X T A P (eXpandable Transient Analysis Program)
Ver. 2.41

Test Cases of XTAPCRIEPI - Central Research Institute of Electric Power Industry
July 2023

eXpandable Transient Analysis Program
XTAP Ver. 2.41

Test Cases of XTAPCentral Research Institute of Electric Power Industry - CRIEPITest Cases of XTAP
1. Basic Test Cases
EDU-01-A

PU Calculation - 1 ....................................................................................... 1

EDU-01-B

PU Calculation - 2 ....................................................................................... 5

EDU-01-C

PU Calculation - 3 ....................................................................................... 9

EDU-01-D

PU Calculation - 4 ..................................................................................... 13

EDU-01-E

PU Calculation - 5 ..................................................................................... 18

EDU-02

Calculation of Three-Phase Circuit PQVI ................................................. 22

EDU-03Comparison with Y Method, which is an Effective Value Analysis MethodLoad P Attached ....................................................................................... 26

EDU-04Comparison with Y Method, which is an Effective Value Analysis Method– Change in Impedance Z ........................................................................ 30

2. Power Electronics and Control System
PE-01

Simulation of Three-phase PWM Inverter Circuit .................................... 34

DLL-01

Basic Operations using DLL Control Block Components ........................ 41

3. HVDC, FACTS
HVDC-01

Separately Excited HVDC Model ............................................................. 45

HVDC-02

Self-Excited HVDC Model ....................................................................... 55

HVDC-03

Self-Excited HVDC Model (Half Bridge MMC) ...................................... 63

HVDC-04

Self-Excited HVDC Model (Full Bridge MMC) ....................................... 71

FACTS-01

Simulation Start of STATCOM ................................................................. 78

FACTS-02

STATCOM Model ..................................................................................... 83

FACTS-03

Separately Excited SVC Model ................................................................. 91

4. Power Quality Analysis
PQ-01

Simulation of Flicker Suppression Using Separately Excited SVC .......... 96

PQ-02-A

Calculation of Excitation Inrush Current under Pressurization of Customer- i -Transformer ............................................................................................ 101

5. Protection Relay
RELAY-01

Distance Relay Modeling ........................................................................ 108

RELAY-02

Coordinated Study of the Fault Current Interrupting Arcing Horn andTransmission Line Protection Relay ...................................................... 117

6. Machine
MAC-01-A

Synchronous Generator (Single-Line 3 LG-O Fault and Constant Excitation
Occurring in a Single-Machine Infinite Bus System) .......................... 129

MAC-01-B

Synchronous Generator (Single-Line 3 LG-O Fault Occurring in a SingleMachine Infinite Bus System with AVR) ............................................. 140

MAC-01-C

Synchronous Generator (Torsional Torque Simulation) ......................... 144

MAC-01-D

Synchronous Generator (Considering Magnetic Saturation) (Single-Line 3
LG-O Fault Occurring in a Single-Machine Infinite Bus System with AVR
Control) ................................................................................................ 150

MAC-02

3 LG-O Simulation of Induction Motor ................................................ 154

7. Power System Analysis
PS-01-A

Simulation of the WEST 10 System of the Institute of Electrical Engineersof Japan .................................................................................................. 159

PS-02-A

Simulation of EAST 10-Machine Power System of the Institute of ElectricalEngineers of Japan (IEEJ) ...................................................................... 165

8. Circuit Breaker
CB-01

Current Breaking Simulation of 66-Kv Fault Current Interrupting Arcing
Horn ....................................................................................................... 170

CB-02

Analysis of Terminal Short-Circuit Breaker and Capacitive Current
Switching in 154-Kv System ................................................................. 177- ii -

9. Switching Surge
SLD-01

Flashover Analysis of 6.6-kV Distribution Line...................................... 187

SLD-02

Lightning Surge Aspects Occurring in Secondary Side of a Pole-mounted
Transformer During Lightning Strike on a Distribution Line ................ 197

SLS-01

Lightning Surge Calculation of 500-kVAir-insulated Substation ........... 205

SSW-01

Switching Surge Calculation of a 500-Kv Vertically Arranged DoubleCircuit Transmission Line ...................................................................... 213

SSW-02

Switching surge Analysis of a 275-kV System........................................ 221

SSW-03

Analysis of Input Surge Owing to Block Energization during Large-scale
System Restoration ................................................................................ 238

SSW-04Comparison of the Ground-fault Surge Analysis of a 500-kV Power
Transmission Line with the Actual Measurements ................................ 247

SSWD-01

Ground-fault Surge Propagation Analysis of a 6.6-kV Power Distribution
Line ........................................................................................................ 261

STL-01

Simulation of Surge Propagation of a 500-kV Horizontally-arranged Single
Circuit Transmission Line ...................................................................... 272

STL-02

Simulation of Surge Propagation in a 500-kV Vertically Arranged Doublecircuit Transmission Line ....................................................................... 279

STL-03

Flashover Calculation using Arcing Horn Flashover Model for Short-tailWaves (Detailed Model, AC Superimposition not Possible) ................. 287

STL-04Flashover Calculation Using Arcing Horn Flashover Model (Simplified
Model, AC Superimposition Possible) for Short Tail Waves. ................ 297

10. Photovoltaic Systems
PV-01

Simulation of Single-phase Inverter Circuit for Solar Power Generation................................................................................................................ 304

PV-02-A

Simulation of a system fault using a basic transient analysis model of PCS
for utility-scale solar power station (Switching model) ......................... 309

PV-02-B

Simulation of a system fault using a basic transient analysis model of PCS
for utility-scale solar power station (average value model) ................... 317- iii -Examples of XTAPNo.

EDU-01-A

ExamplePU Calculation - 1

Name
Field

Basic calculation (power systems)
Japan

Literature

Electric

Association,

NEC

Technical

Standards

Committee:“Gridinterconnection code JEAC 9701-2006,” p. 196, 4th edition of issue dated June 30,
2006
In the Y method, which is effectively applied to a value analysis, the system
calculation is conducted in the PU system. In the present example, the checking method
using the PU system through an instantaneous value analysis is shown. The ultimate
aim is the target learning regarding the relationship of five cases, EDU-01-A to 01-E,
and the parasitic-/system-based PU using actual examples.

Overview

The following system is the specific target of this study.
Power - impedance - transformer - impedance - 3 LG fault
・In the PU system, the capacity and voltage are specified. Although the current can becalculated from the relationship between the two, the specific lesson learned here is that
this is a short circuit current (when Z = 1 pu).
Short-circuit current: Power - impedance - 3LG current at the time of the fault!
Through this, we also learned that the PU is extremely useful for a system check.

1Analysis Circuit and Conditions
Figure 1 shows the analysis circuit.

Ideal voltage
source
Close

three

SWs and go to
3 LG

Z≒0
Z≠0

Z≒0

Figure 1 Analysis circuit[System voltage]
Upper system: Line voltage of 66 [kV].

Distribution system: Line voltage of 6.6 [kV]Standard Z = V ^ 2 [kV] / S [MVA][Calculation of R and L]
Zbase = 6.6^2/1000 = 0.0436 [Ω] = 0.138655 [mH]. This value of Z corresponds to 1 pu
① In the literature, this is written as R + jX = 2.80 + j 7.50% (10 MVA).
② This becomes R+ jX = 2.80 + j7.50% (10 MVA) = 0.028 + j0.075 pu (10 MVA) = 2.80 pu + j7.50 pu(1000 MVA).
Here, a 1000 MVA base is generally used in the trunk transmission system. The line impedanceof the 6.6-kV system becomes a large value when we consider the trunk transmission system (※). This
image is important. A system fault on the 6.6-kV side becomes distant when viewed from the trunk
transmission system. This is in line with the image. If this image is viewed with the impedance (pu)
specifically, it becomes the image of ※.
③ R + jX = 2.80 pu + j7.50 pu (1000 MVA) = 0.12208 [Ω] + j1.0399 [mH] These are the input data.
2.80pu×0.0436 [Ω]7.50pu×0.138655 [mH]2

Branch current [A]Analysis Results

6.6 [kV] side

Time [s]Three-phase current waveform [A]Because I = V / Z, the short-circuit current at 3 LG is calculated using I = 1 [pu] / Z [pu].
Here, Z[pu] ＝ R + jX = 2.80 + j7.50 [pu] (1000 MVA) → |Z| = 8.0 [pu]At the time of 3 LG, the short circuit current I = V/Z = 1 [pu] / 8 [pu] = 0.125 [pu] (1000 MVA, 6.6
kV)
The current of 1 [pu] of the 1000 MVA based on the 6.6-kV side is
I = S/V = {(1000＊10^6)/3} / {6600/sqrt(3) } = 87477.31 [rms, A] = 123711.6 [peak, A],
because S = V * I from the relationship I = S / V
The short circuit current is calculated as I = 0.125 [pu] = 15463.95 [peak, A].

It was checked whether the line constant was set in accordance with the assumptions.
→ PU is very useful for a system checkEnd

3

Change log
Date

Example fileContent changes

version2014/11/19

2.0

2013/10/02

1.3

2012/07/19

1.2

Modified for XTAP Version 2.00
Example name change to EDU-01-A from EDU-01 consequent
to the addition of the pu calculation example
Modified for XTAP Version 1.20
Modified for XTAP Version 1.11

2011/10/18

1.1

Voltage probe and current probe that had been used for the useof control output changed for the use of XPLT output

2010/09/02

1.0First edition created (for XTAP Version 1.10)

4

Examples of XTAPNo.

EDU-01-B

ExamplePU Calculation - 2

Name
Field

Basic calculation (power systems)
Japan

Literature

Electric

Association,

NEC

Technical

Standards

Committee:“Gridinterconnection code JEAC 9701-2006,” p. 196, 4th edition of issue dated June 30,
2006
In the Y method, which is effectively used in a value analysis, the system calculationis conducted in the PU system. In the present example, the checking method using thePU system through an instantaneous value analysis is shown. The ultimate aim is the

Overview

target learning regarding the relationship of five cases of EDU-01-A to 01-E andparasitic-/system-based PU using actual examples.
The following system is the specific target of this study.
Power - impedance - transformer - impedance - 3 LG fault
・The present example focuses on transformers.

5Analysis Circuit and Conditions
Figure 1 shows the analysis circuit.

Circuit breaker+ SW1aIdeal voltage
sourceCircuit breaker+ SW1b

Close

three

SWs and go to
3 LG

Z≠0

Z≒0

Circuit breaker+ SW1cZ≒0

Because a GUI is parasitic-capacitance based rather than in mH,
the voltage was set to 7.5% of the 10 MVA transformer capacity base

Figure 1 Analysis circuit[System voltage]
Upper system: Line voltage of 66 [kV].

Distribution system: Line voltage of 6.6 [kV]Standard Z = V ^ 2 [kV] / S [MVA][Calculation of transformer leakage Z: Considering 66-kV primary side]
Zbase = 66^2/1000 = 4.36 [Ω] = 13.8655 [mH] This Z value corresponds to 1 pu
① In the literature, this is written as % Z = 7.5% (10 MVA).
② Because a GUI is parasitic-capacitance based rather than in mH, the voltage was set to 7.5% of the
10 MVA transformer capacity baseSince %Z = 7.5% (10 MVA) = 7.50 pu (1000 MVA),

7.50pu×13.8655 [mH]R + jX = 0.0pu + j7.50 pu (1000 MVA) = 0.0 [Ω] + j103.99 [mH]The same analysis result is obtained even if the leakage Z of
the transformer is set to the minimum value (for example,
0.0001%) and the Lback is set to 103.99 mH.

6

Branch current [A]Analysis Results

6.6 [kV] side

Time [s]Three-phase current waveform [A]The system voltage is 1 [pu] of the line voltage 66 [kV]. Here, only the 66-kV side of the transformer is set
up.
Because I = V/Z , the short-circuit current at the time of 3 LG is calculated using I = 1 [pu] / Z [pu].
Here, Z [pu] ＝ R + jX = 0.0 + j7.50 [pu] (1000 MVA) → |Z| = 7.5 [pu]Short-circuit current at 3 LG I = V/Z = 1[pu] / 7.5 [pu] = 0.133 [pu] (1000 MVA, 66 kV)
The current of 1 [pu] of the 1000 MVA base on the 66 kV side is
I = S/V = { (1000＊10^6)/3 } / {66000/sqrt(3) } = 8747.731 [rms, A] = 12371.16 [peak, A],
because S = V * I from the relationship I = S / V.
The short circuit current is calculated as I = 0.133 [pu] = 1645.36 [peak, A].
It was checked whether the line constant was set in accordance with the assumptions.
→ PU is very useful for a system check.
End

7

Change log
Date

Example fileContent changes

version2014/11/19

2.0

2013/10/02

1.3

2012/07/19

1.2

Modified for XTAP Version 2.00
Example name change to EDU - 01 - B from EDU - 02
consequent to the addition of the pu calculation example
Modified for XTAP Version 1.20
Modified for XTAP Version 1.11

2011/10/18

1.1

Voltage probe and current probe that had been used for theapplication of the control output was changed for the use of the
XPLT output

2010/09/02

1.0First edition created (for XTAP Version 1.10)

8

Examples of XTAPNo.

EDU-01-C

ExampleName

PU calculation - 3

Field

Basic calculation (power systems)
Japan

Literature

Electric

Association,

NEC

Technical

Standards

Committee:“Gridinterconnection code JEAC 9701-2006,” p. 196, 4th edition of issue dated June 30,
2006
In the Y method, which is an effective value analysis method, the system calculationis conducted in the PU system. In the present example, the checking method using thePU system through an instantaneous value analysis is shown. The ultimate aim is thetarget learning regarding the relationship of the five cases of EDU-01-A to 01-E and theparasitic-/system-based PU using actual examples.
The following system is the specific target of the study.
Power - impedance - transformer - impedance - 3 LG fault
・In the PU, the capacitance and voltage are specified. Although the current can becalculated from the relationship between the two, the specific lesson learned here is that
this is a short-circuit current (when Z = 1 pu).

Overview

Short-circuit current: Power - impedance - 3LG current at the time of the fault!
What we would like to prove through this example is that “the short-circuit current
does not change with the pu.” The advantage of a PU is that it can be used uniformly
irrespective of changes in the voltage class, and this fact has been widely
acknowledged in several publications. In other words, the base Z [Ω] changes with achange in voltage class, and hence the ampere of the short circuit current changes;
however, it remains the same in the PU, which is the greatest merit of the PU system.
That is, the short-circuit current [pu] does not depend on the voltage class in the PU
system, and it is therefore very convenient to measure the strength (1/X) of the line!

9Analysis Circuit and Conditions
Figure 1 shows the analysis circuit.

Ideal voltage
source

Close

SWs and go to

Z≠0

3 LG

Z≠0
Z≠0

Figure 1 Analysis circuit[System voltage]
Upper system: Line voltage of 66 [kV].

threeDistribution system: Line voltage of 6.6 [kV]Standard Z = V ^ 2 [kV] / S [MVA]10

Branch current [A]Analysis Results

6.6 [kV] side

Branch current [A]Time [s]6.6 [kV] side

Time [s]Three-phase current waveform [A]Because I = V / Z, the short-circuit current at 3 LG is calculated using I = 1 [pu] / Z [pu].
Here, Z[pu] ＝ (0 + j2.5) + (0.0 + j7.5) + (2.80 + j7.50) [pu] (1000 MVA) = 2.80 + j17.50 → |Z| =17.72[pu]At the time of 3 LG, the short circuit current I = V/Z = 1[pu] / 17.72 [pu] = 0.0564 [pu] (1000 MVA,
66 kV)
The current of 1 [pu] of the 1000 MVA base on the 66 kV side is
I = S/V = { (1000 * 10^6) / 3 } / {66000/sqrt(3) } = 8747.731 [rms, A] = 12371.16 [peak, A]
because S = V * I from the relationship I = S / V
The short-circuit current is calculated as I = 0.0564 [pu] = 697.73 [peak, A]

Calculate on your own

In contrast, the current of 1 [pu] based on 1000 MVA on the 6.6 kV side isand try to understandthe PU system

I = S/V = { (1000 * 10^6) / 3 } / {6600 / sqrt(3) } = 87477.31 [rms, A] = 123711.6 [peak, A],since S = V * I from the relationship I = S / V.

The short-circuit current on the 6.6-kV low-voltage side is calculated as I = 0.0564 [pu] = 6977.3 [peak, A].
Moreover, the voltage class comes into view when converting the value of the secondary side to theParasitic

primary side (MKS system).
This appears when converting the value of the primary side to the secondary side.

base

be

converted to the systembase

In the analysis of the PU system, the voltage class becomes irrelevant on both the primary and secondarysides, and the effective and instantaneous value analyses are the same. Consequently, the relaxed
responses of the fundamental waves in the results of the effective and instantaneous value analyses

11

must

must coincide. If this does not occur, then either the modeling or the PU system is considered to be flawed.

End

Change log
Date

Example fileContent changes

version2014/11/19

2.0

2013/10/02

1.3

2012/07/19

1.2

Modified for XTAP Version 2.00
Example name changed to EDU- 01-C from EDU-03
consequent to the addition of the pu calculation example
Modified for XTAP Version 1.20
Modified for XTAP Version 1.11

2011/10/18

1.1

Voltage and current probes used for the change in control
output for use of XPLT output

2010/09/02

1.0First edition created (for XTAP Version 1.10)

12

Examples of XTAPNo.

EDU-01-D

ExamplePU Calculation - 4

Name
Field

Basic calculation (power systems)
Japan

Literature

Electric

Association,

NEC

Technical

Standards

Committee:“Gridinterconnection code JEAC 9701-2006,” p. 196, 4th edition of issue dated June 30,
2006
Regarding the PU, (a) the conversion of the secondary side impedance to the primaryside (when the voltage classes are different) and (b) the parasitic capacitance andsystem bases (when the standard capacities are different) have been dealt with in detailin various publications. The aim of this example is to learn (a) above using actual
examples.
The following systems are the specific targets of the present study.
Power - impedance - transformer - impedance - 3 LG fault

Overview

・It is mentioned at the end of EDU-01-C that the “voltage class” comes into view
when converting the ○ side to the △ side. The correct expression is converting theamount in the MKS system of the ○ side to that of the △ side. The important aspect
here is that, having conducted the study using an equivalent circuit of the transformer,
the term a ^ 2 * L is used when converting the series circuit L on the secondary side to
the primary side. When converting the parallel circuit M on the secondary side to theprimary side, the term a ^ 2 * M is used. However, this has nothing to do with the pu
system.

13

Analysis Circuit and Conditions

Sw1b transformer

Circuit breaker+ Sw1bCircuit breaker+ Sw1cSw1a transformer

Circuit breaker+ Sw1aConversion here

Sw1a transformer

Figure 1 shows the analysis circuit.

Converts inductance L of 6.6-kV side to 66-kV side

Figure 1 Analysis circuit

Figure (a) shows the inductance L on the 6.6-kV side surrounded by the dotted line.
Figure (b) shows a circuit in which the inductance L is converted into the 66-kV side.

In Figures (a) and (b), the currents passing through the inductance L surrounded by the solid line arecompared. Because only the inductance L is converted, the waveforms are the same.

14

Branch current [A]Analysis Results

Time [s]The graph was plotted by superimposing the current of phase A on the 66-kV side. This indicates that theconversion was carried out satisfactorily. The method of conversion is as follows.
The inductance at 6.6 and 66 kV increases 10- or 1/10-fold. Because the inductance is a^2*L, as in thecase of a transformer conversion, it should be increased 100-fold.
L on the 6.6-kV side ＝ 1.0399 [mH]

L on the 66-kV side ＝ 103.99 [mH]

⇒ It has been ascertained that the amount on the ○ side in the MKS system has been converted
to that of the △ side. There is no reference to the pu.

The discussion regarding the pu is as follows.
Let us look at an expression such as “Let the leakage reactance of the transformer be 10%.” Regardlessof whether X is placed on the primary or secondary side, X = 0.10 pu. This occurs because the expression
essentially means “there is a 10% voltage drop when a rated current flows through the transformer.”
Instead, suppose there is a specification stating that the leakage reactance is 10% at the transformer
parasitic capacitance base (10 MVA). Suppose we try to convert X using this as the data of the analysis
program of the system base, namely, 1000 MVA. It was previously stated in the example of EDU-01-A that“R + jX = 0.028 + j 0.075 pu (10 MVA) = 2.80 pu + j 7.50 pu (1000 MVA).” It was also mentioned that“the line impedance of the 6.6-kV system becomes a large value when we consider the trunk transmission
system.” Taking this into consideration, we have X = 10% (10 MVA) = 0.1 pu (10 MVA) = 10 pu (1000
MVA) (※).
We will now give a summary of the parasitic and system bases in the impedance of the device. The
generator rear X is parasitic-capacitance based and is set to 1.8 pu, for example. When connecting thesynchronous generator to the power system, it is converted to the system base and to the system itself (seethe figure below). As an image, the large-capacitance generator strengthens the system. Although this has
the same value as the parasitic-capacitance based pu value, it means that, as the capacitance increases, X

15

becomes smaller when converted into the system. This is the same as an image with improved stability. It issufficient to consider the fact that, “even if the pu value that is based on the parasitic capacitance is thesame, the impedance when converted into the system becomes larger as the capacitance becomes smaller,”and remembering the formula is not quite essential.

Generator
発電機
Generator
背後X
rear X

無限大
Infinite bus
母線

背後電圧E
Rear voltage E

FigureConnection of synchronous generator to thepower system

End

16

Change log
Date

Example fileContent changes

version2014/11/19

2.0

Modified for XTAP Version 2.00

2014/07/29

1.0First edition created (for XTAP Version 1.21)

17

Examples of XTAPNo.

EDU-01-E

ExamplePU Calculation - 5

Name
Field

Basic calculation (power systems)
Japan

Literature

Electric

Association,

NEC

Technical

Standards

Committee:“Gridinterconnection code JEAC 9701-2006,” p. 196, 4th edition of issue dated June 30,
2006
Regarding the PU, (a) the conversion of the secondary side impedance to the primaryside (when the voltage classes are different) and (b) the parasitic capacity and systembases (when the standard capacities are different) have been dealt with in detail inprevious publications. The aim of this example is to learn (b) above using actual
examples.
The following system is the specific target of the study.

Overview

Power - impedance - transformer - impedance - connecting the load
・The target circuit is a EDU-01-C circuit connected with a load without a systemfault.
・ The impedance is described in EDU-01-B. The difference between the system andparasitic bases will be described through a graphical representation of theactive/reactive power P-Q and current I in the pu system. It is assumed that the example
of a “PQVI calculation of a three-phase circuit” has been implemented.

18

Analysis Circuit and Conditions
Figure 1 shows the analysis circuit.

Pu
systemFigure 1 Analysis circuit

A resistive load of 1 MW is attached to the 6.6-kV system. The boxed area is subjected to the pu system.
The analysis is conducted under a steady state, and the difference between the system base and the parasitic
base is explained through the creation of a graphical representation of the active/reactive power P-Q andcurrent I in the pu system.

19

Analysis Results
If the system base is the same, Ipu remains the same regardless of thechange in voltage class, which is the advantage of the pu.

Ipu(66 kV,1000 MVA)
Control noControl no

Ipu(6.6 kV,1 MVA)

Ipu(6.6 kV,1000 MVA)
Time [s]Time [s]Because the load is 1 MW, this is naturally because the load is 1 MW, this is naturally 1.0 pu in
0.001 pu in the 1000 MVA base.

the 1 MVA basis. Here, V = 1 pu.

Ppu(66 kV,1000 MVA)

Ppu(6.6 kV,1 MVA)
Control noControl no

If the system base is the same, Ppu remains
the same regardless of the change in voltage
class.

Ppu(6.6 kV,1000 MVA)
Time [s]Time [s]The same meaning is implied.

The same meaning is implied.

When the voltage is converted into the pu, despite the change in voltage
class, all the values are 1 pu in the case of the rated voltage, regardless of thechange in the system base.

Control no

Vpu(66 kV,1000 MVA)
Vpu(6.6 kV,1000 MVA)
Time [s]Vpu(6.6 kV,1 MVA)

It can be seen that even in the pu system, the difference in the A phase voltage occurs
when an active power flows in the transmission line of the inductance owing to therelationship
P = (V 1 * V 2 / X) * sin δ.
Control no

Vapu(66 kV,1000 MVA)
Vapu(6.6 kV,1000 MVA)
Time [s]20

EndChange log
Date

Example fileContent changes

version2014/11/19

2.0

Modified for XTAP Version 2.00

2012/11/15

1.0First edition created (for XTAP Version 1.21)

21

Examples of XTAPNo.

EDU-02

ExampleCalculation of Three-Phase Circuit PQVI

Name
Field

Basic calculation (power systems)
Kenzi, Noda, Takenaka: "Development of Instantaneous Value Simulation Method for

Literature

Induction Motors for Transient Phenomena Analysis,” CRIEP Report R 08022, June2009
It is a well-known fact that a PQ calculation is conducted using a VI bar in the Y
method, which is effectively applied to a value analysis. In this example, an actual
example of a PQVI calculation using the instantaneous value of the voltage/current of
the three phases is shown. In particular, we show the calculation method in both thevolt ampere system and the pu system. The goal is also to become familiar with a dq
conversion (not actual, please refer to the study above) by demonstrating the
calculation method using a αβ conversion.
The following calculation is demonstrated.
１．Units [V], [A], [W], [var](1) Three-phase voltage · current → PQ calculation

Variable names: PP31 and

QQ31

Overview(2) Calculation using the 90° lag information mentioned above

Variable names:

PP 32 and QQ 32(3) abc coordinates → calculation of αβ coordinates of voltage and current
Variable names: LAGVal, LAGVbe, LAGIal, and LagIbe(4) Three-phase voltage ·current → VI calculation

Variable names:

LAGVsimp and LAGIsimp(5) abc coordinates → calculation of αβ coordinates of voltage and current → PQ
calculation

Variable names: PP and QQ

２．Units [pu](1) 1. pu system of (3) and (4): VsimPU, IsimPU, ValPU, VbePU, IalPU, and
IbePU(2) Calculation using amount of voltage in pu and the current mentioned above
Variable names: PPPU and QQPU

22

Analysis Circuit and Conditions
Figure 1 shows the analysis circuit.
Analysis using L and R as parameters

Zbase = 6.6^2/1000= 0.04356 [Ω]1= 0.138655 [mH]2

Ideal voltage
source

Close

three

SWs and go to
3 LG

Figure 1 Analysis circuit・ Case1: L = 0.138655 [mH] = 1.0 pu

R = 0.00004356 [Ω] = 0.001 pu （Figure 1）

・ Case2: L = 0.138655 [μH] = 0.001 pu R = 0.00004356 [kΩ] = 1.0 pu

Analysis result

Branch current [A]Branch current [A]・ Case 1

Time [s]Time [s]Node voltage [M]2Node voltage [M]Node voltage [M]1Time [s]P[W] waveform

Time [s]Q[var] waveform

23

Time [s]V[pu] and I[pu] waveforms

1 ). The voltage is slightly lower owing to
A DC component is superimposed on the current using 3 LG (○2 ). Active power P and reactive power Q fluctuate at 50 Hz owing to the influence
the voltage drop of Tr (○of the DC current. During the fault period, because load L is connected in parallel to the voltage source,
only Q is 1 pu, and P becomes zero following the attenuation of the DC component current (because thevalue of load L is set to 1 pu in the Z-base, namely, Q = 1 pu). Because the green waveform of Q alone iscalculated using the information of a 90° lag, the waveform is shifted by 90° (shown as a negative quantity
for easier understanding).

Node voltage [M]・Case 2Time [s]PQ waveform [pu]During the fault period, because the load R is connected in parallel to the voltage source, P alone is 1 pu,
and Q is zero. Because the only load on the circuit is R, the DC component is quickly attenuated and the
PQ waveform does not oscillate with the system frequency. Here, although the circuit load R is 1 pu, P does
not become exactly 1 pu because of the voltage drop owing to transformer leakage Z (although very small)and because R is slightly larger owing to the ON resistance of the circuit breaker.

A noteworthy point here is that displaying the pu makes for a clearer understanding, and is recommended
because the validity of the analysis result (namely, whether it matches the actual results) can also be easily
checked.

End24

Change log
Date

Example fileContent changes

version2014/11/19

2.0

Modified for XTAP Version 2.00

2013/10/02

1.3

2012/07/19

1.2

Modified for XTAP Version 1.20

2011/10/18

1.1

Modified for XTAP Version 1.11

2010/09/02

1.0First edition created (for XTAP Version 1.10)

Example name change to EDU-02 from EDU-04 consequent tothe addition of the pu calculation example

25

Examples of XTAPExampleNo.

Comparison with Y Method, which is an Effective Value Analysis Method - Load P

Name

Attached

Field

Basic calculation (power systems)

References

EDU-03None
In this example, a comparison of XTAP with the Y method, which is effectively

Overview

applied to a value analysis, is conducted using a simple circuit. The aim is tounderstand through an analysis the difference in voltage phase owing to an exchangeof P. This holds for both the generator and load.

26

Analysis Circuit and Conditions
Figure 1 shows the analysis circuit.
Circuit breaker
Swl(a)

Circuit breaker
Sw1 (b)Attach a loadusing SW

Circuit breaker
Sw1 (c)Figure 1 Analysis circuit
A load is normally floated from the zero-phase circuit using a Y - Δ transformer. For the sake ofsimplicity, in this example the load is attached using SW.

27

Node voltage [V]Branch voltage [V]Analysis Results

Time [s]Time [s]Phase A voltage [V] on the power and load sides

PQ [pu] of loadNode voltage phase angle (deg)Y Method Analysis Results

Voltagephase

difference[deg] on power
and load sides

Active component of transmission sidecurrent (PU) <10**-2>

Time(s)Active load power[pu]Time (s)It can be seen that the load generates a phase difference in the voltage on the power supply and load sides.
The change in the active power is clean owing to the resistive load.
End

28

Change log
Date

Example fileContent changes

version2014/11/19

2.0

Modified for XTAP Version 2.00

2013/10/02

1.3

2012/07/19

1.2

Modified for XTAP Version 1.20

2011/10/18

1.1

Modified for XTAP Version 1.11

2010/09/02

1.0First edition created (for XTAP Version 1.10)

Example name changed to EDU-03 from EDU-05 consequent
to the addition of the pu calculation example

29

Examples of XTAPExampleNo.

Comparison with Y Method, which is an Effective Value Analysis Method – Change in

Name

Impedance Z

Field

Basic calculation (power systems)

Literature

EDU-04None
In this example, a comparison of XTAP with the Y method, which is effectively
applied to a value analysis, is conducted using a simple circuit. The circuit conditions
are as follows and only the impedance is changed. The voltages at both ends remainunchanged (V1, V2, and δ do not change).

Overview

Power supply - impedance - change component ΔZ - power supply
The basic equation of the power transmission is P = V1 * V2/X * sin δ. It can be seenfrom this equation that the active power varies with a change in X. Here, X is a series
element with respect to the power system. Through this example, it was verified that inseries elements, P varies with a change in X.

30

Analysis Circuit and Conditions
Figure 1 shows the analysis circuit.

Series capacitor inserted into
system by opening SW.

Figure 1 Analysis circuit

Node voltage [V]Branch current [A]Analysis Results

Time [s]P[W], Q[Var] in the middle of the line

Time [s]Three-phase current [A] in the middle of the line

31Current active component on the receiving side

Current reactive component on the receiving side

Y Method Analysis Result

PQ [pu] in the
middle of thelineLine current

Time (s)Linecurrent[pu]Time (s)Because the impedance between the power supplies is small, as seen from the relationship P = V 1 * V 2* sin δ / X, P becomes larger after the capacitor is inserted. Further, because P is doubled without a change
in voltage, the active current doubles. It should be noted that the waveform for the instantaneous value isthe combination of the active and reactive components of the current. It was ascertained that the responseof the fundamental wave can be understood from the effective value analysis results, and in the
instantaneous value analysis, an analysis of the power systems including high-frequency components canbe conducted.
Because D = √ {(R / 2 L) 2 - 1 / (LC)} → D = {2 * 0.01890 / (2 * 0.000693)} 2 - 1 / (0.000693 *0.0292) from the equation of applied sinusoidal voltage to the series circuit R - L - C, we have D = -48696.
→ ω = 2πf = √ 48696, from which we have f = 35.12 [Hz] → T = 0.0284 [s]. This is the fluctuating
frequency component of the current, and in terms of PQ, commercial frequencies of 50 Hz ± 35 =approximately 15 and 85 Hz are obtained. The XTAP analysis results show that the 15 and 85 Hz
components were superimposed on the vibration of PQ.

End

32

Change log
Date

Example fileContent changes

version2014/11/19

2.0

Modified for XTAP Version 2.00

2013/10/02

1.3

2012/07/19

1.2

Modified for XTAP Version 1.20

2011/10/18

1.1

Modified for XTAP Version 1.11

2010/09/02

1.0First edition created (for XTAP Version 1.10)

Example name changed to EDU-04 from EDU-06 consequent
to the addition of the pu calculation example

33

XTAP Example Collection

No.

PE-01

ExampleSimulation of Three-phase PWM Inverter Circuit

Name
Field

Power electronics
Technical Report No. 761, “Simulation Technology of Power Electronics System,”

References

the Institute of Electrical Engineers of Japan

In the “Power Electronic System Simulation Technology Cooperative Study Group”
set up by the Industrial Application Division of the Institute of Electrical Engineers,
several benchmark circuits were created with the purpose of evaluating various
analytical programs. In “Benchmark I,” the first benchmark circuit, a three-phase
PWM inverter circuit is applied as a typical power electronic circuit, and its example

Outlineis presented herein.
This example inverter converts a direct current into a three-phase alternating currentof 100 Hz. By combining the control blocks, a PWM pattern corresponding to thethree-phase alternating current is generated, by which the six switches constituting thethree-phase bridge are switched on and off to convert the direct current into analternating current. The above switches and reflux diodes are simulated using
simplified characteristics based on the ON and OFF resistances.

34

Analysis Circuit and Conditions
Figure 1 shows the analysis circuit

s1

s3

s5
30 each

u
10 mH
150 V

iu

v
10 mH

s2

s4

w

iv
10 mH

iw

s6

Figure 1 Analysis circuit[Circuit Operation]
Apply a 150-V DC voltage generated by the DC voltage source on the left side of Figure 1 to the threephase bridge. By switching the six switches of the three-phase bridge, S1, S2, …, S6, ON and OFF according
to the PWM pattern corresponding to a three-phase alternating current, the applied DC voltage is converted
into a three-phase alternating current consisting of u, v, and w phases. However, the PWM patterncorresponding to the three-phase alternating current is generated using a separate control system. The above
three-phase voltage is applied to the load through a 10-mH reactor, which is a higher harmonic filter. The
load is a delta connection of three resistors of 30 [Generation of PWM Pattern]
A PWM pattern is generated using a triangle wave comparison method. As an example, we determine thePWM pattern to generate a sinusoidal wave of the u-phase in Figure 2. The u-phase sinusoidal wave, whichis the command value, is superimposed and plotted on a triangular wave of amplitude 1, called the carrier.
Because the amplitude of the u-phase sinusoidal wave is 0.8, the degree of modulation at this time is said tobe 0.8. When the sinusoidal wave is smaller than the triangular wave, a zero is output, and when thesinusoidal wave is larger than the triangular wave, the output is 1. This is the PWM pattern for the u-phasesinusoidal wave. The PWM pattern remains zero for a lengthy time at locations where the command value
is small, and remains 1 for a lengthy time where the command value is large. In Figure 2, to avoid
complications, the waveform is drawn by lowering the carrier frequency. However, because the carrier
frequency is set much higher than the frequency of the sine wave, the PWM pattern changes more
frequently, and the command value can be faithfully reproduced based on a time change of time ratio (dutyratio) 1.

351u-phase voltage
command value(Sine wave)Carrier wave
(Triangular wave)1
1

PWM Pattern(Upper arm)

Figure 2 Triangular wave comparison method

Switch S1 of the upper arm of the leg corresponding to the u-phase of the three-phase bridge turns ON when
the PWM pattern is 1, and is turned OFF when it is zero (note that one switch is called an arm, and upperand lower arms combine to form a leg). In contrast, switch S2 of the lower arm of the u-phase leg turns ON
when the PWM pattern is zero, and turns OFF when it is 1. However, to prevent the upper and lower armsof the same phase turning on at the same time and short-circuiting the DC power supply, the turn-on timing
is delayed by only a small time-delay Td between the switches in both arms, and the turn-off timing isadvanced by Td, which is called the dead time.

In this example, the parameters related to PWM pattern generation are set as follows:·

Carrier frequency:

5 kHz

·

Frequency of the sine wave:

100 Hz·Degree of modulation:

0.8·

Dead time:

5 s[Switching Element, Reflux Diode, and Snubber Circuit]
In an actual PWM inverter, MOSFETs and IGBTs are used as the switching elements. However, in thisexample, for the sake of simplicity, we conduct a simulation using simplified characteristics such as an ONresistance of 10 mand anOFF-resistance of 1 MThese characteristics are realized by connecting an
ideal switch (ON-resistance of zero, OFF-resistance of ) in-series using a diode with an ON-resistance of
10 mand an OFF-resistance 1 MIn the inverter circuit, the diode is connected anti-parallel to the switching element. This is called a reflux
diode. A reflux diode prevents the occurrence of an overvoltage by blocking the inductive load current. This
refluxing diode is also simulated using an ON-resistance of 10 mand an OFF-resistance of 1 M as
described above.

36

Normally, a protection circuit is provided to prevent the destruction of the switching element from an
overvoltage or overcurrent during the operation of the switching element. This protection circuit is called asnubber circuit. In this example, a CR snubber, which is a series circuit of a capacitor and a resistor, is used.
The capacitor value is set to 0.05 F, and the resistor value is set to 300 .

[Analysis Conditions]
The analysis conditions are as follows.
·

Calculation time step

1 s

·

Calculation start time

0 ms

·

Calculation end time

40 ms

·

Display start time

20 ms

·

Display end time

40 ms

[Example of XTAP Input]
An instance of creating this example using XTAP is shown in Figure 3.

37

Example Name: Simulation of three-phase PWM inverter circuit
No.: PE-01
This is a three-phase PWM inverter circuit of benchmark test 1 created through the “Joint Research Committee on Simulation Technology ofPower Electronics Systems” of the Institute of Electrical Engineers.
In this circuit, using the componentization function, the arms are componentized and used as upper and lower arms of the phases u, v, and w.

Operating conditions of theinverter
Carrier frequency: 5 kHz
Sine wave frequency: 100 Hz
Modulation degree: 0.8
Dead time: 5 s
Analysis conditions
Calculation time step: 1 s
Calculation start time: 0 ms
Calculation end time: 40 ms
Display start time: 20 ms
Display end time: 40 ms

Arm components
The configuration of the arm components can be
confirmed by opening a separate window using a “rightclick → edit componentization diagram.”

Control system(gate pulse output)

Main circuit(a)(a) 主回路(b) 制御系(b) Control system

Figure 3 Example of XTAP input

38

Analysis Results
The results of executing this example using XTAP are shown in Figure 4. Based on the results of thesimulation, it can be confirmed that a direct current is converted into an alternating current of 100 Hz using
an inverter and supplied to the load.

Case 1200

Potential difference between nodes [V]

150
100
50
0
-50
-100
-150
-200
20

21

22

23

24

25

26

27

28

29 30 31
Time [ms]

32

33

34

35

36

37

38

39

40

34

35

36

37

38

39

40

u(Node_u)-u(Node_v) #1(a) Line-to-line voltage between u and v

Case 1
5
4
Branch current [A]

3
2
1
0
-1
-2
-3
-4
-5
20

21

22

23

24

25

26

27

28

i(Lu) #1

29 30 31
Time [ms]i(Lv) #1(b) Load current

39

32

33

i(Lw) #1

Figure 4 Analysis results

END

Revision History
Date

Example fileContent changes

Version2014/11/19

2.0

Update for XTAP Version 2.00

2012/07/19

1.3

Update for XTAP Version 1.20
Because the parameter inside the delay block [delay_v2] in the

2011/10/18

1.2

control system block is a numerical value (5e-6), it waschanged to a variable (Td).

2011/02/14

1.1

Update for XTAP Version 1.11

2010/09/02

1.0

Creation of the first edition (for XTAP Version 1.10)

40

Examples of XTAPNo.

DLL-01

ExampleBasic Operations using DLL Control Block Components

Name
Field

―

Literature―This is an example verifying the basic use of the DLL control block components.

The DLL control block part is a control component used for the control logic, andwas originally described by users in C language or FORTRAN as a control block in
XTAP. The control logic is provided in a dynamic link library (DLL). The DLL
control block component receives input signals (a maximum of ten signals) at theOverview

time of the initial calculation and at the time of the dynamic characteristic
calculation (calculation of each time step), and the result of the operation defined in
the DLL is output as an output signal (a maximum of ten signals). In addition to the
input signal, five types of constants, P1 through P5, may also be specified (given as
parameters at the time the analysis is executed). For details such as the method used
to create the DLL programs, please refer to the Appendix “How to use DLL control
block parts” of the manual.

41Analysis Circuit and Conditions
Figure 1 shows the control block diagram. A DLL control block component is created as shown in Fig. 1,
which executes an operation that integrates input 1 to output 1, and the sum of P1 × input 1 and P2 × input
2 to output 2. The constants P1 and P2 are assumed to be given as parameters during the program execution.

DLL control component

Input 1

Output 1

Input 2

Output 2

Figure 1 Control block diagram[XTAP input]
Because this example has two inputs and two outputs, the third to the tenth input and output terminals are
not used. Therefore, a dummy input value is connected to the input side, and a termination output terminal
is connected to the output side. Further, because the output signal of the DLL control block component is
not directly recorded, a gain constant of 1 is applied, as shown in the figure below.

Input 1

Output 1

Input 2
Output 2

Because outputs 3–10 are not used,

Because inputs 3–10 are not used,a termination output terminal is connected

a dummy input is connected

Figure 2 XTAP input

42Analysis Results
The results of executing this example using XTAP are shown in Figure 3. The following values were
set in the current calculation example: input 1 = 1.0, input 2 = 2.0, P1 = 0.1, and P2 = 0.2. The simulation
results are in agreement with the following numerical values, which can be analytically obtained from theblock diagram of Figure 1.
t( Out1 ) = 1.0dt   t
0Control signal( Out2 ) = 1.0×0.1 + 2.0×0.2 = 0.5

Time [s]Figure 3 Analysis results

End

43

Change log
Date
2014/11/18

Example fileContent changes

version
2.0

First edition created (for XTAP Ver. 2.00)

44

Examples of XTAPNo.

HVDC-01

Name of
Separately Excited HVDC Model

ExampleField
LiteraturePower electronics“DC Power Transmission Engineering,” Tokyo Denki University Press, written andedited by Takahiko Machida
Because this model has basic control and protection systems, it can be applied toother simulations by copying it and changing the rated value.
Separately excited converters are currently used for various applications such as a DC
power transmission, frequency conversion, and BTB. In this example, the simulation
waveforms of the startup of the converter and the AC three-phase ground fault at a

Overview

frequency conversion station are shown. This may also be used for a DC transmission
by adding a DC line to this example.
The circuit of the converter consists of two six-phase Gretz wires connected in-series
in two stages. Because 11th- and 13th-order harmonics and their multiples are generated
in the system, to remove them, an AC filter is connected between the converter andsystem. Further, phase-modifying equipment is connected to compensate for thereactive power, which is consumed by a separately excited converter during operation.

45

Introduction · Using the Model
Refer to the table at the end for the parameters required to use the separately excited HVDC model.
Further, change “MASTER_Start” in “HVDC 1/MASTER” when changing the activation time of theseparately excited HVDC, and change “Pdorder” in “HVDC1/MASTER” when changing the active power
output.

Analysis Circuit and Conditions
Figure 1 shows the analysis circuit created using XTAP, and Figure 2 shows a schematic diagram of theseparately excited HVDC model.
60Hz

50Hz

66 kV

DC 50 kV

66 kV

Frequency
周波数
conversion
変換所
stationFigure 1 Analysis circuit

Separately
excited HVDC model
他励HVDCモデル

AC filter
交流フィルタ直流
DCreactorAC filter
交流フィルタリアクトル

Phase modifying
equipment

調相設備

Phase modifying
equipment

調相設備

AC
system交流系統

ACsystem
交流系統

Convertertransformer
変換器用変圧器

Converter circuit
変換器回路

Figure 2 Schematic diagram of model

The main circuit conditions of the converter are as follows:
Rated capacity of 50 MW, rated DC voltage of 50 kV, and rated DC current of 1000 A
Converter transformer rated voltage of 66/22 kV

46

Analysis Results
Figure 3 shows the voltage and current measurement points.

Transformer
primary
変換器sidecurrent

DC current

直流電流

1次側電流

DC current
系統電流
DC voltage (1)
直流電圧①

直流電圧②
DC
voltage (2)Figure 3 Voltage and current measurement points
Figures 4 and 5 show the DC current and DC voltage waveforms at the startup of the converter. After

Branch current [A]completion of the startup, both the current and voltage are rated values, and a startup is normally performed.

Time [s]Figure 4 DC current waveform at startup

47

Node voltage [V]Time [s]Figure 5 DC voltage waveform at startup (DC voltage ①)
Figure 6 shows the current flowing on the primary side of the transformer, and Figure 7 shows thecurrent flowing in the system. The current flowing through the transformer is close to a square wave, andthe direct current has a chopped waveform. This current, which contains large amounts of harmonics,

Branch current [A]passes through the AC filter where the harmonics are eliminated and the current becomes a sinewave.

Time [s]Figure 6 Transformer primary side current

48Branch current [A]Time [s]Figure 7 Current flowing through the system
Figure 8 shows the waveform of the control angle α at the time of the converter startup. From the figure,the control angle on the 60-Hz side is approximately 20° during a steady state, which is less than 90°,

Control No

indicating that the converter operates as a forward converter.

Time [s]Figure 8 Control angle α

Red line, 50-Hz side; blue line, 60-Hz side

The DC current waveform at the time of the three-phase ground fault on the AC side is shown in Figure
9, the DC voltage waveform is shown in Figure 10, and the control angle is shown in Figure 11. When thesystem detects a drop in AC voltage, the protection operation of the gate shift (GS) → gate block (GB) →restart (RST) is carried out. Because the GS angle is set to 100° in this simulation, the control angle of the49

converter becomes larger than 100° during the gate shift period (in this instance, 100°). Owing to the actionof the GS, the DC current becomes zero, and the converter stops temporarily but restarts after the AC fault
has been eliminated and the AC voltage has recovered, and both the voltage and current of the converter are

Branch current [A]operated again in accordance with the command value.

Time [s]Node voltage [V]Figure 9 Waveform of DC current

Time [s]Figure 10 Waveform of DC voltage

50

Control No.

Time [s]Figure 11 Control angle α

Red line, 50-Hz side; blue line, 60-Hz side

51

The parameters of the model and their respective descriptions are shown below. Note that in the settingsof terminal 2, the variable names are the same as those of terminal 1 except that the 1 at the end is replacedwith 2, and hence the details have been omitted here.
Name

Unit

Item

Rate_PW

Rated power

Rate_Vdc

V

Rated DC voltage

IdcMinp.u.

Minimum value of DC current

CurrentLimitp.u.

Maximum current command value at current limit

ReverseTimes

Power flow reversal speed

ReversePeriodsTime taken to widen the control angle limiter for power flow reversal

LimitOpenDelays

Time from BPP cancellation for startup to opening of control angle limiter

PulseWidthdeg

Ignition pulse width

SSRelayOffTimes

Relay dead time of at the start of simulation

APR_K

-

APR gain

APR_T1sFirst-order lag time constant for APR

APR_T2sFirst-order lead time constant for APR

Limit_APR-

APR limiter (upper limiter, lower limiter)

SystemFrequency1

Hz

System frequency

Rate_S_Trans1

VA

Capacitance of converter transformer (total of Y – Y and Y – D)

L_DCKpu1

H

Standard value of DC reactor inductance (cathode side)

L_DCApu1

H

Standard value of DC reactor inductance (anode side)

R_DCKpu1

Standard value of resistance of DC reactor (cathode side)

R_DCApu1

Standard value of resistance of DC reactor (anode side)

Rate_V1_ltol1

Vrms

Effective value of primary side line voltage of transformer (1)Rate_V2_ltol1

Vrms

Effective value of primary side line voltage of transformer (2)Trans_Rpu1

p.u.

Resistance of converter transformer

Trans_Lpu1

p.u.

Inductance of converter transformer

IdcBiasDeg1

deg

Current bias (approximate value of control angle during REC operation)

VdcBiasDeg1

deg

Voltage bias (approximate value of control angle during INV operation)

Vdc_Order_Lag1

sFirst-order lag time constant when DC voltage command is changed

Idc_Order_Lag1

sFirst-order lag time constant when DC current command is changed

IdcMesLag1

sFirst-order lag time constant for DC current measurement

VdcMesLag1

sFirst-order lag time constant for DC voltage measurement

ACR_K1-

ACR gain

ACR_T11sFirst-order lag time constant for ACR

ACR_T21sFirst-order lead time constant for ACR

AVR_K1-

AVR gain

AVR_T11sFirst-order lag time constant for AVR

AVR_T21sFirst-order lead time constant for ACR

GammaZeroDeg1

deg

Margin angle setting

StartDegree1

deg

Start angle

Start_TConst1sFirst-order lag time constant of current rise at startup

AlimitInvU1

deg

Upper limit of control angle (during INV operation)

AlimitRecU1

deg

Lower limit of control angle (during REC operation)

AlimitInvL1

deg

Lower limit of control angle (during INV operation)

AlimitRecL1

deg

Control angle lower limit (during REC operation)

LimitOpenTimeU1

sFirst-order lag time constant at change of upper limit of control angle

LimitOpenTimeL1

sFirst-order lag time constant at change of lower limit of control angle

52

VAMP_LAG1

sFirst-order lag time constant for AC voltage measurement

FREQ_LAG1sFirst-order lag time constant for frequency detection

K_CF1

-

Gain of commutation failure detection relay (the smaller the value, the higher the sensitivity)

K_DCFRD1-

Gain of DC fault detection relay (the smaller the value, the higher the sensitivity)

T_DCFRD1s

First-order lag time constant of DC fault detection relay

Threshold_UV1

p.u.

AC undervoltage relay threshold

Threshold_OV2

p.u.

AC overvoltage relay threshold

Threshold_OF1

p.u.

Frequency rising relay threshold

Threshold_UF1

p.u.

RecoveryTime1sFrequency drop relay threshold
Lag from fault relay cancellation to actual cancellation of protection operation

GSTime1s

GS execution time (time from start of GS to GB)

BPPTime1s

BPP execution time (time from start of BPP to GB)

GBTime1s

GB execution time (time from start of GB to CBT)

GSDegree1deg

GS angle

BFDegree1

deg

Beta advance angle

BF_IncLag1

sFirst-order lag time constant at the time of canceling beta advance

BF_DecLag1

sFirst-order lag time constant when carrying out beta advance earlier

BFTime1s

Beta advance period

Rate_V_PM1

Vrms

Rated voltage of transformer for phase modifier

R_Trans_GND1

Ground resistance of converter transformer (per unit)

Threshold_PM11

p.u.

Shunt reactor 1 on-off threshold

Threshold_PM21

p.u.

Shunt reactor 2 on-off threshold

Threshold_PM31

p.u.

Stacon 1 on-off threshold

Threshold_PM41p.u.

Stacon 2 on-off threshold

Threshold_PM51

p.u.

Stacon 3 on-off threshold

Threshold_PM61

p.u.

Stacon 4 on-off threshold

Hysteresis_PM1

p.u.

Spu_Tr_PM1

p.u.

Hysteresis amount of on/off phase adjustment
Rated capacity of transformer for phase modifier (denominator is the rated active power of converter)

Lpu_Tr_PM1

p.u.

S_Shant_pu1

p.u.

Capacity per shunt reactor (denominator is the rated active power of converter)

S_SC_pu1

p.u.

Capacity per stacon (denominator is the rated active power of converter)

R_Filter_GND1

Ground resistance of AC filter (per filter)

F_5_pu1

p.u.

Resonance order of fifth-order filter

Q_51

-

S_5_pu1

p.u.

Sharpness order of fifth-order filter
Fundamental wave capacity of fifth-order filter (denominator is the rated active power of converter)

F_11_pu1

p.u.

Resonance order of 11th-order filter

Q_111

-

S_11_pu1

p.u.

Sharpness order of 11th-order filter
Fundamental wave capacity of 11th-order filter (denominator is the rated active power of converter)

F_13_pu1

p.u.

Resonance order of 13th-order filter

Q_131

-

S_13_pu1

p.u.

Sharpness order of 13th-order filter
Fundamental wave capacity of 13th-order filter (denominator is the rated active power of converter)

F_HP_pu1

p.u.

Resonance order of higher-order filters

Q_HP1

-

S_HP_pu1

p.u.

Sharpness order of higher-order filters
Fundamental wave capacity of higher-order filters (denominator is the rated active power of converter)

Rsnubber_pu1-

R setting of RC snubber

Csnubber_pu1-

C setting of RC snubber

Rpressure_pu1-

Voltage divider resistor setting

Lanode_pu1-

Anode reactor setting

DeionTime1s

Thyristor deionization time (see Tdeion section of Thy in the manual)Inductance of transformer for phase modifier (denominator is the rated active power of converter)

53

End

Change log
Date
2018/6/29

Example fileContent changes

version
2.1

Icons of changed parts
Modified for XTAP Version 2.00

2014/11/19

2.0

Modified three-phase voltage source model (the infinite bus
generator component was changed to COS wave voltage
source)2013/04/15

1.3

Changes consequent to modifications in model diagram andparameter input format. Added initialization function.
Change consequent to modification of the separately excited

2012/08/29

1.2

HVDC model
Addition of waveform of 3-phase ground fault on AC side

2012/07/19

1.1

Modified for XTAP Version 1.20

2010/10/18

1.0First edition created (for XTAP Version 1.11)

54

Examples of XTAPNo.

HVDC-02

Name ofSelf-Excited HVDC Model

ExampleField
LiteraturePower electronics“DC Power Transmission Engineering,” Tokyo Denki University Press, written andedited by Takahiko Machida
Because this model has basic control and protection systems, it can be applied toother simulations by copying it and changing the rated value.
This example simulates a DC power transmission system circuit using a self-excitedconverter. In this example, a two-level converter that was also used in the STATCOM
example is connected back-to-back across a DC line, thereby constituting a DC powertransmission system. When compared with the separately excited converter, theOverview

self-excited converter has the following advantages: ① There is no need for phaseadjustment equipment because the converter is capable of outputting reactive power byitself. ② It can also be applied to relatively weak lines with a small short-circuit
capacity. In this example, the simulated waveforms at the time of the converter
activation, power flow reversal, and DC main line ground fault are shown.
Incidentally, if the DC line of this example is ignored, it will have a back-to-back(BTB) configuration and can be used for frequency-conversion purposes.

55Introduction/Model Use
Refer to the table at the end for the parameters required to use the separately excited HVDC model.
Further, “Pdorder” is changed in “HVDC/MASTER” when changing the active power output of theself-exciting HVDC.

Analysis Circuit and Conditions
Figure 1 shows the analysis circuit, and Figure 2 shows a schematic diagram of the model.

潮流反転Power flow reversal

Figure 1 Analysis circuit

Self-exciting HVDC model
自励HVDCモデル
Terminal
端子1
1

端子2

Terminal 2

Figure 2 Schematic diagram of model・Rated equipment capacity = 50 MW (55.9 MVA), primary side rated voltage = 66 kV, secondary side
rated voltage = 56 kV, DC voltage = 125 kV.
・The DC power transmission system is linked with the 66-kV system using the impedance of thesimulated short-circuit capacity of the system.
・ Here, example simulations of the ① startup, ② power flow reversal, and ③ DC main line ground
fault are carried out. The converter begins activation at a simulation time of t = 0.25 s. At t = 0.25 s, thecontrol system is activated, and power is transmitted from terminal 1 to terminal 2. The power command
value is changed at t = 0.5 s and a power flow reversal begins. The ground fault is eliminated by the circuit

56

breaker at t = 0.8 s.

Analysis Results
Refer to the STATCOM model example of FACTS-02 for operation of the converter at startup. Figures 3
and 4 show the voltage and current of the DC circuit during a power flow reversal. Figures 5 and 6 show
the active and reactive power output from terminals 1 and 2, respectively, at the time of a power flow
reversal. Because the reactive power control of each terminal during this simulation continuously controls
the voltage of the AC side (ACAVR), the reactive power is controlled such that the terminal voltage of theControl No.

converter approaches 1.0 pu.

Time [s]Blue line, terminal 1; red line, terminal 2 (the reference voltage is 125 kV)

Branch current [A]Figure 3 DC voltage [pu]Time [s]Figure 4 Direct current [A]57

Control Signal

Time [s]Figure 5 Active power [pu]Blue line, terminal 1; red line, terminal 2 the (reference capacitance is 55.9

Control No

MVA)

Time [s]Figure 6 Reactive power [pu] Blue line, terminal 1; red line, terminal 2 (the reference capacitance is 55.9
MVA)
At t = 0.5 s, the power command value switches from 45 MW (0.8 pu) to -45 MW (-0.8 pu), and thepower flow reversal begins. The rectifier side applies active power control and determines the power to besent to the other systems, whereas the inverter side applies DC voltage control and determines the DC
voltage. Owing to the voltage drop that occurs from the resistance of the DC line when sending power fromterminal 1 to terminal 2, the DC voltage of terminal 2 is 1.0 pu in keeping with the command value (1.0 pu).
However, the DC voltage of terminal 1 is higher than the command value. In contrast, when the power flowis reversed, the DC voltage of terminal 1 follows the command value, and the voltage of terminal 2 ishigher than the command value.
Figure 7 shows the DC voltage, and Figure 8 shows the system side output current at the time of a DC
fault. Because the occurrence of a DC fault at t = 0.8 s amounts to a short circuit of the energy storage
capacitor, the DC voltage becomes zero. Further, once the system detects a DC fault, the converter applies agate block (GB) → circuit breaker release operation, and therefore, despite a large current being output

58

from the converter during the fault, the output current of the system side becomes zero after the circuitControl No.

breaker is opened.

Time [s]Blue line, terminal 1; red line, terminal 2 (the reference voltage is 125 kV)

Branch current [A]Figure 7 DC voltage [pu]Time [s]Figure 8 System side output current [A]59

The parameters of the model and their respective descriptions are shown below. Note that, in the settings ofterminal 2, the variable names are the same as those of terminal 1 except that the 1 at the end is replaced

Common constants

with 2, and hence the details have been omitted here.
Name

Unit

Item

Rate_PW

Rated active power

Rate_Vdc

V

Rated DC voltage

SSRelayOffTimes

Relay dead time at start of simulation

SystemFrequency1

Hz

System frequency

Rate_S1

VA

Rated capacity

Rate_V1_ltol1

V

Effective value of transformer primary side line voltage

Rate_V2_ltol1

V

Effective value of transformer secondary side line voltage

PWMPulseNumber-

Number of PWM pulses (per cycle)

R_TransGND1Ω

Ground resistance of converter transformer

H1sElectrostatic constant

Flag_ACAVR1-

ACAVR flag (1 = used, 0 = not used), Select Flag_AQR

Flag_AQR1-

AQR flag (1 = used, 0 = not used), Select Flag_ ACAVR

Vac_order1

p.u.

AC voltage command value (ACAVR command value)

Q_order1

p.u.

Reactive power command value (AQR command value)

PWMPhase1

degPhase angle of PWM carrier

R_connect_pu1

p.u.

Converter transformer resistance

L_connect_pu1

p.u.

Converter transformer inductanceSpu_filter11

p.u.

Fundamental wave capacity of higher-order filter 1 (denominator is rated capacity of converter)

Q_filter11-

Sharpness of higher-order filter 1

Spu_filter21

p.u.

Fundamental wave capacity of higher-order filter 2 (denominator is rated capacity of converter)

Q_filter21-

Sharpness of higher-order filter 2

Gain_ACR1-

ACR gain

T_ACR1s

ACR time constant

Gain_APR1-

APR gain

T_APR1s

APR time constant

Lag_APR1sFirst-order lag time constant for APR

Gain_DCAVR1-DCAVR gain

T_DCAVR1s

DCAVR time constant

Gain_AQR1-

AQR gain

T_AQR1s

AQR time constant

Gain_ACAVR1-

ACAVR gain

T_ACAVR1s

ACAVR time constant

Vmag1

-

Voltage margin

Pmag1

-

Power margin

Lag_Margin1

s

First-order lag time constant at margin switching (switching speed)

IinvLimit1

p.u.

Maximum current of converter (denominator is the converter rating)

K_DCFRD1-

Gain of DC fault detection relay (the smaller the value, the higher the sensitivity)

T_DCFRD1s

First-order lag time constant of DC fault detection relay

Threshold_OC1

p.u.

Overcurrent relay threshold

Threshold_DCOV1

p.u.

DC overvoltage relay threshold

Threshold_DCUV1

p.u.

DC undervoltage relay threshold

Threshold_OV1

p.u.

AC overvoltage relay threshold60

Threshold_UV1

p.u.

AC undervoltage relay threshold

Threshold_OF1

p.u.

Frequency rise relay threshold

Threshold_UF1

p.u.

Frequency drop relay threshold

VAMP_LAG1sFirst-order lag time constant for AC voltage detection

FREQ_LAG1sFirst-order lag time constant for frequency detection

RecoveryTime1-

Lag from fault relay cancellation to actual cancellation of protection operation

GBTime1s

GB implementation time

NoDetectPeriod1sPartial relay dead time at startup

CalculationTime1sControl cycle

DeadTime1s

Dead time

Flag_DirectSample1-

Voltage, current amount direct detection flag (selected from among three types)

Flag_AverageSample1-

Detection flag using moving average of voltage and current amount (selected from among three

Flag_PeakTopSample1-

Peak top detection flag of voltage and current
types)amount (selected from among three types)

ChargeCircuitCloseTime1sStart time of charging (if starting from a transient state, assign a large negative value)

ChargeTime1s

Charge time

Control_StartDelay1sTime from end of charge to start of control system operation

End

61

Change log
Date

Example fileContent changes

version2018/6/29

2.2

Icons of parts changed

2018/1/26

2.1

Modified the creation logic of HOLD signal

2014/11/19

2.0

Modified for XTAP Version 2.00

2014/07/22

1.4

2013/04/15

1.3

Deletion of snubber capacitor of converter arm and correctionof certain mathematical formulas
Changes consequent to modifications of model diagram andparameter input format
Change consequent to modification of the self-exciting HVDC

2012/08/29

1.2

model
Addition of waveform of direct current main line ground fault

2012/07/19

1.1

Modified for XTAP Version 1.20

2010/10/18

1.0First edition created (for XTAP Version 1.11)

62

Examples of XTAPNo.

HVDC-03

ExampleSelf-Excited HVDC Model (Half Bridge MMC)

Name
Field

Power electronics(1) “DC Power Transmission Engineering,” Tokyo Denki University Press, written andedited by Takahiko Machida(2) T. Kikuma, K. Takenaka, M. Takasaki, “Control Protection Method of DC Power
Transmission System with Modular Multi-Level Converter Using Full Bridge Cell:

Literature

Proposal of a Control Method that can Control Fault Current During DC Fault,” Central
Research Institute of Electric Power Industry Research Report R11021, (2012).
(3) T. Kikuma et al., “Control Method of Double Y-connection Modular Multi-level
Converter,” Journal of the Institute of Electrical Engineers of Japan D, Vol. 133, No. 9,
pp. 917-927 (2013).
Because this model has basic control and protection systems, it can be applied toother simulations by copying it and changing the rated value.
This example simulates the circuit of a DC power transmission system using a

Overview

modular multilevel converter (MMC). It has the same configuration as the self-excited
HVDC (HVDC-02) example, except that the converter circuit is replaced by an MMC
circuit. In this example, the simulated waveforms at the time of a DC main line ground
fault and power flow reversal of the converter are shown.
Several MMC models have been created as examples.

63Simple Explanation of MMC
A brief explanation of an MMC is given below. For a more detailed explanation, please see thereferences. A converter composed of identical cells (modules) connected in multiple stages is called an
MMC. There are cells that use half bridge cells and those that use full bridge (single phase inverter) cells.
The present example deals with a half bridge MMC model. The problem with an MMC is that, as thenumber of cells increases, the number of elements in the circuit becomes enormous, leading to significant
calculation costs. To contend with this problem, an average model made by averaging the switching
operations has been used in certain cases. This example deals with two types of models:① Half bridge MMC averaging model (HVDC-03A)② Half bridge MMC detailed model (HVDC-03B)
In the absence of any specific reason to use the detailed model, the use of the average model isrecommended because it is much faster. The applied control, which is a combination of the controls used in
the references, is used as the control in an MMC.

Introduction/Model Use
Refer to the table at the end for the parameters required to use the MMC model. When changing theactive power output of the MMC, “Pdorder” in “MASTER_VSC 1,” which is located one level below thehighest level of the hierarchy, is changed.

Analysis Circuit and Conditions
Figure 1 shows the analysis circuit.

Power flow reversal

潮流反転Figure 1 Analysis circuit
・Rated equipment capacity = 50 MW (55.9 MVA), primary side rated voltage = 66 kV, secondary side
rated voltage = 56 kV, DC voltage = 125 kV.
・The DC power transmission system is linked with the 66-kV system through the impedance of thesimulated short-circuit capacity of the system.
・In this, example simulations of ① a power flow reversal and ② a DC main line ground fault are

64carried out. At a simulation time of t = 0.5 s, the power command value changes, and a power flow reversal
begins. A ground fault occurs at t = 0.8 s, which is subsequently removed by the circuit breaker.

Analysis Results
Figures 2 and 3 show the DC capacitor voltage and direct current at the time of a power flow reversal.
Figures 4 and 5 show the active and reactive power outputs from terminals and terminal 2 at the time of apower flow reversal, respectively. Because the reactive power control of each terminal during thissimulation continuously controls the voltage of the AC side (ACAVR), the reactive power is controlled

Control No.

such that the terminal voltage of the converter approaches 1.0 pu.

Time [s]Branch current [A]Figure 2 DC capacitor voltage (average value of all cells) [pu] (the reference voltage is 125 kV).

Time [s]Figure 3 Direct current [A]65

Control No.

Time [s]Figure 4 Active power [pu]Blue line, terminal 1; red line, terminal 2 (the reference capacity if 55.9

Control No

MVA).

Time [s]Figure 5 Reactive power [pu]: blue line, terminal 1; Red line, terminal 2 (the reference capacity is 55.9).

At t = 0.5 s, the power command value switches from 45 MW (0.8 pu) to -45 MW (-0.8 pu), and thepower flow reversal begins. The rectifier side controls the active power and determines the power to be sent
to the other system, whereas the inverter side determines and controls the DC voltage.
Figure 6 shows the DC voltage, and Figure 7 shows the system side output current at the time of a DC
fault. A direct current fault occurs at t = 0.8 s. Further, once the system detects a DC fault, the converter
performs a gate block (GB) → circuit breaker release operation, and therefore, despite a large current
being output from the converter during the fault, the output current of the system side becomes zero after
the circuit breaker is opened.

66

Control No.

Time [s]Figure 6. DC voltage [pu]: blue line, terminal 1; red line, terminal 2 (the reference voltage is 125 kV, and is

Branch current [A]the average value of the capacitor voltages of 6 arms)

Time [s]Figure 7 System side output current [A]67

The parameters of the model and their respective descriptions are shown below. Note that in the settings ofterminal 2, the variable names are the same as those of terminal 1 except that the 1 at the end is replaced

Common constants

with 2, and hence these details have been omitted here.
Name

Unit

Item

Rate_PW

Rated active power

Rate_Vdc

V

Rated DC voltage

SSRelayOffTimes

Relay dead time at start of simulation

FlagInitialize-

Initialization flag (used to start the simulation from a steady state)
Input 1 to start from steady state

V_ThirdOrder

p.u.

Third harmonic injection amount for the modulation rate reduction

CellNumPerArm-

Number of stages per arm of MMC cell (in the averaging model, always enter 1)

SystemFrequency1

Hz

System frequency

Rate_S1

VA

Rated capacity

Rate_V1_ltol1

V

Effective value of transformer primary side line voltage

Rate_V2_ltol1

V

Effective value of transformer secondary side line voltage

PWMPulseNumber-

Number of PWM pulses (per cycle)

R_TransGND1Ω

Ground resistance of converter transformer

H1sElectrostatic constant

Flag_ACAVR1-

ACAVR flag (1 = used, 0 = not used), Select Flag_AQR

Flag_AQR1-

AQR flag (1 = used, 0 = not used), Select Flag_ ACAVR

Vac_order1

p.u.

AC voltage command value (ACAVR command value)

Q_order1

p.u.

Reactive power command value (AQR command value)

PWMPhase1

degPhase angle of PWM carrier

R_connect_pu1

p.u.

Converter transformer resistance

L_connect_pu1

p.u.

Converter transformer inductance

L_Arm_pu1

p.u.

Arm reactor inductanceInductance of DC reactor (for upper line)

L_DCK_pu1

p.u.

L_DCA_pu1

p.u.

Inductance value = L_DCK_pu 1 × rated DC voltage 2/rated active power
Inductance of DC reactor (for lower line)

Gain_ACR1-

ACR gain

T_ACR1s

ACR time constant

Gain_APR1-

APR gain

T_APR1s

APR time constant

Lag_APR1sFirst-order lag time constant for APR

Gain_DCAVR1-DCAVR gain

T_DCAVR1s

DCAVR time constant

Gain_AQR1-

AQR gain

T_AQR1s

AQR time constant

Gain_ACAVR1-

ACAVR gain

T_ACAVR1s

ACAVR time constant

Vmag1

p.u.

Voltage margin

Pmag1

p.u.

Power margin68Lag_Margin1

s

First-order lag time constant at margin switching (switching speed)

IinvLimit1

p.u.

Maximum current of converter (denominator is the converter rating)

K_DCFRD1-

Gain of DC fault detection relay (the smaller the value, the higher the sensitivity)

T_DCFRD1s

First-order lag time constant of DC fault detection relay

Threshold_DCOC1

p.u

DC overcurrent relay threshold

Threshold_OC1

p.u.

Overcurrent relay threshold

Threshold_DCOV1

p.u.

DC overvoltage relay threshold

Threshold_DCUV1

p.u.

DC undervoltage relay threshold

Threshold_OV1

p.u.

AC overvoltage relay threshold

Threshold_UV1

p.u.

AC undervoltage relay threshold

Threshold_OF1

p.u.

Frequency rise relay threshold

Threshold_UF1

p.u.

Frequency drop relay threshold

VAMP_LAG1sFirst-order lag time constant for AC voltage detection

FREQ_LAG1sFirst-order lag time constant for frequency detection

RecoveryTime1-

Lag from fault relay cancellation to actual cancellation of protection operation

GBTime1s

GB implementation time

NoDetectPeriod1sPartial relay dead time at startup

CalculationTime1sControl cycle

DeadTime1s

Dead time

Flag_DirectSample1-

Voltage, current amount direct detection flag (selected from among three types)

Flag_AverageSample1-

Detection flag using moving average of voltage and current amount (selected from among three types)

Flag_PeakTopSample1-

Peak top detection flag of voltage and current amount (selected from among three types)

ChargeCircuitCloseTime1sStart time of charging (if starting from a transient state, assign a large negative value)

ChargeTime1s

Charge time

Control_StartDelay1sTime from end of charge to start of control system operation
Time of starting current limit (current zero control)

ManualCL_Time1s

ManualGB_Time1s(used to simulate converter protection)
Time of starting gate block(used to simulate converter protection)
Time of starting circuit breaker release

ManuaCBT_Time1s(used to simulate converter protection)

End

69

Change log
Date
2018/06/21

Example fileContent changes

version
1.0

First edition created70

Examples of XTAPNo.

HVDC-04

ExampleSelf-Excited HVDC Model (Full Bridge MMC)

Name
Field

Power electronics(1) “DC Power Transmission Engineering,” Tokyo Denki University Press, written andedited by Takahiko Machida(2) T. Kikuma, K. Takenaka, M. Takasaki, “Control Protection Method of DC Power
Transmission System with Modular Multi-Level Converter Using Full Bridge Cell:
Proposal of a Control Method that can Control Fault Current During DC Fault,” Central
Research Institute of Electric Power Industry Research Report R11021, (2012).
(3) T. Kikuma, “Control Protection Method of DC Power Transmission System with

Literature

Modular Multi Level Converter Using Full Bridge Cell - Part 2: Development of
Operation Continuous Control and High-Speed Starting Method at AC Voltage Rise,”Central Research Institute of Electric Power Industry Research Report R14016, (2015).
(4) T. Kikuma et al., “Control Method of Double Y-connection Modular Multi-level
Converter,” Journal of the Institute of Electrical Engineers of Japan D, Vol. 133, No. 9,
pp. 917-927 (2013).
(5) T. Kikuma et al., “Control Protection Method of Self-Excited DC Power
Transmission System capable of Suppressing DC Fault Current,” Journal of theInstitute of Electrical Engineers of Japan, Vol. 133, No. 5. pp. 449-456 (2013).
Because this model has basic control and protection systems, it can be applied toother simulations by copying it and changing the rated value.
This example simulates the circuit of a DC power transmission system using a
modular multilevel converter (full bridge MMC). This has the same configuration as

Overviewthat of HVDC-03 (self-excited HVDC model, half bridge MMC), except that the cellsof the MMC are replaced with full bridge cells. For a basic description of the model,
refer to HVDC-03. In this example, the simulated waveforms at the time of a power
flow reversal and a DC main line ground fault are shown.

71

Introduction/Model Use
Refer to the table at the end for the parameters required to use the MMC model. When changing theactive power output of the MMC, “Pdorder” in “MASTER_VSC 1,” which is located one level below thehighest level of the hierarchy, is changed. This example deals with two types of models:②

a full bridge MMC averaging model (HVDC-04A), and② a full bridge MMC detailed model (HVDC-04B).
In the absence of any specific reason to use the detailed model, the use of the average model isrecommended because it is much faster.

Analysis Circuit and Conditions
Figure 1 shows the analysis circuit.

Powerflow reversal
潮流反転Figure 1 Analysis circuit
・Rated equipment capacity = 50 MW (55.9 MVA), primary side rated voltage = 66 kV, secondary side
rated voltage = 56 kV, DC voltage = 125 kV.
・The DC power transmission system is linked with the 66-kV system through the impedance of thesimulated short-circuit capacity of the system.
・In this example, simulations of ① a power flow reversal and ② a DC main line ground fault are
carried out. At a simulation time of t = 0.5 s, the power command value changes and a power flow reversal
begins. A ground fault occurs at t = 0.8 s, after which, the fault removal and restart are carried out through aconverter operation.

Analysis Results
Figures 2 and 3 show the DC capacitor voltage and direct current at the time of a power flow reversal.
Figures 4 and 5 show the active and reactive power outputs from terminals 1 and 2 at the time of a power
flow reversal, respectively. Because the reactive power control of each terminal in this simulation
continuously controls the voltage of the AC side (ACAVR), the reactive power is controlled such that theterminal voltage of the converter approaches 1.0 pu.

72Branch voltage [V]Time [s]Branch current [A]Figure 2 DC voltage [pu]: blue line:, terminal 1; red line, terminal 2 (reference voltage of 125 kV)Time [s]Control signal

Figure 3 Direct current [A]Time [s]Figure 4 Active power [pu]: blue line, terminal 1; red line, terminal 2 (reference capacity of 55.9 MVA)

73

Control signal

Time [s]Figure 5 Reactive power [pu]: blue line, terminal 1; red line, terminal 2 (reference capacity of 55.9)

At t = 0.5 s, the power command value switches from 45 MW (0.8 pu) to -45 MW (-0.8 pu), and thepower flow reversal begins. The rectifier controls the DC current and determines the power to be sent toanother system, whereas the inverter side determines and controls the DC voltage.
Figures 6 and 7 show the DC voltage and current, respectively, at the time of a DC fault. When a direct
current fault occurs at t = 0.8 s, the power transmission is temporarily stopped and is restarted based on the

Branch voltage [V]function of the converter. For details of the full bridge MMC during a DC fault, see the references.

Time [s]Branch current [A]Figure 6 DC voltage [pu]: blue line, terminal 1; red line, terminal 2 (reference voltage of 125 kV)Time [s]4 current [A]Figure 7 Direct74

The parameters of the model and their respective descriptions are shown below. Note that in the settings ofterminal 2, the variable names are the same as those of terminal 1 except that the 1 at the end is replaced

Common constants

with 2, and hence these details have been omitted here.
Name

Un

Item

Rate_PW
it

Rated active power

Rate_Vdc

V

Rated DC voltage

SSRelayOffTimes

Relay dead time at start of simulation

FlagInitialize-

V_ThirdOrderpu

Third harmonic injection amount for the modulation rate reduction

CellNumPerArm-

Number of stages per arm of MMC cell (in the averaging model, always enter 1)

SystemFrequency

Hz

System frequency

Rate_S1
1

V

Rated capacity

Rate_V1_ltol1

V
A
V

Effective value of transformer secondary side line voltage

Rate_V2_ltol1

Initialization flag (used to start simulation from steady state)
Input 1 to start from steady state

Effective value of transformer primary side line voltage

PWMPulseNumbe-

Number of PWM pulses (per cycle)

R_TransGND1r
H1ΩGround resistance of converter transformersElectrostatic constant

Flag_ACAVR1-

ACAVR flag (1 = used, 0 = not used), select Flag_AQR

Flag_AQR1-

AQR flag (1 = used, 0 = not used), select Flag_ ACAVR

Vac_order1

p.u

AC voltage command value (ACAVR command value)

Q_order1

p.u
.
de
.

Reactive power command value (AQR command value)

p.u
g
p.u
.

Converter transformer resistance

PWMPhase1
R_connect_pu1
L_connect_pu1
L_Arm_pu1
L_DCK_pu1

p.u
.
p.u
.
.

Phase angle of PWM carrier

Converter transformer inductance
Arm reactor inductanceInductance of DC reactor (for upper line)
Inductance value = L_DCK_pu 1 × rated DC voltage 2/rated active power

L_DCA_pu1

p.u

Inductance of DC reactor (for lower line)

Gain_ACR1

ACR gain

T_ACR1-.
s

ACR time constant

Gain_APR1-

APR gain

T_APR1s

APR time constant

Lag_APR1sFirst-order lag time constant for APR

Gain_DCAVR1-DCAVR gain

T_DCAVR1s

DCAVR time constant

Gain_AQR1-

AQR gain

T_AQR1s

AQR time constant

Gain_ACAVR1-

ACAVR gain

T_ACAVR1s

ACAVR time constant

Lag_Margin1

s

First-order lag time constant at margin switching (switching speed)

IinvLimit1

p.u

Maximum current of converter (denominator is the converter rating)

K_DCFRD1-.

Gain of DC fault detection relay (the smaller the value, the more the sensitivity)75

T_DCFRD1sFirst-order lag time constant of DC fault detection relay

Threshold_DCOCp.u

DC overcurrent relay threshold

Threshold_OC1
1
Threshold_DCOV

p.u

Overcurrent relay threshold

p.u
.
p.u
.

DC overvoltage relay threshold

Threshold_DCUV
1
Threshold_OV1
1
Threshold_UV1
Threshold_OF1
Threshold_UF1
VAMP_LAG1

p.u
.
p.u
.

DC undercurrent relay threshold
AC overvoltage relay thresholdAC undervoltage relay threshold

p.u
.
p.u
.

Frequency rise relay threshold
Frequency drop relay threshold
First-order lag time constant for AC voltage detection

FREQ_LAG1s.
s

RecoveryTime1-

Lag from fault relay cancellation to actual cancellation of protection operation

First-order lag time constant for frequency detection

GBTime1s

GB implementation time

NoDetectPeriod1sPartial relay dead time at startup

CalculationTime1sControl cycle

DeadTime1s

Dead time

Flag_DirectSampl-

Voltage, current amount direct detection flag (selected from among three types)

Flag_AverageSam
e1
Flag_PeakTopSa
ple1-

Detection flag using the moving average of voltage and current amount (selected from among three types)-

Peak top detection flag of voltage and current amount (selected from among three types)

ChargeCircuitClos
mple1
ChargeTime1eTime1s

Start time of charging (if starting from a transient state, assign a large negative value)s

Charge time

Control_StartDelasTime from end of charge to start of control system operation
Time of starting current limit (current zero control)y1ManualCL_Time1s

ManualGB_Time(used to simulate converter protection)
Time of starting gate blocks(used to simulate converter protection)

1
ManuaCBT_Time

Time of starting circuit breaker releases

1
DCRecoveryTime
DCCLTime1
1
Threshold_UVVDs(used to simulate converter protection)
Delay from DC fault detection relay release until actual cancellation of protection operation(Used to simulate converter protection)
DC current zero restoration execution time at the time of DC fault (time from detection of DC fault to
GB)
Lowering of DC voltage when AC voltage becomes lower than threshold value
DC voltage margin

DCVmag1
1
DCAmag1
OVmag1

DC current margin
AC overvoltage margin(margin for lowering of DC voltage for continuous operation during rise of AC voltage)

End

76

Change log
Date
2018/06/21

Example fileContent changes

version
1.0

First edition created77

Examples of XTAPNo.

FACTS-01

ExampleSimulation Sart of STATCOM

Name
Field
Literature

Basic calculation (power systems)
None
In this example, a calculation of the start simulation of STATCOM is carried out.
STATCOM, the subject of this study, is the most basic two-level converter.
Dismissing with the details, the design of the converter should cope with a DC voltage
fluctuation of ± 20%. Normally, the converter charges the DC capacitor of the converter
from the outside with an initial charging device, and activates it by interlocking withthe power system from the state of a DC voltage of 1 pu. In this example, the DC
capacitor was charged using the diode of the self-excited converter and was then set as
the activating sequence. This start of the simulation is expected to enhance theunderstanding of the converter control.
Vector control based on the dq rotating coordinate system widely used in STATCOM

Overview

for power systems was adopted as the converter control. The PWM method was used
for the pulse generation, which actually applies the switching. In this example, theconversion factors of the vector and PWM controls are multiplied by the rated
modulation factor, which may be found to be useful because such an example has notbeen specifically discussed in textbooks or other studies. Further, the maximum valueof the vector control output (Vdref and Vqref in the example) is determined from the
maximum modulation rate. This aspect also needs to be studied because there are very
few examples available. Regarding the control, the aim is to deepen our understanding
through this example because the actual converter design (equipment specifications
determined based on the requirements of the power system) and control are closely
related.

78

Analysis Circuit and Conditions
Figure 1 shows the analysis circuit.

Circuitt = 0.04 Converter
start-up

PWM ignoring DT

Control system
Pulse

No. 1 startup time of converter

Currently, right term →command value, not voltage
sourceDeviation of Δ and compensation of detection delay of 100 us
Set to zero when checking Vacheck!

Multiply by the rated modulation factor. Important!
Conversion factor for vector control output → to PWM control
Divide using Vdc

Vd, Vq check
Divide using Vdc

Check: Has the voltage returned to its original value?

Figure 1 Analysis circuit[Basic conditions]Power system: 100 kVA, 1.65-kV Zbase = 27.225 [Ohm] = 86.659 [mH]
Converter transformer: 145 [kVA] 1650 [V] R + jX = 2% + j20%
Converter: S = 138 [kVA] Vac = 1650 [V] Vdc = 2000 [V]79

Branch voltage [V]Branch voltage [V]Analysis Results

Line voltage

Ground potential of DC capacitor(upper side)
Ground potential(lower side)

Time [s]Time [s]Converter transformer ab line voltage [V] on secondary sideLine voltage of DC circuit and ground potential[V]○ A direct current line voltage is selected, and an AC voltage source is created on the secondary side of
the converter transformer. Exchange of active/reactive power takes place owing to the phase difference andpotential difference between the voltages of this voltage source and the power system. Because the DC
circuit side is not grounded, the potential as seen from the ground at both ends of the DC capacitor is ±Vdc/2. If grounded with a low resistance, the potential of the capacitor (upper side) is Vdc, and theConverter
start-up

Node voltage [V]Node voltage [V]potential of the capacitor (lower side) becomes zero.

Time [s]Time [s]Primary side current [pu] of converter transformer

Primary voltage [pu] of converter transformer○When AC - CB is closed, the DC capacitor is charged from the system through the diode. The ACvoltage is reduced to approximately 0.7 pu. Because the leakage impedance of the converter transformer isapproximately 20%, a current flow of ΔI = ΔV/Z = 0.7/0.2 = 3.5 pu is created.

80Node voltage [V]Node voltage [V]Time [s]Time [s]○ The active current command (Idref) and active current (Id) are tracked within the range enclosed in red.
At this time, the vector control outputs (Vdref, Vqref) also operate within the limiter range. Because theconverter is assumed to be able to cope with a DC voltage fluctuation of ± 20%, as part of the device
specifications, the trackability is ensured when Vdc falls within this range. Thus, it is clear that suchmatters are not determined solely through converter control, and that the control is closely related to thedesign of the converter (equipment specifications determined based on the requirements of the power

Node voltage [V]Node voltage [V]system).

Time [s]Time [s]○ The AC voltage command (Vacref) and AC voltage (Vac) have been tracked reasonably well. Because
the control circuit is a first-order lag circuit, a steady-state deviation cannot be avoided. The figure on theright shows the waveform as indicated in dq rotation coordinates. In addition, Vd = 1 pu and Vq = 0 pu areachieved under a steady state. Vector control is a theory based on the fact that the control system of the dq
axis current is independent, and PQ is not really independent. Through the moment-to-moment tracking of
the dq axis when Vq = 0 using PLL, PQ becomes independent as seen from the relationship P = Vd * Id +Vq * Iq = Vd * Id.

A simulation analysis was carried out under severe system conditions such as an extremely large voltage

81drop when AC - CB is applied. The ability to judge an analysis result can only come from knowing thefundamentals of electrical circuits and control engineering, and it is therefore important to study the basics.

End

Change log
Date

Example fileContent changes

version2014/11/19

2.0

Modified for XTAP Version 2.00

2012/07/19

1.2

Modified for XTAP Version 1.20

2011/10/18

1.1

Modified for XTAP Version 1.11

2010/09/02

1.0First edition created (for XTAP Version 1.10)

82

Examples of XTAPNo.

FACTS-02

Name of
STATCOM Model

ExampleField
LiteraturePower electronics
None
Because this model has basic control and protection systems, it can be applied toother simulations by copying and changing the rated value.
In this example, the STATCOM circuit is simulated. STATCOM is a type ofequipment that controls the reactive power using a self-exciting semiconductor device.
Because it is able to control the reactive power continuously at a high speed, it can
respond faster to sudden changes in the system voltage.

Overview

The characteristics of STATCOM dealt with in this example are described below.
・The circuit system of the converter is a basic two-level converter.
・Performs a reactive power constant control (AQR) operation that outputs a constantreactive power irrespective of the magnitude of the system voltage.
・A resistance charging system is employed, which charges the device at the time ofthe startup through a resistor.
・A switching pulse applied to the gate of the semiconductor device is created using
the PWM control.

83Analysis Circuit and Conditions
Figure 1 shows the analysis circuit.

STATCOM

Figure 1 Analysis circuit

Capacitor voltage
コンデンサ電圧(DC voltage)(直流電圧)

Switch 1
スイッチ1

Charging resistor
充電用抵抗Converter current
変換器電流

系統電流
電力

System current power

Switch 2
スイッチ2

AC filter
交流フィルタFigure 2 Schematic diagram of model・Rated equipment capacity = 10 MVA, primary side rated voltage = 66 kV, secondary side rated voltage =5.6 kV, rated DC voltage = 12.5 kV・STATCOM is linked with the system (66 kV) through the impedance of the simulated short-circuit
capacity of the system.
・A three-wire ground fault occurs at a simulation time of t = 0.5 s, which is removed by the circuit breakerat t = 0.6 s.

Analysis Results
We first show the waveform at the start of STATCOM. Figure 3 shows the capacitor voltage of
STATCOM, Figure 4 shows the reactive power output by STATCOM, and Figure 5 shows the current
flowing from the converter to the system.

84

Control system

Time [s]Control system

Figure 3 Capacitor voltage for energy storage [pu]Time [s]Figure 4 Reactive power output by STATCOM [pu]85Branch current [A]Time [s]Figure 5 System side output current [A]
Switch 1 is closed at t = 0.05 s and STATCOM becomes linked with the system through the charging
resistor, and the DC capacitor for energy storage is charged to approximately 0.6 pu. Switch 2 is thenclosed at t = 0.25 s, and at the same time at which STATCOM and the system become directly linked, thecontrol system starts to operate. When the control system starts operating, the voltage of the capacitor
becomes 1.0 pu, and the reactive power output of STATCOM conforms to the command value (here, 0.8pu). A steady-state deviation occurs because the reactive power control of STATCOM uses a first-order lag
control, and hence the command value and actual reactive power output of STATCOM differ slightly.
Figures 6 and 7 show the converter current and system side output current under a steady state. Because
STATCOM applies PWM control, the converter current has a waveform that includes a large amount of
harmonics derived from the PWM modulated wave. Because these harmonics are removed using an ACfilter connected between the converter and system, the current flowing through the system is a sinusoidal

Branch current [A]wave.

Time [s]Figure 6 Enlarged view of converter current

86Branch current [A]Time [s]Figure 7 Enlarged view of system side output current

The system voltage at the time of a three-phase ground fault is shown in Figure 8, and the converter
current is shown in Figure 9. In this simulation, when there is a drop in the system voltage, STATCOM
restarts the operation after a gate block (GB). As can be seen from Figure 8, because the converter is
stopped after the GB, the converter current reaches zero. When the fault is removed and the system voltage

Potential difference between nodes [V]recovers, the converter automatically restarts.

Time [s]Figure 8 System voltage

87Branch current [A]Time [s]Figure 9 System side output current

The parameters of the model and their description are shown on the next page.

88

Name

Unit

Item

SystemFrequency

Hz

System frequency

Rate_S

VA

Rated capacity

Rate_V1_ltol

V

Effective value of transformer primary side line voltage

Rate_V2_ltol

V

Effective value of transformer secondary side line voltage

Rate_Vdc

V

Rated DC voltage

PWMPulseNumber-

Number of PWM pulses (per cycle)

R_TransGNDΩ

Ground resistance of converter transformer

HsElectrostatic constant

Flag_ACAVR-

ACAVR flag (1 = used, 0 = not used), Select Flag_AQR

Flag_AQR-

AQR flag (1 = used, 0 = not used), Select Flag_ ACAVR

Vdc_orderp.u.

DC voltage command value (DCAVR command value)

Vac_orderp.u.

AC voltage command value (ACAVR command value)

Q_orderp.u.

Reactive power command value (AQR command value)

PWMPhasedegPhase angle of PWM carrier

R_connect_pu

p.u.

Converter transformer resistance

L_connect_pu

p.u.

Converter transformer inductanceSpu_filter1

p.u.

Fundamental wave capacity of the higher-order filter 1 (denominator is rated capacity of converter)

Q_filter1-

Sharpness of higher-order filter 1

Spu_filter2

p.u.

Fundamental wave capacity of higher-order filter 2 (denominator is rated capacity of converter)

Q_filter2-

Sharpness of higher-order filter 2

Gain_ACR-

ACR gain

T_ACRs

ACR time constant

Gain_DCAVR-

DCAVR gain

T_DCAVRs

DCAVR time constant

Gain_AQR-

AQR gain

T_AQRs

AQR time constant

Gain_ACAVR-

ACAVR gain

T_ACAVRs

ACAVR time constantIinvLimitp.u.

Maximum current of converter (denominator is the converter rating)Threshold_OCp.u.

Overcurrent relay threshold

Threshold_DCOV

p.u.

DC overvoltage relay threshold

Threshold_DCUV

p.u.

DC undervoltage relay threshold

Threshold_OV

p.u.

AC overvoltage relay threshold

Threshold_UV

p.u.

AC undervoltage relay threshold

Threshold_OFp.u.

Frequency rise relay threshold

Threshold_UFp.u.

Frequency drop relay threshold

VAMP_LAGs

First-order lag time constant for AC voltage detection

FREQ_LAGs

First-order lag time constant for frequency detection

RecoveryTime-

Lag from fault relay cancellation to actual cancellation of protection operation

GBTimes

GB implementation time

NoDetectPeriodsPartial relay dead time at startup

SSRelayOffTimes

Relay dead time at the start of the simulation

CalculationTimes

Control cycle

DeadTimes

Dead time

Flag_DirectSample-

Voltage, current amount direct detection flag (selected from among three types)

Flag_AverageSample-

Detection flag using moving average of voltage and current amount (selected from among three types)

Flag_PeakTopSample-

Peak top detection flag of voltage and current amount (selected from among three types)

ChargeCircuitCloseTimes

STATCOM start time = start time of charging(If starting from the transient state, assign a large negative value)
ChargeTimes

Charge time

Control_StartDelays

Time from the end of the charge to the start of the control system operation
充電終了から，制御系動作開始までの時間

89

End

Change log
Date

Example fileContent changes

versionIcons of parts changed

2018/6/29

2.2

2018/1/26

2.1

Modified the creation logic of HOLD signal

2014/11/19

2.0

Modified for XTAP Version 2.00

2013/04/15

1.3

2012/08/29

1.2

Changes consequent to modification of STATCOM model

2012/07/19

1.1

Modified for XTAP Version 1.20

2011/10/18

1.0First edition created (for XTAP Version 1.11)

Ground terminal of transformer added

Changes consequent to modifications in model diagram andparameter input format

90

Example FACTS-03.doc XTAP collection

No.

FACTS-03

ExampleSeparately Excited SVC Model

Name
Field

Power electronics
Technical Report of the Institute of Electrical Engineers of Japan, Electricity and

Literature

Energy Department, Power & Energy Department Stationary Equipment TechnicalCommittee, “Energy Saving Technology of Static Reactive Power Compensator”(2004).
Because this model has basic control and protection systems, it can be applied toother simulations by copying it and changing the rated value.
A separately excited SVC (hereinafter referred to as an SVC) is a device thatcompensates the reactive power through a control using thyristors. Because a type of
SVC that combines a thyristor controlled reactor (TCR) and AC filter (as well as a
stacon) is currently the mainstream type used in Japan, the TCR configuration is used
here. A TCR is a device that controls the reactive power by continuously controlling thecurrent flowing through a reactor using a thyristor. Here, if the control angle is , thefundamental wave component flowing in the reactor changes continuously with theOverview

change in . The reactive power is controlled by utilizing this property.
Because the current waveform output from the TCR contains large amounts ofharmonics, an AC filter is installed such that the harmonic current does not flow out tothe grid. Moreover, by installing an AC filter and a stacon, it is possible to supply
reactive power in addition to consuming it. The TCR part is connected to the AC filter
and the stacon through a delta-delta connection transformer. Most of the harmonics
generated in the TCR section are absorbed by the fifth- and seventh-order filters andstacon. To avoid an inrushing current and suppress the harmonic resonance in thestacon, 6% (or 13%) reactors are inserted in-series. If necessary, the SVC may beconnected to the system through another transformer, but this feature is not included in
the present model.

91Analysis Circuit and Conditions
Figure 1 shows the analysis circuit created using XTAP, and Figure 2 shows a schematic diagram of themodel.

SVC

Figure 1 Analysis circuit
AC filter
交流フィルタStaconスタコン
System系統

Circuit
遮断器
breaker

TCR
SVC(a) Overall view of SVC
Figure 2(b) TCR circuit diagram
Schematic diagram of the model

・The main ratings are as follows: rated capacity of 100 to -100 MW, and rated voltage of 275/20 kV.
・The SVC is linked with the system (275 kV) through the simulated impedance of the short-circuit
capacity of the system.
・A three-wire ground fault occurs at a simulation time of t = 0.5 s, which is removed by the circuit breakerat t = 0.6 s.

Analysis Results
Figure 3 shows the voltage and current measurement points.
AC filter
交流フィルタstaconスタコン
Systemcurrent
系統電流reactive power

無効電力

Systemvoltage
系統電圧

Circuit breaker

遮断器
TCR部current of TCR part
Reactor
リアクトル電流Figure 3 Voltage and current measurement points

92

Figure 4 shows the voltage and current waveforms during a steady state. The harmonic component of the

TCR part reactor current [A]System current [A]System voltage [V]current output from the TCR is absorbed by the AC filter, and a sinusoidal current is output to the system.

Time [s]Figure 4 Voltage and current waveforms during a steady state
Figure 5 shows the waveforms from the occurrence of the three-wire ground fault to the point of returning
to a steady state after removal of the fault. During this simulation, because the system is controlled to holddown the output from the SVC to zero during the fault period, the reactive power output by the SVC

Control angle  [deg]Reactor power [pu]System voltage [V]becomes zero immediately after the elimination of the fault, and rises thereafter.

Time [s]Figure 5 Waveform of each part at the time of three-wire ground fault

93The parameters of the model and their descriptions are shown below.
Name

Unit

Item

SystemFrequency

Hz

System frequency

Rate_V1_ltol

Vrms

Effective value of transformer primary side line voltage

Rate_V2_ltol

Vrms

Effective value of transformer secondary side line voltage

Rate_S_TCR

VA

TCR capacity

Rate_SC

VA

Total capacity of stacon and harmonic filter

F_5_pup.u.

Resonance order of fifth-order filter

Q5-

Sharpness of fifth-order filter

S5_pu

p.u.

Ratio of fifth-order filter (denominator is Rate_SC)

F_7_pup.u.

Resonance order of seventh-order filter

Q7-

Sharpness of seventh-order filter

S7_pu

p.u.

Ratio of seventh-order filter (denominator is Rate_SC)

R_GND



Earth resistance

TCR_Rpu

p.u.

Element resistance of TCR

Trans_Rpu

p.u.

Resistance of TCR transformer

Trans_Lpu

p.u.

Inductance component of TCR transformer

Flag_AQR-

Reactive power control (AQR) flag (1 or 0)

Q_orderp.u.

Command value of AQR (denominator is Rate_S_TCR)

Gain_AQR-

AQR gain

T_AQRsTime constant of AQR

Flag_ACAVR-

AC voltage control flag (ACAVR) (1 or 0)

Vac_order

p.u.

Command value of ACAVR

Gain_ACAVR-

ACAVR gain

T_ACAVRsTime constant of ACAVR

Alpha_mindegMinimum value of control angle

Alpha_maxdegMaximum value of control angle

Control_Starts

Control system start time

SSRelayOffTimes

Relay dead time at start of the simulation

CalculationTimes

Control cycle

VAMP_LAGs

Time constant of first-order lag for AC voltage detection

VAMP_AVE

cycle

Number of cycles of moving average of AC voltage measurement

FREQ_LAGs

Time constant of first-order lag for frequency detection

Limit_Iinv

p.u.

Maximum value of current

CurrentShiftp.u.

Current command value at current shift execution

T_LimitOpensTime constant of limiter opening for cancelling current shift

Threshold_OCp.u.

OC relay threshold

Threshold_OVp.u.

OV relay threshold

Threshold_UV

p.u.

UV relay threshold

Threshold_OFp.u.

OF relay threshold

Threshold_UFp.u.

UF relay threshold

RecoveryTimes

Lag from fault relay cancellation to actual cancellation of protection operation

GBTimes

GB implementation time (time from start of GB up to CBT)

DeionTimes

Thyristor deionization time (see Tdeion section of Thy in the manual)

End

94

Change log
Date

Example fileContent changes

version2018/6/29

2.1

Icons of parts changed

2014/11/19

2.0

Modified for XTAP Version 2.00

2013/04/15

1.1

2012/08/29

1.0

Changes consequent to modifications in model diagram andparameter input format
First edition created (for XTAP Version 1.2)

95

Examples of XTAP

No.

PQ-01

ExampleSimulation of Flicker Suppression Using Separately Excited SVC

Name
Field

Basic calculation (power systems)
N. Tibaki, K. Takenaka, K. Yukihira, K. Huuchi, N. Hatano, Y. Takeuchi, T. Ishikawa,

Literature“Development of a control scheme for separately excited SVC for suppressing flicker,”
CRIEP Report R 0 6010, May 2007
In this example, the calculation of flicker suppression simulation using a separately
excited SVC is carried out.
The aim of this example is to learn the simulation of a separately excited converter
using a thyristor element. The target of the study is the simulation of the conventional
control of SVC suppression of the flicker using a simulated input.

Overview

The ignition of the frequency conversion stations and HVDC is usually conductedusing three-phase batch control. The main difference in the case of flicker suppressionis that an ignition is achieved using three-phase individual control because the reverse
phase component is also subject to compensation. A conventional control system isdesigned with a view to detecting the reactive power of the flicker and canceling it.
Details of the equipment constants and other factors are provided in the references.

96Analysis Circuit and Conditions
Figure 1 shows the analysis circuit.
Current flowing from arc furnace to system side → FF current
Infinite bus voltage

Total current of SVC and FIL

FIL single current

Binding of system and Filter1

Binding of system and Filter2

SVC
Figure 1 Analysis circuit

SVC single current

Figure 2 SVC Circuit configuration (TCR)97

Branch current [A]Branch current [A]Analysis Results

Time [s]Time [s]SVC + filter current [A]Branch current [A]

Arc furnace current [A]Time [s]Control system

Control system

Current flowing in the upper system [A]Time [s]Time [s]Control angle α [deg]Pulse to A-phase upper element

98The arc furnace current and voltage are detected, from which the instantaneous reactive power isdetermined. The SVC acts in such a mannner to cancel this. In actuality, because the TCR can only becontrolled by L, to operate the variation of ± Q, the arc furnace ΔQ is compensated by synthesizing thefundamental wave C of the harmonic filter.
A basic operation of the control is canceling with a feedforward, and hence the gain must be less than 1.
Even if the control effect increases, it operates to amplify the disturbances, should they occur unexpectedly,
and it is therefore necessary to set the gain to less than 1.
By compensating the arc furnace current with the combined current of the SVC + filter, it can be
confirmed that the current flowing to the upper system is smaller.
The control angle α is seen to move corresponding to the Q variation of the arc furnace. Switching is
conducted twice during every cycle using the upper and lower thyristors. The compensable frequency bandis smaller when compared with a self-excited converter, but this is not to say that a flicker can becompensated by a separately excited SVC.

End

99

Change log
Date
2014/11/19

Example fileContent changes

version
2.0

Modified for XTAP Version 2.00
The method of simulating the current flowing from the arc

2014/02/25

1.3

furnace to the system side changed from a mathematical
waveform of the voltage source to a combination of the Cos
wave current source

2012/07/19

1.2

Modified for XTAP Version 1.20

2011/10/18

1.1

Modified for XTAP Version 1.11

2010/08/3

1.0First edition created (for XTAP Version 1.10)

100

Examples of XTAPNo.

PQ-02-A

ExampleCalculation of Excitation Inrush Current under Pressurization of Customer Transformer

Name
Field

Power quality analysis, excitation inrush current calculation
The data of the transformer of the present example were prepared by referring to thefollowing paper.

Literature[1]

K. Tokunaga, “Inrush Current Analysis of Two-Winding Power Transformer

based on Equipment Constants Estimated using Winding Shape Design Procedure,”
Journal of the Institute of Electrical Engineers of Japan, Vol. 128, No. 9, pp. 10751081, 2008.
When the transformer is pressurized, the magnetic flux of the iron core is saturated
owing to the transient phenomenon during pressurization, and the excessive current
flows over many cycles. This is known as the excitation inrush current. From theviewpoint of power quality, a voltage reduction from an excessive current flowing fromthe upper system becomes a problem, and from the viewpoint of protecting theOverview

transformer, the windings are displaced by the electromagnetic force, which generates
an overcurrent flowing through the windings, leading to a concern that the insulator
force may decrease.
In this example, an exciting inrush current calculation is conducted whenpressurizing a customer transformer with a 66-kV reception (high power reception).
This is a basic example without considering the initial magnetic flux (residual magneticflux).

101

Analysis Circuit and Conditions
Figure 1 shows a single line connection diagram of the analysis circuit.

Upper level
current
Short circuit
66 kV
impedancesystemvoltage

Circuitbreaker
Input at t =8.5 msTransformer
66:3.15 kV

50 Hz

Figure 1 Single line connection diagram of analysis circuit[Circuit Operation]
The relationship between the voltage v applied to the transformer and the magnetic flux  generated in
the transformer core is given by

v

d
dt



t

   v ( ) d 
0Therefore, the average value of the vibration of the magnetic flux does not become zero unless the voltage
phase at the time of application is 90°. In particular, when the voltage phase at the time of application isclose to 0°, the magnetic flux vibrates at the rated amplitude around the rated peak value.
The maximum magnetic flux reaches approximately 2 pu. Owing to this excessive magnetic flux, theoperating point on the current-flux curve (the i –  curve) of the iron core enters the saturation region,
resulting in an excessive current flow, which is called an excitation inrush current.
In the circuit in Figure 1, the initial phase of phase a of the system voltage is set to 0º and the circuit
breaker is switched on at time of t = 8.5 ms. In other words, the initial phase is not 90º in any of the phases
a, b, or c, and hence an excitation inrush current is generated.

[Simulation of the 66-kV System]
The 66-kV system is simulated using a voltage source generating 66 kV and the short-circuit impedanceof the upper system. The value of the short-circuit impedance is investigated in the following document:“Current Status and the Corresponding Technology of Power Quality in the Distribution System,”Electric Technology Research Association, Vol. 60, No. 2, March 2005.
Over 210 bus bars were examined in this document, and the average value of the short-circuit impedance
observed from the 6.6 kV bus was found to be j0.0353 Ω, and hence this value was adopted for a shortcircuit. When this value is converted into the value of each phase on the 66-kV side, it becomes 6.49 mH.

102

[Transformer Simulation]
To calculate the inrush current with high accuracy, it is necessary to consider the shape and build of theiron core and winding of the transformer. However, with regard to the excitation impedance portion of thebasic equivalent circuit of the transformer (simulating the turn ratio with an ideal transformer, taking into
account the winding resistance and leakage inductance excitation impedance), there is also a simpler
method for expressing the saturation of the iron core by simulating the excitation inductance with anonlinear inductance, which is the method adopted in this example.
Because the analyzed transformer has a Y - Δ connection, the XTAP three-phase two-turn transformer
component “Y - D (leakage mH/external excitation)” has been used. This is a component in which threesets of basic equivalent circuits of the transformer excluding the excitation impedance are connected in a Y– Δ configuration and the excitation impedance can be connected from the outside. The parameter values
are as follows:・ Primary side line voltage = 66 kV・ Secondary side line voltage = 3.15 kV・ Primary side leakage impedance winding resistance = 6.68 

Leakage inductance = 337 mH

・ Secondary side leakage impedance winding resistance = 0.04  Leakage inductance = 16.1 mH
・ The capacitance-to-ground value is set in accordance with the fault value
The excitation impedance is described next. The excitation impedance in each phase is a parallel
connection of the excitation resistance of 319 kΩ and a nonlinear inductance having the following
characteristics. At up to 1.2-times the rated magnetic flux, the nonlinear inductance is 491 H. When thisvalue is exceeded, it reaches the saturation level of the polygonal line of 592 mH (air-core inductance),
which is specified by the next current-flux curve (i –  curve).

Current [A]Magnetic
flux

Magnetic flux④③

592 mH

[Wb]
①

–1.0

–206.1②–0.4192–205.8491 H

Current③

0.4192

205.8

④

1.0

206.1

592 mH②①

In XTAP, when specifying the nonlinear characteristics of the parts based on polygonal lines, thecharacteristics outside the specified range are calculated through a linear extrapolation.

103

[Analysis Conditions]
The analysis conditions are as follows.
・ Calculation step time:

0.1 ms・ Calculation end time:

200 ms・ Display start time:

0 ms

・ Display end time:

200 ms[Example of XTAP Input]
Figure 2 shows an illustration of executing this example in XTAP

Analysis Results
The results when executing this example using XTAP are shown in Figure 2. The results appear to be ingood agreement with the actual measurement results shown in the literature, except for certain differencesfrom not having considered the initial magnetic flux (residual magnetic flux). For the sake of reference, theb-phase flux waveform (obtained by integrating the b-phase voltage using XPLT) is shown in Figure 3(c).
The average value of the vibration of the magnetic flux does not become zero owing to the voltage phase at
the time of charging.

104

Example title: Calculation of excitation inrush current under pressurization of customer transformer

Systemvoltage

Short-circuit
impedance

No.: PQ-02

Transformer

Circuit breakerCircuit breaker

Customer’stransformer

Circuit breaker

66 kV System

Transformer excitation circuit

Figure 2 Example of XTAP input

105

Phase
componentization

Node voltage [V]Time [s]Branch current [A](a) Primary side voltage of transformer

Time [s]Node voltage phase component value [V-S](b) Excitation inrush current

Time [s](c) b-phase magnetic flux waveform
Figure 3 Analysis result
End

106

Change log
Date

Example fileContent changes

Version2017/12/14

2.1

Change of example file name from PQ-02 to PQ-02-A

2014/11/19

2.0

Modified for XTAP Version 2.00
Change in graph of the b-phase magnetic flux waveform in

2013/05/10

1.3

Figure 3(c) for analysis results consequent to the modificationof the integral calculation of XPLT

2012/07/19

1.2

Modified for XTAP Version 1.20

2011/10/18

1.1

Modified for XTAP Version 1.11

2010/09/02

1.0First edition created (for XTAP Version 1.10)

107

Examples of XTAPNo.

RELAY-01

ExampleDistance Relay Modeling

Name
Field
References

Power system protection
Y. Ohoura, “Protection relay system engineering,” the Institute of Electrical Engineersof JapanIn the field of protection relays, studies on the settings and co-operative judgments
are usually conducted through an effective value analysis. However, in complicatedfaults involving nonlinear phenomena such as a CT saturation, or in the relay

Overview

response analysis of faults in complex systems such as congested systems, aninstantaneous value analysis has been found to be more useful.
To conduct a relay response analysis using XTAP, it is necessary to create a relay
model for an instantaneous value analysis. In this example, we introduce a widelyused distance relay modeling method.

108

1. Application History of Protection Relays
Table 1 shows the protection relay system that is commonly used for protecting a transmission line of 66
kV or more, and for protecting the bus. The PCM current differential relay method is generally used as themain protection for the transmission line, and the distance relay method is used as backup protection.
Table 1 Protection relay system
Protection purpose

Direct grounding system

Resistive grounding system
PCM current differential relay
systemMain

PCM current differential relay

protection

systemLine selection relay
Transmission line

systemprotection(Parallel 2 wire)
BackupDistance relay system

Distance relay system

protection
Current differential relay system (collective) + current differential
Bus bar protection (double bus

relay system (divided)

configuration)

Current differential relay system (collective) + current differential
relay system (divided)

2. Distinction between Effective and Instantaneous Value Analyses
In the analysis of protection relays, it is important to distinguish between an effective value analysis,
which applies a transient stability analysis of the entire system, and an instantaneous value analysis, inwhich the voltage and current waveforms of a part of the system are analyzed and a simulation is carriedout. An effective value analysis is useful when analyzing large-scale systems and in settings with many
parameters to consider or in cooperative examinations. However, it is unsuitable for the simulation ofcomplex faults involving a phase dependency and nonlinear phenomena. An instantaneous value analysis
permits the relay response analysis of complex faults with nonlinear phenomena such as a CT saturation,and the analysis of faults in complex systems such as congested systems. However, it is unsuitable for ananalysis involving large-scale systems or multiple parameters.
jXI

3. Distance Relay
M

Distance relays are widely used as back-up protection oftransmission lines of over 66 kV (in the case of resistive

X2

grounding, however, a ground fault direction relay or aground fault over voltage relay is used as the groundX1

fault protection).
A typical example of a distance relay is described

Φ
0

RI

Figure 1 Conceptual diagram of distance

relay

109

with reference to Figure 1. Distance relays have large Mho and reactance characteristics, in which the mhocharacteristic functions as a direction element and the reactance characteristic functions as a section
selection element. The distance relay can detect a fault only through its own voltage and current input andis characterized based on features such as a relatively reliable selection of the fault section and an easy time
coordination owing to the timed cutoff at each step according to the protection section.

4. Relay Modeling
4.1 Modeling Method
There are two main methods of relay modeling. The first is a method of directly computing a vector-type
principle formula using the inner product or phase shift (direct operation) (often indicated by a vector-type
principle formula). The second is a method of computing using sampling data in the same way as the actualmachine (sampling operation). In the case of a sampling operation, it is essential to accurately know therelay algorithm, without which the data modeling cannot be conducted. In the case of a direct operation,modeling can be conducted if the vector type principle formula is known. The features of both methods areshown in Table 2.
The modeling method of a direct operation is introduced in the present example because it is relatively
easy and can be carried out without resorting to relay algorithms or constants of the analog filter.

Table 2 Comparison of direct operation and sampling operation
Direct operation

Operation
methodFeatures

Sampling operation

・Method of directly calculating inner・ Computing method using sampling values

product of vector type principle formula andin the same way as actual relays

amplitude.
・Can be constructed based on the principle

・Simulate actual relay

formula (the principle formula is common

・Operating time characteristics similar toregardless of the algorithm applied)

actual relays

・Separately set when considering

・All information from analog filter to relay

operation timealgorithm is essential
・Diverse filter compositions and relay
algorithms

1104.2 Modeling Example of Basic Relay Operation
The modeling of most protection relays can be achieved by combining three basic relay operations, as
shown below. An example of the modeling of each relay operation is shown.
Phase shift operation (PSHIFT)
This operation is implemented using a delay component. An advancement can be achieved by reversing

the delay through a curve fitting method (using the property in which the input waveform is a sinusoidalwave). For instance, a forward current of 75° can be achieved by delaying the current by 105° andreversing it (multiplying by -1).

Figure 2 Phase shift operation
Amplitude calculation
The amplitude is calculated using the moving average. Increasing the data window length generates a

calculation delay to the same extent.

Figure 3 Amplitude calculation

Phase difference calculation (inner product)

Each vector is decomposed into its components (x and y components), and the inner product is calculated.
x ∙ y = x1 y1 + x2 y2

Figure 4 Phase difference calculation

111

4.3

Example of Distance Relay Modeling

The modeling of the relay shown below is obtained by combining the relay calculations described above.


Mho relay principle for short-circuit protectionZ I exp j   V  V  0
S

ab

ab

ab

Z S Iab exp j   Vab  Vab  Vab  0 (Note that · indicates the inner product)OP

Vab・VabRES

Iabexp(jφ)・VabIabexp(jφ)

Figure 5 Modeling of Mho elements


Reactance relay principle for short-circuit protectionZ I exp j   V  I exp j   0
S

ab

ab

ab

Z S Iab exp j   Iab exp j   Vab  Iab exp j   0
OPRES

Iabexp(jφ)

Vab・Iabexp(jφ)

Iabexp(jφ)・Iabexp(jφ)

Figure 6 Modeling of reactance elements112

5. Simulation[Objective]
To ascertain the responses of the distance relay (Mho relay and reactance relay) for short-circuit
protection during internal and external faults.

[Analysis conditions]
The analysis conditions are as follows.


Calculation time step: 1 ms

Calculation start time: 0 s

Calculation end time: 0.3 sDisplay start time: 0.1 sDisplay end time: 0.3 s[Fault conditions]

2LS (Phase ab)Internal fault: Fault in the protection section of distance relay step 1External fault: Fault outside the protection section[Preconditions]

The distance relay of phase ab is the target

The current value (the value at the present moment) is used as the polarity voltage of the shortcircuiting mode (the memory effect is not taken into account)A relay locking mechanism in the small current region is not taken into account

The simulated range is restricted to a digital filter and relay operation onlyThe settling values of the relay are as follows:
44X1: 1.097 Ω [80% of the normal phase reactance (approximately 1.367 Ω in secondary side
conversion value) of the self-protection section]
44 M: 4.101 Ω [300% of normal phase reactance of self-protection section]



A timer is not attached, and the operation is instantaneous.
However, relay calculations (window length required for a phase shift or other factor) take thenecessary amount of time.

113

External fault

Internal fault

Impedance of a single section of power transmission line
Z = 17.08Ω (Primary side value)
Z = 1.367Ω (Secondary side value)

Figure 7 Analysis circuit and relay sequence[Analysis Results]
Table 3 shows the results of the “observed relay impedance (ZRY)” and “relay operation judgment” fordistance relay step 1. The results of the observed impedance of the Mho relay (F_M_ab / M_M) and theobserved impedance of the reactance relay (F_X 1 _ ab / M _ X) are shown for the “observed relay
impedance (ZRY).” The distance relay operates when the observed impedance seen from this relay exceeds
the set value.
The result of the observed impedance of the relay shows a slight difference between the behaviors of the
impedance of the Mho relay and the reactance relay. This is because their respective principle formulas, andin particular their maximum sensitivity angles, differ.
In the case of an internal fault, the impedance falls below the set value of the reactance relay (forward
reach of 1.097 Ω) and is within the set value of the Mho relay (forward reach of 4.101 Ω, and backward
reach of 0 Ω), and hence it may be inferred that the distance relay is in operation. In contrast, in the case ofan external fault, the impedance is lower than the set value of the reactance relay and the direction
characteristic, namely, the Mho relay, falls outside the set value (facing the rear side) and hence it may beconcluded that the relay is not in operation.
However, there is a delay from the occurrence of a fault to the relay operation, and it must be understoodthat this is the window length owing to the relay operation.
It is evident from these results that the distance relay is operative in the case of an internal fault andbecomes inoperative in the event of an external fault, thus confirming that it is operating normally.

114

Table: Analysis results

Control Signal

ZRY

External fault
Control Signal

Internal fault

Time (s)on flagControl Signal

OperatiControl Signal

Time (s)Time (s)Time (s)End

115

Change log
Date
2015/05/29

Example fileContent changes

version
1.0

First edition created (for XTAP Version 2.01)

116

Examples of XTAPExample
Name
Field

No.

RELAY-02

Coordinated Study of the Fault Current Interrupting Arcing Horn and Transmission
Line Protection Relay

Overhead power transmission, line protection


Otaka, Kameda, “Coordination between Fault Current Interrupting Arcing Hornand Transmission Line Protection Relay,” Central Research Institute of ElectricPower Research Report H10009, May 2011

Otaka, Iwata, Tanaka, Goda, Central Research Institute of Electric Power Industry
Report H06007, May 2007 “Development of Current Breaking Simulation Model

Literature

for Fault Current Interrupting Arcing Horn (Part 2): Development and Applicationof Arc Conductance-Dependent Model”
Otaka, Iwata, Tanaka, Goda, Tada, Tomonaga, Uemura, Central Research Instituteof Electric Power Industry Report H 05013, June 2006, “Development of Current
Breaking Simulation Model of Fault Current Interrupting Arcing Horn:
Construction of a Basic Arcing Model”
The fault current interrupting arcing horn (referred to hereafter as AH) is a lightning

protection device that instantly interrupts a fault current caused by a lightning strikeand plays a significant role in the reduction of a power failure and instantaneousvoltage drop. The short-circuit current interrupting type AH for 66/77-kV overhead
power transmission lines is capable of interrupting fault currents of up to 10 kA within
a single AC cycle. Typically, if the fault current is interrupted by an AH, there will be

Overview

no need for a transmission line protection relay (Ry hereafter) to shut off thetransmission line by opening the circuit breaker. However, it has been seen that Ry
detects a failure and shuts off the transmission line even when the fault current has beeninterrupted by an AH.
The present example analyzes the operations of an AH and Ry using the parametersof the fault occurrence point (distance from the relay installation point) and the AH
current interruption time (0.5 cycles or one cycle), and aids their coordinated study to
avoid the duplication of interruptions using an AH and Ry.

117

Analysis Circuit and Conditions
Figure 1 shows the single-line diagram of the model system. The model is a 50 Hz, 66-kV system with apower source at one end, and the transmission line is assumed to be an ACSR 520 mm2 cable, with a span
length of 20 km. The back impedance is 12.1 mH and the fault current is 10 kA when a 3LG ground fault
occurs at the nearest end of the substation (equivalent to the maximum interruption current of the AH). The
neutral point of the transformer substation is grounded through a resistor (NGR) with a capacitance of 200
A.

2

ACSR 520 mm 20 km（R = 1.12 ，X = 8.52 ）2L
66 kV，50 Hz
12.1 mH

Fault occurrence at 5, 10, or 15 km

NGR( 200 A）

1L
ab phase
2LG

51S

Fault current interrupting
arcing horn

50S

Figure 1 Single-line diagram of the model system[Analyzed Phenomenon]
Protection relays of a self-end judgment type are often installed for short-circuit/ground faults in thetransmission line protection of a 66/77-kV system without back-up protection. In general, the fault
detection time of Ry is faster for short-circuit faults, and a short-circuit relay is perhaps the only one thatcomes close to the current interruption time of an AH. Hence, the Ry studied in this example has two typesof short-circuit relays applied for the main protection, namely, an overcurrent relay (51 S) and a distance
relay (44 S).
The fault is assumed to be a 2 LG fault in the ab phase, its point of occurrence to be 5, 10 or 15 km fromthe relay installation point, and the AH current interruption time to be either 0.5 cycles or one cycle. The
reason why an AH interrupts a fault current at 0.5 cycles in certain cases and at one cycle in other cases isdue to the variations in the fault conditions (fault current, fault phase) and the surrounding environment(weather, equipment type).
The behavior of an AH and Ry at the time of a fault occurrence is analyzed based on these conditions.

118

[Fault Current Interrupting Arcing Horn Model]
We developed an AH current interrupting simulation model for the short-circuit current interruption of
the 66/77-kV system using the current–voltage waveform obtained through a short-circuit current
interruption test, the details of which are provided in the studies above.
This model combines the Cassie [1] and Mayr [2] models, which have been conventionally used as
macroscopic arc models of circuit breakers, and is capable of accurately simulating the arc conductancegenerated between the arc horns from the time of a fault occurrence to the time of its removal.
Although an AH has variations in the current interruption time, as mentioned above, with XTAP, it iseasier to consider the coordination with Ry if the zero point of the fault current interruption is clearlyspecified. For this reason, in the current interrupting simulation model of an AH, it has been made possible
to select the first or second zero point for the occurrence of the fault current interruption. In specific terms,
the Cassie model is used at the zero point at which the current needs to be continued, and the Mayr modelis used only at the zero point where it is desirable to interrupt the current. An overview of the Cassie andMayr models is provided in [3].

[ Overcurrent Relay Model]
Because the 2 LG current at the line end (20 km point) was 2,670 A, the value of 4,330 A, which is thevalue of the phase current of 2,500 A converted into a line current, (2,500  3), was taken as the settlingvalue of the overcurrent relay. The binary addition method [4] has been applied for an amplitude value
calculation, and because this calculation algorithm includes the filter effect, the application of the digital
filter operation has been omitted.
・Amplitude value calculation (binary addition method)Dm  I m  I m 3 
D51S 

1
k51S

1
I m  I m 3 ·········································· (1)2Dm  Dm 1  Dm  2  ··········································· (2)Here, D51S is the calculated value of the overcurrent relay, k51S is the gain of the amplitude generated in
the process of the calculation, and Im ＝ I1m － I2m (line current). When D51S exceeds the set value of 4,330
A, it becomes an Ry operation. The suffix m－T represents the “time before an electrical angle T  30°.”
For example, m－1 represents a delay at an electrical angle of 30°, and m－3 represents a delay at anelectrical angle of 90°. The analog filter was not taken into consideration, and the confirmation time of the
failure continuity confirmation counter was set to an electrical angle of 60°.

[ Distance Flow Relay Model]
To prevent the operation of a distance relay in response to a fault beyond the protected section owing to

119

an instrument transformer or an analog error of the relay itself, the protection range within which the main
protection can be instantly cut off is set to approximately 80% of the entire protection section [5]. Hence, inthe present example, the point at 80% of the distance from the relay installation point (distance of 16 km)
was set as the protection range, and 0.8 X = 6.81  was set as the settling value. Based on the actualcommonly used relays that we have tested thus far, a product operation with a maximum sensitivity angleof 75° [4] was used for the amplitude value, and the differential filter [4] was used for the digital filter
calculation.

・Amplitude value calculation (product operation with a maximum sensitivity angle of 75°)

D44S 

Vm2  Vm23························· (3)k 44S I m  I m1 Vm3  I m3  I m4 Vm
1・Digital filter calculation (differential filter) y m 1
 x m  x m 6 
2

(4)Here, D44S is the calculated value of the distance relay, k44S is the gain of the amplitude value generated
during the calculation, Im ＝ I1m － I2m (line current), and Vm ＝ V1m － V2m (line voltage). When D44S falls
below the settling value of 6.81, it becomes an Ry operation. In addition, xm is the input value of the digital
filter, and vm is the output value of the digital filter. An analog filter was not taken into consideration, andthe confirmation time of the failure continuity confirmation counter was set to an electrical angle of 60°.

[Simulation of Transmission Lines]
The transmission line was an ACSR 520 mm2 cable, with a span length of 20 km. For the sake ofsimplicity, the “-type equivalent circuit line model” was used in this study.

[Analysis Conditions]
The analysis conditions are as follows.
・Calculation timestep: 0.1 μs
・Calculation start time: 0 ms
・Calculation end time: 50 ms
・Display start time: 0 ms
・Display end time: 50 ms

The fault occurrence time and current interruption time can be specified by double-clicking on the AH
part, and the system frequency and setting value can be specified by double-clicking on the Ry (51S, 44S)
part.

120

[Example of XTAP Input]
The creation of this example using XTAP is illustrated in Figure 2. Because only 1L is used for theovercurrent relays and distance relays, a 2L simulation is omitted. AH and Ry (51 S, 44 S) werecomponentized, and the detailed models can be browsed by editing the component diagrams.

Example name: Coordinated study of the fault current interrupting arcing horn and transmission line protection relay
No: Relay-02 XTAB Example: Ver2.00, Example file: Ver 1.0

SystemvoltageBack
impedance

Instrumenttransformer

Instrumenttransformer

TRV adjustment
resistor/ condenser

Line type:
TACSR 520 mm2

Fault current interrupting
arcing horn

Analysis conditions
・Increments of computation time: 0.1 μs
・Calculation start time: 0 ms
・Calculation end time: 50 ms
・Display start time: 0 ms
・Display end time: 50 ms

Fault current interrupting arcing horn
operation
・Fault occurrence time: 14 ms
・ Current interruption duration: 1 half wave
Settling values of transmission
protection relay
・ Overcurrent relay (51S): 4330A
・Distance relay (44 s): 6.81Ω

Figure 2 Example of XTAP input
121

line

Analysis Results
The analysis results for the case in which the fault occurrence point is 10 km from the relay installation
point and the AH current interruption time is set to 0.5 cycles and one cycle are shown here.
The first result is the case in which the AH current interruption time is set to 0.5 cycles. Figure 3 shows
the current waveform, Figure 4 shows the voltage waveform, Figure 5 shows the 51S calculated value,
Figure 6 shows the 51S output, Figure 7 shows the 44S calculated value, and Figure 8 shows the 44S

Branch current [A]output.

Time [ms]Branch voltage [V]Figure 3 Current waveform at 0.5 cycle interruption

Time [ms]Figure 4 Voltage waveform at 0.5 cycle interruption

122

Control signal

Time [ms]Control signal

Figure 5 51S calculation value at 0.5 cycle interruption (the black line is the settling value)

Time [ms]Control signal

Figure 6 51S output at 0.5 cycle interruption

Time [ms]Figure 7 Calculated value of 44S at 0.5 cycle interruption (the black line is the settling value)123

Control signal

Time [ms]Figure 8 44S output at 0.5 cycle interruption

As can be seen from Figures 5 and 6, the calculated values of ab, bc, and ca phases for the overcurrent
relay exceeded the settling value, resulting in an Ry operation. In contrast, Figures 7 and 8 show that,
although the calculated value of the ab phase is momentarily below the settling value of the distance relay,
it does not exceed the electrical angle of 60°, which is the confirmation time of the fault continuity checkcounter, and hence Ry does not operate.
Next, we look at a case in which the AH current interruption time is one cycle. Figure 9 shows thecurrent waveform, and Figures 10–14 show the voltage waveform, 51S calculated values, 51S output, 44S

Branch current [A]calculated values, and 44S output, respectively.

Time [ms]Figure 9 Current waveform at one cycle interruption

124Branch voltage [V]Time [ms]Control signal

Figure 10 Voltage waveform at one cycle interruption

Time [ms]Control Signal

Figure 11 51S calculation value at one cycle interruption (the black line is the settling value)

Time [ms]Figure 12 51S output at one cycle interruption

125

Control signal

Time [ms]Control signal

Figure 13 Calculated value of 44S at one cycle interruption (the black line is the settling value)

Time [ms]Figure 14 44S output at one cycle interruption

It can be seen from Figures 11 and 12 that, just as in the case of the 0.5 cycle interruption, the calculated
values of the ab, bc, and ca phases for the overcurrent relay exceeded the settling value, resulting in an Ry
operation. Figures 7 and 8 show that, although the calculated value of the ab phase is below the setting
value of the distance relay, it does not exceed the electrical angle of 60°, which is the confirmation time of
the fault continuity check counter, and hence Ry does not operate.
The needless Ry operations, including their occurrence when the fault occurrence point is set at 5 or 15
km from the relay installation point, are classified and shown in Table 1. In the table, ○ indicates that aneedless operation has not occurred, whereas × indicates the occurrence of a needless operation.

126

Table 1 Occurrence of needless operation of transmission line protection relay
Fault
point

Needless operation of 51S relay

Needless operation of 44S relay

ab
Phase

bc
Phase

ca
Phase

Total

ab
Phase

bc
Phase

ca
Phase

Total

0.5 cycles

×

×

×

×

×

○

○

×

1 cycle

×

×

×

×

×

○

○

×

0.5 cycles

×

×

×

×

○

○

○

○

1 cycle

×

×

×

×

×

○

○

×

0.5 cycles

×

○

○

×

○

○

○

○

1 cycle

×

○

○

×

○

○

○

○

AH current
interruption time

5 km

10 km

15 km

From Table 1, it can be seen that, as a whole, the overcurrent relay tends to operate more than the
distance relay. Moreover, if the fault occurrence point is close to the relay installation point, a needless Ry
operation is liable to occur when the current interruption time of the fault current interrupting arcing horn isextended from 0.5 cycles to one cycle.
However, even if a coordinated operation using an Ry may be difficult to achieve, installing an AH is
extremely effective as a countermeasure against lightning damage to an overhead transmission systembecause it helps reduce the equipment damage owing to the continuous current arc at the location of thefault occurrence. We anticipate the widespread use of an AH in the future.

References[1] A. M. Cassie, “Arc Rupture and Circuit Severity: A New Theory,” CIGRE Report, No.102 (1939)[2] O. Mayr, “Beiträge zur Theorie des statischen und des dynamichen Lichtbogen,” Archiv fürElectrotechnik, Vol. 37, pp. 588-608 (1943)[3] Institute of Electrical Engineers of Japan, “Handbook of Electrical Engineering (7th Edition),”Ohmsha Ltd., pp. 886-887 (2013)[4] Protection relay system of Basic Technical Investigation Committee: “Protection Relay System Basic
System Technology,” Technical Report of The Institute of Electrical Engineers of Japan, No. 641(1997)[5] Ohoura: “Protection Relay System Engineering,” Ohmsha Ltd., pp. 174-175 (2002)

End

127

Change log
Date
2015/07/01

Example fileContent changes

version
1.0

First edition created (for XTAP Version 2.01)

128

XTAP Examples
ExampleNo.

MAC-01-A

Synchronous Generator (Single-Line 3 LG-O Fault and Constant Excitation OccurringName

in a Single-Machine Infinite Bus System)

Field

Power system applications
Because this example was created for this training course, there are no references.
※ For the modeling and system characteristics of synchronous generators, refer totextbooks on electrical equipment and electric power systems engineering such as the

Literature

following:・ Technical Report No. 798 of the Institute of Electrical Engineers of Japan,“Application technology of synchronous machine constants,”
・ N. Akira, “Application of Power System Technology Calculation,” Denkishoin
Ltd.,・ C. Concordia, “Synchronous Machine Theory and Performance,” John Wiley andSons, Inc.
The target system of the analysis is a single-machine infinite bus system connectedwith one synchronous generator and an infinite bus by two parallel transmission lines.
This example simulates the phenomenon of the occurrence of a three-phase ground

Overview

fault in one of the lines of a parallel two-line transmission, which is eliminated throughthe operation of circuit breakers at both ends of the line, and analyzes the dynamic
characteristics of the synchronous generator. The field voltage of the synchronous
generator is assumed to remain constant.

129

Analysis Circuit and Conditions
A single-line diagram of the analyzed system is shown in Figure 1. The figures in normal font in the
diagram indicate the value of the impedance, and those in italics indicate the value of the specified flow
condition.

j0.35

Synchronous
generator
同期発電機

G

Vinf=1.00∠0.0° Infinite
無限大母線
bus

0.02

CB1

CB2

j0.15

Pg = 0.9
Vt = 1.01

j0.1750.01

j0.1750.01

SW1100kVA, 220V,
50Hzベース
Missing

Figure 1 Analyzed system[Analysis Phenomenon]
In the model system shown in Figure 1, SW1 is closed at 0.5 s to generate a three-phase ground fault,
and the fault is eliminated by opening CB1 and CB2 at 0.57 s.
The field voltage of the synchronous generator is set to a constant value (gain G2 of the AVR block is set
to zero).
The machine input Tm is assumed to be given by Tm = Tm0/ω with an initial value of Tm0 and an angularvelocity of ω.

[Synchronous Generator Model]
The synchronous generator model was constructed based on the Park equation. The rotor circuit consistsof two axes, a d axis and a q axis. The saturation of the air-gap flux was not taken into consideration.
The generator constants used in this example are shown in Table 1. During the analysis, the circuit
constants of the generator model (resistance and reactance values) are automatically calculated from thegenerator constant and assigned to the circuit elements.
During the steady state initialization of this example, the node to which the synchronous generator isconnected is designated as PV, and the initial power flow on the system side is first calculated. The initialvalue of the internal variable of the generator is then automatically calculated to satisfy the various amountsof generator buses and set as the circuit element.

130

Table 1 Generator constants
Item
Unit inertial constant

Value
[s]

Synchronous reactanceM = 8.05[pu]

Xd = 1.86, Xq = 1.35

Transient reactance [pu]Initial transient reactance

Xd' = 0.44
[pu]

Xd'' = Xq'' = 0.37

Transient time constant of short circuit
Td' = 0.733[s]Transient time constant of initial short
Td'' = Tq'' = 0.048
circuit [s]Armature time constant[s]

Ta = 0.29

Armature leakage reactance [pu]XL = 0.25[Circuit Elements Other than Synchronous Generator Model]
A sine wave power supply is used to simulate the infinite bus voltage. For simplicity, the transmission
line was simulated by connecting the linear resistance and linear inductance in-series, and a linear
inductance was used in the transformer.

[Analysis Conditions]

The analysis conditions are as follows:
Overall

For an enlarged view of the instantaneous value・ Calculation time step

100 µsSame as the value on the left

・ Calculation start time

0sSame as the value on the left

・ Calculation end time

10 s

0.70 s・ Display start time0s0.45 s・ Display end time

10 s

0.70 s・ Output frequency

10

1[Example of XTAP Input]
An illustration showing the creation of the model system of Figure 1 using XTAP is indicated in Figures
2 and 3.

131

Constant field voltage

Single-machine infinite circuit

Figure 2 Model system created in XTAP

Field voltage terminal (EF)(a) d-axis circuit

132

(b) q-axis circuit(c) Armature circuit
Figure 3 Synchronous generator model

133

Analysis Results
The results of executing this example using XTAP are shown in Figure 4.
Figure 4(a) shows the three-phase instantaneous waveform of the generator terminal voltage. The
terminal voltage drops during a 3LG fault. Figure 4(b) shows the three-phase instantaneous waveform of
the generator armature current. The transient DC component is superimposed over the fundamental wave
AC component. Figures 4(c) and 4(d) show the instantaneous active and reactive powers calculated using
an α - β conversion. Owing to the influence of the transient DC component of the armature current, thevibration is superimposed immediately after the 3LG-O fault. In this case, the transient DC component is
attenuated by several hundred milliseconds. Figure 4(e) shows the effective value of the terminal voltage,and Figure 4(f) shows the field voltage. Because the field voltage is kept constant in this example, when thesystem condition changes owing to a single open line, the terminal voltage drops instead of being
maintained at 1.0 pu. Figure 4(g) shows the internal phase angle. Owing to the single open line, theimpedance of the system increases, leading to a large internal phase angle.
In addition, a back-swing phenomenon is observed in the speed deviation of the rotor of the generator(Figure 4 (h)). The relationship between the mechanical torque input and the electromagnetic torque output
at this time is shown in Figure 4(i). The electromagnetic torque output fluctuates owing to the transient DC
component, and as this torque remains greater than the mechanical torque input, the generator rotor
decelerates.

134

Node voltage [V]Time [s]Branch current [A](a) Terminal voltage [V]Time [s]Control No.

(b) Armature current [A]Time [s](c) Active power [pu]135Control No.

Time [s]Control No.

(d) Reactive power [pu]Time [s]Control No.

(e) Effective value of terminal voltage [pu]Time [s](f) Field voltage [pu]136Control No.

Time [s]Control No.

(g) Internal phase angle [deg.]Time [s]Control No.

(h) Speed deviation [%]Time [s](i) Mechanical and magnetic torque [pu]Figure 4 Analysis results using XTAP137

[Output Variables]
As a reference for creating a graph, the output variables set in the synchronous machine model of thisexample are summarized below.
Node voltage・Terminal voltage [V]Branch current
・i(Sync_Gen1/XFL_D)

Field current [puA]・i(Sync_Gen1/XL_A)

Armature current [A]Branch voltage・v(Sync_Gen1/VGA_NDT)

v(Sync_Gen1/EFD)

Field voltage [puV]Control system・s(AVR1/EF_out)

AVR Block output [puV]・s(AVR1/EFS_in)

Field voltage initial value [puV]・s(GOV1/TM_GOV)GOV Block output [puV]・s(Sync_Gen1/AG_DEG)

Internal phase angle [deg]・s(Sync_Gen1/CA_PU)

Armature current (Rms value) [pu]・s(Sync_Gen1/CK_PU)

Braking current [pu]・s(Sync_Gen1/EA_PU)

Terminal voltage（Rms value）[pu]・s(Sync_Gen1/EFD_S)

Field voltage [pu]・s(Sync_Gen1/IMD_S)

D-axis mutual reactance current [pu]・s(Sync_Gen1/IMQ_S)

D-axis mutual reactance current [pu]・s(Sync_Gen1/IG_D)

D-axis current [pu]・s(Sync_Gen1/IFD_S)

Field current [pu]・s(Sync_Gen1/IKD_S)

D-axis damper current [pu]・s(Sync_Gen1/IKQ_S)

Q-axis damper current [pu]・s(Sync_Gen1/IG_Q)

Q-axis current [pu]・s(Sync_Gen1/PG)

Active component of electrical output [W]・s(Sync_Gen1/PG_PU)

Active component of electrical output [pu]・s(Sync_Gen1/QG)

Reactive component of electrical output [Var]・s(Sync_Gen1/QG_PU)

Reactive component of electrical output [pu]・s(Sync_Gen1/SG_PER)

Speed deviation [%]・s(Sync_Gen1/TG)

Electromagnetic torque [pu]・s(Sync_Gen1/TM)

Mechanical torque [pu]・ s(Sync_Gen1/V_NDT)

Terminal voltage (Rms value) [V]

End

138

Change log
Date

Example fileContent changes

version

Modified for XTAP Version 2.00
2014/11/19

2.0

Reconnection of PV designated generator bus parts consequent
to the change in parts
Change of fault occurrence time by improving the initial2013/04/19

1.3

2012/07/19

1.2

Modified for XTAP Version 1.20

2011/10/18

1.1

Modified for XTAP Version 1.11

2010/07/16

1.0First edition created (for XTAP Version 1.10)

calculation value

139

XTAP Examples
ExampleNo.

MAC-01-B

Synchronous Generator (Single-Line 3 LG-O Fault Occurring in a Single-Machine

Name

Infinite Bus System with AVR)

Field

Power system applications
Because this example was created specifically for this training course, there are noreferences.
※ For the modeling and system characteristics of synchronous generators, refer totextbooks on electrical equipment and electric power systems engineering. Example

Literaturereference materials are provided in MAC-01-A.
AVR was created with reference to the model LAT = 1 of “Power System Dynamics
Analysis Program (Y Method).”・Technical Report No. 754 of the Institute of Electrical Engineers “Standard Model ofElectric Power System,” p. 41
The target system of the analysis is a single-machine infinite bus system connectedwith a single synchronous generator and an infinite bus using two parallel linetransmission lines.

Overview

This example simulates the occurrence of a three-phase ground fault in one of thelines of a parallel two-line transmission line that is eliminated through the operation of
circuit breakers at both ends of the line, and analyzes the dynamic characteristics of thesynchronous generator. An automatic voltage regulator (AVR) is used to control thefield voltage.

140

Analysis Circuit and Conditions[Analysis Phenomenon]
An AVR is added to the synchronous generator discussed in the example MAC-01-A. For other analysis
conditions, see this example.
An AVR is a device that controls the field voltage in such a way that the terminal voltage remains
constant. The AVR block used in this example is shown in Figure 1. An AVR can be applied by setting thegain G2 to 0.0 → 1.0.

EAS
+
EA

-

①

②

G1
1+T1s
y0=0.0

+

UL1

G2
1+T2s
y0=0.0 LL1④

EFS/G3
+
+

G4

③

UL2

G3
1+T3s
y0=EFS

EF
LL2

T4s
1+T4s

Example constants: G1 = 1.0, T1 = 0.0, G2 = 1.0 (using AVR control) or 0.0 (constant excitation), T2 =0.2,
UL1 = 100.0, LL1 = -100.0, G3 = 100.0, T3 = 2.0, UL2 = 4.0, LL2 = 0.0, G4 = 0.1, T4 = 0.5(G4 is set to 0.2 when LAT of Y method is 1)

Figure 1 AVR block

When using the control system, it is necessary to set the initial value of each block such that the desired
initial command value is output. In this example, setting the initial value of block ③ to EFS, and theinitial values of blocks ① and ② to zero, makes the initial command equal to EFS. Because the value of
EFS is automatically calculated inside the synchronous generator model and output to the measurement
terminal of the synchronous generator model, the control system can be connected to this terminal.

Analysis Results
The results of executing this example using XTAP are shown in Figure 4.
When compared with the waveform of example MAC-01-A, the field voltage in this example is seen tovary significantly owing to the function of the AVR, and after a 3 LG-O fault, the terminal voltage of the141Control No.

generator returns to the same value as prior to the fault.

Time [s]Control No.

(a) Effective value of terminal voltage

Time [s](b) Field voltage
Figure 2 Analysis results using XTAP

End

142

Change log
Date

Example fileContent changes

version

Modified for XTAP Version 2.00
2014/11/19

2.0

Reconnection of PV designated generator bus parts consequent
to the change of parts

2013/08/21

1.4

Modified AVR diagram of the example title (Figure 1)

2013/04/19

1.3

2012/07/19

1.2

Modified for XTAP Version 1.20

2011/10/18

1.1

Modified for XTAP Version 1.11

2010/07/16

1.0First edition created (for XTAP Version 1.10)

Change in fault occurrence time by improving initial value
calculation

143

XTAP Examples
Example
Name
Field

No.

MAC-01-C

Synchronous Generator (Torsional Torque Simulation)

Power system applications
Because this example was created for this training course, there are no references.
The following references are available as examples of the axis system constants. The
present example was created with reference to [1].
[1] Okamoto et al., “Evaluation of the Effect of Circuit Breaker Opening/Closing

Literature

Control on Generator/Turbine Shaft System,” Document of the Joint Study Group on
Electric Power Technology /Electric Power System Technology, PE-99-101, PSE-9998[2] “First Benchmark Model for Computer Simulation of Subsynchronous Resonance,”
IEEE Trans. Power Apparatus and Systems, Vol. PAS-96, No. 5, pp. 1565-1572, 1977
[3] J.F. Goossens et al, “Full Scale Short Circuit and Other Tests on the Dynamic
Torsional Response of Rodenhuize NR 4-300 MW – 3000 RPM Turcogenerator Part
II,” IEEE Trans. PAS, Vol. PAS-100, No. 9, pp. 4174-4185, 1981
The target system of the analysis is a single-machine infinite bus system, which isconnected to one synchronous generator and an infinite bus using two parallel linetransmission lines.

Overview

This example simulates the phenomenon of the occurrence of a three-phase ground
fault in one of the lines of a parallel two-line transmission, which is eliminated by theoperation of the circuit breakers at both ends of the line, and analyzes the torsional
torque of the synchronous generator.

144

Analysis Circuit and Conditions
The shaft configuration and constants of the generator being analyzed are shown in Figure 1.

High-pressure turbineM:

Low-pressure turbine 1

0.37 sK:

1.69 s21.4

Low-pressure turbine 2

1.83 s150.2

Low-pressure turbine 3

1.71 s

112.5

Generator

2.65 s144.1（shaft stiffness in pu torque/electrical rad）
Figure 1 Shaft configuration and constants[Analysis Phenomenon]
The torsional torque at the time of a 3LG-O fault is simulated using the model “Synchronous generator(d1q1, considering torsional torque),” which considers the axis system.
[Axis System Data]
The axis system data of Figure 1 are input into the synchronous machine model. The input screen is shownin Figure 2.

Figure 2 Shaft data input screen of synchronous generator

145

[Configuration of the Shaft System Model]（１）Calculation procedure
The various amounts of the shaft system of the generator are calculated through the following steps.
① Calculation of pu conversion and initial value of input constant② Calculation of torsional torque and damping torque③ Calculation of input torque into each mass point and output torque from each mass point④ Calculation of angular velocity ω and angle of rotation θ of each mass point（２）Equations of torque
Using the constant after a pu conversion, the equations of torque may be defined as follows. The
reference value for the unit method of torque is the rated capacitance of the generator. Subscript x indicatesthe number of mass points.
・ Inertia torque TQMx[pu]dω xdt
TQS x = CKTPx ⋅ (θ x − θ x +1 )TQDS x = RDSPx ⋅ ω x
TQDM x = RDMPx ⋅ (ω x − ω x +1 )TQM x = LMTx ⋅・ Torsional torque TQSx[pu]・ Self-damping torque TQDSx[pu]・ Mutual damping torque TQDMx[pu](３) Calculation of angular velocity ω and angle of rotation θ of mass points
The difference between the input torque and output torque is the inertia torque at each mass point. Thechange in the rotational velocity ω (speed deviation Sg) can be obtained from this relationship. Here, theinitial value of ω is the rated angular velocity (1 [pu]), and the initial value of Sg is set to zero. The
following equation is obtained by taking the example of turbine 2.
(Input to turbine 2) = (inertia torque of turbine 2) + (output torque of turbine 2)
TQIN2

＝

LMT2
Sg2 =

TQM2

＋

dω 2
= (TQIN 2 − TQOT2 )
dt

dω 2
1
(TQIN 2 − TQOT2 )
=
dt
LMT2

Here, ω and θ can be calculated from the following equations.
・ Angular velocity

ω 2 = S g 2 + ω INIT 2 = S g 2 + 1.0

・ Angle of rotation

θ 2 = ∫ S g 2 dt + θ INIT 2

146

TQOT2

Analysis Results
The results of executing this example using XTAP are shown in Figures 3–5.
Figure 3 shows the angle of rotation of each mass point. During the initial state, the input side (the left sideof Figure 1) is larger than the generator side. After a system fault, the entire shaft system accelerates while
generally maintaining this relationship.
Figure 4 shows the angular velocity of each mass point. Although the mass points rotate at the same rated
angular velocity (1 [pu]) before a system fault, they show different angular velocity fluctuations after thefault.
Figure 5 shows the torsional torque between the mass points. The value of the torque between adjacentmass points on the output side (the right-hand side of Figure 1) is shown for each mass point. The result inthis example is zero because the generator (mass point 7) is a terminal (there is no mass point on its right).
A constant torque is transmitted from the input side to the generator side before the system fault, whereas a
torsional vibration is generated after a fault.
MAC-01-C

1.6

Controlsignal
制御信号

1.5
1.4
1.3
1.2
1.1
1
0.4

0.5

0.6

0.7

Time
[s]
時間 [s]s(SyncGen1/DLTR3) #1

s(SyncGen1/DLTR4) #1

s(SyncGen1/DLTR6) #1

s(SyncGen1/DLTR7) #1

s(SyncGen1/DLTR5) #1

Figure 3 Angle of rotation of each mass point [rad]

147

0.8

Control signal

Time [s]Figure 4 Angular velocity of each mass point [pu]MAC-01-C

2

Control制御信号
signal

1.5

1

0.5

0

-0.5
0.4

0.5

0.6

0.7

0.8

Time
[s]
時間 [s]s(SyncGen1/TQS3) #1

s(SyncGen1/TQS4) #1

s(SyncGen1/TQS6) #1

s(SyncGen1/TQS7) #1

s(SyncGen1/TQS5) #1

Figure 5 Torsional torque of each mass point (between adjacent mass points on the right-hand side in
Figure 1)End148

Change log
Date
2021/11/11

Example fileContent changes

version
2.3

Modified for XTAP Version 2.30
Modified for XTAP Version 2.00

2014/11/19

2.0

Reconnection of PV designated generator bus parts consequent
to a change in parts
Change of fault occurrence time by improving initial value

2013/04/19

1.3

calculation
Combined with the power flow conditions of MAC-01-A, B

2012/07/19

1.2

Modified for XTAP Version 1.20

2010/10/18

1.0First edition created (for XTAP Version 1.11)

149

XTAP Examples
ExampleNo.

MAC-01-D

Synchronous Generator (Considering Magnetic Saturation)(Single-Line 3 LG-O Fault Occurring in a Single-Machine Infinite Bus System withName

AVR Control)

Field

Power system applications
Because this example has been created for this training course, there are no references.
※ For the modeling and system characteristics of synchronous generators, refer totextbooks on electrical equipment and electric power systems engineering. Examples of

Literaturereference materials are provided in MAC-01-A. The AVR was created with reference tothe model LAT = 1 of “Power System Dynamics Analysis Program (Y Method).”・Technical Report No. 754 of the Institute of Electrical Engineers “Standard Model ofElectric Power System,” p. 41.
The target system of the analysis is the single-machine infinite bus system connectedwith one synchronous generator and an infinite bus using two parallel line transmission

Overviewlines.
The need to increase the magnetomotive force by increasing the field current is
ascertained when considering the magnetic saturation.

150

Analysis Circuit and Conditions[Analysis Phenomenon]
The synchronous generator of example MAC-01-B is replaced with parts that consider the magnetic
saturation characteristics. For other analysis conditions, see example MAC-01-B.

Analysis Results

Control Signal

Saturation
absent
飽和なし

Saturation
present
飽和あり

Time [s](a) Terminal voltage

151

Control Signal

Saturation present
飽和あり

Saturation absent
飽和なしTime [s](b) Field current

Figure 1 Analysis results using XTAP

Figure 1 shows the analysis results. Figure (a) shows the terminal voltage of the synchronous generator, andFigure (b) shows the field current.
A terminal voltage of 1.01 [pu] and an active power output of 0.9 [pu] are set as the power flow
conditions.
A comparison of the fluctuations in the transient terminal voltage in Figure (a) indicates that thefluctuation is smaller in the event of magnetic saturation than in its absence.
A comparison of the field currents during the initial state and a steady state after a fault shown in Figure(b) indicates that, for the same terminal voltage and active power output, the need to increase themagnetomotive force by increasing the field current is greater in the presence of magnetic saturation than in
its absence, as the air gap flux decreases owing to a saturation.

End

152

Change log
Date
2015/10/27

Example fileContent changes

version
2.1

Change to the method considering saturation characteristics
Modified for XTAP Version 2.00

2014/11/19

2.0

Reconnection of PV designated generator bus parts consequent
to the change in the parts
Improved initial value calculation and simulation method of

2013/08/21

1.1

magnetic saturation
Combined with the power flow conditions of MAC-01-A, B

2010/07/19

1.0First edition created (for XTAP Version 1.20)

153

XTAP Examples

No.

MAC-02

Example3 LG-O Simulation of Induction Motor

Name
Field

Basic calculation (power systems)
N. Kibo, T. Noda, K. Takenaka, “Development of Instantaneous Value Simulation

References

Method of an Induction Motor for transient phenomena analysis,” CRIEP Report R
08022, June 2009
A3 LG - O simulation of an induction motor is calculated in this example.
The present example has the following objectives: 1) to ascertain that, in the eventof a change in the system voltage owing to a system fault such as a 3 LG, the active
and reactive powers continue to fluctuate even after the removal of the fault, and 2) toOverview

verify through the waveform that, unlike a synchronous machine, an induction motor
is not very likely to turn into the supply source of a short-circuit current, as its current
becomes attenuated during the fault period.
Please look up the references for further details such as the equipment constants. In
addition, please consult the references for a comparison of the results of analog
simulator tests, as well as a comparison of XTAP, EMTP, and Y method results.

154

Analysis Circuits and Conditions
Figure 1 shows the analysis circuit

Circuit breaker +Circuit breaker +Rated output

Circuit breaker +Circuit breaker +Voltage
Circuit breaker +Circuit breaker +Current
Number of poles+ Circuit breaker+ Circuit breaker+ Circuit breaker→Reactive power, power facter

Figure 1 Analysis circuit[Basic conditions]t = 0.01: IM parallel to the systemt = 0.04: Addition of load torque TL = 0.78
t = 0.36: Occurrence of system fault
t = 0.47: Clearing of system fault

Instantaneous drop period 110 ms = 0.11 s

Analysis Results

Systemfault
Control Signal

Control Signal

Toward a steady stateTime [s]Time [s]Induction machine PQ[W, Var]

Terminal voltage of induction machine [V] (Phasevoltage)

155

Control Signal

Control Signal

Time [s]Time [s]Induction machine ωr[pu]Three-phase current of induction machine [A]○ The secondary side of the induction motor is short-circuited and can be viewed as a variable RL loadfrom the standpoint of the electric power system. The three-phase current of the induction motor shows that,
unlike a synchronous machine, the induction motor is not very likely to turn into the supply source of ashort-circuit current as its current becomes attenuated during the fault period. Further, being viewed as aload, the rotational speed decreases when the amount of active power received from the power systemdecreases during the fault period (see ωr). Even after the fault is eliminated, the loss of the rotating bodyincreases owing to the change in slip, and hence ωr does not return quickly to a value of 1 pu (depending

Control Signal

Control Signal

on the load torque characteristic of the induction motor).

Time [s]Time [s]Vd, Vq[pu]Id, Iq[pu]The above shows the voltage/current waveform of the dq axis in the induction motor model. Because thedq axis is selected such that Vq becomes zero, the terminal voltage of the induction machine naturally
becomes the Vd.

156

X TA P
（提案手法）

時 間[sec]

EM TP

検証（３ＬＧ－Ｏ故障）

時 間[sec]

３ＬＧ－ Ｏ故障

1

1

1

0.0

1

1

1

0.20

0.20

0.40

1

0.40

0.20

1

1

0.40

1

0.40

1

1

** IM 3LG-O***

0.0

1** IM 3LG-O***

0.0

1

1

0.20

** IM 3LG-O***

0.0

1** IM 3LG-O***( C) CRIEPI

0.60

1

( C) CRIEPI

0.60

1

( C) CRIEPI

0.60

1

( C) CRIEPI

0.60

1

0.80

1

0.80

1

0.80

1

0.80

1

1.00
時間(秒)

1

1.0秒
0)
時間(

1

1.00
時間(秒)

1

1.00
時間(秒)

1

1.20

1

1.20

1

1.20

1

1.20

1

1.40

1

1.40

1

1.40

1

1.40

1

1.60

1

1.60

1

1.60

1

1.60

1

1.80

1

1.80

1

1.80

1

1.80

1

2.00

#1 1 I M
10 SMM

2.00#1101VTI M

2.00#1101QLI M

2.00#1 1 I M
10 PL

X TA P

シミュレータ試験“ す べ り ” ωrと 違 う

Y法

検証（３ＬＧ－Ｏ故障：Y法）

検証（３ＬＧ－Ｏ故障）

0. 8
0
負荷電
0.20 力有効分0.40 ( P.U.)0.60
0. PL
0
-0
. 20
0.80
負
0.荷電力無20
0.効分(
40 P. U.)0
.6
0
0.Q
00L
-0.20
1.20
1.00
VT
0.40 ﾉｰﾄ
0.ﾞ6電圧0 ( P.
0.U
8.0)
0.20
0.00
25.00
SM
5
. 0 M ｲﾝﾀﾞ
10.0ｸ
0ｼｮﾝﾓｰﾀすべ1
5. 0 り (％
20
. )00

157
0. 0

検証回路（シミュレータ試験回路）

1

Comparison of Simulator Test, EMTP Analysis, and Y Method Results (inJapanese)EndChange log
Date

Example fileContent changes

version2021/11/11

2.30

Modified for XTAP Version 2.30

2014/11/19

2.0

Modified for XTAP Version 2.00

2012/07/05

1.2

Modified for XTAP Version 1.20

2011/09/02

1.1

Modified for XTAP Version 1.11

2010/08/30

1.0First edition created (for XTAP Version 1.10)

158

XTAP Examples
Example
Name
Field

No.

PS-01-A

Simulation of the WEST 10 System of the Institute of Electrical Engineers of Japan

Power systems analysis(1) Committee for Standardization of Benchmark Power System Models, “Standard
Benchmark Power System Models,” Technical Report No. 754, 1999, theInstitute of Electrical Engineers of Japan.
(2) T. Noda, H. Takizawa, T. Nakajima, “A Study of Electromagnetic Transient
Simulations Using IEEJ’s West-10 Benchmark Power System Model,” Proc. ofReferences

International Conference on Power Systems Transients (IPST) 2015, Paper #35,
Cavtat, Croatia, 2015. Downloadable from www.ipstconf.org(3) T. Noda, O. Sakamoto, Y. Aki, “Acceleration of Power System Instantaneous
Value Analysis by Single-Phase Effective Value Simulation of Remote System,”
Journal of the Institute of Electrical Engineers of Japan, Vol. 135, No. 8, pp. 502510, 2015.
The 60-Hz system of Japan is a comb-type linking of the systems of several electric
power companies through 500-kV transmission lines in the east-west direction andextends more than 1,000 km. This system with ten generators is simulated using theWEST 10 system of the Institute of Electrical Engineers proposed in [1]. Although the
WEST 10 system was originally developed as a model system for an effective value

Overview

analysis (transient stability analysis), in the present example, an instantaneous value
analysis using XTAP is attempted. Further, in the presented set of examples, theresults of an instantaneous value analysis using XTAP are compared with the resultsof an effective value analysis based on CPAT developed in our company, and appear
to show a proper agreement. Please consult [1] for details of the WEST 10 system,
and [2] for a comparison with the effective value analysis results.

Analysis Circuit and Results
Figure 1 shows a schematic diagram of the WEST 10 system. The system frequency is 60 Hz. All ten
generators are simulated using a generator model with one damping circuit each for the d and q axes, thegenerator constants of which are typically used for thermal power plants with a large capacitance. The
inertial constant of each

159

G1

A

F

G2

G3

G4

J

G5

G6

G7

G8

G10

G9

Figure 1 Benchmark power system model of the “WEST 10 System” of the Institute of Electrical Engineers

generator model is 7 s, and each generator has a speed governor system model for the thermal and nuclear
power, as well as a rotary exciter model. The impedance of the step-up transformer of each generator is
0.14 pu based on its own rating, with a tap ratio set to 1. The lines connecting the substations are doublecircuit transmission lines of 100 km each, and the power lines connecting the power plants and substations
are double-circuit transmission lines of 50 km each. The power line of G8 alone has a length of 100 km.
The load is simulated using a model with the following characteristics: When the load end voltage is VL, theactive power has a constant current characteristic of VL ≥ 0.7 pu, and a constant impedance characteristicof VL < 0.7 pu. The reactive power is a constant impedance characteristic regardless of VL. Two cases,
namely, a daytime section (under a peak load) and a night section (under a light load), are considered forthe power flow conditions, and the power flow conditions of each generator and each load are specified foreach section. The impedances of each transmission line and transformer in each section are also set.
An instantaneous value analysis requires more detailed model data than an effective value analysis.
Because the WEST 10 system was originally designed for an effective value analysis (transient stability
analysis), the data are insufficient for an instantaneous value simulation. Therefore, the following settings
were made regarding the missing data in the instantaneous value simulation. The rated voltage of thegenerator was set to 22 kV, and the neutral point was grounded through a 0.1 Ω resistor, assuming the
ground resistance of the power plant. The Y - Δ connection was used in the transformer, with the generatorside having a 22-kV Δ connection, and the transmission side having a 500-kV Y connection. The neutral
point of the Y connection was connected to a 0.1 Ω resistor, assuming the ground resistance of thepreceding power station. Regarding the transmission line, although it would be ideal to carry out a realistic
simulation using a distributed constant line model with constant parameters, or a frequency-dependent
distributed constant line model, a π-type equivalent circuit model (1 stage) without interphase coupling wasapplied in this study. A realistic analysis taking into consideration the coupling between phases, thepropagation phenomena of traveling waves, and the frequency-dependence of the line constants are topics
for a future study.
Assuming a one-line three-wire ground fault (fault duration of 70 ms) at points A, F, and J in Figure 1 asa system fault, the daytime and nighttime sections were calculated for each case. The computational time
step was set to 100 μs for XTAP and 10 ms for CPAT.
Figure 2 shows the WEST 10 system created using XTAP for the lines described above.

160

Figure 2 WEST 10 system created using XTAP

Analysis Results
The analysis results are shown in Figures 3–5. In addition to the results of an instantaneous value
analysis using XTAP, the figure also shows the results of the effective value analysis by CPAT forcomparison. The calculation results indicate that the instantaneous value analysis results from XTAP are ingood agreement with the results of the effective value analysis using CPAT for both the daytime andnighttime sections in the cases of ground faults occurring at points A and F. The results of the ground fault
at point J are generally in agreement, but with certain differences. Figure 6 shows the calculation results ofthe fault current waveform using XTAP. Because a DC component is superimposed over the current ofphase b, there is no current zero point (zero missing) up to approximately 70 ms after the fault. Eventually,a zero point is reached for the first time and the fault current cuts off only at approximately 100 ms after theoccurrence of the fault. As a result, in the instantaneous value analysis, the fault duration time of only the bphase is approximately 100 ms, which differs from the results of the effective value analysis. In the case of
the daytime section of the ground fault at point J, shown in Figure 5 (a), a significant difference appears in
that, although generators G 6 through G 9 are not out-of-step in the CPAT results, they are out-of-step in the
XTAP results.

161

Instantaneous value analysis
瞬時値解析
Effective value analysis
実効値解析

Instantaneous value analysis
Effective value analysis

Figure 3 Comparison of the results of instantaneous value analysis using XTAP and effective value analysis
using CPAT
Ground fault: A, (a) daytime and (b) nighttime sections

Instantaneous value analysis
瞬時値解析
Effective value analysis
実効値解析

Instantaneous value analysis
瞬時値解析
Effective value analysis
実効値解析

Figure 4 Comparison of the results of instantaneous value analysis using XTAP and effective value analysis
using CPAT
Ground fault: F, (a) daytime and (b) nighttime sections

162

Instantaneous value analysis
瞬時値解析
Effective value analysis
実効値解析

Instantaneous value analysis
瞬時値解析
Effective value analysis
実効値解析

Figure 5 Comparison of the results of instantaneous value analysis using XTAP and effective value analysis
using CPAT
Ground fault: J, (a) daytime and (b) nighttime sections

Figure 6 Ground fault current waveform at ground point J: (a) daytime and (b) nighttime sections

163

Change Log
Example File
Date

Version2015/09/24

2.0

Content changes

First edition created (for XTAP Version 1.10)
Changed the load model to a model that can acquire power-

2021/03/04

2.1

flow calculation information.
Change the example name.

164

XTAP Exercises
Exercise
Name
Field

No.

PS-02-A

Simulation of EAST 10-Machine Power System of the Institute of ElectricalEngineers of Japan (IEEJ)Power system analysis[1] IEEJ Technical Committee on Standardization Model of Power System,“Standardized Model of Power System,” IEEJ Technical Report No. 754,
1999 in Japanese.
[2] T. Noda, H. Takizawa and T. Nakajima, “A Study of Electromagnetic
Transient Simulations Using IEEJ's West-10 Benchmark Power System

References

Model,” Proc. of International Conference on Power Systems Transients(IPST) 2015, Paper # 35, Cavtat, Croatia, 2015 (available for download fromwww.ipstconf.org)[3] T. Noda, O. Sakamoto, R. Yonezawa, “Enhancement of Electromagnetic
Transient Analysis of Remote Power Systems Using Single-phase Phasorbased Model,” IEEJ Transactions B, Vol. 135, No. 8, pp. 502-510, 2015.
The core of Japan’s 50-Hz grid is a 500-kV transmission line loop. The IEEJ's
EAST 10-machine power system described in reference [1] simulates this grid bymeans of a 10-machine power system. While the EAST 10-machine powersystem is a model grid primarily intended for effective value analysis (transient

Summarystability analysis), this exercise involves its use for transient analysis using the
XTAP simulator.
More information about the modeling method of power system for transient
analysis is provided in references [2] and [3] and in XTAP Example PS-01.
Details of the EAST 10-machine power system are provided in reference [1].

165Analysis Circuit and Conditions
Figure 1 shows the circuit diagram of the EAST 10-machine power system modeled on XTAP. The powersystem operates at 50 Hz. The models for all 10 generators include damper circuits on the d and q axes
respectively, with generator constants specified for thermal, hydro, and nuclear power generators. The inertia
constants are 8 seconds for the thermal and nuclear power generator models and 10 seconds for the hydrogenerator models.

IEEJ’s EAST 10-Machine Power System Model: Daytime conditions
Exercise PS-02-A (Exercise file: Ver 3.30)

Figure 1 IEEJ’s EAST 10-Machine Benchmark Power System Modeled on XTAPThe step-up transformers for all 10 generators have impedances of 0.14 pu (based on machine capacity)
but different tap ratios. Dual transmission lines are used throughout and the loop length is 1,960 km. The
loads are modeled as having constant impedance for both active and reactive powers. (This differs fromreference [1] where the loads are modeled as having constant current for active power above a certain
voltage.) Daytime load flow conditions were assumed.

166

Transient analysis requires more detailed model data than is needed for effective value analysis. Thismeans that, because the EAST 10 power system model was primarily developed for the purpose of effective
value analysis (transient stability analysis), it lacks some data needed for transient analysis modeling. Thus,
as was done for XTAP example PS-01 for the WEST 10 power system, this missing data was specified asfollows.

･

Generator rated voltage was set as 22 kV.

･

The step-up transformers were specified as Y−Δ-connected, with a 22 kV Δ connection on the generatorside and a 500 kV Y connection on the transmission line side. Also, a 0.1 Ω resistance was connected to
the Y-connected neutral to replicate the power plant ground resistance.

･The 500 / 275 kV transformers were specified as Y−Y connected with a 0.1 Ω neutral resistance.

･For the transmission lines, while a more realistic simulation using a constant-parameter or frequencydependent distributed constant model might be desirable, a (single-stage) π equivalent circuit model
without mutual interaction between lines was used in this example.

･

A grid fault was simulated as a three-phase ground fault on the N36 bus of one transmission line
(duration: 70 ms) and this was calculated for each case under daytime and nighttime conditions. The
calculation time step in XTAP was set at 100 μs.

･

The loads were modeled as a resistance and inductance (or capacitance in the case of capacitive loads)
in series, with the respective resistance and inductance (or capacitance) values being simulated by amodel in which the load was controlled to have constant impedance characteristics for both active andreactive powers regardless of the voltage across the load (VL).

167

Analysis Results
Figure 2 shows the time waveforms for the rotor angles in generators 1, 7, 8, and 10 (relative
to generator 6, which was set as the swing node). Along with the results of transient analysis using
XTAP, the results of effective value analysis by CPAT are also provided for comparison. The
results show good agreement between the XTAP tansient analysis and CPAT effective value
analysis. As this case involved generator step-out (loss of synchronism), the CPAT analysis
terminated at approximately 5.8 seconds. Accordingly, while the CPAT and XTAP results diverge
from 5 seconds onwards, their agreement prior to this time is good.

Figure 2 Comparison of Results of Transient Analysis Using XTAP and Effective Value
Analysis Using CPAT

168

Revision History
Date

Exercise File
Version2021/03/04

1.0

Revision Details
Initial document

169

Examples of XTAP

No.

CB-01

ExampleCurrent Breaking Simulation of 66-Kv Fault Current Interrupting Arcing Horn

Name
Field

Overhead power transmission, power qualityCentral Research Institute of Electric Power Industry Report H05013“Development of current breaking simulation model of fault current interrupting
arcing horn: Construction of a basic arcing model”

LiteratureCentral Research Institute of Electric Power Industry Report H06007“Development of current breaking simulation model for fault current interrupting
arcing horn (Part 2): Development and application of arc-conductance dependent
model”
A 66-kV fault current interrupting arcing horn is a lightning protection device thathelps control a power outage by cutting off short-circuit currents of up to 10 kA within
a single AC cycle. However, unlike lightning arresters used for a power transmissionthat utilize the nonlinear resistance characteristics of zinc oxide elements, a shortcircuit current circulates with almost the same magnitude despite being within a single

Overview

AC cycle and impacts the power system by leading to an instantaneous drop in voltage.
This example assumes a case in which a three-phase ground fault (3 LG) has beencleared using a fault current interrupting arcing horn and analyzes the current–voltage
waveform at the location where the arcing horn is installed and at the substation serving
as a power source. In addition, because a fault current interrupting arcing horn may
interrupt the current at the first (0.5 AC cycle) or second (one AC cycle) zero point, thisanalysis was designed to specify the zero point at which the current is cut off.

170

Analysis Circuit and Conditions
Figure 1 shows a single-line diagram of the model system conducting a current breaking simulation of a
fault current interrupting arcing horn; the poles of the 66-kV transmission line are as shown in Figure 2.

2L

50 Hz，66 kV

66kV Transmission line:

2

66 kV 送電線 ：TACSR 510 mm ，10 km

TACSR510mm2, 10km

12.1 mH

1L

3 LG fault
三相地絡故障（3LG）

Figure 1 Single-line diagram of the model system

1.8 m

Overhead ground
架空地線
AC 90 wiremm2

35.9 m2.2 m

31.0 m2.8 m

28.2 m

Power
line
電力線
TACSR 520 mm2

2.4 m

25.8 m

亘長：10 km
Span length
大地抵抗率：100 Wm
Earth resistivity

Figure 2 Poles of the 66-kV transmission line

The model is a 50 Hz, 66 kV system with a power source at one end, where the transmission line isassumed to be an ACSR 520 mm2 cable, with a span length of 10 km. Although it is a double-circuit
transmission line, the circuit breaker on the 2L side is open, and only the 1L side is used. The back
impedance is 12.1 mH, and the short-circuit current is 10 kA, when a three-phase ground fault occurs at thenearest end of the substation (equivalent to maximum cutoff current of fault current interrupting arcinghorn).

[Analyzed Phenomenon]
In this example, a three-phase ground fault (3 LG) occurring in the terminal tower of transmission line
1L is simulated. The tower is fitted with a fault current interrupting arcing horn, and it is assumed that thecurrent of each phase is cut off at the second zero point (one cycle of AC) after the occurrence of a fault.

171

Please refer to the two documents mentioned above for the current breaking simulation model of thefault current interrupting arcing horn. As shown in Figure 3, it is possible to simulate the arc conductance
between the arc horns from the occurrence of a fault up to its clearing using a combination of the Cassie
and Mayr models, which are widely used as macroscopic arc models of circuit breakers.

Cassie
線路側
Line sideホーン
horn

Ground接地側
side
horn
ホーン

Line side
線路側
horn
ホーン

Ground
side
接地側
horn
ホーン

Mayr

Crossfeed
クロスフェード
switch切り替え

MayrモデルModel
Mayr

Short circuitcurrent
短絡電流

Cassie Model
モデル

時間
Time

0

Figure 3 Current breaking simulation model of fault current interrupting arcing horn
[Creation of a Transmission Line Model]
As shown in Figure 2, the transmission line studied consists of a 90 mm2 AC overhead ground wire and a
520 mm2 TACSR power line. For simplicity, a “fixed-parameter distributed constant circuit” is used in thisexample.
[Analysis conditions]
The analysis conditions are as follows.
・Calculation timestep: 0.1 μs
・Calculation start time: 0 ms
・Calculation end time: 50 ms
・Display start time: 0 ms
・Display end time: 50 ms
Further, the time of occurrence of a fault and the current interruption time can be specified by doubleclicking on the part (FCIAH) of the fault current interrupting arcing horn. In the present example, thefollowing settings are applied.
・Time of fault occurrence:14 ms
・Current interruption time: 2 half waves (current is cut off at the second zero point from the fault
occurrence)[Example of XTAP Input]

172

The creation of this example using XTAP is illustrated in Figure 4.

Example name: Current breaking simulation of
66-kV fault current interrupting arcing horn

No.: CB-01

Line type: TACSR 520 mm2
Span length: 10 km
TRV adjustment resistor.
Condenser

Constant parameter

XTAP Example: Ver. 2.00 Example File: Ver. 2.1

Constant parameter

Fault current interrupting
arcing horn

Line

Line

Line

GND

GND

GND

Back impedance
System frequency
System voltage
Analysis conditions
・Increments of computation time：0.1 μs
・Calculation start time：0 ms
・Calculation end time：50 ms
・Display start time：0 ms
・Display end time：50 ms

Fault current interrupting arcing horn・Time of fault occurrence: 14 ms
・Current interruption time: 2 half waves

Figure 4 Example of XTAP input

Analysis Results
First, the execution results of XTAP for the present example are shown in Figure 5.

173

Short-circuit current (kA)(a) Short-circuit current

Bus voltage (kV)Time [ms](b) Bus voltage of substation(c) Voltage between arc hornsof fault current interrupting arcing
horn

Voltage between arc horns [V]Time [ms]Time [ms]Figure 5 XTAP execution result (current interruption time of fault current interrupting arcing horn: 2 halfwaves)

Next, the execution results from XTAP when the current interrupting time of the fault current
interrupting arcing horn is set to 1 half wave (when the current is interrupted at the first zero point from theoccurrence of a fault) are as shown in Figure 6.

174

Short-circuit current (kA)(a) Short-circuit current(b) Bus voltage of substation

Bus voltage (kV)Time [ms](c) Voltage between arc hornsof fault current interrupting arcing
horn

Voltage between arc horns [V]Time [ms]Time [ms]Figure 6 XTAP execution result (current interruption time of fault current interrupting arcing horn: 1 half
wave)

End

175

Change log
Date

Example fileContent changes

version2015/07/29

2.1

Name of example changed from PQ - 03 to CB - 01

2014/11/19

2.0

2014/08/19

1.2

Rectified certain problems

2012/07/19

1.1

Modified for XTAP Version 1.20

2011/10/18

1.0First edition created (for XTAP Version 1.11)

Modified for XTAP Version 2.00
Recalculated line constant consequent to change of XTLC176

XTAP Examples
Example

No.

CB-02

Analysis of Terminal Short-Circuit Breaker and Capacitive Current Switching in 154-

Name

Kv System

Field

Basic calculation (electric power system)


Noda, Ohtaka, Yonezawa, Matsumoto, Matsubara, Kuzuna, “Investigation into
Instantaneous Value Analysis and Transient Phenomena Analysis Methods forElectric Power System and Analysis Example using XTAP (Part 1): Analysis ofswitching surge overvoltage,” Central Research Institute of Electric Power
Research Report H12005, June 2013

LiteratureApplication Guidelines for the Revised Standards of AC Circuit Breakers andSpecial Committee for the Investigation of Cut Off Duty in Electric Power
Systems, “Circuit Breaker Standard JEC-2300 Application Guide: Survey on
Interruption of Electric Power System and Background of Revision of Standards,”
Technical report of The Institute of Electrical Engineers of Japan, No. 1200,
September 2010
When the circuit breaker installed in a substation interrupts the current, a transient

recovery voltage (TRV) having an oscillating component ranging from several hertz toseveral kilohertz is generated between the poles of the circuit breaker. The TRV is avery important parameter in discussing the current interrupting performance andinsulation recovery characteristics of a circuit breaker. Further, because a power system

Overview

exhibits various behaviors depending on its configuration and constants, aninstantaneous value analysis is effective in studying the TRV.
In this example, an analysis of the “Break Terminal Fault interruption” (CB-02BTF.xsf), which breaks a fault occurring within the vicinity of the circuit breaker in a
154-kV system, and an analysis of the “Small Capacitive Current Breaker” (CB-02 CCS.xsf), which shuts off the capacitor bank connected to the substation tertiary
circuit, are carried out.

177

Analysis Circuit and Conditions
A single-line diagram of the model system is shown in Figure 1. The current value in the figure
represents the value of the short-circuit current flowing from the power source. The target of the analysis is
the substation T shown in the figure.

275 kV△
T 変電所
Substation
T

Ｙ
Ｙ

△

Ｙ
Ｙ

△

Ｙ
Ｙ

21.0 kA

154 kV
0.2 kA × 2 Lines
回線

Lines
0.8 kA × 22回線

24.1 km回線
22 Lines

26.4 km回線
22 Lines

154 kV
T－1 変電所
Substation
T-1 △154 kV
Ｙ
Ｙ

△

77 kV

Ｙ
T－2
変電所
Substation
T -2
Ｙ
77 kV

Figure 1 Single-line diagram of model system[Phenomenon Being Analyzed]
As mentioned above, when the circuit breaker interrupts the current, a TRV with a vibration component
ranging from several hertz to several kilohertz is generated between the poles, which is a very importantparameter in discussing the current interrupting performance and insulation recovery characteristics of acircuit breaker. Based on the breaking phenomenon, various test methods have been defined for an
evaluation of the TRV. These are described in detail in JEC-2300-2010 “AC circuit breaker” in accordancewith the standards of the Japanese Electrotechnical Committee of the Institute of Electrical Engineers ofJapan (JEC), and in IEC 62271-100 “High-voltage switchgear and control gear - Part 100: Alternatingcurrent circuit-breaker” in accordance with the standards of the International Electrotechnical Commission(IEC).
The breaking of a fault occurring within the vicinity of the circuit breaker, as shown in Figure 2, is calleda terminal short-circuit fault interruption (hereinafter referred to as a breaker terminal fault (BTF)). Because

178a BTF has the maximum current flowing from the rear power source, it may be stated that it has the main
responsibility of protecting the power system from ground and short-circuit faults. Here, it is assumed thata fault occurring within the vicinity of the substation T is broken by the circuit breaker installed on its
secondary side.

Circuitbreaker
遮断器

Rear powersource
背後電源

Line
線路

故障
Fault

Figure 2 Illustration of terminal short-circuit fault

A case in which the capacitor bank is cut off, as is shown in Figure 3, is called capacitive current
switching. Figure 4 shows the current and voltage waveforms at the time of a capacitive current switching.
When the circuit breaker interrupts the current, the crest value of the circuit breaker terminal voltage UTremains as the DC voltage UL on the load side (line or capacitor bank side) and a 1-cos type voltage, whichis the difference between UT and the voltage US of the power supply side, which appears as the interelectrode voltage UCB of the circuit breaker. Because an interrupting current is small compared with a shortcircuit current and ground fault current, a current interruption may be completed even when the breakerpoles are not sufficiently opened. If a high recovery voltage occurs in such a state, it may exceed thedielectric strength between the poles of the circuit breaker, leading to the possibility of a re-ignition. In
particular, when the breaker re-ignites near the crest value of the inter-electrode voltage, a very highovervoltage may occur, and caution is therefore necessary. Here, it is assumed that the circuit breakerinstalled on the tertiary side of substation T cuts off the capacitor bank connected at that end.

線路
Line
Rear powersource
背後電源

Circuitbreaker
遮断器

コンデンサ
Capacitor bank
バンク

Figure 3 Illustration of capacitive current switching

179

i
t
v
UCB
UL

UT

t
US
Figure 4 Illustration of TRV waveform of capacitive[Power Supply Simulation]
Based on [1], the power source, resistance, and inductance of the primary side of substation T; thesecondary side of substation T-1; and the secondary side of substation T-2 were set to match the shortcircuit current values. Because the power factor of the power supply circuit is not known, the resistance andinductance were adjusted such that the decay time constant of the DC component (τ = L / R) will be 45 ms.
In addition, the resistances and capacitances were inserted in parallel such that the TRV of the blocking firstphase corresponds to JEC-2300-2010.

[Transformer Simulation]
The capacity, winding resistance, and leakage inductance were set based on [2], and the capacitance wasset based on [3]. The neutral point is directly grounded on the 275-kV side and resistance grounded on the154- and 77-kV sides.

[Line Simulation]
The type of electric wire used was determined based on the data of the 154-kV standard transmission
line listed in [4].
A 810 mm2 TACSR wire was used for the power lines (two conductors) and a 90 mm2 AC wire was used
for the overhead ground wire, the ground resistivity of which was 100 Ωm. The simulation was conductedusing a frequency-dependent line model, the sample range of which was 0.1 Hz to 10 MHz, and thesamples size was 400.

[Busbar Simulation]
Because the ground capacitance of individual devices was not known, each bus was collectively
provided with a ground capacitance of 40,000 pF.

180[Capacitor Bank Simulation]
A 60-MVA capacitor bank was installed in the tertiary circuit of the transformer of substation T. In linewith JIS C 4902 (1998) “High pressure and special high-voltage phase advancing capacitor and accessory
equipment,” a 6% series reactor was also installed for the capacity of the capacitor bank.

[Analysis Conditions]

The analysis conditions are shown in Table 1.

Table 1 Analysis conditions

Analysis of Break Terminal
Fault interruption

Analysis of capacitive current
switching

Increment in calculation time

0.1 μs1 μs

Calculation start time

0 ms

0 ms

Calculation end time

200 ms

100 ms

Display start time

40 ms

20 ms

Display end time

160 ms

80 ms[Example of XTAP Input]
An illustration showing the creation of this example using XTAP is provided in Figure 5.

181

Circuit breaker

Secondary

Circuit breakerCircuit breaker

T11
Frequency dependent

Circuit breaker

182

Tertiary

Primary Secondary

Tertiary

Primary Secondary

Figure 5 XTAP input

Tertiary

Primary

Tertiary

Primary

Secondary

Circuit breakerCircuit breakerCircuit breakerCircuit breakerCircuit breakerCircuit breaker

Secondary

Tertiary

Primary

Circuit breakerCircuit breakerAnalysis Results
First, the analysis of a BTF interruption in a line from substation T to substation T-1 was carried out. Let
us consider a case in which a short-circuit failure occurs at 50 ms in a one-line three-phase ground fault
near substation T, and the circuit breaker of the substation opens at 150 ms. Figure 6 shows the current
waveform of the circuit breaker, and Figure 7 shows the TRV waveform of the first phase of interruption (cphase). The TRV was a four-parameter waveform of multiple frequencies, where the crest value uc was 249

Branch current (kA)kV, and the crest time t2 was 307 μs.

Time [ms]Branch current (kA)

Figure 6 Current waveform of circuit breaker

Time [ms]Figure 7 TRV waveform at the time of BTF interruption

183

Next, an analysis of the capacitive current switching in the tertiary circuit of the transformer of substation
T was conducted. We considered a case in which the circuit breaker on the tertiary side of the transformer
opens at 50 ms and interrupts the current flowing through the capacitor bank. Figure 8 shows the current
waveform flowing through the circuit breaker, and Figure 9 shows the TRV waveform generated in thecircuit breaker. The breaking current was 719 A. The TRV crest value of the first phase of interruption (cphase) is 64.2 kV, which is 2.40-times higher than the ground voltage crest value, namely, 26.8 kV, of the

Branch current (kA)tertiary circuit of the transformer.

Time [s]Branch current (kA)

Figure 8 Current waveform at the time of capacitive current switching

Time [s]Figure 9 TRV waveform at the time of capacitive current switching184

References[1] IEEE Working Group 15.08.09, “Modeling and analysis of system transients using digital programs,”
IEEE PES Special Publication, TP-133-0, 1998
[2] Japan Electrical Engineers Association Home Page, http://www.jeea.or.jp/course/contents/03102/, asof January 2013[3] Modeling research committee for sophistication of lightning surge evaluation of The Institute ofElectrical Engineers, “Modeling for Sophistication of Lightning Surge Evaluation,” Technical Reportof The Institute of Electrical Engineers of Japan, No. 704, November 1998
[4] Lightning Protection Design Committee, Power Generation and Transformation Subcommittee,“Lightning Protection Design Guide on Transformer Substation and Underground Transmission Line,”
Central Research Institute of Electric Power Industry Report General Report, T 40, December 1995

End

185

Change log
Date

Example fileContent changes

versionName of example changed from STRV-01 to CB-02

2015/07/29

2.1

2014/11/19

2.0

Recalculated line constant consequent to change of XTLC

2013/08/08

1.0First edition created (for XTAP Version 1.21)

Fixed bugs in three-phase three-wire transformer

186

XTAP Example Collection
Example
Name
Field

No.

SLD-01

Flashover Analysis of 6.6-kV Distribution Line

Lightning surge analysis, overvoltage analysis
For general information on the lightning protection design of medium-voltage lines,
refer to the following.
¾

References“Distribution Line Lightning Protection Design Guide,” Comprehensive Report
T69, Central Research Institute of Electric Power Industry

For information on the analytical model of concrete poles, refer to the following.
¾“Development of Distribution Line Lightning Surge Analysis model for Steep
Wave Lightning Current,” Research Report H07002, Central Research Instituteof Electric Power Industry
Flashover may occur due to lightning overvoltages caused by direct lightning

strikes on medium-voltage distribution lines. Hence, it is necessary to protect themedium-voltage distribution line from lightning by effectively installing lightning
protection equipment, such as an overhead ground wire and surge arrestor. In thisexample, we show the lightning surge analysis model to calculate the overvoltage

Outlineacross the insulators when lightning directly strikes a medium-voltage power
distribution line with the overhead ground wire.
With XTAP, because the concrete pole, phase conductor, pole-mounted transformer,
and other models necessary for a lightning surge analysis of power distribution lines
are registered as components, through their combination, a lightning surge analysis of
medium- and low-voltage power distribution lines becomes possible.

187Analysis Circuit and Conditions
Figures 1 shows the analysis circuit and arrangement of each conductor. The analysis line is a single-circuit
6.6 kV medium-voltage distribution line with a diameter of 40 m. Both ends of the line are terminated withmatching resistors to prevent the unintended reflection. The calculations assume a lightning strike on a utility
pole located in the center of the line.

Striking
point
雷撃箇所

0.3 m

0.3 m

40m
400m

0.6 m

11 m 12 m

Matching
整合抵抗
resistanceMatching
整合抵抗
resistance
Ground wire
：架空地線
Phase
：相導体
conductor

Figure 1

Topology of examined line.

Each equipment of the distribution line is modeled in the following way.
¾

Phase conductor
Calculate the line constants of each conductor shown in Figure 1 by means of XTLC, a line constantcalculation program provided with XTAP.
For the case of focusing only on flashover, as in this example, the calculation frequency of the line
constants can be set to about 100 kHz using the constant parameter distribution constant parameterline model (CP-Line), which is accurate enough for practical use.
In this example problem, the calculation frequency and ground resistivity are set to 100 kHz and 100
Ωm, respectively.
Note that in cases where a long calculation time is required, such as the surge arrester failure
calculation, it is necessary to use the calculation using the frequency-dependent line model (FD-Line).

¾

Concrete pole model
For direct lightning surge analysis, the grounding conductor and the reinforcing bar of the concretepole are often modeled by a single-phase lossless constant-parameter line model.
In this example, a constant parameter line model (CP-line) is used to simulate the surge impedanceof the concrete pole and the grounding conductor¾

Lightning current waveform
Lightning current waveforms are often modeled by a triangular waveform as shown in Figure 2(a).
On the other hand, in recent research current waveforms represented by the Heidler function (Figure2(b)), are also often used as a more realistic waveform.

188

Heidler function is expressed by the following function.

(n=1，η1=0.92，n1=10，τ11=4.6×10-6，τ12=95.1×10-6 in Figure 2(b) )
/
1
/
1

exp

/
exp

/

/

/

1

/

‥

/
12

12

10

Current[kA]
波高値 [kA]

10

Current[kA]
波高値 [kA]

exp

/

8
6
4

8
6
4
2

2

0

0
0

20

40
60
時間 [μs]

80

0

100

(a) Triangular waveform
Figure 2

20

40
60
時間 [μs]80

100

Time [μs]Time [μs](b) Heidler function waveform

Lightning current waveform(Wavefront duration: 2 μs，Wavetail duration: 70 μs）

In this example, lightning channel impedance is assumed to be 1000 Ω in this example.
¾

Flashover of insulator
The integral method model is used in this example to simulate the flashover of a distribution insulator.
In this model, flashover occurs when the integral value B of the portion of the voltage waveformgenerated in the insulator above V0 exceeds a threshold value. Although the integral method model isinferior to the constant voltage model in terms of simplicity, it is suitable for multiphase flashover
analysis because it can take into account the influence of insulator voltage waveform. In this example,
the flashover model constant (V0 = 135 kV, B = 0.065) is set based on the lightning impulse test results.

(a) Constant voltage model
Figure 3(b) Integral Method model

Overview of insulator flashover model

189

Flashover voltage [kV]
フラッシオーバ電圧350
実験結果
ExperimentCalculation
計算結果

300
250
200
150
100
0

Figure 4246Time to flashover [µs]フラッシオーバ時間

8

10

Example of V-t characteristics of a distribution insulator.

In this example. a surge arrester for medium voltage distribution line is represented by a nonlinear
resistor with series gap. The discharge voltage of series gap is set to be 29 kV and the V-I
characteristic of nonlinear resistor is set as shown in Figure 4.

4030
Voltage [kV]

20
10
0
-10
-20
-30
-40
-10000

-5000

0

5000

10000

Current [A]Figure 5 V-I characteristics of surge arrester.

Figure 6 shows this example problem configured on XTAP under the above calculation condition.

190

Concrete pole model
To Overhead
ground wire

To phase
conductors

surge arrester

Concrete pole and grounding conductor
Grounding resistancePole withsurge arresters

Pole withsurge arresters

Figure 6

Pole withsurge arresters

Overview of calculation model in XTAP.

191

Calculation results
Figures 7 and 8 show the calculated results of the lightning current waveforms and the voltage across theinsulators on poles 1 through 5. Figures 7(b) and 8(b) show the voltage between the insulators without surgearresters. The insulator voltage obtained by the triangular waveform is larger than the insulator voltage
obtained by Heidler function waveform. From this, it can be understood that when evaluating the conservativeside (where the insulator voltage is calculated to be larger), the strike current should be modeled as atriangular wave. As shown in Figs. 7(c) and 8(c), the insulator voltage at the lightning struck pole (V_insu13)
is suppressed by the installation of surge arresters.
Figure 9 shows the flashover voltage waveform when the lightning current is a triangular wave. The
flashover current value increases from 15 kA to 50 kA due to the installation of surge arresters. Also, thelocation of the flashover point has changed from Pole No. 1 to Pole No. 3.
In this example, only 1-phase flashover is calculated, but since the 6.6 kV distribution line in Japan is an
ungrounded neutral system, lightning outages do not usually occur with 1-phase flashover. For this reason,in lightning outage rate calculations, multi-phases flashover is often considered to be lightning outage.

192

(a) Lightning current waveform(b) Insulator voltage from Pole. No.1–No.5 (without surge arrester)(c) Insulator voltage from Pole. No.1–No.5 (with surge arrester)Figure 7

Calculation results obtained by the triangular lighting current waveform

193

(a) Lighting current waveform(b) Insulator voltage from Pole. No.1–No.5 (without surge arrester)(c) Insulator voltage from Pole. No.1–No.5 (with surge arrester)Figure 8

Calculation results obtained by the Heidler function lighting current waveform

194

(a) Case without surge arrester (Peak value of lightning current: 15 kA, Flashover at Pole No. 1)(b) Case with surge arrester (Peak value of lightning current: 50kA, Flashover at No. 3)
Figure 9

Insulator voltage when the flashover occurs obtained by the triangular lightning current

waveform.
END

195

Revision History
Date2022/12/13

Example File

Content Changes

Version
2.1

Corrected the flashover current value to 50kA when lightning
arrestors are installed.
Revised example contents with addition of analytical models

2022/08/03

2.1

such as lightning current waveforms represented by Heidler
functions and a flashover model represented by the integration
method.
Update for XTAP Ver. 2.00

2014/11/19

2.0

2012/08/30

1.2

Update for XTAP Ver. 1.20

2011/09/08

1.0

Creation of first edition (for XTAP Ver. 1.11)

Recalculation of line constants following the change in XTLC196

XTAP Example Collection
ExampleNo.

SLD-02

Lightning Surge Aspects Occurring in Secondary Side of a Pole-mounted Transformer

Name

During Lightning Strike on a Distribution Line

Field

Lightning surge analysis, overvoltage analysis
For information on the analytical models of pole-mounted transformers, service lines
and indoor wiring, see the following.
·

Honda, Noda, Asakawa, Shindo, Yokoyama, Abiko, “Improvements to a
Transient Simulation Model of Pole-Mounted Distribution Transformers,”

References

Journal of the Institute of Electrical Engineers of Japan, B, Vol. 124, No. 9,
pp. 1169-1176 (2004)·

Matsuura, Noda, Nakamura, Sakai, “Modeling of Service Drop-Wires andInterior Wiring Cables for Lightning Surge Analysis,” Journal of the Instituteof Electrical Engineers of Japan, B, Vol.130, No. 2, pp. 246-258 (2010)

With the recent progress of our highly information-oriented society, damage toconsumer equipment from lightning is on the rise. Even in power distribution
systems, the number of electronic devices that are driven at low voltages isincreasing. It is therefore important to accurately grasp the aspects of a lightning
overvoltage (current) occurring in low-voltage distribution lines and consumer
appliances after the watt hour meter (WHM) at the time of a lightning strike on a

Outlinedistribution line. In this example, we calculate the lightning overvoltage occurring on
the WHM and the lightning current flowing into consumer appliances when lightning
directly hits a distribution line.
With XTAP, because the power line, pole-mounted transformer, and other models
necessary for a lightning surge analysis of low-voltage distribution lines are registered
as components, through their combination, a lightning surge analysis of low-voltage
distribution lines becomes possible.

197Analysis Circuit and Conditions
In this example, we examine the occurrence of a lightning surge on the secondary side of a pole-mounted
transformer when a lightning strike occurs on a concrete pillar equipped with a pole-mounted transformer.
It should be noted that, despite there being two types of grounding methods for pole-mounted transformers
and lightning arresters, namely, “individual grounding” in which each grounding is separately secured, and“shared grounding” in which the grounding is combined, herein, we examine “shared grounding” only.

High
Lightning arrestor

voltageline

Service line
Consumer’shouse

Figure 1 Conceptual diagram of lightning surge analysis from the secondary side of pole-mounted
transformer·

Analytical Model of Concrete Pillar

In the lightning surge analysis of a distribution line, because a concrete pillar and pulled grounding line
are often regarded as a single conductor, they are usually simulated using a constant single-phase
distributed line. In this study, we adopt a simple concrete pillar model in which a constant single-phase
distributed line is divided into three positions (pole head, high-voltage cross-arm, and pole-mounted
transformer equipment positions). (Because a concrete pillar model capable of accurately calculating thelightning over-voltage of a high-voltage distribution line up to the order of n seconds is registered in
XTAP, its use is sufficient when calculating high-voltage distribution lines only.)
Further, for overhead ground lines and high-voltage distribution lines, a model that calculates the line
constants with the arrangement shown in Figure 2 is generally prepared, and can be used in thecalculations by simply determining the line length.

198

GW：
22 mm22
High-voltage
OC 60 mm：main line

0.8 m

0.8 m

12.75 m11.4 mFigure 2 Conductor arrangement in high-voltage distribution line·

Analytical Model of a Pole-mounted Transformer

For pole-mounted transformers, an analytical model that can accurately calculate the transition voltage
from the primary side to the secondary side has been prepared in XTAP (there are two types of models
with different winding structures). The pole-mounted transformer model has connection points at the
primary side terminal, secondary side terminal, and the chassis ground, and can be used in thecalculations by simply connecting them.

+ Terminal+ Terminal

Secondary side

Primary side
0 Terminal- Terminal- Terminal

Exterior terminal

Figure 3 Example of the analytical model of pole-mounted transformer·

Analytical Model of Service Line and Indoor Wiring

In the DV electric wires used for service lines and VVF cables used for indoor wiring, because thedistance between conductors is small, the mutual coupling becomes strong. It is therefore difficult toobtain the accurate values of the line constants through a simple calculation using XLTC. In XTAP, aservice line model in which the line constants are obtained when considering this mutual coupling has
been prepared and can be used for the calculations by simply determining the line length.
·

Analytical Model of WHM

199

There are two types of WHM: mechanical and electronic. Between them, a mechanical WHM isusually omitted because it does not have an influence on the lightning surge analysis. In contrast, in anelectronic WHM, a varistor (ZnO lightning protection element) is installed between the lines, and isoften simulated by simply inserting a non-linear resistance between the lines simulating the V-I
characteristics. In this example, both mechanical and electrical WHMs are studied.

Voltage [V]Current [A]Figure 4 Example of V-I characteristics of varistor between the lines of electronic WHM·

Analytical Model of Consumer Home

Because the detailed circuit of a consumer home is unknown in most cases, it is difficult to create astandard model. Hence, in this analysis, as a simplified simulation, we conducted the calculation by
assuming a house connected to two grounding devices (an air conditioner and a washing machine,namely, an inter-line varistor and a ground varistor) at the end of a VVF cable (10 m).

Voltage [V]Voltage [V]Current [A]Current [A](a) Inter-line varistor(b) Ground varistor

Figure 5 Example of the V-I characteristics of varistor used in consumer equipment·

Other Analysis Conditions·

Lightning current: 1/70s Ramp wave, Crest value: 10 kA, Lightning impedance: 1000

·

Ground resistance of the lightning arrestor: 30 (in a lightning-struck pillar, the grounding of thelightning arrestor and pole-mounted transformer is shared)·

Calculation step time: 1 ns·

Calculation start time: 0 s

200

·

Calculation end time: 10s

Figure 6 shows the construction of this example using XTAP under the above conditions.

201

Pole-mounted Transformer Model
Lightning Current Waveform

Branchcurrent [kA]

Time [us]Source text - Translation
マッチング – Matching
配電線 – Distribution line
バリスタ - Varistor

Lightning
current model

タイプ type

Highvoltage
cross-arm
position

Position oftransformer
facility

Consumer home

WHM Model (omitted in case of mechanicalConsumer Equipment Model

type)

Figure 6 Example of XTAP input

Analysis Results

202

Potential difference between thenodes [kV]

Mechanical watt-hour
meterElectronic watt-hour meter

Time [us](a) WHM Inter-line Voltage

Unit is [A]Control signal

The calculation results of mechanicaland electronic type overlap

Time [us](b) Current flowing into the consumer home (three lines together)
Figure 7 Analysis results using XTAPFrom the calculation results, it can be seen that with an electronic WHM, the lightning overvoltage
generated between lines can be suppressed to the limit voltage by the varistor. However, unlike
conventional mechanical measuring instruments, because electronic WHMs have inbuilt circuit boards suchas a CPU that are vulnerable to lightning (see the references below for examples), from this result alone, itcannot be stated that the lightning protection performance of electronic instruments is better than that of
mechanical instruments; rather, it is necessary to examine factors other than the overvoltage (induction
voltage generated inside the instrument) obtained from the circuit calculation.
·

Studies on the Lightning Protection Performance of Electronic Watt-hour Meters

Asakawa, Furukawa, Takahashi, Ishimoto: “The Factors and Counter Measures of Failures inElectronic Watt-Hour Meters Caused by Lightning: An Experimental Study on Electromagnetic
Interference Prevention Method,” Journal of the Institute of Electrical Engineers of Japan, B, Vol. 131,
No. 9, pp. 793-800 (2011)
END

203

Revision History
Date

Example File

Content changes

Version2014/11/19

2.0

2012/10/03

1.1

2012/09/18

1.0

Revision for XTAP Version 2.00
Correction owing to an error in the simulation method ofelectronic instruments
Creation of first edition (for XTAP Version 1.20)

204

XTAP Example CollectionNo.

SLS-01

ExampleLightning Surge Calculation of 500-kV Air-insulated Substation

Name
Field

Lighting surge analysis of substations, overvoltage analysis
Central Research Institute of Electric Power Supply Industry, Comprehensive

References

Report T40, “Lightning Protection Design for Power Stations, Substations, andUnderground Transmission Lines”
A lightning surge overvoltage is one of the important factors in the insulation
design of transmission lines and power generation substation equipment.
In this example, we calculate the surge propagation pattern for the case of alightning strike occurring on the first steel tower of a 500-kV air-insulated
substation (the first tower when viewed from the substation) and the lightning
current entering the power line from a reverse flashover, until it reaches the maintransformer in the substation. This corresponds to the first case (circuit symbol A)
of Table III-3-1 (p.84) of the lightning protection design guide mentioned in thereference column above. The modeling methods for all equipment in the substation,

Outlinethe transmission lines, and the system voltage are as recommended by the lightning
protection design guide (however, for the transmission line, a fixed constant line
model is used instead of a frequency-dependent line model).
For XTAP Ver. 1.11, all models described in the lighting protection design guide areprovided as components (which can be downloaded from the support site) for eachvoltage class, and in Ver. 1.2 and later, they are expected to become standard
equipment. By combining these components, it is possible to easily create datacases for a lightning surge analysis. In this analysis example, although only circuit
symbol A of Table III–3–1 (p. 84) of the lightning protection design guide is
described and provided as an example file SLS-01A.xsf, circuit symbols B throughV are also provided as example files SLS-01B.xsf through SLS-01V.xsf.

205

Analysis Circuit and Conditions
In this example, we calculated the surge propagation for a case of lightning striking the first steel tower of a
500-kV air-insulated substation, and the lightning current entering the power line from a reverse flashoverreaching the main transformer in the substation.
Specifically, the first case (circuit symbol A) in Table III-3-1 (p. 84) of the lightning protection design
guide mentioned in the reference column above is described as an example.
The first steel tower is the first tower when seen from the substation. Because the overvoltage in thesubstation becomes the most severe when a lightning strike occurs on this tower, for a typical analysis, alightning strike of the first steel tower is assumed. Figure 1 shows the conceptual diagram of the lightning
surge calculation in the air-insulated substation. From here, we describe the phenomenon until the lightning
current reaches the main transformer, as shown in the figure.

Lightning strike

Flashover

Penetrating lightning
surge
Lightning arrestor

Circuit breaker
Transformer

Figure 1 Conceptual diagram of lightning surge calculation in air insulated substation[Reverse Flashover Phenomenon]
As shown in Figure 1, if a lightning strike occurs on a steel tower, the potential of the body of the tower
increases owing to the impedance of the steel tower itself (called the surge impedance of the tower) and theground impedance of the tower-foot, and becomes higher than the potential of the power line. When thepotential difference between the upper phase arm and the power line exceeds the flashover voltage of thearcing horn, a flashover occurs in the upper phase, and the lightning current penetrates the power line.
Because a flashover occurs owing to the arm potential becoming higher than that of the power line, this iscalled a “reverse” flashover. In addition, because the flashover of an arcing horn is a phenomenon in whicha discharge path is formed owing to a leader progressing between the horns, it occurs within a short time of
microseconds or less. However, the horns will be gradually short-circuited within a finite time. It should be
noted that, owing to the branching of the current into the overhead ground wire, the current entering thesteel tower decreases, thus suppressing the potential increase of the steel tower. At the same time, the206lightning current branching toward the overhead ground wire increases the potential of the power line
through electromagnetic induction, further suppressing the voltage between horns.
To accurately simulate the potential increase of a steel tower described above, a steel tower is simulatedusing a four-stage model created based on a comparison with the measurement results. The groundresistance is set to a severe condition of 10 which is the upper limit of the control value. The
transmission line, including the overhead ground wire, is simulated using a distributed constant line model,
and four spans between the first and fifth towers are simulated. For locations away from the fifth tower, acondition in which the line continues infinitely is simulated through matching with a polyphase resistanceequivalent to the characteristic impedance (surge impedance) of the transmission line. Although the
influence of steel towers far from the fifth tower is ignored, because the traveling waves propagated beyond
the fifth steel tower never return through a reflection within the observation time, even if ignored, noinfluence on the analysis results will occur. Further, the system voltage is simulated by connecting a 500kV three-phase voltage source through a polyphase resistor for matching. Because the system voltage
influences the voltage between the upper-phase horns, it also influences the flashover. The flashoverphenomenon is simulated through a smooth short-circuiting of the horns by connecting a leader model
consisting of a combination of an inductance and a switch between the horns.

[Surge Propagation at the Substation]
The lightning current entering the power line owing to a reverse flashover becomes a traveling wave
(surge) and propagates to the substation through the service line, as shown in Figure 1. In the substation,
because lightning arrestors, circuit breakers, potential transformers, current transformers, and bushings areconnected to the busbar, the traveling wave propagating through the busbar is influenced by the impedanceof these devices, eventually reaching the main substation. Progressive wave components higher than thelimit voltage are negatively reflected at the arrestors and positively reflected at the terminal end of thebusbar. Further, in the potential transformers, current transformers, and bushings, only a steep portion of the
waveform is negatively reflected owing to the capacitance. Thus, a complex reciprocal reflection
phenomenon occurs on the busbar. It should be noted that, a portion of the lightning current branches
toward the overhead ground wire at the top of the first steel tower, which travels up to the anchoring steel atthe entrance of the substation and becomes negatively reflected.
The busbar is simulated using a single-phase distributed constant line, and the lightning arrestor issimulated using a nonlinear resistance having its voltage–current characteristics (v–i characteristics). The
electrostatic capacitances of the potential transformers, current transformers, bushings, and maintransformer are simulated along with the capacitance. Although a transformer is often considered toexperience inductance, when conducting an analysis on the order of microseconds, such as a lightning
surge analysis, the impedance of inductance L becomes an extremely large value, which can be ignored,
and the inductance of the capacitance 1/C between the windings and iron core, and between the windings,
becomes relatively small and thus dominant. This capacitance is called the penetration capacity of thetransformer. The circuit breaker is simulated as a short distributed constant line if it is in a closed state and207simulated with a capacitance in an open state. The anchoring steel is also simulated as a distributed
constant line.

[Modeling Details]
For the modeling details, refer to the lightning protection design guide (however, for the transmission line,
a fixed parameter line model is used instead of a frequency-dependent line model)[Analysis Conditions]
The analysis conditions are as follows.
·

Calculation time step

1.0 ns·Calculation end time

10.0 s·

Display start time

0.0 s·

Display end time

10.0 s[Example of XTAP Input]
An instance of creating this example using XTAP is shown in Figure 2.

Analysis Results
The results of executing this example using XTAP are shown in Figures 3 and 4.

Remarks
In this analysis example, although only the case of the circuit symbol A of Table III–3–1 (p. 84) of thelightning protection design guide mentioned in the reference column is explained and provided as theexample file SLS-01A.xsf, note that the cases of circuit symbols B through V are also provided as example
files SLS-01B.xsf through SLS-01V.xsf.

END

208

500 kVTransmission line
Constant parameters

Matching Circuit
AC superimposed

Fifth steel tower

Third steel tower

Fourth steel tower

Second steel tower

Lightning modelTransmission line 4 spans
Matching circuit &AC withsuperimposition

Constant parameters

Arrestor
10% decrease

Leader model

Arrestor
10% decrease

Transformer
substation
entranceBushingBushing

Leader method constants (3 folds,
AC superposed)
Bushing

Constant parameters

Center ofbusbarBushing

Arrestor
10% decrease

Bushing

Transformer end

Transformer

Figure 2 Example of XTAP input

209

Branch current [kA]

Time [us]Node voltage [kV](a) Lightning current waveform

Time [us]Potential difference between nodes [kV](b) Tower-top potential rise waveform

Time [us](c) Waveform of voltage between horns

Figure 3 Results of analysis using XTAP (part 1)

210

Node voltage [kV]Time [us]Node voltage [kV](a) Busbar voltage waveform (substation entrance)

Time [us]Node voltage [kV](b) Busbar voltage waveform (center of the bus)

Time [us](c) Busbar voltage waveform (transformer end)

Figure 4 Results of analysis using XTAP (part 2)

211Revision History
Date

Example File

Content changes

Version

Revision for XTAP Ver. 2.00
2014/11/19

2.0

Recalculation of line constants according to the changes in
XTLC

2012/07/19

1.2

Revision for XTAP Version 1.20

2011/12/15

1.1

2011/10/25

1.1

Revision for XTAP Version 1.11

2010/07/16

1.0

Creation of the first edition (for XTAP Version 1.10)

Revision of explanation (addition of cases for circuit symbols
B through V)

212

XTAP – Example collectionNo.

SSW-01

ExampleSwitching Surge Calculation of a 500-Kv Vertically Arranged Double-

Name

Circuit Transmission Line

Field

Switching surge calculation
Because this example was created specifically for this induction course,
there are no reference documents.

References* For general information on the switching surge calculation, pleaserefer to the following.
Central Research Institute of Electric Power Industry, Comprehensive
Report 121, “Study on Switching Over-voltages in Power Systems”
A switching surge overvoltage is one of the important factors indetermining the insulation level of the transmission lines and power
station/substation equipment.
In this example, an input surge calculation is conducted on a 500-kv
vertically arranged double-circuit transmission line, which is commonlyused in Japan.
Assuming a 500-kV transmission line connecting an upper substation
and a lower substation, the waveform of the overvoltage generated at
the start and end of the line is calculated when the circuit breaker of theOutlinefirst line (1L) is closed at the upper substation. In the actual 500-kV
system, to reduce the input surge overvoltage, the resistance input
method is adopted (first, energization is applied with a resistance of afew 100 and the resistance is then short-circuited after
approximately 10 ms to complete the input). However, in this example,
we tried setting the input with no resistance allowing the travellingwave phenomenon to be easily understood. In addition, a frequencydependent line model that takes into consideration the skin effect of theelectric line and ground, as well as their frequency characteristics, is
used.

213

Analysis Circuit and Conditions
Figure 1 shows a skeleton diagram of the system in which an input surge calculation isconducted, and Figure 2 shows the assembly of a 500-kV transmission line.
Upper substation

SystemvoltageLower
substation

Closed at
t = 5 ms

Rear impedance

1L
500-kV transmission line, 200 km

50 Hz

2L

Figure 1 Skeleton diagram of a grid in which an input surge calculation is conducted14.5 mac'b11.2 m b'

GW
IACSR
150

PW
c

79.5 m

11.6 m57.5 m72.0 m

a'

TACSR
410
6

12.0 m43.0 m50 cm

e = 300 m

Line length of 200 km
Earth resistivity of 300

Figure 2 Assembly of 500-kV vertically arranged double-circuit transmission line

Further, the assembly in Figure 2 is same as that of example TL-02.

[Analysis Phenomenon]
When closing the 1L circuit breaker at the upper substation, a voltage is applied to thestarting-edge of the transmission line 1L. If the closing timing is close to 90º of the system214

voltage (sine wave) in any one of the three phases, a large voltage will be applied to thatphase. In this example, assuming a system frequency of 50 Hz, an initial part of phase-a ofthe system voltage of 0º, and a circuit breaker closing time t of 5 ms, phase-a of 1L is set tothe peak value, which is the most severe input condition. Because it is difficult to intuitively
understand the travelling wave phenomenon on a constant multi-phase distributed line, we
tentatively consider only a single-phase distributed constant line, namely, phase-a of 1L. The
voltage applied at phase-a of 1L propagates along the transmission line as a travelling wave
and reaches the remote end after approximately 670 s (as estimated from the line lengthwith the propagation speed being the speed of light). Because the remote end has an open
condition (impedance of ), the traveling wave voltage that reaches the end is specularlyreflected and generates a double overvoltage. This overvoltage is generated between the interelectrode gap of the circuit breaker in the lower substation 1L, which is under an open
condition, and between the insulators of the bus bar and the transmission line within thevicinity, which becomes a problem in terms of the insulation design. The travelling wave isthen reflected back and forth along the transmission line and displays an oscillatory
waveform determined by the line length. In lines other than phase-a of 1L, an induced voltage
is generated, which is also reciprocally reflected.
Considering the above description, at the time of normal operation, although an overvoltageof twice the ground voltage peak value will be generated, in practical terms, the overvoltage
value varies depending on the skin effect of the electric wire and ground, its frequency
characteristics, and the induction phenomena with other phases and other lines. Moreover,when another transmission line is connected to the busbar of the upper or lower substation,
the reciprocating reflection phenomenon occurring in these lines is also superimposed,
resulting in a more complex overvoltage waveform. Because it is almost impossible to obtainsuch an overvoltage waveform through manual calculations, a program such as XTAP isrequired. During the actual input surge analysis, taking into consideration the randomness ofthe circuit breaker input timing, a Monte Carlo calculation (a statistical calculation) is
conducted. Although the current XTAP does not have a function to automate this process,such a function is expected to be added in the near future.
[Line Model Preparation]
A frequency-dependent line model that takes into consideration the skin effect of the electric
wire and ground, as well as the frequency characteristics, is used. Because it is identical to215

the line used in example TL-02, please refer to that example.
[Circuit Elements other than the Line Model]
A “SIN wave voltage generator” is used for the system voltage simulation. By calculating theinductance equivalent as 5%, taking 50 Hz, 500 kV, and 1000 MVA as the base, theimpedance behind the upper system is set to 39.8 mH. For the circuit breaker, a “switch witha circuit breaker logic” is used, and its input is set at time t = 5 ms.

[Analysis conditions]
The analysis conditions are as follows:ꞏ

Calculation time step

10.0 sꞏ

Calculation end time

20.0 msꞏ

Display start time

0.0 msꞏ

Display end time

20.0 ms[XTAP Input Example]
Figure 3 shows an example of creating this example using XTAP.
Analysis Results
The results of this example executed using XTAP are shown in Figures 4 and 5.

END

216

Example Name: Switching surge calculation of a 500-kv vertically arranged double-circuit transmission line,
no. SWS-01

Upper substation

System voltage andrear impedance

Circuit breaker
to be closed
Circuitbreaker

Lower substation
Circuitbreaker

Power transmission line

Circuitbreaker

Closed at
t = 5ms

Frequency dependence

5% with 1000 MVA
as the base

Analysis conditions

Line constants are calculated using XTLC under the following conditions* Frequency-dependent line modelCalculation time step:10.0s* Frequency samples 0.1 Hz to 10 MHz(400 points)

Calculation end time:20.0ms* For the pillar assembly, refer to the example table

Display start time:0.0ms
Display end time:20.0ms

Figure 3 Example of XTAP input

217

500
400

Branchvoltage[kV]
ﾌﾞﾗﾝﾁ
電圧[kV]

300
200
100
0
-100
-200
-300
-400
-500

0

0.01
Time
[s]

0.02

時 間 [s ]v(Ea) #1

v(Eb) #1

v(Ec) #1(a) System voltage waveform
1000
800

Nodeﾉｰﾄﾞ電圧
voltage[kV]
[kV]

600
400
200
0
-2 0 0-4 0 0-6 0 0-8 0 0-1 0 0 00

0 .0 10 .0 2Time
時 間[s][s ]u (Se n d _ 1 a ) # 1

u (Se n d _ 1 b ) # 1

u (Se n d _ 1 c) # 1(b) 1L Starting-edge voltage waveform

1000
800

Node ﾉｰﾄﾞ電圧
voltage [kV]
[kV]

600
400
200
0
-200
-400
-600
-800
-1000

0

0.01

0.02

Time
時 間[s][s ]u(R ec_1a) #1

u(R ec_1b) #1

u(R ec_1c) #1(c) 1L Remote-end voltage waveform
Figure 4 XTAP analysis results (1L voltage waveform)

218

500
400

Branchvoltage[kV]
ﾌﾞﾗﾝﾁ
電圧[kV]

300
200
100
0
-1 0 0-2 0 0-3 0 0-4 0 0-5 0 00

0 .01

0 .0 2Time
時 間[s][s ]v(Ea ) #1

v(Eb) # 1

v(Ec) # 1(a) System voltage waveform
2 00
1 50
1 00

ﾉｰﾄﾞ電圧[kV]

Nodevoltage
[kV]

50
0
-50
-1 00
-1 50
-2 00

0

0 .01

0.0 2

Time
[s][s ]時間

u (Se n d _2 a ) #1

u (Se n d_ 2 b ) # 1

u (Se nd _ 2 c) # 1(b) 2L Starting-edge voltage waveform
2 00

Node
voltage
ﾉｰﾄﾞ電圧[kV] [kV]

1 50
1 00
50
0
-50
-1 00
-1 50
-2 00

0

0 .01

0 .0 2Time[s][s ]時間u(R e c_2 a ) #1

u(R e c_2 b ) #1

u(R e c_ 2 c) #1(c) 2L Remote-end voltage waveform
Figure 5 XTAP analysis results (2L voltage waveform)

219

Revision History
Date

Example fileContent changes

VersionUpdate for XTAP Version 2.00
2014/11/19

2.0

Recalculation of line constants owing to changes in
XTLC

2012/07/19

1.2

Update for XTAP Version 1.20
Update for XTAP Version 1.11

2011/10/18

1.1

The line constant file was modified because thenumber of valid digits of XTLC was modified.

2010/07/16

1.0

Initial version created (for XTAP Version 1.10)

220

XTAP Example Collection
Example
Name
Field

No.

SSW-02

Switching surge Analysis of a 275-kV System

Switching surge calculation, switching surge calculation, overvoltage analysis“Review of Simulation Methods for Electromagnetic Transients in Power Systems

References

and Simulation Cases Using XTAP (Part 1) Switching Transient Overvoltage
Simulations,” Survey Report H12005, Central Research Institute of Electric Power
Industry
After elucidating the concept of a switching surge, the principle of occurrence, andOutlinethe simulation method of each element of the power transmission system, we
introduce an example switching surge analysis of a 275-kV system using XTAP
(excerpted from the above reference).

1 Switching surge
1.1 Concept [5]When newly starting the operation of a power transmission line or restarting the operation after construction,
voltage is applied to the power line by keeping the transmission line without a load and closing the circuit
breakers on the power supply side.
Further, when ground faults occur in the transmission lines from lightning strikes or similar events, the circuit
breakers on both sides of the line are temporarily opened to remove the continuous current arc owing to a ground
fault, and the circuit breakers are closed again to resume the power transmission (called reclosing). Thus, when
applying voltage to a power transmission line, it is often kept without a load. When applying voltage to a no-loadtransmission line, an overvoltage called a switching surge occurs in the line. Therefore, the insulation of thetransmission line must be designed to withstand the overvoltage owing to this switching surge.
The magnitude of the overvoltage owing to a switching surge depends significantly on the supply voltage at
the moment when the circuit breaker is closed. Because a circuit breaker is a device that applies the opening andclosing of the electrodes using a mechanical mechanism, its operation is slower compared to the commercial
frequency and the variation is also large. That is, because the phase of the supply voltage at the moment thecircuit breaker is closed cannot be predicted, it is reasonable to think that it is determined (randomly) based on
probability. Therefore, in the study of the actual switching surge, the switching surge calculations are conducted
a number of times by randomly changing the closing timing of the circuit breaker, and the result is displayed in221the form of a cumulative frequency distribution ofway, it is possible to evaluate the probability ofoccurrence of a certain overvoltage. A method for
obtaining the statistical results based on many
simulations conducted when randomly changing theparameters described above is called the Monte
Carlo method. Although the maximum value of thegenerated voltage tends to increase as the number ofsimulations increases, the value at which the

Cumulative frequency distribution [%]the generated voltage, as shown in Figure 1. In thisgeneration probability reaches 2% tends to be aconstant regardless of how frequently the number ofsimulations is greater than or equal to a certain value.
Hence, this value is often used as the representative
figure. The value at which the generation probability
Generated voltage [pu]becomes 2% is called the “2% value.” Thecumulative frequency distribution of the generated

Figure 1 Cumulative frequency distribution ofvoltage is often plotted on a type of graph called a

voltages generated by switching surge

normal probability plot, as shown in Figure 1. In thistype of plot, the vertical axis is scaled such that theplotted points become a straight line in the case of anormal distribution.
An overvoltage from a switching surge is often represented as a multiple of the nominal voltage when the groundwave peak value is 1, which is called the overvoltage multiple. It is important to note that, although pu is used as
the unit, it is unrelated to the pu used in a system analysis (where the effective value of line voltage is set to 1).
Strictly speaking, the term “switching surge” must be used as a generic name for not only for a switching surge
but also for a surge generated by the opening and closing of a switch, such as a cut-off surge or capacitor
switching surge. However, the term “switching surge” is commonly used in the sense of a switching surge.
Hence, the above-mentioned overvoltage multiple is also called as a switching surge multiple.

1.2 Principle of Switching Surge Generation
An overvoltage from a switching surge can be broadly classified into two types: one occurring as an oscillation
phenomenon of the inductance of the power supply side and the capacitance of the transmission side, and theother caused by a specular reflection of the progressive waves propagating on the transmission line at the open
far end. The former occurs when the line can be treated as a lumped constant capacitance because the inductanceof the line is small whereas the capacitance is large, such as with a cable line. In contrast, the latter occurs when
the distribution constant characteristics of the line become remarkable, such as in a long-distance overhead222

power transmission line. From now, we will explainL

t=0
Transmission
line

these circuit phenomena.
First, we will describe the switching surge
phenomenon

occurring

as

the

S

oscillation

e(t)

C

v

phenomenon of the power source side inductanceand transmission line capacitance [1]. We assume
that, in a power station or a substation, a no-loadtransmission line is energized by closing the circuit
breakers. The power source on the upper side of the

Figure 2 Equivalent circuit for energization of acircuit breaker is simulated using a sine wave

no-load transmission line

voltage source e(t) = E sin(t + ) and an inductance(lumped constant capacitance approximation)Lconnected

in-series.

Assuming

that

the

transmission line can be regarded as the capacitanceof the lumped constants, it is simulated with
illustrated in Figure 2. When the circuit breaker S isv / E [p.u.]

capacitance C. The equivalent circuit of the above isclosed at t = 0 under the conditions of the initial
phase of the power supply  at 90° and the residual
voltage of the transmission line at zero, the voltage v

equation.

E
1 

t
 cos t  cos
2
1   LC 
LC 




of the transmission line is given by the followingv(t ) 



(1)




time [ms]Figure 3 Voltage waveform when energizing a
no-load transmission line

In the above equation, the first term corresponds to a(lumped constant capacitance approximation)

steady supply voltage, and the second term
corresponds to the oscillation owing to L and C. It
should be noted that, in the above calculations,because the losses in the power supply side and the transmission line are ignored, the second term does notattenuate. In reality, however, it attenuates because of the losses. As an example, when calculated according toequation (1) for the case of  = 2 60, L = 15 mH, and C = 2 F (for an underground cable of a few kilometers),
Figure 3 is obtained as the waveform of the transmission line voltage v (voltage value normalized using E). It
can be observed that the voltage owing to an LC oscillation is superimposed on the supply voltage. Usually,
because the value of 2LC in equation (1) is much smaller than 1, it is understood that the peak value of v
reaches approximately 2E (that is, 2 pu).
In equation (1), the amplitude of the term owing to the LC oscillation of the second term is approximately equal
to the voltage applied before closing the circuit breaker S. This holds well even when there is a residual voltage
in the transmission line. For instance, assuming that the circuit breaker was opened at the moment when the223power supply reached a negative peak value, namely, E, if the circuit breaker is closed at  = 90° from the statein which the residual voltage of V0 = E exists in the transmission line, as the voltage waveform v of thetransmission line starts from E and oscillates at an amplitude of approximately 2E, it takes a peak value ofapproximately 3E (3 pu).
Next, we describe the switching surge phenomenon occurring from the specular reflection of progressive
waves propagating along the transmission line at the open far end. In the same manner as above, the no-loadtransmission line is energized by closing the circuit breakers in the power station or substation. The power sourceon the upper side of the circuit breaker is simulated using a sine wave voltage source e(t) = E sin(t + ) and an
inductance L connected in-series. The transmission line is simulated using a no-loss distributed constant circuitwith a line length l, propagation velocity c, and characteristic impedance Z0. The equivalent circuit is shown in
Figure 4. We assume that the circuit breaker S is closed at t = 0 under the conditions of the initial phase of thepower supply = 90°, and the residual voltage of the transmission line is zero. In this case, because the voltageof the power supply increases in a step form, it is considered a step wave with amplitude E immediately after t =0. Because the waveform of voltage v1 at the starting edge of the transmission line is obtained by dividing thevoltage of the step wave by L and Z0, it becomes a stepped waveform with an amplitude E vanishing at a time
constant of L / Z0. This waveform v1 propagates along the transmission line as a traveling wave and reaches the
remote end after a propagation time of  = l / c. According to the distributed constants circuit theory[6], because a
voltage wave is specularly reflected at the open end producing twice the voltage, a voltage of 2E (2 pu) is
generated at the remote end of the transmission line. Further, after time , the specular reflected traveling wave
reaches the starting edge. Because the impedance of the power supply side is only L, which is small, the voltage
wave is almost negatively reflected at the starting edge, and thereafter this reciprocating reflection repeats.
Although the supply voltage was considered a step wave immediately after t = 0, because it changes to asinusoidal wave when viewed over the long term, the voltage actually generated on the transmission line is thesinusoidal supply voltage superimposed through the above-mentioned reciprocating reflection. As an example,
assuming an overhead power transmission line equivalent to 100 km, the waveform of the remote-end voltage v2
is calculated using  = 2 60, L = 50 mH, l = 100 km, c = 300 m/s, and Z0 = 400 when applying theBergeron method [2], the results of which are shown in Figure 5 (the voltage values are normalized by E). The
reciprocating reflection resulting in a peak voltage of approximately 2 pu at the remote end is superimposed overthe supply voltage. In the above calculation, because the losses in the power supply side and the transmission
line are ignored, the reciprocating reflection continues. However, considering the losses, the reciprocatingreflection will gradually attenuate. Therefore, the peak voltages occurring near 9 and 17 ms are actually muchsmaller, and the first peak near 1 ms is often the maximum value. The spike-like voltage waveform of the portion
surrounded by the dotted line in Figure 5 is generated owing to a sudden change in the current flowing throughthe inductance L when the traveling wave returning from the remote end reaches the starting edge. It should be
noted that, because the transmission line is treated herein as a distributed constant circuit, in the initial phase of
the reciprocating reflection of the traveling wave, oscillations with a period of 4 were observed owing to thereciprocating reflection phenomenon. In contrast, as the frequency decreases, the behavior of the open-ended
distributed constant circuit approaches the capacitance. In other words, when the high-frequency components are224lost by repeated reciprocating reflections, thet=0LTransmission
line

oscillation period approaches a value determined byS

the inductance and capacitance of the transmission

v1

e(t)

line added to the inductance of the power source sideLine length l, propagation speed c
Characteristic impedance Z0v2(for details of this phenomenon, refer to Section
2.3.2 of [5]).
Because the concept of a residual voltage is the

Figure 4 Equivalent circuit for energization of asame as that of the switching surge, which occurs as

no-load transmission line

the oscillation phenomenon of the inductance of the(simulation using a no-loss distributed constant

power supply side and transmission line capacitancecircuit)described above, it is omitted herein.

1.3 Residual Voltage and Input Conditions


overvoltage from a switching surge varies dependingv2 / E [p.u.]

As mentioned in the previous section, an

on the magnitude of the residual voltage. Therefore,
in an actual analysis, to reproduce a realistic residual

voltage, it is necessary to set the input conditions

according to the operation of the actual equipment.



The input conditions used in the analysis of [7] areshown below.

time [ms]Figure 5 Remote-end voltage waveform when

(1) Three-phase input – Under this analysis

energizing a no-load transmission line

condition, the input is simulated in a state where(simulation using a no-loss distributed constant

there is no residual voltage in the transmission

circuit)line, such as when newly starting the operationof a power transmission line or restarting theoperation of the power transmission line after
construction. Using a circuit breaker of the load-

end under an open condition, the three phases of the circuit breaker on the supply side are closed
concurrently. When parallel lines are in operation, the induced voltages from these lines will besuperimposed.
(2) Three-phase reclosing – Under this analysis condition, we simulate the cutoff and reclosing of all threephases of the circuit breaker together in the transmission lines where accidents have not occurred. Aftercutting off all three phases at both ends of the transmission line, using the circuit breaker at the load-end in
an open state, the three phases of the circuit breaker at the power supply side are all closed concurrently.
Normally, because a large residual voltage remains during all three phases, it gives the most severe225overvoltage. Because this analysis condition includes the cutoff and reclosing of the transmission lines
where accidents have not occurred, such a voltage does not exist during an actual switching operation.
However, it is used to obtain the most severe overvoltage as a reference value.
(3) Selective reclosing – This analysis condition is for ultra-high voltage transmission lines where both ends ofthe transmission line are cut off only for the phase that caused the accident. The circuit breaker at the load
end is kept as is, whereas the circuit breaker at the supply side is switched on.

Furthermore, when analyzing a specific case, the analysis is conducted by setting the residual voltage assumed
for the case without being limited to the above three conditions.

1.4 Simulation of the Power Supply Side

Both ends of power transmission lines are usually connected to a power generation station, a substation, or aswitching station. Among the two ends of the power transmission line, the circuit breaker at the power supply
side is closed and a switching surge is generated. When considering an exclusion of the lightning arrestors andtransmission lines of other routes, which are described later, if the power supply side is a power plant, there will
be a power generator and a step-up transformer behind the circuit breaker. If the power supply side is a
substation, there will be a transformer behind the circuit breaker and transmission lines further spreading thesystem. If the power supply side is a switching station, there will be transmission lines directly behind the circuit
breaker from where the system spreads.
If the power supply side is a power station or a substation, the system behind the circuit breaker is often
simulated using an equivalent circuit, as shown in Figure 6. This equivalent circuit is a simplified circuit thatsimulates the voltage generated by the system behind the circuit breaker and its impedance. The three-phasevoltage sources, namely, ea(t), eb(t), and ec(t), simulate the sinusoidal voltage generated in the busbar of thepower source side with the circuit breaker to be closed under an open condition. Further, the impedances Z1 andZn simulate the impedance of the system behind the circuit breaker. In contrast, when the power source side is aswitching station, the transmission line behind the circuit breaker is simulated as a transmission line model using
the method described in section 1.5, and the power station before or the substation after is simulated using theequivalent circuit shown in Figure 6. Fo the tip of a transmission line or a switching station, we first simulate thetransmission line up to the power station or substation as the transmission line model. For a switching station, thereason for simulating the transmission line behind it as a transmission line model is that the transmission line at
the instant the circuit breaker is closed appears to have a characteristic impedance corresponding to a resistanceof several hundreds ohms, which may result in an error in the equivalent circuit shown in Figure 6.

226

Although the amplitude of the three-phase voltage

Z1source can be easily obtained from the operating
value, the resistance and impedance of the

ea(t)

R1

L1

eb(t)

R1

L1

ec(t)

R1

L1

Zn

impedances Z1, and Zn will become a problem. In theRn

Lnguidelines[8] compiled by the IEEE Working Groupon a switching surge analysis, Zn is omitted (shortcircuited), and the resistance and impedance of Z1
are determined from the short-circuit capacity of thesystem behind the circuit breaker. In other words,

Fig. 6Equivalent circuit of the power supply

the short-circuit impedance is obtained from thesidepositive-phase equivalent circuit of the systembehind the circuit breaker. Taking its real part as theresistance and the value obtained by dividing theZ1

imaginary part by namely, 2f, where f is theea(t)commercial frequency) as the inductance, we obtain

Zn

Z1. However, in the guidelines [9] compiled by theeb(t)CIGRE Working Group, the impedance of the rearsystem is obtained by adding the impedances of theec(t)generator, transformer, and transmission lines thatare present behind the circuit breaker, taking into
consideration the zero-phase impedance. Because

Figure 7 Equivalent circuit of the power supply

the equivalent circuit of the generator is complex, itside considering its frequency dependence

is simulated by approximating the d-axis subtransient reactance Xd". In addition, because thepositive-phasecomponent

impedance

obtainedthrough an integration corresponds directly to Z1, theresistance and inductance can be obtained from the real and imaginary parts. Taking the zero-phase impedance
obtained through an integration as Z0, the value of Zn is obtained from1
Z n  ( Z 0  Z1 )3

(2)

The resistance and inductance values are obtained from its real and imaginary parts. Further, in the CIGRE
guidelines, the frequency-dependent effect, that is, the effect of the increase in the resistance component of thegenerator and transformer, as the frequency increases, is taken into consideration, and the typical frequency
characteristics of the generator and transformer are indicated. Because transformers, transmission lines, andgenerators are present behind the circuit breaker, it is not easy to obtain their comprehensive frequency
characteristics. Moreover, it is also not easy to obtain the frequency characteristics of each generator ortransformer. However, if the frequency characteristics of the system behind the circuit breaker are roughly227known, the frequency-dependent effect can be approximated [3], [10] using an equivalent circuit similar to thatshown in Figure 7.
In equation (2) above, the imaginary part of Z0  Z1 may become negative in certain cases, and a simulation
using the inductance may become impossible. To handle this problem, a simulation method using an idealYtransformer for coupling is introduced in [9]. If the neutral point of the transformer behind the circuit
breaker is directly grounded, it becomes Z0  Z1, and Zn can be omitted (short-circuited).
Reference [5] compares the calculation results of the case in which the value of inductance L is only the leakage
inductance of the transformer just behind the circuit breaker, and the case in which it is the inductance obtained
from the short-circuit capacity of the system behind the circuit breaker, with the measured results. An actual
measurement was carried out at the Shikoku Electric Power Nagagawa trunk line and Kansai Electric PowerOkutataragi trunk line, and in both cases, it was concluded that the calculation results closer to the actual
measurements can be obtained if L is simulated using only the leakage inductance of the transformer just behindthe circuit breaker. Although this conclusion may not be necessarily consistent with the IEEE or CIGRE
guidelines, as an important research result, we introduce it here. From this, conducting a calculation by changing
the inductance as a parameter is considered realistic in an actual analysis.
When conducting a detailed analysis, instead of using the equivalent circuits of Figures 6 and 7, the systembehind the circuit breaker may also be assembled and analyzed.

1.5 Simulation of Transmission Line

In a switching surge calculation, it is necessary to accurately reproduce the phenomena from the peak voltage
observed immediately after closing the circuit breaker until it settles at the supply voltage of the commercial
frequency after the attenuation of the reciprocating reflection and LC oscillation. In other words, because it isnecessary to reproduce within the line model characteristics of a wide frequency range, from the commercial
frequency to a high frequency corresponding to the peak voltage, it is standard to use a frequency-dependent line
model. However, when a switching surge occurs as an oscillation phenomenon of the power supply sideinductance and transmission line capacitance described in the first half of Section 1.2, such as when the line to beenergized is a comparatively short underground cable, a sufficient level of accuracy can be obtained using a
constant parameter line model. Because the capacitance of the transmission line does not have a frequencydependent effect, in such a case, the frequency used to calculate the line constants does not significantly affectthe calculation results.
Because two or more lines assembled in the same tower will have a relatively large mutual induction between
them, they are simulated as a single transmission line model. When considering the mutual induction between
two or more transmission lines running parallel within the neighborhood, all lines are integrated into a single
transmission line model. In a switching surge analysis, to simulate the line from the start to end as a single line
model, the line constants are usually calculated when assuming an average pillar assembly over the entire span.
Further, when the presence of the overhead ground wire is considered, it is eliminated from the matrix of the line
constants. In XTAP, frequency-dependent line models with span lengths of 25, 50, 75, and 100 km are prepared228

for each standard pillar assembly of each voltage class introduced in [4]. When the pillar assembly and span
length differ from these, the analyst must create a frequency-dependent model using the line constant calculation
routine, XTLC, included in XTAP. At that time, it is recommended to set the samples of the line constantcalculation frequency to 400 points within the range of 0.1 Hz to 10 MHz.

1.6 Simulation of Lightning Arrestor

In recent years, zinc-oxide type lightning arrestors have been used not only to suppress a lightning surge
overvoltage, but also a switching surge overvoltage [7], [11]. In addition to the non-linear voltage-current
characteristics, which are the basic characteristics of a zinc-oxide type lightning arrestor, the limiting voltage
slightly increases for sharp waves within an order ofs. However, even in a lightning surge analysis targeting asteeper phenomenon as compared to a switching surge analysis, because the influence of the sharp wave
characteristics is insignificant in the analysis of a system of 500 kV or less, it is ignored. That is, the zinc-oxidetype lightning arrestor is simulated using only a non-linear resistance that reproduces the non-linear voltage–current characteristics [4]. In the lightning surge analysis of a 1,000 kV system, there are examples in which the
sharp wave characteristics are taken into consideration [12], [13].
As the target phenomenon of the analysis is gentle in the case of a switching surge analysis as compared to alightning surge analysis, it is natural to ignore the sharp-wave characteristics of a zinc-oxide type lightning
arrestor and simulate it using only the nonlinear resistance. The representative voltage–current characteristics ofa zinc-oxide type lightning arrestor are summarized in [4], and their models are available in XTAP. Grasping theprocessing energy of the arrestor that follows the switching surge is also an important item of an analysis.
Because the available voltage–current characteristics may be the upper limit of the actual arrestor characteristics,
it is necessary to take this into consideration when analyzing the processing energy.

1.7 Simulation of Transmission Line on Another RouteIn an analysis of a switching surge overvoltage, because the circuit breaker at the far end of the transmission
line to be analyzed is open, there is no need to simulate the remote system. If there is a zinc-oxide type arrestor
at the service line entrance of the substation at the far-end side, it may be simulated. Meanwhile, in the startingedge of the transmission line, it is necessary to simulate the following equipment. First, the system present in the
power supply side of the circuit breaker to be closed, in other words, the system behind the circuit breaker, issimulated using the equivalent circuit shown in Figure 6 or 7, as described in Section 1.4. If there is a zinc-oxidetype arrestor in the line side of the substation at the starting-edge, it must be simulated. In addition, if a powertransmission line of another route is connected to the busbar on the line side of the substation (which ishereinafter simply referred to as the “transmission line of another route”), it must be simulated.

229

The simulation of the transmission line of another route is not easy. Because the analysis is carried out under
the assumption that the transmission line of another route is normally operated, the problem is to determine how
far the system extending further from its far end must be simulated. Although an example that deeply examines
this problem has not been found, a simulation is usually conducted through the following procedure. First,
following the transmission line of another route and examining the type and length of the transmission line, thesubstation whose influence on the analysis results is large is determined, and the simulation is conducted indetail up to that substation. The transmission lines falling under the scope of this detailed simulation are
simulated as they are as transmission line models using the procedure described in Section 1.5. The transformers
in the substations along the way are simulated, including the secondary side, using the procedure described inexample SSW-03. Further, if there is a lightning arrestor installed at the substation, it needs to be simulated.
Systems far from the substation located at the end of the detailed simulation range are simulated using theappropriate equivalent circuits, while taking into consideration whether the system seen from there is a powersource or a load. For example, in the case of the power source, the equivalent circuit in Figure 6 or 7 is often
used. The values of the source voltage E and the impedances R and L in the equivalent circuit are obtained using
the method described in Section 1.4. In contrast, if there is a load far from the terminal substation, the equivalentcircuit shown in Figure 8 is used in the IEEE guidelines [8]. In this equivalent circuit, the impedance of thetransmission line and the transformer up to the actual load is expressed by Rs and Ls, the effective power of theload is expressed by the resistance RP, and the reactive power is expressed by the inductance LQ (if the value ofthe reactive power is negative, it is simulated using capacitance CQ). The simulation method stated here is just an
example, and in practice, the analyst makes a judgement individually and then conducts a simulation. Further,because it is not easy to specify the range of the system that must be simulated in detail, in reality, trial switchingsurge calculations are conducted to observe the influence. Hence, this is often determined through trial and error.

RP

Rs

Rs

Rs

Ls

Ls

Ls

LQ

CQ

RP

LQ

CQ

RP

LQ

Figure 8 Equivalent circuit of a load

230

CQ

1.8 Simulation of Circuit Breaker and Statistical Calculation

As described in Section 1.1, the magnitude of the switching surge overvoltage is statistically evaluated fromthe results of the multiple switching surge calculations conducted by randomly changing the closing timing of
the circuit breaker. Therefore, the circuit breaker is simulated as an ideal switch that can randomly change theclosing timing for each simulation.
First, we describe the simulation of a circuit breaker that does not have an input resistance. As shown in
Figure 9(a), a close command is first given to the circuit breaker. The timing at which this close command isissued is assumed to be evenly distributed between 0° and 360° of the phase of the power supply voltage, and in
the simulation, it is determined using a uniformly distributed random number. Next, taking the point in time after
a certain time T has elapsed from the close command as a reference, the time until the electrodes inside thecircuit breaker really connect is assumed to fluctuate according to the normal distribution with a standard
deviation of . This fluctuation is generated using a random number following a normal distribution. Taking theinherent variation of the simulation resulting from the random number generation as  the ideal switch thatsimulates the circuit breaker will close at the instantT +  has elapsed from the close command. In the①

above, T is the average value of the time required forthe electrodes inside the circuit breaker to connect①

after the close command is issued. Although a closeClose command

command is simultaneously given to the three-phase

Actual closing time

(0° to 360° uniform distribution

circuit breaker, because  is generated using random
numbers for each phase, the values in the three

Time

T


(fixed)phases differ.
A circuit breaker with an input resistance is(a) No input resistancesimulated as follows. Similar to the case of closing①

with no-resistance, a close command is given to thecircuit breaker, and its timing is determined using a②

random number uniformly distributed between 0°
and 360° of the phase of the supply voltage, as
shown in Figure 9(b). Next, taking the point in time②
Time ofresistance shortcircuit1

Close command

after a certain time T1 has elapsed after the close①
Time ofresistance input

2（0° to 360° uniform
distributioncommand as a reference, the time until the resistanceis actually turned on is assumed to fluctuate
according to the normal distribution with a standard

1

deviation of 1. Further, taking the time point after a

Time

T2
(Fixed)T1(Fixed)2(b) With input resistancecertain time T1 + T2 has elapsed after the close
command as a reference, the time until the resistance

Figure 9 Simulation of circuit breakers

is really short-circuited is assumed to fluctuate

231

according to the normal distribution with a standard deviation of 2. These two fluctuations are generated as
random numbers that follow their respective normal distributions. Taking the inherent variation of the simulation
obtained as a result of a random number generation as 1 and 2, the ideal switch for turning on the resistance isclosed when T1 + 1 has elapsed from the close command, and the ideal switch for short-circuiting the resistanceis closed when T1 + T2 + 2 has elapsed from the close command. Here, T1 is the average value of the timenecessary for turning on the resistance after the close command is issued, and T1 + T2 is the average value of thetime required for short-circuiting the resistance after the close command is issued. Although a close command issimultaneously given to the three-phase circuit breaker, because 1 and 2 are generated by random numbers foreach phase, the values in the three phases differ.
In [5], by conducting experiments on the closing timing of the circuit breakers, it has been concluded that thevariations follow a normal distribution, and there are no differences owing to the type or voltage class of thecircuit breaker. Further, the standard deviation of the variations was between 0.5 and 1.5 ms, and the average wasapproximately 1 ms. However, because a considerable amount of time has elapsed since the experiment was
conducted, whether the above findings can be applied to the latest circuit breaker is unknown.
Because XTAP is equipped with statistical calculation functions (functions applying a Monte Carlosimulation) that automatically conduct numerous switching surge calculations by providing random numbers to
the timing of the close command and the variation of the closing circuit breaker, the calculation of the
cumulative frequency distribution of the maximum value of the generated voltage is automated.

1.9 Example Analysis using XTAP

We introduce an example of calculating a switching surge overvoltage generated when power transmission line 1
is energized in a 275-kV system, as shown in Figure 10. The system on the power supply side of substations 1, 3,
and 4 was simulated using the equivalent circuit shown in Figure 6. In the calculation of the switching surge, allvoltage source components in the equivalent circuit were set to a 275-kV three-phase power supply of the same
phase, and thus the tidal current does not become a problem. If the conditions are set such that the tidal currents
do not occur, a Ferranti effect will occur, and the evaluation will become rigorous. For the impedance part of theequivalent circuit, we calculated the respective inductance values from the assumed short circuit capacities andadded realistic resistance values. These resistance and inductance values are shown in Figure 10. Assuming thatthe transmission lines 1, 2, and 3 are overhead lines with the standard assembly stated in [4] and the earth
resistivity as 100 m, we conducted the simulation using a frequency-dependent line model. The lengths oftransmission lines 1, 2, and 3 were 100, 50, and 70 km, respectively, and the samples of the line constantcalculation frequency at the time of the creation of the frequency-dependent line model were set to 400 points
within the range of 0.1 Hz to 10 MHz. The circuit breaker to be closed was the circuit breaker 1 at thesubstation-1 side of transmission line 1, and the circuit breaker was assumed to have no input resistance, as
shown in Figure 9(a). The average time from the close command to the time of closing T was set to 10 ms, andthe variation in the closing time  was set to 1.5 ms. Moreover, in this analysis, we assumed the condition inwhich there is no lightning arrestor. Figure 11 shows the data created by XTAP .

232

Using these, we conducted 300 simulations by taking the timing at which the close command is issued as arandom number uniformly distributed between 0° and 360° of the power supply phase. The calculation time step
was set to 0.5 s from the propagation time of the transmission line. Figure 12 shows the maximum value of thevoltage generated at the end of transmission line 1 (substation 2 side) in pu plotted on a normal probability plot
(taking the peak of the nominal voltage to ground as 1 pu). Here, the largest among the maximum values of thevoltage waveform of each phase was taken. From Fig. 12, the 2% value, which is the maximum value of thegenerated voltage corresponding to a cumulative frequency of 2%, is approximately 2.7 pu. Figure 13 shows aplot of the voltage waveforms of each phase corresponding to the 2% value. From the figure, it can be confirmedthat the maximum voltage is generated in the b-phase.

233Substation 1

2
T

0.40 ?
275 kV
50 Hz

Transmission line 1

×32.1 mH

Substation

×275 kV, 100 km

Short-circuit capacity

7500 MVA

×

Substation 3

Substation 4
Transmission line 3

Transmission line 2

275 kV, 50 km

275 kV, 70 km

267 mH 3.34 ?
Short-circuit capacity

900 MVA

2.01 ? 160 mH
275 kV Short-circuit capacity
50 Hz 1500 MVA

Figure 10 Example circuit for switching surge calculation

Substation 2

Substation 1
Circuit breaker

Frequency dependent

Circuit breakerCircuit breaker

Substation 3
Frequency dependent

Substation 4
Frequency dependent

Figure 11 XTAP data for switching surge calculation

234

275 kV50 Hz

voltage [kV]






Cumulative frequency distribution [%]








time [ms]













(a) Phase-a voltage

voltage [kV]









Voltage [p.u.]time [ms](b) Phase-b voltage

Figure 12 Cumulative frequency distribution ofthe voltages generated at the end ofvoltage [kV]transmission line 1








T. Kono, “Theory of System Insulation,” Corona
Publishing Co., Ltd. 1984.

[2]

H. W. Dommel, “Digital computer solution ofelectromagnetic transients in single- and multiphasenetworks,” IEEE Trans., Power Apparatus andSystems, Vol. PAS-88, No. 4, pp. 388399, Apr.
1969.


time [ms](c) Phase-c voltage

Reference[1]



Figure 13 Waveforms of the voltages generated
at the end of transmission line 1 in the casecorresponding to the 2% value[3]

T. Noda, “Identification of a multiphase network
equivalent for electromagnetic transient calculations
using partitioned frequency response,” IEEE Trans.,
Power Delivery, Vol. 20, No. 2, pp. 11341142, Apr.
2005.

[4]

Lightning Protection Design Committee, Power station and Substation Sub-working group, “Guide to Lightning
Protection Design for Power Stations, Substations and Underground Transmission Lines,” Central Research Institute ofElectric Power Industry, Comprehensive Report T40, December 1995.
* This document was revised to the following document
Lightning Risk Research Committee, Power Station and Substation Lightning Risk Sub-Working Group, “Guide toLightning Protection Design of Power Stations, Substations and Underground Transmission Lines (rev. 2011),” Central
Research Institute of Electric Power Industry, Comprehensive Report H06, September 2012.

[5]

T. Ono, “Study on Switching Over-voltages in Power Systems,” Central Research Institute of Electric Power Industry,Comprehensive Report 121, January 1985.

235

[6]

A. Ametani, “Distributed Parameter Circuit Theory,” Corona Publishing Co., Ltd., 1990.

[7]

H. Matsubara, T. Ono, H. Motoyama, “Reduction of Switching Overvoltage on Transmission Lines,” Central ResearchInstitute of Electric Power Industry, Report T87085, Aug 1988

[8]

IEEE Working Group 15.08.09, “Modeling and analysis of system transients using digital programs,” IEEE PES Special
Publication, TP-133-0, 1998.

[9]

CIGRE Working Group 13.05, “The calculation of switching surges,” ELECTRA, No. 32, pp. 1742, 1974.

[10] T. Noda, M. Takasaki, “A sub-network reduction method for the calculation of electromagnetic transients in a powersystem (Part 1),” Research Report T03013, March 2004[11] T. Watanabe, S. Funabashi, J. Sadakaji, “A method of switching surge analysis for 1000 kV power transformationequipment and results,” IEE Japan, Switching Protection and High-voltage Research Group SP-94-47, HV-94-118, 1994.
[12] J. Takami, S. Okabe and E. Zaima, “Lightning surge overvoltages at substations due to backflashover with assumed
lightning current waveforms based on observations,” IEEE Trans., Power Delivery, Vol. 25, No. 4, pp. 29582969, Oct.
2010.
[13] J. Takami, S. Okabe and E. Zaima, “Study of lightning surge overvoltages at substations due to lightning strokes tophase conductors,” IEEE Trans., Power Delivery, Vol. 25, No. 1, pp. 425433, Jan. 2010.

END

236

Revision History

Date

Example fileContent changes

version2014/11/19

2.0

2013/08/20

1.0

Update for XTAP Version 2.00
Recalculation of line constants following the changes in XTLCCreation of the first edition (for XTAP Version 1.21)

237

XTAP Example Collection
Example
Name
Field

No.

SSW-03

Analysis of Input Surge Owing to Block Energization during Large-scale System
Restoration
Large-scale system restoration, input surge calculation, switching surge calculation,
overvoltage analysis“Review of Simulation Methods for Electromagnetic Transients in Power Systems andReference

Simulation Cases Using XTAP (Part 1): Switching Transient Overvoltage
Simulations,” Central Research Institute of Electric Power Industry, Study Report
H12005.
After describing the concept, the principle of occurrence, and the simulation method of

Outlineeach element of the transmission system for the case of an input surge owing to block
energization during a large-scale system restoration, we introduce an example
analyzing a 500/275-kV line using XTAP (excerpted from the above document).

1 Input Surge Owing to Block Energization during Large-scale System
Restoration
1.1 Concept , [2]When a power outage occurs across a system of a power company within the entire area (hereinafter referred
to as a “complete outage”), it is necessary to restore the system by starting from the voltage of the adjacentinterconnected power company or from a self-starting generator. When adding voltage to a normal transmissionline, as described in example SSW-02, we conduct an operation of setting the power transmission line to a noload and turning on the circuit breaker at the power supply side for each transmission line. However, when
restoring a complete outage, because a large amount of time is needed until restoration if the lines are energized
one by one, as in the case of a normal voltage addition, we follow a procedure in which the system is divided
into several blocks and the blocks are energized individually [3] (in this case, each block is set to no-load).
Because many transmission lines and substations are included in a single block, the length of the transmission
line becomes long, and its inductance and capacitance also increase. Further, the inductance and capacitance ofthe transformers in the substation add to this. Thus, in the case of energizing each block, as the stored
electromagnetic energy increases owing to an increase in the equivalent inductance and capacitance, problems ofan overvoltage and the processing energy of the arrestors tend to occur more frequently compared to the case ofadding voltage to each line [4]. Therefore, for the purpose of examining the appropriate block division during a238large-scale system restoration, and to grasp the overvoltage and processing energy of the arrestors in thedetermined block division, it is necessary to analyze the input surge phenomenon from a block energization.
1.2 Simulation of Each Element
The method of analysis of the input surge generated through block energization for a large-scale restoration isbasically the same as the analysis method of the normal input surge described in SSW-02. However, considering
the substation equipment included in the energized block, it is necessary to simulate the transformers and phasemodifying equipment not present in the normal input surge analysis. Therefore, beginning from the next section,
we will describe the method for simulating thetransformers and phase-modifying equipment.
R1

1.3 Transformer Simulation
The transformer is simulated using the “basic
Primaryside

equivalent circuit of a transformer,” which is a

L1

Excitationinductance

R2

L2
Secondaryside

ExcitationresistanceR3

L3

well-known circuit available in text books. Figure
Tertiaryside

1 shows its one-phase component. The equivalentcircuit is composed of a winding resistance,
leakageinductance,

excitation

resistance,

excitation inductance, and an ideal transformer. In

Ideal transformer

R1, L1: Resistance and leakage inductance of the primary winding
R2, L2: Resistance and leakage inductance of the secondary winding
R3, L3: Resistance and leakage inductance of the tertiary winding

the case of a two-winding transformer, the leakage
inductance is either equivalently placed on the* The figure above shows a one-phase componentof a three-winding transformer. In the case of a twowinding transformer, the tertiary winding is notneeded.

primary side or divided between the primary andsecondary sides in an appropriate ratio. In the caseof a three-winding transformer, as the leakage

Figure 1 Fundamental equivalent circuit of a

inductances between the primary and secondarytransformer

sides, secondary and tertiary sides, and tertiary andprimary sides are given as the test results, theinductances of the primary, secondary, and tertiarysides are calculated from these data and used. The

Magnetic
flux

leakage inductance on the secondary side often has
a negative value, and when this value is used as is,a numerical instability may occur in the simulationin certain cases [5. In some cases, it is

Current

recommended to input a small positive inductancevalue while conducting an analysis using XTAP.
Assuming the excitation resistance as linear, theresistance obtained from the result of the no-load
test at the rated voltage is set. Simulating the

Figure 2 Hysteresis characteristics of iron core239

excitation inductance as non-linear, saturation
characteristics(current–magnetic-flux

characteristics) of the iron core provided by the
manufacturer is set. If the saturation characteristicsof the iron core provided by the manufacturer also
includethehysteresis

characteristics,

thecharacteristics passing through the center is used as(a)  Connection

shown in Figure 2. In this manner, the equivalentcircuit for the phase of a transformer is obtained.
This is prepared for the three phases, connected inpatterns such as Y or YYand taken as thetransformer model. In XTAP, because models ofthese three-phase transformers have been preparedin advance, an analyst does not have to apply thewiring (however, the excitation inductance issimulated as an exterior non-linear inductance).

(b) Y Connection

1.4 Simulation of the Phase-modifying
Equipment

Figure 3 Shunt reactor models

In the course of a large-scale system restoration,
when energizing the blocks of the system one byone, each block is set to no-load. Therefore, to prevent a steady overvoltage owing to the Ferranti effect after
energization, a shunt reactor, which is a type of phase-modifying equipment, may be added in certain cases [1],[4]. As shown in Figure 3, in the shunt reactor model, each phase of the reactor is simulated using a seriesconnection of the inductance and resistance representing a loss, and these models are connected and used as theactual machine. If the saturation characteristics of the iron core is available, the inductance is taken as non-linear.
Because shunt reactors are connected to the tertiary side of the main transformer, even in the model, they areconnected in the same way as the actual machine. In addition, because the models of the shunt reactors are
already available in XTAP, the analyst does not need to build an equivalent circuit.
1.5 Harmonic Overvoltage
In the block energization during a large-scale system restoration, a sustained harmonic overvoltage with low
attenuation occurs owing to a different principle from the input surge overvoltage explained here, of which
studies are ongoing in Japan and abroad [4], [6]. This is an overvoltage caused by the flow of the excitation
inrush current of the transformer at the timing that strengthens the natural vibration at each cycle when thenatural frequency of the block to be energized matches an integral multiple of the lower order of the commercial
frequency. Because the natural vibration is excited as long as the above-mentioned timing is met, and thedamping is small as long as the block that is energized is on a no-load, the attenuation is small.

240

1.6 Example of Analysis using XTAPConsidering the 500 / 275-kV system shown in Figure 4 as one of the blocks during a complete outage restoration,
we calculate the overvoltage generated at the timeof energizing this block. The power source side of
the circuit breaker to be closed is simulated using

R1

R3

L1

L3

the equivalent circuit of example SSW-02 shown

Tertiaryside

in Figure 7, taking into consideration thefrequency characteristics. Here, assuming thefrequency characteristics to be same as thePrimaryside

R2

characteristics of a transformer with a 500-MVA

L2Secondaryside

capacity indicated in the CIGRE guidelines [7],the constants of the equivalent circuit arecalculated from the assumed short circuit capacity

Excitation
resistance

Ideal transformer(10,000 MVA) (the calculated constants of which

Excitationinductance

R1, L1: Resistance and leakage inductance of the primary winding
R2, L2: Resistance and leakage inductance of the secondary winding
R3, L3: Resistance and leakage inductance of the tertiary winding

are shown in Figure 4).

Transformer* The
above1figure shows one phase in the case of
Shunt-reactor
a three-winding transformer

Figure5 Busbar
Basic
Busbar2
4 equivalent circuit of an
Busbar 5

autotransformer

Busbar 1
Equivalent circuit on the power
supply side

Transmission
line 3500 kV, 100 km

275 kV, 50 km

Circuit breaker

12.8 
60.1 mH

Transmission
line 1T1

1000 

Transformer 2
Transformer 4

T2
500 kV 6.47 mH
60 Hz

Busbar 3
Transmission
line 2500 kV, 75 km

Transformer 3

Figure 4 Example circuit for the calculation of input surge owing to block energization

241

Because the circuit breaker to be closed has the closing resistance of example SSW-02 shown in Figure 9(b), thevalue of the input resistance is set to 1,000 .
The average time T1 from the close command to the resistance input, and the average time T2 from the resistance
input to the resistance short-circuit, are both set to 10 ms, and the variation of the resistance input 1 and thevariation of the resistance short-circuit 2 are both set to 1.5 ms. Transmission lines 1 and 2 are both 500-kV
overhead transmission lines with lengths of 100 and 75 km, respectively. Further, transmission line 3 is a 275-kV
overhead transmission line with a length of 50 km. Assuming the pillar assembly of the transmission lines to be
the standard pillar assembly in [8], and the earth resistivity as 100 m, we simulated the assembly using afrequency-dependent model. The line constant calculation frequency samples for the creation of the frequencydependent model are set to 400 points within the range of 0.1 Hz to 10 MHz. Transformers 1, 2, and 3 are 500-,
275-, and 77-kV autotransformers, respectively, having the same constants, and are simulated using theequivalent circuit shown in Figure 5. Although the basic concept of this equivalent circuit is the same as thebasic equivalent circuit of a transformer described in Section 1.3, compared with Figure 1, this is modified to an
autotransformer connection where the low-voltage side of the primary winding and secondary winding arecommon. Transformer 4 is a three-winding transformer of 275 / 77 / 22 kV, which is simulated using the basic
equivalent circuit of Figure 1. In all transformers, the saturation characteristics of the excitation inductance is set
to a two-line approximation, as shown in Figure 6. Further, it is assumed that shunt reactors are connected to thetertiary side of each transformer, and are simulated using the -connection model shown in Fig. 3 (a). For
simplicity, the inductance and resistance values are calculated from the assumed capacity without considering thenon-linear characteristics of the reactor. The constants of the transformer and the shunt reactor above are shownin Table 1. The lightning was simulated using a non-linear resistor having the characteristics of a 30% reduced
high-performance lightning arrestor described in [8]. Regarding the installation location and quantity, a resistor
was installed on buses 1 through 5. The XTAP data thus created are shown in Figure 7.

242

Using

these

data,

we

conducted

1200

300

800

command is issued as a random number

400

磁束[Wb]
Magnetic
flux [Wb]simulations with the timing at which the close

uniformly distributed between 0° and 360° of thepower supply phase. The calculation timeincrement is set to 0.5 s. Figure 8 shows the

0
-400
-800

maximum value of the voltage generated in bus 5-1200-6000 -4000 -2000

plotted in pu on a normal probability plot (taking

02000 4000 6000

Current[A]
電流[A]

the peak of the nominal voltage to ground as 1pu). Here, the largest among the maximum valuesof the voltage waveforms of each phase was taken

Figure 6 Saturation characteristics of excitation

as the maximum. From Figure 8, the 2% value,

inductancewhich is the maximum value of the generated
voltage corresponding to a cumulative frequency
of 2% is approximately 1.8 pu. A plot of theTable 1 Transformer and shunt-reactor constants

voltage waveform of each phase of the caseused in analysis

corresponding to this 2% value is shown in Figure(a) Transformers 1–3

9. In the case corresponding to the 2% value, it
can be confirmed that the maximum voltage is
generated in phase-b.

R10.152 L1

87.6 mH

R2

0.227 L2

17.8 mH

R3

0.178 L3

4.30 mH

Excitation resistance

400 k

(b) Transformer 4
R1

0.756 L1

60.2 mH

R2

0.237 L2

0.001 mH

R3

0.0484 

L31.93 mH

Excitation resistance

150 k(c) Shunt-reactor
Transformers 1–3 (per phase)
R

0.741 L

590 mH

Transformer 4 (per phase)R

243

0.121 L

96.3 mHFigure 7 XTAP data for the calculation of input surge owing to block energization

244


voltage [kV]






Cumulative frequency distribution [%]











time [ms](a) Phase-a voltage
voltage [kV]















time [ms](b) Phase-b voltage

Generated voltage [pu]voltages appearing in busbar 5voltage [kV]Figure 8 Cumulative frequency distribution of the

[2]





References
[1]












time [ms]M. M. Adibi, “Power system restoration:
Methodologies & implementation strategies,” IEEE
Press, 2000.

(c) Phase-c voltage

M. M. Adibi and L. H. Fink, “Restoration fromcascading failures,” IEEE Power & Energy Magazine,
pp. 68-77, Sep./Oct. 2006.

Figure 9 Waveforms of the voltages appearing atbus 5 for a 2% value[3]

M. M. Adibi and R. J. Kafka, “Power system
restoration issues,” IEEE Computer Applications inPower, Vol. 4, No. 2, pp. 1924, Apr. 1991.

[4]

M. M. Adibi, R. W. Alexander and B. Avramovic,“Overvoltage control during restoration,” IEEE Trans.,
Power Systems, Vol. 7, No. 4, pp. 14641470, Nov.
1992.

[5]

X. Chen, “Negative inductance and numerical instability of the saturable transformer component in EMTP,” IEEE
Trans., Power Delivery, Vol. 15, No. 4, pp. 11991204, Oct. 2000.

[6]

K. Furukawa, T. Funakoshi, T. Kawachino, S. Takasaki, T. Shimojo, K. Hirayama, T. Sogabe, T. Shimamura, S. Nohara,“Transient overvoltage in a high-voltage transformer and countermeasures to prevent overvoltage under applied voltage
at black start,” IEEJ Journal B, Vol 125, No. 1, pp. 53

[7]

CIGRE Working Group 13.05, “The calculation of switching surges,” ELECTRA, No. 32, pp. 1742, 1974.

[8]

Lightning Protection Design Committee, Power station and Substation Sub-working group, “Guide to Lightning
Protection Design for Power Stations, Substations and Underground Transmission Lines,” Central Research Institute ofElectric Power Industry, Comprehensive Report T40, December 1995.
* This document was revised to the following document.

245

Lightning Risk Research Committee, Power Station and Substation Lightning Risk Sub-Working Group, “Guide toLightning Protection Design of Power Stations, Substations and Underground Transmission Lines (rev. 2011),” Central
Research Institute of Electric Power Industry, Comprehensive Report H06, September 2012.

END

Revision History
Date

Example File

Content changes

Version2014/11/19

2.0

2013/09/09

1.0

Update for XTAP Version 2.00
Recalculation of line constants owing to the change in XTLCCreation of the first edition (for XTAP Version 1.21)

246

XTAP Example Collection
Name of the
ExampleField

No.

SSW-04Comparison of the Ground-fault Surge Analysis of a 500-kV Power Transmission
Line with the Actual Measurements
Ground fault surge calculation, switching surge calculation, overvoltage analysis
Central Research Institute of Electric Power Industry, Survey Report H12005,

References“Review of Simulation Methods for Electromagnetic Transients in Power Systems
and Simulation Cases using XTAP (Part 1): Switching Transient Overvoltage
Simulations”
We describe the concept of a ground-fault surge, its principle of occurrence, andOutlinethe method of simulation of each element in the power transmission system, andintroduce an example of the ground-fault surge analysis of a 500-kV transmission
line using XTAP (excerpted from the above reference).

1 Ground-Fault Surge
1.1 Concept [1]-[4]When a ground-fault occurs in the transmission lines from a lightning strike or similar event, an overvoltage
occurs during a healthy phase. In the case of multi-line transmission lines, an overvoltage also occurs in eachphase of the circuits other than the circuit where the fault occurred. This is called the ground-fault surge. If theinsulation level of the transmission line is inadequate, there is a risk of a one-phase ground fault accident
progressing into a multi-phase or a multi-circuit ground fault accident. Hence, from this perspective, an
overvoltage needs to be considered during the insulation design. In particular, if an overvoltage from an input
surge is kept low by a resistance input or by the effect of arrestors in the substation, the ground-fault surge
becomes a problem in the insulation of the transmission lines. Because an overvoltage from a ground-fault surge
is an overvoltage that occurs in transmission lines rather than in substations, the unavailability of an effective
suppression method such as the installation of an arrestor is also a feature.
Herein, we briefly summarize the difference of this concept from the input surge. As described in example
SSW-02, the magnitude of an input surge overvoltage is statistically evaluated from the results of numerous
input surge calculations conducted by randomly changing the closing timing of the circuit breaker. Specifically,in the case of no closing resistance, the magnitude of the input surge overvoltage is determined based on thetiming of the close command as well as the closing timing of the three circuit breakers that is stochasticallydetermined based on this timing. If a closing resistance is present, the magnitude of the input surge overvoltage
is determined based on the timing of the close command as well as the close timing of the six circuit breakers

247stochastically, which is also determined based on this timing. Further, the timing of the close command is also
determined stochastically. Thus, because the magnitude of the input surge overvoltage depends on the multiple
timings that are stochastically determined, the probability of occurrence of the combination of the closing timingthat can cause a maximum overvoltage is extremely low. Therefore, rather than finding the maximum value ofthe overvoltage whose occurrence frequency is
extremely low, it is considered reasonable toconduct a statistical evaluation. In contrast, in thePower transmissionline

acase of a ground-fault surge analysis, because thebproblem is a single-line ground-fault that occupiescthe majority of ground-fault accidents, the phase of
Z0ss

t = t0the power supply voltage that causes the groundfault is gradually changed to obtain the maximum(a) Ground-fault of phase-a at t = t0

value of the overvoltage. Although the timing at
which the event that causes a ground-fault, such as||

lightning, is probabilistic, it does not depend on
stochastically determined multiple timings, as inathe case of an input surge, but depends only on thebtiming at which a single-line ground fault occurs.

cHence, it is considered that the ground-fault at the

Va Vb Vc

phase that gives the maximum overvoltage may besufficient.

Usually,anovervoltage

Powertransmission line

Z0ss

becomes(b) Steady voltage

maximum when a ground-fault occurs just after the
instant when the power supply voltage of the+ground-fault phase attains the peak value.
Powertransmission line

a

1.2

b

Principle of Occurrence of Ground-faultcSurge [1], [2], [4]The mechanism of occurrence of a ground-t = t0fault surge can be easily understood by

Va

Z0ss

considering it as the overlap of a transient
voltage caused by a ground-fault of a steady(c) Transient voltage

voltage from a source voltage. The steady* A ground fault is assumed to be at the center of
the transmission line
* The impedance of the power supply side issimply displayed with only the inductance* Z0ss is the combined value of the characteristic
impedance of the transmission line connected to
the load side substation

voltage corresponds to the voltage before theoccurrence

of

the

ground-fault.

However,considering a transmission line where the source
voltage is virtually zero and the voltage

Figure 1 Mechanism of occurrence of ground-occurring in the ground-fault phase before thefault surge

248ground-fault is V(t), the transient voltage is the voltage occurring in each phase when a voltage of V(t) is
superposed on the ground-fault phase at the time of the ground fault.
Considering the superposition of the steady voltage and the transient voltage, the voltage of the ground-fault
phase will become zero (ground-fault condition) owing to the cancellation of V(t) by V(t), and the voltage ofthe healthy phases will be the sum of the transient voltage generated by applying V(t) to the ground-fault phaseand the steady voltage. Taking the case of a ground-fault of phase-a of a single-circuit transmission line as an
example, Figure 1 shows an illustration of the mechanism of occurrence of the above-mentioned ground-fault
surge.
Here, we discuss the transient voltage component in more detail. Because the largest overvoltage is generated
when a ground fault occurs at the center of the transmission line owing to reasons to be described later, weconsider the ground fault to be at the center of the transmission line. First, the transient voltage component of theground-fault surge is the voltage that occurs in each phase when the voltage in the ground-fault phase before theground-fault is V(t) and when a voltage of V(t) is applied on the ground-fault phase at the time of occurrenceof the ground fault, when considering a transmission line with a source voltage of virtually zero. For example, if
the ground-fault occurs when the voltage of the ground-fault phase has a positive peak value, V(t) rises
stepwise from zero to the negative peak value Vmax, and thereafter varies sinusoidally. However, because theovervoltage in question occurs within a time region shorter than a cycle of the commercial frequency, V(t) canbe approximately regarded as a step wave rising from zero to Vmax without any problem, as shown in Figure2(a). When such a step wave is applied to the ground-fault phase at the center of the transmission line where thesupply voltage is virtually taken as zero, an overvoltage represented by the following equation occurs during theinduction phase.
Vm  

Z0m
Vmax
Z0s

(1)

249

t = t0

However, Z0s is the component of the ground-fault
phase of the characteristic impedance matrix, andTime

Z0m is the mutual component between the induction

V(t)

Vmax

phase in focus and the ground-fault phase of thesame matrix; in addition, although Z0m / Z0s depends
on the type and arrangement of the line, it normally(a) Voltage of ground-fault phase at ground-fault
point

takes a value of between 0.1 and 0.35. From theabove, the waveform of the voltage Vm of theinduction phase at the center of the transmissiont = t0line can be illustrated, as shown in Figure 2(b).
Next, we consider the waveform of this surge when

it has propagated to a point of distance x. It is

Time

Z0m
Vmax
Z0s

known that when a surge propagates along atransmission line, it splits into several line waves(b) Voltage of induction-phase at ground-fault
point

and ground return waves. A line wave is a
propagation mode where the forward and return
paths of the current flow are the phase conductorsof the transmission line and the propagation speedt = t0is almost equal to the speed of light. Precisely,

Time

1

although several line waves exist, for simplicity,

0

we collectively refer to them as a line wave. InVmax

contrast, the ground return wave is a propagation
mode where the three phase conductors of the(c) Voltage of ground-fault phase at point P

transmission line are collectively taken as thecurrent forward path and the ground is taken as thet = t0

return path. The propagation speed of the ground1

return path is approximately 90% of the speed ofthe line wave. Taking the propagation speeds of theline wave and the ground return wave as v1 and v0

Z0m
Vmax
Z0s

Time

0

respectively, the time required for the arrival of the(d) Voltage of induction phase at point P

line wave and the ground wave to point P at
distance x from the ground-fault point is 1 = x / v1
and 0 = x / v0, respectively. That is, at point P, as

Figure 2 Branching phenomenon

shown in Figures 2(c), and 2(d), the line wave
arrives at a time 1 after a ground-fault, after which
the ground return wave arrives after the time  =
0  1 has elapsed. Because the line wave that

arrives earlier takes the ground-fault phase as the forward path and the induction phase as the return path, the250

voltage of the induction phase is opposite in polarity to the ground-fault phase. The earth return wave has thesame polarity because both the ground-fault phase and the induction phase are forward paths. Therefore, during
 when the line wave has arrived, if the induction phase has an opposite polarity to the ground-fault phase, thatis, when V(t) is a step wave that rises negatively, as currently considered, a positive surge occurs. If theunderground cable and many overhead transmission lines are connected to the busbars of the substations (or the
power stations) at both ends of the transmission lines, and if the condition in which the characteristic impedance
seen from the transmission line causing the ground-fault is sufficiently small is satisfied, the line wave surgewith a positive polarity will return to the center of the transmission line as a negative surge owing to a negativereflection at both ends of the busbars. At this time, because the source voltage of the ground-fault phase is near
the positive peak, the source voltage of the induction phase is negative, and hence, a negative overvoltage
determined by the surge impedance shown in Figure 2(b) is superimposed. Here, a negative surge that isnegatively reflected from both sides of the above-mentioned transmission line is superimposed, and further,when the surge is reflected at the ground-fault point, the negative surge of the inductance phase increases. In
detailed surge calculations, the overvoltage of inductance phase can theoretically bridge up to a maximum of
approximately 2 pu. However, this is the most severe case in which the ground fault occurs exactly at the centerof the transmission line, and the condition in which the characteristic impedances of the busbars at both ends ofthe transmission line are extremely small was satisfied. The ground-fault surge occurring in reality will besmaller. The reason why the overvoltage reaches the maximum when the ground-fault occurs at the center of thetransmission line is because the line wave surge negatively reflected from both sides of the transmission line issuperimposed at the same timing.
Because the propagations of multiple line-waves are not the same, the propagation of few tens of kilometerscauses a difference of approximately a few s in the propagation time. Thus, although it depends on the line
length, protrusions on the order of a few s to a few tens of s occur in the voltage waveform, which may
become the peak value. However, because the duration of these protrusion-like voltages is within the time rangeof the lightning surge rather than the switching surge, considering the characteristics (V  t characteristics) where
the insulation performance of the insulator becomes higher as the time of voltage application becomes shorter, it
can be ignored as a switching surge-like overvoltage.

1.3 Simulation of Each Element
The simulation of each element in the ground-fault surge calculation can be basically considered similar to the
input surge calculation described in example SSW-02. Similar to the case of an input surge calculation, thedifficulty is to determine how far the system must be simulated. Basically, we follow each transmission line fromboth ends of the transmission line to be analyzed, and simulate in detail up to those substations that are
determined to have a large influence on the analysis results, taking into consideration the type and length of thetransmission line, as described in example SSW-02 of section 1.7. The systems far from the substation located at
the end of the detailed simulation range are simulated with the appropriate equivalent circuits, considering
whether the system seen from there is a power source or a load. Because it is not easy to specify the range of the251system that must be simulated in detail, in reality, trial ground-fault surge calculations are conducted to observe
its influence. Hence, it is often determined through trial and error. In the calculation of a ground-fault surge,
because the calculation results will be toward the severe side as the impedance of both ends of the transmission
line becomes smaller, upon knowing that the results of the severe side can be obtained, a practical way toconnect a voltage source and a small impedance on both the ends of the line is also possible.

1.4 Example Analysis using XTAP
We introduce an example of calculating the ground-fault surge overvoltage for the case of a ground-fault
occurring in phase-b of the first circuit, exactly at the center of power transmission line A, in the 500-kV powertransmission system shown in Figure.
Here, G1 and G2 are the power stations simulated using the equivalent circuit of Figure 6 in the example
SSW-02. The value of the inductance is calculated from Xd" of the assumed generator and the impedance of thestep-up transformer. Here, G3 and G4 are the contracted representations of the systems far from busbars 3 and 4,
respectively, and are simulated using the equivalent circuit of example SSW-02 shown in Figure 6. Assumingthat the power source frequency is 60 Hz and the tidal current does not affect the overvoltage, the phases of thefour three-phase voltage are assumed to be the same. Assuming all power transmission lines to be standard
columns for 500 kV, as mentioned in [6], and the ground resistivity to be 100 m, the simulation is conductedusing a frequency-dependent line model. Samples of the line constant calculation frequency for creating thefrequency-dependent line model were set to 400 points within the range of 0.1 Hz to 10 MHz. The ground-fault
resistance, which simulates the arc resistance at the time of a ground-fault and the tower foot grounding
resistance, is set to 1 . For the sake of simplicity, the load was not simulated. The XTAP data simulated in thismanner are shown in Figure 4. Using these data, we conducted a simulation by changing the timing at which theground-fault occurs from 0° to 180° of the phase of the power source voltage in increments of 1°, anddetermined the characteristics of the maximum overvoltage with respect to the phase of the power source at thetime of occurrence of the ground-fault. The pu (the peak value of the nominal voltage to the ground is 1 pu) of
the results is shown in Figure 5. Because the protrusion-like voltage waveform generated by the propagation
time difference between multiple line waves is also simulated, the calculation time step is taken as 1 s. From
Figure 5, if the ground-fault occurs when the phase of the power source is 50°, the generated overvoltage attains
the maximum value. Figure 6 shows the voltage waveforms of each phase calculated under this condition. The
maximum overvoltage is generated in phase-c of the first line. The file name of this example is SSW-04-A.xsf.

252

G1
1

G2
23.1
mH

2

75 km

23.1
mH

75 km

Transmission line A

0.263 G3

3

44.0 mH

50 km

100 km100 kmGround fault in phase-b offirst line

Figure 3 Example circuit of ground-fault surge calculation

253

4

0.525 64.9 mH

G4

Frequency dependence

Frequency dependence

Figure 4 XTAP data for ground-fault surge calculation

254


Maximum value
1.52 pu

Overvoltage maximum value [p.u.]























Phase of supply voltage at time of ground fault occurrence [deg]voltage [p.u.]

Figure 5 Maximum voltage versus phase of source voltage at time of occurrence of ground-fault



b




a

c

1.52 p.u.
time [ms]







voltage [p.u.](a) First circuit



b




c

a









time [ms](b)Second circuit

Figure 6 Calculation results at the time of occurrence of ground-fault when the phase of the source
voltage is 50°

255

1.5 Comparison with Actual Measurement
Results

Groundfault pint

ReinanSubstation

To reproduce the results of the ground-fault

75.5 kmKaga
substation

68.5 kmsurge test [5] conducted at the 500-kV Kaga-aReinan line connecting the Hokuriku ElectricbPower Kaga Substation and Kansai Electric Powerc

Reinan Substation, we introduce an exampleasimulation using XTAP. As described in Sectionb

1.2, the transient voltage component generated by ac Second line

ground-fault surge is obtained as the voltage

IG

generated when the voltage of the ground-fault

First line

Impulsegenerator

50 

50 

phase before the ground-fault is V(t) and when avoltage of V(t) is applied to the ground-fault(a) Test circuit

phase at the time of occurrence of the ground-fault,
when considering a transmission line where thesource voltage is virtually zero. Therefore, in this800 F10 

surge test, the source voltage is made zero by

Step
voltagesourcegrounding each phase of the busbars in both endsof the Kaga-Reinan line to an impedance, and a
voltage corresponding to the above-mentioned(b) Equivalent circuit of impulse generator

V(t) is applied to the ground-fault phase at the
ground-fault point such that only the transient
voltage component is generated. The impedance

15.2

Overhead ground wire
14.1 Reinan side IACSR 120 mm2
Kaga side AS 100 mm2

18.4

13.8

30.6

Main line2
11.7 ACSR 410 mm  4
Elementary conductor interva

that grounds each phase of the busbars at both endsof the line simulates the impedance of the system
far from it in a simple manner, and herein wecompare the actual measurement result with thecalculation result when this is set to a resistance of

20.8

50 . Figure 7 shows the test circuit, the equivalent

20.9

circuit of the impulse generator used for thevoltage application, and the conductor arrangement(unit m)of the Kaga-Reinan line. The ground-fault point is(c) Conductor arrangement of Kaga-Reinan line

not the exact center of the line but is only 3.5 km
from the center toward the Kaga Substation side.
Further, the phase of the ground-fault is phase-b of

Figure 7 Ground-fault surge test at Kaga-Reinan

the first line. Figure 8 shows the XTAP simulation

line

data of this test circuit. The transmission lines are

256simulated using a frequency-dependent line model, and the frequency samples for a constant line calculation
were set to 400 points within the range of 0.1 Hz to 10 MHz. The calculation time step of the simulation was set
to 0.5 s. Figure 9 shows the results of the actual measurement, and Figure 10 shows the calculation results.
Including the protrusion-like voltage waveform caused by the difference in the propagation time of multiple line
waves, the calculation results agree well with the actual measured results. The file name of this example is SSW04-B.xsf.

257

Reinan Substation

Kaga Substation

Frequency
dependent

Impulsegenerator

Figure 8 XTAP data reproducing the ground-fault surge test of Kaga-Reinan line

Phase-aof first circuit
Phase-aof second circuit

Maximumvalue

Maximumvalue

Phase-cof first circuit

Phase-bof second circuit

Maximumvalue

Maximumvalue

Phase-cof second circuit

Maximum
value

Figure 9 Measured results of ground-fault surge test at Kaga-Reinan line

258

a相











time [ms]





c相









time [ms]










time [ms]

a相














time [ms]






time [ms]






time [ms]Phase-b
2 回線
of第secondcircuit

b相





Phase-b
Of first circuit
第 1 回線 bphase)
相（印加相）(superimposingPhase-a2 回線
of第secondcircuitPhase-c
1 circuit
回線
of第
first








voltage [p.u.]

voltage [p.u.]voltage [p.u.]

voltage [p.u.]





Phase-a回線of第first1circuit

voltage [p.u.]

voltage [p.u.]





Phase-c
2 回線
of第secondcircuit

c相







Figure 10 Calculation results of Kaga-Reinan line ground-fault surge test using XTAP

References[1]

T. Kono, “Theory of System Insulation,” Corona Publishing Co., Ltd. 1984.

[2]

T. Ono, “Study on Switching Over-voltages in Power Systems,” Central Research Institute of Electric Power Industry,Comprehensive Report 121, January 1985.

[3]

T. Watanabe, S. Funabashi, J. Sadakaji, “A method of switching surge analysis for 1000 kV power transformationequipment and results,” IEE Japan, Switching Protection and High-voltage Research Group SP-94-47, HV-94-118, 1994.

[4]

E. W. Kimbark and A. C. Legate, “Fault surge versus switching surge: A study of transient overvoltages caused by lineto-ground faults,” IEEE Trans., Power Apparatus and Systems, Vol. PAS-87, No. 9, pp. 17621769, Sep. 1968.

[5]

T. Ono, H. Matsubara, “Prediction of Fault Surge with TNA,” Central Research Institute of Electric Power Industry,
Research Report 177042, March 1978

[6]

Lightning Protection Design Committee, Power station and Substation Sub-working group, “Guide to Lightning
Protection Design for Power Stations, Substations and Underground Transmission Lines,” Central Research Institute ofElectric Power Industry, Comprehensive Report T40, December 1995.
* This document was revised to the following document
Lightning Risk Research Committee, Power Station and Substation Lightning Risk Sub-Working Group, “Guide toLightning Protection Design of Power Stations, Substations and Underground Transmission Lines (rev. 2011),” Central
Research Institute of Electric Power Industry, Comprehensive Report H06, September 2012.

END

259

Revision History
Date

Example File

Content changes

Version2014/11/19

2.0

2013/09/09

1.0

Update for XTAP Version 2.00
Recalculation of line constants owing to change in XTLCCreation of first version (for XTAP Version 1.21)

260

XTAP Example Collection
Example
Name
Field

No.

SSWD-01

Ground-fault Surge Propagation Analysis of a 6.6-kV Power Distribution LineGround fault surge analysis(1) Electric Technology Research Association, Advanced Technical Committee onDistribution Automation Technology, “Advancements in Power Distribution

References

Automation Technology,” Vol. 72, No. (2016)(2) Kyushu Electric Power Co., Ltd., Kyushu Technosystems Co., Ltd., The
Method of Calculating Surge Propagation Speed in Accident Point Locating
System, Japanese Patent No. 5085111, 2012.11.28
At present, if a ground-fault accident occurs in a distribution line, it takes asignificant amount of time to locate the accident site. To solve this problem, various
ground fault location methods have been proposed. In this example, we describe theprinciple of standardization using the “ground-fault surge arrival time difference

Outlineanalysis method,” which is one of the ground-fault location standardization
methods. Next, we describe the modeling of various types of power distribution
equipment and finally conduct a ground-fault surge propagation analysis of thesystem created by combining the modeled distribution equipment. Based on the
results, we study the standardization using the “ground-fault surge arrival timedifference analysis method.”

1 Ground-fault Surge Arrival Time Difference Analysis Method
1.1 Standardization Principle
In the ground-fault surge arrival time difference analysis method, the surge wave caused by a groundfault is measured using sensors located at two points between which the ground-fault location issandwiched, and the ground-fault location is determined from the difference in the arrival time [1]. As
shown in Figure 1, when a ground-fault accident occurs in a distribution line, a ground-fault surge
propagates along the distribution line from the fault point to both the power supply side and the load side. A
ground-fault surge is basically a zero-phase component that takes each phase of the electric line as theforward path and the ground (or the overhead ground wire) as the return path. It is assumed that the groundfault occurs in a section of length l [m] sandwiched between two sensor-equipped switches, and that the261

ground-fault surge is detected by these two switches. Taking the time detected by the sensor equipped
switch closer to the ground-fault point (near side) as t1 and the time detected by the sensor equipped switch
far from the ground-fault point (far side) as t2, the difference in the arrival time is t = t2 – t1 [s]. At thistime, if the surge propagation velocity v [m/s] in the section is known, the distance from the far-side
sensor to the ground-fault, that is, the standard distance x [m], can be obtained from equation (1).
xvΔt  l2(1)Here, according to the distributed constant circuit theory, if the distribution line does not have
any incidental facilities, the surge propagation speed will be close to the speed of light (300 m/µs).
However, in actual distribution lines, because the measured surge waveform is complexly distorted
owing to the occurrence of branching, as well as the characteristics of various distribution
equipment and current detection sensors, the speed calculated from the difference in surge arrival time isdifferent from the speed of light. Although this propagation speed is referred to as the “apparent surge
propagation speed,” for simplicity, we refer to it as the “surge propagation speed” in this example.
To know the surge propagation speed within the section that differs from the speed of light, it is
not practical to conduct an artificial ground-fault test in an actual distribution line. Therefore, themethod shown in Figure 2 is used, in which the switch outside the target section is operated, andfrom the difference in arrival time measured when the switching surge caused by this passes through thesection, the surge propagation velocity v [m/µs] within the section is calculated.
Thus, the surge propagation velocity within the section is calculated in advance, and at the time of aground-fault, this surge propagation velocity is substituted in equation (1) to standardize the ground
fault location. Please note that this method is patented [2].
As mentioned above, the surge propagation velocity v, which changes depending on the branching, andthe characteristics of various accessory power distribution equipment and current detection sensors, can beconsidered an important parameter that directly influences the accuracy of the standardization.

262

Surge arrival time differenceSurgearrival time t1

Ground-fault point
Standardized distance
Near-side
sensorarrival time

Section distance l

Substation

〜
〜

Surge

 t = t2 ? t1

velocity

velocity

v

v

t2

x

〜
〜

Far side
sensor

Propagation of ground-fault surge
Fig. 1 Standardization principle of ground-fault surge arrival time difference analysis methodSurgearrival time t2

Surge arrival time difference t = t2 ? t1
Section distance

SubstationSurge
arrival time

t1

l

〜
〜

〜
〜
sensorsensorSection propagation velocity

v = l / t

Propagation of switching
surge

Figure 2 Acquisition of surge propagation speed

1.2 Modeling of Various Distribution Equipment

As mentioned earlier, the surge propagation velocity of the power distribution line varies depending on
the branching, as well as the characteristics of various distribution equipment and the current detection
sensors. Further, this surge propagation velocity has a significant influence on the standardization accuracyof the ground-fault surge arrival time-difference analysis method. Hence, to examine the realistic accuracyof the standardization, it is necessary to reproduce the model of the actual distribution line by reproducing
the branching on the analysis program, and conduct a simulation taking into consideration thecharacteristics of the various distribution equipment and current detection sensors. To do so, an analysis
model must first be created for the various distribution equipment and current detection sensors. In the
ground-fault surge arrival time-difference analysis method, focusing on the point in which thestandardization of the ground fault location is conducted using the zero-phase component of the groundfault surge, we modeled the following distribution equipment that influences the zero-phase circuit by theground capacitance. The modeling of each distribution facility is described in Sections 1.3 and 1.4.
Furthermore, in Section 1.5, we describe the modeling of the current detection sensor used to measure theground-fault surge.
Distribution line (overhead ground wire and cable portion)Pole-mounted transformer and high-voltage receiving transformer263

1.3 Modeling of Distribution LineTo analyze the surge propagating from the ground-fault point, it is necessary to treat the distribution line
as a distributed constant circuit. Therefore, we created the distribution line model, as shown in Figure 3,
using the fixed parameter distributed constant line model. The earth resistivity is taken as 100 m, and theline constant calculation frequency is taken as 100 kHz. For an overhead ground wire, the concrete pillars
are arranged such that the overhead ground wire is grounded once every 200 m or less. The grounding
resistance was set to 40  and an impedance of 12 H of the ground wire was also simulated.
A fixed parameter distributed constant line model was also used for the cable portion, and its line
constant was calculated using XTLC. Similar to the overhead ground wire, one end of the metal sheath in
the cable part is grounded with a grounding resistance of 40 through an inductance of 12 H simulating
the inductance of the grounding wire.
GW: 22 mm21m

0.8 m 0.8 m: ACSR- OC 120 mm2

High-voltage line

11 m

GroundFigure 3 Pillar assembly of high-voltage cable and overhead ground wire

1.4 Modeling of Pole-mounted Transformer and High-voltage Receiving Transformer
From the viewpoint of simulating the surge response corresponding to the zero-phase component, thecapacitance between the primary winding and the casing is considered to be the most important factor in
the modeling of the pole-mounted transformer and high-voltage receiving transformer. Therefore, thiscapacitance was proportionally allocated to the primary side terminals and connected between the casing,and the resistance corresponding to the load on the secondary side was connected between the primary side
terminals. The models of the pole-mounted transformer and high-voltage receiving transformer created inthis manner are shown in Figures 4 and 5, respectively. For the capacitance between the primary winding
and the casing, from the capacitance value actually measured on a pole-mounted transformer with acapacity of 5 to 50 kVA, an approximate equation of the following formula was created, of which thecapacitance value Cs [pF] was determined from the transformer capacity Srate [kVA] (the capacity of thepower receiving transformer is the contract capacity).

264

Cs  21.64  ln( Srate )  22.42(2)

Figure 4 Pole-mounted transformer model

Figure 5 High-voltage receiving transformer model

1.5 Modeling of Current Detection Sensor
For the purpose of measuring the zero-phase current, a zero-phase current transformer (ZCT) is
incorporated in the sensor-equipped switch. Because the primary purpose of the ZCT is to measure thezero-phase current at a commercial frequency, it is not designed to measure high-frequency components
accurately. Therefore, by applying a lowpass filter (LPF) with a cut-off frequency of 100 kHz in the ZCToutput, the frequency characteristic of the ZCT is simulated. Furthermore, by adopting an LPF, variations in
the frequency characteristics based on the type and individual differences of the ZCT can be avoided.

2 Analysis Circuit and Conditions
Figure 6 shows the 6.6-kV distribution system to be analyzed. Here, we assume a case in which theground fault occurs at a location sufficiently far from the distribution substation. In this system, three
sensors, from sensor 1 on the substation side to sensor 3 at the end, are installed uniformly at a distance of
2,000 m from each other, and the section between sensors 1 and 2 is called Section A, whereas the section
between sensors 2 and 3 is called Section B. Using this system, we conducted an analysis of the groundfault surge propagation and examined the standardization using the “ground-fault surge arrival timedifference analysis method.”The other analysis conditions are as follows.
ꞏ

Calculation time step: 10 ns

ꞏGround-fault occurrence time: 1 sꞏGround-fault phase: Phase-aꞏCalculation start time: 0 sꞏCalculation end-time: 40 s

265Distribution substation
system voltage and background impedance

constant parameter

constant parameter

constant parameter constant parameter constant parameter constant parameter

GPT resistance converted to primary side

Sensor1
constant parameterground-fault point
Case2

constant parameter

constant parameter

constant parameter

constant parameter

sensorconstant parameter

constant parameter

pole-mounted Tr

pole-mounted Tr

pole-mounted Tr

pole-mounted Tr

Section A

区
間
A

constant parameter

constant parameter

constant parametergroundingSensor2

constant parameter

constant parameter

constant parameter

constant parameter

constant parameter

sensorconstant parameter

constant parametergroundinggroundinggroundingpole-mounted Tr

pole-mounted Tr

pole-mounted Tr

pole-mounted Tr

Section B

区
間
B

ground-fault point
Case3

constant parameter

constant parameter

constant parameter

constant parameter

constant parameter

constant parameter

constant parametergroundinggroundinghigh-voltage
consumerpole-mounted Tr

Sensor3
constant parameter

constant parameter constant parameter constant parameter

constant parameterground-fault point
Case1

pole-mounted Tr

endsensorFigure 6 Analysis circuit(when the ground-fault point is set to the point at the end of the system, as in Case 1)266

3 Calculation of Section Propagation Velocity
To calculate the standardized distance using the ground-fault surge arrival time-difference
analysis method, it is first necessary to calculate the propagation velocity v of each section from thearrival time difference when a surge is from the outside of the section targeted by the groundfault location standardization sandwiched between the sensors. In an actual distribution line, theabove-mentioned method of operating the switch outside the target section and calculating the
propagation velocity of the section from the arrival time difference of the switching surge will notcause a ground-fault, which can hinder the supply. In a simulation study, it is easy to cause a groundfault. Therefore, propagation velocity v of each section is obtained by causing a ground-fault at the endof the system, indicated as “ground-fault point Case 1,” and determining the ground-fault surge
propagating from the ground-fault point toward the substation side. The instant at which the sensor output
exceeds 0.3 A is determined as the surge arrival at each sensor, and the section propagation velocity is
calculated from the time difference.
Figure 7 shows the analysis results of the case in which the ground-fault occurred at the point indicatedas “ground-fault point Case 1.” From the figure, the propagation velocity vA of Section A calculated fromthe surge arrival time difference of 21.69 µs was approximately 92 m/µs, and the propagation velocity vBof section B calculated from the arrival time difference of 7.49 µs was approximately 267 m/µs,
which indicates that the propagation velocity of the sections differs significantly. The difference is due tothe cable portion. Owing to the influence of the large ground capacitance of the cable, a surge reflection
occurs, and the surge passing through the cable greatly decreases. Therefore, in Section A, which has thecable portion, the propagation velocity decreases significantly.

267

Zero-phase current (3I0) [A]control signal

Sensor1 output
センサ1出力Sensor2 output
センサ2出力Sensor3 output
センサ3出力Time [s]Zero-phase current (3I0) [A]control signal

Y-axis (Zero-phasecurrent) is magnified
y軸（零相電流）を拡大

7.49 s21.69 sTime [s]Figure 7 Output of each sensor and surge arrival time difference when ground-fault occurs at “ground-fault
point Case 1”

4 Standardization of Ground-fault Location

4.1 Standardization of Ground-fault Location in Section A
In this section, using the propagation velocity vA of Section A calculated above, we apply
a standardization for a case in which a ground-fault occurs at the location indicated as “Ground-fault
point Case 2” in Section A. From the analysis results shown in Figure 8, the difference in surge arrival
time ∆t between sensors 1 and 2 is 12.10 µs. From this result, when calculating the standardized
distance x by substituting v (= vA = 92 m/µs) and ∆t (= 12.10 µs) in equation (1), the ground-fault point isdetermined to be at 1,557 m from sensor 2. In the analysis setting, because the location set as the groundfault point (true value) is 1,800 m from sensor 2, the error of the standardized distance is 243 m, and its
error rate is 14%.

268

零相電流（3I） [A][A]Zero-phasecurrent0(3I0)
control signal

Sensor2 output
センサ2出力Sensor1 output
センサ1出力Time [s]Zero-phasecurrent0）(3I0)
零相電流（3I[A] [A]
control signal

Y-axis (Zero-phasecurrent) is magnified
y軸（零相電流）を拡大

12.10 sTime [s]Figure 8 Output of sensors 1 and 2 and difference in surge arrival time when the ground-fault occurs in“ground-fault point Case 2”

4.2 Ground-fault Point Standardization in Section B
In this section, using the propagation velocity vB of Section B calculated above, we apply
a standardization for the case in which a ground-fault occurs at the location indicated as “Ground-fault point
Case 3” in Section B. From the analysis results shown in Figure 9, the difference in surge arrival time

∆t between sensors 2 and 3 is 1.31 µs. From this result, when calculating the standardized distance x
by substituting v (= vB = 267 m/µs) and ∆t (= 1.31 µs) in equation (1), the ground-fault point isdetermined to be 1,175 m from sensor 2. In the analysis setting, because the location set as the groundfault point (true value) is 1,200 m from sensor 2, the error of the standardized distance is 25 m, and its
error rate is 2%.

269

Zero-phasecurrent0(3I
零相電流（3I） [A]0) [A]
control signal

Sensor3 output
センサ3出力Sensor2 output
センサ2出力Time [s]Zero-phasecurrent0）(3I[A]
零相電流（3I0) [A]
control signal

Y-axis (Zero-phasecurrent) is magnified
y軸（零相電流）を拡大

1.31 sTime [s]Figure 9 Outputs of sensors 2 and 3 and the difference in surge arrival time when a ground-fault occurs at“Ground-fault point Case 3”
END

270

Revision History

Date
2017/08/25

Example File

Content changes

Version
1.0

Creation of first edition (for XTAP Version 2.01)

271

XTAP Example CollectionNo.

STL-01Name ofthe

Simulation of Surge Propagation of a 500-kV Horizontally-arranged Single Circuit
Transmission Line

ExampleField

Surge analysis (overvoltage analysis), line model
Central Research Institute of Electric Power Industry, Research Report H07005

References

A Transmission Line Model for Electromagnetic Transient Analysis Based on the
Frequency Region Partitioning Algorithm
In various surge analyses (overvoltage analyses), it is important to accurately
simulate the surge (high-voltage traveling wave) propagating through thetransmission lines and underground cables. For an accurate simulation of a surge,
an accurate verification of the transmission line model used in the instantaneous
value analysis is indispensable.

OutlineIn this example, we compare the results of a surge propagation characteristic test
conducted on a 500-kV horizontally arranged single circuit transmission line
(Azumo main line of Tokyo Electric Power), and the results of a surge propagation
simulation conducted by modeling the transmission line using a frequencydependent line model of XTAP. Because the results agree well, it can be understoodthat the frequency-dependent line model of XTAP has a sufficient analysis
accuracy.

272

Analysis Circuit and Conditions
Figure 1 shows the assembly of the 500-kV horizontally arranged single circuit transmission line to beanalyzed, and Figure 2 shows the results of the surge propagation characteristic test.

22 m
GW
38.5 m

IACSR
120

GW

14 m

40 cm
IACSR
240

a
28.5 mbcLine length of 83.212 km
Earth resistivity of 200 m

Figure 1 Assembly of 500-kV horizontally arranged single line transmission line

415 
10
F

a

83.212 km

b
15 k

c

A resistance of 15 k is inserted
between the starting and ending edgesof the lines in each phase and theground, respectively

Figure 2 Surge propagation characteristic test circuit

273

[Analysis Phenomenon]
When a surge propagates on a n-phase transmission line, its waveform can be broken down into npropagation modes. Such a three-phase transmission line can be broken down into a first line-wave mode, asecond line-wave mode, and a ground return mode. The first line-wave mode is a propagation mode where
both ends of the power line (phase-a and phase-c) are set as the forward and return paths, respectively, andthe second line-wave mode is a propagation mode where both ends of the power line (phase-a and phase-c)
are the forward paths, and the central power line (phase-b) is set as the return path. Because the current
flows through the conductors in both the forward and return paths, the attenuation is small, and the
propagation speed is almost the speed of light. In contrast, in ground return mode, because the three powerlines form the forward path and the ground becomes the return path, the current flows through the ground.
Hence, the attenuation is large, and the propagation speed is less than the speed of light.
As shown in Figure 2, when a voltage charged in a capacitor is applied to the starting edge of a phase, arectangular voltage is applied. As the rectangular voltage propagates along the transmission line, it splits
into the above three modes (the branching phenomenon), and at the far-end, the arrival of the two line-wave
modes can be first observed followed by the arrival of the ground return mode.

[Creation of the line model]
Using the line constant calculation program XTLC included in XTAP, and inputting the pillar assemblyshown in Figure 1, we create a frequency-dependent line model (for the different types of line models, referto the “Modeling of Transmission/Distribution Lines and Underground Cables”). By using the frequencydependent line model, the skin effect of the transmission lines and ground and its frequency characteristics(which causes the branching phenomenon) can be reproduced.
The conditions for creating a frequency-dependent line model are as follows. We input the pillar
assembly in Figure 1 and the parameters below to create the model file of the frequency-dependent line
model (.xmf). At this time, from the graph displayed on the screen, we confirm whether the frequency
characteristics of the line calculated using XTLC from the pillar assembly (frequency characteristics of the
propagation function and the characteristic admittance) match the frequency characteristics of the line
model (Figures 3 and 4).

·

Frequency range:

0.1 Hz to 10 MHz·

Number of frequency samples: 400 points·

Earth resistivity:

200 m274

1

4

magnitude [mS]magnitude [pu]

0.8
0.6
0.4
Characteristics given by line constants

0.2
0

3

2
Characteristics given by line constants

1

Approximated characteristics

Approximated characteristics

-1

10

0

10

1

10

2

3

4

10
10
10
frequency [Hz]

5

10

6

10

0

7

10

-1

10

0

1

10

10

(a) Absolute value characteristics

4

5

6

10

10

7

10


Characteristics given by line constants

Characteristic given by line constants

phase angle [deg]phase angle [deg]3(a) Absolute value characteristics



Approximated characteristics







2

10
10
10
frequency [Hz]

100

101

102

103
104
frequency [Hz]

105

106









107

(b) Phase characteristics

Approximated characteristics

100

101

102

103
104
frequency [Hz]

105

106

107

(b) Phase characteristics

Figure 3 Approximation of frequency

Figure 4 Approximation of frequency

characteristics of the propagation function matrix

characteristics of the characteristic admittancecomponent (1, 1)

matrix component (1, 1)

275

[Surge Propagation Simulation]
Creating the circuit shown in Figure 2 using XTAP, we apply an instantaneous value analysis. We set thetransmission line to a frequency-dependent line model and double-click it to specify the model file (.xmf)
created in advance in XTLC.

[Analysis Conditions]
The analysis conditions are as follows:·

Calculation time step

0.2 s·Calculation start time

0 ms

·

Calculation end time

0.5 ms·

Display start time

0.2 ms·

Display end time

0.5 ms[Example of XTAP Input]
An instance of creating this example on XTAP is shown in Figure 5.
Example Name: Simulation of Surge Propagation in a 500-kV Horizontally Arranged Single-circuit
Overhead Transmission Line, No. TL-01

Frequency dependent

Line type: IACSR240
4 Conductor
Line length: 83.212 km

Line type: IACSR240
4 Conductor
Line length: 83.212 km
The line constant is calculated using TLC under the following conditions
・Frequency-dependent line model
・Frequency sample: 0.1 Hz to 10 MHz
・Number of frequency samples: 400 points
・Earth resistivity: 200 m

Analysis conditions
Calculation time step:
Calculation start time:
Calculation end time:
Display start time:
Display end time:

0.2 s0.0 s
0.5 ms0.2 ms
0.5 msAssembly diagram

Figure 5 Example of XTAP input

Analysis Results
The results (far-end voltage waveform) of executing this example in XTAP are shown in Figure 6.

276

Further, the corresponding measured waveforms are shown in Figure 7. It can be seen that the calculation

Node voltage [V]results agree well with the measurement results.

Time [s]Figure 6 Results of analysis using XTAP

Figure 7 Measurement results

END

277

Change Log
Date

Example fileContent changes

version2014/11/19

2.0

2012/07/19

1.2

Update for XTAP Ver. 2.00
Recalculation of line constants owing to change in XTLC
Update for XTAP Version 1.20
Update for XTAP Version 1.11

2011/10/18

1.1

Change in the line constant file owing to changes in number ofvalid digits in XTLC

2010/07/16

1.0

Creation of first edition (for XTAP Version 1.10)

278

XTAP Example CollectionNo.

STL-02Name ofthe

Simulation of Surge Propagation in a 500-kV Vertically Arranged Double-circuit
Transmission Line

ExampleField

Surge analysis (overvoltage analysis), line model
Central Research Institute for Electric Power Industry, Research Report H07005,

Reference“A Transmission Line Model for Electromagnetic Transient Analysis Based on the
Frequency Region Partitioning Algorithm”
In various surge analyses (overvoltage analyses), an accurate simulation of thesurge (high-voltage traveling wave) propagating through the transmission lines andunderground cables is important.

OutlineIn this example, we conduct the simulation of a surge propagation in a 500-kV
vertically arranged double circuit line commonly used in Japan. A frequencydependent line model that takes into consideration the skin effect of the electric wireand earth, as well as its frequency characteristics, is used.

279

Analysis Circuit and Conditions
Figure 1 shows the pillar assembly of the analyzed 500-kV vertically arranged double circuit
transmission line, and Figure 2 shows the surge propagation simulation circuit.

14.5 mac'b11.2 m b'

GW
IACSR
150

PW
c

79.5 m

11.6 m57.5 m72.0 m

a'

TACSR
410
6

12.0 m43.0 m50 cm

e = 300 m

Line length of 200 km
Earth resistivity of 300 m

Figure 1 Pillar assembly of 500-kV vertically arranged double circuit transmission line280

300 

a

200 km

b
1 pu

c
c'
b'
a'

Figure 2 Surge propagation simulation circuit[Analysis Phenomenon]
When a surge propagates on a n-phase transmission line, its waveform can be decomposed into npropagation modes. Such a six-phase line can be decomposed into five line-wave modes and one groundreturn mode. In line wave mode, because the forward and return paths are power lines and the current flows
through a conductor, the attenuation is less, and the propagation speed is almost the speed of light. In
contrast, in ground return mode, because the six power lines form the forward path and the ground forms
the return path, the current flows through the ground. Hence, the attenuation is large, and the propagation
speed is less than the speed of light. A detailed explanation of the current distribution in each of the fiveline modes is given in the following study.
Ametani, “Distribution Parameter Circuit Theory,” Corona Publishing Co., Ltd.
As shown in Fig. 2, when a rectangular voltage wave is applied to a phase of the starting edge, thevoltage waveform branches into the above six modes (branching phenomenon) as it propagates along thetransmission line, and at the far-end, it is possible to observe the five line-modes arriving first, followed by
the arrival of the ground return mode.

[Creation of the Line Model]
Using the line constant calculation program XTLC included in XTAP, and inputting the pillar assemblyshown in Figure 1, we create a frequency-dependent line model (for a classification of the line models,refer to “Modeling of Transmission/Distribution Lines and Underground Cables”). Using the frequencydependent line model, the skin effect of the transmission lines and the ground, along with its frequency
characteristics (which cause the branching phenomenon), can be reproduced.
The conditions for creating a frequency-dependent line model are as follows. We input the pillar
assembly in Figure 1 and the parameters below to create the model file of the frequency-dependent line
model (.xmf). At this time, from the graph displayed on the screen, we confirm whether the frequency
characteristics of the line calculated using XTLC from the pillar assembly (frequency characteristics of the
propagation function and the characteristic admittance) match the frequency characteristics of the line
model (Figures 3 and 4).
· Frequency range:

0.1 Hz to 10 MHz· Number of frequency samples:

400 points

281

300 m· Earth resistivity:
1

components (1, 1) and (2, 3)
Absolute value characteristics

magnitude [pu]

0.8
0.6
0.4
Characteristics given by line constants

0.2
0Approximated characteristics

10 -1

10 0

10 1

10 2 10 3 10 4
frequency [Hz]

10 5

10 6

10 7

10 6

10 7

10 6

10 7

10 6

10 7

phase angle [deg]Phase characteristics







Characteristics given by line constantsApproximated characteristics



10 0

10 1

10 2

10 3
10 4
frequency [Hz]10 5

(1, 1) Elements
0.5

Absolute value characteristics
0.4
magnitude [pu]Characteristics given by line constants

0.3
Approximated characteristics

0.2
0.1
0

10 -1

10 0

10 1

10 2 10 3 10 4
frequency [Hz]

10 5



phase angle [deg]Phase characteristics
Characteristics given by line constants
Approximated characteristics







10 0

10 1

10 2

10 3

10 4

10 5

frequency [Hz](2, 3) Elements

Figure 3 Results of frequency characteristic
approximation of the propagation function matrix

282

5

magnitude [mS]4
3
2Characteristics given by line constants
Approximated characteristics

1
0

10-1

100

101

102 103 104
frequency [Hz]

105

106

107

(a) Absolute value characteristics
Characteristics given by line constants

phase angle [deg]Approximated characteristics







0

10

1

10

2

10

3

4

10
10
frequency [Hz]

5

10

6

10

7

10

(b) Phase characteristics

Figure 4 Results of frequency characteristic
approximation of the characteristic admittancematrix component (1, 1)

283

[Surge Propagation Simulation]
Creating the circuit of Figure 2 in XTAP, we conduct an instantaneous value analysis. We set thetransmission line to the frequency-dependent line model and double-click it to specify the model file (.xmf)
created in advance in XTLC.

[Analysis Conditions]
The analysis conditions are as follows.
· Calculation time step

0.2 s· Calculation start time

0 ms

· Calculation end time

1.0 ms· Display start time

0.5 ms· Display end time

1.0 ms[Example of XTAP Input]
An instance of creating this example in XTAP is shown in Fig. 5.
Example Name: Simulation of Surge Propagation in a 500-kV Vertically Arranged Double-circuit Overhead
Transmission Line, No. TL-02

Frequency dependent

Line type: TACSR 410 6 Conductor
Line length: 200.0 km
The line constant is calculated using XTLC under
the following conditions
・Frequency-dependent line model
・Frequency sample: 0.1 Hz to 10 MHz
・Number of frequency samples: 400 points
・Earth Resistivity: 300 m
Analysis conditions
Calculation time step:
Calculation start time:
Calculation end time:
Display start time:
Display end time:

0.2 s0.0 s
1.0 ms
0.5 ms
1.0 msAssembly diagram

Figure 5 Example of XTAP input

284

Analysis Results

Node voltage [V]Figure 6 shows the results (far-end voltage waveform) of executing this example using XTAP.

Time [s]Node voltage [V](a) Far-end voltage waveform of the first circuit

Time [s](b) Far-end voltage waveform of the second circuit

Figure 6 Results of analysis using XTAP

END

285

Change Log
Date

Example fileContent changes

version2014/11/19

2.0

2012/07/19

1.2

Update for XTAP Ver. 2.00
Recalculation of line constants owing to changes in XTLC
Update for XTAP Version 1.20
Update for XTAP Version 1.11

2011/10/18

1.1

Change in line constant file owing to changes in the number ofvalid digits of XTLC

2010/07/16

1.0

Creation of first edition (for XTAP Version 1.10)

286

XTAP Example Collection
ExampleNo.

Flashover Calculation using Arcing Horn Flashover Model for Short-tail Waves

Name(Detailed Model, AC Superimposition not Possible)

Field

Lightning surge analysis and overvoltage analysis of substations

References

STL-03Motoyama, “Development of a New Flashover Model for Lightning Surge Analysis,”
Journal of IEEJ, B, Vol. 115, No. 7, pp. 839-846, 1995.
In a lightning surge analysis of the substations, the calculation is conducted under
the assumption that the lightning surge enters the substation through a flashover of theupper phase arcing horn of the first steel tower. Further, in a lightning surge analysisof the transmission lines, a ground fault is estimated based on the presence or absenceof a flashover in each arcing horn. Therefore, an accurate modeling of the flashoverphenomenon of an arcing horn is extremely important for an accurate lightning surge
analysis of the substations and transmission lines.
Conventionally, the leader model proposed by Shindo et al. or its simplified version
represented by an inductor and a switch are being used as the flashover model of anarcing horn. However, in these models, it is necessary to conduct preliminary
calculations beforehand assuming a condition in which the arcing horn does not
flashover. Further, although the actual voltage applied to the arcing horn is a short-tailwave, the model constants were derived from experiments using standard waves.

OutlineHence, improvements are necessary. Therefore, Motoyama developed a new arcing
horn flashover model, where, by calculating the progress of the leader (discharge) at
the same time as the lightning surge calculation, the need for a preliminary calculation
is eliminated, and the model constants are derived experimentally using a short-tailwave.
In this example, we conduct a flashover calculation using the above-mentioned arcing
horn flashover model with a short-tail wave (Motoyama model) and compare it withthe measured results.
* The leader model introduced in this example may not be able to conduct proper
calculations in cases in which an impulse voltage is superimposed over a steady AC
voltage, such as the case of a lightning strike when a system voltage is applied. When
applying to cases in which a steady AC voltage is present, the simplified model
introduced in example no. STL-04 is used.

287

Analysis Circuit and Conditions
The analysis circuit is shown in Figure 1. The left side of the analysis circuit is equivalent to an impulse
generator (IG) that generates a short-tail voltage wave, and the voltage generated is applied to the arcing
horn, which is the specimen used.

Detailed Leader Model(Motoyama Model, AC Superposition Impossible)

Figure 1 Analysis circuit[Arcing Horn Flashover Phenomenon]
Here, idealizing an arcing horn and considering it as a rod-rod electrode, we explain the outline of thearcing horn flashover phenomenon. When a high voltage is applied between the rod-rod electrodes, theelectron avalanche first switches over to a streamer discharge between electrodes, where an ionizing wave
then propagates. However, in the phenomenon described thus far, because the current is extremely small,there is no strong light-emission. Subsequently, a leader accompanied by a strong light-emission is
generated from the tip of both electrodes and progresses toward the direction in which the electrodes can beshort-circuited. Finally, both leaders join near the center of the space between electrodes to create a shortcircuit, thereby completing the flashover. The short-circuited leader then shifts to an arc discharge throughthe heating of ionized gas.
In an arcing horn flashover model, we model the above phenomenon and reproduce the presence of aflashover at the time of a lightning strike and the current flowing during the leader progression. From here,
we will outline the arcing horn flashover model for a short-tail wave using Motoyama and its modeling in
XTAP. For details, refer to the references shown on the cover page of this example.

288

[Calculation of Leader Progress Start Condition]
As mentioned earlier, in the phenomenon that occurs until the start of the leader progression, because thecurrent is extremely small, it only has a slight effect on the lightning surge analysis. Therefore, in
Motoyama’s arcing horn flashover model for a short-tail wave (hereinafter abbreviated as “Motoyama
model”), it is assumed that no current flows until the leader starts to progress. In the Motoyama model,
regarding the voltage between the arcing horns V, the time Ts at which the following equation is satisfied istaken as the leader progression start time, and the simulation of leader progression is started from that time.
For a positive polarity,

1 Ts
V (t )dt > 400 D + 50 [kV]
Ts ∫0

(1)

For a negative polarity,

1 Ts
V (t )dt > 460 D + 150 [kV]
Ts ∫0

(2)

where D [m] is the distance between the electrodes of the arcing horn. The left side of the above equations
is the average voltage applied to the electrodes, and it is assumed that when it exceeds the fixed value
defined by D, shown on the right side, the leader starts to progress. In equations (1) and (2), by adopting the
average voltage value calculated through an integration instead of the instantaneous voltage value, thephysical requirement in which “for the streamer discharge to short-circuit between the electrodes andtransform into a leader, it is necessary to supply a certain value of energy or more between the electrodes,”is approximately reproduced.

Figure 2 Implementation of Motoyama model (detailed model)

289

Figure 2 shows the implementation of the Motoyama model on XTAP (this model can be referred byright-clicking on the “Detailed Leader Model” in the circuit diagram of Figure 1 and selecting the“Component Diagram Edit”). In the figure, the part surrounded by the blue dotted line is the part where theleader progress starting condition is calculated. In the implementation, even if there is a time at the
beginning of the analysis in which no voltage is applied between the arcing horns, to integrate correctly, theintegration is started only after reaching a voltage equal to or higher than a certain level.
Note that, with the above method, a proper calculation may not be possible when an impulse voltage issuperimposed during a state in which a steady voltage is applied, such as when assuming a lightning strikewhen a system voltage is used. When applying to cases in which a steady voltage is present, the model inexample number STL-04 is used.
[Calculation of Leader Progression Process]
When the leader progress starting condition is satisfied, a simulation of the leader progress process isstarted. In this simulation, the leader progression speed vLAVE [m/s] is first calculated through the following
equation.
⎛⎞V
vLAVE = K1 ⎜− E0 ⎟⎝ D − 2 X LAVE⎠

(3)

where K1 [m2/V/s] is the velocity proportional coefficient, XLAVE [m] is the leader length, and E0 [V/m] isthe lowest leader progression electric field. Because the leader length XLAVE is the integral of the leader
progression speed vLAVE, it can be calculated through the following equation.
X LAVE = ∫ vLAVE dt(4)

At this time, because the current IL flowing between the arcing horns is the electric charge per unit time
supplied to the leader progressing at a velocity vLAVE, it can be calculated from the following equation.

I L = 2 K 0 vLAVE(5)

where K0 [C/m] is the charge proportional coefficient.
Because E0 in equation (3) and K0 in equation (5) are constants, the following values are used.
E0 = 750 × 103 [V/m],

K0 = 410 × 10−6 [C/m]Because K1 in equation (3) is a value that varies depending on the leader length, we take
K1 = 2.5 [m2/V/s], when 0 < XLAVE < D/4 andK1 = 0.42 [m2/V/s], when D/4 < XLAVE < D/2.
Equation (3) is an expression that conveniently shows that the slope of the leader progression velocity withrespect to electric field is K1. However, when the leader length exceeds D/4, even if K1 = 0.42 is substituted
in equation (3) as is, the correct leader progression velocity cannot be obtained. It is necessary to retain thevalue of the leader progression velocity at the instant in which the leader length exceeds D/4 and calculate
equation (3) in a form in which the leader progression velocity increases with a slope of K1 = 0.42.
The calculation is continued, and when the leader length satisfies the condition of the following equation,
the arcing horns are short-circuited by the leader and the flashover is complete. Subsequently, the leader
shifts to an arc discharge.

290

X LAVE ≥

D
2

(6)

Further, if the voltage V between the arcing horns decreases and the condition in the following equation issatisfied, the value inside the parenthesis of equation (3) becomes negative, and the leader stops its progress.
V < ( D − 2 X LAVE ) E0

(7)

In the implementation of the Motoyama model shown in Figure 2, the part surrounded by the red dottedline applies the calculation of the above-mentioned leader progress process.
[Equivalent Resistance for Simulation]
In this implementation, assuming that no current flows during the time until the condition in which the startof the leader progression in equations (1) or (2) is satisfied, the equivalent resistance value of the arcing
horn is set to a very large R_init value. When the leader starts progressing, the leader current IL is
calculated from the voltage V between the arcing horns using equations (3) through (5), and its equivalent
resistance is set to
Req =

V
IL

(6)

When equation (6) is established and the arcing horns are short-circuited by the leader, the equivalent
resistance value of the arcing horn is set to a small R_arc value corresponding to the arc resistance.
In the Motoyama model shown in Figure 2, the portion surrounded by the green dotted line is the portion
corresponding to the above-mentioned equivalent resistance.
[Setting of Parameters]
Although various parameters can be set by double-clicking the leader model of Figure 1, the parameters to
be set by the user for each analysis are the inter-electrode distance D of the arcing horn and the arc
resistance R_arc. Other configurable parameters and their default values are shown below (in most cases,
there is no need to change the values of these parameters).
Leader constants
Charge proportionality constant K0 [C/m]K0 = 410e-6

Velocity proportionality constant K1 [m2/V/s] (0 < XLAVE < D/4)
K1_1 = 2.5

Velocity proportionality constant K1 [m2/V/s] (D/4 < XLAVE < D/2)
K1_2 = 0.42

Minimum leader progression electric field E0 [V/m]E0 = 750e3

Voltage dead band [V] for calculating leader progression start time
V_margin = 10e3291

Electrical tolerance of leader bridging determination
eps = 0.05

Equivalent resistance before leader progression [Ω]R_init = 1e6Time constant of first-order lag filter for internal processing [s]tau = 10e-9[Analysis Conditions]
The analysis conditions are as follows.
·

Calculation time step

1 ns

·

Calculation start time

0 μs

·

Calculation end time

5 μs

·

Display start time

0 μs

·

Display end time

5 μs

Analysis Results
Figure 3 shows the calculation results of the analysis circuit of Figure 1, where the distance between theelectrodes of the arcing horn D was set to 3 m and the arc resistance R_arc was set to 5 Ω (and the otherparameters were set to the default values). In addition to the voltage between the arcing horns and theleader current, the leader length and leader progression velocity are shown (the leader length and leader
progression velocity are output with the variable names XLAVE and VLAVE, respectively). The results ofthe calculation agree well with the measurement results shown in Figure 4 (reprinted from the references
shown on the cover page of this example).
Figure 5 shows the calculation results of the V − t curve for the cases of positive and negative polarities bychanging the inter-electrode distance of the arcing horn D to 1, 2, and 3 m, respectively.
The calculation results are in good agreement with the measured results shown in the following study.
M. Aoshima, “Flashover Characteristics of Air Gaps by Lighting Impulse Voltages with Short Tail,”
Research Report T87041, Central Research Institute of Electric Power Industry, April 1988.

292

Potential difference between nodes [kV]

3000
2500
2000
1500
1000
500
0

0

1

2

3

4

5

4

5

4

5

時間
[us]
Time
[us]

(a) Voltage between arcing horns

4500

Branch current [A]

4000
3500
3000
2500
2000
1500
1000
500
0

0

1

2

Time
[us]
時間
[us]

3

(b) Leader current

1.6
1.4

Control signal

1.2
1
0.8
0.6
0.4
0.2
0

0

1

2

Time
時間[us][us]3

Control signal(c) Leader length

Time [us](d) Leader progression velocity
Figure 3 Analysis results293

(a) Voltage between arcing horns(b) Leader current(c) Leader length
Figure 4 Measurement results

294

5000

Flashover [kV]

4000

3000
3m
2000

2m

1000

0

1m

0

1

2

3

4

5

Flashover time [μs](a) Positive polarity

Flashover voltage [kV]

5000

4000

3000

3m

2000
2m
1000

0

1m

0

1

2

3

Flashover time [μs](b) Negative polarity
Figure 5 V − t curve calculation result

295

4

5

Revision History
Date

Example File

Content changes

Version2022/10/18

2.1

Name changed from SLS-02 to STL-03

2014/11/19

2.0

Update for XTAP Version 2.00

2012/07/19

1.2

Update for XTAP Version 1.20

2011/12/06

1.1

2011/11/11

1.1

2011/11/04

1.0

Consistency with SLS-03 was obtained for an explanation of
the example title
Clearly specified in the icon of the model that an ACsuperimposition is not possible
Creation of the first edition (for XTAP Version 1.11)

296

XTAP – Example collection
ExampleNo.

STL-04Flashover Calculation Using Arcing Horn Flashover Model (Simplified Model, AC

Name

Superimposition Possible) for Short Tail Waves.

Field

Lightning surge and overvoltage analyses of power stations and sub-stations.
H. Motoyama, K. Shinjo, Y. Matsumoto, and N. Itamoto, N., “Observation andReferences

analysis of multiphase back flashover on the Okushishiku Test Transmission Linecaused by winter lightning,” IEEE Transactions on Power Delivery, Vol. 13, No. 4,
pp. 1391-1398, 1998.
In this example, we simplify the leader progress start condition of the “arcing horn
flashover model (detailed model, AC superimposition not possible) for short tailwaves,” introduced in example STL-03, and introduce a model that allows asuperimposition of the steady AC voltage. For the arcing horn flashover model for
short tail waves, refer to the explanation of example STL-03.
In the detailed model of example STL-03, the average voltage is obtained byintegrating the instantaneous values of the voltages, and the start of the leader
progression is determined based on whether this value exceeds the fixed value basedon the gap length. This simulates the physical requirement that states, “energy
exceeding a certain value must be supplied between electrodes in order to convert a

Outlinestreamer discharge into a leader by short-circuiting the electrodes.” However, when
the impulse voltage is superimposed over a steady AC voltage, such as the case of alightning strike when the system voltage is applied, the integral value of the impulse
voltage rises owing to the previous steady AC voltage, and the calculation of an
accurate average voltage may not be possible. To solve this problem, a simplified
leader progress start condition is adopted, which states that “the leader starts toprogress when the absolute value of the average electric field between the arcing
horns exceeds 500 kV/m,” and with the simplified model introduced in this example,
an analysis is possible even in cases in which the impulse voltage is superimposedover a steady AC voltage.

Analysis Circuit And Conditions

297

Figure 1 shows the analysis circuit. The left side of the analysis circuit is an equivalent circuit of an
impulse generator (IG), which generates a short-tail voltage wave, and the voltage generated by this isapplied to the arcing horn, which is the test specimen.

Simplified leader model(Motoyama model, AC superposition possible)

Figure 1 Analysis circuit[Calculation of the Leader Progress Start Condition]
In the detailed model of example STL-03, the instantaneous value of the voltage is integrated to obtain the
average voltage value, and the start of the leader progression is determined based on whether this value
exceeds the fixed value based on the gap length. However, in cases in which the impulse voltage issuperimposed over a steady AC voltage, such as the case of an assumed lightning strike when the system
voltage is applied, the integral value of the impulse voltage rises owing to the previous steady AC voltage,
and the calculation of an accurate average voltage may not be possible.
To solve the above problem, in the simplified model of this example, a simplified leader progress start
condition is adopted that states, “if the absolute value of the average electric field between the arcing horns
exceeds Es = 500 [kV/m], the leader begins to progress.” Specifically, if D [m] is the distance between theelectrodes of the arcing horn, and the voltage V (t) [KV] between the arcing horn satisfies

V (t )> Es = 500 [kV] ,D
the leader will start progressing.
Figure 2 shows an implementation of Motoyama’s arcing horn flashover model for a short-tail wave,
which adopts the above-mentioned simplified leader progress start condition on XTAP (this model can bereferred to by right-clicking on the “simplified leader model” in the circuit diagram of Figure 1, andselecting “component diagram editing”).
Inside the figure, the section enclosed by the blue dotted line is the part that calculates the start conditionof the leader progress, and is much simpler compared to the corresponding part of the detailed model inexample STL-03.

298

Figure 2 Implementation of the Motoyama model (simplified model)[Calculation of the Leader Progression Process]Although the calculation procedure of the leader progression process is exactly the same as the detailed
model of example STL-03, the settings of the speed proportional coefficient K1 [m2/V/s] differs slightly. If
the inter-electrode distance D of the arcing horn is 1 m or greater, the value will be the same as that of thedetailed model of example STL-03. However, if D is less than 1 m, it is taken as
K1 = 1.0 [m2/V/s] I when 0 < XLAVE < D/4, andK1 = 0.42 [m2/V/s] when D/4 < XLAVE < D/2.
Apart from the above points, there is no difference from the detailed model of example STL-03. Further,
the calculation of the leader progression process is carried out in the portion encircled by the red dottedline in Figure 2.

[Equivalent Resistance for Simulation]
The part that calculates the equivalent resistance is exactly the same as that of the detailed model ofexample STL-03. The calculation of the equivalent resistance is carried out for the portion encircled by the299

green dotted line in Figure 2.
[Parameter Settings]
Various parameters can be set by double clicking the leader model shown in Figure 1. However, theparameters to be set by the user for each analysis are the inter-electrode distance D of the arcing horn andthe arc-resistance R_arc. Other configurable parameters and their default values are indicated below (inmost cases, a modification of these parameters is not required).
Leader constants
Leader progress starting electric field Es[V/m]
Es = 500e3
Charge proportionality constant Ko[C/m]K0 = 410e-6
Velocity proportionality constant K1 [m2/V/s] (0 < XLAVE < D/4, D < 1 m)
K1_1a = 1.0
Velocity proportionality constant K1 [m2/V/s] (0 < XLAVE < D/4, D ≥ 1 m)
K1_1b = 2.5
Velocity proportionality constant K1 [m2/V/s] (D/4 < XLAVE < D/2)
K1_2 = 0.42
Minimum leader progress electric field E0 [V/m]E0 = 750e3
Tolerance of the leader bridging evaluation
eps = 0.05
Equivalent resistance value before leader progression [Ω]R_init = 1e6Time constant of first-order lag filter for internal processing [s]tau = 10e-9[Analysis Conditions]
The analysis conditions are as follows:·

Calculation time step

1 ns

·

Calculation start time

0 μs

·

Calculation end time

5 μs

·

Display start time

0 μs

·

Display end time

5 μs

300

Analysis Results
The inter-electrode distance D of the arcing horn was set to 3 m, and the arc resistance R_arc was set to 5
Ω (the other parameters were set to their default values). The results of the calculation conducted on theanalysis circuit of Figure 1 are shown in Figure 3. In addition to the voltage between arcing horns and theleader current, the leader length and leader propagation velocity are shown (the leader length and leader
propagation velocity are outputs with the variable names XLAVE and VLAVE, respectively). The
calculation results are consistent with the calculation results of the detailed model described in example
STL-03 and with the actual measured results within the range in which no problem occurs in terms ofpractical use.
Figure 4 shows the calculation results of the V − t curve for the cases of positive and negative polarities bychanging the inter-electrode distance of the arcing horn D to 1, 2, and 3 m. In the simplified model of thisexample, because there is no distinction between the positive and negative polarities, the same results canbe obtained in both cases. The calculation results agree well with the calculation results of the detailed
model and the actual measured results shown in the following reference document within the range inwhich no problems occur in terms of practical use.
M. Aoshima, “Flashover Characteristics of Air Gaps by Lighting Impulse Voltages with Short Tail,”Potential difference between nodes [kV]Research report T87041, Central Research Institute of Electric Power Industry, April 1988.

Time [us]Branch current [A](a) Voltage between arcing horns

Time [us](b) Leader current

301

Control signal

Time [us]Control signal(c) Leader length

Time [us](d) Leader progression velocity
Figure 3 Analysis results

5000

Simplified modelFlashover voltage [kV]

Detailed model
4000

3000
3m
2000

2m

1000

0

1m

0

1

2

3
Flashover time [μs](a) Positive polarity

302

4

5

5000

Simplified modelFlashover voltage [kV]

Detailed model
4000

3000

3m

2000
2m
1000

0

1m

0

1

2

3

4

5

Flashover time [μs](b) Negative polarity
Figure 4 V -t curve of calculation results

Revision History
Date

Example fileContent changes

Version2022/10/18

2.1

Name changed from SLS-03 to STL-04

2014/11/19

2.0

Update for XTAP Version 2.0

2012/07/19

1.1

Update for XTAP Version 1.20

2011/12/06

1.0

Initial version created (for XTAP Version 1.11)

303

XTAP Example Collection
Example
Name
Field

No.

PV-01

Simulation of Single-phase Inverter Circuit for Solar Power Generation

Power Electronics“Development of Inverter Simulation Program (Part 2): Verification of Analysis

References

Accuracy by Comparison of Actual Measurements,”

Research Report R07016(June 2008), Central Research Institute of Electric Power Industry
With an increase in the amount of power electronic equipment and complexity, theimportance of a simulation analysis of the power electronic equipment isincreasing.

OutlineIn this example, we simulate an inverter for a recently introduced solar power
generation. We simulate the solar power generation panels with a DC power supply,
boost it with a boost chopper and connect it to a single-phase 200-V AC systemthrough an inverter. The switches and diodes in the circuit are simulated using thesimplified characteristics of ON- and OFF-resistors.

304

Analysis Circuit and Conditions
Figure 1 shows the analysis circuit.
Inverter
Chopper

AC Filter

DC Power Supply

AC System

Figure 1 Analysis circuit[Circuit Operation]
The DC voltage generated by the DC voltage source on the left side of Figure 1 is boosted to 380 V
using a boost chopper. Subsequently, an AC voltage is output using a single-phase inverter and linked to the
200-V system. In this example, because the chopper applies the current control, the ON/OFF timing of theswitch is adjusted such that the current flowing through the chopper smoothing reactor follows thecommand value. The inverter applies a constant DC voltage control and adjusts the output power to the AC
system such that the average value of the DC voltage also follows the command value. Further, to remove
the higher harmonics superimposed on the voltage and power waveforms output from the inverter, it isconnected to the system using an AC filter.
The main circuit parameters of the analysis circuit are as follows:
Chopper smoothing reactor: 1 mH
DC capacitor for energy storage: 5 mF
Reactor for AC filter: 3 mH
Capacitor for AC filter: 30 μF[Analysis Conditions]
The analysis conditions are as follows:·

Calculation time step

0.2 μs·Calculation start time

0 ms

·

Calculation end time

40 ms

·

Display start time

20 ms

·

Display end time

40 ms

305

[Example of XTAP Input]
An instance of creating this example using XTAP is shown in Figure 2.

Figure 2 Example of XTAP input

Analysis Results
The results of executing this example using XTAP are shown in Figure 4. Based on the results of thesimulation, it can be confirmed that a single-phase AC current flows through the inverter to the system side,and that the current after passing through the AC filter is cleaner than the current waveform before passing
through the AC filter over which the higher harmonic wave was superimposed.

Branch voltage [V]Case 1
300
250
200
150
100
50
0
-50
-100
-150
-200
-250
-300
0.02

(a) AC Voltage Waveform at the AC Filter Outlet

0.03

Time[s]v(V_FILX) #1(a) AC voltage waveform at outlet of AC filter

306

0.04

Branch current [A]Time [s]Blue line, AC current before passing through the filter; red line, AC current after passing through thefilter(b) Inverter current waveform
Figure 3 Analysis results
END

307

Revision History
Date

Example File

Content changes

Version2021/06/23

2.1

Name changed to PV-01 (contents unchanged)

2014/11/19

2.0

Update for XTAP Version 2.00

2012/07/19

1.2

Update for XTAP Version 1.20

2011/10/18

1.1

Update for XTAP Version 1.11

2010/07/16

1.0

Creation of the first edition (for XTAP Version 1.10)

308

XTAP Exercises
Exercise

Number

PV-02-A

Simulation of a system fault using a basic transient analysis model of PCS for utility-

namescale solar power station (Switching model)

FieldsPower system analysis, System interconnection, Power electronics
The main references include:[1] Technical Committee on Massively Interconnected Analysis Model forDistributed Generation, "Massively Interconnected Analysis Model for
Decentralized Generating Plants", I.E.E. Japan Technical Report No. 1487,

Reference
documents

2020.
[2] Yonezawa, Fukushima, Noda, Sekiba, Ito, Misawa, Chida, Yamaguchi,
Nakajima, Utsunomiya, and Takeuchi, "Development of a large-scale PV power
generation system model for electromagnetic transient simulations under gridfaults", I.E.E. Japan Joint Workshop on Electric Power Technology/Power
System Technology, PE-15-151, PSE-15-173, pp. 99-104. 2015-09.
Other reference documents are listed at the end of this document.
This exercise simulates the transient response in a utility-scale solar power station
when a system accident occurs in a high-voltage distribution system where thepower station is interconnected. The model for the utility-scale solar uses a basic
instantaneous value analysis model of PCS for utility-scale solars, and generates a
disturbance assuming 3LGO in the upper level system and 3LS in the high voltage
distribution line as the system accident. The model operates normally by outputting
a constant active power during a normal condition in the system voltage, but oncethe voltage drops due to a system accident, depending on the degree and time of theOutlines

drop, the model continues to operate, or performs operations such as gate block
function to stop the operation of the switching device or performing parallel-off
from the system. Also, when the voltage drops, the control to limit the output current
will work.
Distributed power sources such as utility-scale solars normally have various
requirements stipulated by the system interconnection regulations, but the purposeof this exercise is to simulate the transient response characteristics when an accident
occurs. Therefore, the PCS model uses the basic model as it is and does not include
various system protection functions. The FRT requirement also considers only thesimple LVRT (voltage drop tolerance).

309

Circuit to be analyzed and analysis conditions
Fig. 1 shows the circuit diagram of the high-voltage distribution system simulated on the XTAP and theutility-scale solar connected to it. The system frequency is 50 Hz.
A Distribution substation is simulated as a series connection of 0.0001+j0.08 p.u. impedance and voltage
sources with its transformer for distribution and back system together, and the neutral point is grounded witha resistance of 30 kΩ, which is the value of the limiting resistance of the GPT from the primary side [1].
There are two feeders from the distribution substation and a 500 kW utility-scale solar is interconnected at
the end of one of the feeders (4 km away). A 3 MW resistive load is connected at the end of the other feeder.

Exercise name: Basic transient analysis model of PCS for utility-scale solar system (Switching model)
PV-02-A

Intensity ofsolar radiation
Breaker

1-1)
Voltage drop at 0.3 s (1.0→0.7 p.u.)
Voltage restored at 0.37 s (0.7→0.95 p.u.)
1-2)
Voltage drop at 0.6 s (0.95→0.2 p.u.)
Voltage restored at 0.67 s (0.2→1.0 p.u.)

Breaker2-2) Released at 0.8 s

BreakerLoad2-1) Accident at 0.3 s
Fault point(3LS)

Fig. 1 Circuit diagram of the high-voltage distribution system simulated on
the XTAP and the utility-scale solar system connected to it.

The utility-scale solar consists of PV panels with a DC output voltage of about 350 V and an output of
about 500 kW, a PCS model for basic transient analysis, and a 500 VA step-up transformer to boost thevoltage to high voltage. A three-level NPC bridge is used as a inverter inside the PCS, and the switchingfrequency of the PWM control is set to 4.5 kHz. The PCS employs DCAVR for d-axis current control and
AQR for q-axis current control, and specifies 1.0 for the DC voltage command value and 0.0 for the reactive
power command value. Since the PCS model does not include the MPPT control, the DC voltage command
value does not change even if the maximum power output point of the PV panel changes due to changes insolar irradiation, etc. (To simulate changes in solar irradiation, the MPPT control must be added to the PCS).
For the LVRT, which controls implementation of continuous operation and gate blocking, etc., when thesystem voltage drops, gate blocking starts at 0.3 p.u. (gate block time: 0.1 s), and the PCS is disconnectedfrom the system when gate blocking continues for 0.3 s. The current limit will operate when the system
voltage drops less than 0.9 p.u.

310

The following two types are assumed as the system faults.
(1) Assuming 3LG-O in the upper system, the amplitude value of the supply voltage at the distribution
substation will vary as follows.
<1> 0.3 to 0.37 s ... drops to 0.7 p.u.
<2> 0.6 to 0.67 s ... drops to 0.2 p.u.
(2) Assuming 3LS in the high voltage distribution line, three-phase short circuit occurs at 0.3 s at a point
0.2 km from the distribution substation of a feeder where utility-scale solar is not interconnected, andthen the feeder is shut down 0.5 s later (0.8 s).
The PCS model used in this exercise is a model that actually performs switching, and the calculation time
step is set to 2 μs to ensure sufficient simulation of the switching in semiconductor devices.

311

Analysis results
I. Instantaneous voltage drop assuming 3LG-O in the upper system
Fig. 2 shows the system voltage and output current of each phase on the AC side of the PCS, voltage on
the DC side, the active and reactive power output by the PCS, the d-axis current command value (DCAVR
output), the effective value of voltage to determine the voltage drop, and the gate block signal. Each timezone will be discussed below.
(1) 0 to 0.3 s: Operation start to steady state
The PCS will be initialized since it is not initialized at t = 0 (except for the DC capacitor). The DC voltage
rises as power is supplied from the PV panels, but when the active power is output to the AC side, the DC
voltage gradually decreases and converges to the commanded value of 1.0 p.u. (350 V).
(2) 0.3 to 0.6 s: Time from which the fault occurs (system voltage drops down to 0.7 p.u.) through recovery
Due to the occurrence of a fault (drop in the system voltage amplitude value simulating the fault), thevoltage on the AC side of the PCS drops. This lowers the active power that can be transmitted to the AC side,and the DC power rises. Due to the voltage falling below 0.9 p.u., the current limit is activated to prevent
overcurrent, limiting the d-axis current command. Furthermore, since the voltage is above 0.3 p.u., the PCS
will continue to operate.
The voltage recovers at 0.07 s after the fault, but since the gate block time is set to 0.1 s, the gate block
continues for 0.1 s after the fault and then recovers. After recovery, the active power output is restored and it
returns to the steady state. However, the system voltage of 0.95 p.u. after the recovery means the d-axis
current command value is controlled to a larger value than before the fault.
(3) 0.6 to 1.0 s: Time from which the fault occurs (system voltage drops down to 0.2 p.u.) through recovery
Due to the occurrence of an fault (drop in the system voltage amplitude value simulating the accident), thevoltage on the AC side of the PCS drops. In the output current, a whisker-shaped overcurrent flows into thesystem momentarily, but is quickly suppressed by the fast ACR. Since the system voltage has dropped below
0.3 p.u., the gate block signal is output, and then the active power output to the AC side becomes zero. During
the gate block, the breaker remains turned on, generating transient voltages and currents due to the interaction
between the system interconnection reactor and filter capacitor. Since no active power is output to the system,
the DC voltage rises due to the power supply from the PV panels and reaches about 427 V, which is the
maximum open circuit voltage of the PV panels.
Since the voltage recovers from the accident in 0.07 s, it recovers after the gate block time elapsed similarto 2) and returns to the steady state.

312

Control signal

Branch voltage [V]Time [s]Voltage of each phase on the AC side of the PCS

Active power (blue)and reactive power (red) that(line voltage)

the PCS outputs

Control signal

Branch current [A]Time [s]Time [s]Time [s]Current at each phase on the AC side of the PCS

D-axis current command value (DCAVR output)

430
420
410
390

Control signal

Branch voltage [V]

400
380
370
360
350
340
330
320
310
300
0

0.1

0.2

0.3

0.4

0.5

0.6

0.7

0.8

0.9

1

時間 [s][s]
Time

Time [s]Voltage on the DC side of the PCS

Voltage effective value (blue)and gate block signal(red)

Fig.2. PCS response when an instantaneous voltage drop occurs, assuming 3LG-O in the upper system

II. Instantaneous voltage drop assuming 3LS in the high voltage distribution line
Fig. 3 shows the system voltage and output current of each phase on the AC side of the PCS, voltage on
the DC side, the active and reactive power output by the PCS, the d-axis current command value (DCAVR
output), the effective value of voltage to determine the voltage drop, and the gate block signal. Each time
period will be discussed below (0 to 0.3 s is omitted because it is the same as I.).
(1) 0.3 to 0.8 s: Time from which the fault occurs (system voltage drops down to 0.3 p.u.) through recovery
Due to occurrence of the 3LS fault at a feeder coming out of the transformer for distribution, the voltage
on the AC side of the PCS connected to another feeder in the same bank drops. In this case, since the voltage
drops below 0.3 p.u., the PCS will be gate-blocked.
Since the lowered voltage persists even at about 0.61 s, which is 0.3 s after the gate block, the PCS parallels
off the breaker on the AC side. As a result of this, the PCS will be disconnected from the system.

313

(2) 0.8 to 1.0 s: After accidental feeder shutdown
The feeder where the fault occurred is disconnected and the voltage of the feeder to which the utility-scalesolar is connected will be restored. However, this model does not incorporate the PCS into the recovery
operation once it has been paralleled off, so it remains stopped. To simulate the recovery operation after it

Control signal

Branch voltage [V]has been paralleled off, this control system needs to be added.

Time [s]Voltage of each phase on the AC side of

Active power (blue)and reactive power (red) that

the PCS (line voltage)

the PCS outputs

Time [s]Time [s]Current at each phase on the AC side of the PCS

D-axis current command value (DCAVR output)

Control signal

This model does not take
into account the operationof the control system after it
has been paralleled off.

Branch voltage [V]Control signal

Branch current [A]Time [s]Time [s]Time [s]Voltage on the DC side of the PCS

Voltage effective value (blue)andgate block signal (red)Fig.3. PCS response when an instantaneous voltage drop assuming
an 3LS fault in the high voltage distribution line.

314

Reference Document[1] Investigating R&D Committee on Models for Power System Analysis with Integration of Distributed
Generators, "Models for Power System Analysis with Integration of Distributed Generators", I.E.E.
Japan Technical Report No. 1487, 2020.
[2] Yonezawa, Fukushima, Noda, Sekiba, Ito, Misawa, Chida, Yamaguchi, Nakajima, Utsunomiya, andTakeuchi, "Development of a large-scale PV power generation system model for electromagnetictransient simulations under grid faults", I.E.E. Japan Joint Workshop on Electric PowerTechnology/Power System Technology, PE-15-151, PSE-15-173, pp. 99-104. 2015-09.
[3] Nagashima and Noda, "A Dynamic Voltage Simulation Method for Power Distribution Systems by Using
Electromagnetic Transient Analysis (Part 1): Modeling of Distribution Substation and SVR", Central
Research Institute of Electric Power Industry Research Report H13007, 2013.
[4] Investigating R&D Committee on Control Technology for Power Electronics Equipment, "Control
Technology for Power Electronics Equipment," IEEJ Technical Report No. 1084, 2007.

315

Revision History
Date
2021/06/22

Exercise file

Details of changes

version
1.0

First version created (for XTAP Version 3.3)

316

XTAP Exercises
Exercise

Number

PV-02-B

Simulation of a system fault using a basic transient analysis model of PCS for utility-

namescale solar power station (average value model)

FieldsPower system analysis, System interconnection, Power electronics
The main reference documents include:[1] Investigating R&D Committee on Models for Power System Analysis with
Integration of Distributed Generators, "Models for Power System Analysis with
Integration of Distributed Generators", I.E.E. Japan Technical Report No. 1487,

Reference
documents

2020.
[2] Yonezawa, Fukushima, Noda, Sekiba, Ito, Misawa, Chida, Yamaguchi,
Nakajima, Utsunomiya, and Takeuchi, "Development of a large-scale PV power
generation system model for electromagnetic transient simulations under gridfaults", I.E.E. Japan Joint Workshop on Electric Power Technology/Power
System Technology, PE-15-151, PSE-15-173, pp. 99-104. 2015-09.
Other reference documents are listed at the end of this document.
In Exercise PV-02-A " Simulation of a system fault using a basic transient analysis
model of PCS for utility-scale solar power station (Switching model)", the analysis
was performed using a model that takes into account the switching behavior of thesemiconductor devices inside the PCS similar to the actual PCS. In this model,
however, it is necessary to take a smaller analysis time step to simulate theswitching, resulting in too long calculation time. An average value model that

Outlines

averages the on-off switching of semiconductor devices and simulates the voltageof each arm as a voltage source to accelerate the calculation is known. The averagevalue model does not simulate the harmonics associated with switching, but it takes
a longer time step, enabling faster operation. Therefore, this exercise simulates thecase where the PCS model used in Exercise PV-02-A is replaced by the averagevalue model, and compares its operation with the switching model.
Please refer to the exercise slip for Exercise PV-02-A for the circuit conditions,
etc.

317

Circuit to be analyzed and analysis conditions
Fig. 1 shows the circuit diagram of the high-voltage distribution system simulated on the XTAP and theutility-scale connected to it. The system frequency is 50 Hz. Except for the PCS model being the averagevalue model, it is identical to Exercise PV-02-A.
Exercise name: Basic transient analysis model of PCS for utility-scale solar system (average value model)
PV-02-B

Intensity ofsolar radiation
Breaker

1-1)
Voltage drop at 0.3 s (1.0→0.7 p.u.)
Voltage restored at 0.37 s (0.7→0.95 p.u.)
1-2)
Voltage drop at 0.6 s (0.95→0.2 p.u.)
Voltage restored at 0.67 s (0.2→1.0 p.u.)

Breaker2-2) Released at 0.8 s
Load
Breaker2-1) Accident at 0.3 s
Fault point(3LS)

Fig. 1 Circuit diagram of the high-voltage distribution system simulated on the XTAP and the utility-scale solar
system connected to it.

Fig. 2 shows the internal configuration of the three-phase PCS (average value model). Except for the partin the frame with the red dashed line in the figure, it is identical to the switching model. In the average-value
model, the voltage output of each arm is simulated by a voltage source as described later, PWM is notnecessary, and the command value of the output voltage of each phase is input to the bridges as it is.

318

Three-level NPC bridge PCS model for mega-solar

3-level NPC bridge
Breaker

Breaker

Breaker

Output voltage signal ofeach phaseFig. 2 Internal configuration of the three-phase PCS (average value model)
Fig. 3 shows the average value model for the three-level NPC bridge. A current source is placed on the DCside and a voltage source is placed on the AC side. The current source on the DC side flows current with avalue of the current flowing in the voltage source on the AC side multiplied by a signal k equivalent to theratio of PWM on/off (equal to the output voltage signal), and the voltage source on the AC side applies avoltage of the DC voltage (voltage of the DC capacitor) multiplied by k. This makes the power on the DCside equal to the power on the AC side. Various methods have been proposed to implement the averaging
model for [3] to [5]. In this exercise, in order to simulate the transient characteristics of the three-level NPC
bridge, six current sources are placed in parallel with two capacitors, and six voltage sources are placed on
the AC side, as shown in Fig. 3, where the upper and lower sources correspond to the positive and negativesides of the output voltage signal respectively. By configuring the circuit in this way, as described later,transient characteristics equivalent to those of the switching model can be simulated.
When the gate is blocked, the output of each current source and voltage source is set to zero, and the switch
located on the AC side is opened to simulate the gate block operation in a simplified manner. If it is necessary
to correctly simulate the transient behavior of diodes (rectifier circuits) connected in parallel tosemiconductor devices when the gate is blocked, a model that can take this into account should be used (e.g.,[6] and [7]).

319Fig. 3 Average value model for the three-level NPC bridge.

Since this exercise uses the average value model, the calculation time step is set to 40 μs.

320Analysis results
Here, it is explained that only the operation of the on-delay voltage drop assuming 3LG-O in the upper
system.
Fig. 4 shows the results of the switching model (PS-02-A results) cascaded with the results of the averagevalue model in this exercise for the line voltage Vc,ab between ab-phases closer to the inverter side rather than
the system interconnection reactor of the PCS and the line current Ic of each phase during steady state
operation (0.2 to 0.24 s). As can be seen clearly from the figure, the voltage waveform in the switching model
is square-wave-like (due to three levels, three levels of square waves appear in the phase voltage and fivelevels in the line voltage), while the voltage waveform in the average value model is a clean sinusoidal
waveform with no noise. Regarding the current, it can be confirmed that the harmonics originated fromswitching are superimposed in the switching model, while the average value model exhibits a clean waveform
without these harmonics.

500

Potential difference between nodes [V]

400
300
200
100
0
-100
-200
-300
-400
0.2

0.21

0.22

0.23

Time [s][s]
時間

PCS3phNPC1/Vc_ab: u(PCS3phNPC1/Vca)-u(PCS3phNPC1/Vcb) #1*

PCS3phNPC1/Vc_ab: u(PCS3phNPC1/Vca)-u(PCS3phNPC1/Vcb) #2*

Branch current [A]a) Line voltage between ab-phases

Time [s]b) Line current of each phaseFig. 4 Comparison 1 of the switching model and average value model(Blue: switching model, Red: average value model)

321

Fig. 5 shows the results of the switching model (results for PS-02-A) cascaded with the results of theaverage value model in this exercise for the system voltage between the ab-phases on the PCS AC side, theoutput current of the phase a, the voltage on the DC side, and the active and reactive power output by the
PCS. As can be seen clearly from the figure, the switching model and the average value model of this exercise
coincide well for all voltage waveforms, current waveforms, and control signals (output power). (The

Branch voltage [V]waveforms almost overlap. Please calculate each exercise to verify it by yourself.)

Time [s]Branch current [A]a) AC voltage (between ab phases)

Time [s]Branch voltage [V]b) AC current (phase a)Time [s]Controlsignal
制御信号c) DC voltage

Time [s]d) Active power

322

Controlsignal
制御信号

Time [s]e) Reactive power
Fig. 5 Comparison 2 of the switching model and average value model(Blue: switching model, Red: average value model)
Fig. 6 shows the results of calculation of the DC voltage and active power by each model when the input
signal (solar irradiation) to the PV panels is changed in a stepwise manner from 1000 to 600 at 0.3 s insteadof causing a fault using the same circuit (in an actual PCS, the commanded value of the DC voltage changes
by MPPT control when such change in solar radiation occurs; however, this analytical example does not take
MPPT control into account, resulting in the commanded value of the DC voltage being constant.). As can beseen clearly from the figure, the transient changes in each waveform are generally reproduced by the averagevalue model even for steep changes in input.

440
420
400
380
Branch voltage [V]

360
340
320
300
280
260
240
220
200
0

0.1

0.2

0.3
Time
時間[s]
[s]

0.4

0.5

Control signala) DC voltage

Time [s]b) Active power output by PCS
Fig. 6 Comparison 3 of the switching model and average value model(Blue: switching model, Red: average value model)

323

Reference Document[1] Investigating R&D Committee on Models for Power System Analysis with Integration of Distributed
Generators, "Models for Power System Analysis with Integration of Distributed Generators", I.E.E.
Japan Technical Report No. 1487, 2020.
[2] Yonezawa, Fukushima, Noda, Sekiba, Ito, Misawa, Chida, Yamaguchi, Nakajima, Utsunomiya, andTakeuchi, "Development of a large-scale PV power generation system model for electromagnetictransient simulations under grid faults", I.E.E. Japan Joint Workshop on Electric PowerTechnology/Power System Technology, PE-15-151, PSE-15-173, pp. 99-104. 2015-09.
[3] Task force on modeling and analysis of electronically-coupled distributed resources, “Modelingguidelines and a benchmark for power system simulation studies of three-phase single-stage photovoltaic
systems”, IEEE Trans. on power delivery, vol. 26, no. 2, April 2011.
[4] A. Yazdani, “Electromagnetic transients of grid-tied photovoltaic systems based on detailed and averaged
models of the voltage-sourced converter,” IEEE PES General Meeting, San Diego, CA, 2011.
[5] R. Yonezawa, T. Noda, K. Fukushima, T. Nakajima, Y. Sekiba, K. Utsunomiya, E. Ito, K. Misawa, T.
Chida, N. Yamaguchi, and Y. Takeuchi, “Development of detailed and averaged models of large-scale
PV power generation systems for electromagnetic transient simulations under grid faults,” Proc. of IEEEPES Innovative Smart Grid Technologies Asia (ISGT-Asia), pp. 98-104, 2016.
[6] Kikuma, Okada, "Development of Converter Model for High-speed Transient Analysis - Averaged
Model with Rectifier Circuit -," Central Research Institute of Electric Power Industry Research Report
R15022, 2015.
[7] Sano, Yonezawa, Noda: "Instantaneous value analysis model of grid-connected inverters for dynamic
voltage analysis of power distribution systems," Transactions of the I.E.E. Japan, vol. 138, no. 8, pp.
685-694, 2018.

324

Revision History
Date
2021/06/22

Exercise file

Details of changes

version
1.0

First version created (for XTAP Version 3.3)

325