SIMULTANEOUS EHV AC-DC POWER

TRANSMISSION

SYNOPSIS THESIS SUBMtTTED FOR THE AWARD OF THE DEGREE OF

IBottat of pi]tl0Sf0pl|g IN

ELECTRICAL ENGINEERING

BY

HAFIZUR RAHMAN

DEPARTMENT OF ELECTRICAL ENGINEERING

ALIGARH MUSLIM UNIVERSITY

ALIGARH-202002 (INDIA)

JUNE-2006

^ , A7^,« f.^-

SYNOPSIS

SIMULTANEOUS EHV AC-DC POWER TRANSMISSION

Long EHV ac lines cannot be loaded to their thermal limits in order to keep sufficient

margin against transient instability. The transmission angle between sending end and

receiving end in long EHV ac lines seldom exceeds 30°. In these conditions ac current in

the transmission line conductor is much below its thermal current carrying capacity. But

optimum use of transmission lines requires loading of EHV ac lines close to their thenr al

limits. Flexible AC Transmission System (FACTS) components are used to achieve this.

In this thesis it has been shown to achieve the same goal by simultaneous ac-dc power

transmission in which the conductors are allowed to carry superimposed dc current along

with ac current. Zig-zag connected transformers at both end in simultaneous ac-dc power

transmission are used to avoid saturation of core due to dc current injection. Laboratory

modelled experimental verification as well as simulation results establish the feasibility of

simultaneous ac-dc power transmission.

The present research explores the feasibility of converting a double circuit EHV ac line

to simultaneous ac-dc power flow line to get substantial power upgradation of existing

transmission line. In the proposed study of conversion of existing EHV ac transmission

line to simultaneous ac-dc transmission line, no alterations of conductors, insulator strings

and towers structures of the original EHV ac line are required. Substantial gain in the

loadability of the line is obtained. Major advantage of the composite ac-dc line is that the

transmission angle between two ends of line in steady state may go much beyond 30°-45°

values. This is because of inherently fast controllability of dc link embedded in the

composite ac-dc line. Such a large transmission angle is impossible to achieve in pure

EHV ac transmission line

HVDC transmission lines in parallel with EHV ac lines are recommended to improve

transient and dynamic stability as well as to damp out oscillations in power system. In the

present work it has been demonstrated to get advantage of HVDC in parallel with EHV ac

lines by simultaneous ac-dc power transmission. A single machine infinite bus connected

by a double circuit ac line, which is converted for simultaneous ac-dc power transmission

line has been taken up for, detailed study. For the purpose of comparison, the same double

circuit line with pure ac transmission imder same disturbances is also studied. Various

short circuit faults at number of locations of the transmission line have been created and

simulated responses have been studied. A novel approach to solve the first swing stability

problem by simultaneous ac-dc power transmission has been presented. It has been

demonstrated that the stability of the system can be effectively improved by simultaneous

ac-dc power transmission with fast dc power modulation. AC and DC powers flow

independently and added dc power flow does not cause any transient instability.

It has also been shown that it is possible to tap small amount of power, (up to 10% of

total transfer capability of the simultaneous ac-dc transmission system), to feed remotely

located communities, in the same simple way as tapping in case of EHV ac line. It is

economical as compared to complicated methods of tapping from HVDC line. The results

clearly indicate that the lapping of small amount of ac power from the simultaneous ac-dc

transmission line has negligible impact on the performance of composite ac-dc power

flow.

SIMULTANEOUS EHV AC-DC POWER

TRANSMISSION

THESIS SUBMITTED FOR THE AWARD OF THE DEGREE OF

Bnctnr of piitlnanpiig IN

ELECTRICAL ENGINEERING

BY

HAFIZUR RAHMAN

DEPARTMENT OF ELECTRICAL ENGINEERING

ALIGARH MUSLIM UNIVERSITY

ALIGARH-202002 (INDIA)

JUNE-2006

T67^8

CERTIFICATE

certify that the thesis entitled "SIMULTANEOUS EHV AC-DC POWER

[SSION" submitted by Mr. Hafizur Rahman, Reader, Electrical Engg.,

polytechnic, Aligarh Muslim University, Aligarh, India, for the award of the

Doctor of Philosophy (Ph.D.), is the record of the candidate's own work

under my supervision and has not been submitted elsewhere for a degree.

Date: JUTU i^. Zoo6 Signature of Supervisor

(Prof. B. H. Khan)

2).. ^. ^ , Man Professor Dept. of Electrical Engg. A.IVl.U.,Aligarh-2020G2-India

SYNOPSIS

SIMULTANEOUS EHV AC-DC POWER TRANSMISSION

Long EHV ac lines cannot be loaded to their thermal limits in order to keep sufficient

margin against transient instability. The transmission angle between sending end and

receiving end in long EHV ac lines seldom exceeds 30°. In these conditions ac current in

the transmission line conductor is much below its thermal current carrying capacity. But

optimum use of transmission lines requires loading of EHV ac lines close to their thermal

limits. Flexible AC Transmission System (FACTS) components are used to achieve this.

In this thesis it has been shown to achieve the same goal by simultaneous ac-dc power

transmission in which the conductors are allowed to carry superimposed dc current along

with ac current. Zig-zag connected transformers at both end in simultaneous ac-dc power

transmission are used to avoid saturation of core due to dc current injection. Laboratory

modelled experimental verification as well as simulation results establish the feasibility of

simultaneous ac-dc power transmission.

The present research explores the feasibility of converting a double circuit EHV ac line

to simultaneous ac-dc power flow line to get substantial power upgradation of existing

transmission line. In the proposed study of conversion of existing EHV ac transmission

line to simultaneous ac-dc transmission line, no alterations of conductors, insulator strings

and towers structures of the original EHV ac line are required. Substantial gain in the

loadability of the line is obtained. Major advantage of the composite ac-dc line is that the

transmission angle between two ends of line in steady state may go much beyond 30°-45°

values. This is because of inherently fast controllability of dc link embedded in the

composite ac-dc line. Such a large transmission angle is impossible to achieve in pure

EHV ac transmission line

HVDC transmission lines in parallel with EHV ac lines are recommended to improve

transient and dynamic stability as well as to damp out oscillations in power system. In the

present work it has been demonstrated to get advantage of HVDC in parallel with EHV ac

lines by simultaneous ac-dc power transmission. A single machine infinite bus connected

by a double circuit ac line, which is converted for simultaneous ac-dc power transmission

line has been taken up for, detailed study. For the piupose of comparison, the same double

circuit line with pure ac transmission under same disturbances is also studied. Various

short circuit faults at number of locations of the transmission line have been created and

simulated responses have been studied. A novel approach to solve the first swing stability

problem by simultaneous ac-dc power transmission has been presented. It has been

demonstrated that the stability of the system can be effectively improved by simultaneous

ac-dc power transmission with fast dc power modulation. AC and DC powers flow

independently and added dc power flow does not cause any transient instability.

It has also been shown that it is possible to tap small amount of power, (up to 10% of

total transfer capabilit>' of the simultaneous ac-dc transmission system), to feed remotely

located communities, in the same simple way as tapping in case of EHV ac line. It is

economical as compared to complicated methods of tapping fi-om HVDC line. The results

clearly indicate that the tapping of small amount of ac power from the simultaneous ac-dc

transmission line has negligible impact on the performance of composite ac-dc power

flow.

Ill

ACKNOWLEDGEMENTS

The author is deeply indebted to Dr. B. H. Khan, Professor in Electrical Engineering,

for his valuable suggestions and guidance on the research work leading to this thesis.

The author is sincerely grateful to Dr. K. P. Basu, (Retired Professor, Department of

Electrical Engineering), presently at Faculty of Engineering, Multimedia Universit)'.

Selangor, Malaysia for his kind interest and valuable guidelines given in this work.

The author is thankful to Prof Iqbal Ali, Principal, University polytechnic, Aligarh

Muslim University, Aligarh, for encouraging and permitting me to carry out this work.

The author is sincerely thankful to Prof Prabhat Kumar, Chairman, and Pro. A K

Gupta, Prof MS Beg Ex. Chairmen, Department of Electrical Engineering, for their

interest and extending facilities to carry out this work.

The help and regular encouragement by Prof Ekram Husain, Prof Mukhtar Ahmad,

Prof. Anwaruddin Anwar, Prof Shamshuddin, Prof MSJ Asghar, Prof Pervez Mustajab

(Electronics Engg. Deptt.) and Mr. Abdul Shafey, Reader, are sincerely acknowledged.

The author would also like to acknowledge help receive from Mr. Sheikh Moinuddin,

lecturer, of University polytechnic, in carry out the experimental verification of this

work.

Thanks are also expressed to M/S Munawwar Husain, Najmul Hasan STAs in E.E.D.,

Rashid Hasan, TA and Tahir Iqbal, JLA, E.E.S., in assisting the experimental setup of

this work.

Finally, the author wishes to express his sincere thanks to his parents, wife, daughter

and sons for their cooperation, patience and understanding throughout the tenure of the

work.

IV

LIST OF PRINCIPAL SYMBOLS

Id dc current,

la ac current per phase.

I the total rms current per conductor.

Ith thermal limit of a conductor.

ia_meas on-line measured ac current.

Idmeas on-linc mcasured dc current.

IR, IS sending end and receiving end ac currents respectively.

Vdro, Vdio ideal no load dc voltages of rectifier and inverter respectively

E rms line-to-line voltage of converter transformer primary voltage.

T converter transformer turn ratio.

B number of six-pulse bridges in series.

Es, ER sending end and receiving end ac voltages per phase respectively.

R, L, C line parameters per phase of each line.

R q equivalent resistance of the line conductors seen by dc.

Rcr, Rci conunutating resistance of rectifier and inverter respectively.

Xcr, Xci commutating reactance of rectifier and inverter respectively.

a, y firing and extinction angles of rectifier and inverter respectively.

Xd smoothing reactor

Ps , PR sending end and receiving end ac real powers respectively.

Qs, QR sending end and receiving end ac reactive powers respectively.

Pdr, Pdi dc power of each rectifier and inverter respectively.

Pst, Prt total active power at the sending end and receiving end respectively

Qdr, Qdi dc reactive power drawn by each rectifier and inverter respectively.

Qst, On total reactive power at the sending end and receiving end respectively.

P L Transmission line loss.

Vdr, Vdi dc voltage of each rectifier and inverter respectively.

}j.i, }Xr commutat ion angles of inverter and rectifier respectively.

Vph per phase rms voltage of original ac line.

Va per phase voltage of ac component of converted simultaneous ac-dc line.

Vd pole to ground dc voltage of converted simultaneous ac-dc line.

Vdc pole-to-pole dc voltage of converters.

Vmax max imum phase voltage of original ac line.

P'totai total power transfer through the double circuit line before conversion.

Ptotai total power transfer through the composi te line.

X transfer reactance per phase of the double circuit line.

6 power angle/transmission angle between the voltages at the two ends.

5i power angle/transmission angle between the voltages at two ends of line

before conversion.

82 power angle/transmission angle between the ac voltages at the two ends

of the composite line (except in section 5.2).

M multiplying factor,

lord current order for rectifier.

VI

z per phase per unit length per circuit series impedance of line,

Q/km/ph/circuit.;

y per phase per unit length per circuit series admittance of line,

S/km/ph/circuit.

Pm, Pg input and output power of generator.

Pg' transient output power of generator. Pac' transient power of ac system.

Pdc' transient dc power.

(o ,(OQ actual and rated generator angular speeds respectively.

ts.(o angular speed deviation of generator.

vu

TABLE OF CONTENT Chapter Page No.

i SYNOPSIS

ACKNOWLEDGEMENTS iii

LIST OF SYMBOLS iv

INTRODUCTION 1

1.1 General 1

1.2 Literature Survey 4

1.3 Objectives of the Thesis 1 o

1.4 Outline of the Thesis 11

SYSTEM MODELLING 13

2.1 Introduction 13

2.2 Salient Pole / Round Rotor Synchronous 14

Machine Modelling in EMTD /PSCAD

2.2.1 The Per-Unit System 18

2.2.2 Machine Interface to EMTDC 19

2.2.3 Mechanical and Electrical Control 22

2.2.3.1 Exciters 23

2.2.3.2 Governors 23

2.2.3.3 Stabilizers 24

2.2.3.4 Turbines 24

2.3 Transmission Line Model in PSCAD 25

2.3.1 Selecting the Proper Line Model 26

2.4 HVDC Model 27

2.4.1 Six-Pulse Converter Model Block in PSCAD 27

2.4.2 CIGRE HVDC benchmark in PSCAD 29

2.5 Transformer Modelling 34

SIMULTANEOUS AC-DC POWER TRANSMISSION 36

3.1 Introduction 36

vin

3.2 Basic Concept 38

3.3 Analysis 40

3.3.1 Selection of Transmission Voltages 44

3.4 Preliminar}' Economic Consideration 45

3.5 Protection 46

3.6 Effect of Line Capacitance 47

3.7 Feasibility Test for simultaneous ac-dc transmission 48

3.7.1 Simulation Study 48

3.7.2 Simulation Results 50

3.7.3 Experimental Verification 53

3.8 Conclusion 60

POWER UPGRADATION OF EXISTING AC TRANSMISSION

LINES BY CONVERSION TO SIMULTANEOUS

AC-DC COMPOSITE LINES 62

4.1 Introduction 62

4.2 Power Up-gradation of EHV ac Transmission Line 63

by simultaneous ac-dc power transmission

4.3 Case Studies 65

4.3.1 Power Transfer Capability of a pure ac line 65

(A. 1) Computation 65

(A.2) Simulation 67

4.3.2 Conventional double circuit EHV ac line Converted

to Composite ac-dc Power Transmission Line 68

(B.l) Computation 68

(B.2) Simulation 72

4.4 Conclusion 79

TRANSIENT STABILITY IMPROVEMENT

BY SIMULTANEOUS AC-DC POWER TRANSMISSION 80

5.1 Introduction 80

5.2 Control Principle and Strategy 81

IX

5.3 Description of System Studies 86

5.4 Transient State Simulation 87

Al. Simultaneous ac-dc transmission 87

B. Pure ac transmission with series compensation 93

A2. Simultaneous ac-dc transmission (continued) 97

5. 5 Conclusion 121

6 POWER TAPPING IN SIMULTANEOUS AC-DC

POWER TRANSMISSION 122

6.1 Introduction 122

6.2 Small Power Tapping Station Requirements 123

6.3 The System under Study 123

6.4 Digital Simulation of the proposed Scheme 126

Case A. Equal Power Tapping from each line at

different instants 126

Case B. Unequal Power Tapping from each line at different instants 130

Case C. Tapping Power Station Fault Responses 134

6.5 Conclusion 142

7 CONCLUSION 143

7.1 Conclusion of the Present Work 143

7.2 Scope for Future Work 145

LIST OF PUBLICATIONS FROM PRESENT WORK 146

REFERENCES 147

APPENDIX 154

CHAPTER 1

INTRODUCTION

1.1 General

In recent years, various techno-economic reasons such as:

• steadily growing and geographically uneven demand of electrical power,

• the use of larger generating units to reduce the capital cost per MW and to

improve efficiency,

• utilization of remotely located low cost energy resources,

• interconnection of large power pools for reasons of peak diversity and energy

interchange,

• the growing cost of right of way (RoW),

• environmental concerns

and aesthetic reasons have resulted in an increase of transmission system for

transmission of bulk amount of power over long distances. Large steam power stations

are usually located near the coalfields, whilst the hydroelectric power plants are restricted

to sites where sufficient water head is available. The nuclear power stations are also

located at isolated sites. Long lines find their use in cormecting such remotely located

power plants to the load centers. The wheeling of this available energy fi-om these

remotely located sites through existing long EHV ac lines to load centers has the

following limitations.

1. The power flow in ac long lines is limited by stability considerations. This implies that

the lines operate at power level much below their thermal limit.

2. The lack of fast control in ac lines implies that the normal power flow in line is kept

much below the peak value, which itself is limited by stability consideration (as

mentioned above). This margin is required to maintain system security even under

contingency conditions.

3. The ac transmission system requires dynamic reactive power control to maintain

satisfactory voltage profile under varying load conditions and transient disturbances.

4. The increase in load levels is accompanied by higher reactive power consumption in

the line reactance.

Thus these lines are not loaded to their thermal limit to keep sufficient margin against

transient instability. India covers a very large geographical area, and the distribution of

the fijel energy resources of tlie country is not uniform. As a consequence, the length and

power of considerable number of transmission lines of future unified power system will

be such that HVDC transmission would be used to advantage.

HVDC is used to transmit large amounts of power over long distances or for

interconnections between asynchronous grids. When electrical energy is required to be

transmitted over very long distances, it can be more economical to transmit power using

direct current instead of alternating current. For a long transmission line, the value of the

smaller losses, and reduced construction cost of a DC line, can offset the additional cost

of converter stations at each end of the line. Also, at high ac voltages significant amounts

of energy are lost due to corona discharge, the capacitance between phases or, in the case

of buried cables, between phases and the soil or water in which the cable is buried. Since

the power flow through an HVDC link is directly controllable, HVDC links are

sometimes used within a grid to stabilize the grid against control problems with the AC

energy flow. In a number of applications the advantages of HVDC make it the preferred

option over AC transmission.These are:

• point-to-point bulk power transmission over long distances, to overcome stability

problem of ac line.

• Increasing the capacity of an existing power grid in situations where additional

wires are difficult or expei^ive to install.

• Reducing the profile of wiring and pylons for a given power transmission

capacity.

• Stabilising a predominantly AC power-grid, without increasing maximum

prospective short circuit current.

• Reducing corona losses (due to higher voltage peaks) for HVAC transmission

lines of same power.

• Reducing line cost since HVDC transmission requires less conductor (i.e 2

conductors one is +ve another is -ve)

• HVDC can carry more power per conductor, because for a given power rating the

constant voltage in a DC line is lower than the peak voltage in an AC line. This

voltage determines the insulation thickness and conductor spacing. This allows

e.xisting transmission line corridors to he used to carry more power into an area of

high power consumption, which can lower costs.

• HVDC allows power transmission between unsynchronised AC distribution

systems, it can help increase system stability, by preventing cascading failures

from propagating from one part ol' a wider power transmisj:ion grid to another,

whilst still allowing power to be imported or exported in the event of smaller

failures. This has caused many power system operators to contemplate wider use

of HVDC teclinology for its stability benefits alone.

The present situation demands the review of traditional power transmission theory and

practices, on the basis of new concepts that allows full utilization of existing transmission

facilities without decreasing system availability and security.

Recent developments involving deregulation and restmcturing oi" power industr\' are

aimed at isolating the supply of electrical energy from the service involving transmission

from generating station to load centers. This approach is feasible only if the operation of

ac transmission lines is made flexible by introducing fast acting high power solid state

controllers using thyristors or GTO valves. The advent of high voltage and high power

thyristor valves and digital conUoUers in HVDC transmission has demonstrated the

viability of deploying such controllers for power transmission [1-8].

The FACTS concepts involve the application of high power electronic controllers in ac

transmission network. This enables fast and reliable control of ])ower flov/s, voltage

profile and improves stability. FACTS controllers can enable a transmission line to carry

power closer to its thermal rating [1-8]. Another way to achieve the same goal is

simultaneous ac-dc power transmission in which the conductors are allowed to carr}'

superimposed dc current along with ac current. AC and DC power flow independently

and the added dc power flow does not cause any transient instability [14, III].

1.2 Literature Survey

In recent years, public has become more sensitive to the proliferation of overhead

transmission right-of-ways. Industrial countries are experiencing increasing difficulty in

finding suitable corridors for new overhead transmission lines and in many cases it is

simply impossible. There has been an increasing pressure to provide the substantial

power upgrading of existing ac transmission line corridors.

In references [12, 13] various ways are discussed for substantial power upgrading of

overhead transmission lines v ith changes in tower structures and insulator strings or total

conversion of ac line to dc line using the same right of way.

Hammad A E and Long W F [28] have done comparative study on two alternative

schemes for transmitting bulk power between differently managed or independent

regional ac networks. They have reported that conventional HVDC point to point

transmission system (alternative no.l) is more economical for transmitting 900 MW bi­

directional over distances more than approximately 400 km under certain assumptions as

compared to hybrid ac transmission with back-to-back HVDC links (alternative no.2) at

one end. Further the expansion or upgrading of both schemes is equally feasible but with

slightly lowers investment for alternative no.2.

A power system is in synchronous operation if all its connected machines are in

synchronous operation with the AC network and with each other when system is

subjected to disturbances of small or large magnitude. When synchronous operation of

EHV ac-dc transmission lines was thought of, an HVDC link was supposed to be

operating in parallel with ac link. This was done to enhance the stability of the AC link

by using fast converter control of the HVDC link [1-3,5-8,14,16,18,24,27,29-

30,33,34,40-44,49-55,57,58,61]. This way reactive power balance of the AC line can be

maintained either by transferring the excess reactive power to the HVDC link or by

taking the excess reactive power from the HVDC link using current control of rectifier.

When EHV ac lines covering long distances are planned, the reliability and stability

of the system become most important factor. In such cases it is expedient to add a dc

transmission network [1-3.5,6,14,16,18,27,33,40,50,52-55,61]. Advantage can be derived

from the fact that the power transmitted through dc link is independent of the angular

displacement between its ac terminal voltages and will essentially respond to what its

control demands. No stability problem arises in DC transmission system due to its

inherently fast controllability. Improvement of the stability characteristics of

interconnected power s\ stem by using fast acting converter control on HVDC links in the

system is well established. Uhlmann [8] has considered an idealized two- machine system

with parallel ac and dc links, which uses proportional and integral control using the

frequency signals. He has demonstrated a significant improvement in the stability. These

results were obtained from a linearised system analysis with simplifying assumptions.

However, subsequent work done by Peterson [52,53,55] and Dougherty et al [64]

confirmed the general validit>' of the results even tor large disturbances. The results were

based on a simplified model of the system using digital or analog simulation.

Hammad A E [16] has introduced a generic concept of combining the transient angle

and voltage stability of a parallel ac-dc transmission. The HVDC with classical dc power

control even with supplementary damping signal does not contribute to system

synchronizing torque and may increase the risk of instability. He has demonstrated the

utilization of the inherent short term overload capacity of the dc converter can be

maximized to increase the stability margin of the ac system and allow a higher power

transfer on the parallel ac transmission. Advanced HVDC large signal stabilizing control

strategies can be developed to produce large amount of synchronizing and damping

torques that can effectively stabilize the ac system and damp out all power oscillations on

the parallel ac transmission after faults. Such controls also optimize the use of the HVDC

short-term overioad capacity without need for additional reactive power support. The

increase in parallel ac transmission transient stability MW transfer limit can almost be

equal to the HVDC temporary overload. These features constitute large savings compared

to conventional solutions.

Reference [19] has introduced a new HVDC control concept of combined and

coordinated control method (CCCM) instead of conventional marginal current control

method (MCCM). The author has demonstrated improved dj'Tiamic behavior of the dc

transmission during fault recoveries and disturbance with CCCM.

To K W V et al [24] presented a philosophy based on modem control theories to

control the operation of HVDC transmission in parallel with the ac network with

emphasis on transient and dynamic stability. They developed a robust co-coordinated

control scheme by combining bang-bang and optimal control laws for HVDC in parallel

to the ac network. This control scheme applied on AC/DC power system and provides a

significant synchronizing torque as well as reduction in the peak of the first swings

without causing converter failure.

Smed T and Anderson G [27] have discussed the effect of HVDC on system damping.

The effect of real and reactive power modulations have been analytically studied. They

have shown that real power modulation, with frequency, as feedback signal is most

efficient when applied to a short mass scaled electrical distance to one of the participating

machines. The reactive power modulation, with derivative of the voltage magnitude as a

input signal, is most efficient when a well defined power flow direct exists and the

injection point is close to the electrical mid point between oscillator)' machines.

Routray A et al [32] has demonstrated the effectiveness of self-tuning parameters of PI

controllers for the rectifier side current regulator and inverter side gamma controller in a

point-to-point HVDC system. They have selected current error and its derivative and the

gamma error and its derivative as the principal signal to adjust the parameter of the

corresponding controller. They further made comparative study using EMTDC software

with and without timing to prove its superiority on the proposed point-to-point HVDC

system.

Hsu Y-Y and Wang L [33] discussed the improvement of the dynamic stability of a

parallel ac-dc transmission incorporating PID power system stabilizer and a PID rectifier

current regulator with generator speed deviation as input signal. The relative merits of

these two controllers are compared by eigen value analysis of the system. In order to

demonstrate the effectiveness of the controllers, dynamic response simulations are

performed on a digital computer.

Paper [38] pointed out the application of special (or supplementary) HVDC control for

power system dynamic performance improvements. The paper has discussed in detail on

reactive power regulation and dynamic ac voltage control at the HVDC temiinal,

damping of frequency oscillations and augmenting transient stability of dc power

modulation and special dc controls during faults. Trends and future possibilities are also

summarized in this report.

In reference [39], the tnmsient energy function method in transient stability is

extended to include an AC/DC system with two terminal HVDC. A simple do model is

developed and tested for first swing transient stability studies. The authors have discussed

a procedure for first swing trjinsient stability assessment of an AC/DC system and tested

on 20-generator test system.

Rahim A H M A et al [40] discussed the stabilization aspects of an AC/DC system

through the excitation of s>'nclironous generators and the converters firing angle controls.

The authors turther investigated the stabilization of a power system with these controls

through optimal/sub optimal strategies. A simple single machine system feeding power to

infinite bus via an AC/DC link is considered for evaluating the performance of control

strategies.

Vovos and Galanos [41] have discussed the active and reactive power modulations on

an integrated AC/DC system for the enhancement of transient stability. They have

presented a design of a DC link controller which modulates the active and reactive

powers of the converters in the event of major disturbance and improves the transient

stability of the integrated system.

Pai M A et al [44] have outlined a technique to apply the direct method of transient

stability analysis on a single and multi machine AC/DC system. A single and three

machines examples illustrated the validity of the method and the effect of the dc link in

improving the transient stability.

Vovos N A and Galanos G D [49] have studied tiansient stability of a two area power

system connected in two ways (i) two parallel ac links (ii) one ac line in parallel with dc

line by simulation. Both the configurations were subjected to the same disturbances, and

their responses were assessed and compared with regards to system stability.

Vovos N A and Galanos G D [50] have discussed the damping of power swings in ac

tie line using a parallel dc link operating at constant reactive power control utilizing the

fast electronic power flow characteristics of a DC link. The control strategy was tested

using dynamic solution techniques and the results obtained show that under certain

conditions, the contribution of a psjallel dc link to the stability of ac system is significant.

In the discussion of the paper Grund suggested more realistic models including

synchronous machines with governors and voltage regulators.

Klein et al [51] have discussed dynamic stability of a parallel ac-dc power system

incorporating excitation regulators, power system stabilizers (PSS) and DC rectifier

current regulator systems using eigen value analysis after developing small displacement

model of an ac-dc power system and verified with a detailed hybrid computer

representation of the system. Effect of AVR, PSS and rectifier current regulator was

qualatively discussed but the limits of regulator parameter were not studied. The study

concerned the dynamic performance on an ac-dc system including the interaction

between the controls associated machine and with dc links. This technique can be used to

readily incorporate the dynamic characteristic of a dc link into a digital program for

dynamic stability of an ac-dc power system.

Machinda T [53] discussed the improvement of transient stability of ac system by fast

increase in the reference current of the constant regulator at the rectifier station. The

experimental results were demonstrated on the artificial ac and dc transmission facilities.

He concluded that with increase in the dc power component in steady state compared to

the ac power, the transient stability improves.

For an ac-dc system, Klein et al [57] have discussed the effect of dc modulation on the

dynamic stability of (i) one machine, infinite bus and (ii) two machine, infinite bus

configurations.

To K W V and David AK [58] have pointed out that the conventional design of HV ac

-dc system controllers based on model linearization at operating point can not be

guaranteed for transient performance. They have suggested a multi input multi output

coordinated controller (MIMO) based on models of a parallel ac-dc power system derived

on line using identification techniques. The benefits of MIMO adoptive control scheme

have been examined. The digital simulation on the machines system has shown

substantial improvement in the transient stability with the proposed controller.

Dash et al [59] have presented the design of a simple fuzzy logic controller for HVDC

transmission links for fast stabilization of transient oscillations. They further compared

the results of proposed fuzzy controller with variable structure and adoptive controllers

for a variety of short circuits and system operating changes. The input signal to the

controllers is taken as generator speed deviation and its derivative. The output of signal

from controller is added with the current order of the rectifier current regulator. A variety

of transient simulations reveals that even a simple fijzzy controller based on a minimum

number of rules can produce excellent damping of the ac-dc system states in comparison

with other high performance controllers.

In the past, a number of schemes have been proposed for small power tapping from

HVDC lines. Majority of schemes use forced commutated or line commutated inverters

to tap off the power from the HVDC system. These schemes inherently required

additional commutation circuitry or local generation capacity, which in turn lead to high

cost of installations and operational complications.

Ekstrom and Lamell [67] have proposed a scheme of small power tapping from HVDC

line based on a current source line commutated single-phase thyristor bridge, connected

in series with the HVDC line. To start the tap operation a local dc voltage source is

required in this scheme.

Reference [68] provides transmission line planners with various factors which need to

be considered in evaluating the feasibilit}' of tapping in existing HVDC lines or in

developing alternatives for potential new transmission schemes.

In reference [14] simultaneous ac-dc power transmission was first proposed through

a single circuit ac transmission line to get advantage of parallel ac-dc power transmission.

In this scheme a mono-polar dc transmission system was proposed. But the mono-polar dc

transmission with ground as return path for dc current has limitations. No experimental or

simulation results were reported in this paper. Moreover, the instantaneous value of each

conductor voltage with respect to ground becomes more by the amount of the dc voltage

and more discs are to be added in each string insulator to wdthstand this increased voltage.

10

However there is no change in the conductor separation distance, as the line-to-line

voltage remains unaltered.

1.3 Objectives of the Thesis

From the literature surveyed above, following conclusions may be drawn:

Long EHV ac lines cannot be loaded to their thermal limits to keep sufficient margin

against transient instability. To keep sufficient stability margin, transmission angle

between two ends of ac long transmission line is generally kept low and seldom exceeds

30°. But for optimum use of transmission lines there is a need to load EHV ac lines close

to their thermal limits by using Flexible AC Transmission System (FACTS) components.

Very fast control of SCRs in FACTS devices like Static VAR System (SVS), Controlled

Series Capacitor (CSC), and Static Phase Shifter (SPS) and controlled Braking Resistors

improves stability and damps out oscillations in power system.

HVDC transmission lines in parallel with EHV ac lines are recommended to improve

transient and dynamic stability as well as to damp out oscillations in power system.

Alternatively the conductors of EHV ac line are allowed to carry superimposed dc

current along with ac current to achieve same goal. The added dc power component does

not cause any transient instability problem.

In this thesis simultaneous ac-dc power transmission is taken up for detailed study and

its various aspects are explored.

The main aims of the research thesis are:

• To demonstrate tlie feasibility of simultaneous ac dc power transmission

through same conductor for a double circuit transmission line.

• To demonstrate that the zig-zag connected transformer, provided at each end,

does not saturate due to injection of dc current in its neutral terminal. Digital

simulation study and its experimental verification on a laboratory-sized model

are carried out to validate the theory.

11

• To control the ac and dc component in simultaneous ac dc power transmission

system such that the conductor capacity is fully utilized and effective current

through it equals to its thermal limit.

• To obtain the subsiantial power upgrading of the existing EHV ac long line by

converting it to simultaneous ac-dc power transmission in such a way that no

alteration in the conductor size, insulator strings and tower structures of line is

required.

• To improve the transient stabilit}' of power system by dc power modulation in

case of simultaneous ac-dc power transmission in the same way as achieved by

parallel ac-dc transmission line.

• To explore economical, simple and easy method of small amount of power

tapping to feed remotely located communities that lie alongside the

simultaneous ac-dc power transmission line.

1.4 Outline of the Thesis

The thesis is logically divided into seven chapters.

Chapter 1 discusses the general problems of long distance power transmission

highlighting the advantages and limitations of EHV ac and HVDC power transmissions.

The available literature on the topic is reviewed thoroughly and critically. The objectives

of the present work are clearly identified.

Chapter 2 describes the digital modelling details of synchronous machine, transmission

line, CIGRE first benchmark HVDC model with its power circuit and associated controls

in PSCAD /EMTDC environment.

Basic concept of simultaneous ac-dc power transmission, highlighting other issues such

as preliminary economical consideration and possibilities of basic protection of the system

etc. are covered in this chapter. Power flow equation of simultaneous ac-dc power

transmission imder certain assumptions has been derived. The criterion for current and

voltage level selection for converted simultaneous ac-dc power transmission line has been

12

discussed. The latter part of the chapter demonstrates the feasibility test by simulation as

well as experimental verification of simultaneous ac-dc power transmission especially on

saturation effect of zig-zag connected transformer core due to injection of dc component.

Chapter 4 analyses and demonstrates in details the substantial power upgrading of an

existing double circuit EHV ac transmission line by converting it to simultaneous ac-dc

power transmission line. The results are supported by detailed computational and

simulation studies.

Chapter 5 deals with the improvement of transient stability of power system by dc

modulation in simultaneous ac-dc power transmission in detail. For the purpose of

comparison, the results are compared with identical line with pure ac transmission. The

effects of various types of faults created at different locations and their effects on system

performance with regard to transient stability have been studied.

Chapter 6 suggests an economical, simple and easy method of small power tapping

from the simultaneous ac-dc power transmission line for remotely located small

communities living alongside the line.

Chapter 7 includes main conclusion of the work and possibilities of future extension of

the present work.

13

CHAPTER 2

SYSTEM MODELLING

2.1 Introduction

The analysis, operation and design of complex ac-dc systems require extensive

simulation resources that are accurate and reliable. Analog simulators, long used for

studying such systems, have reached their physical limits due to the increasing

complexity of modem systems. Currently, there are several industrial grade digital time-

domain simulation tools available for modeling ac-dc power systems. Among them, some

have the added advantages of dealing with power electronics apparatus and controls with

more accuracy and efficiency. PSCAD/EMTDC and PSB/SIMULINK are such two

simulators that are being increasingly used in the industry as well as in the universities.

Both programs allow the user to construct schematic diagram of electrical networks, run

the simulation, and produce the results in a user-friendly graphical environment.

Furthermore, several real-time digital simulators use models or the graphical front-end

that are similar to PSCAD/EMTDC and PSB/SIMULINK.

PSCAD/EMTDC is a powerful time-domain transient simulator for simulating power

systems and its controls. It uses graphical user interface to sketch virtually any electrical

equipment and provide a fast and flexible solution. PSCAD/EMTDC represents and

solves the differential equations of the entire power system including both elec­

tromagnetic and electromechanical systems and its control in the time domain [10]. It

employs the well-known nodal analysis technique together with trapezoidal integration

rule with fixed integration time-step. It also uses interpolation technique with

instantaneous switching to represent the structural changes of the system.

PSCAD/EMTDC is an industry standard simulation tool for studying the transient

behavior of electrical networks. Its graphical user interface enables all aspects of the

simulation to be conducted within a single integrated environment including circuit

assembly, run-time control, analysis of results, and reporting. Its comprehensive library

of models supports most ac and dc power plant components and controls in such a way

14

that FACTS, custom power, and HVDC systems can be modelled with speed and

precision. It provides a powerful resource for assessing the impact of new power

technologies in the power network.

Simplicity of use is one of the outstanding features of PSCAD/EMTDC. It has great

many modelling capabilities and highly complex algorithms and methods are transparent

to the user, leaving him free to concentrate his efforts on the analysis of results rather

than on mathematical modelling. For the purpose of system assembling, the user can

either use the large base of built-in components available in PSCAD/EMTDC or to its

own user-defined models.

As the idea of simultaneous ac-dc transmission is a new concept, the well-recognized

PSCAD/EMTDC package has been employed for the study. The system components are

modelled in PSCAD/EMTDC as suggested in references [1,7,10].

2.2 Salient Pole / Round Rotor Synchronous Machine Modelling in EMTD /PSCAD

The modelling of synchronous machine is based on Park's transformation from phase

to dqo quantities. The general equivalent circuit for the synchronous machine is as shown

in Fig. 2.1. A second damper winding on the q-axis is included in this model and hence it

can also be used as a round rotor machine to model steam tiirbine generators. This model

can also be suitable for studying sub-synchronous resonance (SSR) problems as well.

This model operates in the generator mode so that a positive real power indicates

electrical power leaving the machine, and a positive mechanical torque indicates

mechanical power entering the machine. A positive reactive power indicates that the

machine is supplying reactive power indicating it is overexcited.

The generalized machine model transforms the stator windings into equivalent

commutator windings, using the dqO transformation as follows:

cos(fl) cos(9-120") cosfe-240-) sin{fl) sin[8-120') sin{9-240')

y2 1/2 V2 (2.1)

15

Supporting subroutines are included in the machine model library for calculating the

equivalent circuit parameters of a synchronous machine from commonly supplied data.

The d-axis equivalent circuit for the generalized machine is shown in Fig. 2.2. and

Fig. 2.3 illustrates the flux paths associated with various d-axis inductances.

By inspecting Fig. 2.2 and 2.3, the following equations can be written:

C1 - f ¥ q - R, Ip,

U|>2 - R2D ID2

U03 - R30 ID3

= Lo d

'clt

let

ID3 (2.2)

Where,

U © ^M[) * I-23D * '-2Q Lt^o + 1230

^ q " Li IQI + L rtQ (IQI + IQ2 * iQIi)

d^

dt

(2.3)

(2.4)

(2.5)

Similar equations can be \\Titten for the q-axis except the speed voltage term, o4'j, is

positive,

and:

"^d = L , ID, * I - M D ('DI * ' D 2 ^ I D B ) (2.6)

Inversion of equation (2.2) gives the standard state variable form X = AX + BU with

state vector X consist of the currents, and the input vector U that of applied voltages.

That is:

d dt

d dt

I D I '

'D2

ID3

I Q I '

IQ2

.'<^3.

= L,-'

= k-'

'-V Y q - R t 'Dl

- R2D ID2

- R3D ID3

'^ Y d - R i iQi'

~ '^20 l<52

" R X> IQ3

• ^u - '

•

* ' - Q - '

U c

UD2

UD3

UQ,'

JQ2

UQ3

(2.7)

(2.8)

16

+ la

V. ^

^c

+ ^c

^ ,

Stator Windings

'^ Equivalent Commutator

Windinqs

?

Fig. 2.1. Conceptual Diagram of the Three-Phase and d q Windings

Note: All quantities shown in Fig. 2.1 are in per unit.

Where,

k- Amortisseur windings

f- Field windings

abc- Stator windings

d- Direct-Axis (d-axis) windings

q- Quadrature-A.xis (q-axis) windings

17

I D I ^-

Rl Li

U D1

L23D

-MD

lD2

•-20 '2D

AAAr iD5

•5D ^30

•VAr u D3

U D2

Fig. 2.2. d-axis equivalent circuit

-MD

n

23D

-3D 2D

n.

Stator Air Gap Amortisseur Field

Fig. 2.3. Flux paths associated with various d-axis inductances

In the above form, equations (2.7) and (2.8) are particularly easy to integrate. The

equations are solved using trapezoidal integration to obtain the currents. The torque

equation is given a :

T = ^q ioi " 'Pd'QI (2.9)

and the mechanical dynamic equation for motor operation is:

dy • T - V c H - P ' ^

dt J (2.10)

Referring to Fig.2.2 and Fig.2.3, the d-axis voltage UD2 and current ID2 ai-e the field

voltage and currait, respectively. The damper circuit consists of parameters LSD and R3D

18

with UD3 =0. The additional inductance L23D accounts for the mutual flux, which links

only the damper and field windings and not the stator winding. The inclusion of such

flux is necessary for accurate representation of transient currents in the rotor circuits.

The saturation effect only on LMD is considered in this machine and is based on the

magnetizing current,

(2.11) 'MD " 'D1 * 'D2 * 'D3

The matrix ^ is recalculated each time there is a change in the saturation factor.

2.2.1 The Per-Unit System

The per-unit system is based on the preferred system indicated in [1,10]. The base

value of voltage and current in the three-phase system is the RMS phase voltage Vao, and

RMS phase current iao- The same voltage base is used in the dqO system but the base

current is changed to 3/2 iao- The dqO transformation in per unit for the voltage or current

is given in equation (2.1). The inverse transform is:

cos(e) sin((?) f

cosfe-120") sinfe-120') 1

cos((?-240") sin((9-240") 1 (2.12)

Note that unit current in the d-winding produces the same total MMF as unit currents

acting in a balanced fashion in the abc-windings. Unit currents in the different circuits

should produce the same physical effect. In both systems base power is:

(2.13) Po = 3 V ^ o i ^

The associated bases are:

° = Rated frequency

1

L = x r/

Vl/ _ ' oO " 0 ~

fi>o

in radians

Base flux linkage

v„ = • <»o

9 =

PolePairs Base mechanical Speed

PolePairs Mechanical angle

The impedances are given in per unit. The input voltages are divided by Vao and the

incremental time At is multiplied by co^ to provide a per unit incremental time. The per-

unit current is converted to output current by multiplying it by iao after transformation

from the two-axis system.

Care should be taken with the following quantities:

Utilities often specify one per unit field current and voltages as that which produces

rated open circuit voltage on the air-gap line (this implies unit power loss in the field

circuit). The per unit field curren: iD2, must be multiplied by XMDO and then divided by

V2 in order to convert it to the value of field current used in the utilit>' system. The value

of the field voltage is multiplied byRp/X^^oo to give the correct per unit value of UD2

2.2.2 Machine Interface to EMTDC

The machine models represent the machine as a Norton cuirent source into the

EMTDC network. This approach uses the terminal voltages to calculate the currents to

be injected.

Fig. 2.4 shows a synchronous machine model interfaced to the EMTDC program. The

synchronous machine model makes use of the phase voltages calculated by EMTDC to

update the injected currents into EMTDC. It is also shown in this figure that multiplying

the phase currents by an integer N simulates (fi-om the system point of view) N identical

machines operating coherently into the AC system.

Va Vb

EMTDC Network Solution

P Q if

1 \

V

f f - t

dqO

Machine Equations

Uq. id

V,

>..

V

• Exciter / iviodei

'ret —'

dqO

' M CO

Governor ivlodel

"ref - •

la

Fig. 2.4. Synchronous machine model interface to EMTDC.

20

Terminating Resistance

It is impoitant to note that representing machines as a Norton current source can have

drawbacks. For instance, each machine must be computationally 'far' irom other

machines for stable operation. In the past, this was usually achieved by separating

subsystems containing machines by distributed transmission lines (which are essentially

time delays). Since the machine was represented by a simple current source (which was

dependent on voltages from the previous time step), any sudden change in voltage would

cause a current response only in the next time step. Thus, for the previous time step, the

machine looked like an open circuit and spurious spikes appeared on the machine

terminal voltage. The cumulative effect of many machines causing this error

simultaneously in the same subsystem was proven to be de-stabilizing (computationally).

It was found that when the machine neared open circuit conditions a smaller time step

was required to maintain computational stability. Alternatively a small capacitance or

large resistance could be placed at the machine terminals to ground to prevent the

machine from being totally open-circuited. Although the physical meaning of parasitic

21

capacitance or leakage resistance could be applied to these elements, it was not a

satisfactory solution.

This idea led to the concept of terminating the machine to the network through a

terminating 'characteristic impedance' as shown in Fig. 2.5. The effect of this added

impedance is then compensated (corrected) by injecting appropriate current adjacent to it.

Using this technique, the machine model behavior has been uniformly good. It

essentially combines the compensation-based model and the non-compensated model,

while eliminating the restriction of adjacent machines and the necessity of calculating the

network Thevenin equivalent circuit.

0 ic(t) 0 '-(*

1 'ma

Z"

±

AC System

EMTDC Network

Fig. 2.5. Interface with terminating resistance.

Where,

lc(t) = Vc(t-At)

Z" Compensating current

Z"=

Calculated machine current

2L"

^ • ^^ Terminating 'characteristic impedance'

The impedance Z" is calculated where L" is the 'characteristic inductance' of the

machine, N is the number of coherent machines in parallel and At is the EMTDC time

step. This resistance is placed from each node of the machine terminal to ground within

the EMTDC network. Instead of injecting the calculated machine current Im(t), a

11

compensated current I„(t) + I^(t)is injected, where V(t - At) is the terminal voltage in

the previous time-step. Thus, the actual current injected into the network is,

Z" is usually quite large, due to the At in the denominator. Also, for a small time

step V(t - At) = V(t), and thus In,a (t) = !„. (0 •> and the error introduced vanishes in the

limit with a small At. However, for a sudden voltage change, as I^(t) + I^(t)is not

calculated until the next time step, the network sees the impedance Z" for this instant

(instead of the open circuit discussed earlier). This is exactly the instantaneous

impedance it would have seen, had the machine been represented in the EMTDC

program main matrix. Therefore, the network current calculated in this instant is

more accurate, and the spurious spikes discii<?<;ed earlier do not arise. Thus, this

concept of terminating the machine with its 'characteristic impedance and then

compensating it by current injection, is a convenient way for assuring accurate

solutions.

2.2.3 Mechanical and Electrical Control

As shown in Fig.2.4, control blocks used to simulate the excitation and governor

systems of the synchronous machine need also to be modelled. These control systems are

not included internally in the machine model, so they must be interfaced through external

connections.

The exciter and governor systems can either be built using standard control system

building blocks (available in the CSMF section of the PSCAD Master Library) or

standard exciter and governor models available may be used. Following sections briefly

describe the general theory behind them.

23

2.2.3.1 Exciters

Exciters are externally interfaced to the machine models through external signal

connections. Since the exciter models exist as separate components, parameters can be

freely selected and adjusted by the user.

Reference [1], published by the IEEE, categorizes exciters into three different types:

(i) Type DC excitation systems utilize a DC generator with a commutator as the source

of excitation system power

(ii) Type AC excitation systems use an alternator, and either stationary or rotating

rectifiers, to produce the direct current needed for the synchronous machine field

(iii) Type ST excitation systems, where excitation power is supplied through

transformers or auxiliary generator windings and rectifiers

There are many different exciter models classified in one of the above three tj'pes,

which are representative of commercially available exciters. PSCAD comes complete

with each of these standard models.

One problem with modelling exciter systems is the inherently large time-constants

involved. The user should ensure that the system is started as close to the steady state as

possible (unless of course, simulation of start-up transients is required), by properly

setting the machine initial conditions. It is also possible, in the case where the exciter

transfer function has large time constants, to bypass these during the initial phase.

2.2.3.2 Governors

Mechanical transients usually fall in the realm of simulation using stability programs,

where the details of modelling is not as comprehensive as in an electromagnetic transient

program. For the duration of most transient simulation runs, the mechanical system can

often be considered invariant, and users should make sure that they really need a

governor model before they use it.

As mentioned above, most governor transfer functions can be simulated by using

control building blocks; interfacing to the machine by either feeding the computed speed

or the computed torque.

24

Governor models are initialized by continuously resetting internal storage locations to

produce an output that is exactly equal to the mechanical torque output of the machine.

PSCAD comes complete with both hydro and thermal governor models. These models

support variable time step, and have an allowance for initialization at any time through

additional control inputs.

2.2.3.3 Stabilizers

Power system stabilizers are used to enhance the damping of po\ 'er system

oscillations, through excitation control. Conmionly used inputs to the stabilizers are:

1. Shaft speed

2. Terminal frequency

3. Power

Most stabilizer transfer ftmctions can be simulated through the use of control building

blocks. PSCAD comes complete with both single-input and dual-input, po\ver system

stabilizers.

2.2.3.4 Turbines

Mechanical power supplied by turbines can often be considered invariant for the

duration of most transient simulation runs. In some cases however, where provisions are

made for fast-valving or discrete control in response to acceleration, prime mover effects

can be significant even if the phenomena of interest span only for a few seconds. Also, to

ensure the accuracy of longer simulation studies, modelling the turbine dynamics may be

considered necessary.

These models support variable time step, and have an allowance for initialization at

any time through additional control inputs. Most turbine dynamics can be simulated

through the use of control building blocks. PSCAD comes complete with both thermal

and hydraulic turbines.

25

2.3 Transmission Line Model in PSCAD

Overhead transmission line corridor (right-of-ways) are represented in PSCAD as two

main parts: By defining the configuration of the transmission corridor itself, where the

definition can include either the admittance/impedance data or the conductor and

insulation properties, ground impedance data, and geometric position of all conductors

within the corridor. This definition is then interfaced to the rest of the electrical system

through electrical interface components (one at each end).

Three conductor transmission systems of very short length (i.e. less than 15 km for a

50 |is time step) can also be represented using an equivalent PI Section. This is

accomplished through a Master Library component, called the PI Section, where only the

admittance and impedance data of the line segment is entered.

A 15 km line length to a 50 ^s time step is derived assuming that waves propagate

through the line at the speed of light. In general however, wave propagation speed is less

than the speed of light.

Using the data provided by the cross-sectional definition of the corridor, the

transmission lines and cables are modelled using one of three distributed (travelling

wave) models:

1. Bergeron Model

2. Frequency-Dependent (Mode) Model

3. Frequency-Dependent (Phase) Model

The most accurate of these is the Frequency Dependent (Phase) model, which

represents all firequency dependent effects of a transmission line, and should be used

whenever in doubt. When using the Bergeron model, impedance/admittance data can

also be entered directly to define the transmission corridor.

For all of these Frequency Dependent models, detailed conductor information (i.e.

line geometry, conductor radius) must be given.

26

2.3.1 Selecting the Proper Line l\/lodei

There are three types of distributed (i.e. travelling wave) transmission models that may

be selected to represent transmission corridor: The Bergeron model, the Frequency-

Dependent (Mode) Model, and the Frequency-Dependent (Phase) Model. These models

exist as components in the Master Librar>-, and each include adjustable properties. The

requirements for thie specific study will detemiine which of the three models is suitable.

(a) Tlie Bergeron IVIodel

The Bergeron model represents the L and C elements of a PI section in a distributed

manner. It is roughly equivalent to using an infinite number of PI sections except that the

resistance is lumped (i.e. Vi in the middle of the line and '/4 at each end). Like PI sections,

it also accurately represents the fundamental frequency. It also represents impedances at

other frequencies, except that the losses do not change. This model is suitable for studies

where the fundamental frequency load-flow is most important (e.g. relay studies). This

model can be;

(1) Single Conductor Model

(2) Multi- Conductor Model

When using the Bergeron model, it is not always necessary to use a tower component

to represent a transmission line. In case of modelling a three-phase system the line data

can be entered, in admittance or impedance format, directly by substituting the manual

entry of Y,Z component.

(b)The Frequency-Dependent (IVIode) IVIodei

The Frequency-Dependent (Mode) Model represents the fi-equency dependence of all

parameters (not just at the specified frequency as in the Bergeron model). This model

uses modal techniques to solve the line constants and assumes a constant transformation.

It is therefore only accurate for systems of ideally transposed conductors (or 2 conductor

horizontal configurations) or single conductors.

27

(c) The Frequency-Dependent (Phase) Model

The Frequency-Dependent (Phase) Model also represents the frequency dependence of

all parameters as in the 'Mode' model above. However, the Frequency Dependent (Phase)

model circumvents the constant transformation problem by direct formulation in the

phase domain. It is therefore accurate for all transmission configurations, including

unbalanced line geometry.

The need for a transmission line model, which can accurately simulate the undesirable

interactions between DC and AC lines in proxim.ity to one another, has become more

prevalent. Constant transformation matrix models with frequency dependent modes, such

as the Frequency Dependent (Mode) model in EMTDC, have proven to be unreliable in

accurately simulating such situations. In addition, inaccurate representation of

unbalanced line geometry has also been a problem.

The Frequency-Dependent (Phase) model should always be the model of choice,

unless another model is chosen for a specific reason. This model is the most advanced

and acciirate time domain line model in the world.

2.4 HVDC Model

2.4.1 Six- Pulse Converter Model Block in PSCAD

PSCAD/EMTDC provides as a single component a six-pulse valve group, as shown in

Fig. 2.11 with associated Phase Locked Oscillator (PLO) firing control, sequencing drop

and parallel snubber circuit.

The Fig. 2.6 component is a compact representation of a DC converter, which includes

a built in 6-pulse Graetz converter bridge (can be inverter or rectifier), an internal Phase

Locked Oscillator (PLO), firing and valve blocking controls and firing angle

(a)/extinction angle (y) measurements. It also includes built in RC snubber circuits for

each thyristor. The general idea here is to eliminate the complex and tedious process of

building a thyristor bridge, putting together controls, and coordinating the valve firing

pulses involved in modeling an HVDC converter.

28

ac input B

_C.

si^Cc

h * .

+ )mBus

? ^^ - ^ ^

'? P-4L

J_

AM

AO

KB

Fig. 2.6. PSCAD compact model of Graetz Bridge.

The 6-pulse Bridge possesses the following external inputs and outputs:

ComBus: Input signal to the internal Phase Locked Oscillator. This input is

connected to the commutation bus though the Node Loop component.

AO: Input alpha order (firing angle) for the converter.

KB: Input block/deblock control signal

AM: Measured alpha (firing angle) output [radian].

GM: Measured gamma (extinction angle) output [radian].

The Phase Locked Oscillator (PLO) shown in Fig.2.7 is based on the Phase Vector

technique. This technique exploits trigonometric multiplication identities to fonn an error

signal, which speeds up or slows down the PLO in order to match the phase. The output

signal 9 is a ramp synchronized to the Phase A commutating bus L-G voltage.

The block diagram of the PLO used in the 6-Pulse bridge component is shown in Fig.

2.7.

29

I 1

\ : / ""

GP

0)Q —

Gl s

If 4 Vsine

V sin 8

Reset ® 2x

1.2 1_0 s

0.8

-9 - 9

Fig. 2.7. Phase vector type phase-locked oscillator.

2.4.2 CIGRE HVDC benchmark in PSCAD

The full three-phase model of the CIGRE HVDC benchmark system is available in

reference [7,10].

(A) Power Circuit Modelling

1) Converter Model:

The converters (rectifier and inverter) are modeled using six-pulse Graetz bridge

block, shown in Fig. 2.6, which includes an internal Phase Locked Oscillator (PLO),

firing and valve blocking controls, and firing angle (or)/extinction angle {y)

measurements. It also includes built-in RC snubber circuits for each thjTistor. Thyristor

valves are modeled as ideal devices, and therefore, negative turn-off and firing due to

large (dv/dt) or (di/dt) are not considered.

2) Converter Transformer Model:

Two transformers on the rectifier side are modeled by three-phase two winding trans­

former, one with grounded Wye-Wye coimection and the other with grounded Wye-

Delta coimection. The model uses saturation characteristic. The inverter side transformers

use a similar model.

30

3) Filters and Reactive Support: Tuned filters and reactive support are i)rovided at

both the rectifier and the inverter ac sides.

(B) Control System Model

The control model mainly consists of aly measurements and generation of firing

signals for both the rectifier and inverter. The PLO, shown in Fig. 2.7, is used to build the

firing signals. The output signal of the PLO is a ramp, synchronized to the phase-A

commutating bus line-to-groimd voltage, which is used to generate the firing signal for

Valve 1. The ramps for other valves are generated by adding 60° to the Valve 1 ramp.

This process is illustrated in Fig. 2.8. As a result, an equidistant pulse is realized. The

actual firing time is calculated by comparing the a order to the value of the ramp and

using interpolation [7,10] technique. At the same time, if the valve is pulsed but its

voltage is still less than the forward voltage drop, this model has a logic to delay firing

until the voltage is exactly equal to the forward voltage drop. The firing pulse is

maintained across each valve for 120°. The a and / measurement circuits; use zero-

crossing information from commutating bus voltages and valve switching times and then

convert this time difference to an angle (using measured PLO frequency). Firing angle a

(in seconds) is the time when vah'e i turns on minus the zero crossing time fDr valve i.

Extinction angle ;' (in seconds) for valve i is the time at which the commutation bus

voltage for valve i crosses zero (negative to positive) minus the time valve i turns off

The control schemes for both rectifier and inverter of the CIGRE HVDC system are

shown in Fig. 2.9 and Fig. 2.10.

Following are the controllers used in the control schemes:

• Extinction Angle (/Controller),

• DC Current Controller;

• Voltage Dependent Current Order Limiter (VDCOL).

31

Interpolated firing of valve 1

Interpolated firing of valve 2

Fig. 2.8. Firing control for the PSCAD/EMTDC valve group model.

(i) Rectifier Control:

The rectifier control system uses Constant Current Control (CCC) technique. The

reference for current limit is obtained from the inverter side. This is done to ensure the

protection of the converter in situations when inverter side does not have sufficient dc

voltage support (due to a fault) or does not have sufficient load requirement (load

rejection). The reference current used in rectifier control depends on the dc voltage

available at the inverter side. DC current on the rectifier side is measured using proper

transducers and passed through necessary filters before they are compared to produce the

error signal. The error signal is then passed through a PI controller, which produces the

necessary firing angle order a . Fig. 2.9 shows rectifier control scheme.

(ii) Inverter Control:

The Extinction Angle Control or y control and current control have been

implemented on the inverter side. The CCC with Voltage Dependent Ciarrent Order

Limiter (VDCOL) has been used here through PI controllers. The reference limit for the

current control is obtained through a comparison of the external reference (selected by

32

the operator or load requirement) and VDCOL (implemented through lookup table)

output. The measured current is then subtracted from the reference limit to produce an

error signal that is sent to the PI controller to produce the required angle order. The /

control uses another PI controller to produce gamma angle order for the inverter. The two

angle orders are compared, and the minimum of the two is used to calculate the firing

instant. Fig. 2.10 shows the detailed control scheme at inverter.

Angle Order for rectifier

AOR

+ - D 4 ^ AR

P ' CERRR

O

Current (filtered)

CMRS <r

F

^L CORDER.

dc current measured at rectifier

G z 0 1 + sT ^ CMR

current order from inverter

Fig. 2.9. CIGRJE Rectifier control.

33

o o

CO Q : _ I LJJ O fc: or uj 5

i

T 4

u-

<1>

' 1 -O E

<

en

<?=

a>

Fig. 2.10. CIGRE Inverter control scheme.

34

2.5 Transformer Modelling [7,10]

Transformers are represented in EMTDC through one of two fundamental methods:

The qlassical approach and the unified magnetic equivalent circuit (UMEC) approach.

The classical approach should be used to model windings placed on the same

transformer leg. That is, each phase is a separate, single-phase transformer with no

interaction between phases. The UMEC method takes inter-phase interactions into

accoimt: Thus, 3-phase, 3-limb and 3-phase, 5-limb transformer configurations can be

accurately modeled.

Representation of core non-linearities is fimdamentally different in each model type.

Core saturation in the classical model is controlled through the use of a compensating

current source injection across selected winding terminals. The UMEC approach uses a

fully interpolated, piecewise linear O-I ciir\'e to represent saturation.

Zig-zag connected transformers are modelled by these two models separately using (i)

classical approach (ii) the UMEC (Unified Magnetic Equivalent Circuit) approach from

three single phase, three windings, with their primary and secondary windings connected

in delta and zig-zag fashion respectively as shown in Fig. 2.11a and Fig. 2.11b. This

Delta - Star Zig Zag connection gives a zero degree phase shift between the primary and

secondary voltages. The connections are determined based on a simple phasor based

analysis.

35

SA

SB

SC

#1

#1

#1

j - #2 , L - - • - • - -

- *3

: * 2 •

- P3

- #2 ,

- #3

Cj

5.

c

SN 1

Fig. 2.1 la. Connection diagram of Delta- zig-zag transformers (Classical Model).

Vb

Vc

Va

u m s c

Id/3 a1 a2

" 1 - , # 2

. ' L ^ .

*t 1

u m e c

# 1

# 3 3 3 Id/3 a4

u r n 9 c c ** 2 t Id/3 b1

b?

I i b^ L

j Id/3 M

Jd /3 c1

ii * 3 o3 ^

Id/3 *^

Va4

Vb4

.Vc4

Aid

Fis. 2. lib. Connection diagram of Delta- zig-zag transformers (UMEC) Model.

36

CHAPTER 3

SIMULTANEOUS AC-DC POWER TRANSMISSION

3.1 Introduction

In the present power systems, vast majority of power transmission lines are of ac type

with few HVDC links. Power flov in ac lines is automatically determined by Kirchoff s

laws. This is in contrast to HVDC links where power flow is regulated by converter

controls. Piare ac lines have following limitations [1-6].

• Power transmission through ac long lines is limited by stability considerations.

This implies that the lines operate at power levels much below their thermal

limits.

The lack of fast controls in ac line necessitates the normal power flow in lines

to be kept even below its maximum permissible value, which itself is limited by

stabilit)'. This margin is required to maintain system securit>' even under

contingency conditions.

In long EHV ac lines, the production and consumption of reactive power by the

line itself constitutes a serious problem. The ac transmission system requires

dynamic reactive power control to maintain satisfactory voltage profile under

varying load conditions and transient disturbances.

The most economical load on an overhead ac line is usually greater than the

natural load. However, when load is increased beyond the natural load, net

reactive power is consumed by the line and must be supplied from one or both

ends. Thus, the increase in load levels is accompanied by higher reactive power

consumption in the line reactance.

Long-distance ac power transmission is feasible only with the use of series and

shunt compensation, applied at certain intervals along the line. Series

compensation of degree k reduces the effective inductance from L by an

37

amount kL leading to its effective value as (l-k)L and thus decreases the

electrical length from pi to pi y (1 -k) .At the same time it also decreases the

surge impedance and increases the natural load by the same factor. The reactive

power produced by shunt capacitance of the line at light load may still be

excessive, requiring shunt compensation of part h of it. The effective shunt

capacitance is then reduced from C to (l-h)C, and the electrical length is

reduced by the factor /(l - h) or by the total factor ^(1 - k){\ - h). The surge

impedance is altered by the factor •yj{\-k)l{\-h) and may be essentially

unchanged if h=k. The compensation is chosen to limit the transmission angle

to 30" and to limit the voltages at both ends such that at compensation point it is

not more than 1.05 times nominal voltage [2].

Because of above-mentioned limitations of ac transmission, faster dynamic controls are

required to overcome the problem associated with ac transmission system.

Recent developments involving deregulation and restructuring of power industry are

aimed at isolating the supply of electrical energy from the service involving transmission

between generating station to load centers. This approach is feasible only if the operation

of ac transmission lines is made flexible by introducing fast acting high power solid state

controllers using thyristors or GTO valves. The advent of high voltage and high power

thyristor valves and digital controllers in HVDC transmission has demonstrated the

viability of deploying such controllers for power transmission. HVDC transmission lines

in parallel with EHV ac lines are recommended to improve transient and dynamic stability

as well as to damp out oscillations in power system [1-9,14,27,33,34,40,41.43,44,49-

55,61].

The FACTS concepts involve the application of high power electronic controllers in ac

transmission network. This enables fast and reliable control of power flows, voltage

profile and imjHoves stabilit}'. FACTS controllers enable the transmission line to carry

power closer to its thermal rating [1-7,1531]- Another alternative concept proposed

recently to achieve the same goal is simultaneous ac-dc power transmission in which the

38

conductors are allowed to carry superimposed dc current along with ac current [14,111].

AC and DC powers flow independently and the added dc power component does not

cause any transient instability.

In this chapter the simultaneous ac-dc power transmission through a double circuit ac

line is described and various other issues involved are highlighted.

3.2 Basic Concept

In simultaneous ac-dc power transmission system, the conductors are allowed to carry

dc current superimposed on ac current. . AC and DC power flow independently and the

added dc power flow does not cause any transient instability. The network in Fig. 3.1

shows the basic scheme for simultaneous ac-dc power flow through a double circuit ac

transmission line. The dc power is obtained by converting a part of ac using line

commutated 12-pulse rectifier bridge as used in conventional HVDC and injected to the

neutral point of the zig-zag connected secondary windings of sending end transformer.

The same is reconverted to ac by the conventional line conmiutated 12-pulse inverter at

the receiving end. The inverter bridge is connected to the neutral of zig-zag coimected

winding of the receiving end transformer. The line conductors are cormected between the

zig-zag secondary windings of the transformer at both ends. The converted double circuit

ac transmission line carriers both 3-phase ac as well as dc power. Each conductor of a line

carries one third of the total dc current along with ac current. Resistance being equal in all

the three phases of secondary winding of zig-zag transformer as well as the three

conductors of the line, the dc current is equally divided among all the three phases. The

three conductors of the second line provide return path for the dc current. Thus the

converted double circuit transmission line carriers both ac as well as dc power.

Zig-zag connection of secondary windings of transformer at both end is used to avoid

saturation of core due-to dc current. Two fluxes produced by the dc current (Id/3) flowing

through each half of a winding in each limb of the core of the transformer are equal in

magnitude and opposite in direction. So the net dc flux at any instant of time becomes zero

39

"[X

^.^

<3

-m

^m

I — i ^ ^ t > t ^ vwv^

s ^s

AJTpS i I I

7*0 >

^

1 • a

g 3

>-

f

o (_> ,g)

H "

o CI

(r) :) o t) r

£

O 51 «

CD

40

in each limb of the core. Thus the dc saturation of the core is avoided. A high value of

reactor Xd is used to reduce harmonics in dc current.

In the absence of zero sequence and third harmonics or its multiple harmonic voltages,

under normal operating conditions, the ac current flow through each transmission line will

be restricted between the zigzag connected windings and the three conductors of the

transmission line. Even the presence of these components of voltages may only be able to

produce negligible current through the ground due to high value of smoothing reactor Xd.

For a single circuit ac transmission line, the return path for dc current would be ground.

3.3 Analysis

Assuming the usual constant current control of rectifier and constant extinction angle

control of inverter [1-3,7,8,61], the equivalent circuit of the scheme (Fig. 3.1) under

normal steady state operating condition is given in Fig. 3.2. The dotted lines in the figure

show the path of ac retum current only. The second transmission line carries the return dc

current Id and each conductor of the line carries dc current /<y/3 along with the ac current

per phase Ig.

Vdro and Vdio are the niaximum values of rectifier and inverter side dc voltages may and

be expressed as:

3V2 V = BTE

71

Where,

E = RMS line-to-line voltage of converter transformer primary voltage.

T = Converter tiansformer turn ratio.

B = Number of six-pulse bridges in series.

Es, ER are the sending end and receiving end ac voltages per phase respectively. R, L, C

are the line parameters per phase of each line. Rcr, R<;j are commutating resistances and a, 7

are firing and extinction angles of rectifier and inverter respectively.

41

V Cos a —4-M'\H

I V

'^/v->-- \l Cos y

V Cos a

R

T 11

i 0 Xg

1 + 1 / 3 s d

C/2

L

_nnrr\.

- R

V Cos y

- - c/2 r ^9

1

R

•AAAr

I +1 / 3

Fig. 3.2. Equivalent Circuit.

42

Neglecting the resistive drops in the line conductors and transformer windings due to

dc current, expressions for ac voltage and current, and for active and reactive powers in

terms of A, B, C, D parameters of each line may be written as [9]:

I R = ( E S - E R ) / B (3.1)

Es = AER + BIR (3.2)

IS = CER + DIR (3.3)

Ps + jQs = -ESER*/B* + D*Es /B* (3.4)

P R + J Q R = ES*ER/B*-A*ER'/B* (3.5)

Neglecting ac resistive drop in the line and transformer,

the direct current flowing from rectifier to the inverter may be given as [1];

^ ^ ^ V , , c o s a - V , , c o s r ^^^^

^cr +Re<i ~ ^ c i

Where, R«q is the equivalent resistance of the line conductors seen by dc.

The dc power Pdr and Pdi of each rectifier and inverter may be expressed as:

Pdr = VdrId (3.7)

Pdi^Vdild (3.8)

Reactive powers required by the converters are:

Qdr = Pdrtan0r (3.9)

Qdi = Pditan0i (3.10)

cos9r = [cosa + cos(a + \IT)V2 (3.11)

cos9i = [cosy + cos(y + Hi)]/2 (3.12)

Hi and Hr are commutation angles of inverter and rectifier respectively and total active

and reactive powers at the two ends are:

Pst =Ps + Pdr and Pr, = PR + Pni (3.13)

43

Qst =Qs +Qdrand Qrt= QR + Qdi

Transmission loss for each line is:

PL = (PS + Pdr)-(PR + Pdi)

(3.14)

(3.15)

la being the rms ac current per conductor and Id/3 being superimposed dc current per

conductor of the line, the total rms current per conductor becomes:

1 =

1 -

In

In j(i,+i,/3yd«t

J 2 . 1 '

— |(l„sin,yt + l,/3)-d6;t In

After simplifications

I = [Ia' + (Id/3)']"'; (3.16)

Power loss for each line = PL ~ 31'R. (3.17)

The net current I in any conductor is offset from zero. In case of a fault in the

transmission system, gate signals to all the SCRs of rectifier and inverter are blocked and

that to the bypass SCRs are released to protect the rectifier and inverter bridges. The

ciarrent in any conductor is no more offset as dc component is blocked. Circuit breakers

(CBs), if provided on zig-zag secondary side are then tripped at both ends to isolate the

faulty line. CBs connected at the two ends of transmission line thus interrupt the current at

natural current zeroes and no special dc CBs are required.

Now if the conductor is allowed to carry net current through it equal to its thermal limit

(Ith):

Ith= [Ia' + (Id/3)']"^ (3.18)

The equation (3.18) shows that the conductor carries effective current equals to its

thermal limit.

44

3.3.1 Selection of Transmission Voltages

The instantaneous value of each conductor voltage with respect to ground becomes

more in case of simultaneous ac-dc transmission system by the amount of the dc voltage

superimposed on ac and more discs are to be added in each string insulator to withstand

this increased dc voltage. However, there is no change required in the conductor

separation distance, as the line-to-line voltage remains unaltered. Therefore, tower

structure does not need any modification if same conductor is used.

Another possibility could be that the original ac voltage of the transmission be reduced

as dc voltage is added such that peak voltage with respect to ground remain unchanged

and there is no need to modify the towers and insulator strings.

Let Vph be per phase rms voltage of original ac line. Let also Vg be the per phase

voltage of ac component of simultaneous ac-dc line with dc voltage Vd superimposed on

it. As insulators remain unchanged, the peak voltage in both cases should be equal.

V2Vph = Vd+V2Va=V„ax (3.19)

Electric field produced by any conductor possesses a dc component superimposed on a

sinusoidally varying ac component. But the instantaneous electric field polarity changes its

sign twice in a cycle if (Vd /Vg) < V2 is insured Under this condition the higher creepage

distance requirement for insulator discs as required for HVDC lines are not needed in this

case.

Each conductor is to be insulated for Vmax but the line-to-line volt^e has no dc

component and Vixmax = V6Va. Therefore, conductor to conductor separation distance of

each line is determined only by rated ac voltage of the line.

Maximum permissible dc voltage offset is that for which the composite voltage wave

just touches zero in each cycle; requiring:

Va = Vph/2and Vd = Vph/V2 (3.20)

45

So the minimum value of ac phase vohage and maximum value of dc voltage with

respect to groxmd of the converted line are V2 and 1/V2 times that of per phase voltage

before conversion respectively.

Insulation design consideration [2,3,8,12]:

Let us define factor Ki such that

Ki=dc withstand voltage/rms ac withstand voltage

If calculated in straightforward manner for overhead line

Ki=V2

A factor K2 may be defined as;

K2=ac insulation level/rated ac voltage

For overhead line K2 = 2.5

This is because high transient over voltages are possible for ac lines.

Similarly for dc side design, a factor K3 may be defined as;

K3=dc insulation leveL rated dc voltage

For overhead line K3«1.7

The actual ratio of insulation level is (ac/dc):

K=Ki ' '" (3.21) K3V,

Practically the converted ac line voltage may be selected a little higher than Va = Vph/2

to have two natural zero crossings in phase voltage (Va) wave cycle. This would avoid the

need of higher creepage distance requirement for insulator discs as required for HVDC

lines.

3.4 Preliminary Economic Consideration

To get the advantages of parallel ac-dc transmission such as improving stability and

damping oscillations, conversion of a double circuit ac line to simultaneous ac-dc power

flow line has been considered such that no alterations in conductors, insulator strings and

towers of the original line are needed. The minimum values of ac phase to neutral voltage

and maximum dc voltage with respect to ground of the converted line are selected as Y2

and I/V2 times the phase voltage before conversion respectively.

46

The cost of transmission line includes the investment and operational costs. The

investment includes costs of Right of Way (RoW), transmission tower, conductors,

insulators, lobour and terminal equipments. The operational costs include mainly the cost

of losses. Additional costs of compensation and its terminal equipments also influences

the ac line cost. DC transmission line itself requires no reactive power but converters at

both ends consume reactive power from ac system. This reactive power varies with the

transmitted power and is approximately half of the latter at each end. So dc line does not

require compensation but the terminal equipment costs are added due to the presence of

converters and fihers.

Replacement of Y- connected transformer with zig-zag transformer which is required

for conversion of pure ac line to simultaneous ac-dc power transmission line may not

increase its cost. This is due to the fact that the ac power transfer by transformer action is

about 25% that of total power. Moreover, the ac voltage reduces to 50% of the original ac

voltage. However, the neutral point of this transformer needs insulation to withstand dc

voltage. The loadability is observed to be almost getting doubled or more than double with

the simultaneous ac-dc power flow for a line of 500 km or longer than 500 km line.

Compared to system where a separate dc line is used in parallel with ac line, in the

proposed system ac line is converted for simultaneous ac-dc power flow, the additional

investment with dc transmission line and ac line compensation are saved.

3.5 Protection

Preliminary qualitative analysis suggests that commonly used techniques in

HVDC/EHV ac system may be adopted for the purpose of the design of protective

scheme, filter and instrumentation network to be used with the converted line for

simultaneous ac-dc power flow. In case of symmetrical faults in the transmission system

gate signals to all the SCRs are blocked and the bypass valves are activated or force

retardation method is applied (i.e. forcing the rectifier into inversion) to protect rectifier

and inverter twidges. CBs are then tripped at both ends to isolate the faulty system. Any

asymmetrical faults will create inequality in the dc current flowing througli the secondary

of the zig-zag transformer, which will result in saturation of the transformer core.

47

Eventually, the ac current on primary side will increase. Primary side CBs, designed to

clear transformer terminal faults and winding faults, would clear these faults easily at

shortest possible time. Blocking the trigger pulses to the bridges and activating the bypass

valves of the bridge would provide protection on dc side. A surge diverter connected

between the zig-zag neutral and ground would protect the converter bridge against any

over voltage.

DC current and voltages may be measured at zigzag winding neutral terminals by

adopting common methods used in HVDC system. AC component of transmission line

voltage is measured with conventional cvts used in EHV ac lines. Superimposed dc

voltage in the transmission line does not affect the working of cvts. Linear couplers with

high air-gap core may be employed for measurement of ac component of line current. DC

component of line current would not saturate me high air-gap cores.

Electric signal processing circuits may be used to generate composite line voltage

and current waveforms from the signals obtained for dc and ac components of voltage and

current. Those signals are used for protection and control purposes.

3.6 Effect of Line Capacitance

Simultaneous ac-dc transmission has a significant advantage over HVDC transmission

due to its ability to utilize the line capacitance. In pure HVDC system, capacitance of

transmission line cannot be utilized to compensate inductive VAR, as the dc line voltage is

constant with time. The rectifier and inverter bridges consumes lagging VAR (about 50%

to 60% that of active power) for their operation [1-3,7,8]. This VAR requirement increases

with gate firing angle of thyristors. The VAR of the converter in addition to lagging VAR

of load is to be supplied by synchronous condenser or static capacitor.

In simultaneous ac- dc power transmission, the superimposed ac-dc voltage varies

with time and the transmission line capacitance appears as shunt admittance to converter

and in parallel to the load. For example 400KV, 500Km, 3-phase transmission line having

shunt admittance of y=j3.3797*10'* mho/ph/Km, the total leading VAR available is

270.378MVA (3.3797*10"^ *500*400^). For long EHV line when receiving end power is

48

less than its natiiral load there is an excess of line charging; Qs is negative and Qr is

positive. This huge amount of leading VAR compensates partly or fully the lagging VAR

requirement of converter and load. But it remains latent in HVDC transmission

3.7 Feasibility Test for simultaneous ac-dc transmission

3.7.1 Simulation Study

To demonstrate the feasibility of simultaneous transfer of ac-dc power through the

same line, a scheme has been considered for detailed study:

The network depicted in Fig.3.1 was studied using PSCAD/EMTDC. A sjTichronous

machine is feeding power to infmite bus via a double circuit, tliree-phase, originally

designed for 400 KV, 50Hz, 450Km, ac transmission line. The 2750 (5x550) MVA, 24.0

KV, synchronous machine, is dynamically modelled, a field coil on d-axis and a damper

coil on q-axis, by Park's equations with the frame of reference based in rotor [1]. It is

equipped with an IEEE type AC4A excitation system [1,10].

The scheme of Fig. 3.1 has been modelled in a three- phase system in

PSCAD/EMTDC environment [10]. Transformers at sending and receiving ends are

modelled using (i) classical approach and also using (ii) the UMEC (Unified Magnetic

Equivalent Circuit) approach from three single phase, three windings, with their primary

and secondary windings connected in delta and zig-zag fashion respectively as shown in

Fig. 3.3a and Fig. 3.3b. Same type of model is selected both for sending and receiving

end. The study is made with both types of models. Both types of model produce same

results. Delta - Star Zig Zag cormection that gives a zero degree phase shift between the

voltages of the two sides. The connections are determined based on a simple phasor based

analysis. The Bergeron model is chosen for three-phase transmission line, as this model is

most suitable for load flow studies. The parameters are edited to suit 400 kV, 450 Km.

The dc system consists of a bipolar converter-inverter modelled as two six pulse bridge.

Each six pulse bridge is a compact PSCAD representation of a dc converter, which

includes a built in 6-pulse Graetz converter bridge (can be inverter or rectifier), an internal

Phase Locked Oscillator (PLO), firing and valve blocking controls and firing angle

49

(a)/extinction angle ( / ) measurements. It also includes built in RC snubber circuits for

each thyristor. Their control system consists of constant current (CC) and constant

extinction angle (CEA) and voltage dependent current order limiters (VDCOL) control.

The controls used in dc system are those of CIGRE Benchmark [11], modified to suit at

desired dc voltage. Two six- pulse bridges at each end of line generate 11"' and 13"

harmonics and inject these at respective ac buses. AC filters at each end on ac sides of

converter transformers are connected to filter out these ll"" and 13'*" haimonics. These

filters and shunt capacitor supply reactive power requirements of the converters. The dc

power obtained firom existing ac system by rectifier is injected into neutral terminal of zig­

zag connected secondary winding of sending end transformer through smoothing reactor.

From far end zig-zag connected transformer neutral, the dc power is inverted to ac and fed

to infinite bus.

SA

SB

#1

sc

#1

#1

\- #3

— • SC1

SN1

Fig. 3.3a. Connection diagram of delta- zig-zag transformer (Classical Model).

50

Va —

Vb

Vc

Va4

Vb4

Vc4

Aid

Fig. 3.3b. Connection diagram of delta- zig-zag transforaier (UMEC Model).

3.7.2 Simulation Results

Fig. 3.5 shows volt^es between conductor to ground and conductor to conductor

respectively. The upper curve shows instantaneous value of conductor to ground (i.e

across string) voltage having dc voltage component superimposed on ac and crosses zero

two times in each cycle and its maximum value Emax remains same as original line of 400

kV. The magnitude of ac and dc current flowing in the conductor is depicted in Fig.3.4.

By injection of dc at neutral point, all phases are offset by same amount. Therefore, the

line-to-line voltage is not affected by dc injection. This fact is elaborated by the second

curve in the same figure.

Fig. 3.6 shows the waveforms of primary phase current (Ipal, Ipbl, Ipcl), magnetizing

currents (Imal, Imbl, Imcl) and fluxes (Flxal, Flxbl, Flxcl) of sending end transformer.

51

In this transformer, the dc current (Id) has been injected through the zig-zag connected

secondary winding neutral which is divided among all three phases equally (i.e. U /3),

superimposes on ac current and enters into line conductors. Curves Imal, Imbl, and Imcl

indicate the magnetizing current waveform of each phase. It has been observed that the

waveforms of the magnetizing currents remain the same with and without dc current

injection. Two fluxes produced by the dc current (I /3) flowing through each half of a

secondary winding in each limb of the core of a zig-zag transformer are equal in

magnitude and opposite in direction. So the net dc flux at any instant of time remains zero

in each limb of the core. Thus the dc saturation of the core is avoided. Zig-zag

transformers provide very high (magnetizing) impedance to positive and negative

sequence currents. The transformer flax waveforms Flxal, Flxbl and Flxcl indicate that

the cores do not saturate due to injection of dc current into transformer secondary.

Saturation is enabled in first run at primary winding and in second run enabled on

secondary windings. In both runs, there is no difference in responses.

Amrr

Idc

i 2

4,87083

eter - 1

lac

•2 2 kA

0.908349

Fig.3.4. RMS values ac and dc currents injected.

52

Conductor To Ground Voltage

' V.Line-qround

m

> v>

13 o O

CD cn iS "o > o O

o O

5.220 5.240 5.260 5.280 5.300 5.320 5.340

Fig. 3.5. Voltage across the insulator string (V_Line-ground), and conductor-to-

conductor voltage (Vcond-cond).

3 O £• Q.

3

o

c TO

53

0|pa1

AM/' \ /

Olma1 0.150 ^

Dlmb1 Almd

-0.050 - 'Y'4\

-0.100--i"^

-0.150 r

. .A A . A A -A A A - A A A -A- A A A A A

OFIxal 1.00 -r^~ 0.75-i '

• Flxbl AFIxd

0.50 ^ r—y-

P^ / " \ A A A A , '"'•' A A A A niA i •'^ A STN I?"'\ J''

wv/vAA/ a-4.180 4.200 4.220 4.240 4.260

Fig. 3.6. Transformer's primary, magnetizing ciarrents and fluxes.

3.7.3 Experimental Verification

The feasibility of the conversion of the ac line to simultaneous ac-dc line was verified

on a laboratory size model. The circuit diagram for this experimental set-up is shown in

Fig. 3.7. The basic objective is to verify the operation of the transformers, particularly, the

effect on core saturation due to the superimposed ac-dc current flow and the power flow

control.

54

»o V </>

55

The transmission line was modelled as a 3-phase TC- network hiaving L = SSn^and^C =

O.SuF. Transformers having a rating of 6 KVA, 400/ 75/ 75 Volts were usecLat.B ch end.

The secondary of these transformers are connected in zig-zag fashion such that there is

zero phase difference between primary and secondary voltages. The secondary line

voltages at zig-zag terminals are 230V. A supply of 3-ph, 400 V, 50 Hz was given at the

input of delta primary winding and a voltage of 230V is obtained at zig-zag secondary at

sending end transformer and 400V, 0-5.25 kW variable resistive load was connected at the

receiving end. Two identical line-commutated 6-pulse bridge converters wei-e used for

rectifier and inverter. A 10 Amp, 23mH dc smoothening reactor (Xj) was used at each end

in between converters and zig-zag connected neutral points. The dc voltages of rectifier

and inverter bridges were adjusted through ac input and firing angles to vary dc current

between 0 to 6A. AC filters were not connected at converter ac buses.

The power transmission with and witliout dc component was found to be in

conformity with theory. There was no satiation of the transformers core with and without

dc component as evident firom the waveforms.

Experimental results for 3A ac current with superimposed 2/3A <lc current in each line

at 120V dc voltage are depicted in following figures from Fig. 3.8 to Fig. 3.16.

v» f- o j : U.COi

•v_^- ^-—y x._.-^ -i . y

H H i ::oorf,'.,.' C H ? 20OrT;',''

: ; H 2 2yun M S.OOii-is ._^^J

i\/- \/\' A \/ \/ V y -:;:.

Fig. 3.8. Phase to ground voltage with Vdc=120V.

(Voltage Probe setting 200mV = 100V)

56

TeK M P D 5 ; —^KHII.CJJ? M c a 5 u r r '* t - T " t — " • T J f T - T - T—^*i—r—T—r- 'T r' r T"T'"'<' 1 "r'T—»~T—r

A . , , ! , , .

. / ^ /I T v-D r

^ . /

Fig. 3.9. Transformer primary current without dc injection.

(Cuirent Probe setting 100mV= lA)

I8,k 19)1 - »4 w . - j t - — i f ' t 1 ', ! t 1

/ \ r, . -

/ . 1 I

"•W-; ','

Fig. 3.10

/ -

• -

y 1 : ' 0 ^ '

. • i 1 L . i . 1—t i i i a 1 1 > • * • . ^ » . , II . ! ,

M S.aOrHS CHI 7^ -y.QGmV

<10Hz

Transformer primary current with dc injected via zig-zag connected neutral.

(Current Probe settinglOOmV= 1 A).

57

T e R Jl^^..,_,.J»L^::£:£,,y.._.,..-.-,-^:'J:^5u;-^' —1—r-

:mi

Fig. 3.11. Smoothed Vdcr with dc filter at rectifier ouput (Ld=23 niH

Cpiiter lOOOmicro F).

(Voltage Probe setting 200mV= lOOV)

z X: A A A : / \ A . A

.ijcvor

* * ' 1*11 .^ I I * S ' ' ' ' T ^ - - H ' t ' ' ' ^ - - * - - * - J « t I « I . 1 - 1 J. > • J t a t, * * i I > 1 I i t

' H ^ ZJjGm'"' C f"i t ' r ' ' ' ' " ^ t " » "

Fig. 3.12. AC voltage at input to rectifier (ELL )

(Voltage Probe setting 0.2V/Div= lOOV)

58

leK J L. «l»iscc' r l Pc?; -atJ•J.M.l.^ :..L'ft:..-. :-: "•"^—r-1 I 1 I " I r I I—r-1—f—1—!—I—!—r-i—7-^! f—T: !—r—^r-i—»—P-T—T—«—i i <—r-r-?—^i—t—r-*—r—.—

r . . . . . . .1 T':'" *•

r'- • ; / - • • *5 / • v . / 1/ \ / • • V / • • ^ - . - . - . , , . .

i -A. • 1 ^ A \ h- A k i ^m

'V

f\ -

I . . J i ,:

' ' J - .

vUw; '

Fig. 3.13. Conductor line-to-line voltage (VL)

(Voltage Probe setting 0.2v s 1OOV).

• A ••^•:^'; A-^' A ^ <^iSES

,' i

1 « • i 1

f

'•"v.'j

* ; ' I ' _i . . , r ( .- ( 1 I ^ r •• .• *H

i: .w ^ y;..; :..\L..z ;...Nrf,..: r.-; n ;

i 0:C '

I I > i I • ;• 1 1-1 t .^ . ' — ^

CHI f tO.OrnV

Fig. 3.14. Simultaneous ac dc current in Phase A.

(Current Probe setting lOOmV s 1 A)

59

T&jrv ^. ,^ . , ^M^-dl: :J_::.^_.^lJ^.:;.^^.^-,,

!*•"•> iM*,iwai3''

Fig. 3.15. Rectifier ac input line current (111) in Phase A.

(Current probe setting lOOmV = \A)

K ..__ _ .__: '»1.

Fig. 3.16. Rectifier do current (Idcr).

(Current probe setting 1 OOmV = 1 A)

The shape and magnitude of primary current of the zig-zag connected transformer

remains unchanged with and without injection of dc current as shown in Fig. 3.9 and

Fig3.10. The wave shape of line-line voltage, as shown in Fig. 3.13, has no dc offset in

60

conformity with the theoretically predicted result. Rectifier input ac current is stepped

shaped as no filters has been connected on ac side as shown in Fig. 3.15. The waveform in

Fig. 3.14 indicates that the current in the conductor is offset by the amount of the dc

current added in the conductor.

3.8 Conclusion

Simulation as well as experimental results establishes the feasibility of simultaneous

ac-dc power transmission. Various issues involved are highlighted. Main concern of

avoidance of transformer core saturation is satisfactorily addressed. Other related issued

are also discussed qualitatively. The higher creepage distance requirement for insulator

discs as required in case of HVDC lines are not required in this case because electric field

produced by conductor to ground voltage reverses its polarity twice in a cycle. The added

dc power does not interfere with normal fimctioning of ac system; rather it improves the

stability as will be demonstrated in the subsequent chapters. The line can be loaded up to

conductor's thermal current limit.

61

Photograph of Experimental Set-up «*^"*«»*

62

CHAPTER 4

POWER UPGRADATION OF EXISTING AC TRANSMISSION LINES BY CONVERSION TO SIMULTANEOUS AC-DC COMPOSITE LINES

4.1 Introduction

In recent years, public has become more sensitive to the proliferation of overhead

transmission right-of-ways. Industrial countries are experiencing increasing difficulty in

finding suitable corridors for new overhead transmission lines and in many cases it is

simply impossible. There is an increasing pressure to provide the substantial power

upgrading of existing ac transmission line corridors. Two possible suggestions discussed

in reference [12,13] for power upgrading of existing ac transmission are:

(i) appropriate modification to existing ac lines without major new construction

with increased voltage level either ac or dc.

(ii) elimination of the old existing HV/EHV ac lines and their substitution with

new lines of EHV/LHV ac or HVDC. This would amount to major change.

A third possibility suggested in reference [66] is adding more circuits onto existing

right-of-ways. One solution to this pioblem, used by Ontario Hydro and others [66], is to

restructuring existing lower voltage transmission or distribution circuits onto new towers

with new high-voltage circuits. This is forcing many electric utilities to utilize existing

corridors by upgrading existing right-of-way (RoW). which will be far cheaper than going

to underground transmission. In all these suggestions, there are major changes required to

get substantial power upgrading of transmission line.

The novelty of the present work of this thesis demonstrates the substantial power

upgrading of an existing EHV ac lines by simultaneous ac-dc power transmission. No

alterations of conductors, insulator strings and tower structures of the original EHV ac line

63

are required in this case. In this chapter the aspect of power upgradation through

simultaneous ac-dc power transmission has been taken up for detailed study.

4.2 Power Up-gradation of EHV ac Transmission Line by simultaneous ac-dc power transmission

In the proposed study of conversion of existing EHV ac transmission line to

simultaneous ac-dc transmission line, no alterations of conductors, insulator strings and

towers structures of the original EHV ac line are required. The ac and dc voltage for the

simultaneous ac-dc converted line may be selected as:

Let V be per phase mis voltage of original ac line. This line is converted to

simultaneous ac-dc transmission line. Let V be the per phase voltage of ac component

of simultaneous ac-dc line with dc voltage V superimposed ua it. As insulators remain

unchanged, the peak voltage in both cases should be equal. Allowing maximum

permissible voltage offset such that the composite voltage wave just touches zero in each

cycle, the maximum value of dc voltage and minimum ac voltage are:

V

Electric field produced by any conductor possesses a dc component superimpose a

sinusoidally varying ac component. The instantaneous electric field polarity changes its

sign twice in a cycle if (Vd A a) < ^2 is insured Therefore, higher creepage distance

requirement for insulator discs used for HVDC lines are not required.

The total power transfer through the double circuij line before conversion (i.e. original

existing ac line) is;

P'.otai^SVph'SinS./X (4.1)

X is the transfer reactance per phase of the double circuit line (or of single circuit line if

considered for conversion) and 5i is the power angle between the voltages at the two ends.

64

To keep sufficient stability margin, 61 is generally kept low for long lines and its value

seldom exceeds 30°[1]. With the increasing length of line, the loadability of the line is

decreased [1,48]. . ^ approximate value of 5i may be computed from the loadability curve

by knovkdng the values of Surge Impedance Loading (SIL) and transfer reactance X of the

line.

P'totai = 2.M.SIL (4.2)

Where M is the multiplying factor and its magnitude decreases with the length of line.

The value of M can be obtained from the loadability curve [1].

The total power transfer thi-ough the composite ac-dc line;

Ptotal = Pac + Pdc = 3 Va^SinS^/X + 2VdId (4 .3 )

The power angle 62 between the ac voltages at the two ends of the composite line may

be increased to a high value due to fast controllability of dc component of power. For a

constant value of total power, Pjc may be modulated by fast control of the current

controller of dc power converters.

Approximate value of ac current per phase per circuit of the double circuit line may be

computed as:

la = V (Sin5/2)/X (4.5)

The dc current order for rectifier and inverter is adjusted on-line as:

Id=3 |h,''-K" (4.6)

A master current controller (MCC) shown in Fig. 4.1 is used to control the current

order for converters on-line. It measures the conductor ac current, computes the

permissible dc current in the conductor and produces dc current order for inverters and

rectifiers such that the conductor current never exceeds the thermal limit.

Power upgradation in percent or increase in power transfer capability of a composite

ac-dc line may be defined as:

65

% Power Upgradation

_ Power tranfer b)' composite ac - dc line - Power transfer by pure ac line Pov/er transfer by pure ac line

P - P '

P' total

(4.7)

a meas

'a

MCC

l.=3[l,,^-l3T^

i i

ui_set

1 'd

p.u.convesion

Fig. 4.1. Master Current Controller (MCC).

4.3 Case Studies

Double Circuit EHVac Line

4.3.1 Power Transfer Capability of a pure ac line

(A.1) COMPUTATION

The loadabilit}' of a Moose (Commercial Name), ACSR, twin bimdle conductor,

400kV, 50 Hz, 450 km double circuit ac line has been computed.

The parameters of the line are:

z = 0.03252 + jO.33086 Q/km/ph/circuit.;

y=j3.33797xl0"^ S/km/ph/circuit.

Current carrying capacity of each sub-conductor = 0.9 kA.

Current carrying capacity of twin bundle conductor, Ith = 1.8 kA/circuit.

66

The simulated loadability curve [9] of the t\vin bundle conductor, 400kV, 50 Hz, ac

transmission line is shown in Fig.4.2.

SIL = 511.22MW/circuit.

M = 1.1, (from loadability curve Fig. 4.2);

X = 74.4435 n/ph

Using equations (4.1), (4.2), and (4.5), the computed power and conductor current at

receiving end are:

P toui = 1124.68 MW; (at 5, ~ 30°);

and Iph/circuif = 0.803 kA;

There is a huge current margin between the current carrying capacity of twin bundle

conductor and the current it cEirries at this power loadability of the line.

10-

P.U. SIL 6-

Loadability curve 400 KV S.C. Line

Conductor Current Capacity=1.8kA(Twin pond.)

I

Inductance:

Capacitance-0.0107SmicroF/km

Frequency=50 Hz

L = 51:1.2226MW, -delta 4 SOdegrees

1.053 mH/km

. 1 . . I r

Thertral

Theoretical stability

rracneai ime

100 200 300 400 500 600 700 800 900 1000

Line length (Km)

Fig. 4.2. Loadability curve of a 400 kV single circuit line.

67

(A.2) SIMULATION

The scheme shown in Fig. 4.3 has been simulated in steady state in PSCAD/EMTDC

environment. A s>'nchronous machine is feeding power to infinite bus via a double circuit,

three-phase, 400 KV, 50Hz, 450Km, ac transmission line.

PS3

T ^^'^ ««B4F.icrtr

. ^ 1 Jmn ~" ^J T^\

Genlrator

&

--K-

^ B-

T

a.

-^'Z^-^

A^^^;^)-<2; Infin.te

Bus

Fig. 4.3. A synchronous machine feeding power to Infinite bus through double circuit ac line.

The simulated results in steady state ai-e shown in Fig. 4.4 to Fig.4.6 below:

Volt Meters - Powermeter

Vs-rms

/ ^ A 500 kV

411.957

Vr-rms j

500 kV

400.552

Ps

0 1500 MW

1175.79

Pr

0 1500! MWi

1116.55 1

Fig. 4.4. Bus voltages and powers at sending and receiving ends.

Cond.C... - I Anal.

la

^ ' 2

kA

0.805246

Trans_Ang)e

. ^

1

90 I degrees I

29.9971

Fig. 4.5. Conductor ac current. Fig. 4.6. Angle between two ends buses.

68

The computed and simulated results are found to be in close conformit)'. In

computation the both end voltages have been taken as 400kV where as in simulation, these

are taken Vs=412 kV and Vr = 400kV.

The conductor current 0.805246 L\ is much below its thermal limit 1.8 kA.

So there is a huge margin in the conductor current capacity.

4.3.2 Conventional double circuit EHV ac line Converted to Composite ac-dc Power Transmission Line

(B.1) COMPUTATION

The 400 kV ac line is converted to simultaneous ac-dc transmission line as per

equations (3.20) and (3.21. The minimum ac voltage and maximum dc voltage are found

as: 115.473kV and 163.328 kV.

This ac voltage Va has been increased from 115.473kV to 120 kV and \'d has been

decreased from 163.328 kV to 160.0kV to have two zero crossings in voltage wave. Thus:

Va=120kVandVd=160kV

82 is assumed to vary maximum up to 80°, which is commonly used value in lines

controlled by FACTS devices [62].

The computed approximate power transfers and also ac, dc component of currents

along with superimposed ac-dc currents for converted line at various transmission angles

are shown in Table-I

69

TABLE-I. COMPUVHD RESULTS USFNG APPROXIMATE FORMULAS.

Power Angle (5)

Degrees

ac power(M\V)

-3Va^Sin62/X

ac current la (kA)

la = V(Sin5/2)/X

dc Current (kA)

Id= 3 JI,f, - - /„ "

Dc Power Pdc=2Vdi x

Idi (MW)

Ptotal~Pac'*"Pdc

(MW)

30'^

290

0.4166

5.253

1684.8

197)

45'»

410

0.6122

5.078

1624.9

2034

60"

502.61

0.805

4.829

1545.5

2048

75"

560.6

0.98

4.529

1149.4

2010

80"

571.55

1.035

4.418

1413.76

1985

The results tabulated in TABLE-II have been computed for simultaneous ac dc, 450 km

line on MATLAB software by using equations (3.1) io (3.18). The calculated values using

approximate equations given in Table-I are in close proximit>' with the exact values. In

both cases, maximum power transfer at receiving end of the line occurs at a transmission

angle of 60°. The fact is also indicated in Fig. 4.7.

70

TABLE -II COMPUTED RESULTS FOR 450 KM LINE LENGTH FOR DIFFERENT TRANSMISSION ANGLES.

Trans.

Angle

(S)

IS''

30"

45"

60"

75"

80"

Ac

current

la(kA)

0.21641

0.40865

0.60062

0.78414

0.95503

1.0086

; Id/cond.

(kA)

1.7869

1.753

1.6968

i 1.6202

1.5258

1 1.4909

i

Total

dc

Current

Id (kA)

^ .3608

5.259

5.0905

4.8607

4.5773

4.4726

S-£nd

Qacs

MVAR

-60.487

-14.374

68.974

183.88

322.5

372.38

R-End

Qacr

MVAR

31.193

-42.218

-149.01

-281.9

-431.83

-483.85

Pac

MW

152.78

291.4

406.31

495.12

535.77

542.24

Pdc

MW

1715.5

1682.9

1629

1553.1

1464.7

1431.2

Ptoral

= P!1C

+Pdc

MV/

1868.2

1974.3

2035.3

2048.3

2000.5

1973.5

Upgrad

ing of

Line in

percent

by

simult.

ac-dc

66.1%

75.54%

80.96%

82.123

%

77.87%

75.47%

71

Real Power transfer 2500

5 10 15 20 25 30 35 40 45 50 55 60 65 70 75 80 85 90 Transmission Angle in Deg.

Fig. 4.7. AC, DC, and total power at receiving end for different transmission angles.

72

(B.2) SIMULATION

The proposed simultaneous ac-dc power transmission scheme shown in Fig. 3.1 has

been simulated in steady state mode using PSCAD software package. A sj nchronous

machine feeding power to infinite bus via a double circuit, three-phase, 400 KV, 50H2,

450Km, ac transmission line is considered. The 2750 (5x550) MVA, 24.0 KV.

synchronous machine, is djnamicaily modeled with a field coil on d-axis and a dcmper

coil on q-axis, with the frame of reference based on rotor. [1,10] It is equipped with an

IEEE type AC4A excitation system for which the block diagram is shown in Fig. 4.8

[1,10]. Transmission lines are represented as the Bergeron model. It is based on travelling

wave line model with distributed LC parameters and lumped resistance.

The converters on each end of dc link are modelled as line commutated two six- pulse

bridges. Their control system consists of constant current (CC) and constant extinction

angle (CEA) and voltage dependent current order limiters (VDCOL) control. The

converters are connected to ac buses via Y-Y and Y-A converter transfonners. Each

bridge is a compact PSCAD representation of a DC converter, which includes a built in 6-

pulse Graetz converter bridge (can be inverter or rectifier), an internal Phase Locked

Oscillator (PLO), firing and valve blocking controls and firing angle (Qr)/extinction angle

(x) measurements. It also includes built in RC snubber circuits for each th>Tistor. The

controls used in dc system are those of CIGRE Benchmark [7,11], modified to suit at

desired dc voltage. AC filters at each end on ac sides of converter transformers are

connected to filter out 11'*' and 13*'' hannonics. These filters and shunt capacitor supply

reactive power requirements of converters.

A master current controller shown in Fig. 4.1 is used to control the current order for

converters. It measures the conductor ac current, computes the permissible dc current and

produces dc current order for inverters and rectifiers.

73

•^. r J

l + sTc

1 + STB +sT. •F™

Fig. 4.8 IEEE type AC4A excitation system.

The wattmeter and ammeter readings at various points on the system are tabulated in

Table-Ill for Vs=209.0 kV and Vr=208.0 kV.

Pdcloss includes line loss due to dc current and converter losses. Pacloss is line loss

due to ac current only. The toml power Pr at receiving end is the actual net power transfer

after subtracting all losses like circuit breakers, transformers etc.

It has been seen from computaiion as well as simulation that the maximum power

transfer by composite ac-dc line occurs at power angle of 60 '. The same amount of power

transfer through conventional double circuit line would require a power angle of 73.68°,

which is beyond the safe limit for power angle. The corresponding conductor current

(Iph/ckt) is 1.7744 L\.

Table IV shows the simulation results for power transfer through different length of

lines. It has been observed that there is a substantial increase of power transfer capacity of

line by simultaneous ac dc power transmission as compared to ac transmission. The fact is

illustrated in Fig. 4.9.

74

TABLE-III. SIMULATION RESULTS OF 450KM COMPOSITE LINE AT VARIOUS TRANSMISSION ANGLES.

Power Angle (8)

Ps (MW)

Pac (MW)

Transfer

Pdc (MW)

Transfer

Pacjoss

(MW)

PdcJoss(MW)

Ploss_total(MW)

Pr (MW)

Total Transfer

Qsline

( M V A R )

Qrjine (MVAR)

Qrec (MVAR)

Qinv (MVAR)

ac current la (kA)

dc current Id (kA)

Cond. dc current

Id/3 (kA)

conductor current

Isim(kA)

Increase of power

transfer

30»

2306

294.89

1715.5

11.94

280.51

292.45

1988.8

-13.78

39.08

883.6

841.3

0.41577

5.24263

1.74754

1.78587

76.94%

45"

2371.0

411.00

1657.0

30.30

265.88

296.18

2051.14

69.98

146.840

884.36

823.5

0.61123

5.1136

1.70453

1.78264

82.49%

60"

2381.3

495.3

1585.8

54.08

241.17

295.25

2062.0

185.58

280.85

885.29

797.43

0.79684

4.91185

1.6373

1.78281

83.451%

75"

2342.0

541.86

1498.5

81.94

217.61

299.55

2019.36

325.12

431.96

878.1

764.64

0.96952

4.6355

1.5452

1.78641

79.66%

80" !

2318.380

548.43

1467.0

i 91.73

208.53

300.26

1995.00

375.35

484.38

869.48

753.04

1.02383

4.52512

1.5084

1.78833

77.5%

75

TABLE-TV. SIMULATION RESULTS AT 60° TRANSMISSION ANGLE FOR DIFFERENT LINE LENGTHS.

Line Length

in km

M

Pac (MW)

Pdc (MW)

Pt(MW)

la(KA)

Id(KA)

Isim (KA)

Qacs (MVAR)

Qacsr(MVAR)

% Power

Upgade

400

L2

552.05

1538.19

2070.0

0.8978

4.7721

1.78411

221.792

-329.0

68.71%

500

1.0

450.25

1615.48

2065.24

0.715761

4.99377

1.78239

155.2

-240.58

101.96%

600

0.85

383.95

1662.27

2046.22

0.594154

5.15329

1.7828

106.0

-176.5

135.44%

700

0.75

338.0

1650.0

1988.0

0.508

5.1311

1.75544

66.58

-126.39

159.25%

800

0.68

304.0

1600.0

1908.0

0.445254

5.0388

1.72465

33.00

-84.69

173.85%

76

Increased Power Transfer Vs Line lenght

600

Line Lenght in km

800

Fig. 4.9. Percent enhancement in power transfer for various lengths of a double circuit

ac-dc composite line.

Fig. 4.10a to Fig. 4.1 Oh show steady state response of various system performances at

power angle of 60 .

S-3nd and R-end vottaqes — i

o Vs-rms n Vr-rms i> 2 2 5 -,

fO O ' H O a Q o o a o a 2 0 0 -175 -1 5 0 -1 2 5 -100 -

75 -50 -25 -

O -

s®^ 0.0 2.0 4.0 6.0 8.0 10 .0

Fig. 4.10a.Sending (Vs) and receiving (Vr) end bus voltages.

Converter Vottaqes

o Vdcr • Vdci

>

sec

- « ? •

-63- -£3-

4.0 6.0 T 1

8.0

jd__.£ . ._ .

10.0

Fig. 4.1 Ob Rectifer (Vdcr) and inverter (Vdci) dc voltages.

O Idc-r • Idc-

-e«—Q—&e «—r<^e-

sec 0.0 - 1 1 1 r —

2.0 4.0 6.0 8.0 10.0

.1>J Fig. 4.10c. Rectifier (Idc-r) and inverter (Idc-I) current.

Converter Alpha Order

O AlphaR • AlphalN

<X)

E'

Q

sec 0.0

j j Ifiil'iis^iiiiii^^igjii'fiiiiiiiiiiamij •

77

Fig. 4.9d. Rectifier (AlphaR) and inverter (AlphalN) firing angles.

o Q_

a:

Active Power Tranfer

O Ps • Pac A Pdc Ptotal

i

• o"

t — j i r

rq -r-x^ , —

h T ? H — 1 ! I :

o o — ' - t^ i

-4-is

- e -

sec 0.0 2.0 4.0 6.0 8.0 10.0

• i

Fig. 4.10e. Sending end (Ps), ac (Pac), dc (Pdc) and

total transfer (Ptotaltr) power.

Reactive Power

A Q d c r • Qdc i

o

05 a>

sec 1 1-

0.0 2.0 4.0 6.0 8.0 10.0

Fig. 4.1 Of. Line sending (Qac_s), receiving (Qac_s) end,

rectifier (Qrec_dc),inverter (Qinvdc) reactive powers.

78

o AC_Currenrt(... • Idc/cond. AAC+DCCurr...

E

0 )

O

o

s e c 0.0 2.0 4.0 6.0 8.0 10.0

Fig. 4.10g. AC, dc and effective conductor ciurent

79

S

sec

Strinq Phase vottaqe

-50 -i

5.425 5.475 5.325 5.375

Fig. 4.1 Oh. Pha.se voltage across a insulator string.

It has been observed from above tracings that system behaves normally and no

abnormality is seen in the operation of the system when dc is superimposed on ac.

4.4 Conclusion

The merits to convert existing double circuit ac transmission line to composite ac-dc

transmission line for substantial power upgrading has been demonstrated. For the

particular system studied, there is substantial increase (about 83.45%) in the loadabilty of

the line. The loadabilty will further increase for longer lines. The line is loaded to its

thermal limit with the superimposed dc current. The dc power flow does not impose any

stability problem. This stability aspect wall be studied in detail in the next chapter. The

advantages of parallel ac-dc transmission aie obtained in simultaneous ac-dc power

transmission line. DC current regulator may modulate ac power flow. No modification is

required in the size of conductors, insulator strings and towers structure of the original

line.

8(3

CHAPTER 5

TRANSIENT STABILITY IMPROVEMENT BY SIMULTANEOUS AC DC POWER TRANSMISSION

5.1 Introduction

The power transfer capability of long ac transmission lines is usually limited by large

signal stability. Economic factors such as the high cost of long lines and revenue from the

delivery of additional power provide sti"ong incentives to explore all economically and

technically feasible means of raising the stability limit. The development of effective

ways to use transmission system close to its thermal limit has attracted much attention in

recent years [4,20]. The progress m the field of power electronics has already started to

influence the power industry. The emergence of FACTS devices is the outcome of this.

Very fast control of SCRs in FACTS devices like Static VAR System (SVS), Controlled

Series Capacitor (CSC), and Static Phase Shifter (SPS) and controlled Braking Resistors

improve stability and damp out oscillations in power system [1-7].

The transient stability criteria includes the ability of the system to withstand a three-

phase fault at critical locations such as near the heavily loaded generator bus at the line

carrying large amount of power in case of ac transmission [5-6] and near the inverter in

case of HVDC [49]. The statistical data indicate occurrence of single line to ground faults

in more than 90% of the total number of faults compared to three phase faults which

occur only about 1% of the total number of faults. Three phase faults are more severe and

transient stability criteria usually include ability to withstand:

(a) Three phase faults cleared by primary protection.

(b) Single line to ground faults cleared by backup protection (due to stuck primary

breaker).

The following are the various methods presently available to improve the transient

stability of the power system;

(i) High-speed fault clearing.

(ii) Reduction of transmission system reactance.

81

(iii) Regulated shunt compensation.

(iv) Dynamic braking.

(v) Reactor Switching.

(vi) Independent pole operation of Circuit Breakers

(vii) Single pole switching.

(viii) Steam Turbine Fast-Valving.

(ix) Generator Tripping.

(x) Controlled system separation and load shedding.

(xi) High Speed Excitation system.

(xii) Control of HVDC Transmission Links.

HVDC transmission lines in parallel with EHV ac lines are recommended to improve

transient and dynamic stability as well as to damp out oscillations in po^ 'er system.

Improvement of the stability characteristics of intercoimected power system bj' using fast

acting converter control on HVDC links in the system is well established [1-3,5-

8,14,16,18,24^7,29-3033,34,40-44,49-55,57,58,61].

In this chapter, the improvement of transient stability by rapidly modulating the dc

power in case of simultaneous ac-dc line has been taken up. A single machine infinite bus

(SMIB) connected by a double circuit ac line which is converted for simultaneous ac-dc

power transmission has been studied. Fig. 5.1 shows the system under study. For the

purpose of compaiison, the same double circuit line with pure ac transmission was also

studied. The transmission angle is varied up to 80° in case of simultaneous ac-dc power

transmission system. Such a large angle is not possible in pure ac system. Both the

configurations were subjected to the same types of faults at various locations of the line,

and their responses were assessed and compared with regards to transient stability using

PSCAD/EMTDC simulation package.

5.2 Control Principle and Strategy

First swing stability in a power system refers to the ability of a power system to

maintain a connected generator in s>'nchronism after the system has been subjected to a

major disturbance such as transmission system faults. The criterion for the first swing

82

Stability is that the accelerating energy, occurring whenever the generator electrical

power is less than the driving mechanical power, must be counterbalanced by an equal

decelerating energy, whenever generator electrical power exceeds the driving mechanical

power [37].

HVDC links, under traditional controls, do not provide synchronizing or damping

effects in response to disturbance on ac side. However, the controllability of an HVDC

link is inherently fast and this can be used to regulate the power flow for system damping

as well as solving the first swing stability problem.

Fig. 5.2 shows the power-angle characteristics of the system considered in Fig. 5.1

where

Pac Steady state ac power by ac system.

Pdc Steady state dc power by the dc system.

6o Steady state power angle.

5c Power angle at the time of clearing the fauh by opening CBs.

6r Power angle at the time when CBs are reclosed.

62 Maximum angle when transient stability power limit is occurred.

Ai Decrement of ac power (Pac) caused by the fault.

A2 Decrement of dc power (Pdc) caused by the fault (blocked firing pulses to SCRs

and activate bypass valves).

Bii Retarding energy to the generator when only one line in service.

B12 Retarding energy to the generator when second line

reclosed.

=0 when faulty line out of service.

Bl Total retarding energy of the generator( Bii+ B12).

=Bi 1 when faulty line is out of service.

Bdc Increment of dc power required to give sufficient retarding energy to generator.

Ai+A2=Kinetic energy gained by tlie rotor during acceleration.

83

j i

^ •i ^ 5

? 1

<1 >cr^

^-<h K

^

-a

o

3 . • • • " 5

O

CO

0 3 O

JD

O

03

>

un

84

When rapid control to increase clc power is not adopted, Bdc=0. The acceleration area

(A1+A2) becomes larger than deceleration (Bl) and the generator may step-out.

As a counter measure to improve the transient stability and solve the first swing

stability, the dc power is increased by an amount

Bdc= (A1+A2) - (Bl) (5.1)

The stepping out of the generator is thus prevented.

Let Pg' =the transient output of generator,

Pac' = the transient power of the ac system, and

Pdc' =the transient dc power.

Equation (5.1) can be represented as

j (Pdc'-Pdc)d6 &

& <>2

= J {(Pac-Pac')+(Pdc-Pdc')}d&- j (Pac-Pac)d5 (5.2) SO A-

The input Pm and output Pg of generator is related as:

Pm=Pg=Pac+Pdc in Steady state

Pg'=Pac' +Pdc' in transient conditions.

Using above relationship, (5.2) bsconies (5.3) as

^2 82 Sc 5;

J Pdc' d5 = }p„d5 - }p;d5 - ]X, ^6 (5.3) &• 8„ S„ 5;

Assuming Pdc' in (5.3) is constant during transient period, then,

?,,'«?^^1Z^.—L^i fp •d5+ \?Jd6) (5.4)

It must be noted that transient ac powers Pg' and Pac' have non-linear characteristics

and can be determined experimentally.

85

E B in >. (O O <

a> 5:

o Q.

O Q

iE Q. W *- >»

c

Bl / Prefault

^ " ^ ^ - / -^ Post fault ; ^ ; ^ ^ " . » ^ ^ (Single line in

service)

,-- During fault

2 ^ % ^ ^ ^ 5,

Power angle, 5 *

Power angle, 6 *

Fig. 5.2. Power-angle characteristics.

The criterion given in equation (5.1) is implemented by adding a Aide control signal to

the dc current order (lord) of the rectifier current regulator. The control signal (Aide) is

derived from speed deviation signal (Ac;) of the generator using a PI controller as shown

in Fig. 5.3.

The control signal is given by

AIdc=Kp(<y-6>o)+KiJ (d)-co o)dt (5.5)

Where, (o ando 0 are the actual and rated generator angular per imit speeds respectively.

While the first term on the right hand side of equation (5.5) influences the transient

behavior of the system, the second term influences the steady state. The proportional

control has the effect of increasing damping of the system; the effect of integral control is

to increase the sjTichronizing co-efficient. The control signal (Aide) provides retarding

energy to generator as well as damping to the system.

86

P.U. Conversion

Ao)-

AOR

Nr-''^ Control signal

Fig. 5.3. Rectifier Regulator.

5.3 Description of System Studies

The network depicted in Fig. 5.1 is taken up for detailed studies using

PSCAD/'EMTDC. A synchronous machine is delivering power to an infinite bus via a

double circuit ac transmission line. The line considered is a three-phase, 400 KV, 50Hz,

450Km. The 2625.0 MVA, 24.0 KV, synchronous machine, is dynamically modelled

with a field coil on d-axis and a diimper coil on q-a.\is with the frame of reference based

on rotor. It is equipped with the IEEE type AC4A excitation system [1,10] for which the

block diagram is shown in Fig. 5.4.

r J

l + sTc

1 + sTii - » •

1-sT, /

Fig. 5.4. IEEE type AC4 A excitation system.

The mechanical power supplied by turbine is considered invariant for the duration of

transient simulation runs.

Transmission lines are represented as Frequency Dependent (Phase) model [10]. The

Frequency Dependent (Phase) model is basically a distributed RLC traveling wave

model, which incorporates the fi-equency dependence of all parameters. This model is

most accurate and suitable for stability studies.

87

The converters on each end of dc link are modelled as two units of line commiitated

six-pulse bridges resulting into twelve-pulse output. The CIGRE Benchmark [11]

modified to suit at dc voltage and current in the present case has been adopted for the

control of converters. These include constant current (CC) and constant extinction angle

(CEA) and voltage dependent current order limiters (VDCOL). The converters are

supplied from ac bus through Y-Y and Y- A converter transformers. The transformer tap

is assumed to be constant (at the prior to the initiation of the disturbance) as the tap

changer dynamics is ver>' slow (5 to 6 sec/step) [1]. Each bridge is a compact PSCAD

representation of a DC converter, which includes a built in 6-pulse Graet2: converter

bridge (can be inverter or rectifier), an internal Phase Locked Oscillator (FLO), firing and

valve blocking controls and firing angle (a)/extinction angle (y) measurements [7,10]. It

also includes built in RC snubber circuits for each thyristor. AC filters at each end on ac

sides of converter transformers are connected to filter out 11''' and 13'*' harmonics. These

filters and shunt capacitor suppl> reactive power requirements of the converters. The

modelling details are given in chapter 2.

5.4 Transient State Simulation

A time domain simulation study is carried out to verify the effectiveness of proposed

scheme in the environment of PSCAD/EMTDC. For the pur]30se of comparison,

simulation study was repeated for the same line with conventional HVAC system. For

both cases, the real power transfer through the transmission line is kept as same. Initially,

the two schemes, namely, simultarieous ac-dc transmission and pure ac are assiamed to be

operating in steady state conditions. The details of the operating conditions are mentioned

in the Appendix. It has been observed through simulation studies, that simultaneous ac-dc

transmission system without dc modulation becomes unstable in the first swing.

However, with proposed control strategy of dc modulation the system becomes stable

after first swing.

A1. Simultaneous ac-dc transmission

Computation and simulation studies in chapter 3 have shown that in simultaneous ac-

88

dc power flow, line delivers maximum power at about transmission angle of 60°. The

details of EHV ac line converted to simultaneous ac-dc line are given in Appendix.

The simulation study is carried for the following cases:

Case (1): Fault near sending end

Initially, the simultaneous ac-dc system is assumed to be operating in its steady state.

The pre-fault power is 1.88 times loadability of original ac line at a transmission angle of

60°. At time t=0.5 sec a solid three-phase to ground fault occurs near sending end Fl as

shovvn in Fig. 5.1. The occurrence of fault is followed by blocking of tiring pulses to

converter SCRs and activation of kiypass valves. After a period of four cycles (80ms), trip

signals are given simultaneously to circuit breal-cers Bland B2 at both ends of faulty line

to clear the fault. Next, pulses to SCRs are released. Thereafter, CBs are reclosed after a

delay of five cycles (100ms) from the instant of clearing fault. Fig. 5 shows the transient

responses for this fault condition and the sequence of events that follows. > U signal

0.00 1 .00 2.00 3.00 4.00 5.00 6.00

Fig. 5.5a. Control signal (Li_signal).

0.0080 -I

0.0060 - •

0.0040 -

0.0020 - -

0.0000 - -

-0.0020 -

-0.0040-

-0.0060 - -

• Delta Omega

0.00 ' i . 6 o '2.60 '3.60 '4.00 '5.00 '6 .60

Fig. 5.5b. Generator's speed deviation (Delta Omega).

89

OPg D Qg

0.0 2.0 4.0 6 0

Fig. 5.5c. Generator's active (Pg) and reactive (Qg) power output.

sec

2 5 0 -

2 0 0 -

150 -

100 -

50 -

0 -

0.

O V:3

l|

j m s

I 0 10 2.0

3 vr rms

3.0

"' -' ~

Y •

4.0 5.0 6 0

Fig. 5.5d. Sending (Vsnns) and receiving (Vr_rms) end bus voltages.

CO a> Oi 1

Q

120-] 1 0 0 -

8 0 -6 0 -4 0 -2 0 -

0-1

O Transm Anqle

ii 1 / ' ^ ^ \

•-M- X ^ . 1 . ^ _ j _ _ ^O I Q Q O

0.0 1.0 2.0 3.0 4.0 5.0 6.0

Fig. 5.5e. Transmission angle ( Transm_Angle) variation.

O V d r D V d i 500 n

400 -

300 -

100 - — —

iL

1

r? t i

—-^-

_.i

i •

i

_

1

0.00 1.00 2.00 3.00 4.00

Fig. 5.5f. Rectifier (Vdr) and inverter (Vdi) dc voltages.

90

o idc rec D Id inv

sec 0.00 1 00 2.00 3.00 4.00

Fig 5.5g. Rectifier (Idcrec) and inverter (Id_inv) dc current.

ID 0) cn a> Q

1 8 0 T

150

120

90

60

3 0 - ^

0 -

OACR g Aoi

^^ ^^^^__^, ^ - ^

Jtf=^' '^~«-T*—e=

sec 0.0 1.0 2 0 3.0 4.0 5.0 &.0

Fig. 5.5h. Rectifier (AOR ) and inverter ( Aoi) firing angle order.

3.0k-r

2.5k-

OPao • Pinvdc AP tranfer

2.0k-r~* 1.5k

1.0k

0.5k-I ijf

0 0 J ^

•V ^'^.r'sz. -ir-r-

n y—•>

sec 0.0 'i.O 3.0 4.0 5.0 6.0

Fig. 5.5i. Ac (Pac), dc (I'invdc ) and total (P_tranfer) power transfer.

Case (2): Fault near receiving end

Now considering a similar solid three-phase to ground fault with the same sequence of

events near the receiving end F3 of the line (near inverter).

Various transient responses for this fault condition are shown in Fig. 5.6. • U signal

1.20 1.00 0.80 0.60 0.40 0.20 0.00

-0.20

] i

... -

— 4 — i — 1

\jrx J

1

1

1

— \ —

1

___L' i sec 0.00 1.00 2.00 3.00 4.00 5.00 6.00

Fig. 5.6a. Control signal (U_signal).

91

0.0080 0.0Q60 0.0040 0.0020 ^ 0.0000

-0.0020 -0.0040 -0.0060

I O^lta Omega

s®^ 0.00 1 .00 2.00 3.00 4 00 5.00 6 00

Fig. 5.6b. Generator's speed deviation (Delta Omega).

OPg • Qg

sec 0.0 2.0 4.0 6.0

Fig. 5.6c. Generator's active (Pg) and reactive (Qg) power output.

ov-.s ims • Vr rms

sec 0.0 10 2 0 3.0 4.0 5.0 6.0

Fig. 5.6d. Sending (Vs ntis) and receiving (Vr_rms ) end bus voltages.

0) a> a> Q

ssc 0.0 1.0 2.0 3.0 4.0 5.0 6.0

Fig. 5.6e. Transmission angle ( TransmAngle ) variation.

92

O Vdr

g •cs

500 -|

400 -

300 -

200 -

100 -

0 -

-100 -

_

—

"

^ ---

• Vdi

_ J 3 , ^ -ix Q — O

0.00 1.00 — I r— 1 1 1 :.00 3.00 4.00

Fig. 5.6f. Rectifier (Vdr) and inverter (Vdi) dc voltages.

o Mc rec • Id inv

sec 0.00 1.00 2.00 3.00 4.00

Fig. 5.6g. Rectifier (Idcrec) and inverter (Id_inv) dc current.

OAOR

01

to Q

0 1 Jrrhre

sec 0.0 1.0 2.0 3.0 4.0 5.0 6.0

Fig. 5.6h. Rectifier (AOR )and inverter (Aoi) firing angle order.

3.0k

2.5k

2.0k

1.5k

1.0k

0.5k

0.0

opac

- i / ^

^ I ^ \ ^ \ \

\

-^ir^!^

a Pinvcte A P tranfer

•• ; i j

l^u-r—PT " B H B 1 B

/ 1 ' i " - ! ! 1

. , _.!:_.„ "L_; j~ _:

sec 0.0 1.0 2.0 3.0 4.0 5.0 6.0

Fig.5.6i.Ac (Pac), dc (Pinvdc) and total (P_tranfer) power transfer.

93

Comparing the responses given in Fig. 5.5 and Fig. 5.6 for tlie case of sending end and

receiving end faults following main observations are made:

• The magnitude of speed deviation for sending and receiving end faults is

almost same, but speed deviation recovers faster in case of fault at receiving

end.

• The generator power is subjected to much higher peak transient for sending

end fault.

• The transmission angle overshoot is larger in case of sending end fault.

• After the clearance of :he fault the real power oscillations have same peak

values for both fault locations. However, in case of receiving end fault the

oscillations subside earlier.

• The responses of dc 'oltage and current indicate no commutation failure for

both fault conditions.

Thus the overall comparison of simulation results of Fig. 5.5 and Fig. 5.6 indicate that

the fault at the sending end is more severe than that at the receiving end for the system

under study.

B. Pure ac transmission with series compensation

The practical series capacitive compensation does not usually exceeds 75% for a

number of reasons, including load balancing with parallel path, high fault current, and the

possible difficulties of power flow control. In systems where sub-synchronous resonance

is main concern, the compensation is limited to 30% [4]. However, 50% compensation

has been taken in the present study so that the ac system becomes comparable with

simultaneous ac-dc configuration.

For the purpose of comparative study, the same network depicted in Fig. 5.1 is reduced

to pure HVAC by deleting the HVDC part and replacing zig-zag transformer with A-Y

transformer along with addition of compensation as depicted in Fig. 5.1 A. This EHV ac

system is assumed to be operating in its steady state and delivering same real power to

the receiving end as that in tlie case of simultaneous ac-dc system but at a transmission

angle of 30°. The exciter is equipped with the power system stabilizer (PSS). The gain of

94

PSS is assumed to be same as that of PI controller used in simultaneous ac-dc system for

dc modulation. The block diagram the PSS is shown in Fig. 5.IB and its pareimeters are

given in Appendix. Fig. 7 shows the transient responses when pure HVAC compensated

system is subjected to a solid three-phase to ground fault Fl at one of th<; line near

sending end. The fault is detected and circuit breakers (CBs) B3, B4 (Fig. 5.1 A.) of the

faulty line are opened after four cycles (80 ms) to clear the fault. Five cycles (100 ms)

after clearing the fault, the CBs are reclosed.

1 B'

-^Ztfr 2 -7fZ:^:S>W:

A'AlEictef

-3

Tut«» 'rV^-i -OLH^ Synchronou

^-IL-^-fT -f7 B. 3 , I

6

r - x,h

infinite Bus

A(o-

Fig. 5.1 A Pure AC study scheme.

G, '*^T' M + sT, V.

Fig. 5.IB. Power system stabilizer

M Delta Omega 0.0100 -[ r 1 - "—

0.0050 -

^ 0 . 0 0 0 0 - — / \ - - /

^ -0.0050 i-'-

-0.0100-L

sec ""

I 1

L _ . _ .

i i

i T "•"

0.0 2.0 4.0 6.0 8.0 10.0

Fig. 5.7a. Generator's speed deviation (Delta Omega).

I P.TDitI

<s>

0.0 2.0 4.0 6.0 8.0 10.0

Fig 5.7b. Transmission angle (PliDifl) between two ends.

O Pg a Q.3

3 a.

\ \ 4 - I I 1 - I T-^-5ec

"T 1 1 -T

2.0 4.0 6.0 8.0 O.D 2.0 4.0 6.0 8.0 10.0

Fig 5.7c. Generator active (Pg) and reactive (Qg) power (pu).

ops _ 3.0k

2.0k

1.0k

0.0-1

-1.0k

-2.0k-'

=iS

0.0 To 4.0 6.0 8.0 10.0

Fig. 5.7d. Sending (Ps) and receiving (Pr) end power.

95

A similar fault condition with the same sequence of events was created for pure HVAC

line with 30% compensation and 45° transmission angle. The pre-fault pov 'er level is

considered to be same as that for the above cases. The response of transmission angle for

this condition is shown in Fig. 5.8. An oscillatory response with slow damping is

observed. The PSS is not capable of providing an effective damping to the system at this

transmission angle.

96

1 PhDiff

ID Q

sec 0.0 2.0 4.0 6.0 8.0 10.0

Fig. 5.8. Transmission angle variation at 45° and 30% series compensation.

Increasing the transmission angle to 49 ^ and reducing the compensation le\'el in order

to maintain same pre-fault real power delivery, the system is subjected to same

disturbance. The transmission angle response shown in Fig. 5.9 is found to be oscillator,'-

leading to instability. A further increase of 1° in transmission angle makes the system

unstable just after first swing as evident from Fig. 10a and Fig. 10b.

• I PoDiff

CO O)

en

Q

90

60

30 4

0

-30

L ' " N ,

0.0 2.0 4.0 6.0 8.0 10.0

Fig. 5.9. Transmission angle between two buses at angle of 49° after fault.

0.0200 -0.0150-0.0100-

^ 0.0050-d 0.0000-

-0.0050--0.0100-J

• Celt 3 Omega

- - ! • - -

_ 1

-— - /^^V~

_3X-t-_„... 1 - . -

sec 8.0 10.0 0 0 2.0 4.0 6.0

Fig. 5.10a. Generator speed deviation at a transmission of angle of 50°.

97

I PhDiff

CO (D a> i _ a> Q

180-1 150-120-1

9 0 i 60-30-0 J

- -• • • • - — - '

- • • ' ; — -

fe;' sec 0.0 2.0 4.0 3.0 8.0 10.0

Fig. 5.10b. Transmission angle between two buses at an initial angle of 50°.

From the results shown above we conclude that in pure ac system, the transmission

angle between two ends of the line cannot exceed 45°, whereas in simultaneous ac dc

transmission system it could go up-to 80* , which will be demonstrated in the

following two cases, i.e. case (3), case (4) and depicted in Fig. 5.11 and Fig. 5.12.

A2. Simultaneous ac-dc transmission (continued)

Case (3): Three Phase to Ground (ABC-G) Fault near Receiving End/Inverter Responses at Transmission angle ofSCf:

The pre-fault power level is considered to be same as in case (1) and case (2) but at

transmission angle of 80°, a three phase to ground fault occurs with same sequence of

events. Various transient responses for this fauhed condition are illustrated below.

sec 0.0 1.0 2.0 3-0 4.0 5.0

Fig. 5.11 a Control Signal 8.0

98

3 a.

0.0080 -f 0.0060-0.0040-0.0020 -0.0000

-0.0020 4 -0.0040 4 -0.0060 -

' Delta Omega

sec 0.0 1.0 2.0 3.0 4.0 5.0 6.0

Fig. 5.1 lb. Geneiator's speed deviation (Delta Omega).

73

sec 0.0 2.0 3.0 4.0 5.0 6.0

Fig. 5.1 Ic. Generator's active (Pg) and reactive (Qg) power output.

275 n 250-"50^ Z.ZO 200-175-150-125-100-75-50-1

OVs rms

n J . J -u—'•

. . . , •

._.

- . ^ g . ^

jUr I- .. :- -

- - • • -

sec 0.0

• Vr rms

?jo a QB >'~:

1.0 2.0 3.0 4.0 5.0 6.0

Fig. 5.lid. Sending (Vsims) and receiving (Vr_nns) end bus voltages.

9'>

0 Transm_Angle

CD

Q

sec

Fig. 5.

0.0 1.0 2.0 3.0 4.0 5.0 6.0 1 le. Transmission angle ( TransmAngle) vaiiation.

>

• Vdi

•e4=-« r r e

•©-^ tr -e-

sec 0.0 1.0 2.0 3.0 4.0 5.0 6.0

Fig. 5.1 If. Rectifier (Vdr) and inverter (Vdi) do voltages.

7.50-1 7.00-6.50-6.00-5.50-5.00-4.50-4.00-3.50-3.00 -

O Idc rec • Id inv

sec o.o" 2.0 To" 6.0 Fig. S.llg. Rectifier (Idcrec) and inverter (Id_inv) dc current.

100

OAOR DAoi

CO m a> CI a> Q

sec 0.0 1.0 2.0 3.0 4.0 5.0 6.0

Fig. 5.1 Ih. Rectifier (AOR)and inverter (Aoi) firing angle order.

90-1 8 0 -7 0 -6 0 -50 -4 0 -3 0 -2 0 - -1 0 -0-1

GamfTia-lnverter

0.0 2.0 4.0

Fig. 5.1 li. Inverter Gamma.

"eL

5

opoc 2.8k T-

• Pinvdc AP tranfer

sec 0.0 1.0 2.0 3.0 4.0 5.0 6.0

Fig. 5.1 Ij.Ac (Pac), dc (Pinvdc) and total (Ptranfer) power transfer.

en

101

2.0k-

1.5k-

1.0k

0.5k

0 .0-

-0.5k-

-1.0k-

-1.5k-

-2.0k

OQrei^iJc noinvdc AQac-Send •Qac-Rend

^p„ijr-~--^ •G e-

•4S^-Kj^- - •

t»-

0.0 1.0 2.0 3.0 4.0 5.0 6.0

Fig. 5.1 Ik. Reactive Power consumed (Qrecdc) rectifier ,(Qinvdc) inverter,And Qac-

send, Qac-Rend drawn from both ends of line

%

Olac Uldc/cond.

• r f "

A lac+dc

-tr-

l y — ^ g "

e< 0.0 1.0 2.0 3.0 4.0 5.0 6.0

Fig. 5.111. AC, DC, Composite currents in the conductor of other healthy line.

3

102

I Fieli:! C wrent

sec 0.0 1.0 2.0 3.0 4.0 5.0 6.0

Fig. 5.1 Im. Generator's Field voltage and current of generator.

Case (4): Three Phase to Ground (ABC-G) Fault near Sending End at Transmission angle ofSOf:

I U_signal

sec 0.0 1.0 2.0 3.0 4.0

Fig. 5.12a Control Signal 5.0 6.0

103

Fig. 5.12b. Generator's speed deviation (Delta Omega).

3 a.

CPg • Qg

lSy*ss»m.,

- © ^

» e ^ — H »

sec 0.0 1.0 2.0 3.0 4.0 5.0 6.0

Fig. 5.12c. Generator's active (Pg) and reactive (Qg) power output.

250 ^ o Vs tms n Vr rms

sec 0.0 1.0 2.0 3.0 4.0 5.0 6.0

Fig. 5.12d. Sending (Vsjins) and receiving (Vrrms) end bus voltages.

104

CO O) a> en a> Q

Q Trans(n_Angle

70->

sec 0.0 1.0 2.0 3.0 4.0 5.0 6.0

Fig. 5.12e. Transmission angle ( Transm_Angle) vajiation.

73 >

OVdr

:5=iife?:ii==

• Vdi

-'^^O • Q

sec 0.0 ID 2.0 3.0 4.0 5.0 6.0

Fig. 5.12f. Rectifier (Vdr) and inverter (\'di) dc voltages.

o Idc re; 7.0 -,

;

.

Old inv

•

1 ' •H >•-**'•'"* V ' !

!

i

1

I

"1 i 1 1

sec 0.0 1.0 2.0 3.0 4.0 5.0 6.0 Fig. 5.12g. Rectifier (Idcrec) and inverter (Id_inv) dc current.

105

CO

cu Ol (U Q

180

160

140

120

100

80

60

OAOR • Aoi

rt»^r~i3 ^^}-^^ \r-^ -B-

sec 0.0 1.0 2.0 3.0 4.0 5.0 6.0

Fig. 5.12h. Rectifier (AOR )and inverter (Aoi) firing angle order.

in

180 160 140 120 100 80H 60 40 20 0

Gairmj-lnverter

sec 0.0

u*- y\

1 0 2.0 3.0 4.0 5.0 6.0

Fig. 5.12i. Inverter Gamma

opac J Hinvdc AH tranter

5.0 sec 0.0 1.0 2.0 3.0 4.0 5.0 6.0

Fig.5.12j.Ac (Pac), dc (Pinvdc) and total (Ptranfer) power transfer.

106

O Qrecdc n Qinvdc A Qec-Send • Qac-Rend

0.0 1.0 2.0 3.0 4.0 5.0 6.0

Fig. 5.12k. Reactive Power consumed (Qrecdc) rectifier ,(Qinvdc) inverter,> jid Qac-

send, Qac-Rcnd drawn from both ends of line.

aidc/cond A lac+dc

^^ 0.0 1.0 2.0 3.0 4.0 5.0 6.0

Fig. 5.121. Currents in the conductor of other healthy line.

107

3 a.

3 a.

6.0-1

5 .0-

4 .0 -

3.0-

2 .0-

1.0-

0.0-

• Field Voitage

— - —

y__ _ •Fielci Cjrrent

3.40-3.20-3.00-2.80-2.60-2.40-2.20-2.00-1.80-1.60-1

: . : [ - ;

K^^

1'

N K

V^

3ec 0.0 1.0 2.0 3.0 4.0 5.0 6.0

Fig. 5.12m. Generator's Field voltage and ciurent of generator.

It has been again demonstrated thiough the responses of Fig. 5.11 and Fig. 5.12 that the

fault near the sending end is more severe than the fault near receiving end/inverter in

simultaneous ac dc transmission in the case under study.

Case (5): Responses for Other type of Faults in the Simultaneous ac-dc transmission

Other type of faults such as single line to ground, double line to ground faults near

sending and receiving end v ere applied and studied as shown below for the same loading

conditions, sequence of events and fault durations as explained in three phase to groimd

fault case. The time based simulation responses are depicted in Fig. 13, Fig. 14 and

Fig.l5.

108

(i) Single Line to Ground (A-G) Fault Near Sending End at Transmission angle 60":

1.00

0.80

0.60 H

0 4G

0.20 -

0.00

-0.20

-0.40

' Ujsignal

^

s^^ 0.0 1.0 2.0 3.0 4.0 5.0

Fig. 5.13a Control Signal • Delta Cmega

6.0

Fig. 5.13b. Generator's speed deviation (Delta Omega).

OPg DQg

3

a.

sec 0.0 1.0 2.0 3.0 4.0 5.0 6.0

Fig. 5.13c. Generator's real (Pg)and reactive (Qg) power.

109

3 0 0 T

250

200

150-t-

100

50 4

O Vs rms • Vr_rms

ae=B:

2S<^ 0.0 1.0 2.0 3.0 4.0 5.0 6 0

Fig. 5.13d. Sending (Vsims) and receiving (Vr_rms) end bus voltages.

100 90 80

c Tr6n$m_Angle

0)

Q

7 0 - _[ en .^Tr4_ 60 50 4C--3 0 -2 0 -1 0 - -

0-1

sec 0.0 1.0 2.0 3.0 4.0 5.0 6.0

Fig. 5.13e. Transmission angle (Transm_Angle) variation.

ovdr u vai

§

- 4 0 0 - ^

sec 0.0 1.0 2.0 3.0 4.0 5.0

Fig. 5.13f. Rectifier (Vdr) and inverter (Vdi) dc voltages.

no

Old: r2c 7.0 -6.0 T

50 -

4.0-

3.0-

2.0-

1.0-

0.0-

-1.0-1

_^ ..

. .

/ S I I ^

1

1 •.:: •

• Id inv

; J

•

. --.

-

^^'^ 0.0 1.0 2.0 3.0 4.0 5.0

Fig 5.13g. Rectifier (Idcrec) and inverter (Idinv) dc current.

0) 0)

on

Q

180 T

160

140

120 4

100

80-f

60

40

20-f

0

OAOR 3 Aol

-3— • r — E } - \ - ^ 3 13-

y —»—

sec 0.0 1.0 2.0 3.0 4.0 5.0 6.0

Fig. 5.13h. Rectifier (AOR) and inverter (Aoi) firing angle order.

at a> i _ oi a> o

1 6 0 - r

140

1 2 0 -

1 0 0 -

8 0 -

6 0 - -

40

20

0

0.0

I Garrma-lnverter

1.0 2.0 3.0 4.0

Fig. 5.13i. Inverter Gamma order. 5.0 6.0

I l l

o Pac "J Pinvdc AP tranter

A—r- A - — A_, i--a B

- © - -«——e

sec 0.0 1.0 2.0 3.0 4.0 5.0 6 0

Fig. 5.13j. Ac (Pac), dc (Pinvdc) and total (P_tranfer)power transfer

2.5C C lac n idc / cond. A lac+dc

-e -

0.0D

sec 0.0 1.0 2.0 3.0 4.0 5.0 6.0

Fig. 5.13k. Currents in the conductor of other healthy line.

I Field VoHage

3 a.

3 a.

s e c 0.0 1.0 2.0 3.0 4.0 5.0 6.0

Fig. 5.131. Field Voltage and current response.

112

(ii) Double Line to Ground (AB-G) Fault Near Sending End at Tmnsmission angle 60":

0.0100 0.0080

0.0060

0.0040 •

0.0020 -

0.0000

-0.0020

-0.0040

-0.0060

-0.0080

-0.0100

I Detta Omega

ft

sec 0 0 1.0 2.0 3.0 4.0 5.0 6.0

Fig. 5.14a. Generator's speed deviation (Delta Omega). OPg GQcj

3 a.

1.75

1.50

1.25 4

1.00

0.75--

0.50 -

0.25

0.00

-0.25

-0.50 J

V.

—e- -•sc

sec 0.0 1.0 2.0 3.0 4.0 5.0

Fig. 5.14b. Generator's real (Pg)and reactive (Qg) power.

OVs rms • Vr rms

5

sec 0.0 1.0 2.0 3.0 4.0 5.0 8.0

Fig. 5.14c. Sending (Vsims) and receiving (Vr_mis) end bus voltages.

if) CD

cn CD

Q

113

120 -J

100 -

80

60

40 -

20 -

0

O Transm_Anqle

sec 0.0 1.0 2.0 3.0 4.0 6.0

Fig. 5.14d. Transmission angle (Transm_Angle) variation.

o Vdr n Vdi

•a >

sec

Fig. 5. 0.0 1.0 2.0 3.0 4.0 5.0

14e. Rectifier (Vdr) and inverter (Vdi) dc voltages.

7.0 -1

6.0 -

5.0 -

4.0 -

3.0 -

2.0 -

1.0 -

0.0 --1 n _<

o i d c

1

_re<

..__,

^__^

• Id inv

a —

—

0 » :

sec 0.0 1.0 2.0 3.0 4.0 5.0

Fig 5.14f. Rectifier (Idcrec) and inverter (Idinv) dc current.

114

0}

Q

OAOR lOU -1

160-

140-

120-

100-

80-

60-

40 -

20 -n J

'

-

QAoi

" E 3 — I - -U-

• - » -

sec 0.0 1.0 2.0 3.0 4.0 5.0 6.0

Fig. 5.14g. Rectifier (AOR) and inverter (Aoi) firing angle order.

CO 0) a> I

a> 73

1 80 -r- -

160-

140 - -

120-

100-

80-

60-

40-

2 0 - -

0 -

I Gatnma-lriv«;rter

0.0 1.0 2.0 3.0 4.0 5.0 6.0

Fig. 5.14h. Inverter Gamma order

sec

O Pac 3.0k-r

n Pinvdc A P tranter

A. A

—o—l^.c».

- 4 -i

-&•

0.0 1 0 2.0 3.0 4.0 5.0 6.0

Fig. 5.14j. Ac (Pac), dc (Pinvdc) and total (P_tranfer) power transfer

115

2.50 -. Olac

0.00 ->

C3 Idc / cond. A lac+dc

-a- "H « •

-»- -or

sec GO 1.0 2.0 3 .0 4.0 5.0 6.0

Fig. 5.14J. Currents in the conductor of other heahhy line.

3

a.

3 Q.

I Feld Current

5gQ 0.0 I.C 2.0 3.0 4.0 5.0 6.0

Fig. 5.14k. Field voltage and current responses.

116

(iii) Single Line to Ground (A-G) Fault in the Middle of the Line at Transmission angle 60°:

1.00 0.90 0.80 0.70 H 0.60 0.50 0.40 0.30 0.20 0.10 0.00

-0.10 H -0.20

sec 0.0 1.0 20 30 4.0 5.0 6.0

Fig. 5.15a. Control Signal.

0.0100-1

0.0080-

0.IDO6O-

0.0040-

0.0020

0.0000

-0.0020-

-0.0040-

-0.0060--

-0.0080

-0.0100

I Deta Cmega

sec 0.00 100 2.00 3.00 4.00 500 S.OO

Fig. 5.15b. Generator's speed deviation (Delta Omega).

3 £2.

1.50 n

1.25--1 nn .

0.75-

0.50-

0.25-

0.00-

-0.25-

-0.50-1

OPg

i

1 L

DQg

-&r-

__l

-.4

sec 0.0 1.0 2.0 3.0 4.0 5.0 6.0

Fig. 5.15c. Generator's real (Pg)and reactive (Qg) power.

117

300 ^

250

200

150

100

50 H

OVs rms • Vr rms

^ < 0.0 1.0 2.0 3.0 4.0 5.0 6.0

Fig. 5.15d. Sending (Vsjins) and receiving (Vr_rms) end bus voltages.

o Transm_Ancile

CO

Q

s^^ 0.0 1.0 2.0 3.0 4.0 5.0 6.0

Fig. 5.15e. Transmission angle (Transm_Angle) variation OVdr • Vdi

S

>

sec 0.0 1.0 2.0 3.0 4.0 5.0 6.0

Fig. 5.15f. Rectifier (Vdr) and inverter (Vdi) do voltages

l is

oidc rec • Id inv

g U H I Qq e us - © « •

^ • 0.0 1.0 2.0 3.0 4.0 5.0 ti.O

Fig 5.15g. Rectifier (Idc_rec) and inverter (Id_inv) dc current

180-T

160-

OAOR • Aoi

140-1"- I

120-

0)

Q

100

80

60

40 + —

20-f

0

1 -- n*,,-

/ ^

sec 0.0 1.0 2.0 3.0 4,0 5.0 6.0

Fig. 5.15h. Rectifier (AOR) and inverter (Aoi) firing angle order.

0)

a>

160 -

140-

100-

80-

60-

40-

20-

0 -

• Gamma-Inverter

-i

—

— .

—:

-U», J ! ^ -U ._. _

r

; 1

\

, i

;—. i

!

0.0 1.0 2.0 3.0 4.0 5.0 6.0

Fig. 5.15i. Inverter Gamma order.

119

O P a : • Pinvdc AP tranfer

sec 0.0 1.0 2.0 3.0 4.0 5.0 6.0

Fig. 5.15J. Ac (Pac), dc (Pinvdc) and total (Ptranfer) power transfer. C' lac • Idc / cond. A lac+dc

2.50

sec 1.0 2.0 3.0 4.0 5.0 6.0

Fig. 5.15k. Currents in the conductor of other healthy line.

3 a.

Z3 d.

6.0-1

4 .0 -

2 .0-

I Field Voltage

-4.0

-6.0

J: - 2 . 0 - - - « ^

I Field Current

sec 0.0 1.0 2.0 3.0 4.0 5.0

Fig. 5.151. Field Voltage and current responses.

6.0

120

In all these cases, the system is found to be stabilizing in duration, less than that for

the solid three-phase to ground faults considered in previous section.

Comparison of results between simultaneous ac-dc system and pure EHV ac system

for the similar nature and duration of faults indicate:

• The magnitude of generator speed deviation in case of simultaneous ac-dc

system is more. However, this speed deviation reduces to zero at a faster rate

as compared to that in pure ac system.

• At the inception of fault, the generator real power overshoot is lai-ger in case of

pure ac system. The subsequent real power oscillation magnitude as well as

duration is longer in pure ac system. While these oscillations decay much faster

in case of simultaneous ac-dc system.

• The pure ac system becomes oscillatory unstable at transmission angle of 49°.

When angle is increased to 50°, the system looses synchronism after first

swing.

• The PSS through exciter is unable to provide sufficient damping to prevent

generator stepping-out in pure ac system.

• The most important merit of the simultaneous ac-dc system is that it remains

stable even up-to very large transmission angle (80° in the present case) as

demonstrated in case (3) and case (4). However, in the rest of cases, the results

are demonstrated for transmission angle of 60° as the system delivers

maximum power at this angle. The steady state is restored just after the first

swing. In pure ac system, the transmission angle is generally kept low for long

transmission lines and seldom exceeds 30°.

5. 5 Conclusion

A novel approach to solve the first swing stability problem by simultaneous ac-dc

power transmission has been presented. A detailed study has been carried out in

121

PSCAD/EMTDC environment to validate the proposed scheme. It has been demonstrated

that the stabiUty of the system can be effectively improved by simultaneous ac-dc power

transmission with fast dc power modulation.

In pure ac system, the transmission angle is generally kept low for long transmission

lines and seldom exceeds 30°. But in simuhaneous ac-dc system we can go much beyond

30° transmission angle without loosing stability for transient disturbances. Satisfactory

transient performance is demonstrated for transmission angle upto 80° for the network

considered in the present study.

For the purpose of study, worst location of faults (i.e. near the sending end and near

inverter) and with heavily loaded line (i.e. close to thermal limit) is considered. It has

been demonstrated that effective control of dc component in simultaneous ac-dc power

transmission considerably improves the stability of the system.

122

CHAPTER 6

POWER TAPPING IN SIMULTANEOUS AC-DC POWER TRANSMISSION

6.1 Introduction

At present, almost about half of the world populations, especially those in developing

countries live without electricity. The supply of electricity is considered essential to avail

normal facilities of daily life. Its availability is fundamental for economic development

and social upliftment. Large power (steam, hydro, nuclear) stations ai-e usually located far

from load centers. The wheeling of this available electric energy from these remotely

located stations to load centers is achieved either by EHV ac or HVDC transmission

lines. These EHV ac.'UVDC transmission lines often pass over relatively small

communities /rural areas having no access to major power transmission network. It 's

most desirable to find methods of connecting these communities to the main transmission

system to supply cheap and abundant electric energy. However, HVDC transmission

system does suffer a significant disadvantage compared to EHV ac transmission, with

regards to tapping of power from ti-ansmission system.

Techno-economical reasons prevent the tapping of small amount of power from

HVDC transmission lines. This is considered a major drawback due to the fact that in

many instances HVDC transmission lines pass over many rural communities that have

little or no access to electricity.

In the past, a number of schemes have been proposed for small power tapping from

HVDC lines. Majority of them use forced commutated or line commutated inverters to

tap off the power from the HVDC system. These schemes inherently required additional

commutation circuitry or local generation capacity, which in turn lead to high cost of

installations and operational complications.

Ekstrom and Lamell [67] have proposed a scheme of small power tapping from HVDC

line based on a current source line commutated single-phase thyristor bridge, connected

in series with the HVDC line. To start the tap operation a local dc voltage source is

123

required in this scheme.

This chapter proposes a simple scheme of small power tapping from the simultaneous

ac-dc power transmission line along its route. The tapping stations aie assumed to draw

power up to 10% of the total power transfer capacity of the composite line.

6.2 Small Power Tapping Station Requirements

The main requirements of a small power tapping stations are [67]:

• The per unit cost of the tap must be strongly constrained, i.e., the fixed cost must

be kept as low as possible.

• The tap must have a negligible impact on the reliabilit}' of the ac-dc s>'stem. This

implies that any fault in the tap must not be able to shutdown the whole system

and,

• The tap controls should not interfere with the main system, i.e. the tap control

system has to be strictly local. Failure to achieve this leads to complex control

system requirement and thus higher cost of hardware.

• Small tap stations having total rating less than 10 percent of the main terminal

rating has potential applications where small, remote comrnimities or industries

require economic electric power.

The tapping stations considered in this study are of fairly small power rating, up to 10

% of the total transfer capacity of the simultaneous ac-dc power transmission line. Short

interruption of the power supplies should be tolerable at the occurrence of temporary

earth faults on the main simultaneous ac-dc power transmission system. Further, any fault

occurring within tapping station and its local ac network is to be cleared by local CBs

provided. These tapping stations will not depend upon the telecommunication links with

the main composite ac-dc transmission system.

5.3 The System under Study

The networic depicted in Fig. 6.1 has been taken up for the feasibility of small power

tap for remote commimities from the simultaneous ac-dc power transmission system. A

124

s>Tichronous machine is shown to be dehvering power to an infinite bus via a double

circuit three-phase, 400 KV, 50Hz, 450Km ac transmission line. The line considered is

converted to simultaneous ac-dc tr.ansmission line with ac rated voltage of 220 kV and dc

voltage of ±320 kV. This network is same as described in chapter 5, Fig.5.1, for

transient stability study except that in present case two separate power tapping stations

are added to each line of the double circuit line, directly and separately (may be at same

location or at different location along the line) as shown in the Fig. 6.1. Since line-to-line

voltage in case of simultaneous ac-dc transmission carries no dc component, :he line-to-

line voltage is a pure ac voltage as mentioned in chapter 3 and has been demonstrated in

Fig. 6.2. These voltage waveforms indicate neither presence of dc voltage component nor

distortion in waveform though the5;e are derived from the composite line. To tap ac power

from the line, transformer can be directly connected between two conductors of the line

without breaking them.

In this study of simuhaneous ac-dc transmission line, ac line voltage component has

been selected as 220 kV. Each tapping station transformer (rated as 120 MVA, 220/66

kV, A-Y) is connected to the local ac load via circuit breaker (CB). These CBs are

provided for local protection, to clear the fault within the local ac network. The nature of

local load considered here is that of summer time-residential class with following

characteristics [1],

• power factor=0.9,

• voltage index for power((3P/5r)=1.2,

• voltage index for QidQ/dV) =2.9,

• frequency index for power (dP I df) =0.8,

• frequency index for Q (dQ I df) = -2.2.

1?5

m I 1 ? 1

>

^

£ ^

o 1 H'

c a

•o

6

o a) c

Oi

o

CD

'5-

o a.

CD

126

OEad • Etact A Beat

^\/;sGz!iz:iQ^^\/^\

1.570 1.580 1,590 1.600 1.610 1.620 1.630 1640 1.650 1.660

Fig. 6.2. Line-to-line voltages among three phases.

6.4 Digital Simuiatlon of the proposed Scheme

In order to examine the feasibilit}' of the proposed scheme for power tapping, and to

observe the performance of the simultaneous ac-dc power transmission system under

various operating conditions, the digital simulation software package PSCAD/EMTDC

was used. The initial operating conditions of the simultaneous ac-dc power transmission

system before the tapping power is switched on are as follows.

AC power at receiving end, Pac= 603.735MW,

DC power at receiving end, Pdc=l 560.183 MW,

Total Power transfer at receiving end, PtotaI_transfer= 2162.9 MW.

Transmission angle =60**

Case A. Equal Power Tapping from each line at different instants

The system is initially assumed to be operating at above conditions delivering the

scheduled real power to infinite bus, at time, 0.5 s, a load of 50MW was switched on by

closing CB of one tapping station transformer which is directly connected to one of the

double circuit line located in the mid way i.e. at 225km from the sending end. Again at

time, 4.5 s, another load of 50 MW was switched on by closing CB of second tapping

station transformer which is connected directly to second line. The transient responses for

these conditions are shown in Fig. 6.3.

127

•s.

o CL 75

2.0k

1.8k

1.5k

1.3k-

1.0k-

0.8k

0.5k-

OPtI ciPac A Pacs • Pinvdc

0.3k-

0.0

-it—

•is -T-rtr--e——fi-

-«——2ir-

_a e ^o-

e e 0 -^i :—e 9- jo o e ©^—e B e—^-e a e-

sec 0.00 1.00 2.00 3.00 4.00 £.00 6.00 7.00 8.00 9.00 " 10.00

Fig.6.3a. Tap power (Ptl), receiving end (Pac) ac powers, sending end (Pacs) ac powers, and receiving end dc power (Pinvdc)

Active Power Transfer _ i

O Pac n Pinvdc A P_tranfer 2.3k

2.0k-

1.8k-

1.5k

i>

/s*—IX- - j V -

1.3k-|--^—.--^i-...i-

1.0k

0.8kH 1 — - , — - 1 - . . ^ - .

0.5k

0.3k

0.0

• i » I O' -B—pe-;—e ;©• •:h&Z77.

—el . ; .•... - e 3 — T — a -

l"

,q_4-- Ql P -, - P . c>—-—9-

- i J i_-

sec 0.00 ' 1.00 " 2.00 ' 3.00 " 4.00 " 5.00 ' 6.00 ' 7.00 " 8.00 ' 9.00 10.00

^lii»33B88ililg |gi^JHggJiilB»iglJiaiillJjigPiiflP£ ^'r' I •!

Fig. 6.3b. Ac (Pac), dc (Pinvdc) and total (Ptranfer) power transfer at receiving end.

128

o Qrecdc c Qinvdc

-1.0k-i

sec 0.0 2.0 4.0 6.0 8.0 10.0

Fig. 6.3e. Qrecdc, Qinvdc Reactive power drawn by rectifier and inverter, Qac-send, and Qac-Rend, reactive power drawn by the lines from each ends.

The Fig. 6.3a and Fig. 6.3b show the time-based responses of real power delivery by

the system to tapping station and xo receiving end. These indicate that whenever the tap

station is switched on it draws ac power from tlie system. The receiving end ac power is

reduces by the same amount. The total power deliver}' by the line at receiving end is also

reduced by the same amount. The time-based responses of reactive powers drawoi by the

converters and lines are shown in Fig. 6.3c. It indicates that the lines draw additional

reactive power from the sending end to meet the tap station reactive power requirement.

Bus Vottaqe

• Vr_rms OVs rms

te^p &Q—B- ?fi:r

sec 0.00 1.00 2.00 3.00 4.00 5.00 6.00 7.00 8.00 9.00 10.00

Fig. 6.3d. Sending (Vs_rms) and receiving (Vr_nns ) end bus voltages.

180-

160

140

120

100

80

60 •

40-

20-

0

OAOR • Aoi

-e !—B-

«a o '^ o o u>i ( o t> o | i j a f3 w

*®«= 0.0 2.0 4.0 6.0 8.0 10.0

Fig. 6.3e. Rectifier (AOR) and inverter (Aoi) firing angle order.

129

450-r 400-

3 5 0 - '

OVdr • Vdi

i 73 >

300-

2 5 0 - -

200-

150-

100-

5 0 -

0

sec

05

Q

sec

f9- !• • O S f-iJ—!—Q 19-

O " " - g -i9 a——e e~ -e-—e~

>-Q-t"

! T I — 1 .

-a :

0.00 1.00 2.00 3.00 4.00 5.00 6.00 7.00 3.00 9.00 mOO

Fig. 6.3f Rectifier (Vdr) and inverter (Vdi) dc voltages.

OTransm Anale 8 0 T - • - - . - ; - ^ •? i

70

60

50

40

o a -e-

30 J

- e — Q o o - e — r e ©• - e a e e -

0.0 1.0 2.0 3.0 4.0 5.0 6.0 7.0 8.0

Fig. 6.3g. Transmission angle (TransmAngle). 9.0 10.0

130

1.60 -r

1.40

1.20 +

1.00

0.80 -h

0.60

0.40 +--

O P g n Qq

0.20

0.00 -"-

^Mgn . i 3 13- - e - -o- a esi

- a - —fa-

s e c -I 1 —

2.0 1 r-

4.0 0.0 2.0 4.0 6.0 8.3 1 D.O

Fig. 6.3h. Generator's active (Pg) and reactive (Qg) power output.

Generator Speed Diffrence

0.0100 n

0.0050-

O.uOUO -

-0.0050 -

.0 0100 -

• Delta Omega

1 i , ;

1 : i

! I i

1 • • '

•

; 1 I

2.0 4.0 6.0 8.0 10.0

Fig. 6.3i. Generator's speed deviation (Delta Omega).

The above time-based simulated responses of various system parameters show no

deterioration of the performance of ac-dc composite system except for the small transient

period at the time of switching on of tap stations.

Case B. Unequal Power Tapping from each line at different instants

The system is initially assumed to be operating at above conditions delivering the

schedtiled real power to infinite bus. At time, 0.5 s, a load of lOOMW was switched on by

closing CB of one tapping station transformer which is directly connected to one of the

131

double circuit line located mid way i.e. at 225kin from the sending end. Again at time,

3.5 s, another load of 50 M^' was switched on by closing CB of second tapping station

transformer which is connected directly to second line. The transient responses for these

conditions are shown in Fig. 6.4.

OR1 C3Fac APacs «• Pinvdc i 2.Oka

1.8k-

1.6k

1.4k4-

1.2k

1.0k

0.3k-

O.Sk-'

0.4k- -

0.2k-

00-

a3

o Q.

- i » - — ^

•or ""-tr - T S - — ' 1 ^ " "

"VJ-- 1 5 -

sec 0.00 0.50 1.00 1 50 2.00 2.50 3.00 3.50 4.00 4.50 5.00

Fig. 6.4a Tap power (Ptl), receiving end (Pac) ac powers, sending end (Pacs) ac powers, and receiving end dc power (Pinvdc)

5 s

2.5k-.

2.3k-;

2.0k-

1.8k-

1.5k-

1.3k

1.0k

0.8k

0.5k

0.3k

0.0

opac

Active Power Transfer

• Pinvdc APtotal transfer

is-..— ijs^^ir- - I t -

- e —

z:t::3:: --r-'

• i z r ? ^ . , ^ . rrSt.

T

r S p • • •D D . mr-

f-• 9 =

1

i

I ::ev—-e-

i

-^H

2 ^ 0.00 0.50 1.00 1.50 2.00 2.50 3.00 3.50 4.00 4.50 5.00

Fig. 6.4b. Ac (Pac), dc (Pinvdc) and total (Ptranfer) power transfer at receiving end.

132

1.3k-r

1.0k-'

0,8k--

0.5k

0.3k-

Keactive Power converter

o Qrecdc n Qjnvdc A Qac-Send • Qac-Rend

<»• - o —i j ' " " »——e o ^ *^ j» •i-'s ' Q

0.0 - r~=f :

-0.3k

J * - -TiK -•b- i!r-

-0.5k-

- 0 . 8 k 4 ^ , _ . _ ; : e — ^ ^ -

-1.0k' ^ f ^ - ^ 3 -

s e c 0 0 1.0 2.0 3.0 4.0 5.0

Fig. 6.4c. Qrecdc, Qinvdc Reactive power drawn by rectifier and inverter, Qac-send, And Qac-Rend, reactive power drawn by line from each ends.

Bus VoHaqe

O Vs rms n V'r rms 2 5 0 - r -~ - ~

200

150-

100-

5 0 -

0 -

n " trs'-^'^'=§^^^>w=^t=f^^—ig—Qs~g—D^'-'^^''=^g-H—c^-g^=

sec 0.0 1.0 2.0 3.0 4.0 5.0

Fig. 6.4d. Sending (Vsrms) and receiving (Vrrms) end bus voltages.

Q

180

160

140

120

100 -

80

60

40 -

20

O

O AOR Converter Aloha Order

• Aoi

—a B i -B- -B — Q -

o s.—(J o • 6 o ' o o e o |_ .Q i o— p <»'

0 00 1.00 2.00 3.00 4.00 5.00

Fig. 6.4e. Rectifier (AOR ) and inverter (Aoi) firing angle order.

133

OVdr 450 T

400

Converters Voltaqe

DVdi

350 - --B—

300

250

200 --

150-

1 0 0 - -

5 0 - -

0

-e-T*'>»* i e—_o o q g...ig g^T'^y'..o ..p. c': a e B-! B -—H --E3 E3—

sec 0.0 1.0 2.0 3.0 4.G

Fig. 6.4f. Rectifier (Vdr) and inverter (Vdi) dc voltages.

5.0

3 a.

1.60 1

1.40 - -

1.20

1.00 t

0.80-

0.60-

0.4C

Q.20

OPg DQg

ei<»—vs. ^ gD a e e e <g>^-'0 e e-

-e— H 2 ,

0.00 -I

®®^ 0.0 1.0 2.0 3.0 4.0 5.0

Fig. 6.4g. Generator's active (Pg) and reactive (Qg) pov\er output.

Generator Soeci Ditfrence

0.0010 -r

0.0005

O.OOOG

-0.0005

-0.0010

sec o 5 "

I De>a Ome-aa

L

1.0 2.0 3.0 4.0 5.0

Fig. 6.4h. Speed deviation of generator (Delta Omega).

134

The above time-based simulated responses of various system parameters show no

deterioration of the performance of ac-dc composite system except for the small transient

period at the time of switching on of tap stations.

Case C. Tapping Power Station Fault Responses

(i)Three-Phase to earth (ABC-G) Fault between Taf> Transformer and Load

Initially, the simultaneous ac-dc system with small power tap is cissumed lo be

operating in its steady state with the following conditions.

Equal Power Tap from each line from the mid way of the line. Pt=(50+50) = 100 MW.

Receiving end ac power, Pac==542.MW.

Sending end ac power, Pacs=694.62MW,

DC power at receiving end. Pdc^l558MW,

Transmission angle =60°

At time, t= 0.5 s a solid three-phase to ground fault occurs between tap transformer

and the local load. After a period of three cycles (60ms), trip signal is given to the local

circuit breaker to clear the fault. Thereafter, CBs are reclosed after a delay of three cycles

(60ms) from the instant of clearing fault. Fig. 6.5 shows the transient responses for this

fault condition.

I cc

2.0k-

1.8k-

1.5k-

1.3k-

1.0k-

0.8k-

0.5k-

0.3k-

opt-: D P a c APacs ' Pinvdc

0.0 -• " ""k'

— .. . i . - ..

• ^

"

- • ! - - -

..-.i......

„ . .p . .

°

- -•- -

1

! A

°

-

„ . ^ ...- Q i

0.00 0.50 1.00 1.50 2.00 2.50 3,00 3.50 4 0 0 4.50 5.00

Fig.6.5a. Tap power (Ptl), receiving end (Pac) ac powers, sending end (Pacs) ac powers, and receiving end dc power (Pinvdc)

2.3kn

2.0k

1.8k

1.Sk

1.3k

1.0k

0.8k

0.5k

0.3k

0.0

O Pac D Pinvdc

ts— -iS-

A P _tranfef

-his 1—ss is-

^rv , - - - i3 - "^Tr^T'^ -»-t-

Q \/*^'-• 0\ _o Q — e - -©- "•€):-; © -e' -Q Q-

135

s e c 0.00 0.50 1.00 1.50 2.00 2.50 3.00 3.50 4.00 4.50 5 00

Fig. 6.5b. Ac (Pac), dc (Pinvdc) and total (P_tranfer) power transfer at receiving end.

OEact • Ebct AEcat

0.500 0.550 0.600 0.650 0.700

Fig. 6.5c. Line voltages variations across at CB's input. Tap Transformer Primary Currents

O High Side Current, A a High Side Current, B A High Side Current. C

0.750

3.00-

2.50-

-t.UU *

1.00-

0.50-

U.UU 1

-3.00-

In

if

fif)i i l vC t

i l

I '

jl i !

?l M I " IT

i

1 i

[. !

1 i

j - i r - r - r - T — j

'

1

J

—r-7-

r • ! --

^ i 1

\ A. '^

1

' r

1 A O I (

0.500 0.600 0.700 0.8i30 0.900 1.000 1.100

Fig. 6.5d. Transformer input line currents variations measured at CB's input.

136

Anqles

O Transm Anqle 8 0 T ^ L . • • - • . -

70-

60

S' 50-f Q

40

30

-f^^^/^/y^^-^er-^ o 'o - 9 e e o o — - e e ©-

sec 0 0 1.0 2.0 3.0 4.0 5.0

Fig. 6.5e. Transmission angle ( TransmAngle ).

s:.

250

200 4

150

100 4

50

0-"

OVs rms C Vr rms

sec 0.0 1.0 2.0 3.0 4.0 5.0

Fig. 6.5f. Sending (Vs_nns) and receiving (Vr_rms) end bus voltages.

180 -t-

160

140

120 -

100

80

60

40 -{—

20

0

O AOR

Converter Alpha Order

• Aoi

J • r f v -

•g-:I\^\/'i^^C^^?t''-!'?l_—o_: g

sec 0 00 1.00 2 00 3.00 4.0O 5.00

Fig. 6.5g. Rectifier (AOR )and inverter (Aoi) firing angle order.

137

>

ovdr 450 T

400

350-F

300

250- -

200-

150-

100-

50-

0

Converters: Vottaqe

• Vdi

^^'^ 0.0 1,0 2.0 3.0 4.0 5.0

Fig. 6.5h. Rectifier (Vdr) and inverter (Vdi) dc voltages.

0.0010 -r

0.0008-

0.0006 - -

0.0004--

0.0002- -

^ 0.0000 —

" -0.CO02--

-0.0004 - -

-0.0006 - -

-0.0008 - -

-0.0010--

sec '~

I Delta Omeqa

V v" '*

0.0 1.0 2.0 3.0 4.0 5.0

Fig. 6.5i. Generator's speed deviation (Delta Omega).

1.60

1.40

OPg iQg

.20--

3 £2.

0.60

0.40

0.20

0.00 ->

1 00 - —oiAAP;'V'^g^ig^^»"~^"a—tr*

0.80

'0—0—g"

a "'igr T sec 0.0 1.0 2.0 3.0 4.0 5.0

Fig. 6.5j. Generator's active (Pg) and reactive (Qg) power output

138

Generator Field

I Field Voltage

3 a.

6.0

4.0 r

0.0 [-

.2.0 ; .

-4 .0 - ••

.6.0 J - ^

3 a.

2.50 j

2.00 -"

1.50 - -

1.00 - -

0.50 -

0.00- •

I Field Current

0.0 1.0 2.0 3.0 4.0 5.0

Fig. 6.5k. Generator's Field Voltage and current.

(ii) Single-phase to earth (A-G) Fault between Tap Transformer and Load

Now considering similar initial conditions as discussed in case (i) but the fault

considered is solid single-phase to ground type instead of three- pha^e to ground with the

sanie sequence of events.

Various transient responses for this fault condition are shown in Fig. 6.6.

a:

2.0k

1 8k

1 5k

1.3k

1.0k

0.8k

0.5k

0.3k

0.0

-0.3k J

OPt1 n p a c A Pac:s • Pin>/cic

-e—«.« ,<—»•

T - * -

0.0 1.0 2.0 3.0 4.0 s° Fig. 6.6a. Tap power (Ptl), receiving end (Pac) ac powers, sending end (Pacs) ac powers, and

receiving end dc power (Pinvdc)

139

2.5k T

2.0k

1.5k-'

1.0k-

oPac

Active Power Transfer

Q Pinvdc A Plotal transfer c

!

1

. , , ' I A i i

0.6k- -Ci.iv»——6—r^-e-r—(».• J"O., . o o , o — o , a—;-3——a a-

0.0 -

-0.5k

sec QOO 0.50 1.00 1.50 2.00 2.50 3.00 3.50 4 00 4.50 Tx)0

Fig. 6.6b. Ac (Pac), dc (Pinvdc) and total (Ptranfer) power transfer at receiving end.

OEact AEcat

0.500 0.550 0.600 0.650 0.700 0.750

Fig. 6.6c. Line voltages variations across at CB's input.

O High Side Current, A Cl HIgli Side Current, B A Hiqii Side Current, C

0.800

sec 0 500 0.600 0.700 0.800 0.900 1.000 1.100

Fig. 6.6d. Transformer input line ciirrents variations measured at CB's input.

140

o v s rms

Bus Voltaqe

n Vr rms

0 -i

sec 0 0 1.0 2.0 3.0 4.0 5.0

Fig. 6.6e. Sending (Vsrms) and receiving (Vr_nns ) end bus voltag,es.

at Q

130 -r

160 -

140

120

100 -

80 -

60

40 -

20 -•

O

O AOR

Converter Alpha Order

a Aoi

i 1...

HsAPy*'

1

! i-

i r • 1 -

0.00 1.00 2.00 3.00 4.00 5 00

Fig. 6.6f. Rectifier (AOR )and inverter (Aoi) firing angle order. Converters Voltaqe

OVdr DVdi

S

>

sec 0.0 1.0 2.0 3.0 4.0

Fig. 6.6g. Rectifier (Vdr) and inverter (Vdi) dc voltages. 5.0

141

Generator Speed Dif'frence

0.0010-r

0 .0005 -

I Delta Ome ja

Z3

a.

0.0000

-0.0005

-0.00)0-L

sec

-M^v.

0.0 1.D 2.0 3.0 4.0

Fig. 6.6h. Generator's speed deviation (Delta Omega). O Pg • Qg

1.60-r-

1,40

1.20

3

1 00 - —oj^^i^cu-osg^

0.80 -

0.60-

0.40-

0.20

0.00-i

-€)~

. -B r^.-s-'^^^^'-B^ = -a-

s®< 0.0 1.0 2.0 3.0 4.0 5.0

Fig. 6.6i. Generator's active (Pg) and reactive (Qg) power output.

142

Generator Field

6.0 J 4.0-1-2.0

' Field Voltage

3 0.0

-2.0 -F

-4.0

-6.0 J

3 O.

2.50 -r

2.00 - -

1.50 -

1.00 -

0.50 -

0.00 - ~

" Field Current

CO 1.0 2.0 3.0 4.0 5.0

Fig. 6.6j. G(;nerator's Field voltage and currenx.

The Fig. 6.5 and Fig. 6.6 are the responses after the occurrence of faults in local ac

network of small tapped power for rural communities from the simultaneous ac-dc

transmission system. These local faults cause only minor disturbance for short transient

period.

6.5 Conclusion

The simulated results show that small amount of power, (up to 10% of total transfer

capability of the simultaneous ac-dc transmission system), to feed remotely located

communities, can be tapped in the same simple way as tapping in case of EHV ac line. It

is also economical as compared to complicated methods of tapping from HVDC line. The

resuhs clearly indicate that the tapping of small amount of ac power from the

simultaneous ac-dc transmission line has negligible impact on the dc power fransfer.

143

CHAPTER 7

CONCLUSION

7.1 Conclusion of the Present Work

Long EHV ac lines cannot be loaded to their thermal limits in order to keep

sufficient margin against transient instability. The transmission angle between sending

end and receiving end seldom exceeds 30". HVDC transmission lines in parallel with

EHV ac lines are recommended to improve transient and dynamic stability as well as

to damp out oscillations in power system.

The present work demonstrates that it is possible to load EHV ac line closed to its

thermal limit by allowing dc component along with ac component in th(j same

conductor of line. Simulation as well as experimental results establishes the feasibility

of simultaneous ac-dc power transmission. Zig-zag connection of secondary windings

of transformer at both end is used to avoid saturation of core due to dc current. Tv 'o

fluxes produced by the dc current (Id /3) flowing through each half of a winding in

each limb of the core of the transformer are equal in magnitude and opposite in

direction. So the net dc flux at any instant of time becomes zero in each limb of the

core. Thus the dc saturation of the core is avoided.

A double circuit EHV ac line is converted to simultaneous ac-dc power

transmission line requiring no alteration in the original transmission line. The higher

creepage distance insulator discs as needed in case of HVDC lines is no more required

in the suggested composite ac-dc line because electric field produced by conductor to

ground voltage reverses its polarity twice in a cycle. The voltage levels are set such

that minimum ac phase to ground voltage and maximum dc voltage with respect to

ground of the converted composite ac-dc line are 1/2 and l/v2 times the phase to

ground of the original EHV ac line in order to insure voltage reversal.

The major advantage of the composite ac-dc line is that transmission angle between

two ends of line in steady state may go up to very large values (80° in the system

144

considered). This is because of inlierently fast controllability of dc link embedded in

the composite ac-dc line. It is found from the study that the composite ac-dc line

delivers maximum real power transfer at receiving end at a transmission angle of 60°.

Such a large transmission angle is impossible to achieve in pure EHV ac transmission

line. It has been shown that for the composite ac-dc line length of 450 km, there is

substantial increase (about 83.45%) in the loadability of the line as compared to EHV

ac line. Further studies show that the power upgrading of the existing ac line gets

doubled or more than double with the simultaneous ac-dc power flow for a line length

of 500 km or more than 500 km line.

A novel approach to solve the first swing stability problem by simultaneous ac-clc

power transmission has been presented. A detailed study has been carried out to

validate the proposed scheme. It has been demonstrated that the stability of the system

can be effectively improved by simultaneous ac-dc power transmission with fast dc

power modulation. Satisfactory tiansient performance is demonstrated for transmission

angle up to 80° for the network considered in the present study. Fcr the pur]3ose of

study, worst location of faults (i.e. near the sending end, near inverter) and with

heavily loaded line (i.e. close to thermal limit) is considered. It has been demonstrated

that effective control of dc component in simultaneous ac-dc power transmission

considerably improves the stability of the system.

Though phase to neutral voltage in case of simultaneous ac-dc transmission has a

dc offset, the line-to-line voltage carries no dc component. In simulation studies the

voltj^e waveform indicates neither presence of dc voltage component nor distortion in

waveform though it is derived from the composite ac-dc line. To tap small ac power

from the line, transformer can be directly connected to the conductors of the line

without breaking them. The tapping transformer primary input currents exhibit no dc

component though these are drawn directly from the composite ac-dc line.

It has been demonstrated that it is possible to tap small amount of power, (up to

10% of total transfer capability of the simultaneous ac-dc transmission system), to feed

remotely located communities along the line, in the same simple way as in case of

145

EHV ac line. It is also economic£.l as compared to complicated methods of napping

from HVDC line. The results clearly indicate that the tapping of small amount of ac

power from the simultaneous ac-dc transmission line has negligible impact on the dc

power transfer.

7.2 Scope for Future Work

As results of the investigations carried out in the present work, a number of ideas

ha\'e emerged which can be pursued further extension of the work.

Following issues may be interesting topics to be taken as further extension of the

present work.

• Detailed cost analysis of simultaneous ac-dc power transmi ssion.

• Design of suitable protective scheme for simultaneous ac-dc power

transmission.

• Analysis and detailed st?jdy of sub-syncltronous resonajice (SSR) due ro

presence of HVDC controllers in simultaneous ac-dc power transmission.

• Extensive field-testing on actual physical system.

• To explore the possibility and nature of transient dc offset on ac side.

• As the idea is new, there may be large number of other related issues, which

need to be further investigated.

146

LIST OF PUBLICATIONS FROM PRESENT WORK

The following is the list of contributions by the author related with present work:

I. "Power Upgrading of Double circuit ac Transmission Line By Simultaneous ac-

dc Power Transmission" The IEEE Proceedings, PES, Power India

Conference, 2006, New Delhi, India, paper no.330:

II. "Feasibility Study of Conversion of Double circuit ac Transmission Line for

Simultaneous ac-dc PoAver Transmission" The IEEE Proceedings, PEDS 2005,

Conference, Kaulampur, Malaysia, pp 972-976.

III. "Enhanced Power Transfer by Simultaneous Transmission of i'\.C-DC: a new

FACTS Concept" The lEE Proceedings, PEMD 2004, Edinburgh. UK,

Publication No. 498, Vol. L pp: 186-191.

IV. "Possibilities of Simultaneous AC-DC Power Transmission", All India Jamia

Electrical Engg. Conference Proceedings (JEEC-2003), The lEE (Delhi

chapter), New Delhi, .pp: 35-42.

REVISED MANUSCRIPT SBMITTED:

V. "Power Upgrading of Transmission Line by Combining ac-dc Transmission"

IEEE Transaction on Power System, Paper ID NO. TPWRS-0864-2005.

VI. " Transient Stability Improvement of Power System By Simultaneous AC-DC

Power Transmission" IEEE Transaction on Power System, Paper ID NO.

TPWRS-0883-2005.

147

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[4] N.G. Hingorani,and L.K. Gyugyi "Understanding FACTS - Concept and Technology

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[5] K.R. Padiyar, "Power System Dynamics Stability and Control" BS Publications,

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[13] Clerici A, Valtorta G and Paris L, "AC and/or DC substantial power upgrading of

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148

[15] Gholipour E.,and Saadate S., "ImDroving Transient Stability of Pov/er System

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pp. 1677-1681.

[16] Hammad A.E., "Stability ajid Control of HVDC and AC Transmission in Paralkr'

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[17] Daneshpooy A et al, "Fuzzy Logic Control for HVDC Transmission" IEEE

Transaction on Power Deliver)', Vol. 12, No.4, October 1997, pp. 1690-1697.

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APPENDIX

A.l A B C D constants A, B, C, D constant of the long ac transmission line may be computed as:

A = cosh}',l, B = Z^5/nhy^]

C=—sinhy,! , D = cosh>',l

;', = slzy = Propagation constant

z.= , Characteristic impedance of the line vy

A.2 DC Converters Terminal Equations

The Graetz bridge circuit has been universally used for HVDC converters as it

provides better utilization of the converter transformer and lower voltage across the valve

vv'hen not conducting.

The representation of the converters is based on the following assumplions:

a) The direct current Id is ripple free.

b) The ac system at the inverter and the rectifier consist of perfectly sinusoidal,

constant frequency, balance voltage sources behind balanced impedances.

c) The converter transformers do not saturate.

For rectifier on constant current control (CCC) and inverter at constant extinction angle

(CEA) control mode:

Inverter equations are written as:

V = ^ B T E

V<„=V^„(cosY^)--X,B.I„,

As y^„ and lord are known. The transformer turns ratio Ti can be adjusted to give Vdi

within the desired range.

155

Rectifier equations are as:

K

a - cos ' ' V X I

Vdro V2E3„T,

AC current draA\-n by convener. T ; -— T, •' 3C d

Voltage £ ^ is known; the iransformer turns ratio T is adjusted to give a within the

desired range.

Converter transformer rating. Si = \ ^ T E ^ I (rated)

A.3 Parameters of the power s 'stem used as an application example

Synchronous Generator: Parameters in p. u. on its own base:

Direct axis svnchronous reactance, Xd =1.81,

Quadrature axis synchronous reactance, Xq=1.76,

Direct axis transient reactance Xd'=0.3,

Direct axis sub-transient reactance, Xd"=0.23,

Quadrature axis subtransient reactance, Xq"=0.25,

Stator leakage reactance, Xi =0.15,

Stator resistance, Ra=0.003,

Direct axis transient open circuit constant, Tdo'=8.0 s.

Direct axis subtransient open circuit constant, Tdo"=0.03 s.

Quadrature axis subtransient open circuit constant, Tqo"=0.07 s.

Inertia constant, H=3.5 s

Excitation System:

KA=200.0, TA=0.015, TC=1.0, TB=12.0, VRMAX=5.64,

VRMIN=-4.53, KC=0, VIMAX=1.0, VIMIN=-1.0

Generator Transformer (GT):

Positive sequence leakage reactance=0.15 pu.

No load loss=0.05 pu, Copper loss==0.05 pu.

156

Converter Transfonners (CT):

Positive sequence leakage reactance=0.10 pu

Rectifier Transformer, 220/150KV, 1200 KVA

Inverter Transformer, 220/133KV, 1025 KVA

DC Links:

Rectifier firing angle (a) limit (minimum) =5°

Inverter extinction angle (y ) limit (minimum )=15°

DC current=5.4kA (Maximum)

Smoothing Inductor= 0.6 H

AC Filters:

11'** Hormonic C-15.0^F, L=0.0055582H, R=0.480a

13"" Hormonic C=15.0nF, L=0.004H, R=0.408 Q

Power System Stabilizer:

^T I I ^ 1 + sT, I

Gl=1.0.Tw=0.05s,

Lead Time Constant=0.248 s,

Lag Time Constant==0.062 s.

A.4 Initial simultaneous ac-dc steady state conditions for Chapter 5:

Generator, Pg= 0.99 pu, Qg=0.233 pu, Vg=l.lpu

Ac transmission voltage Vac=220kV,

Dc transmission voltage Vd=160 kV,

i.e. Bipolar DC Voltage Vdc =±320 kV,

Receiving end power, ac power (Pac)=503 MW,

dc power (Pdc)=1604 MW,

total power (Pr)=Pac+Pdc=2107 MW,

Initial steady state conditions for pure ac system with compensation:

Transmission Voltage 400 kV

Generator, Pg=0.995 pu, Qg=0.133 pu. Vg^l.l

Receiving end power Pr=2100 MW.

Capacitance value for 50% compensation, C=42.78microF