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
REFERENCES
[I] p. S. Kundur, ^""Power system stability and control". New York: McGrawhill Inc.,
1994.
[2] E.W. Kimbark, ''Direct Current Transmission", Vol-L Wiley, New York, 1971.
[3] K.R. Padiyar, "HVDCpower transmission system", Wiley Eastera, New Delhi, 1993.
[4] N.G. Hingorani,and L.K. Gyugyi "Understanding FACTS - Concept and Technology
of Flexible A.C. Transmission Systems", IEEE Press, 2000
[5] K.R. Padiyar, "Power System Dynamics Stability and Control" BS Publications,
Hyderabad, 2002
[6] M.A Pai., D.P.S Gupta and K.R Padiyar. "Small Signal Analysis Power System",
Narosa Publishing House, New Delhi, 2004
[7] J. Arillaga and N. R Watson., Computer Modelling of Electrical power systems, John
Wiley & Sons, England, 2003.
[8] Uhlmaiin, "Power Transmission by direct Current" Springer-Verlag Berlin
Heidelberg New York 1975.
[9] Hadi Saadat, "Power System Analysis", Mc Graw Hill, 1999.
[10] PSCAD/EMTDC, "User's Guide", Manitoba-HVDC Research Centre, Manitoba,
Canada, January2003.
[II] Szechtman M., Wees T. and Thio C.V, "First Benchmark Model for HVDC
Control Studies" Electra.No.l35, April 1991.
[12] Clerici A, Paris L and P. Danfors, "HVDC conversion of HVAC line to provide
substantial power upgrading," IEEE Transaction on Power Delivery, 6(1), 1991,
pp.324-333.
[13] Clerici A, Valtorta G and Paris L, "AC and/or DC substantial power upgrading of
Existing OHTL Corridors," C DC Conference, London, 1991, pp.220-225.
[14] Basu K. P. and Khan B. H., "Simultaneous ac-dc Power Transmission",
Institution of Engineers (India) Journal-EL, Vol 82, June 2001, pp. 32-35.
148
[15] Gholipour E.,and Saadate S., "ImDroving Transient Stability of Pov/er System
Using UPF' IEEE Transaction on Power Delivery, Vol.30 No.2, April2005,
pp. 1677-1681.
[16] Hammad A.E., "Stability ajid Control of HVDC and AC Transmission in Paralkr'
IEEE Transaction on Power Delivery, Vol.l4,No.4, October 1999. pp. 1545-1554.
[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.
[18] Reeve J,and Uzunovic E., "'Study of Power Transfer Capability of DC System
Incorporating AC Loads and a Parallel AC Line'' IEEE Transaction on Power
Delivery, Vol. 12 No.l, Januaiy 1997, pp.426-434.
[19] Karlecik-Mair, " A New Ciosed Loop Control Method for HVDC Transmission",
IEEE Transaction on Power Delivery, Vol. 11, No.4, October 1996, pp. 1955-1960.
[20] Mihalic R et al, "ImproA'ement of Transient Stability Using Unified Power
Controller" IEEE Transaction on Power Delivery, Vol. 11, No.l, January 1996,
pp.485-492.
[21] Gole A.M. et al, "Firing Angle Modulation for Eliminating Transformer DC
Current in Coupled AC-DC Sj'stem" ", IEEE Transaction on Power Delivery, Vol.
10, No.4, October 1995, pp.2040-2047.
[22] L.K. Gyugyi et al., "The unified power flow controller; A new approach to power
transmission control", IEEE Transaction on Power Delivery, vol. 10, No. 2, April
1995, pp. 1085-1097.
[23] Reeve J, and Sultan M, "Robust Adoptive Control of HVDC System" IEEE
Transaction on Power Delivery, Vol. 9, No.3, October 1994, pp.1487-1492.
[24] To K W V et al, "A Robust Coordinated Control Sheme for HVD Transmission
with Parallel AC Systems" IEEE Transaction on Power Delivery. Vol. 9, No.3, July
1994,pp.l710-1716.
[25] Gyugyi L., "Dynamic Compensation of AC Transmission Lines by Solid State
Synchronous Voltage Source". IEEE Transaction on Power Delivery, vol. 9, No. 2,
April 1994, pp. 904-9911.
149
[26] Morin G et al., "Modeling of Hydro-Quebec- New England HVDC System and
Digital Control with EMTP", IEEE Transaction on Power Delivery, vol. 8, No. 2,
Aprill993, pp. 559-565.
[27] Smed T and Anderson G., "Utilising HVDC to Damp Power Oscillations", IEEE
Transaction on Power Deliver}', vol. 8, No. 2, April 1993, pp. 620-626
[28] Hanmiad A.E. and Long W F, "Performance and Economic Comparisons
Between Point to Point HVDC and Hybrid Back to Back HVDC/AC Transmission""
IEEE Transaction on Power Delivery, Vol. 5,No.2, April 1990, pp.1137-1144.
[29] Larsen E W et al, "Parallel AC/DC Tnansmission Lines Steady-State Induction
Issues " IEEE Transaction on Power Delivery, Vol. 4, No.l, January 1989, pp.667-
673.
[30] Lee R.L. et al, "Enhancement of AC/DC System Performance by Modulation of
Multi-terminal DC System in Southwest U.S." IEEE Transactior, on Power Delivery,
Vol. 3, No.l, January 1988, pp.307-316.
[31 ] Fujita H. et al, "Control and Analysis of a Unified Power Flow Controller." IEEE
Transaction on Power Electrofiics, Vol. 14, No.6, November 1999, pp.1021-1027.
[32] Routray A et al, "A Fuzzy Self-Tuning PI Controller for HVDC Links" IEEE
Transaction on Power Electronics, Vol. 11, No.6, September 1996, pp.669-679.
[33] Hsu Y Y and Li Wang, "Damping of Parallel AC-DC Power System Using PID
Power System Stabilizers and Rectifier Current Regulator" IEEE Transaction on
Energy Conversion, Vol. 3, No.3, September 1988, pp.540-546
[34] Lee Y-C et al, "Damping of Power System Oscillations with output Feedback
and Strip Eigenvalue Assignment" IEEE Transaction on Power System, Vol. 10.
No.3, August 1995, pp.1620-1626.
[35] Klein M et al, "Fundamental Study of Inter-Area Oscillations in Power System"
IEEE Transaction on Power System, Vol. 6. No.3, August 1991, pp.914-921.
[36] Arabi et al, "Small Signal Stability Program Analysis of SVC and HVDC in AC
Power System" IEEE Transaction on Power System, Vol. 6. No.3, August 1991,
pp.1147-1153.
150
[37] Edris A.A, "Enhancement of First Swing Stability Using a High-Speed Phase
Shifter" IEEE Transaction on Power System, Vol.6 No.3, August 1991, pp. 1113-
1118.
[38] IEEE Repon, "HVDC Control for System Dynamic Performance" IEEE
Transaction on Power System, Vol. 6. No.2, May 1991, pp.743-7.')2.
[39] Ni Y X and Foud A A, "Simplified Two Terminal HVDC Model and its Use in
Direct Transient Stability Assesment" IEEE Transaction on Power System, Vol. 2.
No.4, November 1986, pp. 1006-1012.
[40] Rahim A H M A and Amin M E., "Stabilization of High Voltage ACDC Power
System I. Evaluation of Control Strategies" IEEE Transaction on Power Apparatus
and System, vol. PAS-104, No. 11, November 1985. pp. 3084-3091.
[41] Vovos N A and Galnos, 'Enhancement of the Transient Stability of Integrated
AC/DC Systems Using Active and Reactive Power Modulations" IEEE Transaction
on Power Apparatus and System, vol. PAS-104,, No. 7, Julyl985, pp. 1696-1702.
[42] M.A. Chaudhry and D.P. Caroll, "Coordinated active £ind reactive power
modulation of multiterminal HVDC system"', IEEE Trans. Power App. System, Vol.
PAS-103,1989, pp. 1480-1485.
[43] Grund C E et al, "Control Design of an Active and Reactive Pov/er HVDC
Modulation System with Kalman Filtering" IEEE Transaction on Power Apparatus
and System, vol. PAS-101, No. 10, October 1982, pp. 4100-4108.
[44] Pai M A et al, "Transient Stability Analysis of Multi- Machine AC/DC Power
Systems Via Energy Function Method" IEEE Transaction on Power Apparatus and
System, vol. PAS-100, No. 12, December 1981, pp. 5027-5035.
[45] Foud A A and Stanton S E, "Transient Stability Analysis of Multi- Machine
Power Systems Part I: Investigation of System Trajectory" IEEE Transaction on
Power Apparatus and System, vol. PAS-100, No. 7, July 1981, pp. 3408-3416.
[46] Foud A A and Stanton S E, "Transient Stabilit>' Analysis of Multi- Machine
Power Systems Part II: Critical Transient Energy" IEEE Transaction on Power
Apparatus and System, vol. PAS-100, No. 7, July 1981, pp. 3417-3423.
151
[47] IEEE Committee Report, "Dynamic Performance Characteristics of North
American HVDC System For Transient and Dynamic Stability Evaluations" IEEE
Transaction on Power Apparatus and System, vol. PAS-100, No. 7, Jul> 1981, pp.
3356-3364.
[48] Dunlop R D, Gutman R. and Marchenco R.P. "Analytical Development of
Loadability Characteristics for EHV and UHV Transmission Lines"', IEEE
Transaction on PAS,vol.PAS-98.N0.2 ,1979, pp:606-617.
[49] Vovos N.A and Galanos G.D., "Transient stability of ac-dc system", IEEE
Transaction on Power Apparatus and System, vol. PAS-98, Vol.4 , July/Augustl979,
pp.1375-1383.
[50] Vovos N.A and Galanos G.D., " Damping of power swings in AC Ties Using a
parallel DC Link Operating at Constant Reactive PoM'er Controf, IEEE Transaction
on Power Apparatus and System, vol. PAS-98, Vol. 2 , March /April 1979, pp.416-
425.
[51] Klein J.W., Kranse P.C. ;ind Femandes L.A., '"Dynamic Stability Assessment
Model of a parallel AC-DC pC'Wer system" , IEEE Transaction on Power Apparatus
and System, vol. PAS-96, Vol.4 , July/August 1977, pp.1296-1304.
[52] Peterson H.A. and Krause P.C, "A direct and quadrature axis representation of a
parallel AC and DC power system", IEEE Transaction on Power Apparatus and
System, vol. PAS-85,, No. 3, March 1966, pp. 210-225.
[53] Machida T., "Improving Transient Stability of AC System by Joint Usage of DC
System" IEEE Transaction on Power Apparatus and System, vol. PAS-85,, No. 3,
March 1966, pp. 226-231.
[54] Peterson H.A., Krause P.C, and Thomas C H "Analog Computer Study of a
parallel AC and DC power system", IEEE Transaction on Power Apparatus and
System, vol. PAS-85, March 1966, pp. 191-209.
[55] Peterson H.A. and Krause P.C, "Damping of power swings a parallel AC and DC
power system", IEEE Transaction on Power Apparatus and System, vol. PAS-85,
December 1966, pp. 1231-1239.
152
[56] Cai H et al, "Determination of the power transfer capacity of UPFC with
consideration of the system and equipment constraints and of installation locations",
lEEProc-GTD, Vol. 149, No.L , January2002 , pp. 114-664120.
[57] Klein J.W., Kranse P.C. and Femandes L.A., "Investigation of dc Modulation in a
parallel AC/DC power system" , IEEE Winter Power Meeting, paper A78235-4,
1978.
[58] To K W V and David A K., '-Multivariable adoptive Control of AC-DC System",
lEEProc-GTD, Vol. 14J, No. 6. , November 1994, pp. 658-664.
[59] Dash P K, Liew A C and Routray A, "High performance Controllers for HVDC
Transmission Links", lEE Proc-GTD, Vol. 141, No.5, September 1994, pp. 422-428.
[60] Gyugyi L.K., "Unified power flow concept for flexible A.C. transmission
systems", lEEproceedings-C, Vol.139, No,4, July 1992, pp. 323-331.
[61] Goudie D B , "Steady state stability of parallel HV AC -DC Power Transmission
System", lEEproceedings-C, Vol. 119, No. 2, February 1972, pp. 216 224.
[62] Stella M., Dash P.K. and Basu K.P., "A Neuro-Sliding Mode Controller for
STATCOM", Electric Power Components and Systems, Vol. 32, pp.131-147,
Feb.2004.
[63] Bruer GD, et al., " Studies of Large AC/DC system on Digital Computer" , IEEE
Transaction on Power Apparatus and System, . PAS-85, Vol.11 , November 1966,
pp.1107-1114.
[64] Dougherty J J and Caleca V, " The EEI AC/DC Transmission Model", IEEE
Transaction on Power Apparatus and System,. PAS-87, February 1968, pp.504-512.
[65] Gless G E, " Direct Method of Liapunov Applied to Transient Power System
Stability", IEEE Transaction on Power Apparatus and System, . PAS-85, Vol.2,
February 1966, pp.l59>168.
[66] Brierley R H et al, "Compact Right of Ways with multi-Voltagr Towers" IEEE
Transaction on Power Delivery, Vol.6 No.4, October 1991, pp. 1682-1689.
[67] Ekstrom A, and Lamell P, "HVDC Tapping Station: Power Tapping from a DC
Transmission Line to a Local AC Network," AC-DC Conference, London, 1991,
PP.126-13L
153
[68] Task force on Small HVDC Taps, working Group, "Integration of Small Taps into
(Existing) HVDC Links" IEEE Transaction on Power Delivery, Vol. 10 No.3, July
1995,pp.l699-1706.
154
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