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July, 2010
Overvoltage Protectors – A Novel
Concept for Dealing with Overbuilt
Distribution Circuits
Daniel J. Ward
Fellow, IEEE
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IEEE TRANSACTIONS ON POWER DELIVERY, VOL. 25, NO. 3, JULY 2010
1971
Overvoltage Protectors—A Novel Concept for
Dealing With Overbuilt Distribution Circuits
Daniel J. Ward, Fellow, IEEE
Abstract—This paper deals with a simple solution to a
problem that has plagued overbuilt distribution circuits for
decades—namely, protection for sustained overvoltage events.
When two overhead circuits of different voltages exist on the same
poles, the possibility exists that contact between the two circuits
will occur. When this takes place, damage and failures occur to
utility equipment on the lower voltage circuit and to customers’
end use equipment served from the lower voltage circuit. Traditional approaches have met with limited success largely due to
their high costs or operational difficulties. The author’s preferred
solution utilizes polymer housed station class arresters on the
lower voltage circuit to limit the resulting overvoltage. They are
intended to be employed in an expendable mode (i.e., they need to
be replaced following a sustained overvoltage event). A field trial
was arranged to demonstrate the effectiveness of the approach.
This trial proved to be quite successful and cost effective, leading
to the deployment of many more station class arresters for this
function on the Dominion distribution system. A screening technique was devised to aid in applying the solution to a range of
distribution voltage combinations and breaker/recloser clearing
times.
Index Terms—Arresters, overbuilt construction, overvoltage
protection, power distribution lines.
I. INTRODUCTION
O
VERBUILT distribution circuits refers to a practice
where two or more overhead circuits are constructed
on the same poles. If the circuits are of different operating
voltages, overvoltage problems can occur when contact is made
between the two circuits.
Should a short circuit occur between the two circuits, the
lower circuit will experience an overvoltage, resulting in many
customers with damaged home appliances, electronics equipment, HVAC, lighting, etc. Additionally, some arresters and
transformers on the lower circuit will be damaged from the
overvoltage event.
The fundamental problem is that most equipment on the lower
voltage circuit and the end-use equipment served from the lower
circuit cannot withstand the duty from this type of overvoltage
event. On a system of any appreciable size, multiple sustained
overvoltage events can be expected over the course of a year
Manuscript received November 06, 2009; revised February 17, 2010, March
03, 2010, and March 12, 2010. First published April 26, 2010; current version
published June 23, 2010. Paper no. TPWRD-00822-2009.
The author is with Dominion Virginia Power, Richmond, VA 23261 USA
(e-mail: [email protected]).
Color versions of one or more of the figures in this paper are available online
at http://ieeexplore.ieee.org.
Digital Object Identifier 10.1109/TPWRD.2010.2046343
(Dominion averages 13 a year). The problem is not unique to
Dominion; it is an industry problem.
Traditional solutions are quite expensive and at least three
different ones have been tried or considered by different electric
utilities. Each has its own drawbacks.
First, one of the circuits can be relocated or even undergrounded (i.e., placed underground). Very often additional
right-of-way is not available and easements to construct underground facilities may not be available while overhead facilities
are still in place. Even if one could overcome these obstacles,
undergrounding is a very expensive option.
Next, the lower circuit can be converted to be the same operating voltage as the upper circuit. The result—should contact
between the two circuits occur—is a voltage no higher than what
is normally supplied. This option may range from just changing
out the transformers associated with the lower circuit to additional equipment replacement like line insulators. In addition,
the phase spacing may have to be reworked in many cases.
Finally, spacer cable can be used in place of bare wire on the
lower circuit. Insulated wire has greater weight and sag than
what was there before and, as a result, requires shorter span
lengths and more poles for support. A similar concept which can
be retrofitted over existing lines involves insulating the phase
wires of the lower circuit using a medium-voltage line cover.
Like the spacer cable option, should a conductor with a line
cover fall on the ground, there will be very little fault current.
So they both have the disadvantage in that their use can result
in a high impedance fault situation with safety concerns to the
public.
None of the aforementioned methods is a panacea for dealing
with overvoltage events from overbuilt circuits. Most of the solutions have not been widely adopted by utilities partly due
to the high cost and partly due to the operational difficulties
with some of them. A simpler and less expensive solution is
needed—particularly areas where repeat sustained overvoltages
have occurred.
II. INITIAL CONCEPT DEVELOPMENT
The initial objectives of this work were twofold, namely:
1) to find a more cost-effective solution to the sustained overvoltage problem. The solutions need to consider impacts to
both customer as well as utility equipment;
2) to utilize an off-the-shelf solution rather than one customized for each area to be protected.
For Dominion, the overbuilt distribution construction most
susceptible to sustained overvoltage events involves an operating voltage of 34.5 kV for the top circuit and either 13.2 kV or
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IEEE TRANSACTIONS ON POWER DELIVERY, VOL. 25, NO. 3, JULY 2010
Fig. 1. Basic shorted overbuilt circuit—single phase equivalent.
Fig. 3. Overvoltage portion of the ITI Curve (2000 version).
Fig. 2. Fast-acting switch solution.
12.5 kV for the lower circuit. The available fault current at our
34.5 kV distribution substations is 10 kA or less in most cases.
A single phase equivalent of an overbuild circuit is shown in
Fig. 1. The length of common distance between the two circuits
represents the exposure for possible contact between the circuits
and is important in helping establish whether one solution or
multiple solutions will be needed for the exposure zone. For
Dominion, the exposure distance ranges from as little 0.8 km
(0.5 mi) or less to 3.2 km (2 mi) or more.
Any means of rapidly detecting the overvoltage and interrupting the higher voltage source might represent a solution.
One approach would require voltage sensing on the lower
voltage circuit, communications back to the recloser or circuit
breaker on the top circuit and rapid fault clearing by that protective device. Concerns about lengthy overall clearing times as
well as for temporary overvoltages, which would either result
in an unnecessary interruption for a temporary overvoltage
or would result in an imposed delay which would render the
overall scheme ineffective, caused this approach to be dropped
from further consideration.
A simple modification to Fig. 1 shows the addition of a fastacting switch, connected from phase to ground on the lower
voltage circuit, as one means of dealing with the overvoltage
issue (see Fig. 2). As shown, this normally open switch is used
to bypass the affected region of the lower voltage circuit when
sustained overvoltages occur. The switch would have to be able
to respond to overvoltages and be fast enough to limit or prevent
equipment damage.
A series of tests [1] using different distribution arresters, including normal duty and heavy duty, gapped and non-gapped,
with and without isolators simulated dropping one phase of a
46 kV line onto a phase of a 12.5 kV line. The available fault
current for these tests was either 1 kA or 2 kA, so it represented
a fairly weak source. It can be expected that the condition of
the distribution arresters would fare much worse under a higher
fault current environment like 6 kA or so since the heating effect
is proportional to
. The author doubted whether distribution
arresters—with or without isolators—represented a viable solution to the overvoltage problem because reclosing is commonly
employed for distribution circuit breakers and reclosers and the
additional fault duty from reclosing would likely destroy a distribution class arrester. Nonetheless, it was of interest from their
experiments that the secondary voltages were kept within the
overvoltage limits of the 1996 CBEMA curve.
The Information Technology Industry Council or ITI developed a later version of what was previously called the CBEMA
(Computer and Business Equipment Manufacturers Association) curve. The overvoltage portion of the ITI curve [2] is
shown in Fig. 3 below. Survivability of electronic equipment
is more likely if the coordinates for the voltage and time falls
below the solid curve. Electronic devices generally tend to
perform better than indicated by the curve, not so much in
magnitude but in the duration of the event that can be withstood
before the device is damaged.
When 19.92 kV comes in contact with 7.62 kV, the maximum
voltage on the 7.62 kV phase is 2.61 times normal without the
effect of arresters. A maximum value of approximately 23 kV
was shown for a non-gapped 10 kV distribution arrester in [1].
This is 2.1 per unit based on the crest of 7.62 kV. For microprocessor-based equipment to be protected, Fig. 3 shows that the
time must be limited to just under 1 ms for a 2.1 pu voltage. We
will revisit this point later in this paper.
Arrester manufacturers generally favored a solution that
would not only survive the overvoltage event but also be
available for future events. While this permanent protection
approach was understandable, the customized nature of this
approach seemed likely to be a good deal more expensive and
more difficult to apply than one where the arresters would be
allowed to fail and then simply be replaced. The author favored
one that did not have to be customized for each location. As a
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WARD: OVERVOLTAGE PROTECTORS—A NOVEL CONCEPT FOR DEALING WITH OVERBUILT DISTRIBUTION CIRCUITS
1973
Fig. 4. SLGF current versus distance for a 3.7 km (2.3 mi) section.
result, the permanent protection approach was not considered
further.
This thought process led to the idea of employing a polymerhoused station class arrester as a fast switch for this type of
protection. The concept was to use them as sacrificial devices
designed to be replaced following a sustained overvoltage event.
Station class arresters, particularly those with polymer housings, have high energy handling capability and are readily available as standardized products. Installing these on the distribution lines with their proximity to the public made polymer housings a requirement. In addition, station class arresters do not
have isolators and, as mentioned earlier, isolators would have
rendered this protection ineffective because of reclosing.
The existing distribution arrester ratings were based on IEEE
C62.22 [3]. The decision was made to employ the same voltage
rating for the station class arresters as was found on the existing
distribution arresters. The assumption is that the station class
arrester with its lower clamping voltage would hog the energy
from nearby distribution arresters and prevent them from failing.
Insulation coordination issues are covered in [10] and [11].
Furthermore, [11] uses the term sacrificial arresters in reference
to a scheme in which several sets of transmission arresters are
switched in to deal with the overvoltages due to load rejection
on a large 735 kV transmission network. So, from a high level
perspective, sacrificial arresters may have been used before at
transmission voltages, but they have not been employed on distribution circuits.
III. FIELD TRIAL
To demonstrate the feasibility or proof of concept, it was decided to perform a field trial. On August 9, 2007, Dominion
crews installed station class arresters on crossarms below an
East Richmond 13.2 kV circuit. Since the 10 kV arresters were
fairly short in height, wildlife guards were added.
The 34.5 kV overbuilt circuit has an available fault current
of 9,376 A at the substation for a single line to ground fault
(SLGF). The circuit exits the substation via underground cable
for 0.55 km (0.34 mi) and then goes overhead for another 3.2
km (2 mi). The 34.5 kV circuit conductor size is 336.4 kcmil
AAC. Fig. 4 shows the fault current versus distance for the first
3.7 km (2.3 mi) of the circuit.
One set of arresters was installed at the source end of the overhead section (0.55 km from substation in Fig. 4) and another set
1.6 km (1 mi) away. The circuit was selected from other potential overbuild situations because it had 3 sustained overvoltage
Fig. 5. Overvoltage protector (OVP) installation.
TABLE I
DIGITAL RELAY DATA ON FAULT CURRENT MAGNITUDES
events in the previous 4 year period. Fig. 5 shows a picture of
the installation at one of the arrester locations.
On April 14, 2008, a vehicle struck a utility pole near the 1.6
km (1 mi) location causing one of the phase wires from the 34.5
circuit to break and come in contact with a phase of the 13.2 kV
circuit. The station class arresters on the affected phase operated
as planned at both arrester locations.
Both circuits employed two shot reclosing and had digital
relays installed on them so there was data provided (Table I)
about the sequence of operations along with the fault current
magnitudes.
The data from the digital relays indicates that the 13.2 kV
circuit breaker tripped first, followed a second later by the 34.5
kV breaker. The 34.5 kV breaker reclosed 10 seconds later and
tripped out again. A second later, the 13.2 kV breaker reclosed
and tripped out again. Some 34 seconds later, the 34.5 kV
breaker reclosed and then locked out. A second later, the 13.2
kV breaker reclosed and locked out. The clearing times of the
circuit breakers were estimated from the relay settings and the
measured fault current.
As mentioned earlier, two arresters failed from the sustained
overvoltage event. Fig. 6 shows one of the failed arresters. As
can be seen from the figure, the polymer housing split open.
There is evidence of arcing internally and externally. It is likely
that the sustained overvoltage shorted the blocks and the internal
pressure buildup caused the polymer housing to rip open. The
wrapping material around the blocks also split open revealing
a stack of MOV blocks which were intact and in the normal
alignment. The construction was sufficiently rugged to keep the
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IEEE TRANSACTIONS ON POWER DELIVERY, VOL. 25, NO. 3, JULY 2010
Fig. 8. Equivalent circuit for maximum power transfer.
Fig. 6. One of the 2 failed station class arresters.
the higher voltage circuit, the circuit breaker and recloser operating times, the number of reclosures and the impedance of a
failed arrester. Although an MOV surge arrester is a nonlinear
device, once it is subjected to duty beyond its TOV (temporary
overvoltage) capability, the model reverts to a fundamental frequency one in which the resistance of the failed arrester can be
treated as a fixed resistance value.
With an R-L circuit, it can be shown that the maximum power
transfer occurs when the arrester resistance
(1)
Fig. 7. Another failed arrester had expelled a disc.
arrester in the shorted mode throughout the reclosing duty of
both circuit breakers.
Fig. 7 shows the other failed arrester. It showed damage similar to the companion arrester; however, one of the MOV discs
had been expelled.
Approximately 300 residential customers were affected by
the incident. Although there was an interruption, there was no
reported damage at any of the customers’ residences. Usually
affected customers contact our claims group within a few days,
but no claims were ever reported for this event. In addition, there
was no damage to any other Dominion equipment including
the distribution arresters on the 13.2 kV circuit. The operations
crews reported that the restoration process went smoothly. Once
the failed arresters were isolated and the phase wire repaired,
service was restored more rapidly than on any of the previous
overvoltage events. The trial was a huge success.
IV. ENERGY HANDLING CONSIDERATIONS
To better understand what range of voltages and fault currents
this method could be applied to, the author investigated the energy transfer aspect of the problem.
The main factors involved with energy transfer are the operating voltages of the two circuits, the system impedance for
and
are the positive and zero sequence impedances
where
and represents the magnitude of the resulting impedance
See Appendix A for more details on the derivation of (1).
Given the available sequence impedances, permits one to
calculate the
portion of the maximum power delivered by
a given circuit. The time factor accounts for the total duration
of fault current that the arrester sees. The resulting equivalent
circuit is shown in Fig. 8.
When contact occurs from one of the top phases to the underbuilt circuit, the station class arrester on the underbuilt phase
quickly shorts out and, since the arrester is grounded, a single
line to ground fault is initiated on the upper circuit. References
[6], [9], and [12] describe a failure mode in which an MOV arrester when subjected to excessive energy fails due to adiabatic
heating in under 1 ms. This is consistent with the field tests and
the earlier discussion of why the station class arrester approach
permits the overvoltage limits of the ITI curve to be met since
the failure results in the valve elements shorting out almost instantaneously to a fairly low value.
The shorted arrester is subjected to the fault duty for the durations associated with the initial fault plus the subsequent reclosures of the 34.5 kV circuit and the reclosing duty of the 13.2
kV circuit.
For the circuit used in the field trial, the contact between the
two circuits occurred at a location nearly 2.15 km (1.3 mi) from
the substation. At this location, the critical value of resistance
to permit maximum power transfer is 3.32 ohms. At this
value of , the calculated fault current would be 3,923 A. The
corresponding energy through the arrester would be calculated
from
(2)
Applying the data of Table I to (2), the maximum energy delivered is 263,313 kJ. Since the term is fairly high compared
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WARD: OVERVOLTAGE PROTECTORS—A NOVEL CONCEPT FOR DEALING WITH OVERBUILT DISTRIBUTION CIRCUITS
1975
Fig. A1. Special case when load is purely resistive.
Fig. 9. Per unit fault current and I Rt versus the 3r=Z ratio.
to what is expected for a failed arrester, a parametric study was
conducted in which the value was varied over a wide range to
better understand how fault current and delivered energy were
affected.
The effect of arrester resistance on fault current and
is
shown in Fig. 9 with the results given in per unit of the maximum
values. A
value of 1 corresponds to the following values:
an arrester resistance
value of 3.32 ohms, a fault current
value of 3,923 A and an energy value of 231,362 kJ.
Note the breakpoint on energy at a
value of 1 as predicted by (1). At higher values, the fault current and energy
values are less; at lower values, the fault current increases, but
the energy values rapidly drop off.
Since the measured initial fault current from the 34.5 kV circuit (6,066 A) agrees with the calculated value (6007 A), one
can conclude that the actual resistance value for a failed arrester
is more like 0.03 ohms or less. For that resistance value, the energy delivered to the failed arrester is 2920 kJ.
As long as the resistance of the failed arrester is sufficiently
low, there is a high chance of success with the overvoltage protector (OVP) concept.
With the expected energy handling requirement calculated,
one next needs to compare this to the manufacturers’ data on
energy handling. Manufacturers typically rate their arrester handling capability relative to its MCOV rating. With a 10 kV arrester rating (8.4 kV MCOV) for 13.2 kV circuits, the energy
handling requirement from the previous calculations is 348 kJ
per kV of MCOV.
References [4]–[6] show there is really no meaningful arrester
energy handling rating for sustained overvoltage events [4]–[6].
What exists are single shot and two shot energy ratings, but
ones applicable to impulse events and not sustained overvoltage
events. It is not even clear whether different arrester manufacturers rate the energy handling capability on the same basis or
use the same amount of conservatism in their ratings. Nonetheless, the author used the published single shot impulse ratings as
a means of selecting and ranking arresters from competitive offerings. A minimum single shot rating of approximately 9 kJ/kV
of MCOV was used to select an arrester for this application.
While this rating is almost 40 times lower than what was required based on the calculations, this led to trying the concept
in the field with two sets of arresters and hope that the resistance
assumptions were conservative.
V. CONCLUSIONS
The OverVoltage Protector was a simple idea employing
high-energy station class arresters in sufficient quantities (or
with sufficient spacing) to limit the energy absorbed by an
underbuilt distribution circuit during a sustained overvoltage
event. To the author’s knowledge, this concept has never been
tried before.
The successful field demonstration was based on calculations
of maximum power transfer during the most common contact
situation involving overbuilt circuits.
The critical assumption for the overall application was how
much resistance is represented by a failed station class arrester.
Although the author considered his assumptions to be reasonable, tests at a high power laboratory would provide more confidence in the results under varying short circuit conditions.
A screening technique was developed to assess the potential
of using this approach for other combinations of voltages and
source impedances.
The OVP concept is much more cost-effective than any of
the other more traditional approaches undertaken by utilities
dealing with sustained overvoltage events from overbuilt construction.
Since the success of the initial pilot demonstration, Dominion
has installed OVPs in a 30 different locations covering 34.5
kV//13.2 kV and 34.5 kV//12.5 kV overbuild circuit combinations and more are planned.
APPENDIX A
MAXIMUM POWER TRANSFER
With a simple series R-L circuit, if the load impedance is
purely resistive as shown in Fig. A1, then [7] shows that maximum power transfer will take place if the load resistance, ,
is equal to the magnitude of the series impedance (i.e., where
). This is an important result employed in the arrester energy handling calculation.
The situation involving contact between two circuits was analyzed for a single line-to-ground fault occurring immediately
following the instant in which the arrester exceeded its temporary overvoltage (TOV) capability. In other words, the initial
contact between the two phase wires causes the arrester to suddenly change from a relatively high impedance element to a relatively low impedance path which absorbs a considerable amount
of energy until the fault is cleared.
For a single line to ground fault, the symmetrical component
solution (see Fig. A2) shows the series connection of positive,
negative and zero sequence diagrams, all of which are in series
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IEEE TRANSACTIONS ON POWER DELIVERY, VOL. 25, NO. 3, JULY 2010
TABLE II
IMPEDANCE AND FAULT CURRENT VALUES
The total sequence impedance
is equal to
for the circuit of Fig. 4
(6)
APPENDIX B
IMPEDANCES AND FAULT CURRENT
The positive and zero sequence impedance values along with
the calculated fault current values for a single line to ground
fault are given in Table II. The impedances are given in ohms at
34.5 kV.
Fig. A2. Sequence diagram interconnection for a SLGF.
ACKNOWLEDGMENT
The author would like to thank J. Woomer of Dominion for
giving the author the freedom to try a new and promising concept, D. Sullivan of Dominion for his insight on operational issues affecting the installation, S. Barnard of Dominion for his
help with the mounting and installation requirements, and J.
Woodworth (who was with Cooper Power Systems at the time
and is now with ArresterWorks, Inc.) for his continued encouragement before the field trials.
Fig. A3. Simplified version of Fig. A2.
REFERENCES
with 3 times the fault resistance. In this case, the fault resistance
represents the arrester resistance.
By rearranging this circuit to look more like that of Fig. A1
and with
, becomes equal to:
(3)
The resulting simplified equivalent form of Fig. A2 is shown
in Fig. A3.
Thus, maximum power transfer will occur when
(4)
or when
(5)
From the available sequence impedances, the critical value of
resistance will permit one to calculate the maximum energy that
can be delivered from a given circuit. The resistance value for a
failed arrester will then determine the required energy withstand
rating for the arrester.
[1] G. L. Goedde, E. S. Knabe, and L. A. Kojovic, “Overvoltage protection of distribution and low-voltage equipment experiencing sustained
overvoltages,” in Proc. IEEE Power Eng. Soc. Winter Power Meeting,
1999, vol. 2, pp. 1202–1207.
[2] Information Technology Industry Council. Washington, DC, 2005.
[Online]. Available: http://www.itic.org/technical/iticurv.pdf
[3] Guide for the Application of Metal-Oxide Surge Arresters for Alternating-Current Systems, IEEE C62.22-1997, 1997.
[4] J. J. Woodworth, Nov. 4, 2008, Understanding the arrester energy handling issue. [Online]. Available: www.ArresterWorks.com
[5] K. G. Ringler, P. Kirkby, C. C. Erven, M. V. Lat, and T. A. Malkiewicz,
“The energy absorption capability and time-to-failure of varistors used
in station-class metal-oxide surge arresters,” IEEE Trans. Power Del.,
vol. 12, no. 1, pp. 203–212, Jan. 1997.
[6] G. St.-Jean and A. Petit, “Metal-oxide surge arrester operating limits
defined by a temperature-margin concept,” IEEE Trans. Power Del.,
vol. 5, no. 2, pp. 627–633, Apr. 1990.
[7] A. L. Shenkman and M. E. Zarudi, Circuit Analysis for Power Engineering Handbook. Berlin, Germany: Springer, 1998.
[8] A. G. Yost, T. J. Carpenter, G. F. Lincks, H. O. Stoelting, and R. W.
Flugum, “Transmission-line discharge testing for station and intermediate lightning arresters,” IEEE Trans. Power App. Syst., vol. PAS-84,
no. 1, pp. 79–87, Jan. 1965.
[9] M. Bartkowiak, M. G. Comber, and G. D. Mahan, “Failure modes and
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[10] A. R. Hileman, Insulation Coordination for Power Systems. New
York: Marcel Dekker, 1999.
[11] Computational Guide to Insulation Co-ordination and Modelling of
Electrical Networks, IEC 60071-4, 2004.
[12] J. A. Martinez-Velasco, Power System Transients: Parameter Determination. Boca Raton, FL: CRC, 2010.
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WARD: OVERVOLTAGE PROTECTORS—A NOVEL CONCEPT FOR DEALING WITH OVERBUILT DISTRIBUTION CIRCUITS
Daniel J. Ward (F’04) received the B.E. degree from
Stevens Institute of Technology, Hoboken, NJ, and
the M.S.E.E. degree from Union College.
He is a Principal Engineer with Dominion Virginia
Power, Richmond, VA. Prior to joining Dominion
Virginia Power, he spent 21 years at General Electric
(GE), Schenectady, NY, and completed GE’s Power
Systems Engineering Course. At GE, he undertook
numerous studies for GE’s distribution business
units related to distribution transformer, capacitor,
network protector, and meter applications. He also
developed and taught a distribution course as part of GE’s Power Systems
Engineering Course and participated in teaching other GE courses related to
engineering economics and loss evaluation. At Dominion Virginia Power, he is
the Lead Engineer involved in distribution studies aimed at system reliability
improvements. He also manages Dominion’s R&D activities in the distribution
1977
area and is involved with equipment applications and power-quality investigations. He has taught distribution engineering courses at Virginia Polytechnic
Institute and State University, Blacksburg, and Virginia Commonwealth
University, Richmond.
Mr. Ward chairs the ANSI C84.1 Committee on Voltage Standards for Electric Power Systems and Equipment, is past Chair of the IEEE Distribution Subcommittee and the Herman Halperin Award Committee, and past Vice Chair
of the Power Quality Standards Coordinating Committee. He is a member of
the Transmission and Distribution Committee, the Insulated Conductors Committee, and is a registered Professional Engineer in the State of Virginia. He has
authored and coauthored more than 30 technical papers, including four prize
papers. He revised the distribution chapter for the 2006 edition of the Standard
Handbook for Electrical Engineers (McGraw-Hill) and received the Power Engineering Society’s Award for Excellence in Power Distribution Engineering in
2000.
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