Cold Start of a 240-MVA GSU Transformer Filled With Natural Ester

256
IEEE TRANSACTIONS ON POWER DELIVERY, VOL. 30, NO. 1, FEBRUARY 2015
Cold Start of a 240-MVA Generator Step-Up
Transformer Filled With Natural Ester Fluid
Steven P. Moore, Senior Member, IEEE, William Wangard, Kevin J. Rapp, Member, IEEE,
Deanna L. Woods, Member, IEEE, and Robert M. Del Vecchio, Member, IEEE
Abstract—A 240-MVA, 165-GRDY/20-kV generator step-up
transformer was insulated with Envirotemp FR3 fluid. The
transformer was energized during a cold temperature period in
mid-December 2010. This is believed to be the largest generator
step-up transformer filled with a natural ester fluid that was
energized at low temperatures. The lack of cold start experience of
natural ester-filled power transformers created questions and at
least some uncertainty as to the outcome. This paper describes the
instrumentation used and data collected during the cold startup
procedure. In addition, a computational fluid dynamics analysis of
this transformer design was performed that provided calculated
temperatures and elapsed times until steady-state temperatures
were achieved. Comparisons were made between the actual and
calculated values. The results indicate that the cold startup procedures for a power transformer are the same whether filled with
mineral oil or natural ester fluid of similar characteristics as used
for this paper.
Index Terms—Cold temperature startup procedures, generator
step-up transformer, hottest-spot winding temperature, natural
ester dielectric fluid, top-oil temperature.
I. INTRODUCTION
T
RANSFORMERS containing natural ester fluids are
growing worldwide. The advantages of natural ester
fluid in transformer technology are measured by significant
improvements in fire safety, environmental performance, and
sustainability [1]. However, the performance of the solid-liquid
insulation at operating temperatures encountered during extremes in weather needs to be understood. Natural ester fluid
technology has evolved to include power class transformers
with ratings as high as 400 kV and 240 MVA, with some units
operating since 2004. This paper documents the cold temperature startup of the new 240-MVA (GSU) transformer near
Burlington, IA in mid-December 2010. The idle transformer
Manuscript received January 10, 2014; revised May 01, 2014; accepted June
06, 2014. Date of publication July 02, 2014; date of current version January 21,
2015. Paper no. TPWRD-00030-2014.
S. P. Moore is with the SPX Transformer Solutions, Waukesha, WI 53186
USA (e-mail: [email protected]).
W. Wangard is with the Engrana LLC, Evanston, IL 60201 USA (e-mail:
[email protected]).
K. J. Rapp is with the Cargill Industrial Specialties, Brookfield, WI 53005
USA (e-mail: [email protected]).
D. L. Woods is with the Alliant Energy, Muscatine, IA 52761 USA (e-mail:
[email protected]).
R. M. Del Vecchio is with SPX Transformer Solutions, Inc., Fremont, CA
94538 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.2014.2330514
and surrounding outdoor temperatures at the startup event were
near the pour point of the fluid.
High viscosity of fluids near their pour points causes reduced
fluid flow, resulting in less heat transport. New transformers
with natural ester fluid, especially the GSU type, were generally started during warm seasons or periods of warmer temperatures. Thus, no cold start experience and data were gathered
that related to the actual temperatures inside a large power transformer. The main question is does the cellulose winding insulation in a fully loaded transformer with natural ester fluid at or
near its pour point overheat?
The insulation in power transformers is typically a combination of liquid and solid dielectric materials. The dielectric
strength of natural ester fluid in comparison with other types
of insulating fluids has been measured to 50 C and found
to be equal or better. Previously published work showed that
a single-phase 167-kVA distribution transformer, energized at
full-rated load after equilibration at 30 C in a cold temperature chamber, maintained temperatures well within acceptance
standards [2]. For the distribution transformer study, the insulation materials aged within the industry-accepted rules and the
full unit life expected was maintained.
The three-phase GSU transformer under the cold start conditions in this field study provided important data that were compared to output from computational fluid dynamics (CFD) analysis, using design data in a model of the transformer. Agreement
between the temperature of the field GSU at full expected load
and the CFD provided confidence in the accuracy of the model.
II. PROPERTIES OF NATURAL ESTER FLUIDS
AT COLD TEMPERATURES
Natural ester insulating fluids for transformers are made from
vegetable oils which are a combination of different triglyceride
compounds. Each of the roughly eight major compounds has
unique physical characteristics, such as melting and crystallization or freezing point temperatures. As the temperature of the
fluid decreases, the individual compounds eventually reach their
freezing point and crystallize. The freezing, gelling, crystallizing, and melting of natural ester fluids are processes that take
place over an extended time period and temperature range. As
the solid crystals form, the fluid becomes increasingly opaque or
cloudy. The solid crystals build as more individual natural ester
compounds in the combined mixture crystallize. The increasing
numbers of crystals slowly increase the viscosity of the fluid
because, as the crystals grow, they form larger networks until
they impede the flow so much that the pour point of the fluid is
0885-8977 © 2014 IEEE. Personal use is permitted, but republication/redistribution requires IEEE permission.
See http://www.ieee.org/publications_standards/publications/rights/index.html for more information.
MOORE et al.: COLD START OF A 240-MVA GENERATOR STEP-UP TRANSFORMER FILLED WITH NATURAL ESTER FLUID
reached. The fluid pour point is defined as 3 C above the temperature of fluid in a tube tilted at 90 where the fluid does not
flow within 5 s of time [3].
The natural ester fluid used in the 240-MVA GSU transformer
has a 21 C pour point. The fluid in the cooling ducts would
only cease to flow if held at 21 C for greater than 72 h or
in less time if the fluid gets colder. Thus, the fluid in this temperature range is in an opaque, highly viscous condition, but
still flows slowly to transport heat away from the interior of the
transformer. The transformer loses heat from the interior outward to the tank wall, so the fluid at the tank walls cools first. As
the fluid cools, it slowly crystallizes and becomes more viscous
and acts like a thermal insulator which helps to further reduce
heat loss through the fluid layer at the tank wall. This maintains
the interior temperature of a transformer.
257
TABLE I
COMPARISON OF THE PROPERTIES OF THE NATURAL ESTER FLUID USED AND
CONVENTIONAL TRANSFORMER MINERAL OIL. THE PROPERTIES ARE 25 C
UNLESS SPECIFIED
TABLE II
STEADY-STATE TEMPERATURES IN C AT 200 MVA
III. GSU TRANSFORMER DESIGN AND INITIAL STEADY-STATE
COOLING ANALYSIS
The GSU transformer has a three-phase ONAN rating of 145
MVA and an ONAF rating of 240 MVA. It has inner and outer
HV windings connected in series and an LV winding sandwiched between them. Cooling of the windings is accomplished
by means of directed oil flow using oil-flow washers. The flow
is thermally driven by differences in oil buoyancy at different
temperatures. The external cooling is by means of radiators
with or without fans. In addition, there is some cooling through
the tank walls.
The cooling design for this transformer was based on an
in-house cooling simulation program. This was used previously
in a comparison study of several transformers where mineral
oil and natural ester fluid were used in the same transformer
[4]. In that study, the cooling program results were compared
with test data and temperature differences between the two
coolants were reasonably well predicted. The feature of the
cooling program which makes this possible is that it uses the
thermal parameters of the coolant as an input, including their
temperature dependences. The parameter which differs most
significantly between the two fluids is the viscosity.
The model is based on fluid flow through ducts. The ducts
involved are ducts in the coils through which the coolant is directed and ducts in the radiators. The bulk oil mixing in the
transformer tank is treated by assuming a linear temperature distribution between the cold bottom oil out of the radiators to the
hotter top oil out of the windings. Thus, the coolant properties
of interest are the viscosity, density, thermal expansion coefficient, and thermal conductivity as well as their temperature
dependences. These properties, in turn, affect the fluid friction
through the ducts and the surface heat-transfer coefficient from
the winding surfaces to the coolant. Fluid friction and the surface heat-transfer coefficient are determined by established correlations through unitless numbers, such as the Reynolds, Nusselt, and Prandtl [5].
The heat is generated in the windings by
losses and eddy
current losses. There are also eddy current losses in the structural components, such as the tank and winding clamping structure which must be considered. All of these losses as well as the
coolant properties are temperature dependent. The heat is dissipated through the tank walls and radiators to the air, also using
established correlations for the heat-transfer coefficients. Thus,
an iterative solution method is used to arrive at a steady-state
solution where the input losses are equal to the dissipated losses
for a given power input to the transformer. The temperatures at
this steady state in the windings and coolant can then be printed
out.
Thermal parameters of the natural ester fluid and mineral oil
are given in Table I for comparison purposes. The natural ester
fluid parameters were obtained from the manufacturer [6], [11],
[12], and the mineral oil parameters are taken from [7].
The thermal expansion coefficient affects the density as
a function of temperature T via
(1)
where is the density at temperature . This temperature dependence is important for buoyancy driven flow. The thermal
conductivity and specific heat are fairly constant in the temperature range of interest. However, the temperature dependence
of the viscosity is quite pronounced. For the natural ester fluid,
the viscosity data were fit to (2) where is in
and T is
in
.
(2)
At 10 C, the viscosity of the natural ester fluid is 4.6 times
higher than that of mineral oil and is almost that much higher at
more normal temperatures.
The cooling program was run at the ONAN and ONAF ratings to check the adequacy of the cooling system at these ratings
and to check that the guarantees would be met. However, for the
cold start test conducted in this paper, the transformer was run
under OA conditions (without fans) at a three-phase megavolt
amperes of 200 and an ambient temperature of 3.0 C. The
temperatures of interest are given in Table II.
These steady-state results will be compared with the steadystate results achieved by the more detailed finite-element fluid
flow analysis and with the test results described later.
258
IEEE TRANSACTIONS ON POWER DELIVERY, VOL. 30, NO. 1, FEBRUARY 2015
IV. COMPUTATIONAL FLUID DYNAMICS AND THERMAL
ANALYSIS METHODOLOGY
A. Overview
A fluid dynamics study is conducted for the cold startup at
a 200-MVA load for the transformer containing a natural ester
coolant. The convective heat transfer and fluid motion of the
coolant in the transformer is modeled using a commercial computational fluid dynamics (CFD) software tool. The geometry of
the system is simplified to an approximate 2-D-axisymmetric
domain, including detailed flow through windings and the external heat exchanger. The natural ester coolant fluid is modeled
as a Newtonian fluid.
The thermal energy sources in the transformer arise from
and eddy current losses in the windings, core losses, and tank/
clamp losses. The sources are obtained from an electromagnetic
(EM) model of the windings for the transformer operating at a
reference condition. In the physical transformer operating at a
power level different from the reference condition, scaling laws
are used to compute the instantaneous thermal sources.
Fig. 1. CFD model of the transformer.
B. System Geometry and Computational Domain
The transformer core and windings are confined inside a rectangular tank to which external heat exchanger units are attached.
The heat exchangers or radiators consist of an array of plates that
span vertically between inlet and outlet header pipes that emerge
from the side of the tank. The fluid flow is driven by buoyancy
effects and is not pumped. The heat exchanger has external fans
to increase heat transfer to the environment but they are not used
during startup. Therefore, the external heat transfer to the environment is driven by free convection, conduction, and, to some
degree, by radiation.
The external tank is made of mild steel and is 10 mm thick.
The transformer has three legs, each with the same winding pattern. It is assumed that heat generation from each winding is
identical. Thus, instead of modeling all three legs, only one leg
is modeled, and the system volume is equal to one-third of the
actual tank volume.
The three-phase core is made of high permeability laminations. The windings are made of paper-covered copper wire.
The winding disks are separated by pressboard spacers to enable
horizontal oil flow for cooling. This flow is directed by suitably
spaced washers to force a serpentine flow pattern. Between the
coils are cylindrical pressboard barriers, which isolate flow between the various windings.
A representative image of the cross-section of one of the
transformer legs is shown in Fig. 1. This figure shows the external tank, the heat exchanger, yoke, core, press plates, and
windings. It does not show the flow collars or winding barriers.
C. Physical Heat Exchanger
The physical heat exchanger (PHE) consists of a set of parallel plates all of the same length, fed by an inlet header which
exits the top of the transformer tank. The plates are parallel and
face each other with a gap separation that is fixed. Heat is transferred to the environment by air movement between the plates.
For free convection, the flow is primarily vertical in nature. Radiative heat transfer is negligible. Due to high conductivity of
TABLE III
PARAMETERS OF THE PHYSICAL HEAT EXCHANGER
the radiator plates and tank, conductive heat flow can be ignored.
The heat exchanger consists of
plates of dimensions ,
, and , where
is the length of the plate between inlet
and outlet headers,
is the plate width, and is the hydraulic
diameter of the cross section. Each plate has a lateral total area
given by
. The factor of two accounts for both
plate sides. The dimensions are listed in Table III.
D. Computational Heat Exchanger
Our main problem is that we have to ensure that the 2-Daxisymmetric heat exchanger matches the flow characteristics
of the physical experiment. The computational heat exchanger
(CHE), shown in Fig. 2, is a series of concentric cylindrical
annular channels with length having a uniform gap width
and radial separation . To match the flow characteristics of
the PHE to the CHE, the CHE must have matching hydraulic
characteristics. is not equal to
. must be determined so
that the total surface area of the CHE matches the PHE.
The length of the CHE plates is set equal to the length of
the plates in the PHE, to match the pressure drop for a given
length. Thus,
. Similarly, for the friction factor to match,
the hydraulic diameter of the model must match the physical
experiment. Thus,
. The gap between the model plates
is . The value of must be greater than the plate thickness:
in order to avoid meshing problems.
MOORE et al.: COLD START OF A 240-MVA GENERATOR STEP-UP TRANSFORMER FILLED WITH NATURAL ESTER FLUID
TABLE IV
HEAT-SOURCE DATA.
259
145 MW
has (radial) length
and axial height
. In the experiment,
the radius of the yoke is . Thus, it is can be shown that an
equivalent porosity of the model zone is
(5)
Using table values, the porosity of the yoke zone is
Fig. 2. Model of the CHE.
The total area of the CHE is calculated by summing up the exposed area of a set of circular annuli of incrementally increasing
radii. For a set of plates, the total surface area is given by
(3)
where is the inner radius of the innermost channel and is yet
unknown, but will be determined shortly.
The total area of the PHE is used to transfer heat from all
legs of the transformer. Since our CHE only models one leg, set
. Substituting into (3) yields an equation in terms
of
and . To avoid meshing conflicts, impose the solution
, where
is the radius of the model tank. The
resulting quadratic equation has the solution
(4)
brackets indicate that the solution is rounded
where the
down to the nearest integer. The model tank radius
is
calculated based on the total volume of the tank. From these
values and the expressions developed in the previous section,
. Solving for the number of channels and the
innermost radius of the first channel yields
9 and
1.4639 m.
E. Core Yoke
The CFD model of the yoke must be made porous to accommodate axial flow of the coolant while simultaneously being
able to capture thermal transients. The porosity of the CFD
model is calculated so that the total mass of the solid phase of
the yoke zone matches the experiment. In the model, the yoke
0.832.
F. Model Equations
The equations are modeled using a finite volume formulation
using the commercial software ANSYS Fluent. The equations
are solved using a second-order transient formulation that includes the effects of buoyancy using the Boussineq approximation. See [8] for details of the methods and equations.
In the solid zones, there is only pure convection and heat
generation. The heat generation terms are calculated from the
EM losses in the windings and the core. For each component,
the thermal source has been calculated for a reference operating
condition.
The formula for the thermal source
(in watts) is given by
(6)
is the transformer
where P is the transformer power and
power at the reference operating condition. The load and “no
load” subscripts refer to the components of power at the reference conditions. Table IV gives the load and “no load” thermal
sources and zonal volumes for each region.
The heat generation term in the model is the thermal source
divided by the volume of the respective zone, and is assumed to
be spatially uniform.
G. Fluid Properties
The density of the natural ester fluid is given by the linear
relationship
(7)
where
Equation (7) is a linearized version of (1), using slightly different notation and parameter values.
260
IEEE TRANSACTIONS ON POWER DELIVERY, VOL. 30, NO. 1, FEBRUARY 2015
TABLE V
EFFECTIVE COPPER DENSITIES
TABLE VI
THERMAL PHYSICAL PROPERTIES OF THE COMPONENTS
Fig. 3. Average temperature near Burlington, IA—NOAA Weather 2010.
K. Computational Mesh
The mesh consists of 310 421 quadrilateral cells. The collars
and washers are modeled using impermeable walls. Nonconformal meshes are created between the winding zones to maintain mesh quality. The outer tank thickness is modeled as a solid
zone.
The viscosity of the natural ester fluid is given by (2). The
heat capacity of the natural ester fluid is given by
(8)
where the temperature is in Kelvin.
H. Windings and Other Solid Materials
In the CFD model, the volume occupied by the copper windings includes the paper wrapping, and its effect on the total
mass of the modeled zone cannot be neglected since the mass
of the winding directly affects the thermal response. For each
winding, we calculate the effective solid density as the mass
of the copper divided by the total volume of the zone. Table V
lists the values for the winding zones. The thermal conductivity of copper is
, and its heat capacity is
. Thermophysical properties of the other solid
materials are shown in Table VI.
L. Solution Methodology
The CFD model is performed using ANSYS Fluent v14.0,
which utilizes a finite volume approach to the solution of the
conservation laws [9]. Time stepping is performed using a
second-order algorithm. The convective discretization terms
are second order for momentum and energy. The pressure
discretization utilizes the body-force weighted method, and the
pressure-velocity coupling uses the PISO method. The fluid is
modeled as a Boussinesq fluid with the parameters indicated in
the previous section.
Time steps were limited, based upon stability, to about 0.25
s and the solution was run for about four days. Thus, a total of
about 350 000 time steps were modeled. Each time step converged within 5–10 iterations. The default convergence criteria
were used.
V. COLD STARTUP PROCEDURE AND TIMELINE
A. Weather Station Data for Burlington, IA, USA
I. Boundary Conditions
The bottom of the tank rests on a concrete pad. It is assumed
that heat transfer to the ground is negligible. The heat transfer
from the sides, top, and heat exchanger surfaces is a convective
boundary condition. The local heat flux normal to the boundary
is given by
The outdoor temperature in December 2010 as measured at
the NOAA weather station near Burlington, IA is shown in
Fig. 3. The installed transformer cold startup was performed on
December 13, 2010, which was the coldest average temperature for the entire month. Alliant Energy measured temperatures
and transformer load during the installation and startup period
as summarized in Table VII.
(9)
where the external convection coefficient is given by
. The ambient temperature
varies with time and
its value is given by local meteorological conditions.
J. Initial Conditions
When the transformer is energized, the fluid velocity is zero
and the initial temperature is equal to the ambient temperature.
B. Transient Load and Experimental Temperature Rise of
Transformer Components
On December 12, 2010, the GSU at the Burlington Generating Station was energized after being assembled and vacuum
filled about two weeks before. The unit was allowed to soak
for about 17 h until 6:00 A.M. on December 13, 2010. The criteria was that the unit should soak until the top-oil temperature
reached 5 C at which time, the load was ramped up slowly in
MOORE et al.: COLD START OF A 240-MVA GENERATOR STEP-UP TRANSFORMER FILLED WITH NATURAL ESTER FLUID
261
TABLE VII
TIMELINE AND TEMPERATURE READINGS
Fig. 6. CFD results from the cold start case. Gross transformer load, ambient
temperature, and calculated results of maximum component temperatures.
VI. COMPUTATIONAL RESULTS COMPARED WITH TEST DATA
A. Cold Startup Case
Fig. 4. Transformer at its site.
Fig. 5. Sample thermographic image.
30- to 50-MW increments until reaching the desired capacity of
200 MW.
Had this GSU been filled with mineral oil, the startup procedure would have been the same except for the 5 C criteria.
This additional criterion was added as a unit of this size filled
with natural ester fluid had not been energized from a cold state
before.
Thermographic images were taken of the GSU during the
startup procedure. Temperatures for the tank wall and outermost
radiator panels were recorded and are shown in Table VII. A picture of the transformer is shown in Fig. 4 along with a sample
thermographic image in Fig. 5.
The CFD simulation was run for about four days. Time zero
corresponds to core energization. Load power was applied at
0.693 days at which time the ambient temperature was
13.8 C. This was the start of the computational analysis.
The ambient, maximum winding, and top-oil temperatures
are shown in Fig. 6, along with the gross load. Steady state at 200
MVA was reached after 3.5 days. The low-voltage (LV) winding
had the highest peak temperature of 80 C. The next highest of
75 C occurred in the outer high-voltage winding (HVO). The
inner high-voltage winding (HVI) had a peak temperature of
60 C. The steady-state winding model predicted a peak temperature of 83.2 C for the LV winding, 76 C for the HVO
winding, and 65.5 C for the HVI winding.
In Fig. 7, the temperature of the top-oil sensor is compared to
the CFD model. During the initial temperature rise, prior to the
onset of circulating flow through the heat exchanger, the errors
in the calculation are about 10 C. However, the onset of steady
flow through the heat exchanger stabilizes the flow. After about
1.5 days, the error is a maximum of 4 C and drops to about 1 C
at the end of three days. The top oil reaches 57 C in three days.
The steady-state top-oil calculation is 61.6 C.
The GSU described in this paper is equipped with four fiberoptic (FO) temperature sensors. Two are embedded in the LV
winding and two in the HVO winding at the estimated hotspot locations. Unfortunately, data from these probes were not
recorded at the time of the cold startup. Normally, the transformer is monitored hourly for top-oil temperature, and a simulated winding hot-spot temperature via a winding temperature
indicator (WTI) was calibrated to indicate the LV hotspot in this
case, and fiber-optic (FO) probes if available.
Ideally, the CFD results should be compared with the FO
data. However, the WTI data can serve as a substitute with
some caveats: “It is shown that in case of rapid load change,
the methods used by classic WTI can indicate a lower temperature by more than
even if they are properly adjusted for
steady state conditions.” [10].
262
Fig. 7. Comparison of temperatures predicted at the top-oil location by the
CFD simulation and the recorded measured temperature.
Fig. 8. Average of the two LV winding fiber-optic probe temperatures minus
the WTI temperature versus the rate of change of the WTI temperature.
In April 2013, the GSU was de-energized for three weeks and
then re-energized with a sudden load. For this run, data from
the FO and WTI probes were recorded. The LV FO probe data
minus the WTI data are plotted versus the rate of change of the
WTI temperature in Fig. 8. This plot supports the observations
made by [10]. It shows that the temperature of the WTI can be
as much as 9 C below the FO temperature for the most rapid
temperature change. This plot can be used to correct the WTI
temperatures to indicate the true hot-spot temperatures when the
temperatures are changing rapidly.
The CFD model results for the LV hot-spot temperature are
plotted along with the WTI temperature versus time in Fig. 9.
The temperatures in Fig. 9 track each other fairly well through
the dips and peaks of the power input but are a bit off as steady
state approaches. This could be an artifact of the WTI indicator
as mentioned previously, especially since the upper branch of
the curve in Fig. 8 shows a discrepancy of about C as steady
state is approached on the left. If we add 10 C to the WTI information shown in Fig. 9, the maximum winding temperature
is still only 45 C for a fluctuating load and only 78 C for a
steady-state load. So for this worst case assumption, temperatures are still well below the maximum nameplate operating
temperature. Thus, the cold startup strategy currently employed
is sufficient to avoid overheating the cellulose insulation and
IEEE TRANSACTIONS ON POWER DELIVERY, VOL. 30, NO. 1, FEBRUARY 2015
Fig. 9. LV winding hot-spot temperature comparison of the CFD model calculation and WTI data.
Fig. 10. CFD results from the very cold, sudden start case. Gross transformer
load, ambient temperature, and calculated results of maximum temperatures for
various components.
natural ester fluid for the 105 C insulation system used per the
standard [13].
During temperature rise tests at the factory, the calculated
hottest spot rise for the HV winding was 65.6 C and the measured hottest spot temperature rise using fiber optics was 64.4
C. For the LV winding, the calculated value was 69.6 C and
the measured one was 70.9 C.
B. Very Cold Startup/Sudden Power CFD Calculation
In this case, the initial transformer and ambient temperatures
are fixed at 25 C. At time zero, the transformer power is raised
suddenly to 200 MVA. Initially, the fluid viscosity is quite high
and it is thought that this may inhibit the formation of natural
convection currents to cool the transformer windings.
Results of the simulations are shown in Figs. 10 and 11. In
Fig. 10, the winding and top-oil temperatures and gross load are
plotted versus time. The winding temperatures ramp up almost
immediately and overshoot the steady-state temperatures. It is
not until a significant amount of heating has occurred that there
is enough density gradient in the fluid to push fluid through the
heat exchanger and start the thermo-siphon. This occurs at about
6 h as shown in Fig. 11.
MOORE et al.: COLD START OF A 240-MVA GENERATOR STEP-UP TRANSFORMER FILLED WITH NATURAL ESTER FLUID
263
[13] IEEE Standard for the Design, Testing, and Application of Liquid-Immersed Distribution, Power, and Regulating Transformers Using HighTemperature Insulation Systems and Operating at Elevated Temperatures, IEEE Standard C57.154–2012, 2012.
Steve P. Moore (SM’90) received the B.S. Engineering Technology-Electrical degree from the
Milwaukee School of Engineering, Milwaukee, WI,
USA.
He has been with SPX Transformer Solutions,
Waukesha, WI, USA, and all of its previous company names. Much of his 38-year career was spent
on marketing and sales of new power transformers,
turn key projects, and field service. His current
activities are in the areas of load-tap changers and
the application of natural ester fluid in medium and
large power transformers.
Fig. 11. Heat exchanger mass flow rate for the very cold, sudden start case. The
thermosiphon flow has become fully developed and reached steady state after
8 h.
VII. CONCLUSION
In conclusion, we found that the same cold startup procedures
that are used for power transformers filled with mineral oil can
be used when they are filled with natural ester fluid with similar
physical characteristics, particularly viscosity, to the fluid used
in this transformer. There were no dielectric issues observed
during energization and startup.
The calculated temperatures from the CFD model were close
to the measured temperatures. Even if the load was ramped up
much faster than normal, the hottest spot temperature is not expected to go above the 105 C rating of the insulation system.
REFERENCES
[1] D. Bingenheimer, L. Franchini, E. Del Fiacco, J. Mak, V. Vasconcellos,
and K. Rapp, “Sustainable electrical energy using natural ester technology,” presented at the CIRED 21st Int. Conf. Elect. Distrib., Frankfurt, Germany, Jun. 6–9, 2011.
[2] K. J. Rapp, G. A. Gauger, and J. Luksich, “Behavior of ester dielectric
fluids near the pour point,” presented at the IEEE Conf. Elect. Insul.
Dielectr. Phenomena, Austin, TX, USA, Oct. 17–20, 1999.
[3] Standard Test Method for Pour Point of Petroleum Products, PA 5.01,
ASTM Standard D97, ASTM International, West Conshohocken, PA,
USA, 2012.
[4] R. M. Del Vecchio and R. Ahuja, “Comparison of the thermal performance of FR3, a natural ester based coolant, with transformer oil,” in
Proc. CIGRE, Brugge, Belgium, 2007.
[5] R. M. Del Vecchio and P. Feghali, “Thermal model of a disk coil with
directed oil flow,” in Proc. IEEE Transm. Distrib. Conf., New Orleans,
LA, USA, Apr. 11–16, 1999, pp. 914–919.
[6] K. J. Rapp and J. Luksich, “Transformer design and natural ester
fluids,” in Proc. 12th Int. Elect. Insulation Conf., Birmingham, U.K.,
May 29–31, 2013, pp. 48–52.
[7] F. Kreith and W. Z. Black, Basic Heat Transfer. New York, NY,
USA: Harper & Row, 1980.
[8] F. White, Fluid Mechanics. New York: McGraw-Hill, 1986.
[9] “ Fluent 14.0 User Manual,” ANSYS, Inc., 2013.
[10] J. N. Berbrue, B. L. Broweleit, and J. Aubin, “Optimum transformer cooling control with fiber optic temperature sensors,” 2007.
[Online]. Available: www.neoptix.com/literature/v1165r04_Art_Optimum_Cooling.pdf
[11] Cargill, Cargill Dielectric Fluids reference R2120, Cold Start Recommendations for Envirotemp™ FR3™ Fluid Filled Transformers, May
2013. [Online]. Available: www.environmentalfluids.com
[12] Cargill, Cargill Envirotemp™ FR3™ Fluid reference R2020, Envirotemp FR3 Fluid Behavior in Cold Temperature Environments, May
2013. [Online]. Available: www.environmentalfluids.com
William Wangard received the Ph.D. degree in mechanical engineering from Colorado State University,
Fort Collins, CO, in 2001
He was a Senior Consulting Engineer with Ansys,
Inc. (formerly Fluent, Inc.), Evanston, IL, USA. He
founded Engrana LLC, Evanston, IL, USA, in 2009.
Engrana provides thermal and fluid-flow simulation
consulting services for clients spanning many industrial sectors.
Kevin J. Rapp (M’10) received the B.S. degree
in chemistry from the University of Wisconsin,
Parkside, WI, USA, IN 2003 after completing undergraduate research in lipid-cellulose interactions
chemistry.
He began his career at the Thomas A. Edison Technical Center, Cooper Power Systems, Franksville,
WI, USA, in 1976. In 2012, the Envirotemp FR3
Fluid, which Kevin co-invented, became part of
a new dielectric fluids business unit at Cargill,
Brookfield, WI, USA, where he is currently Senior
Chemist. He is involved in standards work at ASTM, West Conshohocken,
PA, USA, as Chairman of the D27.15 and D27.91 subcommittees, and is
the Technical Advisor for the U.S. National Committee, and for IEC TC10
regarding insulating fluids for electrotechnical applications. He holds many
U.S. and international patents and has published many papers.
Deanna L. Woods (M’11) is a Substation Manager
of Construction, Maintenance, and Engineering for
Illinois Power, Decatur, IL, USA. Currently, she is a
Senior Engineer with Alliant Energy, Madison, WI,
USA, responsible for the substation predictive maintenance program and mentoring substation engineers.
Robert M. Del Vecchio (M’78) received the Ph.D.
degree in physics and the M.S. degree in electrical
engineering from the University of Pittsburgh, Pittsburgh, PA, USA, in 1972 and 1978, respectively .
After serving in various academic positions,
he then joined the Westinghouse R&D Center,
Pittsburgh, in 1978, where he worked on modeling
magnetic materials and electrical devices. He joined
North American Transformer (now SPX Transformer
Solutions), Milpitas, CA, USA, in 1989, where he
developed computer models and transformer design
tools. Currently, he is a Consultant.