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Improved Gapped-Core CT Dimensioning Algorithm Considering Relay and System Requirements Mahdi Davarpanah, Student Member, IEEE, and Majid Sanaye-Pasand, Senior Member, IEEE
Abstract—Current-transformer (CT) core saturation leads to relay maloperation. Meanwhile, the increment of core area and subsequently the total weight of cores not only leads to a heavy and expensive CT, but also it is one of the most important restrictions in manufacturing CTs. Gapped-core CTs, especially the TPY-type CT, are used to decrease the core area and weight, and to reduce the core remanent flux and possibility of CT saturation. Designing a TPY-type CT includes proper determination of influencing parameters and the utilization of an appropriate analytical approach. For this purpose, a new algorithm to design the TPY-type CT, and some analytical methods to compute the required parameters of gapped-core CTs are proposed in this paper. It is indicated that the CT dimensioning based on the proposed approach results in a much lower expense CT while it satisfies the IEC standard and relay requirements. Index Terms—Current-transformer (CT) saturation, distance relay, gapped-core CT, transient dimensioning factor.
I. INTRODUCTION
A
CURRENT transformer (CT) should supply protective relays by stepping down the fault current level, while preserving its original waveshape. If the fault current is a perfect sinusoid, the appropriately designed CT maintains high accuracy, even if the current amplitude is very high [1]. However, fault currents often contain a considerable amount of decaying dc component which generates a dc flux in the CT iron core and may result in CT saturation. The current reproduced by a saturated CT may be significantly different from the real fault current. Therefore, it can lead to relay maloperation or an additional operational delay [2]–[7]. Furthermore, modern power systems may operate close to their stability limits. In this condition, accurate transmissionline relay operation and high-speed fault clearing are required to prevent undesirable line outage and to preserve transient stability. Since the proper CT performance is a vital requirement for correct operation of protective relays, the CT knee-point voltage should be computed accurately based on the relay and power system requirements. Two strategies have been developed to properly step down the fault current including: 1) Manuscript received March 10, 2012; revised October 31, 2012; accepted December 12, 2012. Date of publication January 29, 2013; date of current version March 21, 2013. This work was supported by the University of Tehran under Grant 8101064-1-05. Paper no. TPWRD-00251-2012. The authors are with the Electrical and Computer Engineering School, College of Engineering, University of Tehran, Tehran 14395-515, Iran (e-mail:
[email protected];
[email protected]). Color versions of one or more of the figures in this paper are available online at http://ieeexplore.ieee.org. Digital Object Identifier 10.1109/TPWRD.2012.2234485
selecting a proper CT knee-point voltage and 2) utilizing electronic current transducers (ECTs) (e.g., optical current sensor [8] and Rogowski coil [9]). However, such transducers require an external power supply and are not widely used for ac system protection. On the other side, dimensioning the conventional CT to transfer the worst possible fault current without any distortion may result in a very high amount of , especially due to the remanent flux which can be as high as 80% of the saturation flux for closed-core CTs [10]. The cost and size of such a CT can also become very high and impractical. The most practical method to decrease the remanent flux is to use an air gap in the core magnetic path of CT [10]–[12]. TPY-type CT is a gapped-core CT which usually has been utilized for transmission-line distance relays and its remanent flux should be limited to 10% of the saturation flux [13]. This char. It should be acteristic considerably decreases the required noted that TPZ-type CTs, with a large air gap and negligible remanent flux, can also be used as a protective CT [14]. However, due to poor accuracy, it is not usually recommended by many relay manufacturers [15], [16]. Although the required core area of a TPY-type CT is usually much less than that of a closed-core CT, this CT may still be large and heavy. In addition, the total weight of cores is one of the most important restrictions in manufacturing a top-core-type CT in which the cores are placed at the top of the CT bushing [17]. In other words, the application of this type of CT is not preferable when it includes two or more weighty cores, especially in the earthquake-prone regions. Meanwhile, only topcore-type CTs are manufactured by some companies for extra high voltage (EHV) applications due to some insulation considerations and the ability to pass a high amount of rated and short-circuit currents. Therefore, decreasing the CT dimensions is extremely important for top-core-type CTs. For tank-type CTs, in which the cores are placed in a tank close to the ground, the core weight is not a major restriction in the production of CT. However, CT in some busbar arrangements (e.g., one-and-a-half breaker busbar) may include seven cores consisting of four gapped-cores and three closed cores [18]. In this case, not only using a large core area results in a heavy and expensive tank-type CT, but also the expenses may considerably increase, because the CT may require a special design and type test. Although the prevention of CT saturation is theoretically possible by selecting a CT with enough large core area [10], this CT may practically have an excessive core area and weight. Consequently, industrial relays should work properly under a defined CT saturation condition in order for a lower expense CT to be manufactured. An acceptable degree of CT saturation should be
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DAVARPANAH AND SANAYE-PASAND: IMPROVED GAPPED-CORE CT DIMENSIONING ALGORITHM
determined by the relay manufacturer which is usually specified as the required time after the fault instant during which the CT must not be saturated. To calculate the CT knee-point voltage, the required parameters and equations are defined by the IEC standard [13]. The transient performance of a TPY-type CT for a fault containing a decaying dc component is explained in [11], [12], and [17] based on the IEC recommended methodology. However, the CT design algorithm and the appropriate values of parameters, which influence the CT dimensioning, have not been discussed comprehensively in the available documents. Modified power system operational techniques such as: 1) a widely utilized single-pole autoreclose instead of a three-pole one; 2) applying the controlled switching of transmission lines; and 3) developing digital relays with acceptable performance to some extent for CT saturation, impose new requirements on some system equipment, such as CTs. In other words, the parameters values, which influence CT dimensioning, must be modified considering the power system changes. To perform this subject and to investigate the effects on CT dimensioning, an improved algorithm to design the TPY-type CT is proposed in this paper. In addition, some analytical equations are suggested to formulate the designing restrictions defined by [13]. Finally, the appropriate values of influencing parameters, which can significantly affect the CT dimensioning, are discussed based on the relay, autoreclosing scheme, and power system requirements. II. PROPOSED EQUATIONS FOR THE TPY-TYPE CT DESIGN Some analytical equations are proposed in this section which are required to design the TPY-type CT. A. Secondary Loop Time Constant The amount of secondary loop time constant as follows [13]:
Fig. 1. Typical hysteresis curve of a closed-core magnetic core.
where is the mean length of the magnetic path, is the total length of the flux path, , is the magnetizing current, and is the effective applied field intensity. The permeability of the magnetic core is generally defined as the derivative of the magnetic flux with respect to the field intensity for a single-valued B-H curve. Meanwhile, it can also be used in (5) for the region adjacent to the remanent flux point on the hysteresis curve [20] (5) where is the remanent flux of the magnetic core for the closed-core CT. indicates an operating point on the hysteresis curve in the region adjacent to the remanent flux point (Fig. 1). Therefore, flux density in the magnetic core is concluded as follows. Since no residual flux is present in the gap of a TPY-type CT, its flux density should be computed by
is determined (1)
where , , and are the CT magnetizing inductance, secondary side winding resistance, and burden resistance, respectively. Due to presence of the gap, the effect of the magnetic core on the value is negligible. Therefore, can be practically estimated as a linear inductance by [13], [19]
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(6) where and indicate an operating point for the gap. If the flux fringing in the gap is ignored, is equal to (i.e., ). Combining (4)–(6) , results in (7) (7) Simplification of (7) leads to
(2)
(8)
where , , , and are the number of secondary winding turns, core area, gap length, and permeability of free space, respectively. can be derived from (1) and (2) as
where and are the equivalent permeability and residual flux of a TPY-type CT which are defined as
(3)
(9)
B. Calculation of the Remanent Flux
(10)
For TPY-type CTs, the flux magnetic path includes ferromagnetic material and two or more gaps filled by nonmagnetic materials of relative permeability equal to 1. In this case, the Ampere’s law is expressed by [19] (4)
Consequently, the ratio of the gapped-core remanent flux to the closed core one can be calculated by (11)
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assumed equal to . As is normally stated in minutes, (15) can be rewritten as follows: (16)
Fig. 2. Current transformer: (a) circuit diagram and (b) phasor diagram.
where , which is one of the magnetic core intrinsic features, can be measured based on the method recommended by [13]. Subsequently, can be computed using (11) based on the TPY-type CT dimensions and the measured .
in minutes. According to the IEC standard, maxwhere is imum-acceptable for TPY-type CT is 60 min [13]. Based on (16), should be more than 182 ms when equal to or less than 60 min is required. 3) Required Gap Length: Due to the remanent flux, the CT may be saturated before operation of high-speed protective relays. In order to overcome or reduce this problem, gapped-core CTs can be used. As shown in (11), the remanent flux of TPYtype CT is a function of the gap length, the permeability of the magnetic core, and the magnetic path mean length. To measure the remanent flux of the utilized magnetic core, a closed-core CT is examined. Then, the remanent flux is considered equal to 80% of the saturation flux [10]. In other words, , where is the saturation flux. According to (11), to satisfy the requirement of , the gap length should be greater than the following value: (17)
C. Formulation of TPY-Type CT Restrictions Some restrictions must be considered in the design of a TPYtype CT based on the IEC standard requirements, which are discussed in this subsection. 1) Instantaneous Error: The total instantaneous error of the TPY-type CT determined by (12) should be less than 10% [13]. Otherwise, a smaller gap length should be utilized to satisfy this restriction
For the studied CT in this paper, which uses oriented steel as the core material [21], the measured permeability of the magnetic core adjacent to the point is equal to about 0.0314 (corresponding to a relative permeability of 25000). Consequently, the required CT gap length must be greater than the following value to satisfy the requirement: (18)
(12) where is the rated transient dimensioning factor. 2) Phase Displacement and Corresponding : The magnetizing current of a TPY-type CT is normally much higher than the core-loss resistive current. Moreover, the burdens imposed on the CT by modern relays and instruments are usually resistive. Consequently, as shown in Fig. 2, CT phase displacement can be calculated by (13) where is the fault current at the CT secondary side, and is the magnetizing current. , as shown in Fig. 2, is derived as (14)
III. IEC METHOD FOR CT DIMENSIONING The IEC standard recommends a widely utilized method to calculate the required for the TPY-type CT which is explained in this section. Afterwards, an algorithm for improved design of this CT type is proposed. A. Calculation of the Required
(15)
The dc component in the asymmetrical fault current causes the flux in the CT core to rise faster than it would rise with a symmetrical fault current. The following equation is recommended by [13] to consider the effect of the maximum decaying dc component on CT dimensioning:
is the secondary side time is close to zero, can be
(19)
The integration of (1), (13), and (14) leads to
where is in radians and constant in milliseconds. Since
It should be noted that a quite small gap on the order of 0.0001 to 0.0003 p.u. of the magnetic path length is recommended by [11] to achieve the remanent flux of 10%. This is in good agreement with (18) and verifies the proposed equation for gap length calculation.
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where is the decaying dc time constant of the fault current. is the system angular frequency, and is the duration of the fault current flow during which the CT specified accuracy must be maintained (i.e., the CT must not be saturated). Equation (19) is utilized to calculate the transient dimensioning factor when the remanent flux is ignored. However, most transmission lines are employed with the autoreclose scheme which results in some degree of remanent flux in the CT core prior to fast reclose of the circuit breaker (CB). In such conditions, (20) should be utilized to describe for a resistive burden when a single-shot autoreclose is applied [13], [22]
(20) and are the dc offset factors of the first and where second fault current flows. and are the duration of first fault current flow, and the autoreclose dead time, respectively. and are transient dimensioning factors caused by the first and second fault current flows, respectively. is a factor to calculate the remanent flux reduction during the autoreclose dead time. Furthermore, is the duration of the fault after CB reclose for which CT must not saturate. Meanwhile, is usually considered equal to by the relay manufacturers. Finally, the required considering the fast autoreclose is calculated by (21) is the ratio of the symmetrical short circuit current where to the CT rated primary current. and should be provided by the CT manufacturer, and the other parameters should be calculated based on the CT requirement of the relay and the system to which the relay and CT are connected.
Fig. 3. Flowchart of the TPY-type CT design algorithm.
B. Proposed Approach for Improved TPY-Type CT Design The flowchart of the proposed TPY-type CT design algorithm is illustrated in Fig. 3. According to (3), is dependent on the gap length, core area, and winding resistance. Moreover, the core area should be computed based on the calculated . Since and, therefore, depend on , and is also dependent on , the gapped-core CTs should be designed using an iterative approach. In this method, some parameters such as and are recalculated in each iteration. The initial values of the iterative approach include: 1) 0.32 mm; 2) ; 3) 1.7 T, equal to the knee-point flux-density of the CT magnetic core; and 4) to calculate the initial value of . Under such conditions, the initial values of and are computed by
where is the number of CT secondary winding turns. Based on comprehensive studies for numerous designed CTs, the considered initial core area decreases the number of required iterations and guarantees the convergence of the proposed iterative approach. Subsequently, should be calculated based on the core area, the number of winding layers, and the assumed conductor diameter of about 1 mm for a CT with the rated secondary current of 1 A, or about 2 mm for a rated current of 5 A. Then, can be approximately calculated using (3). Afterwards, at which reaches its peak value, , should be calculated by [13] (23)
(22)
(24)
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If than
reaches its peak value before (i.e., is smaller ), should be substituted for . Afterwards, can be computed using (20). Then, and the new amount of core area should be computed by (21) and (25), respectively
REQUIRED
TABLE I SOME INDUSTRIAL RELAYS
FOR
(25) Convergence criteria of the proposed approach include: 1) minor variations of the CT core area in two consecutive iterations and 2) fulfillment of all IEC standard restrictions. If the relative difference between the new calculated core area and the old one is more than 0.1% (i.e., 0.1), the new core area should be substituted for the old one, and the previous calculations should be repeated for the new core area. The iterative procedure is converged as soon as the condition 0.1 is met after which the proper core area and are calculated for the corresponding gap length. The IEC standard restrictions include: 1) limitation of remanent flux which is represented by the gap-length; the mean length of the magnetic path for the studied CT is 1.15 m; therefore, based on (18), a gap length of at least 0.32 mm is required to fulfill the remanent flux restriction; therefore is considered equal to 0.32 mm; 2) instantaneous error which should be less than 10%; and 3) the phase error, represented by , should be less than 182 ms. To consider enough construction margins, a safety factor for and should be taken into account which is assumed about 20%. Therefore, for the two last restrictions, 8% and 228 ms are, respectively, considered. It should be noted that the greater the gap length, the less the values of , , , and . Therefore, the maximum-acceptable gap length results in the minimum core area which is a low expense CT. Consequently, in the proposed algorithm, the gap length is increased with steps of 0.1 mm for calculation of the maximum-acceptable for which the designing restrictions are fulfilled. The maximum gap length, for which both of the restrictions are satisfied, is denoted by and the corresponding core area is considered as . IV. KNEE-POINT VOLTAGE INFLUENCING PARAMETERS A. Required Saturation-Free Period for Industrial Relays Modern relays can tolerate CT saturation to some extent. This amount is usually capability which decreases the required specified by the manufacturer as a required period after the fault inception during which the CT must not saturate. The required neglecting the remanent flux is expressed by (26) for many industrial distance relays. In this equation, is concluded by simplification of (19), where is much higher than and regardless of the remanent flux
(26) is usually specified by the relay manufacturer based where . As a on numerous tests performed on the relay for various for result, some manufacturers specify different amounts of as given in Table I [15], [16], [23]. Although , various
recommended by the relay manufacturer, should be utilized to calculate the lowest possible value of the required , if the relay manufacturer is unknown, the value of 10 ms would be recommended. This is a proper typical value for all investigated modern digital distance relays for close-in faults. It should be noted that distance relay manufacturers not only specify the CT requirements for a close-in fault, but also another required saturation-free period is usually given for a fault at the end of zone-1. Therefore, should be calculated for both of these cases and the maximum value must be considered. Since the modifications recommended in this paper for the second case is similar to those of the first one, the CT calculations are only accomplished for close-in faults due to the paper length limitation. B. DC Offset Factor Fully asymmetrical fault current occurs when a short circuit fault takes place at the zero voltage instant. Some investigations have shown that 95% of the network faults occur when the voltage angle is between 40 and 90 . In addition, fully asymmetrical fault current accompanied by the maximum remanent flux in the same direction can unlikely occur, simultaneously. Therefore, some CT manufacturers consider partial dc offset factors and , which are usually less than 0.7, to reduce the CT dimensions [15], [24]. On the other hand, controlled switching of transmission lines is these days employed to reduce detrimental electrical transients, especially during fast re-energizing. Based on this technique, which has been utilized for many modern transmission lines, both energizing and re-energizing instants are controlled in a way that make the current when the instantaneous phase-toground voltages close to zero [25], [26]. Accordingly, if the line circuit breaker is equipped by a controlled switching relay, it is proposed that full dc offset factor be considered after autoreclosing (i.e., ). Although, according to [15], may be assumed equal to 0.6, it is recommended to consider to take the worst possible case into account. C. Rated Burden The higher the rated burden of a closed-core CT is considered, is obtained. Similarly, it seems that the increase in the more the rated burden of a gapped-core CT always leads to a higher . However, if the value is increased, secondary loop time . constant will decrease which may result in a lower value of In this case, the CT will saturate earlier than the relay required time if the reduction of is more than the increment of the and ). total secondary loop resistance (i.e., the sum of For example, for a TPY-type CT with parameters given in Table II, the applied burden is about 2 , while the rated burden is 10 . This CT is designed based on the proposed algorithm
DAVARPANAH AND SANAYE-PASAND: IMPROVED GAPPED-CORE CT DIMENSIONING ALGORITHM
TABLE II TPY-TYPE CT PARAMETERS
TABLE III CALCULATED PARAMETERS OF THE DESIGNED TPY-TYPE CT
Fig. 4. Calculated
for the designed TPY-type CT.
which results in the computed parameters shown in Table III. As illustrated in Fig. 4, the calculated of the gapped-core CT for is almost 5% less than that for 2 , whereas it was assumed that increase in the rated burden leads to a higher value of . Therefore, unlike closed-core CTs, the rated burden of the TPY-type CTs should be considered equal to the applied burden. This rule must be considered especially when a CT is designed based on long autoreclose dead time and low values. D. Autoreclose Dead Time Autoreclosers are widely applied for restoring transmission lines to service subsequent to operation of their associated protective relays. An autoreclosing scheme can be employed as three-pole or single-pole. These are usually applied after clearing a phase-phase and single-phase to ground faults due to zone-1 distance relay operation, respectively. Single-pole autoreclosing scheme provides an efficient method of improving transient stability, reducing torsional impact on generator shafts and reducing switching voltage transients [27], [28]. Therefore, it is an attractive operating practice for many power utilities. The deionization time is dependent on several factors such as voltage level, fault current magnitude, wind velocity, air humidity, capacitive coupling to live phases, etc. [27]. The autoreclose dead time must be selected to provide sufficient time to deionize the arc channel after opening the faulted phase CB. Typical single-pole autoreclose dead time is in the order of 800–2000 ms whereas a dead time of about 300–400 ms can be sufficient for three-pole autoreclose [29]. Consequently, in this study, autoreclose dead time is considered equal to 800 and 400 ms for single-pole and three-pole reclosings, respectively.
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E. Proper Values of Influencing Parameters 1) EHV Transmission Line Autoreclosing Conditions: On EHV transmission lines (i.e., the rated voltage higher than 345 kV, for which the conductor spacing is increased), the percentage of multiphase faults decreases. For instance, the performed studies shows that only 3% of total short-circuits in the 525-kV transmission lines include three-phase or phase-phase-to-ground faults [30]. Not only are phase-phase-to-ground and especially three-phase faults rare on EHV lines, but also they are unlikely to be of a transient nature. Since, three-phase faults very often result from ground straps left in place after breaker maintenance or downed line structures, blocking of autoreclosing should be considered in this case [27]. This is true even for phase-phase to ground faults. In other words, only the single-phase and phase-phase faults are recommended to be considered in autoreclosing applications as probable transient faults. It should be noted that the extremum amount of parameters which influence the CT dimensioning is usually occurred in the case of three-phase faults. Meanwhile, as explained above, it is not necessary to consider three-phase faults. Therefore, to calculate the required amount, proper values of influencing parameters should be considered for both cases of phase-phase and single-phase faults. Then, out of the two calculated amounts, the maximum value must be considered as by which the CT core area is computed. 2) Affected Parameters Values: The parameters which influence the CT dimensioning and also are affected by reclosing scheme are discussed in this subsection. a) Single-phase fault: For a single-phase fault on the EHV transmission line, the current passing through the CT may be more than the three-phase fault current, because of proximity to generators. Moreover, the CT secondary lead resistance should be doubled [15]. On the other hand, a smaller and a greater are expected for the single-phase fault that usually leads to a lower than that of calculated based on the three-phase fault. b) Phase-phase fault: For a phase-phase fault, the current passing through the CT is 0.866 multiplied by the three-phase fault current [18]. Meanwhile, the autoreclose dead time should be considered similar to the three-phase fault, because of performing the three-pole reclose. As smaller and values are expected for phase-phase faults, the calculated is also smaller than that obtained for three-phase faults. As a result, neglecting the three-pole reclosing for a threephase fault, which is not a practical case for EHV lines, results in a lower required . Therefore, is computed only based on single-phase and phase-phase faults in the next section. V. CASE STUDIES In order to investigate the proposed design algorithm and also the suggested parameters values, one of the most important 400-kV substations of the Iran national grid, Jalal substation, is studied. This substation is located near some power plants. Table IV shows the Thevenin equivalent impedances of this substation which are acquired by simulation of the Iran grid using DIgSILENT software [31]. In this section, not only is the proposed algorithm to design a gapped-core CT evaluated, but also
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TABLE IV SYMMETRICAL THEVENIN IMPEDANCES OF THE STUDIED SUBSTATION
TABLE V TYPICAL CT PARAMETERS
Fig. 6. Calculated (a) core area, (b) instantaneous error, and (c) phase displacement of a TPY-type CT based on the parameters of Table V.
TABLE VI FAULT CURRENTS FOR THE STUDIED TRANSMISSION LINE
Fig. 5. Decreasing
as a result of increasing gap length.
the designed CT based on the widely utilized typical parameters values is compared with a CT which is designed based on the proposed modifications in determination of the parameters values. A. CT Design Using Typical Parameters Values Table V shows the typical parameters of a gapped-core CT which are widely utilized for TPY-type CT dimensioning of Iranian substations. Since many of these high voltage substations have been designed by reputable international companies, utilization of these parameters in other countries is also expectable. As an example, a CT with these parameters and the rated turn ratio of 1500/1, equal to 27, equal to 75 ms, and the rated burden of 6 are designed based on the proposed algorithm. According to (3), the more the gap length, the less the results and, therefore, a smaller is obtained. The calculated for a short-circuit fault with 90 ms, 400 ms, and 40 ms is shown in Fig. 5 for different gap lengths. As shown in Fig. 5, the calculated values are 15.3, 12.6, and 9.9 for gap lengths of 3, 2, and 1 mm, respectively. On the other hand, a larger gap length leads to a greater and , as illustrated in Fig. 6. Consequently, the best design of TPYtype CT is obtained by selecting the largest gap length which provides minimum required core area and acceptable and amounts. As shown in Fig. 6, if the maximum possible gap length of 3 mm is utilized, the calculated , , and , will be equal to 583.4 ms, 3884 V, 61.4 , 7.96% and 28.6 min, respectively. In this case, and are less than their permissible thresholds (i.e., 8% and 48 min). B. CT Design Using Proposed Parameters Values As discussed previously, only two cases of single-phase and phase-phase faults may be considered for TPY-type CT dimen-
sioning. Table VI shows the calculated current amplitude and the corresponding values for various fault types at the beginning of a 400-kV transmission line in the Jalal substation. According to Table IV, the ratio of X/R for positive-sequence impedance is the largest. Therefore, asymmetrical fault currents have a lower value compared with that of the three-phase fault as given in Table VI. This decreases the required when CT dimensioning is performed considering fast autoreclose. Moreover, the proper value of is considered equal to 10 ms based on the capability of modern relays. Furthermore, the dc offset factors are assumed equal to 1 due to utilization of controlled switching in the studied transmission line and to consider the worst possible case. Accordingly, the TPY-type CT is designed based on the modified influencing parameters. Table VII shows the effect of the proposed parameters values on the calculated which is denoted by as compared with the calculated value based on the typical parameters value, . It should be noted that the CT design procedure is shown in several steps in Table VII to more clearly investigate the effect of each parameter. As given in Table VII, in the first step, is calculated only based on the modified dc offset factors which results in the of 2.97 times the base value calculated in the previous subsection (i.e., 3884 V). In the second step, in addition to the new values of and , the modified amount of is also considered for designing the CT. Although modified dc offset factors leads to a larger core area, modified CT saturation time ( ) results in a smaller CT core area compared with the base value. This indicates the importance of proper selection of the value. In the last two steps, required and amounts are computed based on the parameters corresponding to the single-
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TABLE VII TPY-TYPE CT DESIGN BASED ON THE MODIFIED PARAMETERS
Fig. 9. Impedance trajectory for the fault at 1% of the line length.
Fig. 10. Secondary and scaled primary currents for the fault at 40%. Fig. 7. Schematic circuit of the studied system.
Fig. 8. Magnetizing characteristic of the designed gapped-core CT.
phase and phase-phase faults, respectively. In these steps, previously discussed modified parameters values are also considered. It should be noted that out of these two cases, the maximum calculated must be considered for CT dimensioning. As shown in Table VII, the required based on the proposed influencing parameter values is only 0.4 times of the base value. In the other words, the weight of the TPY-type CT core computed by the proposed parameters is 2.5 times less than that calculated by the typical values. Meanwhile, the results of other designed CTs based on the proposed algorithm and modified parameters values show that the CT weight reduction of up to 3 times is also achievable. These results indicate the necessity of modifying the TPY-type CT influencing parameters values as proposed in this paper. C. Distance Relay Performance To evaluate the proposed strategy for designing gapped-core CTs, the performance of a modeled distance relay under CT saturation conditions is investigated. Fig. 7 shows the studied system which is implemented in the PSCAD/EMTDC software. The positive and zero sequence impedances of the transmission line are and , respectively. Due to the low remanent flux of the designed TPY-type CT (i.e. 1.4%, its hysteresis characteristic is very similar to the magnetizing curve). Thus, the nonlinear characteristic of the CT core is modeled based on the magnetizing curve (Fig. 8). Fig. 9 shows the impedance trajectory for a phase-to-ground fault at 1% of the line length considering the CT remanent flux
Fig. 11. Impedance trajectory for the fault at 40% of the line length.
generated by autoreclosing, as one of the most severe CT saturation cases. In such an adverse condition, the impedance remains in the zone-1 of the distance relay and thus the relay correctly operates. Fig. 10 shows the CT secondary current as compared with the scaled primary current for a phase-to-ground fault at 40% which is distorted 27 ms after autoreclosing. Fig. 11 shows that the impedance trajectory remains inside the zone-1 which results in the proper relay operation. Consequently, these two severe cases show that the distance relay operates correctly when the gapped-core CT is designed based on the proposed strategy. One very important finding of this study, is that the core area of the gapped-core CT can be reduced without endangering proper operation of the protection system. VI. CONCLUSION Proper selection of the TPY-type CT parameters, which influence CT dimensioning, was scrutinized in this paper. This issue was investigated due to the new changes in power systems, such as: 1) capability of modern digital relays in proper operation under some extent of CT saturation; 2) reliable single-phase reclosing systems which are widely employed in some countries as the only applied autoreclosing scheme; 3) utilizing three-pole autoreclose only in the case of phase-phase faults; and 4) applying switching control relays for energizing and re-energizing the transmission-line CBs for reduction of the switching transient overvoltages. In order to investigate the effects of the proposed parameters values on , an analytical approach is required to result in an improved CT design algorithm considering
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the IEC standard restrictions. Therefore, initially some analytical equations were proposed to formulate the restrictions imposed by the IEC standard for gapped-core CTs. Subsequently, an algorithm to design the TPY-type CT was suggested which results in the minimum CT core area with proper selection of the gap length. At last, proper values of TPY-type CT parameters were discussed considering the relay, autoreclosing, and power system requirements. CT dimensioning in the studied substation indicated that the of the improved designed CT using proper parameters values is 2.5 times less than that of the designed CT based on the widely used typical parameters values. REFERENCES [1] A. Hooshyar, M. Sanaye-Pasand, and M. Davarpanah, “Development of a new derivative-based algorithm to detect CT saturation,” Inst. Eng. Technol. Gen., Transm. Distrib., 2012, accepted for publication. [2] H. W. Lee, Y. C. Kang, S. Jang, and Y. G. Kim, “Distance relay suitable for use with a measurement type current transformer,” in Proc. IEEE Power Tech Conf., Lausanne, Switzerland, 2007, pp. 1176–1181. [3] S. Miao, P. Liu, and X. Lin, “An adaptive operating characteristic to improve the operation stability of percentage differential protection,” IEEE Trans. Power Del., vol. 25, no. 3, pp. 1410–1417, Jul. 2010. [4] H. Dashti, M. Sanaye-Pasand, and M. Davarpanah, “Fast and reliable CT saturation detection using a combined method,” IEEE Trans. Power Del., vol. 24, no. 3, pp. 1037–1044, Jul. 2009. [5] Z. Gajić, B. Hillström, and F. Mekić, “HV shunt reactor secrets for protection engineers,” in Proc. 30th Western Protect. Relay. Conf., Washington, DC, Oct. 2003, pp. 1–30. [6] S. Nam, J. Park, S. Kang, and M. Kezunovic, “Phasor estimation in the presence of DC offset and CT saturation,” IEEE Trans. Power Del., vol. 24, no. 4, pp. 1842–1849, Oct. 2009. [7] C. S. Yu, “Detection and correction of saturated current transformer measurements using decaying DC components,” IEEE Trans. Power Del., vol. 25, no. 3, pp. 1340–1347, Jul. 2010. [8] S. Kucuksari and G. G. Karady, “Complete model development for an optical current transformer,” IEEE Trans. Power Del., vol. 27, no. 4, pp. 1755–1762, Oct. 2012. [9] M. Chiampi, G. Crotti, and A. Morando, “Evaluation of flexible Rogowski coil performances in power frequency applications,” IEEE Trans. Instrum. Meas., vol. 60, no. 3, pp. 854–862, Mar. 2011. [10] IEEE Guide for the Application of Current Transformers Used for Protective Relaying Purposes, IEEE Standard C37.110, 1966. [11] B. Bozoki, H. J. Calhoun, and C. M. Gadsden et al., “IEEE Committee Report, Gapped core current transformer characteristics and performance,” IEEE Trans. Power Del., vol. 5, no. 4, pp. 1732–1740, Oct. 1990. [12] Y. C. Kang, J. Y. Park, J. Y. Lee, B. E. Jang, S. I. Kim, and Y. G. Chonbuk, “Compensation of an air-gapped current transformer considering the hysteresis characteristics of the core,” in Proc. DPSP Conf., Glasgow, U.K., 2008, pp. 495–500. [13] Instrument Transformers- Part 6: Requirements for Protective Current Transformers for Transient Performance, IEC 60044-6, 1992. [14] E. Lesniewska and W. Jalmuzny, “Influence of the number of core air gaps on transient state parameters of TPZ class protective current transformers,” Inst. Eng. Technol. Sci., Meas. Technol., pp. 105–112, Mar. 2009. [15] “Application Manual of REL521,” Jul. 2001. [Online]. Available: www.abb.com, Doc. ID: 1MRK 506 111-UEN
[16] Burdens & current transformer requirements of micom relays. Appl. Notes, Publ. Code: B&CT/EN AP/A11, 2007. [Online]. Available: www.areva-td.com [17] “ABB Outdoor Instrument Transformers Application Guide,” 2.1 ed. 2005. [Online]. Available: www.abb.com [18] H. Gremmel, Switchgear Manual, 10th ed. Berlin, Germany: ABB, 2001. [19] D. K. Cheng, Field and Wave Electromagnetics. Reading, MA: Addison-Wesley, 1983. [20] Working Group C-5 of the Systems Protection Subcommittee of the IEEE Power System Relaying Committee, “Mathematical models for current, voltage, and coupling capacitor voltage transformers,” IEEE Trans. Power Del., vol. 15, no. 1, pp. 62–72, Jan. 2000. [21] Cogent Power Ltd., Technical Document of Unisil Electrical Steel Grain Oriented. 2002. [Online]. Available: www.cogent-power.com [22] P. K. Gangadharan, T. S. Sidhu, and G. J. Finlayson, “Current transformer dimensioning for numerical protection relays,” IEEE Trans. Power Del., vol. 22, no. 1, pp. 108–115, Jan. 2007. [23] SIPROTEC Numerical Protection Relays Technical Document of Relays, Siemens Co., 2003. [24] CIGRE Working Group B5.02, “Co-ordination of relays and conventional current transformers,” 2003. [25] M. Sanaye-Pasand, M. R. Dadashzadeh, and M. Khodayar, “Limitation of transmission line switching overvoltages using switchsync relays,” presented at the IPST05, Montreal, QC, Canada, Jun. 2005. [26] “Controlled Switching Application Guide,” 1st ed. ABB, 2004. [27] IEEE Guide for Automatic Reclosing of Line Circuit Breakers for AC Distribution and Transmission Lines, IEEE Standard C37.10-2002. [28] H. Khorashadi-Zadeh and Z. Li, “Design of a novel phasor measurement unit-based transmission line auto reclosing scheme,” Inst. Eng. Technol., Gen., Transm. Distrib., pp. 806–813, Apr. 2011. [29] “Protection Application Handbook,” ABB Co., 1999, Rev. 0. [30] J. Esztergalyos, J. Andrichak, D. H. Colwell, D. C. Dawson, J. A. Jodice, T. J. Murray, K. K. Mustaphi, G. R. Nail, A. Politis, J. W. Pope, G. D. Rockefeller, G. P. Stranne, D. Tziouvaras, and E. O. Schweitzer, “Single phase tripping and auto reclosing of transmission lines,” IEEE Trans. Power Del., vol. 7, no. 1, pp. 182–192, Jan. 1992. [31] “User’s Guide of DIgSILENT 14.0 software,” DIgSILENT company, Germany, 2010. [Online]. Available: www.digsilent.de Mahdi Davarpanah (S’12) received the B.Sc. degree in electrical engineering from the Power and Water University of Technology, Tehran, Iran, in 2002, the M.Sc. degree from The University of Tehran, Tehran, in 2005, and is currently pursuing the Ph.D. degree at the University of Tehran. He is a Visiting Scientist at the University of Toronto, Toronto, ON, Canada. His research interests include power system protection, control, and transients.
Majid Sanaye-Pasand (M’98–SM’05) received the B.Sc. degree in electrical engineering from The University of Tehran, Tehran, Iran, in 1988, and the M.Sc. and Ph.D. degrees in electrical engineering from The University of Calgary, Calgary, AB, Canada, in 1994 and 1998, respectively. Currently, he is a Professor with the School of Electrical and Computer Engineering, University of Tehran, Tehran. His research interests include power system protection, control, and transients.
