Cable Links Designing in HVAC and HVDC Submarine Power Grids – Selected Issues

Published by Mirosław PAROL1, Sylwester ROBAK1, Łukasz ROKICKI1, Jacek WASILEWSKI2,
Warsaw University of Technology, Institute of Electrical Power Engineering (1)
PSE Innowacje Sp. z o.o. (2)

Abstract. Submarine (offshore) HVAC and HVDC cable power grids are becoming more and more important. The establishment of submarine transmission interconnections between various national power systems and the development of offshore wind farms are main reasons for that. This paper presents the concept of offshore cable power grids, taking into account the issues of the injection of a power generated in offshore wind farms to onshore power grids and the intersystem power exchange. Some general power system requirements with regard to the rules, determining how a submarine cable link and an onshore power grid should cooperate, have been defined. The main subject of this paper is the designing process of HVAC and HVDC submarine cable links. Detailed guidelines on how to design both kinds of submarine cable links, also taking into account some major challenges and practical obstacles have been given. Last but not least, some final conclusions have been drawn and included.

Streszczenie. Podmorskie kablowe sieci elektroenergetyczne HVAC i HVDC stają się coraz ważniejsze. Budowa podmorskich połączeń przesyłowych między systemami elektroenergetycznymi różnych państw oraz rozwój morskich farm wiatrowych są głównymi przyczynami tego stanu rzeczy. Artykuł ten prezentuje koncepcję podmorskich kablowych sieci elektroenergetycznych, biorąc pod uwagę zagadnienia wprowadzenia mocy generowanej w morskich farmach wiatrowych do lądowych sieci elektroenergetycznych i międzysystemowej wymiany mocy. Zdefiniowane zostały niektóre ogólne wymagania systemowe odnoszące się do reguł, jak powinny ze sobą współpracować podmorskie łącze kablowe i lądowa sieć elektroenergetyczna. Głównym tematem tego artykułu jest proces projektowania podmorskich łączy kablowych HVAC i HVDC. Zostały podane szczegółowe wytyczne, jak projektować obydwa rodzaje podmorskich łączy kablowych, biorąc także pod uwagę niektóre główne wyzwania i praktyczne przeszkody. W końcu zostały wyciągnięte i zawarte wnioski końcowe. (Projektowanie łączy kablowych w podmorskich sieciach elektroenergetycznych HVAC i HVDC – zagadnienia wybrane).

Keywords: HVAC and HVDC power transmission, submarine (offshore) power grids, designing of cable link.
Słowa kluczowe: przesył mocy w technologii HVAC i HVDC, podmorskie sieci elektroenergetyczne, projektowanie łącza kablowego.

Introduction

The subject-matter presented in this paper is very important from the practical point of view (i.e. designing practice). It results from the fact that, importance of submarine (offshore) HVAC and HVDC power grids is expected to grow with time. In particular, it will be a result of offshore wind farms being developed and submarine transmission interconnections between various national power systems being established. As a consequence, proper procedures outlining the design of HVAC and HVDC submarine cable links (lines), composing respectively of HVAC and HVDC submarine power grids, need to be formulated. Every single HVDC submarine cable link is composed of a HVDC submarine cable line and converter substations situated on both sides of the line.

The most advanced review of different issues related to HVAC and HVDC submarine power cables has been presented in [1, 2]. Additionally, a few sessions on this subject have been organized during JICABLE’15 Conference [3]. It is worth paying attention to the fact that in the time of planning, apart from different possible power transmission technology (HVAC, HVDC, HVAC-DC systems), one can distinguish several typical classes of submarine power grids structures.

The most important information concerning issues such as: concept of HVAC and HVDC submarine cable power grids, information about existing and planned submarine cable links, short characteristics of these cables, general power system requirements for these cable links, and designing rules of these power links will be presented in consecutive sections of this paper. Some of them are explicitly associated with the topic of designing HVAC and HVDC submarine power grids. Issues discussed in this publication were among others, a subject of [4].

Concept of submarine cable power grids

Offshore wind power, marine gas & oil industry is planned for worldwide development. Additionally, increase in the integration of power systems of individual seaside countries is expected. This creates need for advanced research among others on defining technical standards for planning the submarine (offshore) cable power grids. A planning process of submarine cable power grids should be based on multi-stage optimization problem solving, i.e. determining the time series of power grid structures within the assumed time horizon. Such a task is a complex decision problem in which more than one criterion and a set of limitations (technical, functional, locational, environmental, reliability, and legal and administrative requirements) should be considered. Also the conditions of risk and uncertainty should be taken into account. Existing power infrastructure have a significant influence on the development of submarine power grids, such as connections between seaside power systems (for energy international exchange purposes) and links between offshore installations and the nearest power systems. One of the important steps of the planning process of submarine power grids is to determine target topology requirements of power networks. Structural classes such as: radial, ring, topology with international coordination and meshed topology can be considered in the planning process [5, 6].

The class of international coordination is an evolution of the star topology, in which the existing international power cable link is introduced into the offshore nodal substation. This solution can be used for a relatively short distance between the nodal substation located in the sea and the interconnection submarine cable link.

Within the meshed topology of submarine power grid structure the nodal offshore substations are built in the first order to aggregate offshore installations. Next, the purpose of the submarine interconnection cable is to import/export power and energy. The submarine cable links can form a multi-nodal loop power network or an open (tree) structure. The cost-effectiveness of using a meshed class strongly depends on the planned power and energy exchange between individual domestic power systems.

Different variations of the aforementioned topology classes of submarine power networks can be considered such as: radial-ring, radial-star, meshed with ring aggregation, etc. Each topology of a submarine power grid can be classified as an open (radial, star) or closed one (ring, meshed). The class of open topology ensures only one power flow path between any two grid nodes (within the meaning of onshore or offshore power substations). At least two power flow paths exist in the closed topology class. Different techniques and strategies of power flow control are associated with the particular grid topologies [7]. It is intuitive that closed power grid structures contribute to higher reliability of energy delivery from/to offshore power grid participants. Considering the type of voltage used for transmission or distribution of electric power: AC or DC, it is essential to refer to the following classes of elements of a submarine power grid: power transmission lines and offshore power distribution/conversion nodes.

A selection of specific type of voltage technology (HVAC or HVDC) for the submarine power grids is primarily dependent both on the final grid topology and planned location of the power substations. The latter can function in the submarine power subsystem as: collection, hub (cluster), system connector, system interconnector or mixed. The distances between the nodes of a submarine power grid determine a set of solutions regarding the type and level of applied voltage for transmission cable lines. Low distances between the grid nodes favor the use of HVAC technology. In turn, higher distances enforce the use of HVDC systems. The installed power of offshore generation/load is a second variable affecting the decision whether HVAC or HVDC technology should be used [8].

The selected type and level of voltage for the offshore transmission lines is an input to the selection process of a type and level of voltage for the distribution/conversion nodes. This determines both the number and type of equipment needed to be installed in the offshore and onshore substations. They are i.a.: switchgears, transformers, devices for reactive power compensation, power electronic converters AC/DC and DC/AC (including harmonic filters).

Other inputs to the selection process of a type and level of voltage (including possible combination of voltage types for power transmission lines and distribution nodes HVDCAC) applied to the planned submarine power grid are assumptions formulated by the transmission system operators as well as the available technology of power flow control. The states of the submarine power system mainly depend on the power load and generation at the offshore energy consumers and wind power plants respectively, as well as the international power exchange. The issues associated with control process of power flow depend strongly on both the choice of AC/DC and DC/AC converter technologies used in the submarine power grids and the grid topology [9].

Overview of existing and planned submarine cable links

In the past, submarine cable links have been used primarily to supply individual objects located at a short distance from the shore. Nowadays submarine power cables can provide electricity for distances exceeding 100 km. Due to the increasing popularity of offshore wind farms, submarine cables can be used as export cables and submarine power lines forming submarine power grids. Currently, many running submarine cable connections operate in radial structure. Examples of such solutions are given in [10]. Basic technical data of some existing submarine cable links have been shown in [4]. Schematic of sampled export cable (London Array wind farm export cable) is shown in Fig. 1.

Fig. 1. Schematic of London Array wind farm export cable; based on [11].

In order to fully exploit the potential of offshore wind energy, many power system operators decide to build a multi – node submarine power grid. The development process is likely to rely on the integration of existing radial connections into larger structures, managed and funded by several AC transmission system operators, or specially appointed, independent submarine grid operator. Examples of such solutions are [12 – 14]:

• Polish submarine power grid,
• Kriegers Flak Combined Grid Solution (KFCGS),
• North Sea Countries Offshore Grid Initiative (NSCOGI).

Offshore power substations Essential are elements of the proposed concept of Polish submarine power grid. They are points of connection of wind farms, as well as cables exporting power to Polish, German, Danish, Swedish and Lithuanian power systems. At the present stage of the project it has not been decided whether the system will be developed as a direct (HVDC) or alternating (HVAC) current grid [12]. Concept of Polish submarine power grid is shown in Fig. 2.

Fig. 2. Concept of Polish submarine power grid; based on [12].

Kriegers Flak Combined Grid Solution anticipates commissioning of three offshore wind farms with total capacity of 938 MW. These offshore wind farms will be combined with the power systems in Germany, Denmark, and Sweden [13]. During the planning of Kriegers Flak network development, concepts of radial connection of each wind farm to the mainland or to the common node on the Baltic Sea were considered. During the research of KFCGS HVAC, HVDC and mixed power grid solutions were analyzed. After completing the analysis the Swedish power grid operator withdrew from the project with a stipulation it will be able to join it at a later time.

Studies on integrated submarine power grids, connecting transmission systems of North Sea countries (NSCOGI) were launched at the end of 2010 [14]. Actions were taken to develop a long – term plan for the development of offshore wind farms in the North Sea and the submarine power grid to allow delivery of power to customers in different countries participating in the project. Due to the relatively large distances between network nodes NSCOGI plans to build necessary infrastructure in HVDC technology. Connection of mainland transmission systems and wind farm generators will be implemented by means of VSC type power electronic converters.

General characteristic of HVAC and HVDC submarine power cables

During the last few dozen years many different types of submarine power cables were designed and manufactured. A very good survey of these cables can be found in [1].

The most important information concerning the following issues related to submarine power cables:

• type of insulation (XLPE, oil-filled paper insulation, paper-mass impregnated insulation (MI), gas-filled insulation);
• current-carrying conductors design (copper, aluminum);
• types and shapes of conductors (solid conductor, stranded round conductor, profile wire conductor, profile wire hollow conductor, segmental (Milliken) conductor, segmental hollow conductor);
• conductors number in three-phase systems (single-core
cables, three-core cables);
• types of cables in DC systems (single-core cables, two-core
(two conductors) cables, coaxial cables);
• water-blocking sheaths;
• armor;
• non-metallic outer sheath;
• optical fibers;
• cable accessories (cable joints, cable terminations, other cable accessories); have been described in [1, 4].

In respective manufacturers’ catalogues, for example in [15 – 18], detailed information on the design of HVAC and HVDC submarine power cables can be found.

General power Grid requirements for submarine cable links

General requirements

Power grid security requirements should be fulfilled by onshore as well as offshore power grids. Hence, submarine cable links should not adversely affect the operation of the existing onshore power grids. Requirements for onshore grids are formulated by authorities responsible for security and reliability. An example of such requirements can be found in [19]. At present, the requirements for offshore power grids are being increasingly developed [20]. When HVAC or HVDC submarine cable link is planned the following issues affecting system security and operation have to be made more specific:

• Network topology. Usually the radial reference scenario of connecting a submarine cable link to the source is considered. For meshed networks or in a cable link with intermediate load, requirements have to be adapted as appropriate.

• Type of generation source. In general, submarine cable links are assumed to be used for connecting wind sources (offshore wind farms). However, it may not always be the case. Due to the variety of marine energy resources (gas resources, energy of sea waves), the cable power line can provide a link for different types of sources (e.g. an offshore gas plant).

• Choice of cable cross-section. For radial systems, for the purpose of offshore wind farms, cable cross-section should be selected with the assumption that the power output of the offshore wind farm is equal to the total sum of installed powers of individual wind turbines comprising the wind farm.

• Maximum link capacity. This parameter is closely related to the capacity of the power grid to balance the loss of power infeed. Thus, this parameter results from the features of the load-frequency control system, and in particular on the control reserve level (seconds, minutes and hours).

• Maximum length. The maximum length of a cable link through which an offshore energy source is connected to the onshore grid is limited, which is due to large cable capacitance. This problem worsens as cable rated voltage increases. With large cable lengths, compensation reactors may prove to be ineffective. Therefore, with some approximation, an assumption can be made that the maximum length of a submarine cable line with rated voltage above 200 kV is 100 km.

• Effect on the onshore power grid. The cable link should be selected at least for normal states (N-0) and for states with single outages (N-1) in the onshore grid and taking into account at least planning contingencies in the onshore grid [21].

• Service life. Due to huge difficulties associated with the construction and repair of submarine cable lines, such systems should be highly reliable. In addition, for the construction of such system to be economically viable, an appropriate service life of the cable line is necessary. A minimum service life of 40 years meets such expectation.

• Occupational safety and health. Working on the sea as well as staying on an offshore substation is associated with an increased exposure to harmful health factors: physical, chemical, biological and psychophysical [22].

Specific features of HVAC submarine cable links

The specific features of HVAC submarine cable links require that particular attention is paid to voltage conditions in onshore grid nodes related to reactive power flow. Reactive power compensation of cable lines can be realized by variable shunt reactors. For optimal operation and maximum utilization of transmission capacity of HVAC cables various reactive control actions can be required [23, 24].

For the compensation of the reactive power of cable lines, variable shunt reactors should be adopted. A selection of reactors parameters are based on cable parameters and load flow analysis. The following options for the location of shunt reactors can be distinguished:

• Case 1 – compensation reactor installed at only one termination of the cable line (in an onshore substation).
• Case 2 – compensation reactors installed on both terminations of the cable line.

Case 1 should be considered as the preferred one.
Case 2 should be considered only where Case 1 does not ensure proper operation of the cable line.

The dynamics of power flow in cable links, due to the variable power output from offshore wind farms, may require fast-reacting systems, such as SVC or STATCOM shunt FACTS devices, to maintain the standards for acceptable voltage performance of power grids. Moreover, FACTS devices like SVC and STATCOM applied in an offshore grid can improve overall multi-machine system stability [25].

Another issue related to energy quality of an onshore grid which can be affected by the offshore grid (along with the sources installed in it) is how to eliminate high-order harmonics. To maintain proper system operating standards, it will be necessary to verify if additional AC harmonic filters should be installed in onshore substations to which the offshore grid is to be connected [26]. It should be noted that, stability is an essential precondition of power grid operation [27, 28].

Specific features of HVDC submarine cable links

Distinct requirements regarding HVDC offshore grids or HVDC offshore transmission systems can be formulated using the following parameters: rated power, rated voltage, frequency, power losses, reliability, availability and maintenance [29]. Additional criteria can include weight or dimensions of devices and systems [30].

LCC and VSC HVDC system technologies are currently available. The technological features analyses show that HVDC VSC can be used for offshore grids [30, 31], and HVDC LCC is the appropriate technology for point-to-point connections between strong systems.

Due to the requirements of the system security and operation, the following issues are particularly important:

• scope of ancillary services (e.g. participation in frequency control),
• fault ride through capability,
• black-start capability.

The onshore grid should make it possible to implement the function of black-start of the HVDC system. This procedure requires relatively low active power and an appropriate voltage in the AC node of the inverter substation. At the same time, the HVDC system should have technical capability to energize the bus-bars of a remote AC substation located on the other termination of the HVDC system.

Designing rules of HVAC and HVDC submarine cable power links

General rules for designing HVAC submarine power cables

The process of designing HVAC submarine power cables consists of several mostly electrical tasks, as well as some other ones, regarding to different research fields. In Fig. 3, an algorithm of designing HVAC submarine power cable lines has been presented.

The first step of the designing process is an investment program analysis. That kind of analysis in the case of submarine cable link needs to be carried out taking into consideration a wide spectrum of different conditions related to different kinds of possible risk, as well as having in mind various probable expansion scenarios (for instance possible variants of cable link expansion in the direction of offshore power grids). After that, detailed technical, functional, economic, location and reliability requirements of cable link are formulated.

Proper location of the cable line route, both in the offshore and onshore parts, needs to be chosen. It will allow for recognizing geomorphological soil parameters. The soil characteristic is very important, because it allows identification of the final cable routes, as well as identification of difficulties in their laying and heat dissipation conditions. Determining the likely cable line route makes it possible to calculate predicted cable length, both in the onshore and offshore parts.

In the next part of the subsection we will focus only on a case of designing a HVAC cable link, dedicated for the task of exporting the power from offshore wind farm (OWF) to onshore electric power substation.

Attention needs to be paid to correct calculation of cable current-carrying capacity and its relation to cable long-term current loading. An algorithm of cable current-carrying capacity calculation is presented in Fig. 4. Other, detailed information about the process of designing HVAC submarine power cable line can be found in [4].

It needs to be noticed that significant parts of the proposed general rules of how to design HVAC cable power lines (links) is still relevant for the case of designing the offshore cable power grids. In the case of HVAC submarine cable power grids, differently than in the case of connecting OWFs to the electric power system, the problem of reactive power management should also be taken into consideration [4]. Making the use of shunt reactors of constant inductance may not be sufficient to ensure demanded reactive power flows and voltage levels in grid nodes. Flexible control and regulation of both voltage and reactive power in HVAC offshore cable power grids may require FACTS devices, like SVC or STATCOM, to be applied.

It should also be expected that values of transmitted active powers via particular cable lines may change. They will not only be the consequence of generation power values in OWFs, but also the subject of TSO regulation policy. Therefore, the right decision on which cable conductor’s cross-sections should be selected will depend on determination of the values of active and reactive components of currents which flow through submarine power grid branches for different possible operation variants of this network.

HVAC power cable lines composing the offshore power grids will be exposed to greater short-circuit powers (currents) appearance and influence, originating from electric power systems, which are connected to these grids, and to some point also from OWFs [4]. OWFs will be mostly connected to the grid via power electronic converters and thus they will not represent significant short-circuit power sources. On the basis of calculated short-circuit currents in offshore power grid branches, it will be possible to determine maximum equivalent thermal short-circuit currents and then check if the criteria concerning short-circuit capacity of conductors and sheaths of selected submarine cables are satisfied.

General rules for designing HVDC submarine power cables

It needs to be emphasized that so far any standards concerning HVDC submarine power cables for rated voltage of above 5 kV with extruded insulation and accessories for them have not been elaborated or published. This situation is very inconvenient and as a consequence, when one is faced with the task of land-based or submarine HVDC links designing, he needs to take advantage of some previous experiences gained during the process of designing land-based and submarine HVAC cable power lines [4].

In this paper HVDC submarine power lines connected to converter substations, which are based on topology of VSC (Voltage Source Converters), are analyzed. In the case of HVDC submarine power links constructing, which serve the task of transmitting the energy from OWFs, arrangements with VSC topology based converters are the only ones that will find application and have practical meaning. Advantages of VSC technology and drawbacks of LCC (Line Commutated Converters) technology have been concisely described in [4]. Because of that, some general rules of how to choose cable cross-sections for a HVDCVSC link (i.e. with VSC topology based converters) will be described later in this paper.

Like in the case of HVAC power cables, designing HVDC power cables, both submarine and land-based ones, includes several purely technical (electrical and mechanical) tasks, as well as some other ones, related to different areas, like for instance the aspect of economic calculations or the issue of natural environment protection.

Designing a HVDC cable power link, being a component of a submarine link, is, similarly as in the case of HVAC submarine power cable designing, the process, that takes place after carrying out investment program analysis, and a task strictly connected to it. The main subject of this paper is the HVDC cable power link serving for the purpose of exporting the power from OWF to onshore electric power substation.

The steps necessary when designing HVDC submarine power cables reflect the ones appropriate for HVAC submarine power cables designing task, which have been described above. Additionally, we also need to select proper AC/DC converter substations. However, the following steps (relevant only for HVAC power cables) need to be omitted: cable line charging power determination, choice of devices for the task of capacitive reactive power compensation.

Other important issues, like environmental aspects, including requirements given in location and environmental decisions, issued by respective administrative authorities, should also be the subject of analysis.

One of the substantial tasks we also need to focus on is the issue of correct calculation of current-carrying capacity of HVDC power cables. It should be underlined that the choice of HVDC cable conductor cross-section differs from the cross-section choice for HVAC cable conductors. Because of the fact that converter substations are characterized by small overloading capacity (about 5 – 10%), short-circuit currents in quasi steady-states (on the link’s DC side) are not being analyzed. However, short-circuit currents should be taken into consideration in transient states (because of overvoltage aspects).

The minimal steps of the building permit design process include:

• develop a design concept (select the routing of the cable line),
• obtain formal legal permits/agreements required for submitting the application for building permit,
• develop building permit design, containing, in particular, the selection of the cable cross-section, cable accessories, cable coding, etc.

Assumptions for the construction of submarine cable lines – onshore segment

Basic requirements

In special cases (river crossing, wetlands, etc.), the parameters of the cable line should be individually selected and agreed with the transmission system operator.

Protection of cables at crossings

It is worth to emphasize that every crossing of existing infrastructure (e.g. pipeline) has its own characteristics and should be designed separately.

Each crossings of cable lines with various assets should be executed is such a way as to:

• cable lines do not hinder the operation of the existing infrastructure, and vice versa,
• repair and maintenance conducted on any existing infrastructure do not generate difficulties on the other asset,
• crossing does not provide any hazards to the surroundings.

Assumptions for the cable line construction – offshore segment

Survey of cable line routing

In the case of power cable links routing requirements regarding the telecommunications cables can be adopted. The minimum requirements can be found in [32]. If the routing of a cable line is being developed, basic design requirements should be taken into consideration (e.g. the location of the offshore power substations). Other factors which can affect cable line routing include: marine operations related to the cable laying and burial, construction stages or depth of burial.

Fig. 3. Procedure of designing HVAC submarine power cable line.

When selecting the subsea cable entry point to onshore area the following factors are important:

• already designated cable corridors,
• coastal conditions, including the conditions of the soil and its stability,
• other already existing civil or military infrastructures.

Submarine cable crossing and parallel laying of cables

General recommendations for design of crossings between power cables and the existing or new telecommunications assets are presented in [33] and [34].

Vertical clearance between the cable and other assets (e.g. gas pipe), as well as horizontal distance between HVDC/HVAC cables belonging to different installations should be maintained [33 – 35].

For the construction of a cable line with the parallel cable layout, the minimum horizontal distance should be 50 meters [36].

Protection against cable mechanical exposures

Burial is the primary method to protect submarine cables. Protection of cable through burial may include: jetting, ploughing, mechanical cutting, and open trench dredging. General burial depth requirements are presented in [4].

For selecting of burial technique the following factors should be analyzed [36]:

• water depth, marine conditions,
• soil / rock properties,
• environmental impact,
• cable length, mechanical properties and specific weight,
• burial depth requirement,
• technique of cable laying,
• potential burial equipment (and support vessel).

Fig. 4. Block diagram of cable current-carrying capacity calculation.
Summary and conclusions

Currently, the problem of designing submarine cable links is becoming more and more important. It results from development of offshore wind power which is driven by the renewable energy sector.

The submarine cable power grids concept, the review of submarine cable links, both existing ones and the ones which are planned to be built, general requirements for submarine cable links imposed by power system, and rules of how to design HVAC and HVDC submarine cable power links have been presented among others in the paper. Many different types of submarine power cables were designed and manufactured. Currently, the role of HVAC and HVDC cables is most often performed by cables with extruded insulation (XLPE). Current-carrying conductors are made of copper or aluminum. As for submarine cables conductors, most of them are made of copper.

The process of designing HVAC and HVDC submarine power cables includes several tasks related to different research areas. Designing of a line, which is supposed to be a component of a submarine link, is a procedure most often consisting of many iterations, taking place once the investment program analysis has been carried out, and a task, which is strictly connected to it.

The particular submarine HVAC and HVDC cable links need to be designed and constructed in such a way, that all the electrical, mechanical and environmental requirements for the cable link can be met. Secure cable link cooperation with the power system is also expected to take place, in accordance with mandatory regulations and state of the art of how electric power lines should be designed, imposed by technical standards (recommendations, guidelines). Safety of people, animals and property also needs to be ensured, when constructing, operating and exploiting the submarine cable links. The technical requirements, concerning HVAC and HVDC submarine cable links, should be treated as an integral part of the economic analysis. This analysis leads to the definition of the optimal link structure, taking into account some assumptions regarding to both environmental and economic criteria. Stated technical requirements related to submarine cable links should also take into account submarine power grids (also multi-terminal) which are planned for the future.

Choice of HVDC cable conductor cross-section differs from the choice of HVAC cable conductor cross-section. Because of the fact that converter substations are characterized by small overloading capacity, short-circuit currents in quasi steady-states (on the DC side of the link) are not taken into consideration.

In order to fully exploit the potential of offshore wind energy, many power system operators decide to build a multi – node submarine power grid. The development process is likely to rely on the integration of existing radial submarine connections into larger structures.

A significant part of discussed general rules referring to HVAC and HVDC cable power lines (links) designing is still relevant also for the case of designing the offshore cable power grids. HVAC cable power lines being components of offshore power grids will be exposed to appearance of greater short-circuit currents, originating from connected to these grids electric power systems and to a lesser extent from OWFs.

There is a quite large set of possible topologies of submarine power grids providing both the international power and energy exchange and introducing or receiving energy to/from the offshore installations. Typical classes (topologies) of submarine cable power grids include: radial topology, ring topology, star topology, topology with international coordination, and meshed topology. The final power grid topology as well as the grid development process should be a subject of detailed studies and analyses.

Acknowledgments: This research was supported by the Polish power transmission system operator PSE S.A.

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Authors: prof. dr hab. inż. Mirosław Parol, Politechnika Warszawska, Instytut Elektroenergetyki, ul. Koszykowa 75, 00-662 Warszawa, E-mail: miroslaw.parol@ien.pw.edu.pl; dr hab. inż. Sylwester Robak, Politechnika Warszawska, Instytut Elektroenergetyki, ul. Koszykowa 75, 00-662 Warszawa, E-mail: sylwester.robak@ien.pw.edu.pl; mgr inż. Łukasz Rokicki, Politechnika Warszawska, Instytut Elektroenergetyki, ul. Koszykowa 75, 00-662 Warszawa, E-mail: lukasz.rokicki@ien.pw.edu.pl; dr inż. Jacek Wasilewski, PSE Innowacje Sp. z o.o., Al. Jerozolimskie 132, 02-305 Warszawa, E-mail: jacek.wasilewski@pse.pl;

Source & Publisher Item Identifier: PRZEGLĄD ELEKTROTECHNICZNY, ISSN 0033-2097, R. 95 NR 3/2019. doi:10.15199/48.2019.03.02

Study on Transformer Tank Vibration Characteristics in the Field and its Application

Published by Ji SHENGCHANG, Zhu LINGYU, Li YANMING,
State Key Laboratory of Electrical Insulation and Power Equipment, Department of electric engineering, Xi′an Jiaotong University

Abstract. Vibration signal analysis method is one of effective methods to monitor the condition of transformer windings and core. In this paper, the multi-channels vibration measurement system is set up and the vibration signals on the oil-tank surface of running transformer in the field are measured. The influences of transformer type and sensor positions on the measured vibration signals are analyzed. The results show that, the same type of transformer’ oil tank vibration signals have almost same characteristics, and the vibration signals measured on a transformer tank surface have similar characteristics only at the same position of phase “A” (or “a”) and “C” (or “c”). Moreover, the attachment position of each sensor should be marked on the tank surface with paint, which is the reference position in the next measurement. It was recommended that the permissible error in the position of the sensor is within 5cm. The vibration characteristics acquired in the field are applied to diagnose a transformer suffered three-phase short circuit, and the results presented show the accuracy and efficiency of the acquired transformer tank vibration characteristics.

Streszczenie. Analiza sygnału wibracji jest jedną z metod monitorowania stanu uzwojeń i rdzenia transformatora. W artykule zastosowano wielokanałowy analizator wibracji do badania pracującego transformatora. Analizowano wpływ pozycji czujnika dla różnych typów transformatora. Stwierdzono, że sygnały wibracji miały podobny charakter dla jednego rodzaju transformatora i że sygnały są podobne dla tej samej pozycji czujnika. Rekomendowane jest więc zaznaczanie pozycji czujnika (z tolerancją 5 cm) dla otrzymania powtarzalności pomiarów. Otrzymane sygnały wibracji mogą być wtedy wykorzystywane do diagnostyki. (Analiza sygnału wibracji transformatora i jej wykorzystanie)

Keywords: Transformer, oil tank, vibration characteristics, in the field, windings and core
Słowa kluczowe: transformator, wibracje, diagnostyka.

Introduction

Throughout transformer’s life, mechanical shocks during transportation and installation, insulation aging, repeated thermal processes and multiple short-circuit forces will cause winding deformations or core clamping pressures drop. It leads, particularly, to the reduction of the capability to withstand future short circuit electro-mechanical stresses, to the increase of the winding vibration and, consequently, to the increase of the solid insulation mechanical fatigue. In this way, the isolation can be degraded and short circuits between turns will appear. The defects mentioned can create PDs and combustible gases, thus converting the mechanical defect into the problem of insulation [1, 2]. From these considerations, the relevance of an early detection of winding deformations or core clamping condition is clear. Some techniques, such as frequency-response analysis (FRA) [1, 3] or leakage reactance measurement （LRM） [4], are widely used to detect changes in transformer geometry, especially winding deformations. Although, in recent years, some publications have appeared reporting some online applications of these techniques [5, 6], at the moment they are used only in offline tests. Steady-state vibration recorded on the tank surface provides essential information about running conditions both for windings and the magnetic core in power transformer [7]. Therefore, vibration signal analysis method is one of effective methods to monitor the condition of transformer windings and core, it is a complementary technique to FRA or LRM having the advantage that it can be used for on-line monitoring and, thus, catastrophic failures can be avoided between successive maintenance outages.

In recent years, much research has been devoted to the vibration signal analysis method for monitoring power transformer [7-17]. In [7-10], different vibration models have been developed to calculate tank vibrations, taking into account transformer operating conditions, such as on-load current, applied voltage, temperature, and etc. There are also many measurements taken on the test transformer in laboratory or on the power transformer in manufactory [11-17], and some tank vibration characteristics were presented, such as the relationship between the vibration amplitude and load current. In [14], it is presented that the acquired vibrations must be identified and related to the state change of the monitored transformer by means of comparisons either with a similar new transformer or the averaged values of a set of transformers of similar type and age. But this conclusion has no data supporting. Until now, only [18, 19] presents some vibration data measured on the running transformer in the field, but in-depth study has not been done, such as comparison of vibration signals for different type of transformer, comparison of vibrations at different positions on the transformer tank. Therefore, it is necessary to study the transformer′ oil-tank surface vibration characteristics in the field, which can guide the application of vibration signal analysis method.

In this paper, based on the multi-channels vibration measurement system, the tank vibrations of some running power transformers in the field are measured, the influence of transformer type and sensor position to the vibration signals are analyzed deeply. Furthermore, the transformer tank vibration characteristics is applied to diagnose a transformer which has suffered three-phase short-circuit.

Measurement system

The multi-channels vibration measurement system developed for the running transformer is shown in Fig.1. Each of these vibration accelerative sensors has a sensitivity of 100mv/g, its scope is 50g (g is acceleration of gravity) as well as the frequency response within 0.4~10kHz range. The function of charge amplifier is to transform the charge signals exported by the vibration sensors to voltage signals and magnify it synchronously. The A/D card with USB ports is chosen to convert the analogy voltage signals to digital signals with a resolution of 12 bit, maximum sampling frequency is of 200 kHz. Then the digital vibration signals are exported and processed by the notebook computer (NB).

Fig.1. Vibration measurement scheme for transformer

The recommended sensor positions are shown in Fig.2 [17]. Usually 12 points are used, 6 at high voltage bushing side of the transformer tank and 6 at low voltage bushing one (further HV and LV side, respectively), approximately equally distant from monitored elements. The simultaneous installation of all sensors is not necessary; even one sensor can be used, being sequentially installed on all points. But in this paper, vibration signals of 12 points are acquired synchronously. Temporary sensor installation is easily achieved through magnetic fixation. Points of sensor installation have not been chosen near manholes, pipelines or stiffening ribs. To permit repetitive sensor installation just at the same positions, they can be marked on the tank surface with paint. In [17], it was recommended that the permissible error in the position of the sensor is of 15-30cm, however, there has no vibration data measured in the field for supporting such recommendation, which is worthy of further study.

Fig.2. Recommended sensor positions on a transformer tank

Fig.3 shows the on-scene picture of multi-channels vibration measurement system which is used to measure the vibration signal of transformer tank in the field. In the dotted circle, it is one of vibration sensors installed on the tank through magnetic fixation, and its amplified image is shown in the real-line circle. Other instruments shown in Fig.3, such as charge amplifier, A/D sampling card and notebook computer, are also illustrated via the dotted arrows.

Fig.3. Transformer vibration measurement in the field
Transformer Vibration signal measurement and analysis

A. Same position for the different transformer

1) Different type For different types of transformers, the vibration signals measured at the same positions of oil-tank surface are shown in Fig.4 (a) and Fig.4 (b). The transformer type is SSZ9-M-50MVA/110kV and SFZ-20MVA/110kV respectively.

Fig.4 shows that to the different types of transformers, the vibration signals measured at the same surface positions present different characteristics. The main vibration frequency of SSZ9-M-50MVA/110kV transformer is of 200Hz, however, SFZ-20MVA/110kV transformer′ is of 300Hz. Furthermore, for the vibration amplitude of each frequency, the difference of two types is also very obvious. The nonlinearity of core magnetostriction leads to the existence of high order harmonic. But why the magnitude of 200Hz or 300 Hz is the largest? The reason is that the core has the primary natural vibration frequency of about 200Hz or 300Hz. With the excitation of core magnetostriction, resonance takes place. For different types of transformers, the cores are clamped with different pressure. Therefore, their primary natural vibration frequency is different, which leads to above measurement results. It also means that the vibration signals measured on different types of transformers’ tank surface are not comparable, even though which are acquired at the same position.

Fig.4. Comparison of Vibration signals for different types of transformers

2) Same type The vibration signals on the oil-tank surface of two same types of transformers (Type: SFZ-240MVA/345kV, it is called as 1# and 2# transformer respectively) are measured and recorded. When two transformers’ on-load current is equal, the frequency spectrums of vibration signals are shown in Fig.5. Fig.5 shows that spectrum characteristics of two same types of transformers are almost similar. However, at the main vibration frequency (300 Hz), the vibration magnitude of 1# and 2# transformer is 11.4mV and 12.7mV respectively, that is to say that 2# transformer’s vibration magnitude is greater than 1# transformer’ about 10.2%. Furthermore, at the other frequency, the vibration magnitude of 2# transformer is slightly greater than that of 1# transformer. The reason leading to above phenomenon is that 1# and 2# transformer’s tap position is on -2.5% and 2.5% respectively when vibration measurement is taken on. In [15], it has been presented that the vibration magnitude of each frequency component is almost proportional to square of open-circuit voltage. Such oil tank vibration characteristics acquired in this paper conform to results in [15]. If two transformers′ vibration signals are normalized according to square of applied voltage, not only their spectrum characteristics are similar, but also their vibration magnitude is almost equal. For example, at the primary frequency (100Hz), the vibration magnitude of 1# and 2# transformer is 1.89mV and 2.14mV respectively. The 2# transformer’s vibration magnitude is 1.94mV when it is normalized by square of 1.05 (namely, (1+2.5%)/(1-2.5%)), which is almost equal to that of 1# transformer (1.89mV).

Fig.5. Vibration spectrums of two same types of transformers at equal on-load current

When 1# transformer’s tap position (-2.5%) is as same as 2# transformer’, the vibration signals of 1# and 2# transformer’ oil-tank surface are measured and recorded at current of 310A and 350A respectively, the frequency spectrums of vibration signals are shown in Fig.6.

Fig.6. Vibration spectrums of two same types of transformers at different current

Fig.6 shows that spectrum characteristics of two same types of transformers are almost similar at different current. However, at the fundamental frequency (100Hz), the vibration magnitude of 1# and 2# transformer is 1.89mV and 2.45mV respectively, that is to say that 2# transformer’s vibration magnitude is greater than 1# transformer’ about 29.6%. The 2# transformer’s vibration magnitude is 1.92mV when it is normalized by square of 1.13 (namely, 350/310), which is almost equal to that of 1# transformer (1.89mV). The reason leading to above phenomenon can be explained as follows.

• The fundamental frequency component consists of vibration signals of the windings and that of the core. The high frequency vibration of tank surface is mainly caused by magnetostriction of core and independent of windings vibration [15].

• The core vibration is mainly dependent on the voltage applied to the primary winding which is independent of the load [15].

• Because 1# transformer’s tap position is as same as that of 2# transformer, then the vibration signal’ harmonic components of two transformers are almost equal. Furthermore, the fundamental frequency component caused by core vibration should be equal.

• The windings’ fundamental frequency vibration signal was proportional to the square of the loading current [15]. Therefore, when two transformers′ vibration magnitudes of fundamental frequency are normalized by square of loading current, they will be almost equal.

The above results mean that when there are no historical data, same types of transformer’ vibration signals normalized according to square of applied voltage and loading current (only for fundamental frequency component) can be compared with each other.

B. Different position for the same transformer

1) Top and bottom position at side of same phase From Fig.2 it can be seen that for the vibration sensors installed at the top and bottom of the same phase′s oil tank surface, such as position “1” and “2”, position “3” and “4”, they are at the symmetrical position. For a transformer, the spectrums of vibration signals measured at position “1” and “2” are shown in Fig.7.

Fig.7. Vibration spectrums at the top and bottom of tank

Fig.7 shows that for the vibration signals at top-bottom symmetrical position, their spectrums are almost similar. However, there exists obvious difference at some frequency, such as 100Hz, 300Hz, 600Hz, and etc. This phenomenon can be explained as follows.

• The main sources of tank vibration are forces appearing in the winding and the core.

• Winding vibrations are due to electro-dynamic forces caused by the interaction of the current in a winding with leakage flux. These forces are proportional to the current squared. Obviously, the leakage flux of the winding bottom position is almost equal to its top position′s. On the other hand, the current flowing through the winding is independent of position. Thus, for the top and bottom position, the winding vibrations are similar.

• The core vibration caused by magnetostriction forces is proportional to squared voltage [15]. Thus, for the top and bottom position, the core vibrations are similar.

• Core and winding vibrations interact and transmit through the oil and the transformer supporting elements to the tank. For the top-bottom symmetrical positions, the oil tank structures are different, for example, at the top of tank there are HV, LV bushings and oil reservoir, but the bottom of tank is fixed on the ground, which means that the tank surfaces at the top and bottom position have different nature vibration characteristic.

• Therefore, although the vibrations of core and winding are almost similar, the vibration spectrums at top-bottom symmetrical position still have some differences.

The above results and analysis mean that for the top-bottom symmetrical position shown in Fig.2, such as position “1” and “2”, position “3” and “4”, the measured vibration signals can not be compared with each other.

2) HV and LV side In this case, the vibration signals are measured at two sides of the tank: at high voltage bushing side and at low voltage bushing one. From Fig.2 it can be seen that these positions are left-right symmetrical. For a transformer, the spectrums of vibration signals measured at position “5” and “7” are shown in Fig.8, which presents that the vibration signals’ magnitudes of position “5” and “7” are equal at almost all frequencies, except at frequency of 200Hz, 300Hz and 700Hz. Such results also demonstrate that vibration signals of HV and LV side can not be used to estimate the condition of windings and core by comparing with each other.

Fig.8. Vibration spectrums on the tank of high and low voltage sides

The analysis about difference of vibration characteristics at the top-bottom symmetrical position can explain the phenomenon shown in Fig.8. The structures of HV and LV side oil tank are different, especially, the HV and LV bushing have different structural dimension, which makes the HV and LV side tank have different natural vibration characteristics. Thus, although activated by the same vibration source (winding and core vibration), the left-right symmetrical positions at the tank of HV and LV side have different vibration response.

3) The same position at side of different phase For a transformer, the spectrums of vibration signals measured at position “1”, “3” and “5” are shown in Fig.9.

Fig.9 shows that at position “1” and “5”, the vibration signals measured on the oil tank surface present same characteristics, which are different from that measured at position “3”. The above measurement results are easy to explain. Each phase’s tank vibration will be influenced by that of the other two phases. For example, the position “1” will be influenced by position “3” and “5”. On the other hand, the influence effect is related to the distance between each other. The position “1” and “5” are symmetrical, that is to say, the structure of tank, vibration source and influencing of other phases are both same at these two positions, which leads to their same vibration characteristics. However, the position “3” is at the middle of oil tank, and the vibration signal measured at this position would be influenced by the other two phase’s winding and core vibrations, moreover, the distances from position “3” to position “1” and “5” are equal, which leads to its vibration characteristics be different from the sideward tank’s.

Fig.9. Vibration spectrums on the tank of HV sides

The above measurement results means that the vibration signals measured at the same position on oil-tank of phase “A” (or “a” ) and “C” ( or “c” ) have same characteristics, therefore, each phase’s windings and core condition can be diagnosed by comparing their vibration signals. However, the above method can not be used to vibration signals measured on tank surface of phase “B” (or “b”).

4) Around the measured position When the vibration signals on the transformer oil tank are measured in the field, a sensor attachment position is regarded as the referred point, and then the installing position of this sensor is moved up, down, left and right 5cm and 10cm respectively. The vibration signals at the original position and changed position after sensor is moved are all recorded. When sensor is at the original position and position is moved up 5cm, the measured vibration signals′ spectrums are shown in Fig.10.

Fig.10. Vibration spectrums at primary position and upward 5cm

From Fig.10 it can be seen that when the attachment position of sensor is moved up 5cm, the vibration signal has slightly change compared with the original position’. Only the vibration magnitude at the main frequency increases from 12.9mV to 13.4mv, which enlarges about 3.9%. According to [5] and [7], when the relative value of the transformer vibration magnitude changed above 20%, the windings and core were deemed having serious fault and the transformer must be out of running. This means that the difference of 3.9% is within the acceptable scope and will not lead to wrong discrimination.

When the sensor is moved down, left and right 5cm respectively, the same result can be drawn as the above, which is that the vibration signal measured at the new position has slightly difference compared with the original position ‘.

When the sensor is moved up 10cm from its original position, the vibration signals are shown in Fig.11.

Fig.11. Vibration spectrums at primary position and upward 10cm

From Fig.11 it can be seen when the sensor is moved up 10cm, the vibration signal measured at the new position has little change at some frequencies, such as 100Hz, 200Hz and 400Hz. However, at other frequencies there is obvious difference compared with that of the original vibration signal, such as at the main vibration frequency (300Hz), the vibration magnitude enlarges from 12.9mV to 14.8mV, and increases about 14.7%. Furthermore, such great change is caused by the displacing of vibration sensor’s attachment position but not the fault of windings or core. That is to say, at this condition, any decisions can not be made about the compression of windings or core.

When the sensor is moved down, left and right 10cm respectively, the same result can be drawn as the above, which is that the vibration signal measured at the new position has great difference compared with that of the original position.

Also, in the field the vibration signals were measured when the attachment position of sensor were changed at the range of 6~9cm. Compared with the original vibration signal, the signal measured at the new position demonstrate that farther the distant of sensor are moved, greater the vibration signal changes.

With respect to the above results, it was recommended that the permissible error in the position of the sensor is within 5cm, which is different from the conclusion drawn in [17].

Application of vibration characteristics

In the HeXing substation, Hangzhou power bureau, China, 1# transformer suffered three-phase short-circuit at Nov. 2009, the short-circuit fault lasted for about 90s. The transformer type is OSFPSZ9-150MVA/220kV. Then, the vibration signals were acquired by the multi-channel vibration measurement system. According to the Part III, B(3), the vibration signals measured at the same position on oil-tank of phase “A” (or “a” ) and “C” ( or “c” ) have the same characteristics and they can be used to diagnose each phase’s windings and core condition by comparing with each other. The results of the measurement at position “7” and “11” are shown in Fig. 12.

Fig.12. vibration spectrums at the top and bottom of tank

From Fig.12 it can be seen that spectrum characteristics of vibration signals at position “7” and “11” are almost similar. However, at the fundamental frequency (100 Hz), the vibration magnitude of position “7” and “11” is 3.78mV and 2.58mV respectively, which means that the vibration magnitude of position “7” is greater than that of position “11” about 46.5%.

According to [14], [17] and [19], the fundamental frequency component consists of vibration signals of the windings and that of the core, and the high frequency vibration of tank surface is mainly caused by magnetostriction of core and independent of windings vibration. Therefore, the high frequency vibration measured on the tank surface can be used to diagnose the condition of core directly. In this case, for the harmonics with the frequency of 200Hz and above, the magnitudes of vibration signals at position “7” and “11” are almost equal, which means the condition of 1# transformer core is well. Thus, it can be concluded that the winding at position “7” has occurred serious fault which leads to the great increase of the magnitude of fundamental frequency vibration.

In the same substation, there is a 2# transformer, whose type and age are as the same as that of 1# transformer. In order to verify the above conclusion, we also measured the vibration signal at the same position as 1# transformer’ (“7”). Two transformer’ load current and tap changer position are also same. The result is shown in Fig.13.

According to Part III, A(2), same types of transformers vibration signals normalized according to square of applied voltage and loading current can be compared with each other. From Fig.13 it can be seen that spectrum characteristics of 2# transformer’ vibration signal at position “7” is as the same as that of 1# transformer at position “11”. For 2# transformer, the fundamental frequency component of vibration signal at position “7” is 2.6mV, which is almost equal to that of 1# transformer at position “11”. In the other word, at the same position (“7”), the fundamental frequency component of 1# transformer is larger than that of 2# transformer about 46.5%. Thus, it can be concluded that 1# transformer’s windings at position “7” has occurred serious fault, which further verifies above diagnostic result. Therefore, it is decided that 1# transformer must be shut down and overhaul must be performed in the manufactory.

Fig.13. 2# transformer vibration spectrum at position “7”

Based on above diagnostic result deduced from vibration signal analysis method, the transformer is isolated from the system and transported to the transformer manufactory, where the overhaul of windlass cover is performed, the visual inspection results are shown in Fig.14.

Fig.14. Photos of windings having deformation

From Fig.14 it can be seen there is a severe deformation in the upper disks of low voltage windings of phase “c”. The results verify above diagnostic result. In the other word, the transformer tank vibration characteristics acquired in the field can be used to monitor the condition of transformer.

Conclusions

Vibration signal analysis method is one of effective methods to monitor the condition of transformer windings and core, which is a complementary technique to FRA or LRM and has the advantage of on-line monitoring. However, until now, in-depth study has not been done on transformer tank vibration characteristics in the field, which limits its effective application. In this paper, the vibration signals on the oil tank surface of power transformer are measured in the field, and the vibration characteristics of transformer tank are studied, which is outlined as follows.

1) The vibration signals on the oil-tank surface of different types of transformers have great difference and they can not be compared with each other. On the other hand, the same types of transformers’ oil-tank vibration signals have almost same characteristics, which mean that if there are no historical data, diagnosis can be carried out by comparing same types of transformers’ vibration signals at the same position. It must be stated that when comparison is made, vibration signals should be normalized according to square of applied voltage and loading current (only for fundamental frequency component).

2) For the vibration signals measured on the oil-tank surface of a transformer, at the top-bottom symmetric positions of same phase and the left-right symmetric positions of HV and LV side, the vibration characteristics have obvious difference at some frequencies. Therefore, the vibration signals measured at these positions can not be compared with each other to diagnose the condition of transformer windings and core.

3) The tank surface vibration signals at the same position of phase “A” (or “a”) and “C” (or “c”) have similar characteristics, and they can be used to diagnose each phase’s windings and core condition by comparing with each other. However, the above method can not be used to vibration signals measured at the same position of phase “B” (or “b”), whose characteristics are different from that of phase “A” (or “a”) and “C” (or “c”).

4) When vibration signals of transformer oil-tank surface are measured, the attachment position of each sensor should be marked on the tank surface with paint, which is the reference position in the next measurement. It is recommended that the permissible error in the position of the sensor is within 5cm, which is different from the conclusion drawn in [17].

At last, the transformer tank vibration characteristics presented in this paper was applied to diagnose a transformer suffered three-phase short circuit, and it was deduced that windings at position “7” has occurred serious fault. The overhaul of windlass cover verified the validity of diagnostic result. The conclusions presented in this paper will guide the application of vibration signal analysis method in the field.

Acknowledgement: This paper is supported by the National Natural Science Foundation of China (Grant No.50907046).

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[4] A. Babare, F. Cannata, G. Caprio, S. Sacchetti, and G. Zafferrani, “Ennel-diagnossis of on- and off-line large transformers”, in Proc. Cigré Symp., Berlin, Germany, pp. 110–04, 1993.
[5] T. Leibfried and K. Feser, “Monitoring of power transformers using the transfer function method”, IEEE Transactions on Power Delivery, vol. 14, no. 4, pp. 1333–1339, Oct. 1999.
[6] S. Birlasekaran and F. Fetherston, “Off/On-Line FRA condition monitoring technique for power transformer”, IEEE Power Eng. Rev., vol. 19, pp. 54–56, Aug. 1999.
[7] J.C. Lavalle, “Failure Detection in Transformer Using Vibrational Analysis”, MS dissertation, MIT, Cambridge, MA, USA, 1986.
[8] S.H Lee, “Modeling of Winding Vibration Amplitude for Diagnosis of Power Transformer”, MS dissertation, Soongsil University, 1994.
[9] C. Booth, J. R. McDonald, and R. Aresi, “The use of neural networks for the estimation and classification of vibration behavior in power transformers,” in Proc. Amer. Power Conf., 1995, pp. 1132–1135.
[10] Belén García, Juan Carlos Burgos, and Ángel Matías Alonso, “Transformer Tank Vibration Modeling as a Method of Detecting Winding Deformations—Part I: Theoretical Foundation”, IEEE Transactions on Power Delivery, Vol.21, No.1, pp. 157-163, 2006.
[11] M.A Sanz-Bobi, A. Garcia-Cerrada et al, “Experiences Learned from the On-line Internal Monitoring of the Behavior of a Transformer”, in Electric Machines and Drives Conference Record, 1997. IEEE International, TC3/11.1-TC3/11.3, 1997.
[12] Masato Mizokami, Masao Yabumoto, Yasuo Okazaki, “Vibration Analysis of a 3-Phase Model Transformer Core”, Electrical Engineering in Japan, Vol.119, No.1, pp. 1-8, 1997.
[13] Chan-Soo Chung, Chi-Hyoung You, and et al, “Fault Discrimination of Power Transformers Using Vibration Signal Analysis”, in Integrating Dynamics Condition Monitoring and Controlling Conference 21st Century, Rotterdam, Holand, pp.523-529, 1999.
[14] Cipriano Bartoletti, Maurizio Desiderio, Danilo Di Carlo, et at, “Vibro-Acoustic Techniques to Diagnose Power Transformers”, IEEE Transactions on Power Delivery, Vol.19, No.1, pp. 221-229, 2004.
[15] Ji Shengchang, Cheng Jin, Li Yanming, “Research on Vibration Characteristics of Windings and Core of Oil –filled Transformer”, Journal of Xi’an Jiaotong University，Vol.39, No.6, pp. 616~619, 2005
[16] Belén García, Juan Carlos Burgos, and Ángel Matías Alonso, “Transformer Tank Vibration Modeling as a Method of Detecting Winding Deformations—Part II: Experimental Verification “, IEEE Transactions on Power Delivery, Vol.21, No.1, pp. 164-169, 2006.
[17] Ji Shengchang, Luo Yongfen, Li Yanming, “Research on Extraction Technique of Transformer Core Fundamental Frequency Vibration Based on OLCM”, IEEE Transactions on Power Delivery, Vol.21, No.1, pp. 1981-1988, 2006
[18] Mechefske CK, “Correlating Power Transformer Tank Vibration to Winding Looseness”, Insight—J. Non Destruct. Test. Cond. Monitor., Vol. 37, no. 8, pp. 599-604, 1995
[19] Z. Berler, A. Golubev, V. Rusov, and et al, “Vibro-Acoustic Method of Transformer Clamping Pressure Monitoring”, in Conference Record of the 2000 IEEE International Symposium on Electrical Insulation, Anaheim, CA USA, April 2-5, pp.263~266, 2000.

Authors: Ji Shengchang is with Xianning west road 28#, State Key Laboratory of Electrical Insulation and Power Equipment, Xi′an Jiaotong University, Xi′an, ShannXi, Province, Republic of China. (email: jsc@mail.xjtu.edu.cn) Li Yanming is with Xianning west road 28#, Department of electric engineering, Xi′an Jiaotong University, Xi′an, ShannXi, Province, Republic of China. (email: ymli@mail.xjtu.edu.cn)

Source & Publisher Item Identifier: PRZEGLĄD ELEKTROTECHNICZNY (Electrical Review), ISSN 0033-2097, R. 87 NR 2/2011

Testing of LV Switchgear for Powering Cellulose Fiber Breaking Systems

Published by Adam SMOLARCZYK1, Tadeusz DASZCZYŃSKI1, Sławomir FISZER2
Warsaw University of Technology, Electrical Power Engineering Institute (1), Elektroteam Sp. z o.o. (2)

Abstract. The article describes an innovative low voltage switchgear for use in the pulp and paper industry. The switchgear is an innovative structure designed to supply single-phase systems with significant currents above 1 kA. The switchgear uses a special transformer with windings connected in the V system. The scope of the tests of the secondary and primary circuits of the switchgear and their selected results are presented.

Streszczenie. W artykule opisano innowacyjną rozdzielnicę niskiego napięcia do zastosowania w przemyśle celulozowo-papierniczym. Rozdzielnica stanowi innowacyjną konstrukcje przeznaczoną do zasilania jednofazowych układów o prądach przekraczających 1 kA. W rozdzielnicy zastosowano transformator specjalny z uzwojeniami połączonymi w układzie V. Przedstawiono zakres zaproponowanych, w ramach projektu badawczego, badań obwodów wtórnych i pierwotnych rozdzielnicy oraz wybrane ich wyniki. (Badania rozdzielnicy nn do zasilania układów rozbijania włókien celulozy).

Keywords: LV switchgear, switchgear testing, pulp and paper industry.
Słowa kluczowe: rozdzielnica niskiego napięcia, testowanie rozdzielnic, przemysł celulozowo-papierniczy.

Introduction

The wood and paper industry includes the sawmill industry, the board and plywood industry, the furniture industry, and the pulp and paper industry. The pulp and paper industry in Poland is developing well. It is a branch of industry that arouses great interest among foreign investors. The production of cellulose and paper requires large amounts of water. Therefore, the largest plants are located on large rivers. The largest paper producers include plants in Świecie, Kwidzyn, Ostrołęka, Kostrzyn, Kielce, Szczecin, Krapkowice and Klucze [1].

The functional requirements for the tested low voltage switchgear (LV) have been defined for the purposes of supplying equipment and production lines of the domestic pulp and paper production plants. These plants use techniques for the production of pulp and paper, as well as derived products such as, for example, cellulose sponges. Manufacturing processes require significant amounts of thermal energy obtained through electricity. The control of the production process requires the supply of electricity with specific voltage and current parameters in the required power cycles.

The pulp and paper industry is characterized by a high degree of production diversification, so the technologies of power devices (including switchgears) should be flexible in design (adaptation to the needs of a given production line at the design stage) and operation (adapting the supply conditions to the requirements of a given manufacturing process).

In Poland, there are many pulp and paper plants subject to the IPPC Directive (Integrated Pollution Prevention and Control) [2]. Pursuant to this directive, the volume of pollutant emissions in manufacturing processes is subject to BAT standards (Best Available Techniques) [3], which define emission limit values and are used in larger plants in the European Union. Power equipment should enable manufacturing processes to be adapted to the directive. The functional requirements of the LV switchgear used to supply the cellulose fiber breaking systems are:

– low supply voltage below 1 kV,
– significant current of more than 1 kA,
– possibility of cyclical supply of loads with various current levels,
– modularity of power devices enabling configurability and adaptation of power systems to the specifics of a given production line

The innovativeness of the switchgear (compared to other LV switchgears) is based on the use of a special transformer in each cell supplying cellulose cooking systems, the windings of which are connected to a V system.

Testing of LV switchgears

According to [4] the features of low-voltage switchgear should ensure compatibility with the rated data of the circuits to which it is connected, and the installation conditions should be declared by the switchgear manufacturer. All devices, electrical apparatus and low-voltage switchgear circuits should be so arranged as to facilitate operation and maintenance, and at the same time to maintain an appropriate degree of safety.

It should be noted that, in accordance with the [4] standard, there are three methods of verification (Fig. 1): tests, comparisons, calculations – they are considered equivalent. It does not mean, however, that each of the points in the standard can be verified in any way by one of the three methods. It has been precisely defined how (using an appropriate method) individual requirements can be verified. In practice, this means that e.g. the short-circuit withstand requirements cannot be verified by calculation, but have to be verified by a test. It should be remembered that testing the switchgear under short-circuit conditions is a destructive test.

According to [4] the tests of impulse withstand voltage, temperature rise limits, short-circuit withstand test, EMC and mechanical operation should be carried out. There is no more reliable method of verification than hardware laboratory tests, therefore, although the standard allows for some points to be verified by e.g. calculations, taking into account the safety of operation and powered devices, as well as the correctness of the switchgear parameters verification, the Ordering Party accepted the test at all points using the most appropriate reliable method, that is, through research.

There are some methods that can be used for verification of construction of a switchgear like FEM methods [6]. The process can be divided into two-stage simulation approach which includes electromagnetic and CFD analysis coupled together. The output of electromagnetic simulation is heat loss generated as the results of Joule heating and induction of eddy current on sheet metal parts of the enclosure. Heat loss in an input for further CFD simulation. CFD simulation is used to calculate radiation and natural convection.

Fig. 1. Responsibility for tests according to standard PN-EN 61439 [5]
Switchgear construction

The LV switchgear was tested in the Laboratory of Electrical Apparatus and Switching Process in Electrical Power Engineering Institute at Warsaw University of Technology. The switchgear was manufactured by Electroteam Sp. z o.o. and the tests were carried out from May till end of June 2020.

Based on the analysis of the technical documentation of the LV switchgear, it was found that it consists of a power supply section and six outflow sections supplying cooking cells. (Fig. 2) [7].

Fig. 2. The appearance of the six sections of the LV switchgear

The switchgear delivered for testing consists of the main power supply section equipped with the LS Susol AN-16C3- 16A main switch (QG) and one outlet section supplying the cooking cell [7]. The cooking cell feeding section consists of:

– LS Susol TS 800N section compact switch (Q1),
– two power contactors (K1.1, K1.2) Metasol MC-800a by LS for switching the primary terminals of the TR1 transformer of the cooking cell,
– a transformer (TR1) type 3FR AN with a electrical power of 242 kVA by BREVE,
– Socomec three-phase network parameters analyzer (AS1) DIRIS A10 installed on the upper voltage side of TR1 transformer,
– digital overcurrent relay (SEP1) Sepam 10 B 43E by Schneider Electric for low voltage side circuits of transformer TR1,
– a single-phase converter of network parameters (PV1) P30P by Lumel installed on the lower voltage side of the TR1 transformer,
– digital transformer TR1 (ZT1.1) temperature control relay of the TR-100 type by Novatek-Electro.

In addition to the above-mentioned elements, the switchgear provided for testing includes elements such as fuse switch disconnectors, installation switches, lamps, switches, current transformers, a fan, a 24 V DC power supply, auxiliary relays.

Fig. 3 shows a single-line diagram of the LV switchgear delivered for testing. It shows all the main above-mentioned primary elements as well as protection and measurement systems used in the switchgear.

Fig. 3. Single-line diagram of the switchgear provided for testing [8]

An important element of each outlet section of the switchgear is a transformer, the windings of which are connected in a V system (the use of this type of transformer proves the switchgear innovation). A special transformer of this type is used as intermediary devices to reduce the unfavourable load asymmetry in a three-phase supply network in the case of supplying a single-phase load with high power from this network. The traditional supply of high-power single-phase loads with phase-to-phase voltage from a three-phase network causes a strong asymmetry in the electrical network by loading only two phases. The essence of the issue is explained in Fig. 4.

Fig. 4. Explanation of the method of increasing the uniformity of the load in a three-phase network by using a transformer with windings connected in a V system [9]

The switchgear uses a special transformer type 3FR AN with an electrical power of 242 kVA by BREVE (Fig. 5)

Fig. 5. The appearance of the special 3FR AN transformer before installation in the switchgear

The transformer is supplied from a three-phase LV network with phase-to-phase voltage of 400 V. The voltage (and current) is changed on its secondary side by applying phase-to-phase voltages to the appropriate taps (installed in phases L1, L3) of the transformer’s primary side. The range of obtainable voltages and currents on the secondary side of the transformer are shown in Table 1

Table 1. 3FR AN transformer available secondary voltage and current

V60708090100110120130
kA2,52,52,52,52,32,22,01,85
.

It should be noted that the transformer loads the network in phases L1 and L3 with a current of approx. 349 A, and in phase L2 with a current of approx. 698 A.

The scope of the research carried out

After reviewing the technical documentation of the switchgear and comparing its features with devices available on the market (task 1 of the research project [7]), the following switchgear tests were carried out (or planned):

– current path tests and checking the current conduction system (task 2),
– performance of voltage resistance tests (task 3),
– performance of functional short-circuit tests (task 4),
– testing of protection automation systems (task 5).

Tests of current circuits and checking the switchgear current conduction system

The laboratory setup for testing the current load capacity was located in the Laboratory of Electrical Apparatus and Switching Processes of the Institute of Electrical Power Engineering, Warsaw University of Technology. The test circuit consisted of: inductive regulator, short-circuit transformers, switchgear under test. The system was powered by mains voltage, the short-circuit transformers were connected in a triangle on the primary side. The short-circuit transformers were star-connected on the secondary side, and the voltage on this side of the short-circuit transformers was 110 V AC. The tested LV switchgear was connected to short-circuit transformers with 2x YKY 1×150 mm2 cables per phase. The currents on individual phases were measured with a CMP-2000 clamp meter.

Impulse withstand voltage test

The following tests were performed:

– test with withstand voltage at the mains frequency of main, auxiliary and control circuits connected to the main circuit,
– withstand voltage test at the mains frequency of auxiliary and control circuits not connected to the main circuit,
– test with impulse withstand voltage of the main, auxiliary and control circuits connected to the main circuit,
– test with impulse withstand voltage of auxiliary and control circuits not connected to the main circuit.

All tests were made in accordance with [4]. Voltage levels depending on the voltage of the main, control and auxiliary circuits.

Short-circuit withstand strength test

According to [4], for switchgears where the contractual short-circuit current exceeds 10 kA, the rated short-circuit currents must be tested and verified. When verifying, the following should be used:

– if the distribution system under verification covers several variants, the least favorable one should be selected,
– if the tested kits are the least favorable variants from a wider group of products in the distribution system, then the test results can be used to evaluate similar variants without performing tests.

Fig. 6. View of the distribution of protection and measurement automation systems in the switchgear cell
Tests of protection automatics and measurement systems used in the switchgear

The protection automatics and measurement systems of the tested switchgear (secondary circuits) include (Fig. 6):

– temperature control relay (ZT1.1) TR-100,
– overcurrent relay (SEP1) Sepam 10 B 43E,
– network parameters analyzer (AS1) DIRIS A10,
– network parameters transducer (PV1) P30P.

Fig. 3 shows a diagram of the connection of individual elements of the switchgear secondary circuits. Elements such as: TR-100 relay, DIRIS A10 analyzer and P30P transducer exchange data (temperature, currents, voltages, powers) with the supervisory system using the Modbus RTU protocol and telecommunications link in the RS-485 standard.

As part of task five of the research project [8] tests were carried out:

– analysis of possible disturbances in the operation of the tested system,
– analysis of the control and signaling system,
– analysis of the correct selection of protection functions and their interaction within the protected device,
– analysis of the correctness of the selection of settings for protection functions used in the tested system,
– laboratory tests of the functions of the SEPAM 10 B device,
– laboratory tests of the TR-100 device functions,
– tests of correctness of data exchange between switchgear devices and the supervisory system.

Most of the tests (forcing currents and voltages) were performed using a CMC 256plus microprocessor tester [10] and a precise resistance decade MDR-93-6b (simulation of Pt100, KTY83 temperature sensors).

Fig. 7. Diagrams of the set characteristics of the THERMAL, I>, I>>> functions of the Sepam 10 B relay in relation to the exemplary thermal characteristics of the transformer
Fig. 8. Graphs of theoretical and actual characteristics of temperature measurements by TR-100 relay

On the basis of the conducted tests, it was determined whether the tested device behaves as expected, e.g. whether it provides (or not) a signal to its binary outputs after simulating a specific type of disturbance. For example, Fig. 7 shows the resultant characteristic t(I) of the current functions activated and set in the Sepam 10 B relay. It is also shown how the characteristics of the overcurrent functions set in the relay should follow the theoretical thermal (heating) characteristics of the TR1 transformer, so that it is not damaged.

Fig. 8 shows the R(temp) charts of the theoretical and real characteristics of the TR-100 relay (channels 2 and 3). As you can be seen, the actual graphs of temperature measurement by the TR-100 relay are in line with the theoretical graph of temperature changes suitable for the Pt100 sensor.

Conclusions

The switchgear provided for testing is innovative due to the specificity of its application (pulp and paper industry) and the placement of a special transformer with windings connected in the V system in its cell. The tests performed as part of the research project were divided into several tasks.

Comprehensive tests, according to [4], to which the switchgear was given concerned: (a) tests of current paths and checking the current conduction system, (b) voltage withstand tests, (c) functional short-circuit tests, (d) tests of protection automatics and measurement systems.

The switchgear tests were positive. In most cases, the tested switchgear systems (primary and secondary circuits) behaved as expected. In the case of protection systems, the behavior of the Sepam 10 B and TR-100 relays was, in some cases, not as expected.

REFERENCES

[1] Bulletin of Association of Polish Papermakers, 2016-2018, No 11, ISSN 1436-2517
[2] Council Directive 96/61/EC of 24 September 1996 concerning integrated pollution prevention and control (IPPC)
[3] M. Michniewi cz i inni, Najlepsze Dostępne Techniki (BAT). Wytyczne dla branży celulozowo-papierniczej, Raport sfinansowany ze środków Narodowego Funduszu Ochrony Środowiska i Gospodarki Wodnej na zamówienie Ministra Środowiska, Ministerstwo Środowiska, Warsaw, August 2005
[4] PN-EN 61439-1:2011 Low-voltage switchgear and controlgear assemblies – Part 1: General rules
[5] Technical Application Papers No.11 Guidelines to the construction of a low-voltage assembly complying with the Standards IEC 61439 Part 1 and Part 2, ABB SACE A division of ABB S.p.A. L.V. Breakers, 06/2016
[6] T . Kądziołka, M. Kryś , A. Wier tek, Analysis of power loss and temperature distribution in low voltage switchgear – coupled analysis EM + CFD in ANSYS software, Przegląd Elektrotechniczny, 94 (2018), No 1
[7] T . Daszc z yński, M. Szewczyk, A. Smolarczyk, S. Stoczko , Zadanie 1. – Wykonanie przeglądu dokumentacji technicznej i porównanie cech z urządzeniami istniejącymi, Raport z pracy badawczej, Nr projektu POIR.02.03.02-10-0023/19, Warsaw, March 2020
[8] A. Smolarczyk, K. Kurek, R. Kowal i k, M. Januszews k i , M. Szewczyk, T. Daszczyński , Zadanie 5 – Wykonanie testów układów automatyki zabezpieczeniowej, Raport z pracy badawczej, Nr projektu POIR.02.03.02-10-0023/19, Warsaw, June 2020
[9] M. Łukiewi cz, Transformator z uzwojeniami w konfiguracji Vv firmy Elhand, Elektrosystemy, September 2006
[10] OMICRON elec t ronic s, https://www.omicronenergy.com/en/ products/cmc-256plus/, on-line access: 2.06.2020

Authors: dr hab. inż. Adam Smolarczyk, Warsaw University of Technology, Electrical Power Engineering Institute, ul. Koszykowa 75, 00-662 Warsaw, Poland, E-mail: adam.smolarczyk @ien.pw.edu.pl, dr inż. Tadeusz Daszczyński, Warsaw University of Technology, Electrical Power Engineering Institute, ul. Koszykowa 75, 00-662 Warsaw, Poland, E-mail: tadeusz. daszczyński@ien.pw.edu.pl, Sławomir Fiszer, Elektroteam Sp. z o.o., ul. Brzozowa 8A, 97-400 Bełchatów, Poland, E-mail: s.fiszer@elektroteam.com.pl

Source & Publisher Item Identifier: PRZEGLĄD ELEKTROTECHNICZNY, ISSN 0033-2097, R. 96 NR 11/2020. doi:10.15199/48.2020.11.38

Harmonic Measurement Data Evaluation

Published by Electrotek Concepts, Inc., PQSoft Case Study: Harmonic Measurement Data Evaluation, Document ID: PQS1001, Date: March 15, 2010.

Abstract: Utility power system harmonic problems can often be solved using a comprehensive approach including site surveys, harmonic measurements, and computer simulations.

This case study presents a harmonic data analysis for a utility 12.47kV substation monitoring location for a two-week period. The analysis included trends of the rms voltage and current and statistical summaries of the voltage and current distortion values. The results of the analysis showed that the harmonic distortion levels were below the IEEE Std. 519 voltage limits.

INTRODUCTION

A harmonic measurement analysis case study was completed for a 12.47kV utility substation bus. The two-week monitoring period was from May 17, 2009 thru June 1, 2009. The power quality instrument used to complete the harmonic measurements was the Dranetz-BMI Encore SeriesTM. The instrument samples voltage at 256 points-per-cycle, current at 128 point-per-cycle, and follows the IEC 61000-4-3 method for characterizing harmonic measurement data. This involves analysis of continuous 200msec samples and storing aggregated 10-minute minimum, average, and maximum trend data. The measurement and statistical analysis was completed using the PQView® program (www.pqview.com).

MEASUREMENT RESULTS

Figure 1 shows the measured rms voltage regulation trend on the 12.47kV substation bus during the two-week period. Various pole-mounted distribution feeder capacitor banks (e.g., 600 kVAr) are switched on-and-off each day using time clock controls in an attempt to maintain a relatively constant voltage. Statistical analysis of the measurement data yields a minimum rms voltage of 12.27kV, an average voltage of 13.02kV, and a maximum voltage of 13.56kV. In addition, the CP95 value was 13.39kV. CP95 refers to the cumulative probability, 95th percentile of a value.

Figure 1 – Measured Substation Bus Voltage Trend

Figure 2 shows the corresponding measured voltage distortion trend and histogram during the two-week period. Statistical analysis of the measurement data yields a minimum distortion of 1.13%, an average distortion of 1.73%, and a maximum distortion 2.83%. The CP95 value was 2.30%.

Figure 2 – Measured Voltage Distortion Trend and Histogram

Figure 3 shows the corresponding rms current trend. Statistical analysis yields a minimum current of 378A, an average current of 690A, and a maximum current of 1165A. CP95 value was 1014A.

Figure 3 – Measured Substation Current Trend

Figure 4 shows the current distortion trend. Statistical analysis yields a minimum distortion of 4.39%, an average distortion of 7.89%, and a maximum distortion 14.24%. CP95 value was 11.73%.

Figure 4 – Measured Current Distortion Trend

Figure 5 shows the corresponding statistical summary of total harmonic voltage distortion and number of individual harmonics. The analysis shows that the predominate harmonics for the measured substation bus voltages were the 3rd, 5th, 7th, and 9th. The measured values were below the IEEE Std. 519 voltage distortion limits, which are 5% THD and 3% for any individual harmonic.

Figure 5 – Measured Statistical Summary of Voltage Distortion and Harmonics

Figure 6 shows the corresponding statistical summary of total harmonic current distortion and number of individual harmonics. The base current for the statistics summary was 1082A.

Figure 6 – Measured Statistical Summary of Current Distortion and Harmonics

Figure 7 shows one sample calculated harmonic current waveform from the measured harmonic spectrum data. The waveform was created using an inverse DFT with 256 points per cycle. The fundamental frequency current value was 469A, the rms current value was 472A, and the current distortion was 11.8%.

Figure 7 – Example Calculated Substation Current Waveform
SUMMARY

This case study presents a harmonic data analysis for a 12.47kV substation monitoring location for a two-week period. The analysis included trends of the rms voltage and current and statistical summaries of the voltage and current distortion values. The results of the analysis showed that the harmonic distortion levels were below the IEEE Std. 519 voltage limits.

REFERENCES

1. Power System Harmonics, IEEE Tutorial Course, 84 EH0221-2-PWR, 1984.
2. IEEE Recommended Practice for Monitoring Electric Power Quality,” IEEE Std. 1159-1995, IEEE, October 1995, ISBN: 1-55937-549-3.
3. IEEE Recommended Practices and Requirements for Harmonic Control in Electrical Power Systems, IEEE Std. 519-1992, IEEE, ISBN: 1-5593-7239-7.

RELATED STANDARDS
IEEE Std. 519-1992
IEEE Std. 1159-1995

GLOSSARY AND ACRONYMS
ASD: Adjustable-Speed Drive
CF: Crest Factor
DPF: Displacement Power Factor
PF: Power Factor
PWM: Pulse Width Modulation
THD: Total Harmonic Distortion
TPF: True Power Factor

New Idea in Power Networks Development. Selected Problems

Published by Bożena MATUSIAK, Anna PAMUŁA, Jerzy S. ZIELIŃSKI,
Katedra Informatyki Uniwersytetu Łódzkiego

Abstract. Smart grids and thereafter open grids development implies changes in organizations, technical equipment Power Networks (especially ICT). These new grids influence also on organization and operation of the energy markets, for which it is very important to know correct forecast (mostly short-term) of an electrical energy demand. Excessive requirements concerning accuracy of the forecasting results implied new tools of informatics, especially artificial intelligence application.

Streszczenie. Sieci inteligentne, a potem sieci następnych generacji, spowodują zmiany w organizacji, wyposażeniu technicznym (szczególnie ICT) sieci elektroenergetycznych. Te nowe sieci wpływają także na organizację i działanie rynku energii wymagającego znajomości prawidłowej prognozy (głównie krótkoterminowej) zapotrzebowania na energię elektryczną. Wygórowane wymagania odnośnie do dokładności wyników prognozy spowodowało zastosowanie nowych narzędzi informatyki, głównie sztucznej inteligencji. (Nowa koncepcja rozwoju sieci elektroenergetycznych. Zagadnienia wybrane)

Keywords: smart grids, ICT, energy market, artificial intelligence.
Słowa kluczowe: sieci inteligentne, ICT, rynek energii, sztuczna inteligencja.

Introduction

Power Networks development is upon influence of Smart Grid idea. It does not exist one definition of the Smart Grid; in [30] one can find three definitions: from USA, Europe and China. In opinion of the authors the first one is most suitable for this paper and for that reason it will be citation from [30]:

– “It is self-healing (from power disturbance events).
– It enables active participation by consumers in demand response.
– It operates resiliently against both physical and cyber attacks.
– It provides quality power that meets 21st-century needs.
– It accommodates all generation and storage options.
– It enables new products, services and markets.
– It optimizes asset utilization and operating efficiency.”

The above definition of the Smart Grid determines following problems being considered in the paper:

– Basic problems in developing smart grids.
– Microgrids.
– Selected problems of an energy market in smart grids.
– Load- and price forecasting.

Basic problems in developing smart grids [35]

Implementation of the smart grids idea needs new transmission- and distribution grids, large capacity storing devices and number of measurement, monitoring and control devices.

Distribution smart grids additionally must implement Advanced Distribution Management System (ADMS), Advanced Metering Infrastructure (AMI) and Wide Area Monitoring (WAM). All the above mentioned systems imply new problems that are to be solved by Information Communication Technology (ICT).

For example in Germany exist nearly about 3.5·106 measurement points and number of data stored necessary for market information growth 2 TeraBytes/year. It is foreseen generation 22 GigaBytes/day/106 consumers. Storing such number of data is nonsense and data management require data inspection in real time to discover future disturbances.

When consider the smart grids development it is necessary take into account following challenges:

– Dynamic external environment unable exact determination of work completion.
– Replacing existing systems with new offering more functionality may be not accepted option from the operational – as well as economical point of view.
– Implementation projects are to be accepted by all partners.
– Dynamic external environment unable exact determination of work completion.
– Replacing existing systems with new offering more functionality may be not accepted option from the operational – as well as economical point of view.
– Implementation projects are to be accepted by all partners.

All the above mentioned influence on prediction time of smart grid implementation which seems be far away. Contemporary practice is development step-by step solving separate projects and implementing them.

Microgrids [24]

One of assumptions in smart grids is necessity utilization of Renewable Energy Sources (RES), what implies intrusion in the grid Dispersed Generation (DG) and Dispersed Storage (DS). Operation of the grid with great number of singular DG is difficult and much more easy is to operate a group of DG – Microgrid.

Microgrids it is interconnection of small modular generation to Low- or Medium- voltage distribution systems. Microgrids can be connected to the main power network or be operated islanded, in a coordinated, controlled way [13]

Intrusion of Microgrid into the distribution network (in future smart distribution grid) needs creation of Active Distribution Network (ADN) passing following stages [24]:

– remote monitoring and control of DG and RES,
– determination of great number of DG and RES management,
– full active power management together with real time communication and remote control.

ADN operation implies necessity of application one of two different strategy: microgrids or virtual consumers. Concept of virtual consumer (virtual energy market) is adaptation of a model similar to information and business ability of Internet. Electrical energy bought from conventional generators, RES or storage devices, according to demand is delivered to agreed nodes. The system would use new ICT technologies as well as advance power electronics and storing devices.

Diversity of RES and storage devices as well as architecture and collaboration with power system implies necessity to define control strategy in operation. “Building Network “ strategy emulate “vacillatory source” in islanded network . DER unit realizing this strategy controls voltage in the connection with the system node setting up the system frequency.

Power and energy management strategy is very important in islanded microgrid and it is more critical than in power system because of specific characteristics of the microgrid.

New ICT needs for an energy market in smart grids

Some countries in Europe such as Italy, UK, Germany and Spain have been largely implemented in the AMI and developed new business processes in the energy market also taking into account the distributed generation and distributed energy sources, including renewable energy. (the results of recent European projects such as EUDEEP, FENIX, MORE MICROGRIDS, SEESGEN-ICT ).

The development of distribution networks in the direction of smart grids in Poland needs investment in infrastructure and ICT tools Some software and existing applications which are already adapted to the new needs of the energy market and intelligent networks in Europe will be summarized as follows in several main groups:

1. Energy flow calculation and market integration tools.
2. Power system analysis tools.
3. Customer portfolio and levels simulation tools.
4. Simulation and optimization tools for DR, DG and energy storage operation.
5. Forecasting and information systems tools. In table 1 there are some ICT samples (you can see and compare: http: //www.dconnolly.net/tools.html)

Table 1. Some samples:

Name of toolShort Characteristic
Ad 1: Wilmar planning toolA Strategic planning tool for analyzing the integration of renewable power technologies to be applied by system operators, power producers, potential investors in renewable technologies and energy authorities. The model optimizes power markets based on a description of generation, demand and transmission between defined model regions and derives electricity market prices from marginal system operation costs. The model is a stochastic linear programming model with wind power production as the stochastic input parameter. The model optimizes unit commitment taking into account trading activities of different actors on different energy markets. As a result the simulated output by different production forms, marginal price on each region, and others.
Ad 1: EMPS (multi-area power market simulatorThe EMPS model is a stochastic model designed for long-term optimization and
simulation of hydro-thermal power system operation. It allows the simulation
of large hydro systems with a relatively high degree of detail.. The EMPS model is widely used in the Nordic countries for price forecasting. Large producers can directly employ EMPS in their scheduling decisions.
Also thermal plants can be included. The time step is one week and planning horizon is up to several years.
Ad2: Siemens PSSE – Transmission System Analysis and PlanningPSSE is an integrated, interactive program for simulating, analyzing, and optimizing power system performance. It provides the user methods in many technical areas, including:
Power Flow, Optimal Power Flow, Balanced or Unbalanced Fault Analysis, Dynamic Simulation, Extended Term Dynamic Simulation, Open Access and Pricing, Transfer Limit Analysis and others
Ad 2: Powerworld SimulatorPowerWorld Simulator is an interactive power systems simulation package
designed to simulate high voltage power systems operation on a time frame ranging from several minutes to several days. Potential applications: Transmission Planning, Power Marketing, Simulation of Electricity Markets, Operator Training to improve operators’ knowledge of the system and response to unexpected events, Real-Time System Monitoring, Planning and Operations
Ad3: UPV FlexmodThe tool can calculate the available load reduction and the following payback
peak as a function of time when certain load control strategy is used (such as load reduction during morning peak, with allowed temperature drop of 1 °C). The results are specific to certain customer. Each customer is modeled separately.
Ad 3: DER-CAMDER-CAM ((Distributed Energy Resource Customer Adoption Model) is an economic model of customer DER adoption implemented in the General Algebraic Modeling System (GAMS) optimization software. This model has been in development at Berkeley Lab since 2000. The objective of the model is to minimize the cost of operating on-site generation and combined heat and power (CHP) systems, either for individual customer sites or a micro grid
Ad4: FlexprofFlexprof has been developed at VTT for assessing the revenues of the aggregation of demand flexibility, integrated with RES in the electricity market.
Flexprof tries to simulate trading on the spot market, taking account the possibility of flexibility calls. The situation with and without flexibility can then be compared. It can dynamically allocate the flexibility calls based on market price forecasts. Flexibility allocation is done with linear programming,
and the final flexibility calls are obtained with stochastic programming.
Any time period can be used in the simulation. One year’s simulation with six
customer types takes about one hour. The model has so far been adapted to the English and German market.
Ad4: OffpeakOffpeak tool can be used for profitability assessment of DER aggregator business. Special attention has been paid to the services that DER can provide within the Great Britain power system. The heuristic-based tool can quickly estimate the profits of several years of operation using historical price data.
Ad5: PrevedoVentoPart of PrevedoEnergia, a tool for forecasting power output from variable
renewable energy sources for bidding on power market.
PrevedoSole predicts the power output for each PV device according to the
provincial solar radiation forecast. The individual outputs are then aggregated for
each of seven market zones before bidding as a zonal whole schedule.
Ad5: Inter-Regional Electric Market Model (IREMM)The IREMM model is based on demand/supply precepts, and is not a “traditional” cost-recovery plus pricing model. IREMM provides a broad-based, comprehensive view of competitive electric power markets:
Forecasts market-clearing economy, energy prices, represents all buyers and
sellers within an interconnected system simultaneously, identifies economic energy transactions, analyzes the interaction of supply and demand in a competitive bulk power market, is not a cost-based, franchise area-specific pricing model.
Source: on the basis of Seesgen-ICT internal materials: Jussi Ikäheimo VTT (Finland), 2009
Load forecasting

Energy Market operation needs knowledge on demand of electricity and prices in different intervals of time what implies necessity of application of load- and price forecasting tools. Historically the first method applied time series-based methods, but in 20th century the Artificial Intelligence (AI) tools dominate [34]

One of the first well reported application of Expert System (ES) in Short Term Load Forecasting (STLF) is the paper [29] written by S.Rahman and R.Bhatnagar. From that time number of papers presenting application of ES, Artificial Neural Networks (ANN), Fuzzy Logic (FL) and Hybrid Systems (HS) combining no less than one of AI tools with another models is growing. As an illustration of contemporary state of the art we did review papers printed in the three years of the IEEE Trans. on PWRS (2008, 2009, 2010 – Feb) with following results:

– Short-term load forecasting [1,2,3,4,6,19,20,26,29,31]
– Forecasting another power system problems: [8,9,11,12,28,33].

Taking into account tools used in these papers we van find: Expert Systems, Artificial Neural Networks, Fuzzy Logic, Wavelet Transform, Models, Statistics, Evolutionary Algorithms, Hybrid Systems. It is worth of mention that dominate application of hybrid system where we can find PSO (Particle Swarm Optimization) Algorithm in Hybrid System with Wavelet transform and Artificial Neural Network [4].

Final Remarks

Smart Grids – new idea in electric power need for designing, construction and operation new tools, devices, services and quite new market integration tools. Not all of them are mature what opens necessity of further researches applying new ICT solutions.

REFERENCES

Abbreviations: PE – IEEE Power & Energy; PWRS – IEEE Trans. on Power Systems

[1] Amjad y N. , Ke yn ia F.: Day-Ahead Price Forecasting of Electricity Markets by Mutual Information Technique and Cascaded Neuro-Evolutionary Algorithm. PWRS, Feb. 09, 306-318.
[2] Areekul P., Senjyu T., Toyama H., Yona A.: A Hybrid ARIMA and Neural Network Model for Short-Term Model Price Forecasting in Deregulated Market. PWRS, Jan.10, 524-530.
[3] Bas hi r Z.A. , El -Hawar y m.M.E.: Applying Wavelets to Short-Term Load Forecasting Using PSO-Based Neural Networks. PWRS, Feb. 09, 20-27.
[4] Bessa R.J. , Miranda V., Gama J. : Applying Wavelets to Short-Term Load Forecasting Using PSO-Based Neural Networks. PWRS, Nov. 09, 1657-1666.
[5] Chakrabar t i S., Kyr iakides E., Bi T . , Cai D.: Ter z i ja V.: Measurements Get Together. PE, vol. 7, No.1,4149.
[6] Chen Y. , Luh P.B., Guan C., Zhao Y. , Miche l D., Coolbeth M.A. , Friedland P.B., Rourke S.J . : Short-
Term Load Forecasting Similar Day-Based Wavelet Neural Networks. PWRS, Feb. 10. 322-330.
[7] Dickerman L . , Har r i son J. : A new Car, a New Grid. PE, vol. 8, No. 2, 55-61.
[8] Do Couto Fil ho M.B. , Stacchini de Souza J.C: Forecasting-Aided State Estimation – Part I: Panorama. Nov. 09,1667-1677.
[9] Do Couto Filho M. B., Stacchini de Souza J .C.: Forecasting-Aided State Estimation – Part II: Implementation. PWRS, Nov. 09,1678-1685.
[10] Dr iesen J . , Kat i rei F. : Design for Distributed Energy Resources. PE vol. 7, No. 3, 30-39.
[11] El ias C.N., Hatz ia rgyr iou N.D. : An Annual Midterm Energy Forecasting Using Fuzzy Logic. PWRS, Feb. 09, 469-478.
[12] Gaj b h iye R.K. , Nai k D. , Dambare S., Soman S. A.: An Expert System Approach for Multi-Year Short-Term Transmission System Planning. An Indian Experience. PWRS, Feb. 08, 226-237.
[13] Hat z i argyriou N., Jagoda G., Pamuła A. , Ziel i ński J.S.: Microgr ids. Some Remarks on Polish Experiences in DER Intrusion into Distribution Grids. Large Scale Integration of RES and DG. 25-26 September 2008, Warsaw.
[14] Horowi tz S.H., Phadke A.G. , Renz B. A.: The Future of Power Transmission. PE, vol. 8, No.2, 34-40.
[15] ICT for a Low Carbon Economy. Smart electricity Distribution Networks. EU Commission, Directoriate-General Information Society and Media: ICT for Sustainable Growth Unit. July 2009.
[16] Kat irarei F ., I ravania R., Hatziargyr iou N., Dimeas A. : Microgrids Management, Controls and Operation Aspects of Microgrids. PE, vo. 7, No.3, 54-65.
[17] Ki r s chen D. , Bouf fard F. : Keep the Lights On and the Information Flowing. PE, vol. 7, No.1, 55-60.
[18] Kroposky B., Lasseter R. , Ise T., Morozumi S., Papathanassiou S. , Hatziargyr iou N. : Making Microgrids Work. PE, vol. 7, No. 3, 41-53.
[19] L i vel y M.B. : The Wolf in Pricing, PE, vol. 7, No.1, 61-69.
[20] Mao H. , Zeng X. -J. , Leng G., Zhai Y. -J. , Keane J.A. : Short-Term and Mid-term Load Forecasting Using Bilevel Optimisation Model. PWRS, May 09, 1080-1090.
[21] Marnay Ch. , Asano H., Papathanassiou S., Strbac G. : Policymaking for Microgrids. Economic and Regulatory Issues of Microgrid Implementation. PE, vol. 7, No. 3, 66-77.
[22] Matusiak B. , Pamuła A., Ziel i ński J.S. : Technologiczne i inne bariery dla wdrażania OZE i tworzenia nowych modeli biznesowych na krajowym rynku energii. Rynek Energii nr.4, sierpień 2010, 31-35.
[23] Nou rai A. , Kearns D.: Batteries Included. PE, vol.8, No.2, 49- 54.
[24] Pamu ła A. , Ziel i ń s k i J .S. : Sterowanie i systemy informatyczne w mikrosieciach.. Rynek Energii, I(III), luty 2009, 63-69.
[25] Ph i l ips A. : Staying in Shape. PE vol. 8 No.2, 27-33.
[26] Pindoriya N.M., Singh S.N., Singh S.K.: An Adaptive Wavelet Neural Network-Based Energy Price Forecasting in Electricity Market. PWRS, Aug. 08,1423-1432.
[27] Piwko R., Mi l ler R., Gi rard R.T.G., MacDowe ll J ., Clar k K. , Murdo ch A. : Generator Fault Tolerance and Grid Codes. PE vol. 8, No. 2, 19-26.
[28] Rab iee A. , Shayanfar H.A. , Amjady N.: Reactive Power Pricing. Problems & a Proposal for a Competitive Market. PE, vol.7, No.1, 18-32.
[29] Rahman S. , Bh atn agar R.: An expert system based algorithm for short term load forecast. PWRS, vol. 3, No. 2 1988, 392-399.
[30] Santacana E., Rackl i f f e G., Tang L. , Feng X.: Getting Smart. PE, vol. 8, No.2, 41-48.
[31] Yun Z. , Quan Z., Cai xin S., Shaolan L., Yuming L. , Ya ng S. : RBF Neural Network and ANFIS-Based Short-Term Load Forecasting Approach in Real-Time Price Environment. PWRS, Aug. 08, 853-858.
[32] Venkataramanan G., Marnay Ch. : A Larger Role for Microgrids. PE, vo.7, No.3, 78-82.
[33] Zhao J.H., Dong Z.Y. , Xu Z ., Wong K.P. : A Statistical Approach for Interval Forecasting of the Electricity Price. PWRS, May 08, 267-276.
[34] Ziel i ń s k i J .S. : Artificial Intelligence in power system application. XXX Międzynarodowa Konferencja z Podstaw Elektrotechniki I Teorii Obwodów IC-SPETO 2007, Gliwice- Ustroń 23-26.05.2007, 245-246.
[35] Ziel i ń s k i J .S. : Rola teleinformatyki w środowisku sieci inteligentnych. Rynek Energii nr.1, luty 2010, 16-19.

dr Bożena E. Matusiak, Uniwersytet Łódzki, Wydział Zarządzania, Katedra Informatyki. E-mail: bmatusiak@wzmail.uni.lodz.pl;
dr Anna Pamuła, Uniwersytet Łódzki, Wydział Zarządzania, Katedra Informatyki, E-mail: apamula@wzmail.uni.lodz.pl.
prof. dr hab. Inż. Jerzy S. Zieliński kierownik Katedry Informatyki na Wydziale Zarządzania Uniwersytetu Łódzkiego, uczestnik projektów europejskich: EU DEEP, SYNERGY+, MORE MICROGRIDS, SEESGEN-ICT. E-mail: jzielinski@wzmail.uni.lodz.pl

Source & Publisher Item Identifier: PRZEGLĄD ELEKTROTECHNICZNY (Electrical Review), ISSN 0033-2097, R. 87 NR 2/2011

A Short Review of Energy-Efficient Line-Start Motor Design

Published by Tine MARČIČ1, TECES, Research and Development Centre for Electric Machines (1)

Abstract. The paper provides an overview of design problems and contemporary research progress in the currently very interesting field of energy-efficient line-start motors. Discussed are problems related to induction motors (IMs), line-start synchronous reluctance motors and line-start interior permanent magnet synchronous motors (LSIPMSMs). Emphasis is given on small rated power motors, where the LSIPMSM presents the most interesting alternative for replacing IMs widely used in low-cost single-speed drives with ventilator fans, pumps and compressors.

Streszczenie. Artykuł daje przegląd problemów projektowania i postęp we wspólczesnych badaniach w interesującym obszarze wydajności energetycznej silników bezpośredni włączanych. Dyskutowane są problemy związane z silnikami indukcyjnymi, silnikami synchronicznymi reluktancyjnymi z bezpośrednim włączaniem i takie same z magnesem trwałym. Nacisk został położony silniki małej mocy, które skutecznie zastępują silniki indukcyjne w niskokosztowych napędach w wentylatorach, pompach i kompresorach. (Krótki przegląd efektywności energetycznej silników o bezpośrednim włączaniu)

Keywords: computer aided design, induction motors, squirrel cage motors, synchronous motors.
Słowa kluczowe: projektowanie wspomagane komputerowo, silniki indukcyjne, silniki klatkowe, silniki synchroniczne

Introduction

Nowadays, a large share of electric energy is converted into useless heat by electric drives worldwide. A large portion of these electric drives is represented by single-speed applications with ventilator fans, pumps and compressors. In such drives the electric motors are mostly started and fed directly from line, i.e. without the usage of any power electronics. Therefore, the used line-starting motors must fulfil one fundamental requirement – they have to be able to start from standstill and accelerate the complete drive to the rated speed when they are fed from a constant amplitude and constant frequency voltage source, i.e. the so called line-starting capability. And considering the elevated environmental conscience and global market trends, the used motors have to exhibit the highest possible efficiency also. However, these two requirements are quite contradictory when the actual motor design is considered. Therefore the line-start motor design process is connected with adequately addressing many design compromises.

This paper is devoted to providing an overview of the contemporary progress from available literature and own research results in the currently very interesting field of energy-efficient line-start motors.

Overview of line-start motor topologies

The line-start brushless motor family includes induction motors (IMs) [1], line-start synchronous reluctance motors (LSSRMs) [2] and line-start interior permanent magnet synchronous motors (LSIPMSMs) [3]. Their principal rotor structures are depicted in Fig. 1, whereas the stator structures are the same [4, 5].

Motor designers utilize different designs of the squirrel-cage (SC) in all previously mentioned line-start motor types. The SC provides asynchronous starting capability or the so called “line-starting capability” and damping of dynamic oscillations at fast load changes also. In relation to IMs (also called asynchronous motors), the SC is usually made of electrically conducting bars which are embedded in slots of the rotor’s iron core and connected on both ends with cage-end rings. In large motors, the SC can be die-casted or fabricated [6] by using different materials. But in large volume production of small rated power motors the SC is mostly die-casted from aluminium and its alloys. The IM performance both in transient- and steady-state heavily depends on the SC and rotor slot design [7].

The SC in rotors of LSSRMs is usually constructed as electrically conducting material within the LSSRM’s magnetic flux barriers (FBs) [2], which are accountable for the main torque producing component of a LSSRM in its steady-state synchronous operating region. Furthermore, the LSIPMSM has permanent magnets (PMs) inserted in FBs, thus different authors have presented many different SC, FB and PM arrangements within rotors of LSIPMSMs [8-27] along with their design methods. The evolution of these rotor designs has been in line with the evolution of PM materials and their contemporary price and availability [28]. However, the one mostly used SC design in LSIPMSMs is still the one similar to IMs and the nowadays mostly used PM material in LSIPMSMs is of Nd-Fe-B type. The PMs and FBs are accountable for the torque producing components of a LSIPMSM in its steady-state synchronous operating region. Due to the hybrid nature of LSSRMs and LSIPMSMs, the motor designer has to account for all the different torque producing components in both the asynchronous and the synchronous operation region.

Fig.1. Principal rotor cross-sections of induction motors (IMs), line-start synchronous reluctance motors (LSSRMs) and line-start interior permanent magnet synchronous motors (LSIPMSMs)
Candidates for energy-efficient line-start motors

IMs have been traditionally used in all kinds of applications, mainly due to their low price and robust construction. However, especially for small rated power IMs, their relatively small efficiency and power factor make them inappropriate for markets with strict regulations regarding energy efficiency. Some previous studies were focused on improving the IM efficiency by using expensive cage materials (copper alloys) also in small-sized IMs [1, 29]. The LSSRMs present an alternative only for larger machines because a large portion of rotor material has to be allocated to FBs in order to achieve sufficient torque capability. Thus, LSIPMSMs with buried PMs bellow the SC are currently identified as the most promising design for energy-efficient small rated power applications [30]. Fig. 2 presents a comparison between the measured characteristics of efficiency and power factor and their product for a 1.1 kW four-pole three-phase LSIPMSM and IM with SCs made from aluminium [3]. The efficiency characteristics can be directly compared to measured characteristics of the same rated power IM with the SC made from copper available in [1]. From that comparison it can be seen that LSIPMSMs offer much higher efficiency increase. But on the other hand, for large machines the efficiency increase of LSIPMSMs in comparison to IMs is far from substantial [5].

Fig.2. Comparison of measured efficiency (EFF) and power factor (PF) characteristics of a 1.1 kW three-phase four-pole LSIPMSM and IM, measured at equal voltage 380 V / 50 Hz
LSIPMSM torque components

The electromagnetic torque te equation (1) which is part of a LSIPMSM dynamic model written in the d-q reference frame [31] neatly depicts the torque producing components in all line-start motors. In Eq. (1) the subscripts d and q denote variables in the d- and q-axis, respectively; i denotes stator winding currents, ik denotes SC currents, Ls are stator self-inductances, Lm are mutual inductances, Ψm is the length of the PM flux linkage vector, and p is the number of pole pairs.

.

The (asynchronous) cage torque (due to the presence of a SC) influences the LSIPMSM performance in any operation state, where the slip differs from 0. Thus, the cage torque enables line-starting performance and damping of dynamic load oscillations. Apart from the stator winding design, the cage torque depends mainly on the SC design and material.

The synchronous torque components which are represented by the reluctance torque (due to the presence of FBs) and the PM torque (due to the presence of PMs) influence the LSIPMSM performance in any operation state, where the slip differs from 1. In the synchronous operation region (where the slip equals 0) they represent useful torque components. Contrarily, during the line-starting transient they represent braking torques. Thus, they degrade the total torque which should accelerate the LSIPMSM drive up to synchronism. The reluctance torque depends mainly on the design of FBs, which also have to accommodate the used PM segments. The PM torque depends mainly on the placement, dimensions and type of PM material.

For these reasons, the LSIPMSM’s static torque-slip characteristic in the asynchronous operation region is generally lower than the static torque-slip characteristic of a pure IM with the equal SC design, materials, stator and rotor slots geometry and stator winding design. Fig. 3 shows the impact of the aforementioned PM braking torque and the braking reluctance torque on the LSIPMSM’s static torque-slip characteristic [3]. The comparison of measured torque-slip characteristics of a three-phase IM, LSIPMSM and equal LSIPMSM design without PMs in the rotor which actually represents a LSSRM (all with equal SC design) is presented. The difference between the IM torque-slip curve and the LSSRM torque-slip curve at certain slip points represents the reluctance braking torque. And, the difference between the LSSRM torque-slip curve and the LSIPMSM torque-slip curve at certain slip points represents the PM braking torque in the asynchronous operation region.

Fig.3. Measured static torque-slip curves of a three-phase LSIPMSM, equal LSIPMSM design without PMs (i.e. the LSSRM), and IM with equal SC design, at equal voltage
Design process

As it can be noticed from the previous sections, the design problems of LSSRMs and LSIPMSMs are quite similar and are closely related to IM design problems. And usually the main aim of a new LSSRM or a LSIPMSM design is to replace an existing IM, therefore the new motor has to comply with the following two requirements:

– it has to exhibit line-starting and synchronization capability [2, 16, 20, 32];

– in comparison to the existing IM, it has to exhibit a higher (or at least an equal) torque per unit (stator) current density value and a higher efficiency value in its steady-state synchronous operation region [3].

Different authors have used differently complex approaches in coping with design problems [2, 3, 5, 8-27, 32-35]. Strongly coupled finite element (FE) models [5, 21, 23, 33] are in the author’s opinion computationally too complex to be regularly used by motor designers, thus the design procedure depicted in Fig. 4 has been found to be very useful. It has been founded as a hybrid based on preceding knowledge and experience on design, dynamic modelling and analysis of SC IMs [36-39], (cageless) synchronous reluctance motors [40-44] and (cageless) synchronous PM motors [45-49]. The procedure includes employment of the power balance method based on results from time-stepped FE analyses in the analysis of synchronous performance and employment of lumped parameter dynamic models in the analysis of line-starting performance. The power balance method is employed because FE analyses provide a very detailed image of the geometry and material dependant distribution of magnetic field in the machine region. Thus, the FB design, placement and energy-product of PM material, and their effect on iron core saturation and motor parameters are accounted for in sufficient detail. The employed methods and procedures were described and experimentally validated in [3].

Fig.4. LSIPMSMs or LSSRMs design procedure
Design considerations

Along with all the aforementioned, the following list presents further important LSIPMSM design aspects, which should be kept in mind by the motor designers in order to achieve target motor performance where a lot of compromises are to be made [3].

– The motors’ line-starting transient depends on the supply voltage [3, 35] and frequency, the drive inertia [23, 33], the characteristic of the mechanical load, and also the starting position of the rotor [23]. The initial rotor position of the LSIPMSM influences its responding current and speed line-starting transient. Its effect is much expressed when the motor is started without any load.

– The compromise between a LSIPMSM’s adequate line-starting performance in the asynchronous operating region and efficiency in the LSIPMSM’s synchronous operating region is connected to the stator winding’s number of turns. Therefore, the number of turns often has to be adopted in accordance with the target load characteristic, especially when rigorous starting conditions are expected.

– The SC material plays a vital role in the electromechanical performance of line-start motors. A higher cage resistance causes that the motor exhibits a higher starting torque value. But on the other hand, the slope of the torque-slip curve near the synchronous speed is lowered. In relation to IMs, this produces an increase of losses and motor temperature and thus the IM efficiency in steady-state is degraded. However, the impact of SC material on LSSRM and LSIPMSM performance may be more severe. When e.g. a LSIPMSM is line – started, the SC should accelerate the complete LSIPMSM drive up to a certain speed and if the acceleration is sufficient the rotor should be pulled into synchronism. Thus, the LSIPMSM’s “pull-in” transient into synchronism depends on the slope of the static torque – slip characteristic of the LSIPMSM near the synchronous speed, and consequently the LSIPMSM’s starting and synchronization capability depends on the used SC material. Results from different studies which can be related to the cage resistance are available in [23, 26, 34, 35].

Economic considerations

LSIPMSMs with buried PMs bellow the SC are currently identified as the most promising design for energy-efficient small rated power applications. However, motor manufacturers tend to be very rigid when it comes to the manufacture of new motor designs because the manufacturing tools (lamination punching, SC die-casting, winding tools, …) always present a substantial part of the motor manufacturing cost. Therefore, especially in high-volume production of small rated power motors (e.g. ventilator, pump and compressor motors) the manufacturers tend to use existing tools until they are worn out or there is a change on the demand side. Enforcement of stricter policies for motor efficiency (like the EU directive 2005/32/EC) is going to force the manufacturers to consider new motor designs as well.

The manner in which LSIPMSM rotors are manufactured (die-casting of SCs at relatively high temperatures, which are higher than the Curie temperatures of PMs) makes the placement of magnetized magnetic segments in the rotor’s FBs impossible before the SC is die-casted. On the other hand, pulse magnetization of the whole multi-pole rotor with buried PMs presents also a problem, because the induced currents of the SC limit the depth of magnetic field penetration in the rotor area and the magnetization homogeneity of magnetic segments as well [50, 51]. Therefore, the easiest way to manufacture a LSIPMSM rotor is by simple purchase and insertion of pre-magnetized PM segments, which may be quite costly.

An economic assessment and overview of the LSIPMSM manufacturing related cost increase and on the other hand the potential electric energy savings by using LSIPMSMs in contrast to IMs is discussed in [30].

But on the other hand and as indicated before, the motor designers can take advantage of both of the synchronous torque components (i.e. a combination of the PM torque and the reluctance torque as well) where by designing higher rotor saliency, less PM material or cheaper PM material can be used in order to achieve sufficient target synchronous performance.

Conclusion

This work presented an integral overview of the line-start motor design related problems. Discussed and pointed out were LSIPMSM design aspects which were related to IM and LSSRM design as well. These comprise the motors’ construction, stator winding design and the arrangement of SC, PMs and FBs in the rotor; and the manufacturing related economic considerations also.

This work was supported in part by the Slovenian Research Agency, Project No. L2-1180.

REFERENCES

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Author: Dr. Tine Marčič, TECES, Research and Development Centre for Electric Machines, Pobreška cesta 20, SI-2000 Maribor, Slovenia, E-mail: tine.marcic@teces.si.

Source & Publisher Item Identifier: PRZEGLĄD ELEKTROTECHNICZNY (Electrical Review), ISSN 0033-2097, R. 87 NR 3/2011

New Approach to Protecting Transformers Against High Frequency Transients – Wind Turbine Case Study

Published by Dariusz SMUGAŁA(1), Wojciech PIASECKI(1), Magdalena OSTROGÓRSKA(1), Marek FLORKOWSKI(1), Marek FULCZYK(2), Ole GRANHAUG(3),
ABB Sp.z.o.o., Corporate Research Center, Cracow, Poland (1), ABB Oy, Medium Voltage Products, Vaasa, Finland (2), ABB AS, Medium Voltage Products, Skien, Norway (3)

Abstract. Novel protection method of wind turbine transformers against high frequency transients occurring during switchgear operation is described in this paper. Presented results are continuation of research on Very Fast Transients mitigation methods previously published in literature [8]. Principles of novel suppressing device parameters optimization for windmill transformers are also included. ATP-EMTP simulations results for wind farm application were verified by full scale functional tests.

Streszczenie. W artykule przedstawiono nową metodę ochrony transformatorów turbin wiatrowych przed wysokoczęstotliwościowymi przepięciami mogącymi wystąpić w trakcie ich pracy. Przedstawione rezultaty są wynikiem kontynuacji wcześniejszych badań prowadzonych nad ochroną transformatorów dystrybucyjnych przed zakłóceniami o wysokokiej częstotliwości mogącymi wystąpić w sieci SN [8, 9]. Wyniki symulacji przepięć oraz doboru parametrów urządzeń ochronnych, zostały zweryfikowane w trakcie testów funkcjonalnych. (Ochrona transformatorów turbin wiatrowych przed przepięciami wysokoczęstotliwościowymi).

Keywords: high frequency transients, transformers, protection, wind turbines
Słowa kluczowe: przepięcia wysokoczęstotliwościowe, transformatory rozdzielcze, metody ochrony, turbiny wiatrowe.

Introduction and problem definition

High Frequency (HF) transients influence on wind turbine transformers (especially dry type) has been observed within the research activities presented in this article. They were mainly focused on protection of transformers located at the windmill nacelle. This configuration type is mostly used in practice. Research results presented in this document are study continuation on Fast Transients (FTs) phenomenon overstressing the distribution transformer’s insulation system [2, 4]. Present activities were focused on transients occurrence during normal wind turbine operation [8]. Wind turbines, considering variable operation conditions e.g. wind strength and direction, power network conditions, service, need to be controlled through the breakers operating at relatively high frequency. Presently most of wind turbines, except of air insulated switch-disconnectors, are operated through the Vacuum Circuit Breakers (VCB). Current breaking operations under certain conditions may result in overvoltages generation. During switching operations, interrupters frequently installed into the switchgear (SWG) usually located at the windmill tower bottom, many high frequency transients are generated. In consequence of extremely high voltage steepness (du/dt) occurring e.g. during inductive (e.g. no load) current interruption [3, 5], the insulation system of wind turbine transformers may be overstressed and may lead to pre-mature aging of the insulation material. It may increase transformers failure rate [1, 5, 6]. The high frequency Transient Overvoltages (TOV) problem is dangerous to other connected equipment, e.g. cables and accessories [6]. Generated transients character and overvoltage level depends on wind farm topology and a breaker type. There are two topologies of power network wind turbines connections used in practice (Fig. 1):

– Oil-filled transformer placed at the tower bottom with short connection between the transformer and breaker with long (tens of meters) connection between the transformer and wind turbine with generator,

– Dry-type transformer located at the windmill nacelle with relatively long connection (usually 80÷100 m) with SWG located at the tower bottom.

High frequency transients are generated as a result of relatively short wind turbine cables capacitance (tens of nF, C1,C2 in Fig. 2 and Fig. 3) interaction with low value of dry-type transformer inductance (LT in Fig. 3).

Fig.1. Single wind turbine power network connections topologies

Single windmill power network diagram with dry-type transformer is presented in Fig. 2.

Fig.2. Exemplary wind turbine topology with dry type transformer located at the nacelle (BRK – breaker)

Wind turbine transformer TX1 is connected to the VCB through the cable having capacitance C2. SWG comprising VCB is connected through the cable of C1 capacitance, to the power network connecting point represented by transformer TX2.

Fig.3. Simplified single wind turbine ATP/EMTP power network model: CT, LT – transformer capacitance and inductance, C1, C2 – cable capacitance, LC2 – cable inductance

Additionally, the following factors in MV networks have an influence on the generated transients level:

– switching in/out power network by operating breaker,
– ground faults,
– pre-strikes and reignitions during switching,
– wave reflections if cable surge impedance do not match the transformer impedance.

The cable capacitance results in generated transients filtering but simultaneously in interaction with transformer inductance, can be a reason for overvoltages and HF transients. They are in particular related with:

a. Switching-off of an unloaded transformer with high no-load currents values. High natural circuit frequency resulting in fast Transient Recovery Voltage (TRV) build-up,
b. Switching-on a transformer to a high-capacitance line. In this case the input capacitance charging from the network capacitance is limited only in practice by the typically very low line resistance.

Moreover, the HF impedance mismatch between the transformer input and the line result in potential wave reflections, multiplying the overvoltage build-up at the transformer terminals.

The effect of reignition level, number of strikes or voltage steepness mainly depends on circuit parameters e.g. cable parameters, but also on circuit breaker type, transformer size and type, supplying voltage level, resonant frequency value etc.

The main problem in avoiding the VFTs is a low value of the surge impedance due to a low impedance of power cables. There are cases described in literature, of the transformer failures during current breaking (e.g. when the VCBs are used for operation [7]). It is supposed that the HF transients occurring during the switching are the most likely the cause for that.

The problem of very high du/dt is enhanced in the case of short connections to the surge source. Increasing the impedance of the surge source may be achieved by introducing a suppressing series element upstream the protected equipment [9].

Solution

There are several transients problem solutions existing in the market and described in the literature [8] e.g. surge capacitors, RC snubbers, series resistors or surge arresters, however all of them have some weak points. Surge arresters protect power network (e.g. transformers) against overvoltages, but does not provide sufficient protection against high voltage steepness occurring during VCB operation. Capacitors and RC snubbers with typical values of this capacitance (C=0.1μF ÷ 0.5μF) are combined with resistance (R=5Ω ÷ 25Ω) and connected upstream the protected device. However they generally provide sufficient protection but they are characterized by large weight, size and significant costs which typically limits its applicability. These features exclude this solution to locate into relatively small windmill tower space. A potentially applicable lowpass RC filter is not acceptable, due to power dissipation and significant voltage drop.

Therefore a special construction of a series impedance element was developed [5]. The device described in [10] consist of series impedance element, coupled with small (≈10nF) capacitance connected upstream the protected device (transformer). At 50/60Hz the impedance of the element is close to zero, at higher frequencies the impedance has a resistive character (close to cable connection impedance). For the windfarm application where cable connection between the breaker and transformer is present, the cable connection capacitance is utilized. Required capacitance is provided by capacitance of power cable (80÷100m) connection between the SWG and transformer.

Taking into consideration high short circuit current rating of SWG, the original air coil construction developed previously, was replaced by high frequency magnetic material.

Suppressing device concept

The limitations of the presently used mitigation methods led to the development of a new concept of high du/dt mitigation using an impedance choke connected in series [6]. Increasing the impedance of the surge source may be achieved by introducing a series, impedance element upstream the protected transformer, as shown in Fig. 4.

Fig.4. Series protective impedance element connected upstream the transformer: C – connection capacitance, Lo – TX2 transformer load inductance

The use of series filter as protecting device is a frequently used method, mostly as common mode chokes in various low voltage systems comprising power electronics.

The choke of appropriately designed frequency characteristic allows significantly increase the voltage wave-front rise time and minimize its influence on the equipment under normal, operating conditions. This means that the choke impedance at 50/60Hz frequency must be close to zero.

To eliminate risk of destroying protection device as result of short circuit current which may occur into the wind turbine power network, previous protective device was redesigned and improved. Concept based on parallel connection of inductor and resistance was replaced by combination of high frequency magnetic material located at the main conductor e.g. cable, coupled with connected in parallel resistance (Fig.5) .

Fig. 5. Basic diagram of VFTs suppression device solution with cores placed in series and single damping resistor connected in series with secondary winding

The VFTs suppressing device comprises two windings:

– primary winding made of large cross section conductor (e.g. cable placed within),
– secondary winding with optimized resistance value.

Inductive impedance can be realized by high permeability magnetic cores put on the main conductor e.g. cable. During normal operation when rated current flows through the conductor and for the short circuit current as well, core are saturated so the inductance is negligible.

Tab. 1 Dependence of the voltage escalation process during the inductive current disconnection on the current level [8]

.

Suppressing device self-resistance is equal to the main conductor resistance (several μΩ).

For the higher frequencies magnetic rings combination provides required inductance so resistance at the secondary winding provides necessary damping.

To optimize the choke parameters for various magnetic cores and resistance connected in parallel, ATP-EMTP simulations were performed. During analysis various cross-section value cores were simulated having material permeability value near to 30 000. It was experimentally verified that cores characterized by this permeability value provides the best saturation characteristic for most often used in practice wind turbine network topologies. Finally the tested chokes equipped with magnetic components having cross-section from 40 cm2 to 120 cm2 were characterized by summary inductance from 0,6 to 1,5 mH and were tested with transformer connected to the VCB through the specific Zcable cable connection impedance.

The simulations were performed in simplified circuit for the “worst case”. Single re-ignition during VCB operation were treated, for the simulation needs, as voltage step with the magnitude equal the highest system voltage amplitude (in this case 29 kV).

Figures (Fig.6÷Fig.7) presents selected choke current simulation results for various damping resistance values (R1÷R8) and various cores cross-sections. Choke current simultaneously represents choke core saturation occurrence.

Fig.6. Choke current for 40 cm2 core cross-section

The figure below presents the voltage at the transformer when the step voltage is applied for various resistor values.

Fig.7. Voltage at the transformer terminals simulated for 80 cm2 suppressing device core cross-section

Performed simulations indicated that, the best performance is achieved when the resistor value is close to cable wave impedance – Zcable.

Fig.8. Damping resistance value calculation; Voltage at the transformer terminals as a function of choke impedance

Damping resistance should be calculated to provide the best suppressing effectiveness. Voltage at the transformer terminal was taken as the criterion for proper resistance value calculation (Fig.8). Calculation results [8] confirmed the best performance of suppressing device is achieved for resistor value near to cable surge impedance.

The optimal protection provides combination of the choke with additional small capacitor. Especially for the case when the transformer is connected to the circuit-breaker through the short cable characterized by small cable surge impedance.

Functional tests results

New concept of VFT suppressing device functional tests were performed for typical single wind turbine circuit (Fig.1, Fig.2).

Tests stand comprised the following components:

– 630A rated current Ring Main Unit (RMU) equipped with 24 kV VCB,
– 3×80 m single phase, 25 mm2 cross-section cable with 33 mH/km inductance and 150 nF/km capacitance
– 630 kVA transformer – In Yy 24 kV / 0.24 kV
– 3x suppressing devices located between the RMU and transformer (Fig.4)

Tests were performed for various chokes configurations (various resistance and core cross-sections values):

a. base case, no suppression devices,
b. 40 cm2, 70 cm2 and 120 cm2 cores cross-section
c. Various dumping resistance values from 0.7xZcable to 6xZcable (alternatively no resistor connected) for each version of choke core cross-section

Base case – no active chokes connected
In the base case there were no active chokes connected into power network.

Fig.9. Base case, Voltage at R-S-T phases
Fig. 10. Base case, R – phase
Fig.11. Chokes with 70 cm2 core cross-section at T phase and damping resistors R5 at the secondary winding
Fig. 12. Chokes with 40 cm2 cores cross-section and damping resistors R8
Fig.13. Chokes with 40 cm2 cores cross-section and resistors R5 at the secondary winding, phase T
Fig.14. Chokes with 70 cm2 core cross-section in T phase and damping resistors R8
Fig.15. Chokes with 120 cm2 core cross-section at T phase – with resistors R5 at the secondary winding

The experimental results confirmed the applicability of the series-choke protection concept to mitigate high frequency and high du/dt transients.

In cases when the transformer internal capacitance is low, what corresponds to dry-type transformers case, additional small surge capacitor plays an important role in the transients suppression. It has to be pointed out, that the value of the capacitance used was more than an order of magnitude smaller, than typical the value of the typical snubber’s capacitor.

Conclusions

A new mitigation method against high du/dt overvoltage hazards in a form of a series-connected choke element was developed. It was demonstrated that the use of the choke significantly reduces voltage steepness and number of reignitions generated during transformer operated through the breakers. Additionally noticeable overvoltage reduction was observed.

The use of appropriately designed series choke device can:

• Limit the du/dt values at transformer terminals
• Limit transient overvoltage
• Eliminate wave reflections in cable and HF oscillations (when Zchoke = Zcable)
• Eliminate or reduce the number of re-ignitions (requires C in order to lower oscillation frequency

The problem of potential VFT-related hazard to transformer and other power equipment resulted from switching operations was demonstrated in a practical case. The number of pre-strikes during contact making was reduced and high frequency oscillations were practically eliminated.

High du/dt was over 2x reduced with the use of the chokes only. Further reduction was achieved when a small (10nF) surge capacitors were used. Prototypes of chokes were experimentally tested and confirmed the applicability of the series-choke protection concept to mitigating high du/dt transients resulting e.g. from the VCB switching operations. The resistor value should provide the best suppressing effectiveness. The voltage at the transformer terminal was taken as the criterion. Simulation results demonstrated that the best performance of suppressing device is achieved for resistor value close to cable surge impedance The parameters of the recorded during the tests transients were presented in Tab. 2 and Fig 16.

Tab. 2. Average du/dt values observed during tests

.
Fig.16. VFTs parameters for 40cm2 core cross-section and various damping resistance values

When chokes are inactive and no suppressing device is connected before transformer voltage steepness is few tens of kilovolts per microsecond.

Connecting chokes with larger core cross-section, results both in the voltage steepness reduction and the overvoltage peaks suppression. For relatively large permeability cores with cross-section (40 cm2), du/dt reduction is insignificant then for cross-section larger (70 ÷ 120 cm2) du/dt reduction is very high (up to 10x).

For all tested chokes high degree of oscillation reduction is observed during breaking, especially when additional suppressing resistor is connected at the secondary winding.

Performed experiments demonstrated that, despite that the risetimes of the waveforms observed at the transformer terminals for the base case (with no protection) were relatively long due to the experiment set-up limitations, a significant reduction in the transients amplitudes was observed. Especially, when an appropriate combination of core parameters and the resistance value was applied.

REFERENCES

[1] CIGRE working group A2-A3-B3.21, Electrical Environment of Transformers; Impact of fast transients”, ELECTRA 208, (2005)
[2] Lopez–Roldan J., De Herdt H., Min J., Van Velthove R., Decklerq J., Sels T., Karas J., Van Dommelen D., Popow P., Van der Sluis L., Aquado M., Study of interaction between
distribution transformer and vacuum circuit breaker, Proceedings of 13th ISH (2003), pp. 62÷64
[3] Popov M., Acha E., Overvoltages due to switching off an unloaded transformer with a vacuum circuit breaker, IEEE Trans. on Power Delivery, Vol. 14, No. 4, (1999), pp. 1317÷1322
[4] Burrage L. M., Shaw J. H., McConnell B. W., Distribution transformer performance when subjected to steep front impulses, IEEE Trans. on Power Delivery, Vol. 5, No. 2, (1990)
[5] Piasecki W., Bywalec G., Florkowski M., Fulczyk M., Furgal J., New approach towards Very Fast Transients suppression, Proceedings of IPST’2007
[6] Paul D., Failure Analysis of Dry-Type Power Transformer, IEEE Transaction on Industry Applications, Vol. 37, No. 3, (2001)
[7] Wong S. M., Snider L. A., Lo E. W. C., Overvoltages and reignition behavior of vacuum circuit breaker, Proceedings of IPST’2003
[8] Smugała D., Piasecki W. , Ostrogórska M., Florkowski M., Fulczyk M., Kłys P., Distribution transformers protection against High Frequency Switching Transients, PRZEGLĄD ELEKTROTECHNICZNY (Electrical Review), ISSN 0033-2097, R. 88 NR 5a/2012
[9] Florkowski M.,Fulczyk M.,Ostrogórska M., Piasecki W., Steepness reduction of ultra-fast chopped surges at transformer terminal, ICHVE (2012), pp.145÷149

Authors/Autorzy:
Dariusz Smugała, PhD. Eng., E-mail: dariusz.smugala@pl.abb.com
Magdalena Ostrogórska, MsC.Eng. E-mail: magdalena.ostrogorska@pl.abb.com
Wojciech Piasecki, PhD.Eng., E-mail: wojciech.piasecki@pl.abb.com
Marek Florkowski, PhD,DSc.,Eng E-mail: marek.florkowski@pl.abb.com
ABB Corporate Research Center, Starowislna 13 A Str., 31-038 Cracow, Poland,
Marek Fulczyk, PhD.Eng. E-mail: marek.fulczyk@pl.abb.com
Muottitie 2, 65100 Vaasa, Finland
Ole Granhaug, MSc.Eng. E-mail: ole.granhaug@no.abb.com
ABB AS Amtm. Aallsgate 73, P.O. Box 108 Sentrum, 3717, Norway

Source & Publisher Item Identifier: PRZEGLĄD ELEKTROTECHNICZNY, ISSN 0033-2097, R. 89 NR 10/2013

Transient Performance of Interconnected Wind Turbine Grounding Systems

Published by Blagoja MARKOVSKI, Leonid GRCEV, Vesna ARNAUTOVSKI-TOSEVA, Marija KACARSKA,
Ss. Cyril and Methodius University of Skopje, Faculty of Electrical Engineering and Information Technologies

Abstract. We analyze the transient grounding characteristics of interconnected wind turbine grounding systems, for fast rising current pulses. By increasing the number of wind turbines, influences on harmonic impedance and transient potential are examined for different soil characteristics and different locations of excitation. Simulations are performed using simple model of grounding system that neglects the foundation reinforcement. The influence of such simplification for isolated wind turbines is analyzed in previous papers. Here we extend the previous analysis for interconnected wind turbines and we look at the possibilities for optimization of the transient analysis of extended grounding systems in wind farms.

Streszczenie. Przeanalizowano w pracy charakterystyki stanów przejściowych uziemień turbin wiatrowych dla szybkiego narastania impulsów prądowych. Przy wzroście liczby turbin wiatrowych zbadano wpływu na impedancje harmoniczną oraz stan przejściowy dla różnych charakterystyk gleby i różnych lokalizacji wzbudzenia. Przeprowadzono symulacje przy użycie prostego modelu uziemienia, który zaniedbuje wzmocnienie fundamentów. Wpływ takiego uproszczenia dla pojedynczych turbin przebadany został w e wcześniejszych publikacjach. W tym artykule rozszerzono poprzednia analizę na połączone turbiny oraz skierowano uwagę na możliwości optymalizacji analizy stanów przejściowych rozbudowanych systemów uziemienia na farmach wiatrowych. (Wyznaczenie stanu przejściowego w systemie uziemienia elektrowni wiatrowej ).

Keywords: grounding system, lightning, transient analysis, wind turbine.
Słowa kluczowe: systemy uziemienia, wyładowania atmosferyczne, analiza stanu przejściowego, turbina wiatrowa

Introduction

Recently a number of papers have been devoted to the transient performance of isolated wind turbine grounding systems [1-4]. In practice, wind turbines are often spread across large areas, electrically interconnected by buried medium voltage cables. Metallic armour of such cables and bare wires are bonded to the wind turbine grounding electrodes, forming an extended grounding system. Such connection should provide significant reduction of the overall grounding resistance [5]. Due to the specific construction at exposed locations, wind turbines often suffer direct lightning strikes that may provoke damage or malfunction on the equipment. Therefore the high frequency performance of interconnected grounding systems is of great practical interest.

In this paper we analyze the transient performance of extended grounding system in wind farm. We consider typical grounding systems of wind turbines with spread footing foundations, interconnected with bare wires buried at depth of 0.5 m. Mutual separation between wind turbines will be varied between 75 m and 300 m to analyze the influence of the length of the buried bare conductor that is in direct contact with earth (only in case of three wind turbines in row, in other cases mutual separation is 300 m ). Details of the grounding system geometry are given in Fig. 1. By increasing the number of wind turbines, influences on the harmonic impedance and on the transient potential will be examined for different types of soil and for lightning current pulses related to the first and subsequent return strokes.

Wind turbine arrangement is illustrated in Fig. 2. Two cases of lightning strike, one in the middle and at the end of the row are analyzed separately. Simulations are performed using simplified model of wind turbine grounding system that neglects the foundation reinforcement mesh.

Recent papers have shown that simplified models for isolated wind turbine grounding system lead to significant overestimation of the transient potential and harmonic impedance in the high frequency range [6-7]. Here we extend the previous analysis for interconnected grounding systems. We compare the influence on the harmonic impedance and transient potential for simplified model of the adjacent grounding system and for model that integrates the foundation reinforcement (see Fig. 1).

Rigorous electromagnetic model is used for the computations [8–9], based on a mathematical method developed from the antenna theory and solved by the method of moments. This model is implemented into the Tragsys computer software [10].

Fig.1. Wind turbine grounding system (thick lines) integrated with the foundation reinforcement mesh (thin lines)
Fig.2. Illustration of wind turbines arrangement
Frequency domain analysis

Harmonic impedance is important quantity in transient analysis of grounding systems. It does not depend on the excitation, but solely on geometry and electromagnetic characteristics of the grounding system and surrounding medium. It is equal to the grounding resistance R in the low frequency range and it has larger or smaller values than R in the high frequency range, whether the inductive or capacitive characteristics of the system are dominant.

Current with variable frequency from 100 Hz to 10 MHz is injected in the vertical conductors above earth, in one grounding system of the row (see Fig.2). Analysis are performed for low resistive earth, with ρ = 100 Ωm, and highly resistive earth, with ρ = 1000 Ωm, for excitation in the middle and at the end of the row. We analyze the influence of the buried bare bonding wires and the influence of the adjacent grounding systems. The foundation reinforcement mesh is omitted due to computational efficiency and its influence will be analyzed later.

From Fig. 3 and Fig. 4, it is evident that the buried horizontal bare wires have major influence on the reduction of the harmonic impedance in the low frequency range, while the influence of the adjacent grounding systems is considerably lower. The interconnection of several wind turbines improves the grounding performance for slow varying excitations with low frequency contents, such as fault currents or slow rising current pulses (typical for first lightning strokes) in case of highly resistive earth. However due to the great mutual separations, the adjacent grounding systems do not provide improvement in case of fast rising lightning current pulses (typical for subsequent return strokes). At high frequencies, currents dissipate only locally, near the affected wind turbine grounding system.

Time domain analysis in case of lightning strike

Time domain analysis are important for proper design of protective equipment. We analyze the transient potential (in respect to distant neutral earth) at current feed points, for low resistive earth with ρ = 100 Ωm and for highly resistive earth with ρ = 1000 Ωm. Two standardized lightning current waveforms related to the first and subsequent return strokes are considered. They are reproduced by means of a usual double exponential function:

.

where Im is the peak value of the current pulse. Values of the coefficients k, τ1 and τ2 for the current pulses are given in Table 1

Table 1. Parameters of first and subsequent lightning current pulse

T1/T2 [μs]Im [kA]kτ1 [μs]τ2 [μs]
10/3502000.9510.002110.2485
0.25/100500.9950.0069910.87
.

Fig.5 illustrates the transient potential at current feed points, for lightning current pulse related to the subsequent return stoke, injected in grounding system at the end of the row. Wind turbine interconnection has no influence on the transient potential in the initial surge period, during the current pulse rise, while the horizontal buried bare conductors significantly reduce the transient potential during the pulse decay. The adjacent grounding systems contribute to further reduction only in case of highly resistive earth, after a period of the decay time to half-peak. In case of low resistivity earth the adjacent grounding systems and the horizontal conductors longer than 200 m do not provide improvement of the transient performance

Fig.3. Harmonic impedance of interconnected wind turbines, for excitation at the end of the row: a) ρ=100 Ωm; b) ρ=1000 Ωm.
Fig.4. Harmonic impedance of interconnected wind turbines, for excitation at the middle of the row: a) ρ=100 Ωm; b) ρ=1000 Ωm.
Fig.5. Transient potential in respect to distant neutral earth for current pulse related to subsequent return stroke in low and highly resistive earth: a) ρ=100 Ωm; b) ρ=1000 Ωm.
Fig.6. Transient potential in respect to distant neutral earth for current pulse related to first stroke in low and highly resistive earth: a) ρ=100 Ωm; b) ρ=1000 Ωm.

Fig.6 illustrates the transient potential at current feed points, for lightning current pulse related to the first stoke, injected in grounding system at the end of the row. Transient response in case of slow rising current pulse is primarily governed by the low frequency behaviour of the harmonic impedance. In case of low resistive earth, the adjacent grounding systems provide small reduction of the transient potential during the entire transient period, while the main contribution comes from the buried bare bonding wires. In case of highly resistive earth, adjacent grounding systems provide significant reduction of the transient potential, after a period of the decay time to half-peak.

It is worth noting that for excitations in the middle of the row, harmonic impedance has lower values in the low frequency range (relative to the earth resistivity) than for excitations at the end of the row. Consequently, transient potentials are expected to be considerably reduced for first lightning stroke, and during the pulse decay for subsequent return strokes

Influence of the grounding system models in transient analysis of extended grounding systems in wind farms

Previous analysis have been performed using simplified model of grounding system that neglects the influence of the foundation reinforcement mesh, (see Fig. 1, with thin lines). Such model leads to significant overestimation of the transient potential, since the foundation reinforcement mesh provides additional paths of the lightning current and reduces the inductive behaviour of the harmonic impedance in the high frequency range. To analyze the influence of the grounding system models, we compare three cases. First the excited and the adjacent grounding systems are modelled by simple geometry that includes only the basic grounding electrodes. Next we use complex model for the adjacent grounding system, integrated with the foundation reinforcement. Finally we use complex integrated models for the two interconnected grounding systems.

Fig.7 shows that the complexity of the model for adjacent grounding system has no influence on the low frequency and high frequency performance of the grounding systems. The use of complex model that integrates the foundation reinforcement mesh has significance only for the local wind turbine grounding system, in case of transient analysis for fast rising current pulses. For low frequency analysis the use of simplified model for the local grounding system will not introduce significant errors.

Fig.7. Influence of the model for local and adjacent grounding systems: a) ρ=100 Ωm; b) ρ=1000 Ωm
Conclusion

When wind turbine grounding systems are interconnected by buried bare wires, these wires are most effective in the improvement of the transient performance. Adjacent grounding systems provide negligible improvement of the transient performance in case of low resistive earth. For highly resistive earth, the adjacent grounding systems provide additional improvement of the transient performance, during the pulse decay period that is mostly governed by the low frequency behaviour of the extended grounding system.

The analysis for lightning strikes at the end of the cascade can be considered as worst case analysis for interconnected grounding systems. Lightning strikes to wind turbines in the middle of the row will produce lower transient potentials. This is due to the lower values of the harmonic impedance in the low frequency range for different types of soil (see Fig.4), than in case of excitation at the end of the row (see Fig.3).

The use of complex model of grounding system, integrated with the dense mesh of the foundation reinforcement has significance only for the local grounding system that is directly affected by the lightning current pulse. For low frequency analysis, simple model of grounding system can be used as well. The complexity of the adjacent grounding systems has negligible influence on the transient performance of the extended grounding system. This observation is important for optimization and reduction of the computational times during transient analysis of extended grounding systems

REFERENCES

[1] Y. Yoh, Y. Takuma, FDTD analysis of wind turbine earthning, in Proc. 28-th Int. Conf. Lightning Protection, (2006), 1551- 1556.
[2] S. Pastromas, E. Pyrgioti, Two types of earthning system of lightning protection for wind turbines, in Proc. 29th Int. Conf. on Lightning Protection, (2008), 1-10.
[3] A. Elmghairbi, A. Haddad, H. Griffiths, Potential rise and safety voltages of wind turbine earthning systems under transient conditions, in Proc. 20th Int. Conf. Electricity Distribution (CIRED 2009), (2009), 1-4.
[4] S. Yanagawa, D. Natsuno, K. Yamamoto, Measurements of transient grounding characteristics of a MW class wind turbine generator system and its considerations, in Proc. 31st Int. Conf. on Lightning Protection, (2012), 1-5.
[5] Wind Turbine Generator Systems – Part 24: Lightning protection, IEC Std. 61400-24, (2010).
[6] B. Markovski, L. Grcev, V. Arnautovski-Toseva, Step and touch voltages near wind turbine grounding during lightning strokes, International Symposium on Electromagnetic Compatibility (EMC Europe 2012), (2012), 1-6.
[7] B. Markovski, L. Grcev, V. Arnautovski-Toseva, Transient characteristics of wind turbine grounding, Frequency dependent and soil ionization effects, in Proc. 31st Int. Conf. on Lightning Protection, (2012), 1-6.
[8] L. Grcev and F. Dawalibi, An electromagnetic model for transients in grounding system, IEEE Trans. Power Delivery, 5 (1990), 1773-1781.
[9] L. Grcev, Computer analysis of transient voltages in large grounding systems, IEEE Trans. Power Delivery, 11 (1996), 815-823.
[10] TRAGSYS-software for high frequency and transient analysis of grounding systems. http://www.tragsys.com

Authors: M.Sc. Blagoja Markovski, Ss. Cyril and Methodius University, Faculty of Electrical Engineering and Information Technologies, P.O. Box 574, 1000 Skopje, Macedonia, Email: bmarkovski@feit.ukim.edu.mk;
Prof. Dr. Leonid Grcev, Ss. Cyril and Methodius University, Faculty of Electrical Engineering and Information Technologies, P.O. Box 574, 1000 Skopje, Macedonia, Email: lgrcev@feit.ukim.edu.mk;
Prof. Dr. Vesna Arnautovski-Toseva, Ss. Cyril and Methodius University, Faculty of Electrical Engineering and Information Technologies, P.O. Box 574, 1000 Skopje, Macedonia, Email: atvesna@feit.ukim.edu.mk;
Prof. Dr. Marija Kacarska, Ss. Cyril and Methodius University, Faculty of Electrical Engineering and Information Technologies, P.O. Box 574, 1000 Skopje, Macedonia, Email: mkacar@feit.ukim.edu.mk

Source & Publisher Item Identifier: PRZEGLĄD ELEKTROTECHNICZNY, ISSN 0033-2097, R. 91 NR 6/2015. doi:10.15199/48.2015.06.13

Harmonic Current Cancellation Evaluation

Published by Electrotek Concepts, Inc., PQSoft Case Study: Harmonic Current Cancellation Evaluation, Document ID: PQS1002, Date: March 15, 2010.

Abstract: Utility power system harmonic problems can often be solved using a comprehensive approach including site surveys, harmonic measurements, and computer simulations.

This case study presents the results for a customer step-down transformer harmonic current cancellation evaluation. The analysis was completed using the SuperHarm program. The simulation results show the reduction of the customer primary harmonic current distortion level due to the phase shifting caused by the different step-down transformer connections.

INTRODUCTION

A harmonic current cancellation evaluation was completed for the system shown in Figure 1. The case study was completed using the SuperHarm program. The accuracy of the simulation model was verified using three-phase and single-line-to-ground fault currents and other steady-state quantities.

Isolation transformers can be used with multiple nonlinear loads to create an effective twelve-pulse operation. In an actual twelve-pulse configuration, the front-end rectifier circuit uses twelve diodes instead of six. When properly designed, this configuration practically eliminates the 5th and 7th harmonics. The disadvantages of this design are cost and construction due to the requirement for either a delta-delta/delta-wye transformer pair, or a three-winding transformer to accomplish the 30° phase shifting necessary for proper operation. This configuration also affects the overall drive system efficiency rating because of the voltage drop associated with the transformer requirement.

One possible harmonic current cancellation method is to use a pseudo twelve-pulse configuration that can be achieved by supplying one adjustable-speed drive through a delta-wye connected transformer, and another drive through a delta-delta connected transformer. When the two waveforms are combined on the primary, the resulting waveform injected into the utility system has a much lower current distortion value, primarily because the 5th and 7th harmonics nearly cancel.

For effective cancellation to occur, the customer nonlinear loads must be operated simultaneously and have similar characteristics. This is true for the system shown in Figure 1 where multiple adjustable-speed drives are being operated in pairs. Figure 2 shows the simulated current waveform and harmonic spectrum (single phase shown) for the 250 hp, 480 volt dc drive running at full load and 75% power factor. The current has a fundamental frequency value of 297 amps, an rms value of 315 amps, and a THD value of 35.2%. The utility source impedance may be approximated using the following expression:

XSC ≈ kVφφ2 / MVA3φ = 12.52 / 200 = 0.781Ω

where:
kVφφ = system rms phase-to-phase voltage (kV)
MVA3φ = three-phase short circuit capacity (MVA)
XSC = system short circuit reactance (Ω)

Figure 1 – Illustration of Oneline Diagram for Harmonic Current Cancellation Evaluation
Figure 2 – Customer DC Drive Current Waveform and Spectrum
SIMULATION RESULTS

This case study evaluates the effect of applying transformers with different connections to determine the harmonic current distortion levels on both the primary and secondary sides of the customer transformers. Case #1 involved two delta/delta transformers. Case #2 involved one delta/delta and one delta/wye transformer. Case #3 involved two delta/wye transformers. The simulation results are summarized in Table 1. The results show that the only case that achieves the harmonic cancellation is Case #2, which includes the required 30-degree phase shift introduced by the delta/delta, delta/wye configuration.

Figure 3 shows the simulation results for Case #1, where the current distortion on the primary of the customer transformer is not reduced due to the harmonic current cancellation.

Figure 4 shows the simulation results for the Case #2, which is the mitigation solution case highlighting the effect of the 30-degreee phase shift.

Table 1 – Summary of Simulation Results for the Harmonic Cancellation Evaluation

Case #TX #1TX #2IPCCIISO1IISO2
1Δ ΔΔ Δ33.8%33.8%33.8%
2Δ ΔΔ Y8.9%33.8%33.8%
3Δ YΔ Y33.8%33.8%33.8%
.
Figure 3 – Simulation Results for Case #1
Figure 4 – Simulation Results for Case #2
SUMMARY

This case study summarizes the results for a customer step-down transformer harmonic current cancellation evaluation. The results show the reduction of the customer primary harmonic current distortion level due to the phase shifting caused by the different transformer connections.

REFERENCES

1. Power System Harmonics, IEEE Tutorial Course, 84 EH0221-2-PWR, 1984.
2. IEEE Recommended Practice for Monitoring Electric Power Quality,” IEEE Std. 1159-1995, IEEE, October 1995, ISBN: 1-55937-549-3.
3. IEEE Recommended Practices and Requirements for Harmonic Control in Electrical Power Systems, IEEE Std. 519-1992, IEEE, ISBN: 1-5593-7239-7.

RELATED STANDARDS
IEEE Std. 519-1992
IEEE Std. 1159-1995

GLOSSARY AND ACRONYMS
ASD: Adjustable-Speed Drive
CF: Crest Factor
DPF: Displacement Power Factor
PF: Power Factor
PWM: Pulse Width Modulation
THD: Total Harmonic Distortion
TPF: True Power Factor

Understanding Power Flow and Naming Conventions In Bi-directional Metering Applications

Published by Michael Bearden

Introduction

In 1982 Raymond “Ray” Stevens published a paper “Power Flow Direction Definitions for Metering of Bi-directional Power”. This paper helped reduce the confusion in the terminology associated with the naming of power quantities based on the direction of active and reactive power flow. The paper did not address leading and lagging current and Power factor. The illustration below came from the paper and has been adopted as a standard and is used in the Handbook for Electricity Metering.

Over the last 24 years we have all used and continue to use the same terms (words) when we are talking about bi-directional power flow. However over this period of time, we have made up some new terms and continue to have issues with some old terms “lead and lag”. The big issue is that when two people are using the same terms they do not mean the same things. The intent of this paper is to help us not only say the same thing but to mean the same thing.

Illustration 1 – Power Values
Review

In using the above illustration to base the definition of power and the direction of the power flow, it is important that after the source and load have been defined for the metering point they can not be changed because active power flow changes direction.

Active Power (True Power): Watts

When the Active Power (Watts) flow from the “SOURCE” through the metering point and into the “LOAD” we say the Active Power (Watts) are being DELIVERED. Therefore when the Active Power is being supplied by the “SOURCE” into the load it will be referred to as Delivered Power (Watts) and has a positive sign.

When the Active Power (Watts) flow from the “LOAD” through the metering point and into the “SOURCE” we say the Active Power (Watts) are being RECEIVED. Therefore when the Active Power is being supplied by the “LOAD” into the source it will be referred to as Received Power (Watts) and has a negative sign.

Note: The Active Power is always on the x axis and does not fall into any of the four quadrants. Active power does not lead or lag it is delivered or received.

Reactive Power: Vars

When the Reactive Power (Vars) flow from the “SOURCE” through the metering point and into the “LOAD” we say the Reactive Power (Vars) are being DELIVERED. Therefore when the Reactive Power is being supplied by the “SOURCE” into the load it will be referred to as Delivered Reactive Power (Vars) and have a positive sign.

When the Reactive Power (Vars) flow from the “LOAD” through the metering point and into the “SOURCE” we say the Reactive Power (Vars) are being RECEIVED. Therefore when the Reactive Power is being supplied by the “LOAD” into the source it will be referred to as Received Reactive Power (Vars) and have a negative sign.

Note: The Reactive Power is always on the y axis and does not fall into any of the four quadrants. Reactive power does not lead or lag it is delivered or received.

Terms (Standard from the above illustration):

Delivered Kw/Kwh
Received Kw/Kwh
Delivered Kvar/Kvarh
Received Kvar/Kvarh

Apparent Power: Volt-amps

When the Active Power (Watts) flows from the “SOURCE” through the metering point and into the “LOAD” (Delivered Power) and the “LOAD” is resistive (No Vars) the Apparent Power (VA) will be on the x axis with Watts. Apparent Power has no sign or defined direction it is a vector quantity.

When the Active Power (Watts) flows from the “SOURCE” through the metering point and into the “LOAD” (Delivered Power) and the “LOAD” is inductive (Vars are present) the Apparent Power (VA) will be in Quadrant 1. Apparent Power has no sign or defined direction it is a vector quantity.

When the Active Power (Watts) flow from the “SOURCE” through the metering point and into the “LOAD” (Delivered Power) and the “LOAD” is capacitive (Vars are present) the Apparent Power (VA) will be in Quadrant 4. Apparent Power has no sign or defined direction it is a vector quantity.

When the Active Power (Watts) flows from the “LOAD” through the metering point and into the “SOURCE” (Received Power) and the “LOAD” is inductive (Vars are present) the Apparent Power (VA) will be in Quadrant 2. Apparent Power has no sign or defined direction it is a vector quantity.

When the Active Power (Watts) flows from the “LOAD” through the metering point and into the “SOURCE” (Received Power) and the “LOAD” is capacitive (Vars are present) the Apparent Power (VA) will be in Quadrant 3. Apparent Power has no sign or defined direction it is a vector quantity.

Terms (Standard from the above definition):

Kva/Kvah
Kva/Kvah Quadrant 1
Kva/Kvah Quadrant 2
Kva/Kvah Quadrant 3
Kva/Kvah Quadrant 4

Note: We also like to group our Kva/Kvah values based on the direction of power flow and then refer to them as Delivered and Received Kva/Kvah even though the Kva/Kvah has no sign or direction.

If we could (would) stop at this point there would be little to NO confusion. When we were talking about delivered and received power we would all be using the same words and mean the same thing. Therefore, when possible the above definition should always be used when referring to Active, Apparent and Reactive Power.

The paper did not address Leading and Lagging current or talk about Power Factor.

We will address the terms leading and lagging along with power factor later in this paper.

What’s New
Additional Kvar/Kvarh Quantities

Some of the confusion comes when terms are used that fall outside of the standard power flow definitions. The following is a list on non-standard terms term that are used routinely.

Terms (non-standard or made-up names):

Kvar/Kvarh (absolute Del. + Rec.) with Delivered Power (Kw)
Kvar/Kvarh (absolute Del. + Rec.) with Received Power (Kw)
Kvar/Kvarh (Net, Del. – Rec.) with Delivered Power (Kw)
Kvar/Kvarh (Net, Rec. – Del.) with Received Power (Kw)

The first question maybe, what are the quantities and what are they used for? After talking to a number metering people, I found out they are quantities that are used to comply with billing (needs) tariffs and to get the same results as in the past using electro-mechanical meters. Names were then created (made up) that best describe the quantity or requirement.

The next question maybe, where (how) are the quantities being used? The two most common uses for the absolute values are for calculating Delivered and Received Kva/Kvah which is a made-up name for an undefined electrical quantity. The other application is in billing Kvar/Kvarh (absolute value) based on the flow of active power. The two most common uses for the net values are for calculating Kqh to be used with Delivered and Received Kw/Kwh. The other application is in billing Net Kvar/Kvarh (this is where the customer is given an equal credit for Received Kvar/Kvarh) based on the flow of active power.

The reason for the confusion is the terms (delivered and received) are being used interchangeable between non-standard and the standard names for Reactive Power. This allows one person to be talking about Delivered Vars as per Illustration 1 and the other person think he is talking about the nonstandard terms (Kvar with Delivered Power).

The other term that causes confusion and the meaning is unknown when talking about Kvar/Kvarh is Leading and Lagging. The reason for the confusion is, Vars do NOT Lead or Lag, The quantities (term) for Vars is Delivered and Received. This is very important in bi-directorial applications. The only quantity that Leads or Lags is current and it Leads and Lags in reference to voltage.

We will look at leading and lagging current later in this paper.

Additional Kva/Kvah Quantities

Some of the confusion comes from terms that we use that fall outside of the standard power flow definitions and naming convention as described above. The following is a list on non-standard terms that are used routinely.

Terms (That are non-standard or have made-up names):

Kva/Kvah with Delivered Power (Kw)
Kva/Kvah with Received Power (Kw)
Kva/Kvah Quadrant 1 Only
Kva/Kvah Quadrant 3 Only

Please note the non-standard terms are not defined electrical quantities. They are one electrical value (Kva/Kvah) that is being referenced to, in conjunction with a standard defined electrical quantity (Kw/Kwh). The Quadrant 1 and 3 go back to the old electro-mechanical metering days when the reactive metering package was made-up of one Kwh and one Kvarh meter and a phase shifting transformer. The Kva/Kvah was then calculated base on the results of the Kwh and Kvarh meter.

Additional Terms (Lead and Lag)

The terms lead and lag should only be used in reference to current. The current will be in phase with the voltage or it can lead or lag the voltage depending on the device taking active power. The terms leading or lagging current is always viewed from the perspective of the point which is supplying the active power. We are going to look at the terms leading and lagging current in conjunction with the same illustration (see illustration 1) that was used to define power terms as before. To help understand the labels for leading and lagging current and help reduce the confusion I have changed the reference from Load to IPP and Source to System (see illustration 2). The meter has been connected to register delivered active power when the IPP is taking power from the system and the meter will register received power when the system is taking active power from the IPP.

Reference 1

In the first set of examples (conditions) the IPP is seen as the load by the system which is providing the active power. When the IPP is taking active power from the system (which is 95% of our metering installations) we say the power is being delivered. The three conditions which follow should help us to understand leading and lagging current when the active power is in the delivered direction.

Condition 1;

When the IPP appears as a resistive device to the system, this will cause the current to be in phase (not leading or lagging) with the voltage and the Kva will be on the x axis with the active power (Kw delivered) from the system.

Condition 2;

When the IPP appears as an inductive device to the system, this will cause the current to lag the voltage and the Kva will move into quadrant 1. The IPP is now taking both active power (Kw delivered) and reactive power (Kvar delivered) from the system.

Condition 3;

When the IPP appears as a capacitive device to the system, this will cause the current to lead the voltage and the Kva will move into quadrant 4. The IPP is now taking active power (Kw delivered) from the system and sending reactive power (Kvar received) back to the system.

Reference 2

In the second set of examples (conditions) the System is seen as the load by the IPP which is providing the active power. When the IPP is sending active power to the system (which is 5% of our metering installations) we say the power is being received. The three conditions which follow should help us to understand leading and lagging current when the active power is in the received direction.

Condition 1A;

When the System appears as a resistive device to the IPP, this will cause the current to be in phase (not leading or lagging) with the voltage and the Kva will be on the x axis with the active power (Kw received) from the IPP.

Condition 2A;

When the System appears as an inductive device to the IPP, this will cause the current to lag the voltage and the Kva will move into quadrant 3. The System is now taking both active power (Kw received) and reactive power (Kvar received) from the IPP.

Condition 3A;

When the System appears as a capacitive device to the IPP, this will cause the current to lead the voltage and the Kva will move into quadrant 2. The System is now taking active power (Kw received) from the IPP and sending reactive power (Kvar delivered) back to the IPP.

Illustration 2 – Leading and Lagging
Additional Values (Power Factor)

Power Factor, is another value that I hear people sticking on the terms lead and lag. Power Factor is the ratio between true and apparent power. The ratio will always be between 0.0 and 1.0 and will not have a sign. The following terms are commonly used (naming conventions) for Power Factor.

Power Factor
Delivered Power Factor
Received Power Factor
Average Power Factor
Average Delivered Power Factor
Average Received Power Factor

Power Factor: is the ratio between true and apparent power (normally the instantaneous value).

Delivered Power Factors: is the power factor associated with delivered power (normally associated with a maximum demand value).

Received Power Factors: is the power factor associated with received power (normally associated with a maximum demand value).

Average Power Factor: What is average power factor? The best answer that I found, was average power is the ratio between the accumulated kwh and Kvah over some period of time, normally between demand resets.

Avg. Del. Power Factors: is power factor associated with delivered power.
Avg. Rec. Power Factors: is power factor associated with received power.

Summary

Power Values (Kw/kwh, Kvar/Kvarh and Kva/Kvah): The best naming convention would be to stay with Delivered and Received (Illustration 1).

Leading and Lagging: Use these terms for describing the relationship of current to voltage. Leading and Lagging is normally viewed from the perspective of the supplier of active energy (Illustration 2).

Power Factor: Is the ratio between true and apparent power. Power factor does not lead or lag and has no sign. Power factor is normally viewed from the perspective of the supplier of active energy.