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Bulk power transmission by superconducting DC cable
Electric Power Systems Research, 1 ( 1 9 7 7 / 7 8 ) 41 - 4 9
41
© Elsevier Sequoia S.A., L a u s a n n e - - P r i n t e d in t h e N e t h e r l a n d s
Bulk Power Transmission b y Superconducting DC Cable
P. C H O W D H U R I a n d F. J. E D E S K U T Y
Los Alamos Scientific Laboratory of the University of California, Los Alamos, New Mexico 87545 (U.S.A.) (Received M a r c h 24, 1 9 7 7 )
SUMMARY
It is estimated that the total electric system load in the U.S.A. will increase five- to sevenfold b y the year 2000. It is reasonable to assume that the transmission line capacities will also increase at least b y the same amount. For energy conservation, future transmission lines must be more efficient than those of today. There will also be mounting public pressure to transmit electric power underground because o f increasing awareness of environmental quality. There are considerable research activities to develop new technologies to transmit large blocks of power more efficiently b y underground cables. The goal of the DC superconducting power transmission line (SPTL) program at the Los Alamos Scientific Laboratory is to develop such a cable. The DC SPTL has the advantage that it has very high p o w e r capability (10 GW and more) and no c o n d u c t o r or dielectric losses. The only "loss" of a DC SPTL is in the refrigeration requirement. It has no systems constraints, such as the stability or reactive compensation requirements of an AC system. The paper discusses the conceptual Los Alamos design, some o f the technical problems associated with the development and the approaches to their solution.
INTRODUCTION
Increasing demand for electric power, increased cost and restrictions on rights-ofway of transmission lines, lower i2r-loss requirements and electrical-system stability criteria are some of the major reasons for building extra-high-voltage (EHV) power transmission lines. Table 1 shows clearly the increased power density over the right-of-way
for higher voltage power transmission lines [1]. EHV overhead power lines may be objectionable for several reasons [ 2 - 8 ] : (1) lack of reliability because of fiashover of polluted insulators, breakage of line or towers during storms and hurricanes; (2) radio-frequency interference; (3) possible effects of electric fields on man, flora and fauna; (4) visual pollution. Moreover, it may be impossible to build overhead lines in an urban area because of unavailability of rights-of-way and prohibition b y statutory regulations. Pressing needs appear to exist for underground transmission [2 - 6]. There are three major types o f application of underground cables: (1) transmission of power between generating station (power park) and load center; (2) transmission link within a power pool; (3) transmission link between two power pools. It is generally good design practice to connect one cable to one generator. Therefore, for the first type of application, the required cable rating will increase from 1.2 GW in 1970 to 1.5 GW in 1980 to 2 GW in 1990 [3]. For power-pool connections, it is expected that the desirable cable ratings will be 2 GW in 1980, 4 GW in 1990 and 8 GW in 2000 [3]. The major problems of a high-power highvoltage underground cable are thermal dissipation of i2r losses and thermal aging of the dielectric material. Although a forced-cooled SF6 insulated cable has much promise [ 3 ] , a superconducting p o w e r transmission line (SPTL) seems to be more suitable for transmission of higher power electrical energy. Its c o n d u c t o r loss is negligible and its dielectric aging should be minimal because of the low operating temperature.
42 TABLE 1 Characteristics of UHV and EHV overhead lines Line voltage (kV)
Surge impedance loading
Right-of-way (m (ft))
Power/width (MW/m (MW/ft))
91.4 83.8 61.0 54.9 48.8 38.1 30.5
80.1 (24.4) 44.3 (13.5) 31.2(9.5) 16.4 (5.0) 8.9 (2.7) 3.6 (1.1) 2.0 (0.6)
(GW) 1400 1100 765 500 345 230 138
7.3 3.7 1.9 0.9 0.4 0.14 0.06
(300) (275) (200) (180) (150) (125) (100)
A DC SPTL is ideal for transmitting large blocks of electric power, because (1) it has no system constraints, such as stability, load flow, short-circuit current level and reactive compensation; (2) it has negligible c o n d u c t o r losses and no dielectric losses -- therefore, it is very efficient; and (3) its current density capability is very high and one coaxial DC SPTL can be equivalent to the three cables for the corresponding three-phase AC system -therefore, it is less expensive.
LOS ALAMOS DC SPTL DESIGN
The DC SPTL development at the Los Alamos Scientific Laboratory (LASL) has been concerned primarily with the design of a flexible, compact, 100 kV, 5 GW cable. The voltage and p o w e r levels were arbitrarily selected to show the feasibility of a low-voltage high-power DC SPTL. Figure 1 shows the conceptual design of the LASL DC SPTL [9, 1 0 ] . The innermost
VACUUMJ A C K E T ~ - ~
SPIRAL T
A
P
E
~
//_!_ RAPPO,A
SUPERCONDUCTOR J DIELECTRIC (MULTIFILAMENTARY Nb5 Sn IN Cu)
INSULATION -ARMOR
Fig. 1. Conceptual design of the LASL DC SPTL.
member of the cable is a spiral support. Upon this support is w o u n d the inner conductor which consists of subcables made up of wires of multifilamentary NbaSn superconductor e m b e d d e d in copper matrix. Each subcable is made up of several wires. The dielectric layer is composed of wrapped paper or plastic tape. The outer, return-current c o n d u c t o r is similar to the inner conductor, except in the number of subcables and wires. An outer flexible armor completes the cable. The cryogenic enclosure is rigid and consists of a helium container and a vacuum jacket with multilayer thermal insulation within the evacuated annulus. The enclosure will be built in sections, roughly 30 m long, to be joined and laid in trenches. Lengths of flexible cable up to 900 m long will then be drawn into the enclosure. The cable plus enclosure thus provide both complete electrical and refrigeration circuits with the " g o " and "return" helium coolant flow paths separated b y the cable armor. The refrigeration system will utilize line refrigerators with the final expansion engine at the far end of each line refrigeration section. Thus the cable itself becomes the final heat exchanger of the refrigerator [11]. The operating temperature range of the cable will be 10 - 12 K, which is significantly higher than that for most other applications of superconductivity. This factor, in addition to the small enclosure diameter needed for the DC SPTL, will significantly reduce the refrigeration loads. Spacing between the refrigerators along the line could vary from 5 to 30 km, depending u p o n the actual installation situation. Figure 2 shows a schematic representation o f the DC SPTL system.
43 •
2700 m
R
CONTRACTION
1~ POTHEAD POTHEAD REFRIGERATOR
,oo:
MANHOLE
mj~JOINT
, ,vl
I I ENCLOSURE JUNCTION BOX JOINT
JOCTiONO,
'"<----'.U<,HEAOj
i ~ 1 l, JUNCTION BOX
LI
_/
t.
I I JUNCTION BOX
JUNCTION BOX
. . . .
~ E R A T O R
~EXPANDER
i
[
)
_f=--=L
I I JUNCTION BOX
?,ANOE: JUNCTION BOX
i a--~
BULKHEAD
JUNCTION BOX
POTHEAD REFRIGERATOR
POTHEAD
Fig. 2. Schematic representation of the DC SPTL system.
D C SPTL D E V E L O P M E N T
Development of a DC SPTL is a multidisciplinary program. It requires the simultaneous development of the cable, the cryogenic envelope, and the refrigeration system. In addition, it is necessary to study the interaction of the DC SPTL with external electrical systems, both to determine the steadystate behavior and the response to electrical transients and faults as well as to estimate its economic competitiveness vis-d-vis alternative technologies.
If this heat is n o t removed, the critical temperature limit for the operating condition may be exceeded, driving the superconductor "normal". This would then be followed by destructive localized power dissipation because the normal-state resistivity o f most superconducting alloys is around 4 × 10 -7 ~2 m (20 times the resistivity of copper at room temperature). A superconductor is stabilized b y controlling all temperature excursions and b y adding closely coupled parallel normal c o n d u c t o r (cryoconductor or cryostabilizer), such as copper with a lowtemperature resistivity of 10 - l ° ~2 m (200
Superconductor development
A superconductor experiences no i2r losses under DC conditions. However, the current density in the superconductor is limited by its temperature and magnetic flux density (Fig. 3). These three parameters define a critical surface. Superconductivity can only exist inside the volume defined b y that surface, which is different for each superconducting alloy or c o m p o u n d . The critical temperature o f Nb3Sn , the superconductor selected in our work, is 17 - 18 K. If properly prepared and arranged, Nb3Sn can support a critical direct current density of 15 k A / m m 2 at 14 K at a self-field of 0.5 T (5 kG). Although a superconductor offers zero resistance to DC, it experiences power dissipation with AC and electrical and mechanical transients because of magnetic flux movements inside the superconductor. These losses c a u s e localized heating in the superconductor.
Jl A
,I \
I
I
t, Fig. 3. Phase diagram for the s u p e r c o n d u c t i n g state. J = current density; H = magnetic field; T = temperature.
44 times lower than copper at room temperature), so t h a t the composite can continue to carry the required currents, at least temporarily [12]. Figure 4 shows the cross-section of a stabilized composite superconducting wire to be used in the LASL DC SPTL [13]. Under transient conditions, the current can easily transfer from the superconductor to the copper matrix.
( \
I
Fig. 4. Cro~-seetion of the Nb3Sn multifilamentary superconducting wire cryogenically stabilized by copper matrix. The Nb3Sn is formed by the reaction of the Nb tube with the Sn of the bronze rod, leaving a thin layer of Nb as diffusion barrier between Nb3Sn and the Cu matrix. The use of multifilamentary superconducting wires instead of superconducting tapes has several advantages: (1) it conforms with the accepted cable manufacturing techniques; (2) it avoids the handling of the highly anisotropic and somewhat fragile Nb3Sn tapes; (3) the necessary superconductor stability criteria can be more easily met.
Dielectric development Apart from being adequate under electrical stress, the dielectric should also be mechanically and thermally suitable for handling at room temperature and operation at a cryogenic temperature. It should be borne in mind that the electrical requirements of a dielectric under DC stress are different from those under AC stress. For AC operation, the corona inception voltage and the dielectric losses will be the limiting constraints. For DC applications, the polarity-reversal stress will be the limiting factor. For long-term stress applications, dielectric behavior depends upon two entirely independent phenomena: the statistical nature of breakdown and aging. It is expected
that the aging mechanism of a DC SPTL will be slower than that of an AC SPTL because of the absence of dielectric losses under DC stresses. The breakdown process may also be different under flowing and stationary cryogen [14]. Mechanically, the cable should be able to withstand bending and unbending at room temperature w i t h o u t the dielectric layers buckling. During cool~lown, the dielectric should be able to withstand the mechanical stress created by the contraction. It will also have to continuously withstand the cryogenic operating temperature w i t h o u t cracking. Many data are available on cryogenic dielectrics under AC stress, but very few under DC stress or under polarity reversal. Oilimpregnated cellulose paper will be ideal from the manufacturing standpoint because it is the standard dielectric for all high-voltage cables. It has also been found that the strengths of the oil-impregnated paper at liquid nitrogen temperature with DC, impulse and polarity reversal are significantly higher than those at room temperature [15 - 16]. However, no data are available at supercritical helium temperatures, i.e. greater than 5.2 K. There is also the likelihood that the oil vapor might contaminate the helium during the cooldown process. It should be evident that the selection of a suitable dielectric material for the DC SPTL is a painstaking process which requires careful long-range planning. The LASL has embarked upon such a program. The dielectric development program has been divided into two parts: initial screening tests and cable sample tests. Nine dielectric materials have been selected for the initial screening tests (Table 2). DC breakdown tests will be conducted on 76 pm (3 m i l ) t h i c k sheets of these dielectric materials at 12 K and 1 MPa (10 atm) of helium pressure to determine the effects of polarity and rate of voltage application. A series of tests will also be performed on cellulose paper impregnated with a specially selected mineral oil. Three dielectric materials will be selected for cable sample tests after analysis of the statistically designed screening tests. During the cable sample tests, the three selected dielectric materials will be tested with DC, impulse, switching surge and polarity reversal in addition to step-stress tests and corona tests under
45 TABLE 2 Dielectric materials selected for screening tests Trade name Cellulose paper Cryovac Kapton Makrofol Mylar Nomex Teflon Udel Valeron
Type of material Cross-linked medium-density polyethylene Polyimide Polycarbonate Polyester Fibrous-spunbonded polyamide Polytetrafluoroethylene Polysulfone Polyethylene
long-term DC stress. A p r o t o t y p e cable will be manufactured with the finally selected dielectric to determine whether all specifications can be met in this realistic configuration. All nine dielectric materials (Table 2) were wound on a 44 mm (laA in.) o.d. aluminum mandrel and cooled down to 4 K in order to determine whether they fracture because of mechanical stress caused b y thermal contraction. The dielectric tapes were 19 mm (3A in.) wide and 76 ~m (3 mil) thick. Two layers of tapes were w o u n d with a 33/67 registration and 3.2 mm (l/s in.) b u t t gap on the aluminum mandrel with a hand tension o f a b o u t 2.2 kg. An additional cool-down test was performed on the cellulose paper impregnated with a special mineral oil. None of the dielectric tapes showed any signs of damage after the cool-down tests. No residue was observed on the oil-impregnated cellulose paper. Electrical characteristics
A DC SPTL has to be designed so that it will operate satisfactorily under normal and abnormal system conditions. The areas of concern are: ripple current produced by the converters, overcurrents produced b y faults, and overvoltages produced by lightning and switching. The ripple current will cause steady-state power losses in the conductor, requiring the refrigeration system to supply cooling power to compensate for these losses. A transient overcurrent will cause transient power losses in the conductor. This may drive the superconductor "normal", requiring, under the worst condition, power interruption to the load. As shown in Fig. 1, the LASL design of the DC SPTL contains four electrically conducting cylinders: the inner
and outer current-carrying conductors, and the inner and outer cylinders of the cryogenic enclosure. Therefore, the usual analysis of surge propagation performed for a twoconductor coaxial cable system will not apply. For the LASL DC SPTL, multivelocity surges will appear on all four conductors and will need special analysis. Studies on the effects of ripple current and fault currents on the DC SPTL have been completed; the overvoltage study is in progress. Harmonic currents on the DC SPTL produced both b y the normal and abnormal harmonic voltage sources have been computed for a 100 km long DC SPTL under various operating conditions. The computations of the normal harmonics are based on the standing-wave theory of long transmission lines. The computations of the abnormal current harmonic were performed by representing the DC SPTL as an equivalent n-network because the line length at these frequencies is less than the quarter-wavelength. The DC SPTL was assumed to be terminated at either end b y a smoothing reactor. No other filter network is connected to the DC SPTL. The results of the analysis are shown in Tables 3 and 4 for the normal and the abnormal harmonics, respectively. The most severe fault current through the DC SPTL will occur if the line-end bushing of the inverter-side smoothing reactor flashes over. The fault current will consist of two components. The first c o m p o n e n t will be a traveling wave caused by the discharge of the cable. The second c o m p o n e n t will be driven by the voltage source on the AC side of the rectifier. The magnitude and duration of the cable discharge current cannot be controlled b y any external means, such as valve control of the converter or a DC circuit breaker. Its magnitude is given b y il = V / Z c
(1)
where V is the system voltage and Z c the surge impedance of the cable. Its duration r at the point of flashover is given by r = 2l/v
(2)
where l is the length of the cable and v the velocity of propagation in the cable. The duration diminishes monotonically along the DC SPTL and is zero at the rectifier end. For
46 LT
TABLE 3 Maximum normal harmonic currents along DC SPTL* Frequency (Hz)
Max. harmonic current (Arms)
720 1440
262.4 34.0
'I
*Line length = 100 km; 12-pulse operation; firing angle delay = 15 ° ; overlap angle = 22.6 °. TABLE 4 Abnormal harmonic currents along DC SPTL* Frequency (Hz)
Max. harmonic current (A rms)
60 120 180 240 300
70.3 35.1 23.4 15.3 8.6
(a)
I
4Lc
LT
L~.
(b)
Fig. 5. Representation of a 12-pulse rectifier bridge with a fault at the inverter end. (a) Schematic diagram; (b) equivalent circuit.
*Line length = 100 km; 12-pulse operation; firing angle delay = 0 - 90°; firing angle tolerance = ±0.25 °.
v~
v b v bv
a 100 k V , 5 GW, 1 0 0 k m l o n g DC S P T L , il = 1 3 . 1 0 k A a n d T = 1 . 1 2 ms. The second c o m p o n e n t of the fault current will b e m a i n t a i n e d b y t h e v o l t a g e sources on t h e AC side o f t h e rectifier. T h e r e f o r e , t h e AC-side r e a c t a n c e s , t h e s m o o t h i n g r e a c t o r a n d t h e DC S P T L r e a c t a n c e will limit this c o m p o nent of the fault current. The profile of the s e c o n d c o m p o n e n t o f t h e f a u l t c u r r e n t will b e dependent upon the means of interruption, such as valve c o n t r o l o f t h e c o n v e r t e r or a DC circuit b r e a k e r . C o m p u t a t i o n s o f t h e s e c o n d c o m p o n e n t of the fault current were made with the following assumptions: (1) t h e f a u l t is i n i t i a t e d at t h e b e g i n n i n g o f commutation; (2) t h e r e is n o c o m m u t a t i o n o v e r l a p ; (3) t h e f a u l t c u r r e n t is i n t e r r u p t e d b y b l o c k i n g t h e firing o f t h e s u b s e q u e n t valves; (4) firing angle d e l a y is z e r o ; (5) c o n v e r t e r s o p e r a t e in t h e 12-pulse mode. T h e 12-pulse s y s t e m w i t h its e q u i v a l e n t circuit f o r t h e c o m p u t a t i o n o f t h e f a u l t c u r r e n t is s h o w n in Fig. 5. In this s y s t e m , t h e f a u l t c u r r e n t has a c h a n c e o f i n t e r r u p t i o n e v e r y ~ / 6 radians o f t h e v o l t a g e wave. This will p r o d u c e a half-sinusoidal s e c o n d c o m p o n e n t o f t h e f a u l t c u r r e n t , as s h o w n in Fig. 6.
6,
A I I @ ~ ~
(a)
00002 ~',: e,
,, -
_.
e
cb V' ObV'aCV'bc
V'
A
2
1
~
A22
(b)
I I I I I
(c) 0
tm ---~ t
Fig. 6. Voltage across 12-pulse rectifier bridge and current profile under fault conditions. (a) Voltage across bridge 1; (b) voltage across bridge 2; (c) fault current. T h e p e a k m a g n i t u d e a n d d u r a t i o n o f t h e fault c u r r e n t are given b y (n + 1 ) ( A l l + A21) + AI2 + A22 imn =
4Lc + LT +/-q
(3)
47
and nn tmn
--
6c~
7n (4)
+ ~
12~
where n is the number of subsequent valves which fail to block; A1z, A12, A2~, A22 are the areas under the voltage wave, as shown in Fig. 6. Lc is the leakage inductance of the converter transformer; Lr is the inductance of the smoothing reactor; and/~ is the inductance of the DC SPTL. The fault currents (second component) of a 100 kV, 5 GW, 100 km long DC SPTL are shown in Fig. 7.
current (imt), including the prefault load current. The current starts to decay once it is transferred to the non-linear resistor of the DC circuit breaker. The characteristics of the non-linear resistor were assumed to be (5)
Va = K i b a
where K and b u , : ~ ; ~ a n t s , va is the voltage across the non-linear resistor, and ia is the current through the non-linear resistor. The decaying total current is given by i(1--b) = i(lmt b) .__ ( K / 4 L c
2.0
[
i
r
f
[
r
I
i
+ L T +/d)(l
--b)t
(6)
I
The constant K was derived on the assumption t h a t Va = 1.7 VDC at imt, and b was assumed to be 0.05. Figure 8 shows the fault current profiles for t d = 5 and 10 ms.
I-Z LU 1.5 (.3
._1 0 F-
hO OT
0
n=O
I
I
I
I
[
IO
n=l
I
, , , i , , , , i
n=2
J
I
I
2 oL ' 20
TIM E (ms)
Fig. 7. Fault currents o f a 100 kV, 5 GW, 100 k m long DC SPTL, cleared by valve control (12-pulse operation). The c o n v e r t e r transformer leakage reactance is 0.18 p.u. and t h a t o f the s m o o t h i n g reactor is 2 p.u.
DC power circuit breakers are n o t commercially available, b u t three types are being investigated [ 17 - 20]. Although the modes of operation of these three types of circuit breakers are significantly different, t h e y are similar in two basic principles: (1) the fault current is n o t limited during the initial period t d of sensing and contact opening; (2) the total current is diverted to a nonlinear resistor during the final period of interruption. Computations o f the fault current were made for two values of td, 5 and 10 ms. During this period, the fault current analysis is the same as t h a t with the valve control except that no valve is blocking. The current at the end of td is the peak of the total
/
/
Q.
>o v
/
/
1,5
\
1.0
\ 0.5 ._J
\ 0
0
I0 TIME (ms)
20
Fig. 8. Fault current i and circuit breaker voltage v o f a 100 kV, 5 GW, 100 k m long DC SPTL, cleared by a DC circuit breaker with delay times t d = 5 and 10 ms. The c o n v e r t e r transformer leakage reactance is 0.18 p.u. and that o f t h e s m o o t h i n g reactor is 2 p.u.
48 APPLICATIONS OF DC SPTL A DC SPTL is particularly attractive for transmitting large blocks of power (2.5 GW and more) from a pow e r park to the urban load center. A DC SPTL may also be attractive to bring power to a large m et r opol it an area from outlying AC transmission networks. Both the economic and technical considerations have to be taken into a c c o u n t in selecting a particular transmission scheme. For the DC SPTL, the cost of t he terminal converter stations has t o be considered in addition to the cost of the DC SPTL and its installation. Comparing it with the AC alternatives, the DC SPTL o f LASL design has only one cable compared with three cables for AC systems. However, the DC terminal equipmen t will be considerably m or e expensive than the AC terminal equipment. Therefore, a cross-over distance exists for each application b e y o n d which t he DC SPTL will be more economical. If, for technical reasons, a DC transmission is chosen, th en the DC SPTL will have to c o m p e t e with the ot her DC alternatives. In this case, one cable o f the LASL DC SPTL will c o m p e t e favorably with t w o cables for the alternative DC transmission schemes. Moreover, in m a ny cases, several pairs of conventional DC cables will be required to transmit the same pow e r that can be transmitted by one DC SPTL of the LASL design. In addition, it is e x p e c t e d t ha t the DC SPTL will offer further savings in converters. As the DC SPTL is inherently a high-current system, the rated voltage of the DC SPTL will be substantially lower than t ha t of the alternative DC transmission schemes for the same transmitted power. It has been r e por t ed that the converter cost increases by a bout 4% for every 100 kV increase in the rated voltage [ 2 1 ] . As converter cost is a significant part of the total transmission cost, this will result in substantial cost savings for the DC SPTL system. As the land area for a lower voltage converter station is smaller than that for a higher voltage station [ 2 2 ] , additional savings are e x p e c t e d to result from converter stations dedicated to th e DC SPTL. Recently, a study was com pl e t e d to estimate the cost of transmitting 10 GW p o wer f r o m a power park to a city center over a 100 km distance. The first contingency
rating of the cable system was also 10 GW, while the second contingency rating was 7.5 GW for 4 hours. Analysis showed that the voltage rating of 300 kV will be the most economical for this application (Fig. 9). The steady-state losses caused by ripple current were f o u n d to be negligible, 2.12% being the total estimated losses of the system, including those o f the converter stations. Fault current calculations showed that the thyristor converters will be able to block in ample time before the specified limit of 2 K t e m p e r a t u r e rise is reached. The relative costs of the various c o m p o n e n t s of the DC SPTL system are shown in Table 5. These costs do n o t include the land cost for the refrigeration system and t hat for converters. I
l
l
75
l
[
l
l
l
GW DC SPTL
"~ i o o o "--
A
8
500 I
J
I
0 LINE
1
l 500
VOLTAGE
I
I
l
I 1000
( k V )
Fig. 9. Cable and enclosure cost of a 7.5 GW DC SPTL as a function of system voltage. Curve A represents today's technology, curve B the expected technology in about 1985, and curve C the expected technology in about 1990. TABLE 5 Relative costs of components of a 7.5 GW DC SPTL system Item
Relative cost (%)
Cable and enclosure Refrigeration system Other Installation Converter Capitalized losses Demand cost for losses Total
11.32 5.54 4.74 10.15 43.65 16.55 8.05 100
49
CONCLUSIONS The DC S P T L is efficient a n d cost effective f o r t r a n s m i t t i n g large blocks o f p o w e r . T h e e c o n o m i c l o w e r limits o f p o w e r a n d transmission distance m u s t be d e t e r m i n e d f o r each a p p l i c a t i o n . The electrical s y s t e m requirem e n t s , such as ripple losses and o v e r c u r r e n t , can be m e t b y a well~lesigned DC SPTL.
ACKNOWLEDGEMENT The w o r k r e p o r t e d in this p a p e r was s u p p o r t e d b y t h e U n i t e d States E n e r g y Research and D e v e l o p m e n t A d m i n i s t r a t i o n .
REFERENCES 1 Federal Council on Science and Technology Energy R & D Goals Study: Report of Technical Group on Electrical Transmission and Systems, 14 July 1972. 2 A . F . Corry, Proc. 1972 Applied Superconductivity Conf., IEEE Publ. No. 72-CHO 682-5TABSC, 1972, p. 160. 3 J. Nicol, Proc. 1972 Applied Superconductivity Conf., IEEE Publ. No. 72-CHO 682-5-TABSC, 1972, p. 165. 4 A. F. Corry and E. Kasum, Proc. 1972 IEEE Underground Transmission Conf., IEEE Publ. No. 72-CHO 680-0-PWR, 1972, p. 1. 5 A. S. Brookes, Proc. Amer. Power Conf., 35 (1973) 1145. 6 J. Nicol, in K. Mendelssohn (ed.), Proc. Fifth Int. Cryogenic Engineering Conf., IPC Science and Technology Press, Whitstable, England, 1974, p. 68.
7 G. C. Knickerbocker (ed.), IEEE Publ. No. 78CHO 1020-7-PWR, 1975. 8 J. E. Bridges, IEEE PES Winter Meeting, New York, 1977, Paper No. F 77 256-1. 9 J.W. Dean and H. L. Laquer, Proc. 1976 Underground Transmission and Distribution Conf., IEEE Publ. No. 76-CH 1119-7-PWR, 1976, p. 417. 10 H. L. Laquer, J. W. Dean and P. Chowdhuri, IEEE Trans., MAG-13 (1977) 182. 11 J.W. Dean and J. E. Jensen, in K. D. Timmerhaus and D. H. Weitzel (eds.), Advances in Cryogenic Engineering, Vol. 21, Plenum Press, New York and London, 1976, p. 197. 12 J. K. Hoffer, E. C. Kerr and H. L. Laquer, IEEE Trans., PAS-94 (1975) 2008. 13 H. L. Laquer, Proc. Amer. Soc. Metals 1976 Conf. on Manufacture of Superconducting Materials, Amer. Soc. Met., Metals Park, Ohio, in press. 14 L. Centurioni, G. Motinari and A. Viviani, Rev. Gen. Elec., 84 (1975) 583. 15 A. A. Hossam-Eldin and B. Salvage, Proc. Int. Conf. on High Voltage DC and/or AC Power Transmission, IEE Conf. Publ. No. 107, 1973, p. 47. 16 Z. Iwata, N. Ichiyanagi and E. Kawai, IEEE-PES Winter Meeting, New York, 1977, Paper No. F 77 186-0. 17 A. N. Greenwood and T. H. Lee, IEEE Trans., PAS-91 (1972) 1570. 18 A. N. Greenwood, P. Barkan and W. C. Kracht, IEEE Trans., PAS-91 (1972) 1575. 19 G. A. Hofmann, G. L. LaBarbera, N. E. Reed and L. A. Shillong, IEEE Trans., PAS-95 (1976) 1182. 20 A. Ekstrom, H. Haertel, H. P. Lips, W. Schultz, P. Joss, H. Holfeld and D. Kind, Conf. Int. des Grands R~seaux Electriques ~ Haute Tension, paper No. 13-06, Paris, 1976. 21 N. G. Hingorani and F. J. Ellert, Amer. Power Conf., 38 (1976) 411. 22 P. Lips, IEEE Trans., PAS-95 (1976) 894.
41
© Elsevier Sequoia S.A., L a u s a n n e - - P r i n t e d in t h e N e t h e r l a n d s
Bulk Power Transmission b y Superconducting DC Cable
P. C H O W D H U R I a n d F. J. E D E S K U T Y
Los Alamos Scientific Laboratory of the University of California, Los Alamos, New Mexico 87545 (U.S.A.) (Received M a r c h 24, 1 9 7 7 )
SUMMARY
It is estimated that the total electric system load in the U.S.A. will increase five- to sevenfold b y the year 2000. It is reasonable to assume that the transmission line capacities will also increase at least b y the same amount. For energy conservation, future transmission lines must be more efficient than those of today. There will also be mounting public pressure to transmit electric power underground because o f increasing awareness of environmental quality. There are considerable research activities to develop new technologies to transmit large blocks of power more efficiently b y underground cables. The goal of the DC superconducting power transmission line (SPTL) program at the Los Alamos Scientific Laboratory is to develop such a cable. The DC SPTL has the advantage that it has very high p o w e r capability (10 GW and more) and no c o n d u c t o r or dielectric losses. The only "loss" of a DC SPTL is in the refrigeration requirement. It has no systems constraints, such as the stability or reactive compensation requirements of an AC system. The paper discusses the conceptual Los Alamos design, some o f the technical problems associated with the development and the approaches to their solution.
INTRODUCTION
Increasing demand for electric power, increased cost and restrictions on rights-ofway of transmission lines, lower i2r-loss requirements and electrical-system stability criteria are some of the major reasons for building extra-high-voltage (EHV) power transmission lines. Table 1 shows clearly the increased power density over the right-of-way
for higher voltage power transmission lines [1]. EHV overhead power lines may be objectionable for several reasons [ 2 - 8 ] : (1) lack of reliability because of fiashover of polluted insulators, breakage of line or towers during storms and hurricanes; (2) radio-frequency interference; (3) possible effects of electric fields on man, flora and fauna; (4) visual pollution. Moreover, it may be impossible to build overhead lines in an urban area because of unavailability of rights-of-way and prohibition b y statutory regulations. Pressing needs appear to exist for underground transmission [2 - 6]. There are three major types o f application of underground cables: (1) transmission of power between generating station (power park) and load center; (2) transmission link within a power pool; (3) transmission link between two power pools. It is generally good design practice to connect one cable to one generator. Therefore, for the first type of application, the required cable rating will increase from 1.2 GW in 1970 to 1.5 GW in 1980 to 2 GW in 1990 [3]. For power-pool connections, it is expected that the desirable cable ratings will be 2 GW in 1980, 4 GW in 1990 and 8 GW in 2000 [3]. The major problems of a high-power highvoltage underground cable are thermal dissipation of i2r losses and thermal aging of the dielectric material. Although a forced-cooled SF6 insulated cable has much promise [ 3 ] , a superconducting p o w e r transmission line (SPTL) seems to be more suitable for transmission of higher power electrical energy. Its c o n d u c t o r loss is negligible and its dielectric aging should be minimal because of the low operating temperature.
42 TABLE 1 Characteristics of UHV and EHV overhead lines Line voltage (kV)
Surge impedance loading
Right-of-way (m (ft))
Power/width (MW/m (MW/ft))
91.4 83.8 61.0 54.9 48.8 38.1 30.5
80.1 (24.4) 44.3 (13.5) 31.2(9.5) 16.4 (5.0) 8.9 (2.7) 3.6 (1.1) 2.0 (0.6)
(GW) 1400 1100 765 500 345 230 138
7.3 3.7 1.9 0.9 0.4 0.14 0.06
(300) (275) (200) (180) (150) (125) (100)
A DC SPTL is ideal for transmitting large blocks of electric power, because (1) it has no system constraints, such as stability, load flow, short-circuit current level and reactive compensation; (2) it has negligible c o n d u c t o r losses and no dielectric losses -- therefore, it is very efficient; and (3) its current density capability is very high and one coaxial DC SPTL can be equivalent to the three cables for the corresponding three-phase AC system -therefore, it is less expensive.
LOS ALAMOS DC SPTL DESIGN
The DC SPTL development at the Los Alamos Scientific Laboratory (LASL) has been concerned primarily with the design of a flexible, compact, 100 kV, 5 GW cable. The voltage and p o w e r levels were arbitrarily selected to show the feasibility of a low-voltage high-power DC SPTL. Figure 1 shows the conceptual design of the LASL DC SPTL [9, 1 0 ] . The innermost
VACUUMJ A C K E T ~ - ~
SPIRAL T
A
P
E
~
//_!_ RAPPO,A
SUPERCONDUCTOR J DIELECTRIC (MULTIFILAMENTARY Nb5 Sn IN Cu)
INSULATION -ARMOR
Fig. 1. Conceptual design of the LASL DC SPTL.
member of the cable is a spiral support. Upon this support is w o u n d the inner conductor which consists of subcables made up of wires of multifilamentary NbaSn superconductor e m b e d d e d in copper matrix. Each subcable is made up of several wires. The dielectric layer is composed of wrapped paper or plastic tape. The outer, return-current c o n d u c t o r is similar to the inner conductor, except in the number of subcables and wires. An outer flexible armor completes the cable. The cryogenic enclosure is rigid and consists of a helium container and a vacuum jacket with multilayer thermal insulation within the evacuated annulus. The enclosure will be built in sections, roughly 30 m long, to be joined and laid in trenches. Lengths of flexible cable up to 900 m long will then be drawn into the enclosure. The cable plus enclosure thus provide both complete electrical and refrigeration circuits with the " g o " and "return" helium coolant flow paths separated b y the cable armor. The refrigeration system will utilize line refrigerators with the final expansion engine at the far end of each line refrigeration section. Thus the cable itself becomes the final heat exchanger of the refrigerator [11]. The operating temperature range of the cable will be 10 - 12 K, which is significantly higher than that for most other applications of superconductivity. This factor, in addition to the small enclosure diameter needed for the DC SPTL, will significantly reduce the refrigeration loads. Spacing between the refrigerators along the line could vary from 5 to 30 km, depending u p o n the actual installation situation. Figure 2 shows a schematic representation o f the DC SPTL system.
43 •
2700 m
R
CONTRACTION
1~ POTHEAD POTHEAD REFRIGERATOR
,oo:
MANHOLE
mj~JOINT
, ,vl
I I ENCLOSURE JUNCTION BOX JOINT
JOCTiONO,
'"<----'.U<,HEAOj
i ~ 1 l, JUNCTION BOX
LI
_/
t.
I I JUNCTION BOX
JUNCTION BOX
. . . .
~ E R A T O R
~EXPANDER
i
[
)
_f=--=L
I I JUNCTION BOX
?,ANOE: JUNCTION BOX
i a--~
BULKHEAD
JUNCTION BOX
POTHEAD REFRIGERATOR
POTHEAD
Fig. 2. Schematic representation of the DC SPTL system.
D C SPTL D E V E L O P M E N T
Development of a DC SPTL is a multidisciplinary program. It requires the simultaneous development of the cable, the cryogenic envelope, and the refrigeration system. In addition, it is necessary to study the interaction of the DC SPTL with external electrical systems, both to determine the steadystate behavior and the response to electrical transients and faults as well as to estimate its economic competitiveness vis-d-vis alternative technologies.
If this heat is n o t removed, the critical temperature limit for the operating condition may be exceeded, driving the superconductor "normal". This would then be followed by destructive localized power dissipation because the normal-state resistivity o f most superconducting alloys is around 4 × 10 -7 ~2 m (20 times the resistivity of copper at room temperature). A superconductor is stabilized b y controlling all temperature excursions and b y adding closely coupled parallel normal c o n d u c t o r (cryoconductor or cryostabilizer), such as copper with a lowtemperature resistivity of 10 - l ° ~2 m (200
Superconductor development
A superconductor experiences no i2r losses under DC conditions. However, the current density in the superconductor is limited by its temperature and magnetic flux density (Fig. 3). These three parameters define a critical surface. Superconductivity can only exist inside the volume defined b y that surface, which is different for each superconducting alloy or c o m p o u n d . The critical temperature o f Nb3Sn , the superconductor selected in our work, is 17 - 18 K. If properly prepared and arranged, Nb3Sn can support a critical direct current density of 15 k A / m m 2 at 14 K at a self-field of 0.5 T (5 kG). Although a superconductor offers zero resistance to DC, it experiences power dissipation with AC and electrical and mechanical transients because of magnetic flux movements inside the superconductor. These losses c a u s e localized heating in the superconductor.
Jl A
,I \
I
I
t, Fig. 3. Phase diagram for the s u p e r c o n d u c t i n g state. J = current density; H = magnetic field; T = temperature.
44 times lower than copper at room temperature), so t h a t the composite can continue to carry the required currents, at least temporarily [12]. Figure 4 shows the cross-section of a stabilized composite superconducting wire to be used in the LASL DC SPTL [13]. Under transient conditions, the current can easily transfer from the superconductor to the copper matrix.
( \
I
Fig. 4. Cro~-seetion of the Nb3Sn multifilamentary superconducting wire cryogenically stabilized by copper matrix. The Nb3Sn is formed by the reaction of the Nb tube with the Sn of the bronze rod, leaving a thin layer of Nb as diffusion barrier between Nb3Sn and the Cu matrix. The use of multifilamentary superconducting wires instead of superconducting tapes has several advantages: (1) it conforms with the accepted cable manufacturing techniques; (2) it avoids the handling of the highly anisotropic and somewhat fragile Nb3Sn tapes; (3) the necessary superconductor stability criteria can be more easily met.
Dielectric development Apart from being adequate under electrical stress, the dielectric should also be mechanically and thermally suitable for handling at room temperature and operation at a cryogenic temperature. It should be borne in mind that the electrical requirements of a dielectric under DC stress are different from those under AC stress. For AC operation, the corona inception voltage and the dielectric losses will be the limiting constraints. For DC applications, the polarity-reversal stress will be the limiting factor. For long-term stress applications, dielectric behavior depends upon two entirely independent phenomena: the statistical nature of breakdown and aging. It is expected
that the aging mechanism of a DC SPTL will be slower than that of an AC SPTL because of the absence of dielectric losses under DC stresses. The breakdown process may also be different under flowing and stationary cryogen [14]. Mechanically, the cable should be able to withstand bending and unbending at room temperature w i t h o u t the dielectric layers buckling. During cool~lown, the dielectric should be able to withstand the mechanical stress created by the contraction. It will also have to continuously withstand the cryogenic operating temperature w i t h o u t cracking. Many data are available on cryogenic dielectrics under AC stress, but very few under DC stress or under polarity reversal. Oilimpregnated cellulose paper will be ideal from the manufacturing standpoint because it is the standard dielectric for all high-voltage cables. It has also been found that the strengths of the oil-impregnated paper at liquid nitrogen temperature with DC, impulse and polarity reversal are significantly higher than those at room temperature [15 - 16]. However, no data are available at supercritical helium temperatures, i.e. greater than 5.2 K. There is also the likelihood that the oil vapor might contaminate the helium during the cooldown process. It should be evident that the selection of a suitable dielectric material for the DC SPTL is a painstaking process which requires careful long-range planning. The LASL has embarked upon such a program. The dielectric development program has been divided into two parts: initial screening tests and cable sample tests. Nine dielectric materials have been selected for the initial screening tests (Table 2). DC breakdown tests will be conducted on 76 pm (3 m i l ) t h i c k sheets of these dielectric materials at 12 K and 1 MPa (10 atm) of helium pressure to determine the effects of polarity and rate of voltage application. A series of tests will also be performed on cellulose paper impregnated with a specially selected mineral oil. Three dielectric materials will be selected for cable sample tests after analysis of the statistically designed screening tests. During the cable sample tests, the three selected dielectric materials will be tested with DC, impulse, switching surge and polarity reversal in addition to step-stress tests and corona tests under
45 TABLE 2 Dielectric materials selected for screening tests Trade name Cellulose paper Cryovac Kapton Makrofol Mylar Nomex Teflon Udel Valeron
Type of material Cross-linked medium-density polyethylene Polyimide Polycarbonate Polyester Fibrous-spunbonded polyamide Polytetrafluoroethylene Polysulfone Polyethylene
long-term DC stress. A p r o t o t y p e cable will be manufactured with the finally selected dielectric to determine whether all specifications can be met in this realistic configuration. All nine dielectric materials (Table 2) were wound on a 44 mm (laA in.) o.d. aluminum mandrel and cooled down to 4 K in order to determine whether they fracture because of mechanical stress caused b y thermal contraction. The dielectric tapes were 19 mm (3A in.) wide and 76 ~m (3 mil) thick. Two layers of tapes were w o u n d with a 33/67 registration and 3.2 mm (l/s in.) b u t t gap on the aluminum mandrel with a hand tension o f a b o u t 2.2 kg. An additional cool-down test was performed on the cellulose paper impregnated with a special mineral oil. None of the dielectric tapes showed any signs of damage after the cool-down tests. No residue was observed on the oil-impregnated cellulose paper. Electrical characteristics
A DC SPTL has to be designed so that it will operate satisfactorily under normal and abnormal system conditions. The areas of concern are: ripple current produced by the converters, overcurrents produced b y faults, and overvoltages produced by lightning and switching. The ripple current will cause steady-state power losses in the conductor, requiring the refrigeration system to supply cooling power to compensate for these losses. A transient overcurrent will cause transient power losses in the conductor. This may drive the superconductor "normal", requiring, under the worst condition, power interruption to the load. As shown in Fig. 1, the LASL design of the DC SPTL contains four electrically conducting cylinders: the inner
and outer current-carrying conductors, and the inner and outer cylinders of the cryogenic enclosure. Therefore, the usual analysis of surge propagation performed for a twoconductor coaxial cable system will not apply. For the LASL DC SPTL, multivelocity surges will appear on all four conductors and will need special analysis. Studies on the effects of ripple current and fault currents on the DC SPTL have been completed; the overvoltage study is in progress. Harmonic currents on the DC SPTL produced both b y the normal and abnormal harmonic voltage sources have been computed for a 100 km long DC SPTL under various operating conditions. The computations of the normal harmonics are based on the standing-wave theory of long transmission lines. The computations of the abnormal current harmonic were performed by representing the DC SPTL as an equivalent n-network because the line length at these frequencies is less than the quarter-wavelength. The DC SPTL was assumed to be terminated at either end b y a smoothing reactor. No other filter network is connected to the DC SPTL. The results of the analysis are shown in Tables 3 and 4 for the normal and the abnormal harmonics, respectively. The most severe fault current through the DC SPTL will occur if the line-end bushing of the inverter-side smoothing reactor flashes over. The fault current will consist of two components. The first c o m p o n e n t will be a traveling wave caused by the discharge of the cable. The second c o m p o n e n t will be driven by the voltage source on the AC side of the rectifier. The magnitude and duration of the cable discharge current cannot be controlled b y any external means, such as valve control of the converter or a DC circuit breaker. Its magnitude is given b y il = V / Z c
(1)
where V is the system voltage and Z c the surge impedance of the cable. Its duration r at the point of flashover is given by r = 2l/v
(2)
where l is the length of the cable and v the velocity of propagation in the cable. The duration diminishes monotonically along the DC SPTL and is zero at the rectifier end. For
46 LT
TABLE 3 Maximum normal harmonic currents along DC SPTL* Frequency (Hz)
Max. harmonic current (Arms)
720 1440
262.4 34.0
'I
*Line length = 100 km; 12-pulse operation; firing angle delay = 15 ° ; overlap angle = 22.6 °. TABLE 4 Abnormal harmonic currents along DC SPTL* Frequency (Hz)
Max. harmonic current (A rms)
60 120 180 240 300
70.3 35.1 23.4 15.3 8.6
(a)
I
4Lc
LT
L~.
(b)
Fig. 5. Representation of a 12-pulse rectifier bridge with a fault at the inverter end. (a) Schematic diagram; (b) equivalent circuit.
*Line length = 100 km; 12-pulse operation; firing angle delay = 0 - 90°; firing angle tolerance = ±0.25 °.
v~
v b v bv
a 100 k V , 5 GW, 1 0 0 k m l o n g DC S P T L , il = 1 3 . 1 0 k A a n d T = 1 . 1 2 ms. The second c o m p o n e n t of the fault current will b e m a i n t a i n e d b y t h e v o l t a g e sources on t h e AC side o f t h e rectifier. T h e r e f o r e , t h e AC-side r e a c t a n c e s , t h e s m o o t h i n g r e a c t o r a n d t h e DC S P T L r e a c t a n c e will limit this c o m p o nent of the fault current. The profile of the s e c o n d c o m p o n e n t o f t h e f a u l t c u r r e n t will b e dependent upon the means of interruption, such as valve c o n t r o l o f t h e c o n v e r t e r or a DC circuit b r e a k e r . C o m p u t a t i o n s o f t h e s e c o n d c o m p o n e n t of the fault current were made with the following assumptions: (1) t h e f a u l t is i n i t i a t e d at t h e b e g i n n i n g o f commutation; (2) t h e r e is n o c o m m u t a t i o n o v e r l a p ; (3) t h e f a u l t c u r r e n t is i n t e r r u p t e d b y b l o c k i n g t h e firing o f t h e s u b s e q u e n t valves; (4) firing angle d e l a y is z e r o ; (5) c o n v e r t e r s o p e r a t e in t h e 12-pulse mode. T h e 12-pulse s y s t e m w i t h its e q u i v a l e n t circuit f o r t h e c o m p u t a t i o n o f t h e f a u l t c u r r e n t is s h o w n in Fig. 5. In this s y s t e m , t h e f a u l t c u r r e n t has a c h a n c e o f i n t e r r u p t i o n e v e r y ~ / 6 radians o f t h e v o l t a g e wave. This will p r o d u c e a half-sinusoidal s e c o n d c o m p o n e n t o f t h e f a u l t c u r r e n t , as s h o w n in Fig. 6.
6,
A I I @ ~ ~
(a)
00002 ~',: e,
,, -
_.
e
cb V' ObV'aCV'bc
V'
A
2
1
~
A22
(b)
I I I I I
(c) 0
tm ---~ t
Fig. 6. Voltage across 12-pulse rectifier bridge and current profile under fault conditions. (a) Voltage across bridge 1; (b) voltage across bridge 2; (c) fault current. T h e p e a k m a g n i t u d e a n d d u r a t i o n o f t h e fault c u r r e n t are given b y (n + 1 ) ( A l l + A21) + AI2 + A22 imn =
4Lc + LT +/-q
(3)
47
and nn tmn
--
6c~
7n (4)
+ ~
12~
where n is the number of subsequent valves which fail to block; A1z, A12, A2~, A22 are the areas under the voltage wave, as shown in Fig. 6. Lc is the leakage inductance of the converter transformer; Lr is the inductance of the smoothing reactor; and/~ is the inductance of the DC SPTL. The fault currents (second component) of a 100 kV, 5 GW, 100 km long DC SPTL are shown in Fig. 7.
current (imt), including the prefault load current. The current starts to decay once it is transferred to the non-linear resistor of the DC circuit breaker. The characteristics of the non-linear resistor were assumed to be (5)
Va = K i b a
where K and b u , : ~ ; ~ a n t s , va is the voltage across the non-linear resistor, and ia is the current through the non-linear resistor. The decaying total current is given by i(1--b) = i(lmt b) .__ ( K / 4 L c
2.0
[
i
r
f
[
r
I
i
+ L T +/d)(l
--b)t
(6)
I
The constant K was derived on the assumption t h a t Va = 1.7 VDC at imt, and b was assumed to be 0.05. Figure 8 shows the fault current profiles for t d = 5 and 10 ms.
I-Z LU 1.5 (.3
._1 0 F-
hO OT
0
n=O
I
I
I
I
[
IO
n=l
I
, , , i , , , , i
n=2
J
I
I
2 oL ' 20
TIM E (ms)
Fig. 7. Fault currents o f a 100 kV, 5 GW, 100 k m long DC SPTL, cleared by valve control (12-pulse operation). The c o n v e r t e r transformer leakage reactance is 0.18 p.u. and t h a t o f the s m o o t h i n g reactor is 2 p.u.
DC power circuit breakers are n o t commercially available, b u t three types are being investigated [ 17 - 20]. Although the modes of operation of these three types of circuit breakers are significantly different, t h e y are similar in two basic principles: (1) the fault current is n o t limited during the initial period t d of sensing and contact opening; (2) the total current is diverted to a nonlinear resistor during the final period of interruption. Computations o f the fault current were made for two values of td, 5 and 10 ms. During this period, the fault current analysis is the same as t h a t with the valve control except that no valve is blocking. The current at the end of td is the peak of the total
/
/
Q.
>o v
/
/
1,5
\
1.0
\ 0.5 ._J
\ 0
0
I0 TIME (ms)
20
Fig. 8. Fault current i and circuit breaker voltage v o f a 100 kV, 5 GW, 100 k m long DC SPTL, cleared by a DC circuit breaker with delay times t d = 5 and 10 ms. The c o n v e r t e r transformer leakage reactance is 0.18 p.u. and that o f t h e s m o o t h i n g reactor is 2 p.u.
48 APPLICATIONS OF DC SPTL A DC SPTL is particularly attractive for transmitting large blocks of power (2.5 GW and more) from a pow e r park to the urban load center. A DC SPTL may also be attractive to bring power to a large m et r opol it an area from outlying AC transmission networks. Both the economic and technical considerations have to be taken into a c c o u n t in selecting a particular transmission scheme. For the DC SPTL, the cost of t he terminal converter stations has t o be considered in addition to the cost of the DC SPTL and its installation. Comparing it with the AC alternatives, the DC SPTL o f LASL design has only one cable compared with three cables for AC systems. However, the DC terminal equipmen t will be considerably m or e expensive than the AC terminal equipment. Therefore, a cross-over distance exists for each application b e y o n d which t he DC SPTL will be more economical. If, for technical reasons, a DC transmission is chosen, th en the DC SPTL will have to c o m p e t e with the ot her DC alternatives. In this case, one cable o f the LASL DC SPTL will c o m p e t e favorably with t w o cables for the alternative DC transmission schemes. Moreover, in m a ny cases, several pairs of conventional DC cables will be required to transmit the same pow e r that can be transmitted by one DC SPTL of the LASL design. In addition, it is e x p e c t e d t ha t the DC SPTL will offer further savings in converters. As the DC SPTL is inherently a high-current system, the rated voltage of the DC SPTL will be substantially lower than t ha t of the alternative DC transmission schemes for the same transmitted power. It has been r e por t ed that the converter cost increases by a bout 4% for every 100 kV increase in the rated voltage [ 2 1 ] . As converter cost is a significant part of the total transmission cost, this will result in substantial cost savings for the DC SPTL system. As the land area for a lower voltage converter station is smaller than that for a higher voltage station [ 2 2 ] , additional savings are e x p e c t e d to result from converter stations dedicated to th e DC SPTL. Recently, a study was com pl e t e d to estimate the cost of transmitting 10 GW p o wer f r o m a power park to a city center over a 100 km distance. The first contingency
rating of the cable system was also 10 GW, while the second contingency rating was 7.5 GW for 4 hours. Analysis showed that the voltage rating of 300 kV will be the most economical for this application (Fig. 9). The steady-state losses caused by ripple current were f o u n d to be negligible, 2.12% being the total estimated losses of the system, including those o f the converter stations. Fault current calculations showed that the thyristor converters will be able to block in ample time before the specified limit of 2 K t e m p e r a t u r e rise is reached. The relative costs of the various c o m p o n e n t s of the DC SPTL system are shown in Table 5. These costs do n o t include the land cost for the refrigeration system and t hat for converters. I
l
l
75
l
[
l
l
l
GW DC SPTL
"~ i o o o "--
A
8
500 I
J
I
0 LINE
1
l 500
VOLTAGE
I
I
l
I 1000
( k V )
Fig. 9. Cable and enclosure cost of a 7.5 GW DC SPTL as a function of system voltage. Curve A represents today's technology, curve B the expected technology in about 1985, and curve C the expected technology in about 1990. TABLE 5 Relative costs of components of a 7.5 GW DC SPTL system Item
Relative cost (%)
Cable and enclosure Refrigeration system Other Installation Converter Capitalized losses Demand cost for losses Total
11.32 5.54 4.74 10.15 43.65 16.55 8.05 100
49
CONCLUSIONS The DC S P T L is efficient a n d cost effective f o r t r a n s m i t t i n g large blocks o f p o w e r . T h e e c o n o m i c l o w e r limits o f p o w e r a n d transmission distance m u s t be d e t e r m i n e d f o r each a p p l i c a t i o n . The electrical s y s t e m requirem e n t s , such as ripple losses and o v e r c u r r e n t , can be m e t b y a well~lesigned DC SPTL.
ACKNOWLEDGEMENT The w o r k r e p o r t e d in this p a p e r was s u p p o r t e d b y t h e U n i t e d States E n e r g y Research and D e v e l o p m e n t A d m i n i s t r a t i o n .
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