from Fig. 2b, the value of tON should be possibly high. In the case that we consider here it should be close to 2.53 ms with some margin necessary to allow for the spread of parameters of the used components, the temperature effects and the allowed variation of the grid voltage 230 VAC.

If we assume that the regulators IA and IB operate properly with a negligible voltage drop across them and the RMS value of the line voltage is 230 V, the total power drawn by the lamp and the power transferred to the LEDs with the series current regulator IB are given by (8) and (9):

.

The determination of the optimum length of the LED string is more critical here than for lamps without the filtering capacitor [8]. This length (number of diodes) should be smaller than the exact value resulting from the conduction voltage of a string for a given value of tON:

.

First of all, we should prevent the situation where the decrease in the voltage across C will cause a change in the current IB (the capacitance should be preferably large). Such a situation will not cause any significant change in the waveform of the supply current and the requirements specified in Section 7.4.3 of the standard [1] may remain satisfied. However, in the previously stable luminous flux a variable component will appear, increasing considerably an important parameter known as Percent Flicker [5, 6]. To prevent that, the number of LEDs initially calculated as a maximum number from (10) should be decreased by taking into account voltage drops across the current regulators IA and IB (the minimum values necessary for correct operation), the permissible reduction of the 230 VAC grid voltage (-10% according to the binding standard [9]), the effect of temperature on the forward voltage drop for the LEDs used in the lamp (ΔUF/ΔT) and the ripple voltage across the capacitor.

Summary

The main objective of this paper has been to prove that a simple solution of the supply circuit of LED lamps used as replacements for conventional 230 VAC bulbs (Fig. 2a), which is different from the commonly used ones, has a number of advantages. First of all, the conformance to the binding EMC standard [1], according to Section 7.4.3, has been based on the third set of requirements rather than on the less rigorous second one. The most important advantage is the possibility of using a large filtering capacitor thanks to which a stable luminous flux is guaranteed in a simple way. What is more, the inrush current is eliminated when the lamp is being turned on (Fig. 4).

Fig.4. Waveform of the supply current of a 13 W modified LED lamp from the moment of its turn-on (as voltage across a series 1Ω resistor)

The calculations and plots of harmonics contents have been verified by measurements. It should be noted that the calculations were carried out for an ideal rectangular waveform of the supply current (Fig. 2b), which is a worse case than the real one. In a real circuit it is easier to satisfy the requirements of the standard, because the slopes of the edges in the supply current waveform are finite thanks to which the amplitudes of higher harmonics are reduced. There is also no need to excessively reduce the variable component of the luminous flux. Percent Flicker equal to about 5–10% will satisfy even the rigorous requirements of the future standards referring to lighting installations that are expected to come in force soon (e.g. [6]) and is imperceptible by the human eye. The measured values of Percent Flicker for various sources of light available on the market have been shown in [10].

Finally, we should mention another effect resulting from the proposed method of supplying LED lamps. As the current charging the capacitor is limited by the regulator IA, the lamp will be turned on with some delay. After connecting the supply the capacitor needs a few periods of the line voltage until it becomes charged with a relatively small current IA to the voltage at which the string of LEDs can be turned on. This delay, which is dependent on the capacitance C and the current IA, will never be longer than just a few hundred milliseconds, so it should not be considered by any means as a disadvantage.

The results presented in this contribution are an outcome of statutory activities of the Department of Electronics, Electrical Engineering and Microelectronics

REFERENCES

[1] Standard EN IEC 61000-3-2:2019 Electromagnetic Compatibility (EMC) – Part 3-2: Limits for harmonic current emissions (equipment input current ≤ 16 A per phase)
[2] Li S., Tan S-C., Lee C-K., A Survey, Classification, and Critical Review of Light-Emitting Diode Drivers, IEEE Transactions on Power Electronics, 31 (2016), n.2, 1503-1515
[3] Ning N., Chen W.B., Yu D.J., Feng C.Y., Wang C.B., Selfadaptive load technology for multiple-string LED drivers, Electronics Letters, 49 (2013), n.18, 1170-1171
[4] Kim J., Lee J., Park S., A soft self-commutating method using minimum control circuitry for multiple-string LED drivers, Proceedings IEEE International Solid-State Circuits Conference, San Francisco, USA, 2013
[5] Gao Y., Li L., Mok P.K.T., An AC input switching-converter-free LED driver with low-frequency-flicker reduction, IEEE Journal of Solid-State Circuits, 52 (2017), n.5, 1424-1434
[6] Castro I., Vazquez A., Arias M., Lamar D.G., Hernando M.M., Sebastian J., A review on flicker-free ac-dc LED drivers for single-phase and three-phase ac power grids, IEEE Transactions on Power Electronics, 34 (2019), n.10, 10035-10057
[7] Rymarski Z., The analysis of output voltage distortion minimization in the 3-phase VSI for the nonlinear rectifier ROCO load, Przegląd Elektrotechniczny, 85 (2009), n.4, 127-132
[8] Filus Z., Checinski J., Some remarks on design of LED lamps and their AC direct drivers, Elektronika ir Elektrotechnika, 25 (2019), n.3, 39-42
[9] Standard PN-EN 60038:2012 CENELEC STANDARD VOLTAGES
[10] Chęciński J., Filus Z., Stabilność strumienia światła lamp LED zasilanych z sieci 230 VAC, Przegląd Elektrotechniczny, 94 (2018), nr.8, 5-8

Authors: Jacek Chęciński, PhD, Zdzisław Filus, PhD, DSc, prof. SUT, Silesian University of Technology, Department of Electronics, Electrical Engineering and Microelectronics, ul. Akademicka 16, 44-100 Gliwice E-mails: jacek.checinski@polsl.pl, zdzislaw.filus@polsl.pl

Source & Publisher Item Identifier: PRZEGLĄD ELEKTROTECHNICZNY, ISSN 0033-2097, R. 97 NR 2/2021. doi:10.15199/48.2021.02.08

Improving the Operation Characteristics of Non-Insulated Overhead Power Lines

Published by 1. Oleksandr KOZLOVSKYI1, 2. Dmitro TRUSHAKOV1, 3. Oleksandr SAVCHENKO2, 4. Serhiy RENDZINYAK3, 5. Vasyl KORUD3, Central Ukrainian National Technical University (1), Kharkiv Petro Vasylenko National Technical University of Agriculture (2), Lviv Polytechnic National University (3)
ORCID: 1.0000-0001-6885-5994; 2.0000-0003-0326-2383, 3.0000-0002-6401-0852, 4.0000-0003-4544-4871, 5.0000-0002-1289-3534

Abstract. Methods of protection of wires of overhead power lines of group “anti-icing” are analyzed. The expediency of applying the winding method of protection of wires of overhead lines in operation for a long time is shown. Based on numerical simulations of the hydrodynamic problem, the degree of influence of the protective coating on the wind load of the wire is analyzed. The research results showed a decrease in the modified wire’s drag coefficient with increasing wind speed.

Streszczenie. Przeanalizowano metody ochrony przewodów linii napowietrznych z grupy „przeciwoblodzeniowa”. Pokazano celowość zastosowania metody uzwojenia do ochrony przewodów linii napowietrznych eksploatowanych od dłuższego czasu. Wykorzystując symulacje numeryczne zagadnienia hydrodynamicznego analizowany jest stopień wpływu powłoki ochronnej na obciążenie żyły kabla wiatrem. Wyniki badań wykazały spadek współczynnika rezystancji zmodyfikowanej żyły wraz ze wzrostem prędkości wiatru. (Poprawa charakterystyk pracy nieizolowanych napowietrznych linii elektroenergetycznych)

Keywords: overhead power line, wind load, drag coefficient, numerical simulation.
Słowa kluczowe: linia napowietrzna, obciążenie wiatrem, współczynnik oporu, symulacja numeryczna.

Introduction

In 2020, the power distribution system of Ukraine included more than 330 000 km of overhead power transmission lines (OHTL) with a voltage of 6-35 kV [1]. The vast majority of these lines were built using all aluminium conductors (AAC) and aluminium conductor steel reinforced (ACSR) grades A and AC (European analogs, respectively, A and A/S [2]). These wires are a rope-like structure obtained by twisting individual cylindrical wires around the central one. Experience of operation of such OHTLs has shown their reduced reliability to extreme environmental influences, particularly wind and ice-wind loads. That is primarily due to the design features of the wire, due to which it has a greater drag coefficient under certain conditions than the cylinder of equivalent diameter, and its shape contributes to the accumulation of ice-frost deposits, which leads to increased mechanical loads on the OHTL.

In the 1980s, to eliminate the drawbacks of AAC and ACSR wires, the development of protected wires type SAX has begun in Finland. Today, the SAX system includes wires insulated with cross-linked polyethylene, the necessary linear armature, lightning protection, and vibration protection devices. Implementing the SAX system allowed to increase the reliability of overhead lines by reducing the adhesion of ice-frost deposits to the polymer insulation of the wire, reducing the drag coefficient, and eliminating the possibility of interphase shortcut circuits due to touching of wires and tree branches.

Today, Ukrainian manufacturers also manufacture shielded wires of type SIP-3 for 6-35 kV overhead lines, but they have not yet become widespread in the power grid. This situation appears because of high initial capital costs for the construction of new OHTLs; for example, only the most protected wires have about 1.5 times higher cost than wires of types AAC, ACSR. At the same time, a significant part of the overhead lines of 6-35 kV distribution networks have exhausted their normal service life but are in operation due to a lack of resources for their reconstruction and replacement. Therefore, there is a need to develop specific ways to improve the design of transmission line wires of distribution networks in operation.

Review of publications

All known methods of protection of OHTL wires from ice loads are divided into two large groups: “de-icing” (DI) and “anti-icing” (AI) [3].

The methods of the DI group consist in the OHTL wire release from ice-frost deposits when the values of mechanical loads on their linear elements close to the critical ones are reached. Typically, these goals bases on using the melting process of ice by artificial short-circuit currents. However, this method has not been widely used in the distribution networks of electricity companies. First of all, this happens because of the electrical equipment operation at melting ice process in modes close to the faulty ones. It can occur in some cases of forced technological breaks or interruptions in electricity supply to consumers, with the need to attract additional investment, for example, specialized equipment needed to adapt the power system with the OHTL [4, 5], as well as in the absence of a guarantee in the success of a particular smelting. In general, the power system consists of both components with lumped parameters and components with distributed parameters. Modelling the state of such a system requires the development of complex algorithms, for example, based on diakoptic approaches, and the use of modern software [6]. Thus, the methods of DI group are of little use for distribution networks that have a significant service life.

The methods of AI group AI are designed to wholly or partially prevent the formation of ice-frost deposits on the wire of the OHTL by creating a protective coating on it. Among these methods, the most common is a subgroup of so-called passive methods, i.e. those based on creating protective coatings that do not require an external source of energy for their work. The technology of applying a protective coating on the wire can be classified as follows: plasma spraying, extrusion, dyeing (lubricating), and winding [7-11].

Plasma spraying methods [7] and extrusion can be implemented only in production conditions and dyeing and winding [8-11] in the field conditions.

The painting method consists of applying unique chemical compounds to the wire through pneumatic spraying, lubrication, immersion, etc. The protective coatings created in this way can reduce the adhesion of water and ice droplets to the wire, and enter into an exothermic reaction with water droplets with intense heat release.

The main reasons that have interfered with the widespread implementation of these technologies in electric distribution networks are:

– a limited number of working cycles of “freezing-melting” of chemical compounds, and hence the need to reapply them before the start of the ice season;

– high-quality application of some protective coatings requires multi-stage preparation of the wire surface, which is difficult to implement in the field;

– the effectiveness of many protective coatings is reduced due to their interaction with natural or technogenic pollutants.

By shell methods, a rigid protective shell on the transmission line conductors is created, which has hydro-icephobic properties or promotes self-cleaning of the wire from ice-frost deposits during wet growth. A variant of the shell method is winding, consisting of a spiral winding on the surface of the tape wire, which performs a protective function [8-11].

The main advantage of the winding method is the possibility of its implementation on overhead transmission lines during operation. Technical implementation of this method can achieve lower costs by using robots based on Line-Scout technology [12].

Creating a protective shell on the surface of the wire will change its aerodynamic drag and hence change the wind load on the overhead transmission line in general. Many theoretical, laboratory and field studies were conducted to assess the influence of roughness on the drag coefficient of the OHTL wire form. As a result, new types of wires with reduced aerodynamic drag were developed [13-15]. However, they are designed for high and ultra-high voltage overhead transmission lines. Therefore, the problem of estimating the change in the load on the modified OHTL wire of the distribution power networks from wind is actual till now.

Basic researching principles

To more effectively solve the stated goal, the problem of determining the influence in wind load on the modified wire was divided into two more straightforward tasks:

1) determination of the drag coefficient CD of the modified wire;

2) assessment of the change in load P from wind pressure on the modified wire.

Wires of grades A (AC) are a rope-like structure consisting of one or more layers of wires spirally wound around the central core. The adjacent wires inside one winding touch each other at two diametrically opposite points. It is a peculiarity of the wires of these types. As a result, the outer contour of the wire consists of the sum of the upper arcs of the upper winding wires (Fig. 1, a). Therefore, the transmission line wire has a complex geometric shape, so the study of CD changes should be carried out by computer simulation.

Fig.1. General type of prototypes: a) AC-50/8 wire; b) modified wire

Statement of the first problem

Homogeneous air flow with a constant velocity v is fed perpendicular to the longitudinal axis of the sample. The flow rate of the test samples is such that the Mach number is M << 1. It is necessary to determine the drag coefficient of the shape of each test sample (Fig. 1) when changing the airflow rate in the range from 6 m/s to 35 m/s; to equate the calculated values of the component of the coefficient of aerodynamic drag along the axis x for a circular cylinder with the data of physical experiments.

The hydrodynamic problem was solved by numerical simulation in the Fluent software module of the ANSYS software package. The motion of the fluid is described by the Reynolds-Averaged Navier-Stokes (RANS) equations [16].

The following assumptions were made during the simulation:

– air is an incompressible, viscous liquid;

– fluid regime is turbulent;

– point contact between the ends of the arcs of the envelope curve of the wire AC-50/8 is replaced by arcs of fixed size;

– the shell of the modified wire is solid and has a thickness of 0.5 mm;

– the circular cylinder’s diameter is equal to the diameter of the wire AC-50/8: dc = dw = 9.6 mm;

– the effect of the sample length on the CD is neglected;

– the flow is directed perpendicular to the wire and the arc segment of the modified wire (Fig. 2, b, c).

The solution to this problem is performed in a two-dimensional flat formulation, so appropriate profiles were pre-built for each prototype.

The wire profile is based on a section of steel aluminum wire AC-50/8 (Fig. 2), which has been in operation for more than 15 years. A section usage made it possible to determine the actual height of the arcs of the wire leading round. Due to the deformation of the wires during the retraction of the wire under operation conditions, their contact points were shifted from the center of the wire by an average length of 0.26 mm.

Fig.2. Wire of the AC-50/8 type: a) section; b, c) 2D models of profiles

The modified wire profile is obtained by connecting the vertices of the rounding arcs with segments, followed by the displacement of the resulting hexagon with smoothed angles at a distance of 0.5 mm from its center.

Fig.3. Calculation domain and boundary conditions

The computational domain of the model was developed considering the experience of conducting similar numerical

experiments [15, 17, 18] and is presented in Fig. 3. The distance from the test sample to the sidewalls of the domain was assumed to be equal to 25dw. The area occupied by the test sample was excluded from the calculation.

In numerical modeling, the accuracy of calculating the drag coefficient CD of the shape of the sample is determined by the correctness of the choice of turbulence model and construction quality of the calculation grid.

In solving the problem, the Spalart-Allmaras model (SA) was chosen to consider fluid turbulence [16].

The Spalart-Allmaras model is a relatively simple one-parameter model that describes the entire flow region, including the boundary regions. It gives good results for moderately complex flows in the boundary layer under a pressure gradient in external aerodynamics problems. Its features are fairly good stability, reliability, and relatively low requirements for computing resources.

The chosen turbulence model is sensitive to the details of the calculation grid, especially in the boundary layer area. Therefore, the thickness of the first layer of elements (cells) of the grid significantly impacts the accuracy of calculations. To save computing resources, curved structured (regular) grids were used. Their quality was controlled by the dimensionless parameter of the near-wall layer y+, which was in the range of y+ ≤ 1.0, according to the recommendations presented in [16]. Thus, grid sets were created for the profile of each prototype. The obtained sets of grids were also checked for regularity by increasing their frequency until the grid convergence of the mathematical problem solution was achieved.

Solvers of the second order of accuracy were chosen to solve the model equations.

The following initial and boundary conditions were set during the simulation:

– pressure and speed communication scheme: coupled;

– the test sample is stationary during the flow;

– flow temperature (air), ta = 273.16°K;

– flow density, ρt=0 = 1.293 kg/m3;

– input condition: inlet (velocity), vx = {6, 10, 15, 20, 25, 30, 35} m/s;

– output condition: pressure outlet, pG = 0 Pa;

– condition on the surface of the sample: wall.

The drag coefficient of the shape of the samples is calculated by the well-known expression:

.

where: F is the total force acting in the direction of flow on the test sample, N; ρ is air density, kg/m3; v is air velocity, m/s; S is the characteristic area of the sample perpendicular to the airflow, m2; φ is the angle between the direction of wind flow and the surface of the wire, under the condition of the task φ = 90°, so sin φ = 1.

The numerical experiment was performed for sections: the circular cylinder with a diameter of 9.6 mm; the wire was of AC-50/8 type; modifications of the wire AC-50/8* with a wall thickness of the protective coating equal to 0.5 mm.

The mathematical model setup was performed by determining the aerodynamic characteristics of the circular cylinder and comparing the obtained results with the data defined from physical experiments, the results of which are given in [19, 20].

All prototypes: the cylinder, the overhead line wire, the hexagon with rounded faces (modernized wire) are inconveniently streamlined bodies, forming a large separation area with back-circulating flows due to a significant increase of pressure (Fig. 4, Fig. 5) [21]. The Reynolds number (Re) for all samples under selected conditions is in the range of 4.3ꞏ103-4.0ꞏ104.

Fig.4. Speed distribution in the cross-section of the wire of AC-50/8 type (v = 25 m/s)

Fig.5. Speed distribution in the cross-section of the modified wire of AC-50/8* type (v = 25 m/s)

The wrapping structure of the AC-50/8 wire in the cross-section was obtained based on the developed numerical model is presented in Fig. 4. On its front part, the zone of stagnation with zero speed is formed. On the bow, the pressure distribution corresponds to the theoretical pressure distribution when flowing through the flow of a non-viscous liquid. On the reverse side, there is a large area of separation with reverse circulation. In the depressions between the wires of the wire are formed reverse flows of low speed. The change in flow structure leads to a significant decrease in pressure. The main flow separation is observed in the upper part of the wire. At breakpoints in the depressions, reverse currents of low velocity not more than 1 m/s are formed.

In Fig. 5 the distribution of speed in the cross-section of the wire AC-50/8* is given. Compared to the AC-50/8 wire, which has six wires in the upper twist, the modified wire has no small circulation zones between the individual wires. Although it also has discrete vortices that distort mainstream flow lines. They have slightly less power. Two stagnant zones are formed on the sides of the modified wire (bottom and top), leading to flow separation.

Fig.6. Dependence CD = f(v) of the experimental samples

The change in the drag coefficients of the mold for all test specimens depending on the flow velocity is presented in Fig. 6. For the cylinder, this parameter is nonlinear and tends to local growing. Absolute error between the calculated value of the drag coefficient of a smooth cylinder shape according to the developed model and the experimental data presented in [13] varies between 3-8%.

The peculiarity of the wire of the OHTL of the conventional design is that its phenomenon of the drag crisis corresponds to smaller values of the Reynolds number than for a smooth cylinder of equivalent diameter. In this case, the obtained calculated data differ significantly from the experimental ones, so there is no decline in the obtained dependence of the CD = f(v) in the speed range 11-15 m/s (Fig. 6). The calculated values of the drag coefficient for the AC-50/8 wire at the Reynolds number above the transition number are in the range of 0.97-0.985 (Fig. 6), which are close to the results of experimental studies presented in [13] (CD = 0.98-1.0) and to the generally accepted normative value of 1.2 for maximum wind loads [22]. The drag curve of the shape of the modified wire changes linearly and is descending.

Statement of the second problem

Standard and modified wires of the overhead transmission line of failure class of 2KB type [22] of the distribution network are under load because of the wind pressure. It is necessary to determine the change in wind load when replacing the standard wire with a modified one in the range of wind speeds from 6 m/s to 27 m/s.

In the general case, the wind load on the wire can be determined from (1):

.

Then the change in wind load on the wire will be

.

or in expanded form

.

where indices 1, 2 are concerning, respectively, to the parameters of standard and modified wire; ΔS is the increase in the characteristic area of the wire due to the creation of a protective sheath of the wire ΔS = S2 – S1.

It follows from (2) that the reduction of the wind load on the wire can be achieved by reducing the roughness of its surface (CD1 > CD2) and the use of the minimum allowable wall thickness of the protective coating (ΔS → min).

The dependence of the change in load on the wire from the airflow flow calculated using the formula (2) is presented in Fig. 7.

Fig.7. Dependence of change of wind load on the wire from wind speed

Discussion

Analysis of the obtained data of the numerical experiment (Fig. 7) shows that the creation of a protective coating on the wire AC-50/8 changes the wind load on it in the range from –7.0% to +2.5%, so at wind speeds from 15 up to 25 m/s it decreases, and at higher speed begins to increase. This negative load change can be offset by optimizing the protective coating profile. It is expected that with increasing wire diameter to be modified, the effect from a decrease of CD will increase and can be increased to 30% after the point of drag crisis (wind speed range 25-40 m/s) [13].

Further research aims to determine the critical wall thickness of the protective coating (critical diameter of the modified wire), that is, the value of the thickness of the protective sheath of the wire, to which the wind load on its does not increase.

Conclusions

1. The expediency of using the winding method in creating passive protective coatings on uninsulated wires of the OHTL, which are in operation based on LineScout technology, is substantiated.

2. It is established that due to the change of the aerodynamic profile of the standard wire AC-50/8 after its modification, the coefficient of resistance of the form decreases with increasing wind speed.

3. It is established that the creation of a protective coating on the wire AC-50/8 using the winding method at low wind speeds (up to ≈25 m/s) is decreased the wind load on it to 7.0%. At high wind speeds (25-35 m/s), on the contrary, it is increasing to 2.5%. So, this method is sufficiently effective for overhead lines in operation.

REFERENCES

[1] Performance report of Energy and Utilities the National Regulatory Commission in 2019 the year (2020, May 27). Retrieved from https://www.nerc.gov.ua/data/filearch/Catalog3/Richnyi_zvit_N KREKP_2019.pdf
[2] IEC 61089, “Round wire concentric lay overhead electrical stranded conductors,” Published 1991, 65 p.
[3] Farzaneh, M., Atmospheric Icing of Power Networks, Springer Netherlands, 2014, 381 p.
[4] O l e k s a n d r А . S a v c h e n k o , О l e k s a n d r О . M i r o s h n y k , S t a n i s l a v V . D y u b k o , T a r a s S h c h u r , P a w e ł K o m a d a , K a n a t M u s s a b e k o v , Justification of ice melting capacity on 6-10 kV OPL distributing power networks based on fuzzy modeling, Przegląd Elektrotechniczny, 95 (2019), nr 5, 106-109, doi: 10.15199/48.2019.05.26
[5] K o z l o v s k y i , O . , T r u s h a k o v , D . , R e n d z i n y a k , S . , Temperature influence of load current of overhead electrical distribution networks in difficult weather conditions, Acta Technica CSAV (Ceskoslovensk Akademie Ved), 63(2018), No. 5, 701-708
[6] S t a k h i v , P . , R e n d z i n y a k , S . , H o h o l y u k , O . , Modeling of electric power systems based on diakoptic approach and parallel algorithms in modern computer tools, Przegląd Elektrotechniczny, 86 (2010), nr 1, 115-117
[7] Pat. 3276280 Japan, ICI5 H01 B 7/28, H 01 B 7/34. Snow and ice sticking reducing / light and heat shielding type coated power transmission line / Toyoda Minoru, Tomizawa Nobuo, Matsumoto Kiyoshi and et.; applicants and owners KANSAI Electric Power Co; Soko Seiren KK; Mitsubishi Heavy IND Ltd. No JP19910276280; fil. 30.10.1991; pub. 16.4.1993.
[8] H i g u c h i N . , Snow Accumulation Prevention Methods, Hokkaido Electric Power, 1972.
[9] V . N . S e d u n o v , A . V . N a c h a l o v , N . G . T s a r a n o v , e t c . , Proposals for the use of fluoroplastic surface coatings to reduce the mass of ice-frost deposits on wires and lightning protection cables, Electro, 2005, No. 5, 46-48 (in Russian)
[10] Pat. 138930 Ukraine, H02G 7/16 (2006.01). Method of forming anti-icing coating on uninsulated wires and lightning protection cables of overhead power transmission line / O. A. Kozlovskyi; applicants and owners Central Ukrainian National Technical University; fil. 03.06.2019; pub. 10.12.2019 (in Ukrainian)
[11] K o z l o v s k y i , O , T e l i u t a R . , Improvement of design of the Non-insulated wires of the operating overhead power lines, Visnyk Kharkivsʹkoho natsionalʹnoho tekhnichnoho universytetu silʹsʹkoho hospodarstva imeni Petra Vasylenka, 2019, issue 201, 21-22
[12] N . P o u l i o t , P . R i c h a r d a n d S . M o n t a m b a u l t , LineScout Technology Opens the Way to Robotic Inspection and Maintenance of High-Voltage Power Lines, IEEE Power and Energy Technology Systems Journal, March 2015, Vol. 2, No. 1, pp. 1-11, doi: 10.1109/JPETS.2015.2395388
[13] N a o s h i K i k u c h i , Y u t a k a M a t s u z a k i , H i d e o B a n s e , T a k a o K a n e k o , A k i h i r o Y u k i n o a n d H i r o t a k a I s h i d a , Development of Conductors with Reduced Wind Drag and Wind Noise for Overhead Power Transmission Lines, Furukawa Review, 2002, No. 21, 50-55
[14] D o n g Q i n g L i , Z h e n L i , Z h e n L i u , L o n g L i u , C h a n g L o n g Y a n g , We i F a n , L u Y u Y a n g a n d J i a J u n S i , Application Research on Drag Reduced Conductors for Electric Power Transmission Lines in Strong Wind Areas, Proc. of 2016 International Conference on Electronic, Information and Computer Engineering, ICEICE 2016, 26-27 April 2016, Vol. 448, March 2016, Article number 01093, doi: 10.1051/matecconf/20164401093
[15] C h a o M . , J u n Z . , M i n g n i a n W . , Y a o j u n M . , Large Eddy Simulation of Flow over a New Type of Low-Wind- Pressure Conductor Using WALE Model, Proc. of 2019 16th International Bhurban Conference on Applied Sciences and Technology, IBCAST 2019, 8-12 January 2019, 13 March 2019, pp. 811-815, Article number 8667236, doi: 10.1109/IBCAST.2019.8667236
[16] ANSYS Fluent User’s Guide R.19.2, ANSYS Inc., Canonsburg, PA, August 2018.
[17] R a v i n d r a n , M a g e s h R . , Computational Study for Analysis of the Potential for Drag Reduction for Flow around a Circular Cylinder and Cactus-Shaped Cylinders, Journal of Mechanical and Civil Engineering, 2015, 13-24
[18] H y u n A . S o n , S u n g s u L e e a n d J o o y o n g L e e , Numerical Analysis of Drag Force Acting on 2D Cylinder Immersed in Accelerated Flow, Water, Vol. 12, No. 6, 2020, A.N. 1970, doi: 10.3390/w12061790
[19] M i l t o n v a n D y k e , An Album of Fluid Motion, Moscow, Mir, 1986,184 p. (in Russian)
[20] R o n a l d L . P a n t o n , Incompressible flow, 4th Edition, Wiley, 2013, 912 p.
[21] C . D e m a r t i n o , F . R i c c i a r d e l l i , Aerodynamics of nominally circular cylinders: A review of experimental results for Civil Engineering applications, Engineering Structures, 137 (2017), 76-114, doi: 10.1016/j.engstruct.2017.01.023
[22] Regulations of arrangement of electrical installations, Kharkiv, Fort, 2017, 760 p. (in Ukrainian)

Authors: assoc. prof. Oleksandr Kozlovskyi, Central Ukrainian National Technical University, pr. Universytetskyi 8, 25006 Kropyvnytskyi, Ukraine, E-mail: kozlovskyioa@gmail.com; Assoc. prof. Dmitro Trushakov, Central Ukrainian National Technical University, pr. Universytetskyi 8, 25006 Kropyvnytskyi, Ukraine, Email: dmitro.trushakov@gmail.com; assoc. prof. Oleksandr Savchenko, Kharkiv Petro Vasylenko National Technical University of Agriculture, ul. Alchevskih 44, 61002 Kharkiv, Ukraine, E-mail: savoa@khntusg.info; D.Sc., prof. Serhiy Rendzinyak, Lviv Polytechnic National University, ul. Bandery 12, 79013 Lviv, Ukraine, E-mail: serhii.y.rendziniak@lpnu.ua; assoc. prof. Vasyl Korud, Lviv Polytechnic National University, ul. Bandery 12, 79013 Lviv, Ukraine, E-mail: vasyl.i.korud@lpnu.ua.

Source & Publisher Item Identifier: PRZEGLĄD ELEKTROTECHNICZNY, ISSN 0033-2097, R. 98 NR 2/2022. doi:10.15199/48.2022.02.06

The Latest Superconducting Short Current Limiters – Review of Selected Solutions

Published by 1. Joanna KOZIEŁ1, 2. Damian GNAT1, 3.Michał MAJKA1, 4.Yerbol Turgynebekov2,
Lublin University of Technology, Department of Electrical Engineering and Electrotechnologies, Faculty of Electrical Engineering and Computer Science(1) Taraz State University after M.Kh.Dulaty, Kazakhstan(2)
ORCID: 1.0000-0003-1682-2589, 3.0000-0002-7153-040X

Abstract. In the article after the introduction, containing the idea and principle of operation of the SFCL, selected prototypes and scientific projects (Korean and Russian) of superconducting short-circuit current limiters were described due to the dynamic development of the latest technologies and 2G HTS tapes used to build superconducting short-circuit current limiters. The article presents the construction of selected superconducting short-circuit current limiters (Korean) – SFCL 500V / 50A, and the Russian superconducting current limiter 220kV / 50kA, also the R-SFCL 10kV / 1.8kA superconducting current limiter is described. The article contains information on a SFCL with electric drive.

Streszczenie. W artykule po wstępie zawierającym ideę i zasadę działania SFCL, opisano wybrane prototypy i projekty naukowe (koreański i rosyjski) nadprzewodnikowych ograniczników prądu zwarcia ze względu na dynamiczny rozwój najnowszych technologii i taśm 2G HTS, stosowanych do budowy nadprzewodnikowych ograniczników prądu zwarcia. W artykule przedstawiono budowę wybranych nadprzewodnikowych ograniczników prądu zwarcia (koreański) – SFCL 500V/50A, oraz rosyjski nadprzewodnikowy ogranicznik prądu 220kV/50kA, opisano również nadprzewodnikowy ogranicznik prądu R-SFCL 10kV/1,8kA. Zawarto w artykule informacje odnośnie SFCL z napędem elektrycznym.(Najnowsze nadprzewodnikowe ograniczniki prądu zwarcia – przegląd wybranych rozwiązań).

Słowa kluczowe: nadprzewodnikowe ograniczniki prądu zwarcia, rezystancyjny SFCL, YBCO, RBS.
Keywords: superconducting fault current limiter, resistive SFCL, YBCO, RBS,

Introduction

Superconducting fault current limiters (SFCL) limit the surge current and the periodic component of the fault current. As a result, not only the thermal effects are reduced, but also the dynamic effects of the fault current on the devices in the network. The use of superconducting fault current limiters allows to limit the fault currents exceeding the rated currents 10 ÷ 20 times, to values not greater than 3 ÷ 6 times the rated current [1], [2], [3].

The dynamic effects of the fault currents and the values of the accompanying mechanical forces are greatest when the fault current reaches its first maximum after a fault, i.e. within 0.005s at a frequency of 50Hz. If the fault circuit is interrupted or we increase its impedance very quickly, i.e. in a time much shorter than 0.005s, the dynamic forces will not reach their first maximum, protecting the system and the power devices working in it against excessive stresses and damage [4], [5], [6].

Fig.1. Concept of operation of a superconducting current limiter [1]

During normal operating conditions of the power system, a superconducting fault current limiter containing a nonlinear element should not generate losses or their value should be minimal, while the impedance value should oscillate around zero. An SFCL protects the secured circuit against thermal and dynamic effects of the emergency current flow, introducing a high impedance into it when the permissible safe value of the current in the circuit is exceeded [1], [18]. The SFCL operation time is almost instantaneous, the emergency current amplitude does not reach the first, most dangerous maximum. After the failure of the emergency current, the limiter quickly returns to its original state characterised by a negligibly low impedance [8].

Korean superconducting fault current limiter SI-SFCL 500V/50A

In South Korea, research and an experimental project on a superconducting current limiters is conducted by the Korea Electric Power Corporation [7], [8]. The saturated iron-core superconducting fault current limiter (SI-SFCL) can significantly reduce the fault current and the load on circuit breakers in DC networks.

The main assumption of the project was to build a limiter with which it would be possible to achieve a fault current limitation of 70%.

Fig.2. Structure of SI-SFCL [7],[8]

There are three main parts in the composition of the SI-SFCL: one magnetic iron core, a copper primary coil (CPC) and a superconducting secondary coil (SSC). Simulation tests confirming the properties were performed with a 3D simulation model using the Finite Element Method. Based on the results of simulation tests, a design of the SI-SFCL on a laboratory scale was developed, intended to work in a DC power supply system of 500V, 50A. In the research experiment, the performance characteristics of each winding and the characteristics of fault current limitation by the SI-SFCL were analysed. The structure of the SI-SFCL is shown in Figure 2. The surge arrester has two coils and a magnetic core. The CPC secondary coil is made of copper and connected to a DC power system, while the SSC primary coil is made of a superconducting material connected to a DC voltage source.

The CPC and SSC generate magnetic fields in opposite directions [7], [8]. Technical data of the limiter are presented in Table 1.

Table 1. Specifications of the lab-scale SI-SFCL [7],[8]

.

While designing the core of the limiter, efforts were made to minimise the losses in the core, including eddy currents and hysteresis losses during fault current limitation. Such losses in the core cause high reluctance in the magnetic path, losses of magnetic field and energy, reduce the effect of the limiter on the fault currents. The core is made of thin laminated silicon steel which has high permeability and can reduce unwanted losses. The resistivity of the silicon steel is high and the steel laminates used are insulated from each other by a thin layer of Kapton tape, thus increasing the overall resistance of the iron core to prevent eddy current flow. 50PN470 silicon steel was used to build the core. The structure of the SI-SFCL is shown in Fig. 3 [7], [8].

Fig.3. Design of SI-SFCL [7],[8]

Russian superconducting fault current limiter 220 kV/50 kA

Another superconducting fault current limiter is the Russian 220kV/50kA limiter. The city of Moscow in Russia today has a population of 12.6 million. Between 2000 and 2019, energy consumption increased by over 59%. Reliable electricity supplies are a prerequisite for the city’s development.

The situation in Moscow is specific, because apart from the rapid increase in electricity consumption there is a short distance between the electricity producer and consumer, as many power plants are located in the city. The Moscow energy company has installed a resistive superconducting fault current limiter in its substation (220 kV class with a fault current limiting capacity of 50 kA) to verify the effectiveness of the SFCL.

Fig.4. a) Front , b) top view of the 220 kV SFCL at the substation in Moscow [9]

The superconducting fault current limiter was installed in a series with the 12 km long 220 kV cable line, which connects the western and south-western part of the municipal network and ends at gas-insulated switchgears with a switching capacity of 50 kA at Mnevniki and 63 kA at Ochakovo [9].

Table 2. Technical parameters of the limiter 220kV/50kA [9]

.

Research on the superconducting current limiter project started in 2015 and was carried out in three main stages: the engineering phase was completed in 2016, the procurement phase was finished in 2016-2018, and the construction phase was carried out in 2018 and 2019 (Fig. 4). The first launch and experimental tests were performed at the end of 2019 [9].

The superconducting fault current limiter consists of three external single-phase apparatus with a reservoir (hereinafter referred to as “phases”, Fig. 4) and a cooling system, which is located in a separate building with the dimensions of 7x14x10m3. The area occupied by the three phases of the surge arrester is similar to the area occupied by three air chokes (Fig. 4 b). The main technical data of the limiter are presented in Table 2 [9], [10].

Each of the SFCL limiter phases consists of a cryostat, high-voltage busbars and a set of superconductors (Fig. 5). The superconductor assembly consists of a set of current-limiting modules connected to high-voltage shields (corona rings).

Fig.5. The scheme of 220 kV SFCL phase [9]

The corona rings help to evenly distribute the electric field inside the device, and the second-generation high-temperature superconducting tape serves as a switching resistor [9].

Fig.6. Cooling system for 220 kV/50 kA [9],[10]

The SFCL cryostats were double-walled tanks with vacuum insulation. The walls were made of stainless steel to withstand the internal pressure. Each cryostat was leak tested with helium up to 15 bar. The special feature of these cryostats are two hatches removable from both sides, enabling maintenance, connection and disconnection of superconducting assemblies from the sleeves and other technical activities. The diameter of the holes was limited to 0.5 meters, which allowed access to the interior of the cryostat with a reasonable level of engineering difficulty. Each phase of the SFCL was equipped with its own cooling subsystem. The diagram of the limiter’s cooling system is shown in Figure 6 [9]. The cooling subsystem consists of a pressure builder, a Turbo-Brayton NeoKelvin cryochamber (rated cooling capacity 2kW at 70K, built by Taiyo Nippon Sanso), a circulation pump (Cryozone) and a pipe (Nexans) for transporting liquid nitrogen. Redundancy is ensured by bypassing the separate phases: even if one subsystem fails, the SFCL remains fully operational.

Superconducting fault current limiter R-SFCL 10 kV/1.8 kA

The project of a resistive superconducting fault current limiter 10 kV/1.8 kA was developed in Karlsruhe, Germany, on the basis of projects [11], [12]. The simulation tests of the R-SFCL application were based on the German medium voltage grid (Fig. 7) [13].

Fig.7. Medium-voltage network topology in Germany with superconducting R-SFCL current limiter installed [13]

Table 3 presents the detailed parameters of the MV network fragment, which was tested with the use of the RSFCL superconducting current limiter [13].

Table 3. Parameters of medium voltage electrical network [15]

.

In the construction of the R-SFCL 10 kV/1.8 kA superconducting fault current limiter, the YBCO coil design was used in accordance with the ENSYSTROB design [11], [12]. The limiter consists of 25 coils connected in a series. Each coil consists of YBCO (Super Power SF12100) superconducting tapes and a stainless steel shunt resistor switched on in parallel. The superconducting coil is made of six tapes arranged in an anti-parallel arrangement. By designing the coils from six parallel tapes, the critical current IC R-SFCL reaches 1.8 kA (Ic for each tape is approximately 300 A).

Table 4. Performance of superconducting R-SFCL[13]

.

The SC winding of the limiter is cooled in liquid nitrogen at the temperature of 77 K. The main parameters of the R-SFCL subjected to simulation tests are presented in Table 4 [13].

Superconducting limiter for electrically powered aircraft

A very interesting design solution is the SFCL used in electrically powered aircraft [14]. The power of the system consists of three main units, including the aircraft’s electric propulsion system (EPAS), the main DC bus, and other receivers. The electric system consists of turbines, HTS generators, HTS cables and a superconducting fault current limiter [14], [15].

Fig.8. Schematic of electric propulsion in an electric aircraft using superconducting devices [14]

Together with the circuit break, the SFCL protects the system power against possible high fault current to maintain a high standard of safety in electric aircraft [14], [16].

All the bifilar coil designs to date have been dominated by the BCP design, where the inner and outer windings are connected in parallel. The result is a high current loss and a greater heat load for the cryogenic cooling system. In order to reduce losses and improve the ability to limit fault currents, a helical bifilar coil has been proposed, in which the inner and outer windings are connected in a series so that they conduct the same currents (reduction of the voltage difference between the turns near the terminals compared to the BCP type), and it will also guarantee an equal current distribution in the windings [17].

Fig.9. a) Scheme of bifilar coil with parallel connected windings BCP, b) Scheme of bifilar coil with series connected windings BCS [17]

Summary

The progress in this area compared to previous years is increasing and promising. Currently, work on limiters is carried out in many institutes and research centres. Due to its technical parameters, simple structure and efficiency of limiting the fault current, the main direction of development of the superconducting fault current limiter is the resistive type. Resistive SFCLs are used in electric aircraft due to advantages such as: compactness, lightness and high reliability [1], [16]. Resistive SFCLs are configured as bifilar type to obtain non-inductive or low-inductance structures. This is very beneficial for power systems by avoiding poor voltage regulation and degraded power quality.

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Authors: dr inż. Joanna Kozieł, Department of Electrical Engineering and Electrotechnologies, Faculty of Electrical Engineering and Computer Science, Nadbystrzycka 38A, 20-618 Lublin, e-mail: j.koziel@pollub.pl,mgr inż. Damian Gnat, Department of Electrical Engineering and Electrotechnologies, Faculty of Electrical Engineering and Computer Science, Nadbystrzycka 38A, 20-618 Lublin, e-mail: damian.gnat@pollub.edu.pl, dr hab. inż. Michał Majka, LUT Professor, Department of Electrical Engineering and Electrotechnologies, Faculty of Electrical Engineering and Computer Science, Nadbystrzycka 38A, 20-618 Lublin, e-mail: m.majka@pollub.pl mgr inż. Yerbol Turgynebekov, M. Kh. Dulaty Taraz State University, Kazakhstan, e-mail: bosik90@mail.ru

Source & Publisher Item Identifier: PRZEGLĄD ELEKTROTECHNICZNY, ISSN 0033-2097, R. 98 NR 2/2022. doi:10.15199/48.2022.02.36

Performance Comparison of Harmonic Filters in an Industrial Power System for Harmonic Distortion Reduction

Published by 1. Estifanos Dagnew Mitiku1, 2. Gebrie Teshome Aduye2, 3. S. Abdul Rahman3, 4. Mahilet Mentesinot Abuhay4, Electrical & Computer Engineering Department, Institute of Technology, University of Gondar, Ethiopia (1, 2, 3), Industrial Engineering Department, Institute of Technology, University of Gondar, Ethiopia (4) ORCID. 1. 0000-0002-6613-3979, 2. 0000-0003-1184-9184, 3. 0000-0003-0134-8028, 4. 0000-0001-7554-8562

Abstract: The paper presents performance comparison of harmonic filters for the reduction of harmonic distortion in an industrial power system, i.e., National Tobacco Enterprise, an industry found in Ethiopia. The performance comparison is done for single tuned, double tuned, high pass and Ctype harmonic filter. As a result, the double tuned harmonic filter provides a better performance than the other type of filters by giving a reduced Total Harmonic Distortion for the voltage and current waveforms. MATLAB/SIMULINK simulation results are presented for validating the analysis.

Streszczenie. W artykule przedstawiono porównanie wydajności filtrów harmonicznych do redukcji zniekształceń harmonicznych w przemysłowym systemie elektroenergetycznym, tj. National Tobacco Enterprise, przemysł znajdujący się w Etiopii. Porównanie wydajności jest dokonywane dla pojedynczego strojonego, podwójnie strojonego, górnoprzepustowego i typu C filtra harmonicznego. W rezultacie podwójnie dostrojony filtr harmonicznych zapewnia lepszą wydajność niż inne typy filtrów, zapewniając zmniejszone całkowite zniekształcenie harmoniczne dla przebiegów napięcia i prądu. Przedstawiono wyniki symulacji MATLAB/SIMULINK w celu weryfikacji analizy. (Porównanie wydajności filtrów harmonicznych w przemysłowym systemie zasilania w celu redukcji zniekształceń harmonicznych)

Keywords: Harmonic Filters, Harmonic Distortion, Total Harmonic Distortion, MATLAB/SIMULINK.
Słowa kluczowe: fitr, harmoniczne, THD

Introduction

Power quality (PQ) is an issue to both utilities and electricity consumers at all levels of usage. One of the PQ problems is harmonics which is a sinusoidal component of a periodic wave or quantity having a frequency that is an integral multiple of the fundamental frequency, i.e., 50Hz or 60Hz [1]. Harmonic voltages occur as a result of current harmonics, which are created and drawn by non linear electronic loads, injected to the supply system [2]. National Tobacco Enterprise (NTE) is one of the largest industries which are located in Ethiopia, supplied through a 15kV feeder emanating from the nearby substation. It has two, 630KVA transformers that supply power to linear and nonlinear loads. The non-linear loads cause harmonic distortion resulted in loss of data, overheating or damage to sensitive equipment and overloading of capacitor banks. As a result, single tuned, double tuned, high pass and c-type harmonic filters are employed independently for the reduction of harmonic filters and their performance is compared. The paper in [3], present how to suppress distortions by using the non-linearity current index to determine the shunt single-tuned passive filter compensator value. The paper in [4], describes design of passive filter to reduce the harmonics emitted by power electronic devices in a hybrid, micro-grid network with nonlinear load, energy storage, wind turbine and solar cell. The paper in [5], presents selection of tuning and quality factor to design a single-tuned passive harmonic filter. The paper in [6], describes the effect of single tuned harmonic filter on grid connected PV system and its impact on harmonics and power factor. The paper in [7], aims to design a single tuned filter and testing it in simple model using ETA P to analyze harmonics variation of the harmonic filter in a power system. The study in [8], discusses the use of single tuned passive filters in reducing harmonics in the plastics processing industry. The focus of the paper in [9], was to study the performance of single tuned passive harmonic filter by manipulating the Q-factor in the design. Various types of damped double tuned filters are designed for different nonlinear loads for harmonic reduction has been discussed in [10]. A new method of designing double-tuned filter is proposed based on resonance frequency, by using the relationship that the impedance of double-tuned filter and two parallel single tuned filters is equal and the resonance frequency of single tuned filter is the zero of the impedance of double-tuned filter is presented in [11]. The paper in [12], presents a double tuned passive filter was designed with the parameters of two parallel single tuned passive filters and the results are simulated using MATLAB/SIMULINK software. The paper in [13], presents PQ improvement based on high pass filter for a nonlinear RL-load connected to a single phase ac supply which can simultaneously improve the PQ and control the reactive power requirement of the load. The study in [14], uses passive harmonic filter to mitigate the harmonic voltage distortion in a power system that contained the roughing mill (RM) and finishing mill (FM) electric motor drives. The paper in [15], presents the design of two passive filters to reduce the current harmonics produced by nonlinear loads in industrial power system using MATLAB /Simulink software. The research work in [16], consists of harmonic simulation based on optimal design of C-type filter using Proteus software and hardware implementation with C-type filter and second order high-pass filter. The paper in [17], presents the new C-type high harmonic power filter design process to reduce the total harmonic distortion in an industrial power supply system. The article in [18], presents a method for selecting the elements of a C-type filter working with a conventional LC-type filter for compensating reactive power and filtering out higher harmonics generated by arc furnaces and ladle furnaces. The performance of a single tuned filter and C-type filter is compared in [19], to reduce harmonic content in an electric power distribution network system in the cement industry. To mitigate the other types of PQ problems different techniques have been applied [20]–[27].

From the above literature survey it could be observed that single tuned, double tuned, high pass and c-type harmonic filters were applied independently for the mitigation of harmonic distortion. In this paper, performance comparison of single tuned, double tuned, high pass and c-type harmonic filters in industrial power systems for harmonic distortion reduction is analysed and the THD value of each of the harmonic filters is used for comparison.

Harmonic distortion measurement at the industry

To assess the level of harmonic distortion in the industry power system, monitoring has to be performed at the service entrance points of the industry at the Points of Common Coupling (PCC), the tapping point on the 15kV feeders. However, as the distance and line impedance, from PCC to the primary of the transformer is negligible, the primaries are taken as PCC as shown in figure 1 [2].

Fig.1. Monitoring location and point of common coupling.

The harmonic distortion in the industry is measured while all the machines are working at the same time, to observe the cumulative characteristics of the industrial loads and data collection is accomplished through direct measurement. Having made suitable analysis, the collected data are computed and compared with acceptable values set by standards of IEEE recommended practice for harmonic control [28]. Table 1: Current distortion limits for general distribution systems (120V to 69 000V) [2].Fig. 1: Monitoring location and point of common coupling. The harmonic distortion in the industry is measured while all the machines are working at the same time, to observe the cumulative characteristics of the industrial loads and data collection is accomplished through direct measurement. Having made suitable analysis, the collected data are computed and compared with acceptable values set by standards of IEEE recommended practice for harmonic control [28].

Table 1. Current distortion limits for general distribution systems (120V to 69 000V) [2].

.

Short circuit current and rated current of 15kV feeder at the PCC are averaged to be 10kA and 1000A respectively, which gives ISC/IL ratio in the range of <20. Then, the TDD values of the current harmonics should not exceed 5% at the PCC. The requirement of the utility to provide good quality of voltage is listed in table 2.

Table 2. Voltage distortion limits [2].

.

The maximum voltage and current harmonic contents of the industry when it works at full load are shown in table 3.

Table 3. Maximum voltage and current harmonics level at NTE

.

It is observed from the recorded data that the dominant harmonic currents are the 5th and 7th harmonics and the total current THD value for the three phases is 18.10%, 18.33% and 18.44%, respectively at the PCC which is beyond the IEEE standards, i.e., 5%. Since the current THD values are beyond the standard values, single tuned, double tuned, high pass and c-type harmonic filters with their designed parameters are applied to reduce the harmonic distortion and their performance is compared based on the current THD percentage values when each of the harmonic filters are applied.

Principle of operation

The circuit used for the reduction of harmonic distortion is shown in figure 2. It consists of switches, single tuned, double tuned, high pass and c-type harmonic filters and the load. Due to the non-linear loads of the industrial power system, as harmonic currents are injected to the system the voltage and current waveforms are distorted. Harmonic filters with their designed values and respective switches are applied independently to compare the performance of the harmonic filters and they are switched ON and switched OFF to reduce the harmonic distortion. At first, the single tuned filter switch is ON while the other switches are OFF. Then, the high pass filter switch is ON while the other switches are OFF. Next, the double tuned filter switch is ON while the other switches are OFF. Finally, the C-type filter switch is ON while the other switches are OFF. For simplicity, only one phase is considered out of the three phases as power is taken from the same phase to mitigate the harmonic distortion to show the control circuit.

Fig.2. Control Circuit to show harmonic filters application

Simulation Results

The harmonic distortions have been simulated using MATLAB/SIMULINK software. The harmonic filters such as single tuned, double tuned, high pass and c-type filters are simulated with the designed values in order to compare their performance in the reduction of harmonic distortion occurred in the industrial power system due to non linear loads. The voltage and current waveforms both from the source side and the load side before and after applying the harmonic filters are presented below for comparison. The industry has non-linear loads which are supplied by the utility from the nearby substation and non-linear loads draw harmonic currents of a distorted waveform. The simulation is performed with the designed values of the harmonic filters with a nominal line to line voltage of 400V, operating frequency of 50Hz, nominal reactive power of 50KVAR, quality factor of 7 and the tuning frequency is set for the 5th and 7th harmonics.

The voltage and current waveforms before applying the harmonic filters is shown in figure 3.

Fig.3. Voltage and Current Waveforms without harmonic filters

Fig.4. FFT of the harmonic current before filtering

The industry has non-linear loads which are supplied by the utility and these loads draw a distorted harmonic voltage and current waveform as shown in figure 3 and from the Fast Fourier Transform (FFT) analysis as shown in figure 4 the THD value is 18.33%, which is above the IEEE standard acceptable limit, i.e. 5% for this study. Therefore, to reduce the harmonic distortion, single tuned, double tuned, high pass and c-type harmonic filters are designed and placed at the appropriate location at the industrial power system; consequently, the harmonic distortion is reduced and their performance is compared.

Single Tuned Harmonic Filter
Fig.5. Voltage and Current Waveforms with single tuned filters

Fig.6. FFT of the harmonic current with single tuned filter

High Pass Harmonic Filter
Fig.7. Voltage and Current Waveforms of high pass filter

Fig.8. FFT of the harmonic current using high pass filter

Double Tuned Harmonic Filter
Fig.9. Voltage and Current Waveforms of double tuned filter

Fig.10. FFT of the harmonic current using double tuned filter

C-Type Harmonic Filter
Fig.11. Voltage and Current Waveforms with c-type filters

Fig.12. FFT of the harmonic current using c-type filter

Conclusion

In this paper, performance comparison of harmonic filters in an industrial power system for harmonic distortion reduction is studied. The harmonic distortion drawn by the industry non-linear loads provides a THD value of 18.33%. And when single tuned harmonic filter is applied the THD is reduced to 5.08%. Whereas, the application of high pass harmonic filter gives a THD of 5.09%. A double tuned harmonic filter provides 4.00% THD value. Finally, using c-type harmonic filters 4.68% THD value is obtained. As a result, from the simulation results and FFT analysis, it is observed that the double tuned harmonic filter gives a better performance than the single tuned, high pass and c-type harmonic filters in reducing the harmonic distortions to the acceptable magnitude set by IEEE standards, i.e. less than 5% for this study and it is recommended for the industry to use the double tuned harmonic filter at the secondary of the transformer to get rid of additional heating, false tripping and equipment malfunction due to harmonics which causes production loss to the industry.

REFERENCES

[1] IEEE, IEEE 1159 1995 Recommended practice for monitoriong electric power quality. 1995.
[2] W. Are et al., ‘Electrical Power Systems Quality , Second Edition’.
[3] M. S. Almutairi and S. Hadjiloucas, ‘Harmonics Mitigation Based on the Minimization of Non-Linearity Current in a Power System’, 2019.
[4] A. Bagheri and M. Alizadeh, ‘Designing a Passive Filter for Reducing Harmonic Distortion in the Hybrid Micro-grid Including Wind Turbine , Solar Cell and Nonlinear Load’, no.12, pp. 10–13, 2019.
[5] Y. Cho, H. Cha, Y. Cho, and H. Cha, ‘Single-tuned Passive Harmonic Filter Design Considering Variances of Tuning and Quality Factor Single-tuned Passive Harmonic Filter Design Considering Variances of Tuning and Quality Factor’, vol. 8972, 2014.
[6] R. Dua and A. Agrawal, ‘Impact Of Single Tuned Filter on Grid Connected PV System’, vol. 08, no. 06, pp. 213–217, 2019.
[7] M. Awadalla, M. Orner, and A. Mohamed, ‘Single-tuned filter design for harmonic mitigation and optimization with capacitor banks Single-tuned Filter Design for Harmonic Mitigation and Optimization with Capacitor Banks’, no. September 2015, 2019.
[8] I. O. P. C. Series and M. Science, ‘Harmonic reduction by using single-tuned passive filter in plastic processing industry Harmonic reduction by using single-tuned passive filter in plastic processing industry’, 2018.
[9] Y. K. Haur, T. J. Son, L. K. Yun, W. Jee, and K. Raymond, ‘Design of Single-Tuned Passive Harmonic Filter to Meet Ieee-519 Standard By Means of Quality-Factor Manipulations’, vol. 29, no. 1, pp. 1364–1379, 2020.
[10] K. R. Cheepati, S. Ali, and A. Rangampet, ‘Overview of Double Tuned Harmonic Filters in Improving Power Quality under Non Linear Load Conditions’, vol. 10, no. 7, pp. 11–26, 2017.
[11] H. E. Yi-hong and S. U. Heng, ‘A New Method of Designing Double-tuned Filter’, no. Iccsee, pp. 206–209, 2013.
[12] K. R. Cheepati, S. Ali, and S. K. M, ‘Performance Analysis of Double Tuned Passive Filter for Power Quality’, vol. 9, no. 7, pp. 3295–3305, 2016.
[13] P. Control and G. Mishra, ‘Design of Passive High Pass Filter for Hybrid Active Power Filter Applications Department of Electrical Engineering National Institute of Technology Design of Passive High Pass Filter for Hybrid Active Power Filter Applications’.
[14] B. Park, J. Lee, H. Yoo, and G. Jang, ‘Harmonic Mitigation Using Passive Harmonic Filters : Case Study in a Steel Mill Power System’, 2021.
[15] Z. A. Memon, M. A. Uqaili, and M. A. Unar, ‘Harmonics Mitigation of Industrial Power System Using Passive Filters’, no. July, 2016.
[16] I. A. Shah and R. K. Ali, ‘Design of a C-type Passive Filter for Reducing Harmonic Distortion and Reactive Power Compensation’, vol. 4, no. 12, pp. 38–47, 2016.
[17] R. Klempka, ‘A New Method for the C-Type Passive Filter Design’, no. 7, pp. 277–281, 2012.
[18] A. Furnaces, ‘Selection of C-Type Filters for Reactive Power Compensation and Filtration of Higher Harmonics’, 2020.
[19] A. Muchtar and W. M. Muttaqin, ‘Comparison between single tuned filter and c-type filter performance on the electric power distribution network Comparison between single tuned filter and c-type filter performance on the electric power distribution network’, 2019.
[20] S. A. Rahman, ‘Direct Converter Based DVR to Mitigate Single Phase Outage’, no. September 2019, 2021.
[21] A. Rahman, ‘Mitigation of Voltage Sag , Swell and Outage without Converter’, no. October 2019, 2021.
[22] S. A. Rahman, S. B. Mule, E. D. Mitiku, G. T. Aduye, and C. Gopinath, ‘Highest Voltage Sag and Swell Compensation using Single Phase Matrix Converter with Four Controlled Switches’, no. 4, pp. 134–138, 2021.
[23] C. Reads, ‘Realization of Single Phase Matrix Converter Using 4 Controlled Switches’, no. October 2019, 2021.
[24] S. A. Rahman and G. Teshome, ‘Maximum voltage sag compensation using direct converter by modulating the carrier signal’, vol. 10, no. 4, pp. 3936–3941, 2020.
[25] S. A. Rahman and E. Dagnew, ‘Voltage sag compensation using direct converter based DVR by modulating the error signal’, vol. 19, no. 2, pp. 608–616, 2020.
[26] S. A. Rahman, E. D. Mitiku, S. B. Mule, G. T. Aduye, M. A. Huluka, and S. Mesfin, ‘Voltage Sag Mitigation Using Direct Converter Based DVR without Error Signal’, no. 12, pp. 34–37, 2021.
[27] S. A. Rahman, S. Birhan, E. D. Mitiku, G. T. Aduye, and P. Somasundaram, ‘A Novel DVR Topology to Compensate Voltage Swell, Sag, and Single-Phase Outage’, vol. 17, no. 4, pp. 1–10, 2021.
[28] C. R. C. P. Llc, Power quality © 2002. 2002.

Authors: Lecturer, Mr. Estifanos Dagnew Mitiku, Department of Electrical & Computer Engineering, Institute of Technology, University of Gondar, Gondar, Ethiopia, Email: est7eced@gmail.com; Lecturer, Mr. Gebrie Teshome Aduye, Department of Electrical & Computer Engineering, Institute of Technology, University of Gondar, Gondar, Ethiopia, Email: gebrie.415@gmail.com; Associate Professor, Dr. Abdul Rahman, Department of Electrical & Computer Engineering, Institute of Technology, University of Gondar, Gondar, Ethiopia, Email: msajce.abdulrahman@gmail.com; Lecturer, Mrs. Mahilet Mentesinot Abuhay, Department of Industrial Engineering, Institute of Technology, University of Gondar, Ethiopia, Email: mahiletme@gmail.com.

Source & Publisher Item Identifier: PRZEGLĄD ELEKTROTECHNICZNY, ISSN 0033-2097, R. 98 NR 3/2022. doi:10.15199/48.2022.03.12

Test Stand for Testing and Diagnostics of Medium Voltage Vacuum Interrupters

Published by Paweł WĘGIEREK, Michał LECH, Lublin University of Technology, Faculty of Electrical Engineering and Computer Science

Abstract. The article presents the detailed construction and capabilities of a research station for the diagnosis and testing of vacuum interrupters used in medium voltage electrical switching devices. The correct functioning of the stand has been confirmed by conducting a number of tests on the electrical strength of the MV switchgear vacuum interrupter type HVKR 24/400.

Streszczenie. W artykule przedstawiono szczegółową budowę oraz możliwości stanowiska badawczego służącego do diagnostyki oraz badań komór próżniowych wykorzystywanych w elektroenergetycznej aparaturze łączeniowej średniego napięcia. Poprawność funkcjonowania stanowiska potwierdzono przeprowadzając szereg badań wytrzymałości elektrycznej próżniowej komory rozłącznikowej SN typu HVKR 24/400. (Stanowisko badawcze przeznaczone do badań i diagnostyki komór próżniowych średniego napięcia).

Keywords: vacuum interrupters, vacuum diagnostics, dielectric strength, vacuum switchgears
Słowa kluczowe: komory próżniowe, diagnostyka próżni, wytrzymałość elektryczna, próżniowa aparatura łączeniowa

Introduction

Power engineering is a field of economy developing at a surprisingly fast pace. Many factors influence this process. One of them is the constantly growing number of new electricity consumers, and thus increasing the power required in areas that have not been urbanized before. According to the Transmission System Development Plan, by 2030 [1] the total net electricity demand in 2040 will be 204.2 TWh with 159.9 TWh in 2020.

Another factor in the development of the Polish power industry is its current technical condition (Fig. 1). Outdated elements of the power infrastructure force the investments of power companies related to various types of modernizations [2, 3]. A particular problem is visible in the area of high and medium voltage overhead lines, over 75% of which were built more than 25 years ago [4]. MV power lines are one of the most important elements of the distribution system for both technical and economic reasons [5].

Fig.1. Age structure of selected elements of the polish power system [4]

Another factor enforcing the dynamic development of the power industry is a number of legal requirements imposing on power companies to improve their power supply conditions and to move away from equipment using environmentally harmful greenhouse gases [6-11].

The above factors show that there is a strong need to develop new solutions among power equipment, mainly medium voltage. Many companies and scientific entities have faced this problem and developed pro-ecological equipment, using vacuum as an insulating medium.

One of such devices is the innovative EKTOS vacuum switch disconnector, which is the final result of the project entitled: “Development and implementation of an innovative closed cased overhead vacuum switch disconnector dedicated to intelligent medium voltage networks”, carried out by the Lublin University of Technology in consortium with EKTO Sp. z o.o. from Białystok as part of the activities of the National Research and Development Centre. This device is addressed to Distribution System Operators (DSOs) who want to improve the reliability of electricity supply in the areas they manage through investments in new technologies [12].

Methods for diagnosing vacuum conditions

Equipment such as contactors, switches and disconnectors with installed vacuum interrupters are commonly used in the power industry. The ever-increasing number of such devices is directly related to the need for an effective diagnosis of the state of vacuum inside them. A number of existing diagnostic methods are used for this purpose to assess the proper functioning of medium voltage vacuum extinguishing interrupters in terms of dielectric strength. Examples of vacuum interrupters used in switchgear are shown in Figure 2.

Fig.2. Examples of vacuum interrupters used in switchgear [13]

The basic ones include: the Pening method and the Magnetron method [14, 15]. They consist of placing the tested chamber in a magnetic (strongly axial) field, and then applying a DC voltage of 10 ÷ 20kV to open contacts. According to the gas breakdown mechanism, due to the electric field created due to the applied voltage, electrons are emitted from the cathode moving towards the anode. Placing the chamber under test in the magnetic field changes the path of electrons motion into a spiral one, thanks to which the number of collisions with atoms and molecules of residual gas increases. The mentioned methods consist in recording the current of electron emission, which results from the collision ionization occurring in the chamber. In order to assess the vacuum condition, the characteristics determined for the new vacuum chamber must be known for further comparison.

Another method of diagnosing the state of vacuum is the static AC ignition voltage method [14, 16]. It consists in measuring the value of the jump voltage and then comparing it with the Pashen curve for a given chamber.

The static DC ignition voltage method, which is relatively simple, can also be used to check the correct operation of vacuum chambers. It consists in applying a certain voltage value to the chamber in the open state and then measuring the current value in the chamber and comparing it with the maximum allowable value [14, 17].

In high-frequency test systems, a frequently used method of diagnosing the vacuum condition is the method of AC current switching capability [14, 18], consisting in determining the ability to switch off the AC current, which clearly decreases at certain pressure values.

The Fowler-Nordheim dependence is often used for vacuum chamber tests [14, 19]. This method was called the emission current test method [14, 20]. The use of this method requires appropriate testing equipment, allowing the application of high DC voltage to the chamber, as well as measuring equipment enabling the measurement of currents at the microampere level.

A similar method to the one described above is the test method for emission currents with HF current surges [14, 17]. It consists in forcing a high current value of the frequency exceeding 1 kHz to flow through the chamber, which smoothes the contact surfaces of the chamber being diagnosed.

An interesting method of diagnosing the vacuum condition is the measurement of X-ray radiation [14, 18, 20]. The analysis uses the fact of proportionality of its intensity to the emission current. This method is characterized by a significant defect, consisting of interference from background radiation, which is greater than the radiation of the chamber under operating conditions, so that the results can be significantly disturbed.

Another method consists in measuring the arc voltage at direct current of 10A (DC arc voltage method) [14, 20]. This voltage increases its value while extinguishing the electric arc, while it decreases its value while developing new cathode spots associated with, among others with residual gas in the chamber. The higher the value of the peak voltage, the lower the pressure in the tested chamber.

The method based on measuring the value of the voltage initiating the micro-discharge and the voltage initiating the emission current, called the Vd/Ve method, uses the dependence about the inverse of these voltages in relation to the pressure inside the chamber [14, 21].

For vacuum chambers with external access to their screen, a method of measuring the screen potential can be used to assess the vacuum. It uses the phenomenon of changing the chamber screen potential under the influence of emitted electrons from the chamber contacts [14, 22].

Another method is the method of switching off low induction current, which consists in applying overvoltage impulses to the chamber’s contacts and using the phenomenon of power surges [14, 22].

The method of switching off capacitive current is implemented by breaking the circuit in which the capacitive current flows in the oscillating system [14, 22]. Then, the value of voltage appearing at the terminals of the tested chamber is used to assess the state of the vacuum.

There is also a diagnostic method of vacuum chambers based on the measurement of partial discharges that appear in the chamber during operation. However, the effectiveness of this method is visible at high pressures, which indicate complete leakage of the chamber [14, 22].

Test stand

This test stand for the diagnosis and testing of vacuum interrupters used in medium-voltage switchgear has been designed and manufactured on the basis of a stable, mobile construction with a special platform used for the foundation of the vacuum pump set (Fig. 2). Inside the stand there is space for mounting the research object – medium voltage vacuum interrupter. The desire for a comprehensive study of the electrical strength of the vacuum interrupter is associated with the need to be able to change the contact distance in the appropriate range and with appropriate accuracy. In this test stand it was realized by mounting an extraction screw with a 1 mm pitch thread. Thanks to this, with the use of an appropriate reference scale, it is possible to set the inter-contact distance with the accuracy of 0.1 mm.

Fig.2. View of the test stand together with the method of test object assembly

Power supply to the test stand is provided by means of YHAKXS 1x120mm2 power cable terminated with an angular connector head enabling quick and convenient connection of power supply to the test stand. The test set consists of three main elements: high voltage transformer, capacitive measuring divider and control panel (Fig. 3).

Fig.3 Test set with control panel

The nominal parameters of the kit are shown in Table 1. The schematic diagram of the complete test stand is shown in Figure 4.

An important element of the test stand is a vacuum set to obtain the appropriate pressure inside the vacuum interrupter to be tested. This is done by a set consisting of a turbomolecular and rotary vacuum pump operating at a capacity of 90 l/s.

Table 1. Rated parameters of the test set

.
Fig.4. Block diagram of a complete test bench for testing and diagnostics of medium voltage vacuum interrupters

Research facility

In order to verify the correct operation of the test stand, a test object was installed in it, which is the HVKR 24/400 vacuum disconnector interrupter (Fig. 5). Interrupters of this type are used in three-pole medium-voltage switch disconnectors operating in overhead power networks. Rated parameters of the interrupter are shown in Table 2.

Fig.5 Test facility: HVKR 24/400 vacuum interrupter

Table 2. Rated parameters of the HVKR 24/400 chamber

.

The above vacuum interrupter consists of two poles: mobile and fixed, with contacts made of a mixture of tungsten and copper at a ratio of 70% tungsten to 30% copper. An inseparable element of the interrupter is an elastic bellows enabling the movement of the moving pole, as well as a condensation screen catching conductive particles which, if deposited on the interrupter casing, would deteriorate its operating parameters.

Verification of the correctness of the position

Using the test stand described in this article, it is possible to diagnose the vacuum condition of the selected switch extinguishing interrupter. It is necessary to know its reference electrical strength characteristics and then to compare it with the obtained test results.

Verification of the correctness of operation of this method, called as a static AC ignition voltage method, was carried out in laboratory conditions for the contact distance in the range of 1 ÷ 5 mm for pressure from 4×10-4 ÷ 1.2×103 Pa. The test results are presented in Figures 6 and 7.

Fig.6 Relationship of breakdown voltage Ud as a function of pressure p inside the vacuum interrupter under test

Fig. 7. The relation of the voltage breakdown Ud as a function of the contact distance d

When analysing the above characteristics, attention should be paid to the pressure range in which the dielectric strength of the inter-contact interval of the vacuum interrupter under test is kept constant (Figure 6). This creates a certain safety zone which guarantees the reliable operation of a given device with respect to the electrical strength of the vacuum interrupter installed in it. This situation occurs below a pressure of . 5×100 Pa. The recorded breakthrough voltages in this zone are listed in Table 3.

As the pressure in the tested interrupter increases, a sharp drop in strength is visible. Figure 7 shows the dependence of the breakthrough voltage of the tested vacuum interrupter on the contact distance for selected interrupter pressure values. For pressures between 8,0×10-4 ÷ 5.3×10-1 Pa, the breakthrough voltage increases with the increase in the inter-contact distance. When the interrupter is further aerated, the characteristics are flattened. From the pressure value equal to 6.7×100 Pa, the contact distance did not influence the dielectric strength of the electrical interruption. A vacuum interrupter to be diagnosed, for which the measured value of dielectric strength would be within this range, would be diagnosed as defective.

Table 3. Values of breakthrough voltages recorded in the safety zone of the vacuum interrupter under test

.
Summary

The test stand presented in this article provides an opportunity to diagnose standard vacuum interrupters used in medium-voltage switchgear as well as to test them for improvement of operational parameters.

In order to verify the correct operation of the presented test stand, a number of tests were performed and graphical relationships between the selected parameters were obtained. On the basis of the obtained test results, it can be concluded that the stand described in the article was properly designed and made, and thus it is possible to use the static AC ignition voltage method.

The nearest research works will concern the improvement of electric parameters of vacuum interrupters by increasing the electrical strength of the inter-contact break, as well as limiting the negative effects related to the burning process of the electric arc of the inter-electrode space. The research will be supported by modern computer software enabling professional simulation of physical phenomena taking place in vacuum interrupters used in modern medium voltage electrical apparatus.

This work was supported by The National Centre for Research and Development and co-financed from the European Union funds under the Smart Growth Operational Programme (grant # POIR.04.01.04-00-0130/16).

LITERATURE

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[3] Marzecki J.: Modernization and development directions of low and medium voltage rural network, Przegląd Elektrotechniczny, 2/2019, p. 67-70
[4] Power engineering, distribution and transmission, Polish Power Transmission and Distribution Associaton’s Report, Poznań, 2017
[5] Chojnacki A. Ł.: Comparative analysis of indicators and reliability properties of medium voltage overhead and cable power lines, Przegląd Elektrotechniczny, 11/2019, p. 26-30
[6] Konarski M., Węgierek P.: The use of Power restoration systems for automation of medium voltage distribution grid, Przegląd Elektrotechniczny, 7/2018, p. 167-172
[7] Montreal Protocol on Substances that Deplete the Ozone Layer, Montreal, 1987
[8] Kyoto Protocol to the UN Framework Convention on Climate Change, Kyoto, 1997
[9] Regulation (EU) No 517/2014 of the European Parliament and of the Council of 16 April 2014 on fluorinated greenhouse gases
[10] Quality Regulation 2018 – 2025 for Distribution System Operators
[11] Ordinance of the Minister of Economy of 4 May 2007 on detailed conditions of the power system operation
[12] Węgierek P., Staszak S., Pastuszak J.: EKTOS – innovative medium voltage outdoor vacuum disconnector in a closed housing dedicated to the network smart grids, Wiadomości Elektrotechniczne, 11/2019, p. 21-25
[13] http://www.repo.itr.org.pl/energetyka/vc.html, access:18.06.2020r.
[14] Chmielak W.: Review of methods of diagnostics of the vacuum in vacuum circuit breakers, Przegląd Elektrotechniczny, 2/2014, p.213-216
[15] Kuhl W., Schilling W., Schlenk W.: Messung des lnnendruckes in Vakuumschaltróhr, Vakuum-Technik 34. Jahrgang . Heft 2/85 Seite 34 bis 38
[16] Damstra G. C.: Pressure Estimation in Vacuum Circuit Breakers, IEEE ‘Trans. on Dielectrics and Electrical Insulation Vol. 2 No.2, April 1995
[17] Frontzek F.R., Konig D.: Measurement of Emission Currents Immediately After Arc Polishing of Contacts, IEEE Trans. on EI, vol. 28,No. 4, 1993, p. 700-705
[18] Frontzek F.R., Konig D., Methods for internal pressure diagnostic of vacuum circuit breakers, IEEE 18th ISDEIV – Eindhoven-1998, p. 467-472
[19] Kamarol M., Ohtsuka S., Hikita M., Saitou H., Sakaki M.: Determination of Gas Pressure in Vacuum Interrupter Based on Partial Discharge, IEEE Transactions on Dielectrics and Electrical Insulation Vol. 14, No. 3; June 2007, p. 593 – 596
[20] Walczak K., Janiszewski J., Mościcka-Grzesiak H.: Evaluation of internal pressure of vacuum interrupters based on dynamics changes of electron field emission current and X-radiation HV, Eng. Symp. Aug. 1999
[21] Ziyu Z., Shuheng D., Xiuchen J., Naixiang M., Liwen L., Huansheng S., Chongfang L.: Measurement of Internal Pressure of Vacuum Tubes by Micro-discharge and Emission Current XXIII-rd ISDEIV – Bucharest – 2008
[22] Damstra G.C., Merck W.F.H., Bos P.J., Bouwmeester C.E.: Diagnostic Methods for Vacuum State Estimation, IEEE 18th ISDEIV-Eindhoven-1998, p. 443-446

Authors: dr hab. inż. Paweł Węgierek, profesor uczelni, mgr inż. Michał Lech, Politechnika Lubelska, Wydział Elektrotechniki i Informatyki, ul. Nadbystrzycka 38A, 20-618 Lublin, E-mail: p.wegierek@pollub.pl, m.lech@pollub.pl.

Source & Publisher Item Identifier: PRZEGLĄD ELEKTROTECHNICZNY, ISSN 0033-2097, R. 97 NR 2/2021. doi:10.15199/48.2021.02.36

A Statistical Analysis of Wind Speed Probabilistic Distributions for the Wind Power Assessment in Different Regions

Published by 1. Yuly Bay, 2. Nikolay Ruban, 3. Mikhail Andreev, 4. Alexandr Gusev, Tomsk Polytechnic University. ORCID. 1. 0000-0001-9928-408X, 2. 0000-0003-1396-9104, 3. 0000-0002-6420-4374, 4. 0000-0003-0814-2356

Abstract. The penetration of renewable energy sources (RES) into the electricity supply is gaining popularity all over the world, including countries that have large oil and gas reserves, since only the development of alternative energy will help avoid regression and take a green path development, reducing the damage to the environment. According to estimates of the International Energy Agency (IEA), the capacity of RES units built in China in 2016 was 34 GW, and Australia is one of the world leaders in the photovoltaic power plants installation, the share of which in the Australian electricity production exceeds 3%. It should be noted, that the final power generation capacity and stability are stochastic (probabilistic) in nature. Unlike the classical type generator, the output RES characteristics depend on the geographical features of the installation area, the season, and prevailing winds. Risks associated with inaccurate knowledge of the cumulative distribution function (CDF) describing these sources, as well as environmental uncertainties, are the reasons why it is more difficult for distribution network operators (DNO) to take RES into account in the power balance calculations. The wind speed CDF clarification can provide significant assistance in predicting the RES power production.

Streszczenie. Według szacunków Międzynarodowej Agencji Energetycznej (IEA) moc jednostek OZE wybudowanych w Chinach w 2016 roku wyniosła 34 GW, a Australia jest jednym ze światowych liderów w instalacji elektrowni fotowoltaicznych, której udział w australijskiej produkcji energii elektrycznej przekracza 3%. Należy zauważyć, że końcowa moc i stabilność wytwarzania energii ma charakter stochastyczny (probabilistyczny). W przeciwieństwie do generatora typu klasycznego, charakterystyka wyjściowa OZE zależy od cech geograficznych obszaru instalacji, pory roku i dominujących wiatrów. Ryzyko związane z niedokładną znajomością skumulowanej funkcji dystrybucji (CDF) opisującej te źródła, a także niepewności środowiskowe powodują, że operatorom sieci dystrybucyjnych (DNO) trudniej jest uwzględnić OZE w obliczeniach bilansu mocy. Wyjaśnienie prędkości wiatru CDF może zapewnić znaczącą pomoc w przewidywaniu produkcji energii z OZE. (Analiza statystyczna rozkładów probabilistycznych prędkości wiatru do oceny energetyki wiatrowej w różnych regionach)

Keywords: power system, wind speed time series, probability density function, cumulative distribution function.
Słowa kluczowe: energetyka wiatrowa, rozkład statystyczny.

Introduction

The structure and principles of power system management are becoming more and more complicated. Over the past 15 years due to the insufficient capacity of traditional generation sources, in most developed countries, for reasons of ensuring energy, environmental safety, etc., preference is given to RES, which is being actively introduced in China, Europe and the United States, and the total generated capacity is approximately 2195 GW. Due to this, the total RES capacity is expanding, which leads to an increase in power system stochastic processes.

In the classical cases, the electrical power system (EPS) is a «vertically» arranged system, where a number of operating factors and controlled variables are clearly defined and set within a specific way, established by the DNO [1]. However, in cases of renewable generation penetration, especially in large amount, there is a problem of discrepancy between the generated capacity and the electricity demand. Poor predictability associated with the current wind flow strength, which does not coincide in time with the required capacity, leads to mode dispatching problems. The RES, unlike traditional generators, restructuring EPS into a «vertical-horizontal» one [2], adding uncertainties in management that require further research and forecasting.

The ability and accuracy of forecasting is limited by the statistical information quality or methods of its processing. For example, in these works [3], deterministic methods were used to predict power generation, in order to represent RES as classical. In the articles [4], the probability distribution functions were selected for the input and output characteristics by the statistical analysis methods and testing by goodness of fit criteria. There are also studies devoted to the investigation of the power system units probabilistic characteristics, such as the expected value and standard deviation, the calculation of which contributes to the calculation of the optimal RES implementation capacity in order not to loss of steady state and transient stability.

The wind speed probability distribution approximation

The distribution law choice depends on many factors, including the specifics of the problem. To determine the estimated wind speeds of low frequency (dependence on the wind rose chart [5]), the maximum wind speeds possible in a particular area [6]), the main requirement is a reliable coincidence of empirical and theoretical distributions in the high-value range. The approximation itself as applied to the wind speed distribution was initially widely used for statistical extrapolation of the maximum wind speeds [7]. Subsequently, the approximation of the wind speed distribution by the Weibull and Weibull-Goodrich laws has become one of the most widely used [8]. Along with this law, the normal distribution law is often used, but a large sample size is required to reliably estimate the distribution parameters.

There are papers [9] that claim that the probability distribution is also well described by the lognormal distribution. The laws that can be used for modelling the wind speed, as well as their parameters, are given in the Table 1.

Table 1. Expressions of statistical distributions

.

where k – shape parameter, c – scale parameter, Г – gamma function, α,β,η – parameters of distributions, – normal distribution

The form of the distribution law also depends on the set of observations. In such situations, the distributions of the criteria statistics are often unknown, which is a frequent source of incorrect conclusions.

For optimal research, it is necessary to use several methods to determine the possible distribution law, even before using the goodness of fit criteria. Several well-known methods have been used to determine the various distributions parameters, out of which the method of moments, the graphical method, and the maximum likelihood method. In the case of using the graphical method, it has the advantage of simplicity, however, the accuracy of the input parameters estimating can be insufficient [10]. The likelihood method, on the contrary, has good accuracy, but to achieve it, it is required to use iterative methods [11]. The method of moments equates a certain number of statistical moments of the sample with the corresponding population moments [12]. The use of these methods (at least the maximum likelihood method and method of moments) usually implies that there is an assumption of the possible probability laws that are available in the wind time series. However, in the case of considering the unexplored wind time series, it is more logical to use the graphical or brute force method [13], with subsequent evaluation by several goodness of fit criteria.

The goodness of fit tests

The suitability of the chosen theoretical distribution for describing the empirical probability of a given meteorological argument is verified using the goodness of fit criteria. In this article, we will use Pearson’s chi-squared test [14] and Kolmogorov-Smirnov Goodness-of-Fit Test [15, 16], since the first of them is very sensitive to the dissimilarity of the values edges, the second allows us to more accurately assess the differences in the central regions.

Applying both criteria (with a given 5% significance level), the selected theoretical distribution function can be safely used for indirect calculations. For the measure of the difference between the theoretical and empirical distributions, Pearson takes the value X2 determined by the formula:

.

where n – the sample size, mi – the relative frequencies of the empirical distribution, pi – the corresponding theoretical probability densities, k – the gradations number.

Kolmogorov proposed another goodness of fit criteria, which, in contrast to the Pearson criterion, is based on a comparison of experimental and theoretical distributions integral laws.

As a measure of difference, A. N. Kolmogorov-Smirnov test uses the value:

.

where n – the sample size, D – corresponds to the upper bound (the largest value of the difference between the considered and the original sample) |F*(xi) – F(xi)| = δ(xi).

Input wind time series data

For the experiments, three samples of wind time series data with unknown CDF were taken. The sample size is between 9000 and 200000 volumes, depending on the example. The first sample (Fig. 1a) was taken from one of the graphical method experiments to study Weibul’s law parameters, and was randomly generated. The second time series is taken from the small-scale wind turbine power curve study (Fig. 1b) [17]. The third sample (Fig. 1c) is taken from the wind hourly NUTS 2 time series array [18].

Fig.1. Wind time series data

Fig.2. Extracted wind data CDFs

Based on the information provided, preliminary conclusions can be made about the wind values repeatability, maximum observed and average (mean) values. It should be noted, that in this case, all samples are not tied to particular months, but represent the full input data set for all the time [19]. The parameters that can be obtained before calculating the extracted CDF are shown in Table 2.

Table 2. Wind time series parameters

.

Before the process of finding a fitting CDF and checking it with the goodness of fit criteria, it is necessary to process the input wind data. To do this, we extract the unique values occurring in the wind time series, find the number of occurrences of each unique wind speed value, get the total number of measurements and get the cumulated frequency at the finish (Fig. 2).

A graphical analysis of wind speed CDFs

In order to determine the optimal PDLs, we need to estimate the shape and scale parameter of the curves. Using extracted wind data CDFs, we generate the corresponding PDs. According to the obtained PDs, using the graphical method in conjunction with additional ones, all parameters of possible PDLs are determined, to which the studied wind time series may belong. An example is shown in Fig. 3 for the first data array (a). All parameters of possible distributions are given in Table 3.

Fig.3. A graphical wind data analysis

Fig. 3 shows eight PDFs, namely the Gumbel, Exponential, Gamma, Logonormal, Normal, Rayleigh, Uniform, and Weibull, fitted to the wind speed values. Graphically it can be observed that Logonormal PDF gives the best match. The Gamma, Rayleigh and Weibull distributions match the histogram to a lesser degree, and the remaining distributions provide the worst fits.

Similarly, these eight PDFs were also fitted to other two wind series data and it was observed that the Logonormal, Gamma, Weibull, and Rayleigh the best ones for further analyses.

The most widely used distribution of the selected laws is the Weibull distribution. It is easy to use and accurate for most wind conditions that may occur in research. The Rayleigh distribution is a simplified version of the Weibull distribution, characterized by its simplicity due to the use of only one parameter, which negatively affects the quality of the obtained characteristics, and it is not so often suitable. Gamma and lognormal distributions are also two-parameter, they are less common in wind descriptions, but they can be much better suited for a several wind time series [20] (depending on the wind samples specific values repeatability).

Table 3. Wind time series obtained distribution parameters

.
Fig.4. Obtained wind data CDFs

After that, the wind time series is checked using the Pearson’s chi-squared test and Kolmogorov-Smirnov Goodness-of-Fit test according to the laws selected above. For the first sample data, the Weibull distribution meets the goodness-of-fit criteria (Fig. 4a). The second one corresponds to the Rayleigh distribution (Fig. 4b).

For the third sample, the Gambel distribution and the normal distribution were the closest, but neither of them fully satisfied the Kolmogorov test. This may be due to the small number of distribution laws considered, which were proposed in the article, or to the complexity of the original law (multiparameter, multimodal distribution, etc.).

Thus, we can conclude that the tools for finding the probabilistic characteristics of the wind time series presented in this article are extensive, but not always sufficient for the most accurate description of complex laws. For some cases, it may be necessary to use more sophisticated and advanced methods to obtain reliable probabilistic parameters.

Conclusion

The study of the wind speeds CDF was based on real and accurate measurements of these values at three obviously different sites. The results showed that it was possible to fully determine the probabilistic characteristics corresponding to the goodness-of-fit criteria for two of them. Thus, for some investigated wind time series, it will be necessary to expand the initial list of possible CDFs.

The implemented capabilities for modeling the distribution from random variables allow us to model the CDF and PD for the RES active and reactive power of various configurations based on the specific territory wind models.

Acknowledgment – The work was supported by Ministry of Science and Higher Education of Russian Federation, according to the research project № МК-5320.2021.4.

REFERENCES

[1] Zhang, J., M. Cheng, and X. Cai. (2012). Short-Term Wind Speed Prediction Based on Grey System Theory Model in the Region of China. Przeglad Elektrotechniczny, 88 (7a), 67-71.
[2] Strzelczyk, F. (2009). Renewable energy sources in power system. Przeglad Elektrotechniczny, 85 (9), 340-349.
[3] Karaki, S.H., Chedid, R.B., Ramadan R. (1999). Probabilistic performance assessment of autonomous solar–wind energy conversion systems, IEEE Trans Energy Conversion, 14 (3), 766–772.
[4] Kruangpradit P., Tayati W. (1996). Hybrid renewable energy system development in Thailand, Renewable Energy, 8 (1–4), 514–517.
[5] Sohoni, V., Gupta, Sh., Nema, R. (2016). A comparative analysis of wind speed probability distributions for wind power assessment of four sites. Turkish Journal of Electrical Engineering & Computer Sciences, 24, 4724-4735.
[6] Giraldo, J., Castrillon, J., Granada-Echeverri, M. (2014). Stochastic AC Optimal Power Flow Considering the Probabilistic Behavior of the Wind, Loads and Line Parameters. Ingeniería e Investigación, 15, 529-538
[7] Soroudi, A., Aien M., Ehsan, M. (2012). A Probabilistic
Modeling of Photo Voltaic Modules and Wind Power Generation Impact on Distribution Networks. IEEE Systems Journal, 6 (2), 254-259.
[8] Malska, W. and D. Mazur. (2017). Analysis of the Impact of Wind Speed for Power Generation on the Example of Wind Farm. Przeglad Elektrotechniczny, 93 (4), 54-57.
[9] Akyuz, H., Gamgam, H. (2017). Statistical Analysis of Wind Speed Data with Weibull, Lognormal and Gamma Distributions. Cumhuriyet Science Journal, 38, 68-76.
[10] Ross, R. (1994). Graphical Methods for Plotting and Evaluating Weibull Distributed Data. Proceedings of the 4th Int. Conf. Properties and Applications of Dielectric Materials,1, 250 – 253.
[11] Cousineau, D., Brown, S., Heathcote, A. (2004). Fitting distributions using maximum likelihood: Methods and packages, Behavior Research Methods, Instruments, & Computers,36, 742–756.
[12] Prem, Ch., Siraj, A., Vilas, W. (2018). Study of different parameters estimation methods of Weibull distribution to determine wind power density using ground based Doppler SODAR instrument. Alexandria Engineering Journal, 57 (4), 2299-2311.
[13] Dongbum, K., Kyungnam, K., Jongchul H. (2018). Comparative Study of Different Methods for Estimating Weibull Parameters: A Case Study on Jeju Island, South Korea. Energies, 11 (2), 1-19.
[14] Seyit, A., Akdağ, A., D. (2009). A new method to estimate Weibull parameters for wind energy applications. Energy Conversion and Management, 50 (7), 1761-1766.
[15] Çelik, H., Yilmaz, V. (2008). A Statistical Approach to Estimate the Wind Speed Distribution: The Case of Gelibolu Region. Doğuş Üniversitesi Dergisi, 9 (1), 122-132.
[16] Bielecki, S. (2017). Reactive Power Demand – Verification of a Hypothesis of Normal Distribution Values). Przeglad Elektrotechniczny, 93 (9), 20-23.
[17] Loic, Q., Clement, J., Christian. E. (2014). Measuring the Power Curve of a Small-scale Wind Turbine: A Practical Example. Conference Proceedings Paper – Energies “Whither Energy Conversion? Present Trends, Current Problems and Realistic Future Solutions”, pp. 1-11.
[18] González-Aparicio, I., Monforti, F., Volker, P., Zucker, A., Careri, F., Huld, T., Badger, J. (2017). Simulating European Wind Power Generation Applying Statistical Downscaling to Reanalysis Data. Applied Energy, 199, 155-168.
[19] Rosas, P. A. C., Nielsen, A. H., Bindner, H. W., Sørensen, P. E., Lindahl, S. O. R., Nielsen, J. E. & Pedersen, J. K. (2004). Dynamic Influences of Wind Power on The Power System, Technical University of Denmark, Denmark, Forskningscenter Risoe.
[20] Lingfeng, W., Chanan, S., Andrew, K. (2010). Wind Power Systems: Applications of Computational Intelligence, Springer-Verlag Berlin Heidelberg.

Authors: Assistant of Division for Power and Electrical ngineering, Yuly Bay, Tomsk Polytechnic University, 30, Lenin Avenue, Tomsk, Russia, E-mail: nodius@tpu.ru; Associate professor of Division for Power and Electrical Engineering, Nikolay Ruban, Tomsk Polytechnic University, 30, Lenin Avenue, Tomsk, Russia, E-mail: rubanny@tpu.ru; Associate professor of Division for Power and Electrical Engineering, Mikhail Andreev, Tomsk Polytechnic University, 30, Lenin Avenue, Tomsk, Russia, E-mail: andreevmv@tpu.ru; Professor of Division for Power and Electrical Engineering, Aleksandr Gusev, Tomsk Polytechnic University, 30, Lenin Avenue, Tomsk, Russia, E-mail: gusev_as@tpu.ru.

Source & Publisher Item Identifier: PRZEGLĄD ELEKTROTECHNICZNY, ISSN 0033-2097, R. 97 NR 12/2021. doi:10.15199/48.2021.12.14

Frequency Resolution Improvements in Induction Motor Fault Diagnosis : Experimental Validation

Published by Ameur Fethi AIMER1, Ahmed Hamida BOUDINAR2, Mohamed Amine KHODJA2, Azeddine BENDIABDELLAH2, University of Saida. Algeria (1), University of Sciences and Technology of Oran. Algeria (2) ORCID: 1. 0000-0003-4933-109X

Abstract. Induction machines are enjoying growing interest mainly due to their robustness, their weight to power ratio and their manufacturing cost. However, several faults affect the reliability of these machines. In order to identify these defects, the power spectral density, based on the Periodogram technique is used for its simplicity and its short computing time. However, it is limited in frequency resolution in cases of low motor slip (harmonic close to the fundamental), in the case of very noisy signals (false alarms) and in the detection of incipient faults (low amplitude harmonics) which makes the diagnosis inefficient. To improve the frequency resolution of the spectral analysis, we highlight in this paper the impact of the choice of the weighting windows in order to have a reliable diagnosis of induction motor’s rotor faults. The experimental results will then show the properties of each window to improve the frequency resolution and thus correct the Periodogram’s limits.

Streszczenie. Maszyny indukcyjne cieszą się coraz większym zainteresowaniem głównie ze względu na ich solidność, stosunek masy do mocy oraz koszt wykonania. Jednak kilka usterek wpływa na niezawodność tych maszyn. W celu identyfikacji tych defektów wykorzystuje się gęstość widmową mocy, opartą na technice Periodogram, ze względu na jej prostotę i krótki czas obliczeń. Jest jednak ograniczona w rozdzielczości częstotliwości w przypadkach niskiego poślizgu silnika (harmoniczna zbliżona do podstawowej), w przypadku bardzo zaszumionych sygnałów (fałszywe alarmy) oraz w wykrywaniu początkowych usterek (harmoniczne o niskiej amplitudzie), co sprawia, że diagnoza jest nieskuteczna . Aby poprawić rozdzielczość częstotliwościową analizy spektralnej, w niniejszym artykule zwracamy uwagę na wpływ doboru okien ważenia na wiarygodną diagnozę uszkodzeń wirnika silnika indukcyjnego. Wyniki eksperymentalne pokażą następnie właściwości każdego okna, aby poprawić rozdzielczość częstotliwości, a tym samym skorygować granice Periodogramu. (Poprawa rozdzielczości częstotliwości w diagnostyce usterek silnika indukcyjnego: walidacja eksperymentalna)

Keywords: Induction motor; Fault diagnosis; Broken rotor bars; Frequency resolution.
Słowa kluczowe: silnik indukcyjny, diagnostyka, uszkodzenie prętów

Introduction

Nowadays, induction motor is widely used in most electric drives applications, especially at constant speed. Advances in power electronics associated with modern control techniques have led to the consideration of variable speed applications, which were previously limited exclusively to DC motors and synchronous motors. Thus, faced with this growing interest, a general reflection is naturally directed towards the detection of faults and the monitoring of induction machines state. There are several techniques in fault diagnosis; vibration analysis being the most widely used method [1], [2], [3]. This method is mainly used for the detection of mechanical faults.

Motor current signature analysis or MCSA has been used more and more in recent years. Its peculiarity is that the stator current spectrum carries information on almost all of the electrical and mechanical faults that can affect the induction motor [4], [5], [6]. Spectral analysis based on signal processing has been used in recent years in the diagnosis and monitoring of induction machines faults [7]. This technique is well suited to the fault diagnosis insofar as many phenomena result in the appearance of sideband frequencies directly related to the speed of rotation of the motor.

Based on the calculation of the Fourier transform (FT), the power spectral density (PSD) is a widely used tool in research and industry associated with the analysis of stator current [8]. This is justified by the simplicity and the low cost of the current sensors and the harmonic content of the stator current. However, this technique has several drawbacks linked to the problem of frequency resolution. Indeed, the calculation of FT introduces a smoothing effect as well as a negative effect. These effects result in the appearance of sideband lobes in the stator current spectrum [9] and therefore reduce the clarity of the analysis.

When analyzing a signal, it is interested to have a main lobe as narrow as possible and side lobe amplitudes as low as possible, both advantages are impossible to achieve simultaneously. Because of this resolution problem, the PSD find difficulties in detecting faults when harmonic are near to the fundamental (in the case of a low motor slip), of false alarms (in the case of highly noisy signals) and for harmonics of low amplitude (case of incipient faults detection ).

Within this objective, this paper focuses on the choice criteria through experimental tests of the window weights and the impacts of this choice on the detection and localization of induction motor’s rotor faults.

Stator current analysis

The spectral analysis of the stator current knows a growing interest these last years, because of the quantity of information contained in its spectrum on most of the faults which can appear on an induction machine. It is interesting to note that, as in the case of the vibratory analysis, the spectral components of the fault continue to increase with time by the increase of the fault severity [6]. The broken rotor bars faults of the induction motor are considered among the most commonly studied faults because of their simplicity of implementation. This fault induces changes in the spectral components of the stator current and thus generates the appearance of new sideband frequencies in the current spectrum relating to the broken rotor bars fault [7].

Indeed, broken rotor bars give rise to a sequence of sidebands frequencies given by:

.

where: fs is the supply frequency and fc the sideband frequencies associated with the broken rotor bars fault, s is the motor slip and k = 1, 2, 3…

When analyzing the stator current, it is just possible to evaluate the general condition of the rotor. If there are broken rotor bars in various parts of the rotor, the current analysis is not able to provide information on the configuration of non-contiguous broken bars. For example, the frequency component does not exist if broken bars are electrically π/2 radians away from each other.

It should be noted that some experimental studies have demonstrated that both the skewing and non-insulation of rotor bars lead to a reduction of broken rotor bars harmonic components.

Power Spectral Density Calculation

Fourier Transform

The Fourier transform (FT) is a powerful mathematical tool used to extract useful information from a signal in the frequency domain. It is a nonparametric method, which lends itself well to the analysis of stationary phenomena. The FT is given by the following relation [9-10]:

.

where FTx(f) is called the Fourier Transform of the signal x(t), represented in our case by the stator current signal of the induction motor. Of course, it is impossible to analyze the signal over an infinite period. It is therefore necessary to truncate the signal prior to digital processing.

Truncation operation

The signal to be processed must be limited in time, this is said to be truncated. Mathematically, this amounts to do the following operation:

.

where: x(t): is the measured signal; xT(t): is the signal to be processed; ΠT: is the rectangular window; T: is the time length of the window.

However, this truncation operation introduces negative effects on the signal spectrum. Indeed, these effects also known as side lobes appear during this operation. These side lobes result from the brutal impact of truncation of the signal that comes to replace it by zero outside the support of the rectangular window ΠT. These effects reduce the analysis accuracy.

Weighting windows

To resolve the truncation operation effects, we use the weighting windows ωT(t). This implies that the weighted signal xp(t) is processed instead the truncated signal xT(t).

The new signal is given by :

.

While performing a fault diagnosis operation based on peaks detection, it is more suitable to have a main lobe as narrow as possible and side lobe amplitudes very low to avoid false alarms. Unfortunately, it is almost impossible to have both properties in the same time. Thus, the weighting windows are chosen based on the nature of the processed where: signal and the searched compromise.

Table 1. Weighting windows description

.

Table 1 gives the main weighting windows used with a compact support. Therefore, we consider the main lobe width at -3dB defined by the parameter L for the frequency resolution Δf, and the amplitude of the highest side lobe given by the parameter A. These windows are shown in Fig. 1 [11].

Fig.1. Representation of the weighting windows

Discrete Fourier Transform

To determine the Fourier transform of a signal using a digital computer, the number of frequencies obtained is limited due to the limited computing power of the computer. It is therefore necessary to substitute the continuous variable f by a discrete variable.

The operation dedicated to the frequency discretization is based on the replacement of the continuous frequency f by the discrete frequency kΔf (where k is an integer). The obtained frequencies are known as frequencies components of the DFT (Discrete Fourier Transform). Since the FT of a digital signal should be periodic with Fe period, the frequency resolution is given for N samples by the following equation:

.

where: Δf : Frequency resolution; ; Fe : Sampling frequency; N : Number of samples (with which we calculate the DFT).

The frequency discretization is than defined by a sampling operation in the spectral domain. Numerically, the DFT is expressed by:

.

FFT Algorithm

The Fast Fourier Transform, also known as FFT is an algorithm based on fast calculation of the DFT proposed by J.W. Colley and J.W. Tuckey in 1965. The FFT algorithm uses a number of points NTF equal to a power of 2, which results in a computing time gain compared to a classic calculation using the DFT, this gain in time is given by the following equation [11]:

.

If the number of points obtained after the acquisition step is not a power of 2, the record length of the signal is completed with zeros in order to use the FFT algorithm; this procedure is called as the zero padding procedure or the zeros extension step.

Power spectrum

Finally, we define the power spectral density (PSD) as the square modulus of the Fourier Transform. The PSD is independent of the signal phase. In addition, it is always real and positive; it is given by [12]:

.
Experimental tests

The experimental tests presented in this paper are carried out by the DIAGNOSIS group at the LDEE laboratory at the University of Sciences and Technology of Oran, Algeria. The motor used in these practical tests is a three-phase squirrel cage induction motor coupled to a Direct Current generator used as a load. The parameters of the induction motor are given in the appendix.

In this paper, we deal with the broken rotor bar diagnosis issue; this fault is created artificially in our tests. The measurement chain includes three hall-effect current sensors, an anti-aliasing filter, a tachometer and an acquisition card. Finally, a computer is used to process the acquired signals. this test bench is shown in Fig. 2.

The motor operating modes used to validate the diagnosis procedure are:

– Healthy engine operation.
– Motor operating with 01 broken bar at a motor slip of 4.06%.
– Motor operating with 01 broken bar at a motor slip of 2.13%

Fig.2. Experimental setup description

Interpretation and discussion

Figure 3 shows the estimation of the power spectral density PSD using the periodogram in the case of 1 broken rotor bar. For this purpose, the spectral analysis is carried out using the four weighting windows studied in this paper. According to eq. (1), the rotor bars fault is located on both side of the fundamental at a particular frequency. For this test, the motor slip is 4.06% which gives a sideband frequency signatures around 45.94Hz and 54.06Hz for k=1. This frequency signature is repeated for the values of k=2,3…etc. Indeed, the parameter k represents the multiplicity of the fault frequency signatures on the spectrum.

Fig.3. Stator current PSD with various weighting windows for 1 broken rotor bar and a motor slip of 4.06 %

Fig.4. Stator current PSD with various weighting windows for 1 broken rotor bar and a motor slip of 2.13 %

For the rectangular window, the frequencies are barely detectable. Whereas for the other windows, the detection of these frequencies is easier. It should be noted that the Hanning window is distinguished by a larger main lobe compared to the Hamming and Gaussian windows. On the other hand, this same Hanning window gives the sideband frequencies with the greatest amplitude.

For the last test shown in Fig. 4, the power spectral density PSD per Periodogram of the stator current in the case of 1 broken rotor bar is highlighted. In this test, the motor slip is equal to 2.13% which gives sideband frequencies close to the fundamental. These frequencies calculated using eq. (1) are located around 47.87Hz and 52.13 from either side of the fundamental. For this low value of the motor slip, the sideband frequencies of the broken rotor fault are too close to the fundamental.

Under these conditions, localization using the rectangular window is almost impossible due to the position of the sideband frequencies regarding the fundamental. For the Hamming and Gaussian windows, the localization is difficult and less obvious compared to the Hanning window. Indeed, the Hamming and Gaussian windows offer a narrow main lobe and therefore are best suited for cases of low motor slip.

Finally, the Hanning window is more suitable in the case of incipient faults, given the large amplitude of the side lobes.

Conclusion

This paper investigates the influence of the weighting windows choice on the frequency resolution of the stator current spectrum. In this aim, we present three weighting windows used to resolve the resolution problems due to the rectangular window use. Indeed, a proper choice of the weighting window is necessary to study critical cases that may arise (e.g. case of low motor slip and incipient faults).

To assess each window, we take into consideration the study of fault diagnosis of broken rotor bars and its identification using the power spectral density spectrum. Through the study of each window, we searched a compromise between a narrow main lobe width and side lobes amplitude. This compromise was clearly shown by the experimental results presented in this paper. It has been observed that the Hanning window gave side lobes of low amplitude but the main lobe is wider.

Furthermore, windows Gaussian and Hamming offer the possibility of having a narrow main lobe but the side lobes has more amplitude than that obtained with the Hanning.

Finally, we can say that the Hanning window is recommended for the diagnosis of incipient faults and the Hamming window or Gaussian window is more appropriate in the case of faults too close to the fundamental. The next step will be devoted to the development of an adaptive process composed of several weighting windows. This process will be achieved using Artificial Intelligence.

Appendix. Induction motor parameters

.

REFERENCES

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[11] A.F. Aimer, A.H. Boudinar, M.A. Khodja, A. Bendiabdellah, Assessment of windowing effect on the frequency resolution of the stator current PSD for induction motor broken rotor bars diagnosis, IEEE 1st International Conference on Innovative Research in Applied Science, Engineering and Technology IRASET, Meknes, Marocco 16-19 Apr. 2020
[12] M.B. Koura, A.h.Boudinar, A. Bendiabdellah, A.F. Aimer, Z. Gherabi, Rotor faults diagnosis by adjustable window, Przeglad Elektrotechniczny Journal. March 2021. Vol. 97, Issue 3. pp.123-129

Authors: Dr. Ameur Fethi AIMER, Diagnosis Group-LDEE Laboratory. University of Saida, Algeria. Email: fethi.aimer@yahoo.fr Prof. Ahmed Hamida BOUDINAR, Diagnosis Group-LDEE Laboratory. USTO-Oran, Algeria. Dr. Mohamed Amine KHODJA, Diagnosis Group-LDEE Laboratory. USTO-Oran, Algeria. Prof. Azeddine BENDIABDELLAH, Diagnosis Group-LDEE Laboratory. USTO-Oran, Algeria.

Source & Publisher Item Identifier: PRZEGLĄD ELEKTROTECHNICZNY, ISSN 0033-2097, R. 97 NR 11/2021. doi:10.15199/48.2021.11.12