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NBS Technical Note 1204

Calibration of Test Systems

for Measuring Power Losses

of Transformers

Oskars Petersons

S,R Mehta

K^^
Tm he National Bureau of Standards' was established by an act of Congress on March 3, 1901. The ^ Bureau's overall goal is to strengthen and advance the nation's science and technology and facilitate their effective application for public benefit. To this end, the Bureau conducts research and provides: (1) a

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The Bureau's technical work is performed by the National Measurement Laboratory, the National

Engineering Laboratory, the Institute for Computer Sciences and Technology, and the Institute for Materials

Science and Engineering

.

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Ciaiihersburg, MI) 20899. 'Some divisions within the center are located at Boulder, CO 80303. located at Boulder, CO, with some elements al Gailhcrsburg, MD.
NATIONAL EXn?EAU OF STANDARDS

LIBRARY

NBS Technical Note 1204

CMbmtion of Test Systems

for Measuring Power Losses

of Transformers

Oskars Petersons National Bureau of Standards

Gaithersburg, MD 20899

and

S.P. Mehta

ASEA Electric, Inc. Waukesha, WI 53186

National Bureau of Standards

National Engineering Laboratory Center for Electronics and

Electrical Engineering

Elect rosystems Division Gaithersburg, MD 20899

Prepared under the NBS Industrial Research

Associate Agreement

August 1985

,ht of

<^o^ J^ »w ^ \.

*'">fAU 0» "^

U.S. Departmoit of Commerce

Malcolm Baldrige, Secretary National Bureau of Standards

Ernest Ambler, Director
National Bureau of Standards Technical Note 1204

Natl. Bur Stand. (U.S.) Tech. Note 1204, 104 pages (Sept. 1985)

CODEN: NBTNAE

us GOVERNMENT PRINTING OFFICE

WASHINGTON: 1985

For sale by the Superintendent of Documents, U.S. Government Printing Office, Washington, DC 20402
FOREWORD

This technical note reports on the work conducted by ASEA Electric, Inc.,

and the National Bureau of Standards under the NBS Industrial Research

Associate Agreement. The calibration system was built and experiments were conducted by ASEA Electric, Inc., under the design and development guidance of

NBS. Several special calibrations v;ere performed by NBS for this project on a reimbursable basis following the usual NBS practices.

Certain commercial equipment, instruments, or materials are identified in

this paper in order to adequately specify the experimental procedure. Such

identif icaion does not imply recommendation or endorsement by the National

Bureau of Standards, nor does it imply that the materials or equipment are necessarily the best available for the purpose.

The use of the developed calibration instruments might be associated with

hazardous voltages. The calibration instruments must be used in conformance

with the applicable safety codes, standards, and practices.

iii
ACKNOWLEDGMENTS

During the development work and preparation of this report, the authors

received assistance from numerous colleagues at ASEA Electric, Inc., the

National Bureau of Standards, and other organizations. Jeffrey J. Nemec

constructed the equipment and conducted the experiments. Arnold G. Perrey

and Wilbur C. Sze provided valuable guidance in the design of the standard

current transformer and active divider. W. E. Anderson, N, M, Oldham, and

T. M. Souders performed special high accuracy calibrations. C. P. DeWitt of

Holt Instrument Laboratories kindly provided the design information on special

shunts. Betty Meiselman typed the manuscript under tight deadline conditions.

Special thanks to Steven P. Moore and Albert C. Swenton for supporting the

project and for their guidance related to measurement requirements. The

authors gratefully acknowledge the following readers of the manuscript:

B. C. Belanger, N. M. Oldham, S. P. Moore, W.J.M. Moore, K. R. Reed, E. So,

and J. R. Sorrells.

IV
EXECUTIVE SUMMARY

It is estimated that average annual capitalized cost of losses in

transformers in the U.S.A. is around two billion dollars. The economic impact

of power transformer losses is at a state where the cost of evaluated losses

over the lifetime of a power transformer often rivals the initial price of the

transformer.

With increasing cost of electric energy over the past decade, demands for

improvement in efficiency of power transformers has been rising. Improvement

in efficiency is achieved by reduction in transformer losses.

Power factor during the load loss test is a ratio between the effective

resistance (losses) of a transformer and the effective impedance of a transformer. Reduced losses with effective impedance remaining essentially

the same means that the power factor during the load loss test gets smaller as losses are reduced.

Accurate measurement of losses in transformers at low power factor poses

difficult technical problems. At low power factor a small phase angle error anywhere in the test system (such as in voltage transformers, current

transformers, and power transducers) can cause large error in the measurement

of real power. For example, a very small 290 microradian phase error at 1%

power factor produces a 2.9% error in the measurement of real power.

Although accurate measurement of transformer losses is very critical,

requirements for measurement accuracy and traceability of measurements to

higher echelon standards is not addressed in any industry standards.

High phase angle accuracy devices have been developed and utilized in

various laboratories for accurate measurements at low power factors. However,

this technology in general is not uniformly utilized for critical measurements

in industrial environment.

*****

In this report a system is described that was developed to calibrate and

verify the accuracy of test equipment utilized for measuring the losses of

large power transformers. Several standard instruments were either

constructed or acquired from commercial sources. Methodologies were developed

for (a) calibrating and aligning the test systems, (b) for calibrating and

maintaining the standards themselves and (c) for ensuring the traceability of

measurements

.
Specifically this report describes:

1) Development and construction of a two-stage transformer and ancillary

devices having a phase angle uncertainty of ±13 yrad.

2) Development and construction of an active voltage divider and

standard capacitance networks having a phase angle uncertainty of ± 14

yrad.

3) Development of an active phase shifter for the calibration of power

transducers having a phase angle uncertainty of ±20 yrad.

4) Methodology for alignment of voltage, current, and power transducers

in the test system for loss measurement.

5) Methodology for determining measurement system error vs. power

factor.

6) Results of system based verification where simultaneous measurement

of losses with a calibration system and test systems are carried out.

7) Requirements for calibration and maintenance of standards and

alignment of the test system to maintain traceability to the higher

echelon standards.

It is hoped that this report will contribute to the development of

industry standards that include accuracy and traceability requirements for

test systems utilized for the measurement of transformer losses. Only in this

way can there be fair and uniform comparisons of transformer losses by the

purchasers of power transformers.

vl
TABLE OF CONTENTS

Page

FOREWORD iii

ACKNOWLEDGMENTS iv

EXECUTIVE SUMMARY v

TABLE OF CONTENTS vii

LIST OF FIGURES ix

LIST OF TABLES xii

ABSTRACT 1

1. INTRODUCTION 1

2. TEST SYSTEMS FOR TRANSFORMER LOSSES 5

2.1 General Description 5 2.2 Test System 1 7 2.3 Test System 2 9

3. CALIBRATION APPROACH 11

4. CALIBRATION OF CURRENT CHANNEL 15

4.1 General Approach 15

4.2 Two-stage Current Transformer — Theory 19

4.3 Two-stage Current Transformer — Construction 22

4.4 Functional Tests of Current Transformer 25

4.5 Precision Resistors 27

4.6 Calibration Data 27

5. CALIBRATION OF VOLTAGE CHANNEL 31

5.1 General Approach 31

5.2 Standard Voltage Divider 33

5.3 Standard Capacitor Networks 35

5.4 In-House Calibration of Standards 37

6. CALIBRATION OF POWER TRANSDUCERS 41

6.1 Standard Transducer 41

6.2 Standard Phase Shifter 41

7. ALIGNMENTS OF TEST SYSTEMS 43

7.1 Test System 1 43

7.2 Test System 2 47

vn
Page

8. ANALYSIS OF MEASUREMENT ERRORS 51

8.1 General Comments 51

8.2 Standard System 53

8.3 Test System 1 55

8.4 Test System 2 55

9. SYSTEM- BASED VERIFICATION 62

9.1 Experimental Method and Results 62

9.2 Discussion of Experimental Results 62

10. MAINTENANCE OF STANDARD INSTRUMENTS AND TEST SYSTEMS 66

10.1 General Comments 66

10.2 Standard Current Transformer 67

10.3 Standard Shunts 67

10.4 Standard Capacitors 68

10.5 Resistors in the Phase Shifter 69

10.6 Inductive Voltage Divider 69

10.7 Standard Power Transducer 69

10.8 Maintenance of Test Systems 70

11. DISCUSSION AND CONCLUSIONS 72

12. REFERENCES 74

13. APPENDIX 1 — CALIBRATION REPORTS AND DATA 76

13.1 Introductory Comments 76

13.2 Two-Stage Current Transformer 77

13.3 Two-Stage Current Transformer/Shunt Combination 79

13.4 Standard Capacitor 80

13.5 Capacitor Networks 0.01, 0.1, and 0.1 Microfarad 81

13.6 Inductive Divider 83

13.7 Power Transducer 87

14. APPENDIX 2 — SUGGESTED CALIBRATION PROCEDURE AND ACCURACY GOAL ... 90

14.1 Calibration Procedure 90

14.2 Accuracy Goal 91

viil
LIST OF FIGURES

Page

1. Conceptual Circuits for Measuring Transformer Losses 6

(a) Based on voltage and current transformers

and a wattmeter

(b) Based on a current transformer, a high voltage

divider and a power transducer

(c) Based on an RC-type impedance bridge

(d) Based on a transformer-ratio-arm bridge

2. Test System 1 8

Block diagram showing principal components

3. Test System 2 10

Block diagram showing principal components

4. System Based Verification 13

5. Instrumentation for Calibration of Test Systems 14

Test system 2 (console), two-stage current transformer

(on desk of console), high voltage standard capacitor

(at right), null detector (on console at left), standard

power transducer, electronics for active voltage divider

(on top of power transducer), 1000-pF standard capacitor,

precision DVM, two 2-ohm shunts, inductive voltage divider,

three precision capacitive networks (on top of inductive

divider)

6. Calibration of current channel, V^>V 16

7. Calibration of current channel, V>Vx 16

8. Bipolar Buffer Amplifier 18

(a) Schematic circuit diagram

(b) Magnitude and phase angle adjustments; two circuits

are required

9. Two-Winding Transformer 20

(a) Circuit diagram

(b) Equivalent circuit

10. Two-Stage Transformer 20

(a) Circuit diagram

(b) Equivalent circuit

(c) Simplified equivalent circuit

11. Two-Stage Transformer with Separate Burdens 20

12. Details of Standard Two-Stage Current Transformer 23

IX
Page

13. Standard Two-Stage Current Transformer in Enclosure 24

(a) Primary terminals

(b) Secondary terminals

14. Checking of Two-Stage Transformer Against Inductive Divider 26

(a) Taps in the secondary winding

(b) Ratio of two windings

15. Self-Checks of Two-Stage Transformer 26

(a) 0.1 tap against 100-T winding

(b) Adjacent taps against 100-T winding

(c) 10-T winding against 100-T winding

16. Coaxial Shunts, Cross Section 28

17. Coaxial Shunts, External View 28

18. Calibration of Voltage Channel 32

(a) Same polarity

(b) Opposite polarity

19. Active Divider with Controlled Source 34

(a) Open-loop configuration

(b) Closed-loop configuration

20. Implementation of Controlled Source in Active Divider 34

21. Low Voltage Capacitor Network 36

(a) Circuit diagram

(b) Equivalent circuit

22. Components of Active Divider 38

(a) Low voltage capacitor networks

(b) Enclosure containing electronics

23. Capacitance Bridge Based on Inductive Voltage Divider 39

(a) Comparison of small capacitors

(b) Comparison of large capacitors

24. Capacitance Bridge Based on Two-Stage Current Transformer 39

25. Active Phase Shifter 42

(a) For zero power factor

(b) For power factors of 0.00, 0.01, 0.02, 0.05,

and 0.1; shielding omitted

26. Current Channel Alignment, Test System 1 44

2 7. Voltage Channel Alignment (Coarse), Test System 1 45

28. Voltage Channel Alignment (Fine), Test System 1 . . . 46
Page

29. Power Transducer Alignment, Test System 1 48

30. Standard System, Limits of Errors 57

31. Test System 1, Limits of Errors 59

32. Test System 2, Limits of Errors 61

33. System-Based Verification, Test System 1 65

Comparison of results as obtained by the test system 1 and

standard system

XT
LIST OF TABLES

Page

Table 1. Calibration of Standard Current Transformer 29

Table 2. Calibration of Trans impedance of Standard Current

Transformer - Shunt Combination. 30

Table 3. Calibration of Standard Current Transformer,

Taps in Secondary Winding 30

Table 4. Current Channel Alignment, Test System 1 49

Table 5. Voltage Channel Alignment, Test System 1 49

Table 6. Power Transducer Alignment, Test System 1 49

Table 7. Current Channel Alignment, Test System 2 50

Table 8. Voltage Channel Alignment, Test System 2 50

Table 9. Power Transducer Alignment, Test System 2 50

Table 10. Errors of Components of Standard System 54

Table 11. Errors in Current Measurement, Standard System 54

Table 12. Errors in Voltage Measurement, Standard System 55

Table 13. Errors in Power Measurement, Standard System 56

Table 14. Combined Errors in Power Measurement, Standard System 56

Table 15. Errors in Power Measurement, Test System 1 58

Table 16. Combined Errors in Power Measurement, Test System 1 58

Table 17. Errors in Power Measurement, Test System 2 60

Table 18. Combined Errors in Power Measurement, Test System 2 60

Table 19. System-Based Verification, Test System 1 64

Table 20. Comparison of Inductive Divider against

Two-Stage Transformer 86

Table 21. Suggested Goal for Limits of Errors of

Future Test Systems 91

xn
CALIBRATION OF TEST SYSTEMS

FOR MEASURING POWER LOSSES OF TRANSFORMERS

Oskars Petersons

National Bureau of Standards

Gaithersburg, MD 20899

S. P. Mehta

ASEA Electric, Inc.

Waukesha, WI 53186

ABSTRACT

A calibration system for accuracy verification and alignment

of test systems for measuring transformer losses is described.

Methodologies are presented for assessing measurement uncertainties

and for evaluating overall accuracy of test systems. Procedures

are suggested for continuing maintenance and calibration of standard

instruments and test systems to ensure traceable measurements.

Keywords: calibration; loss measurement; low power factor;

power transformers; traceability

1 . INTRODUCTION

This report describes a system that was developed to calibrate and verify

the accuracy of test equipment utilized for measuring the losses of large

power transformers. Several standard instruments were either constructed or acquired from commercial sources. Various methodologies were developed for

calibrating and aligning the test systems, for calibrating and maintaining the

standards themselves, and for ensuring the traceability of measurements to

higher echelon standards. Because of the low power factor, measurement of

load or conductor loss is particularly difficult and subject to significant

uncertainties. Hence, the calibration system is particularly designed to

reduce uncertainties that relate to measurements of load or conductor losses,

but the system can be utilized equally well to verify the test system for

measurement of no-load or core losses.

At low power factors a small phase error anywhere in the test system such

as in the current and voltage transformers, and in the power transducer, can cause a large error in the measurement of the real power. This error increases as the power factor decreases. For example, when measuring the

losses of a transformer having a power factor of 0.01, an error in the phase

angle of one minute (0.29 milliradian) causes an error of 2.9% in the measured

power. If an error of 0.29 milliradian is probable in each critical

component, — voltage, current, and power transducer, — the combined

root-sum-of-squares error in the power measurement becomes 5.0%. While an uncertainty in the phase angle of 0.29 milliradian is not at all difficult to

achieve in laboratory conditions, it may not be so in industrial testing
areas. For example, the standard for instrument transformers, ANSI C5 7.13,

specifies the lowest calibration uncertainty of ±3 minutes for both the

voltage and current transformers that are used in revenue metering

applications. For applications discussed here, ten to fifty times smaller

phase angle errors are required for each of the critical components in the

test system, in order to maintain the overall measurement uncertainty for real

power in the one percent range.

The phase angle error of various laboratory devices — such as special

voltage and current transformers, inductive dividers, gas-dielectric

capacitors, and voltage dividers based on such capacitors — is well within

the ±50 yrad level. However, this measurement and instrumentation technology

is not universally available to the potential users in industry. Also the

calibration infrastructure whereby a test system could be readily verified and

made traceable* [1] to higher echelon physical standards and whereby tests

could be repeated and the results compared among various suppliers and users

is not well established.

While it is acknowledged that accurate measurements of transformer losses

is a difficult technical problem, there are strong economic and technical

motivations for improving the measurement accuracy and for establishing a uniform basis for assessing the measurement accuracy. This relates directly

to the increasing cost of electric energy over the past decade which brought

about a growing demand for improvements in power transformer efficiency. In

order to minimize the estimated 10% energy loss between the points of

generation and the end use of electricity [2], power transformer

specifications almost invariably include evaluation of transformer losses in

determination of the total owning cost of the transformer.

Economic impact of power transformer losses is at a state where over 90%

of inquiries include loss evaluation and the cost of evaluated losses often

rivals the initial price of transformers. Loss evaluations as high as $10,000/kW for no-load loss and $5,000/kW for load loss are not uncommon with

the average being about $4000/kW for no-load loss and about $2000/kW for load

loss. It is estimated that the average annual equivalent capitalized cost of

losses in transformers in the U.S.A. is around two billion dollars. In this

climate, accurate measurement of transformer losses has received considerable

attention in recent years.

*The term "traceability" is frequently used with measurements and

calibrations. However, a universally accepted definition for this term has

evaded formulation [1]. The definition 3 and parts of definition 2 of

reference 1 summarize well the key elements of this term. Quoting from

this reference:

"Traceability means the ability to relate individual measurement results

to national standards or nationally accepted measurement systems through an unbroken chain of comparisons."

"Measurements have traceability to the designated standards

if .. .scientifically rigorous evidence is produced. . .to show that the

measurement process is producing measurement results (data) for which the

total measurement uncertainty relative to national or other designated

standards is quantified."
From the manufacturer's viewpoint, improvements in loss measurement

accuracies are highly desirable because any upgrading in manufacturing

processes, in materials, and in designs intended for reduction in losses would

be difficult to evaluate without accurate measurements of transformer losses.

From the user's viewpoint, the valididy of the whole selection process,

including fair comparison of losses becomes questionable without accurate

measurement of transformer losses.

Standardization ; It is very important that the issue of loss measurement

accuracy be addressed in the industry standards so that uniform comparisons

and evaluations of transformer losses can be realized. At present, industry

standards dealing with transformers do not address the issue of transformer

loss test accuracy. However, work is under way to review the critical factors

that have impact on loss measurement accuracy. Appropriate correction factors

and their proper use are being developed in a working group of the Transformer

Committee of the Institute of Electrical and Electronics Engineers.

In addition to accurate loss measurements, it is equally important that

the accuracy of the test system be independently verifiable and traceable to

nationally accepted standards. The National Bureau of Standards (NBS)

calibrates certain high-accuracy physical standards such as instrument

transformers, high-voltage capacitors, and power transducers, all of which

could be applied to the calibration of the test system for loss measurements.

NBS also receives inquiries from the industry about viable calibration

alternatives for such test systems. Occasionally requests have been received

to perform spot checks of actual on-site test systems. The NBS interaction

with the industry in this area has been hampered by an inadequate measurement

infrastructure: unavailability of stable physical standards that could be

calibrated at NBS and then used to verify the test system; instabilities in

the physical standards making it difficult to transfer the calibration from

NBS to the user and to maintain it at the user's facility for any length of

time; and nonunif ormity in the input and output quantities of the standard

instruments thus making it difficult to calibrate them at NBS without

resorting to special and very expensive calibration setups.

Calibration of the actual test system by NBS on-site has some merits but

also serious disadvantages. Admittedly, such NBS activity could be considered

a welcome addition to the in-house calibration procedures since it would

provide additional assurance of the validity of the procedures used. However,

without proper in-house procedures, NBS on-site calibration might be of

limited value. Time constraints and expense would dictate that only a few

test points be covered; there would be little time for uncovering the causes

of erroneous measurements ; and there would be no assurance of how well the

test system will behave in the future.

In response to requests for advice on calibration procedures, NBS has

recommended that industrial organizations become self-sufficient in

calibration capabilities insofar as it is economically feasible so that only

a minimal number of standard instruments would have to be sent to outside

calibration organizations such as NBS. ASEA Electric was interested in

establishing such a calibration system and sought NBS advice. Arrangements

were made within the NBS' Industrial Research Associate Program whereby NBS

will provide to ASEA Electric design and development guidance for the
establishment of physical calibration standards and procedures. ASEA Electric

will construct and purchase the necessary instruments and will conduct the

experiments. Results of the investigation will be published in the open

literature.

One of the goals was to establish a calibration system that is as much

as possible self-contained, requiring only relatively few and infrequent

calibrations from outside. A second goal was to have the calibration system

traceable. Thirdly, for the two principal components in the calibration

system — the standard current transformer and the voltage divider — the

initial maximum uncertainty goal was -^30 Mrad each, and ±50 yrad for the third

component — the standard power transducer — leading to the combined rss uncertainty of ±66 yrad. Actual results were better than this goal, with an uncertainty of about ±33 yrad. After alignment of the test systems against

the calibration standards, it was found that the factors that limit the

accuracy are predominantly the residual nonlinearities in the transducers and

the associated electronics in the test systems.
2. TEST SYSTEMS FOR TRANSFORMER LOSSES

2. 1 General Description

Test systems for measuring transformer and shunt reactor losses are supplied by several manufacturers. Such systems may differ from each other In

fine detail, however, their general layouts have considerable commonality.

All test systems contain precision devices to scale (step down) the test

voltage for the measurement purpose. The voltage scaling devices can be

either precision voltage (potential) transformers or gas-dielectric

capacitors. Current scaling is accomplished with precision current

transformers, which also isolate the measuring circuit from high voltage.

To accommodate three phase devices, current is usually measured in the high

voltage circuit. The actual power measurement is performed with a wattmeter

(power transducer) designed to have low phase angle uncertainty, or with an impedance bridge.

Conceptually, the test systems for transformer loss measurement are illustrated by the four circuit diagrams in Fig. 1. In the circuit of

Fig. 1(a), the test voltage and current are reduced with precision instrument

transformers and the power is measured with a low-power-factor wattmeter. In

Fig. 1(b), an active voltage divider is employed to scale down the voltage.

The reduction of current is achieved with a current transformer to which an amplifier is connected in order to obtain the signal with the required

amplitude; the outputs of the two amplifiers are voltages, hence a voltage

multiplier-type of transducer is used. In Fig. 1(c), a bridge technique is

illustrated. This particular bridge, employing a high voltage capacitor and a current transformer, has some similarities to the classical Schering and

Maxwell bridges [3]. In Fig. 1(d), another bridge technique employing a high

voltage capacitor, a current transformer, and a transformer-ratio-arm bridge

is shown [4,5,6].

For single-phase measurements the current transformer can be connected in

the ground lead as in Fig. 1(d); in three-phase measurements, the current

transformers are in the high voltage lead as in Figs, (a), (b), and (c).

Special high accuracy current transformers such as zero-flux transformers,

two-stage transformers, or amplifier-aided transformers are necessary.

Similarly, high accuracy, especially for the phase angle, is required in the

voltage transformers, the high voltage dividers, or the capacitors. The

wattmeter or watt transducer should be suitable for measurements at low power

factor and should have small phase angle uncertainty. Methods employing

wattmeters or power transducers are most suitable for power factors in the

approximate range of one percent or higher. For power factors substantially

below one percent as encountered in shunt reactors, bridge techniques may have

to be considered.

ASEA Electric utilizes two independent test systems for measuring

transformer losses. Both systems are of the general type shown in Fig. 1(b)

enabling measurement of losses on a three-phase basis using the

three-wattmeter method. The first of these two systems was placed in service

in 1971, the second in 1980. Following is a brief discussion of the salient

features of these two systems.
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2.2 Test System 1

A simplified schematic representation of test system 1 is shown in

Fig. 2. The system is divided into three principal subsystems — the first

for scaling and measuring the current, the second for scaling and measuring

the voltage, and the third for measuring the power.

The current subsystem (or the current channel) includes scaling of the

primary current by means of a zero-flux current transformer. Three such

transformers, one for each phase, are inserted in the high voltage line. The

transformers are gas insulated for 100 kV. Secondary currents from the

working and control windings are passed through the main and compensating

burdens located in the console. Separate sets of burdens are provided for

each of the four ranges of the primary current — lOA, lOOA, lOOOA, and

10,000A. Compensating burdens are adjusted during the calibration and

alignment process of the test system. The voltage at the main burden is

supplied through a precision buffer amplifier to the rms current transducer,

peak current transducer, and power (watts) transducer. The design of the

amplifiers permits small adjustments in the in-phase and quadrature components

of gain in order to facilitate the alignment of the system.

The voltage subsystem (or the voltage channel) utilizes a gas-dielectric

high voltage capacitor in an operational integrator configuration as shown to

the right in Fig. 2. The three ranges for the primary voltage — 1 kV, 10 kV,

and 100 kV — with a corresponding secondary voltage of 10 V, are achieved by

changing the gain in the ranging amplifier that follows the integrator. Small

adjustments can be made in the gain of the ranging amplifier. The output

voltage of the ranging amplifier is supplied to the rms, average and peak

voltage transducers and to the power transducer.

The power measuring subsystem consists of a power transducer (time

division multiplier) and scaling amplifiers, both at the input and output of

the transducer, to enable operating the transducer near its full scale. The

gains of the output amplifiers are reciprocals of the gains of the input

amplifiers.

The principal accuracy-related specifications as originally supplied by

the manufacturer of the test system are the following:

a. voltage (rms, average, peak): ±0.5% of reading; from 10% of full

scale to full scale; three ranges — IkV, lOkV, lOOkV

b. current (rms, peak): ±0.5% of reading; from 10% of full scale to full

scale; four ranges — lOA, lOOA, lOOOA, 10,000A

c. power: ±5.0% of reading at power factor of 0.01;

±0.5% of reading at power factor of 0.1;

0.1 kW to 10 MW in 72 ranges.

In order to improve the accuracy of power measurements, especially at low

power factors, several of the amplifiers and the power transducer were rebuilt

to take advantage of more stable components and circuit configurations.

However, the overall functional layout of the test system was left unchanged.

The accuracy figures discussed in section 8 are those for the rebuilt system

after it was calibrated and aligned.
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A simplified block diagram of this system after modifications is shown in

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following exceptions: current scaling is accomplished with a two-stage current

transformer; the maximum operating voltage is 230 kV, the maximum current is

lOOOA; as originally built, this system did not include scaling amplifiers in

conjunction with the power transducer.

The principal accuracy-related specifications of the test system before

alignment were as follows:

a. voltage (rms, average, peak): ±0.3% of reading; at 10% of full scale

and above; four ranges ~ 1 kV, 10 kV, 100 kV, and 230 kV

b. current (rms, peak): ±0.3% of reading; at 10% of full scale and

above; four ranges — lA, lOA, lOOA, lOOOA

c. power: ±0.3% of reading at power factor of 1.0;

±0.6% of reading at power factor of 0.1;

±4.0% of reading at power factor of 0.01;

0.1 kW to 100 MW in 48 ranges

Accuracy improvements obtained after the modification, calibration, and

alignment are discussed in section 8.
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10
3. CALIBRATION APPROACH

The two test systems for which the calibration system is designed are of

the type shown in Fig. 1(b), consisting of amplifier-aided capacitive

dividers, special current transformers connected to the amplifiers, and power

transducers (voltage multipliers).

It is desirable to verify the accuracy of a test system on a system basis

by simultaneously measuring voltage, current and power with the test system

and with the calibration (standard) system. Such a comparison can be

conveniently performed on a single-phase basis. A system-based verification

is essential as the final check to ensure that there are no adverse

interactions among the components in the test system and that various

electromagnetic interferences do not cause excessive errors. A system-based

verification by itself is insufficient when large discrepancies are found in

the measurements, since the causes of such discrepancies are more difficult to

identify. Since such discrepancies could be caused by malfunctions in the

test system, it would be an undesirable practice to correct the final result

without correcting the cause of the discrepancy.

In recent years significant developments in accurate measurements of

losses in the power system apparatus at very low power factors have been

reported. Most of these developments have as their basis the

current-comparator-based bridge techniques [4,6,7,8,9]. These developments

are directly applicable for the system-based verification of the test systems

utilized for measurement of losses in power system apparatus.

The approach taken in designing the calibration system described in this

report was to make it suitable not only for a system-based verification but

also for verification of the principal components in the test system — the

high voltage divider, the current transformer, and the power transducer.

Further considerations for selecting this approach were the architecture of

the two test systems at ASEA Electric (it is a commonly used architecture in

other test systms) ; desirability to have a calibration process that is as much

as possible self-contained requiring relatively few calibrations from outside;

realistically achievable and maintainable accuracy in an industrial test

environment; and cost effectiveness of the total approach. The calibration or standard system consists of four basic parts:

1. a two-stage current transformer with precision shunts; for

calibration and alignment of the current transformer and its

ancillary components in the test system, for supplying signal to

current and power measuring instruments;

2. an active voltage divider based on a compressed-gas high voltage

capacitor; for calibration and alignment of the voltage (potential)

transformer or the high voltage capacitor and its ancillary

components in the test system, for supplying signal to voltage and

power-measuring instruments

3. a standard power transducer (voltage multiplier); for calibration of

transducers in the test system; and

4. an active phase shifter also to calibrate the power transducers at

various power factors, e.g., 0.00, 0.01, 0.02, 0.05, and 0.10.

11
With such a calibration system, the test systems in Fig. 1(a) and Fig. 1(b)

and their principal components can be readily calibrated. With a few

additional components, the calibration system can also be adapted for the

verification of the impedance bridges as shown in Fig. 1(c) and Fig. 1(b).

Sections 4 to 6 describe the calibration procedures in general, while

section 7 presents the procedures and calibration results for the two specific

test systems at ASEA Electric.

The components of the calibration (or standard) system can be configured

in a self-contained power-measuring system, thereby enabling verification of

the test equipment on a system basis as shown in Fig. 4.

The calibration system itself is readily verifiable and calibratable

against higher echelon standards. Most of the calibrations can be performed

in-house with minimal help from outside calibration laboratories, including

NBS. Finally, the entire measurement process, including the loss measurement

of the transformer, is traceable.

To facilitate the verification of the calibration instruments in-house,

several other instruments were acquired: an inductive divider, a standard

1000-pF low voltage capacitor, a high quality ac voltmeter, and a null

detector. The two-stage current transformer was designed to serve in several

other calibration roles besides the principal one of calibrating the current

channel in the test system.

The total instrumentation package for calibration of test systems is

relatively small, consisting of 13 instruments and impedance standards. All

are shown in Fig. 5.

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14
4. CALIBRATION OF CURRENT CHANNEL

4. 1 General Approach

The current transformers in both test systems have a single ratio, and

thus operate over a wide dynamic range. For example, the secondary currents

range from 10 mA to 10 A for the primary currents of 10 A to 10,000 A in one

of them. The nominal output voltages of the amplifiers are maintained within

the optimum range of 2 V to 10 V by changing the impedances in the networks

associated with the feedback amplifiers. The errors of the amplifier circuit

enter directly into the measurement error.

In designing the calibration equipment and procedures, the objective was

to keep them simple. Consequently, the physical standards for calibration of

the current channels in the test system consist of a two-stage current

transformer and a precision two-ohm resistor (shunt). The current transformer

has three principal ratios of 1000 A/1 A, 100 A/1 A, and 10 A/1 A, obtained by

changing the number of turns in the primary winding. By using a multi-ratio

current transformer, the dynamic range for the secondary current can be

maintained between 0.1 A and 1.0 A.

The current channels in both test systems are of the general type shown

in Fig. 1(b). The outputs of the current transformers are converted to

voltages by means of trans impedance circuits containing precision resistors

and feedback amplifiers. The trans impedances in the test systems are calibrated and adjusted against the standard current transformer and the

precision shunt by utilizing a circuit as shown in Fig. 6. For simplicity,

the current transformer in the test system at the left is shown as a simple

two-winding transformer. The standard two-stage current transformer is at the

right. Besides the precision resistor R, another resistor Rj having the same nominal value as R has to be employed, but R^ need not be a precision

component. The two-stage transformer and the resistors form a high-current

trans impedance (or shunt), the value of which is R/n, where n is the turns

ratio N2/N1

.

The two trans impedances are compared in a potentiometric circuit. A

seven-decade inductive voltage divider is connected to the output of the

amplifier in the test system. The ratio (d) of the inductive divider is set

to produce the desired trans impedance (^x/^x^ ^^ ^^^ test system when the

detector (D) indicates null.

(Rx/nx)d = R/n, (1)

d = (R/Rx)(nx/n), (2)

where n^^ is the turns ratio of the current transformer in the

test system, and R^ is the resistance associated with the amplifier

circuit in the test system.

In the actual alignment operation of (Rx/"x) » ^^^ effective

resistance and its phase angle are adjusted until the balance condition is

achieved. The phase angle errors of the standard resistor, the two-stage

current transformer, and the inductive divider are negligible.

15
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16
If the voltage at the output of the amplifier is smaller than that at the

standard current transformer, the inductive divider must be connected to the

terminals of the standard current transformer as in Fig. 7. This

configuration introduces an error, predominantly in the phase angle, since the

effective resistance R is modified by the input impedance of the inductive

divider in parallel with R. A correction must be made for this error. Alternatively, a buffer amplifier presenting an infinite input impedance to

the current transformer and zero output impedance to the detector can be

inserted between points A and B in Fig. 7. A simple buffer amplifier with

bipolar output and unity gain is shown in Fig. 8(a). The buffer amplifier can

be readily aligned for exact unity gain by using the null detector in the two

connections, as in Fig. 8(a), and injecting a small current from the circuit

in Fig. 8(b).

Since relatively small phase angle errors could lead to large errors in

the test system when measuring power, precautions must be exercised so that

the calibration process itself does not introduce significant errors. The errors in the calibration circuits most frequently arise from grounding

problems and from pick up of extraneous 60-Hz signals by the leads and the

equipment. Typically, the calibration is performed with 1 to 2 volts on R.

An extraneous signal of 100 to 200 microvolts could result in a phase angle

error of 100 to 200 microradians. It is prudent to observe the following

precautions:

1. Avoid multiple grounds and thus possible significant voltages between

these grounds. Test the voltage between the grounds if multiple

grounding exists. For example, connect the null detector across the

two grounds in the circuit of Fig. 6 when the circuit is energized.

2. Keep the leads short and wiring orderly; avoid high impedances in

critical wiring due to poor connections.

3. Keep the calibrating equipment and wiring away from high

electrostatic and magnetic fields. Keep the loops in the wiring

small so as to minimize the pickup from magnetic fields. Where

applicable, use shielded conductors, e.g., from the standard current

transformer and the resistors to the detector.

4. The instruments used as null detectors are frequently sources of pick

up. Since the calibrations are done at the power line frequency,

battery-energized instruments are preferable. Both inputs of the

detector in Figs. 6 and 7 are at a potential above ground; hence, a detector with a differential input can be used. Check that the common mode rejection is adequate by connecting both inputs to the same point, V or V^, when the bridge is energized. Detectors with

single-ended inputs, with one of the inputs being the chasis, will

generally function properly in this application since the

impedances on both sides of the detector are low, and the

chasis- to-ground capacitance will cause no problems. Check the

interference caused by external sources by observing the indication

of the detector without the circuit being energized.

17
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Figure 8. Bipolar Buffer Amplifier

18
5. Finally, it should be noted that the standard current transformer is

designed only for operation with both windings at or near the ground

potential. Since the current transformer in the test system is

normally inserted in the high voltage lead, it must be ensured that

the same primary current passes through both current transformers.

4 .2 Two-Stage Current Transformer — Theory

A two-stage current transformer is the principal component employed for

the calibration of the current channel in the test system. Such a transformer

was first described more than sixty years ago; in the more recent past it has

been of considerable interest to standards laboratories [10,11], Its

principles are briefly reviewed here by employing equivalent circuits.

For the two-winding transformer of Fig. 9(a), a conventional equivalent

circuit is shown in Fig. 9(b). It consists of an ideal transformer followed

by a T-network of impedances that are expressed with respect to the secondary

winding: Z^ is the magnetizing impedance, and Z\ and Z2 are the equivalent

primary and secondary leakage impedances, respectively. For power frequency

instrument transformers of the type described here, Z^ and Z2 are predominantly the winding resistances, and Zjjj is predominantly the opencircuit inductance combined with the equivalent conductance resulting from the core losses.

The actual no-load voltage/current ratio of the transformer can be

readily calculated from the equivalent circuit.

(Vi/Vz)^.^^^ = -(l2/Ii)^^^Q = (l/n)/(l+Z2/Z^). (3)

Since Z2 /Zjjj is small, typically <0.001, the equation (3) simplifies to

(Vl/V2)j^^Q = -(l2/Ii)^^^Q = (l/n)(l - Z2/Z„). (4)

Equation (4) suggests a reciprocal device for which the forward

short-circuit current ratio is equal to the reverse open-circuit voltage

ratio. Since Z^ is nonlinear, the reciprocity is not valid if very low

ratio uncertainties, of the order of ±0.0001, are desired. For a current

transformer that is operated with a burden, the impedance of the burden must

be added to Z2 in order to calculate the current ratio. Z^ does not affect

the current ratio, but influences the voltage ratio if the voltage transformer

has a burden.

A two-stage transformer is obtained by adding another core and another

winding to a simple transformer. The actual construction is illustrated in

Fig. 10(a). A compensation winding, N3 , having the same number of turns as N2, links only the toroidal core, CI, at its right. For best results, this core and the compensation winding are enclosed in a magnetic shield (MS) to

reduce the leakage flux that otherwise would not be confined in the toroidal

structure. Another toroidal core, C2 , is added and the primary and secondary

windings, N^ and N2 , are wound on the combination of the two toroids.

19
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If the magnetic shielding is perfect and N3 is magnetically coupled to N]^ and N2 exactly in proportion to the number of turns, the equivalent circuit of

Fig. 10(b) is a valid representation of the device of Fig. 10(a). Zjjj^ is

the magnetizing impedance of the core CI, Zm2 of C2, and Z3 is the leakage

impedance of N3 . After performing a "ir-T" transformation and neglecting

insignificant terms, the equivalent circuit of Fig. 10(c) is obtained. In

this conventional T-network, the secondary leakage impedance Z2 is effectively

transfered to the primary side where it does not affect the accuracy of an unloaded transformer. The residual secondary leakage impedance, ZiZ-i^lZ^i^

is reduced, typically, by a factor of 1000 to 10,000 from the original value

and frequently, as in the transformer built in this investigation, can be

neglected.

The expression for no-load voltage/current ratios, analogous to eq. (4),

is given by

(Vl/V2)j^^Q = -(l2/Ii)^^^o = (l/n)[l-(Z2Z3)/Z^lZ^2)]. (5)

For the transformer that was constructed, Z2 = 2 ohms, Z3 = 5 ohms, Z^^ =

2ni2 ~ j50,000 ohms, yielding the error term of 4 x 10" . While this term

will vary with the flux density in the cores, it is completely negligible in

the applications for which the transformer was constructed. The actual

no-load voltage and current ratios are equal to the turns ratio, and the

transformer is reciprocal. The equivalent circuits of Fig. 10 have considered

only the errors that are of magnetic origin. Internal capacitances will

contribute to additional ratio and phase angle deviations, but these

deviations do not exceed one or two parts per million in the transformer that was constructed.

Equation (5) applies only to the no-load condition. The advantages of a two-stage current transformer will be defeated by connecting a common burden

to both windings, N2 and N3 . On the other hand, the advantages will be

retained by connecting separate, nominally equal, impedances to the two

windings as in Fig. 6. Only one of the burden impedances needs to be known to

great accuracy [12],

Consider the circuit with two burden impedances as shown in Fig. 11. The

combined output current is given by

-(I2+I3) = (Ii/n)[l-(Z2+R)(Z3+Ri)/Zin2Zmll- (6)

Even with the added burden resistances the error term remains negligible in

equation (6). The current I2 can be calculated from the considerations of a simple two-winding transformer

-I2 = (Ii/n)[l-(Z2+R)/Z^]. (7)

Equation ( 7) is based on the core at the left and the windings N^ and N2 being

considered the constituents of a transformer. With respect to the core at the

right, the effective primary current is the error component given in equation

( 7) , Hence the current in N3 becomes

-I3 = (Ii/n)(Z2+R)/Zn^[l-(Z3+Ri)/Znji]. (8)

21
Neglecting the product of error terms,

-I3 = (Ii/n)(Z2+R)/Zin2. (9)

The above discussion considers the two-stage current transformer as a combination of two current transformers: (a) the principal current transformer

and (b) the compensation current transformer. The latter compensates for the error of the principal current transformer.

I2 and I3 , when passed through the two resistors R and R^ , respectively,

produce the required voltage V as in Figs. 6 and 7. Thus,

V = -I2R-I3R1. (10)

Let

Rl = R(l+a), (11)

where ot is a small deviation.

Substituting (7), (9), and (11) in (10) and neglecting the products of the error terms yields

V = (Ii/n)R,

V/Ii = R/n. (12)

Thus, the effective trans impedance depends only on the value of the precision

resistor connected to N2 and on the turns ratio of the two-stage transformer.

4.3 Two-Stage Current Transformer — Construction

The components comprising the entire transformer are shown in Fig. 12,

The magnetic cores and principal winding are similar to a range extending

transformer that was described previously, except only one magnetic shield is

used [13]. Starting at the left is the first toroidal core on which a 1000-turn compenstion winding is wound. It is followed by electrostatic and

magnetic shields and a 100-turn compensation winding. Then the second

core, identical to the first one, is added. The primary and secondary

windings are placed on the combined toroid. Additional windings and taps,

beyond those required for the 1000/1, 100/1, and 10/1 ratios, are provided in

order to enhance the flexibility of the transformer for performing

self-checking operations, and to facilitate the use of the transfomer in

impedance bridge circuits. To minimize the winding- to-winding capacitances,

electrostatic shields are provided between secondary and primary windings.

The assembled unit is mounted in a metal-lined phenolic enclosure with readily

accessible terminals as shown in Fig. 13.

A decision to add this winding was made after the shields were constructed. A preferable place would have been below the two shields. This

winding is not critical to the principal operation of the transformer.

22
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24
4.4 Functional Tests of Current Transformer

Properly constructed two-stage transformers are rugged and highly

stable devices. Most of the failure modes, such as open and short circuits,

tend to be catastrophic and are readily apparent. There are no long-term

drifts. Subtle failures causing significant ratio errors are possible, e.g.,

as a result of partial turn-to-turn short circuits in a winding. As a result

of the transformer action, the effects of such short circuits are coupled in

the other windings of the transformer. Thus the integrity of the transformer

can be verified by checking one or two selected ratios without reverting to

complete recalibration. Also, if convenient, the checking can be done in the

voltage transformer mode.

A circuit for checking the transformer against a standard inductive

divider is shown in Fig. 14. An inductive divider can be utilized to check

the ratios of all taps of the secondary winding and the ratios of the

secondary-to-primary windings. The in-phase component is balanced by

adjusting the setting of the inductive divider. The residual phase angle

difference is determined from the reading of the null detector. In this case a tuned microvoltmeter used as a null detector yields information on the

magnitude of the phase angle only. Hence, a phase-sensitive instrument

providing information on the polarity of the unbalanced signal is preferable.

Using a seven-decade inductive divider, this method is entirely satisfactory

at ratios as low as 0.1 and even 0.01. At a ratio of 0.001, the resolution

and the sensitivity may be insufficient.

A self-checking method is outlined in Fig. 15. The 0.1 tap of the

1000-tum winding can be compared to each of the two 100- turn windings. The

operation can be repeated by comparing the output between any two adjacent

taps with that of a 100-tum winding. An inductive divider can be added for

comparing the output of the 10- turn winding to that of the 100- turn. The

circuits of Fig. 15(a) and (b) can be operated either in the voltage

transformer or the current transformer modes. Again, a phase-sensitive

detector is preferable for determining the in-phase and quadrature components

of the error. The tests shown in Fig, 15, while not being self-contained

calibrations, are excellent means for checking the integrity of the current

transformer after a comprehensive calibration has been performed.

25
IVD CT IVD

e r-®—

(a) Taps in the secondary winding (b) Ratio of two windings

Figure 14. Checking of Two-Stage Transformer Against Inductive Divider

t)0

X

100 T ©0 100 T

(a) 0.1 tap against 100-T winding (b) Adjacent taps against 100-T winding

e

100 T 10 T IVD

(c) 10-T winding against 100-1 winding

Figure 15. Self-Checks of Two'^Stage Transformer

26
4.5 Precision Resistors

The two resistors at the output of the standard current transformer are coaxial shunts of the type shown in Figs. 16 and 17. These shunts were

obtained from a commercial source. The shunt consists of a central conductor

surrounded by a concentric resistance element of a low-temperature coefficient

material. This element is in turn surrounded by a concentric outer

conductor. The current is Introduced at the male conductor and passes down

the inner rod to the junction with the resistor. It then folds back down the

resistor and exits via the outer shell of the male connector. The voltage is

sampled at the female connector, the center of which is connected through the

outer tube to the common end of the resistor. Due to this use of concentric

coaxial conductors, the voltage lead is coupled to the resistor in a path

which is essentially free of the magnetic field of the current in the

resistor. The spacing between the resistor element and the central conductor

is kept extremely small to reduce the inductance. The space between the

resistor element and the outer shell is filled with a thermal transfer

compound, and the outer shell is constructed with an Integral heat sink to

assist in the control of self heating effects.

These particular units are two-ohm shunts, adjusted to a tolerance of

±400 ppm over a current range of to 2 amperes. Calculation of the reactance

of the shunt Indicates that the phase shift of the output voltage vs. the

input current at 60 Hz should be less than 10 microradlans. Calibration

confirms that this goal was met.

4.6 Calibration Data

The three principal ratios, 10 A/1 A, 100 A/1 A, 1000 A/1 A, of the

standard current transformer were calibrated in the no-load (short circuit)

condition and with the tubular two-ohm shunts connected in the secondary and

compensation windings. The calibration results are summarized in Table 1; the

actual NBS calibration reports are reproduced in Appendix 1. In all cases,

the measured ratio and phase angle deviations are within 3 ppm and 4 yrad,

respectively. The estimated calibration uncertainties are ±10 ppm and ±10

yrad. Thus the indicated deviations are well within the estimated measurement

uncertainties. The equipment that was used to calibrate the standard current

transformer is designed for operation primarily in the 0,5 A to 5.0 A range of

the secondary current. Hence, the 1.0 A and 0.5 A data in the report are believed to be more reliable than those at lower currents.

The transimpedance of the current transformer-shunt combination was measured for the 100 A/1 A and 1000 A/1 A ratios of the current transformer.

The results are tabulated in Table 2. The resistance of the shunt can be

calculated from the measured transimpedance and the transformer ratio by using

eq. (12) In a slightly modified form,

R - (V/Ii)ni, (13)

where ni is the measured complex ratio of the transformer. As the data in

Table 2 indicate, the computed phase angle of the shunt is nearly Identical

and negligible for both sets of measured data. The resistance of the shunt is

Identical within the resolution of the equipment. It should be noted that the

calibration equipment for transimpedance is more accurate in phase angle than

that for the transformer ratio [12].

27
Cross Section A-A

Inner Conductor

Resistance Material

Outer Conductor

Simplified Cut-away View

Figure 16. Coaxial Shunts, Cross S<?ctlon

Figure 17. Coaxial Shunts, External View

28
During the construction of the transformer, the linearity of taps in the

secondary winding was measured using a special calibration arrangement. The

results are tabulated in Table 3. The corrections are expressed as fractions

of the input quantity. To obtain the corrections to the ratio or the output

quantity, divide the tabulated corrections by the tap setting. It is noted

that all magnitude corrections in ratio are smaller than 1 ppm, except for the

0.1 tap; all phase angle corrections in ratio are smaller than 1 urad. For

the 0.1 tap the phase angle correction is 1.29 yrad.

Table 1

Calibration of Standard Current Transformer

Nominal

ratio

Burden

resistance

(ohms)

Current ratio

Phase

angle

(mlcroradians)

10/1

100/1

1000/1

2

2

2

10 X 1.000000

1.000000

100 X 0.999997

0.999997

1000 X 1.000000

1.000000

-1-4

-3

+1

-2

Secondary Current lA

Estimated uncertainty: ± 10 x lO"** In ratio

± 10 wrad In phase angle

29
Table 2

Calibration of Trans impedance

Standard Current Transformer-Shunt Combination

Nominal

Transformer

Ratio

Measured trans impedance Computed impedance of shunt

Magnitude

( ohms

)

Phase angle

(microradians)

Magnitude

( ohms

)

Phase angle

(microradians)

100/1

1000/1

0.01999302

0.00199930

-2.6

-1.8

1.999296

1.99930

0.4

0.2

Frequency: 60 Hz

The data in Table 1 is used in computations.

Phase angle is positive if the impedance is inductive.

Estimated uncertainty: ±10" in magnitude of trans impedance

±5 urad

in phase angle of trans impedance

Table 3

Calibration of Standard Current Transformer

Taps in Secondary Winding

Tap Magnitude Quadrature

setting (A) correction (a) correction (b)

0.1 0.06 X

10-^ 0.12 X lO-'' 0.2 0.06 0.11 0.3 0.06 0.01

0.4 0.06 -0.14

0.5 0.06 -0.30

0.6 0.06 -0.44

0.7 0.06 -0.51

0.8 0.05 -0.49

0.9 0.04 -0.34

0.1* 0.00 0.08

-I2/I1 - V1/V2 = A + a + jb

Tap 0.1* is the auxiliary secondary winding.

Estimated uncertainty: ±0.05 x 10" in a ±0.10 X 10"'' in b

30
5. CALIBRATION OF VOLTAGE CHANNEL

5.1 General Approach

The voltage channel of the two test systems is of the type shown In Fig.

1(b). The voltage for the power transducer is generated by means of an active

divider incorporating a gas-dielectric, high voltage capacitor and operational

amplifier circuits. To obtain the dynamic range over three decades — voltage

from 100 V to 100 kV — additional amplifiers and passive networks are utilized to control gain.

For calibration of voltage channels, a standard active divider is

employed in the circuit configuration shown in Fig. 18. If the outputs of

both active dividers are of the same polarity, an inductive divider and a null

detector are inserted between the two dividers as shown in Fig. 18(a). Of

interest is the factor n^ by which the primary voltage is reduced in the

test system.

Vi/n^ = -dVi(Ci/C2),

tiv = -(l/d)(C2/Ci), (14)

where d = inductive voltage divider setting.

Occasionally the amplifiers that follow the integrating stage may be of

the inverting type as shown in Fig. 18(b). In such circumstances, either an inverting type buffer amplifier can be inserted between the standard divider

and the inductive divider, or the inductive divider can be connected between

the two active dividers as in Fig. 18(b). The balance equation becomes

nv = (C2/Ci)(l-d)/d. (15)

As in the case when calibrating the current channel, precautions must be

taken in order to avoid errors that could be introduced during the calibration

process. Most common measurement problems are caused by improper grounding of

various instruments and by electromagnetic interference which could produce erroneous signals in the measurement circuit. Specific precautions mentioned

in section 4.1 should be observed when calibrating the voltage channels. An

important precaution in using standard high voltage capacitors is to have the

low voltage electrode and the conductors leading to it completely shielded.

31
Vi

(Vi/nv) IVD

<2>

C2

(a) Same polarity

Vi

(Vi/nv)

:ci

^^>y^wJ-^^>J

cx:ti>

^H

(b) Opposite polarity

Figure 18. Calibration of Voltage Channel

32
5.2 Standard Voltage Divider

In developing the standard voltage divider, the goal was to produce an instrument whose voltage ratio depends only on the ratio of the two

capacitances, C^ , and C2 , in Fig. 18. The capacitors can be designed to have

negligible phase defect, and in-house means can be provided to check the

capacitors. The schematic operational amplifier circuit shown in Fig. 18 has

imperfections resulting from the finite gain of the amplifier and compensating

networks that have to be added to the amplifier to eliminate dynamic

instabilities and dc drift. The compensating networks for dynamic stability

of high performance amplifiers can become especially complicated [13].

A paper describing a very accurate active divider — both for magnitude

and phase angle — that incorporates two gas dielectric capacitors and a magnetic current comparator was recently published [14]. If only the phase

angle is of primary importance, a relatively simple solution is possible. A

feedback amplifier with low gain and employing a solid-dielectric capacitor in

the feedback path is combined with a manually adjusted source that is

phase-locked with the high voltage source. The operation of such an instrument is illustrated in Fig. 19. In Fig. 19(a) a capacitive divider is

shown having two voltage sources — Vj is an independent source, while V2 a source that is phase-locked to V^ and is manually adjusted in magnitude and

phase. The voltage at the junction of the two capacitors is monitored with a null detector which in turn is preceded by a preamplifier. At the null

condition, the same current passes in both capacitors and the voltage ratio

becomes

(V1/V2) = -(C2/C1). (16)

The need for adjusting V2 within a few ppm can be avoided by including

the preamplifier in the feedback loop as shown in Fig. 19(b). The combined

voltage of the controlled source (V2) and that of the amplifier (V3) are now applied to C2 . The voltages in this circuit are related as follows:

(Vi-V4)Ci = -(V2+V3-V^)C2.

The ratio of voltages on the two capacitors become,

Vi/(V2+V3) = -(C2/Cl)[l-V^/(V2+V3)-(Cl/C2)V4/(V2+V3)]. (17)

The voltages V3 and V^ are related through the gain of the amplifier,

which, in this case, is -100. If V3«V2) the error terms in equation (17)

become negligible and the voltage ratio becomes the capacitance ratio as in

equation (16). For example, if V3 = O.OOOIV2, the error is 1 x 10"^.

In practice, there are several possibilities for implementing the

controlled source: from a tap of the supply transformer, from another high

voltage divider, or from the guard capacitance of the high voltage capacitor.

A very stable controlled source can be produced by using a high voltage

standard capacitor that has two or more shielded capacitances in the same

structure. Commercial units of this type are available. Operational

amplifier circuits are employed, as in Fig. 20, to implement the controlled

source.

33
Vi

I

I

(a) Open- loop configuration

Vi

e

V4

C2

-100

\>

-^

V2

(b) Closed-loop configuration

Figure 19. Active Divider with Controlled Source

Vi

V4 10^-

R

-AAAr

R -V2

-VSAr

N-V3

I—v\Ar-

^[

V2

C2

V2+V3

Figure 20. Implementation of Controlled Source in Active Divider

34
The advantages of the circuit in Fig, 20 are: (1) the voltage at the

summing point of the principal amplifier is monitored and, if necessary, can

be adjusted to a negligible value; (2) the voltage ratio is the reciprocal of

the capacitance ratio, hence, there is no need to calibrate the active divider

as a system, only the two capacitors need to be calibrated, primarily for the

phase angle; (3) the loop gain of the feedback circuit is 100, hence, there

are no problems with dynamic stability; (4) the amplifier comprising the

controlled source is not in the feedback, loop and does not contribute to

dynamic instabilities.

5.3 Standard Capacitor Networks

The ratio of the divider critically depends on the ratio of the two

capacitances, C^ and C2 • High accuracy of about ±5 pradians is required for

the phase angle, but only a moderate accuracy for the magnitude of about

±0,05%. The high voltage capacitor (100 pF) is a guarded, gas-filled unit.

Such capacitors have negligible phase defect and capacitance dependence of

temperature that is typically 20 ppm/°C. For this purpose, such a capacitor

can be considered as ideal.

To obtain the required output voltages of the active divider, three

capacitors, 1.0 mF, 0.1 yF, and 0.01 yF, are required for the low voltage arm. Units having polystyrene film as the dielectric were selected. Their

temperature coefficient of capacitance is -0.01%/°C, and the dissipation

factor (tangent of the phase angle) ranges from 50 ppm to 100 ppm. No special

precautions, except for a moderately accurate calibration, are required for

maintaining the capacitance value. The dissipation factor must be measured

very accurately and preferably should be compensated. Compensation is

feasible, since the three-terminal capacitance rather than the two-terminal

capacitance is of importance.

A compensated low voltage capacitor is shown in Fig. 21(a) [15]. C is

the value of the actual capacitor, and G is its equivalent loss conductance.

The T-network containing the two capacitors and the resistor is the

compensating network. The approximate equivalent T-network of the

compensating network is shown in Fig. 21(b). The effective negative

conductance, G^ , is adjusted to cancel G. The result is that the effective

transadmittance becomes equal to the admittance of a pure capacitor,

-(I2/Vi)v2=0 = >C. (18)

The adjustment of the capacitive networks is best accomplished during the

calibration of these networks. Since the dissipation factor of the capacitors

themselves is very small, typically of the order of 60 ppm, the compensation

need not be very precise. Reduction of the effective dissipation factor to

within 5 ppm is readily achieved. Preliminary indications are that the

compensated dissipation factor remains stable within ±5 ppm. The compensation

is for one frequency only. This causes no problem if the calibrations and

tests are performed at one power frequency. If the test frequency is changed,

e.g., from 60 Hz to 50 Hz, a different setting of the resistor is required.

35
G

-VSAr- 1

I

I

Vi

^e

Ci H^HH

I2

Ci =C/100

1 R =

(a)C

(a) Circuit diagram

r— — ---^^^- — - — -^

Gi=-U)2Ci2R

(b) Equivalent circuit

Figure 21. Low Voltage Capacitor Network

36
The three capacitor networks are enclosed in metal shielding boxes as shown in Fig. 22(a). Two pairs of coaxial terminals are provided — one for

the current, the other for the voltage (the current terminals are in back and

not visible in Fig. 22(a)). The enclosure containing the electronics for the

active divider is shown in Fig. 22(b). The high voltage capacitor is shown in

Fig. 4.

5.4 In-House Calibration of Standards

Means for checking the capacitors on site is highly desirable. This

applies to both the high-voltage capacitor and the special capacitive

networks. Such a capability is established by acquiring one standard 1000-pF

capacitor and an impedance bridge. Two impedance bridges are available — one

that is based on the inductive divider, the other utilizing the two-stage

current transformer.

The inductive divider bridge is shown in Fig. 23(a). It is always

operated with the nominal ratio of 1/11. Thus the following comparisons are

made: 100-pF (high voltage) vs. 1000-pF (standard); 1000-pF vs. 0.01-yF;

0,01-yF vs. 0. 1-yF ;0. 1-yF vs. 1-yF. The bridge is balanced by adjusting the

ratio of the inductive divider. The residual indication on the null detector

is the unbalanced phase angle. It should be negligible for the two

gas-dielectric capacitors and for the capacitive networks after the loss

component has been compensated. When the bridge is balanced,

(C2/C1) = (l-d)/d, (19)

where d is inductive voltage divider setting.

Several precautions are in order. First, the leads from the capacitors to the

detector must be completely shielded, especially when calibrating capacitors

of low value, e.g, 100 pF to 0.01 UF. Excessive capacitive currents, when

passed in the tap of the inductive divider, could cause loading errors. A

test can be performed to determine whether such errors are significant by

inserting a 10-ohm resistor between the bridge ground and the tap of the

inductive divider. The balance should not change. If it does, the capacitive

currents can be balanced with the "Wagner ground," consisting of the resistive

divider and the capacitor connected to its tap. Second, when measuring large

capacitances, 0.1 yF and 1.0 yF, the leads should be short so that the lead

impedances do not add significantly to the impedances of the capacitors and

the inductive divider. Fig. 23(b) is the suggested circuit for comparing the

two largest capacitors; it incorporates the features of a Kelvin bridge.

Greater accuracy can be obtained by using the standard current

transformer in an impedance bridge configuration as shown in Fig. 24. The

compensation winding now consists of a series connection of the 1000-tum and

100-turn windings. This combined winding is connected to the source and to

the current terminals of the series connection of the two capacitors. The

1000-turn secondary winding and the 100-turn primary winding are also

connected in series and then to the potential terminals of the two

capacitors. The inductive divider, connected to the 100-turn auxiliary

secondary winding, and C3 , provides the magnitude adjustment. At balance,

(C2/C1) = (N2/Ni)(l-K).01d), (20)

where d = inductive voltage divider setting.

37
m..ri'|'iTi'|'HnTirjT'rg'''i'n'''i'a'''i''T''i'^r''''''i'''i'!;i'''''''i''''8l''^ «...».« WESTCOTT* Fle«ible Slainlesi- Steel Ruler

^ ^ ^ -«ti«as..j

(a) Low voltage capacitor networks

(b) Enclosure containing electronics

Figure 22. Components of Active Divider

38
IVD

Nt^^

10Q I-±r

v._ •

(a) Comparison of small capacitors

JJ

IVD

^

n' -±r O.2Q

1 mF

(b) Comparison of large capacitors

Figure 23. Capacitance Bridge Based on Inductive Voltage Divider

C3=0.1 Ci

N3 = 0.1 N2

Figure 24. Capacitance Bridge Based on Two-Stage Current Transformer

39
In general, the inductive divider bridge is easier to use but is not as accurate as the two-stage transformer bridge because of inherent uncertainties

of the divider. The inductive divider bridge is an excellent instrument for

monitoring the performance of the capacitors after the initial high-accuracy

calibration has been performed by other means. The three capacitor networks,

before calibration at NBS, were adjusted with the inductive divider bridge.

Good results were achieved with residual deviations in the dissipation factor

of 6 X 10~ , -13 X 10" , and 0. The inductive divider bridge itself can

be calibrated at fixed points, including the 11/1 ratio, by using the

transformer bridge as in Fig. 14(a).

The current transformer bridge can yield phase angle uncertainties that

are within 3 yrad in each measurement. Thus, even with three measurements

involved in the buildup process, the overall uncertainty can be maintained

within ± 9 yrad.

40
6. CALIBRATION OF POWER TRANSDUCERS

6. 1 Standard Transducer

The power transducers in the two test systems are of the time division

multiplier type that multiply and average two voltages (or a voltage and

current) [16]. For their calibration, a special higher accuracy time division

multiplier was purchased commercially. This standard transducer was calibrated at NBS against a dual-channel signal source which serves as a phase

angle standard [17]. The calibration results given in Appendix 1 indicate

that the corrections of the transducer are within one percent (except for one point) of the full-scale reading for the three ranges of the power factor —

1.0, 0.1, and 0.01. The manufacturer's specified uncertainty limit is also

± 1% for the three ranges of the power factor. The calibration at NBS had an uncertainty of ±0.2% for the magnitude and ±50 yrad for the phase angle. To

realize the highest accuracy in the calibration of the transducers in the test

systems, the corrections for the errors of the standard transducer had to be

applied. This approach is adequate; however, a standard with negligibly small

corrections that could be resolved into magnitude and phase angle components

would be more desirable.

6.2 Standard Phase Shifter

A very promising type of calibrator for transducers operating at small

power factor is an accurate 90 degree phase shifter. Such a device was developed and constructed in the later stage of the project, but it was not

available during the alignment of the test systems.

An accurate yet simple 90 degree (or near 90 degree) phase shifter is

obtained by converting the active divider described in section 5.2 and shown

in Fig. 19 into an integrator shown in Fig. 25. The input element is a 26.5

kfl miniature metal film resistor. The feedback element is the 0,1-yF special

capacitance network used with the active divider. The estimated phase angle

error limit for the resistor is ±5 yrad, assuming a stray capacitance of 0.5

pF due to the resistor itself and the electrostatic shielding enclosure. All

the other phase angle errors are as in the active divider. The combined

effect is estimated not to exceed ±20 yrad. The controlled source shown in

Fig. 25 is implemented by a circuit that is analogous to that of Fig. 20.

Calibration of a transducer is outlined in Fig. 25(b). The input and

output voltages of the phase shifter are applied to the transducers, and the

magnitudes of voltages are measured with precision digital voltmeters. To

provide nominal power factors of 0.01, 0.02, 0.05, and 0.1, C2 can be shunted

with the four resistors as shown. Output of the transducer is

P = kViV? cos (21)

where = tan (wRC),

k has the dimension of Siemens

Lagging power factor is obtained by interchanging the connections at the

transducer. It is to be noted that for the feedback resistors it is important

to know accurately their resistance values, not their phase angles.

41
e i"i

hr"

1 jr, F^^-D^^

b^^l

V2

C2

(a) For zero power factor

Vi

e

26.5 k

D>

<^

0.1 pF

^f

o——A/\/sr

o—© WAr

o ^AAr

o ^^r

POWER

TRANSDUCER

V2

2.65 MQ

1.32 MQ

529 kO

264 kQ

PRECISION

DVM

(b) For power factors of 0.00, 0.01, 0.02, 0.05,

and 0.1; shielding omitted

Figure 25. Active Phase Shifter

42
7. ALIGNMENT OF TEST SYSTEMS

7.1 Test System 1

Figure 26 shows a schematic representation of the circuit used in the

alignment of the current channel. The input signal to the buffer amplifier is

nominally 5 volts, corresponding to the primary currents of 10 A, 100 A, 1000

A, and 10,000 A. Nominal voltage is achieved by changing the burdens in the

secondary windings of the zero-flux current transformer. The buffer amplifier

has a fixed gain of two. Small in-phase and quadrature currents can be

injected in the feedback loop of the amplifier in order to produce small

adjustments in gain. The scaling amplifier reduces the signal to

approximately one volt.

The gain of the buffer amplifier is adjusted for the lOO-ampere primary

current. For other currents, the gain of the current channel is aligned by

small ajustments in the burdens of the current transformer. The general

procedure outlined in section 4.1 is followed for the alignment of the current

channel.

It is to be noted that the gain of the current channel is not trimmed for

all the settings of the scaling amplifier. However, as the gain of this

amplifier is changed, the resulting phase angles remain relatively stable

within ± 20 yrad; the changes in the magnitude are larger, approaching ±0.1%.

Still larger changes are observed due to nonlinearities of the zero flux

current transformer. The adjustments are made to produce the smallest overall

error in a particular range of the primary current. After the adjustment,

the residual phase angle and magnitude errors were determined at the full

scale points and 50% of the full scale points. The results are given in

Table 4. These residual errors are included in the assessment of the

uncertainty of the test system, as discussed in section 8.

Adjustment of the voltage channel is a two-step procedure. Initially the

buffer and ranging amplifiers are adjusted in the circuit of Fig. 27. The

final adjustment, for phase angle only, is accomplished by using the circuit

of Fig. 28. Adjustments are made for three full-scale voltage ranges —

1 kV, 10 kV, and 100 kV. Thus, three separate trimming circuits are provided

for the ranging amplifier. The general procedure for alignment described in

Section 5.1 is followed.

As for the current channel, different gain settings of the scaling

amplifier, residual pickup and nonlinearities in the test system prevented

alignment of the voltage channel over the entire operating range. The

magnitudes and phase angles were adjusted to yield the lowest practical error

in a particular range. The residual phase angle and magnitude errors are tabulated in Table 5.

43
> ^^

10 A

1000 A

10,000 A

J—WSr-L»

ZERO FLUX CT

COMPENSATION BURDENS

MAIN BURDENS

STANDARD TWO STAGE

CT

BUFFER

AMPLIFIER

NULL DETECTOR ¥S>r

COO£

INDUCTIVE PRECISION VOLTAGE SHUNTS

DIVIDER BUFFER AMPLIFIER SCALING AMPLIFIER

Figure 26. Current Channel Alignment, Test System 1

44
I

iroo<

a.<o

0.<Zoz<

cr o O 0. m <

til CO c <

Q.

O s

0)

COH

0)

COOcc

00

60

03O>

CNl

0)

i-i a

60

•H

Pi4

45
4-1

CO

4JU

(U H

P! •H

4-) C!I

00

(UbO

cd

4-> iHO>

00

CM

(U

l-i 3

bO •H

46
The third critical subsystem is the power transducer and its associated

circuitry. In order to obtain high accuracy in power measurements at low

power factors, great care must be exercised in selecting this subsystem.

Special transducers for low power factor applications have to be utilized.

The original time-division-multiplier-type transducers were replaced with

improved units of the similar type that were fabricated at ASEA Electric.

Transducers are most accurate near their full-scale values. Hence,

scaling amplifiers are employed that maintain the two input signals between

0.7 of full-scale and full-scale, yielding an output signal that is no lower

than 0.5 of full-scale. To maintain a direct reading feature for the entire

power measuring system, scaling amplifiers are used also at the output of the

transducer, but the gains of these amplifiers are reciprocals of the gains of

the input amplifiers.

The transducers in the test system are aligned using the circuit of

Fig. 29. Dual-channel, variable-phase source supplies two signals to the

transducer under test and to the standard transducer that was calibrated at

NBS. Adjustments are made in the system transducer to minimize its deviation

from the correct value. Because of residual nonlinearities, perfect alignment

could not be made over the entire measurement range. The largest residual

errors for three power factors are tabulated in Table 6.

7.2 Test System 2

The general principles for aligning this test system are similar to those

of system 1. Hence, the circuit diagrams are not repeated.

The current channel in this test system contains a two-stage current

transformer with a fixed turns ratio of 10,000/1. The secondary current is

passed to an operational amplifier that acts as a current-to-voltage

converter. Four ranges of conversion are provided, corresponding to

full-scale primary currents of 1 A, 10 A, 100 A, and 1000 A. The buffer

amplifier was modified from the original design to provide noninteractive

magnitude and phase angle adjustment for each of the four ranges.

As for test system 1, adjustments were made to minimize the error in the

entire range. The residual errors are tabulated in Table 7.

The voltage channel is aligned by a procedure that is similar to that

utilized with test system 1. The residual errors are tabulated in

Table 8.

The power transducer circuit was modified by the addition of scaling

amplifiers. The calibration and alignment are similar to those employed with

test system 1. The results are given in Table 9.

47
SCALING

AMPLIFIERS

(I) VARIABLE

PHASE

60 Hz

SOURCE

(V)

TEST

SYSTEM

POWER

TRANSDUCER

SCALING SUMMING

AMPLIFIER AMPLIFIER

(I) (V)

h-t>—D>-^n

DISPLAY

STANDARD

POWER

TRANSDUCER

Figure 29. Power Transducer Alignment, Test System 1

48
Table 4

Current Channel Alignment

Test System 1

Primary

current

(A)

Phase angle error

(urad, absolute value)

5

1050

100 500

A* B* C*

3870

3796

28

5552

6 75

13

3469

6 92

34

Magnitude error, absolute value

average: 0.13%

maximum: 0.34%

Table 5

Voltage Channel Alignment

Test System 1

Primary

voltage

(kv)

Phase angle error

(yrad, absolute value)

0.5

1.0 5.0

10.0 40.0

A* M c*

1325

440

75

25

662080

106 100

62640

Magnitude error, absolute value

average: 0.10%

maximum: 0.42Z

Table 6

Power Transducer Alignment

Test System 1

Power

factor

(lagging)

Error In power

(percent of reading)

1.00 0.10

0.01

A* B* C* 3*

0.00

+0.11

+0.64

+0.03

+O.02

-0.025

0.00

40.01

+0.04

+0.01

+0.05

+0.22

49
Table 7

Current Channel Alignment

Test System 2

Primary Phase angle error current (lirad, absolute value)

(A)

0.5

A<t) B* C*

10 10 20

1 15 30 15 5 5 3 2 10 3 4 3 50 10 4 16 100 10 5 10 500 8 5 6

Magnitude error, absolute value

average: 0.16Z

maximum: 0.27%

Table 8

Voltage Channel Alignment

Test System 2

Primary

voltage

(kV)

Phase angle error

(wrad, absolute value)

0.5

1.0 5.0

10.0 50.0

A* H C*

3

34

2 20

5

3

4

3 20

4

34

3 20

Magnitude error, absolute value

average: 0.10%

maximum: 0.20%

Table 9

Power Transducer Alignment

Test System 2

Power

factor

(lagging)

Error in power

(percent of reading)

1.00 0.10

0.01

A* B4> C* 3*

-0.04

+0.50

+0.81

+0.22

+0.73

+0.62

. . .

+0.19

+0.71

+0.03

+0.12

+0.65

+0.49

50
8. ANALYSIS OF MEASUREMENT ERRORS

8. 1 General Comments

Errors or uncertainties in transformer loss measurements are contributed

by numerous sources. Several authors in the past have discussed the nature of

these errors [18,19,20]. General guidelines for expressions of errors are well established; a comprehensive summary is available in [21].

The propagation of errors in the measurement of power is apparent from

the general equation for power under sinusoidal conditions

P = VI cos 0. (22)

Taking the partial derivatives of (22) yields,

3P = (I cos 0) 8V + (Vcos 0)3I-(VI sin 0)30. (23)

Dividing (23) by (22),

3P/P = 3V/V+9l/I-tan 30. (24)

Equation (24) indicates how small changes in voltage, current, and phase

angle combine to produce changes in power. While relative changes in the

magnitude of V and I enter directly as changes in power in an additive manner,

the change in phase angle in radians is modified by the tangent of the phase

angle. For phase angles approaching it/ 2 rad (small power factors, <0.1),

tan 0**± 1/cos 0, and therefore equation (24) becomes,

3P/P = 3V/V + 3I/I+ 30/COS0. (25)

When dealing with errors, it is customary to substitute incremental

changes in (24) and (25) for infinitesimal changes. Also a (± ) sign is used

since the errors can be positive or negative. Thus, the equation for

propagation of errors in power measurement for small power factors becomes,

AP/P = ±AV/V ± AI/I+ A0/COS0. (26)

The first two terms on the right of equation (26) are known as the

magnitude errors; the third term is the phase angle error. Each critical

subcomponent in the measurement system must be assessed for the two types of errors. In the case of voltage and current scaling devices, all three

components are separately identifiable. In the case of the power transducer,

the two magnitude components are inherently combined, if the terminal

characteristics of the instrument are considered.

From the discussion in the previous paragraphs, the error in power

measurement at low power factors can be formalized as follows,

Pt = Pi(l ± a ± b/cos 0), (27)

51
where

P^- is the true power,

Pj^ is the indicated power,

a is the fractional magnitude error, and

b is the phase angle error in radians.

When dealing with error components, it is frequently awkward to use

simple fractions. It is much more convenient to express such values in terms

of percent, parts per thousand, or parts per million. It is important that

similar units of value are used for b as for a, e.g., centiradians,

milliradians, microradians.

Both error components, a and b, have contributions from numerous

sources. The methods for combining the contributions are reviewed in the text

that follows. The uncertainties in a measurement process arise from

lack of precision (random errors) and possible offsets or biases (systematic

errors). The random errors can be determined rigorously from the measurement

results; they are amenable to the treatment by statistical methods and are described by such measures as standard error, probable error, and confidence

interval. The effect of random errors can be reduced by increasing the number

of replicate measurements and taking the mean of the increased set of

measurement data.

The assessment of systematic errors is a more difficult and more subjective process. Quoting from reference 21:

"Although a general guideline for the approach to the assessment of

systematic uncertainties can be formulated, there are, unfortunately,

no rules to objectively assign a magnitude to them. For the most

part, it is a subjective process. Their magnitudes should preferably

be based on experimental verification, but may have to rely on the

judgment and experience of the metrologist. In general, each

systematic uncertainty contribution is considered as a quasi-absolute

upper bound, overall or maximum limit on its inaccuracy. Its magnitude

is typically estimated in terms of an interval from plus to minus ^ about the mean of the measurement result. By what method then should

the magnitude of these maximum limits be assigned? It may be based on comparison to a standard, on experiments designed for the purpose [4]

,

or on verification with two or more independent and reliable measurement

methods. Additionally, the limits may be based on judgment, based on experience, based on intuition, or based on other measurements and

data. Or the limits may include combinations of some or all of the

above factors. Whenever possible, they should be empirically derived

or verified. The reliability of the estimate of the systematic

uncertainty will largely depend on the resourcefulness and ingenuity

of the metrologist."

In estimating various systematic errors, different approaches

appropriate to each individual error are utilized: direct calibration and

experimental verification, published data on the stability of components

and materials, design data, and experience with similar standards.

52
Several methods are used for combining the component errors to

obtain the overall uncertainty. The following two are the most commonly

used.

(a) combine linearly the random error at a certain accepted

confidence level and all systematic errors,

(b) combine the systematic errors in quadrature (as the square root

of the sum of squares, rss), and add linearly the result to the

random error.

The first approach tends to overestimate the error, but it does represent a quasi-absolute upper bound for the overall error, even if the probability

of reaching this upper bound is small. The second approach yields an estimated upper bound which is considerably smaller than the maximum possible

limit, if several component errors of similar magnitude are combined. This

method appears to be favored in engineering-type measurements [22,23],

The next three sections summarize the estimated measurement errors for

the two test systems as well as for the standard system. The latter consists

of the standard current transformer and shunt, the standard voltage divider,

and the standard power transducer. The standard system was used for loss

measurement on transformers on single phase basis and, thus, for direct

verification of test system 1. In all cases the random errors are much

smaller than the systematic errors; hence, the former are neglected.

8.2 Standard System

The errors relating to components of the standard system are summarized

in Table 10. In Tables 11, 12, and 13, the errors of the components are combined to yield the overall errors for the measurement of current, voltage,

and power. In power measurement, the error in current and voltage channels

Include the contributions from all the components listed in Tables 11 and 12,

except the rms voltmeter.

A current interaction term is added in Table 13. It is twice the error

of the current measuring instrument which, in this case, is the rms voltmeter. The interaction term does not include the error of the current

channel, which is common to both the current and power measuring instruments

and, therefore, does not contribute an additional error. The interaction term

is required because the load loss is determined and reported as a function of

the test current.

Using the data in Table 13, combined errors are calculated for several

power factors. They are tabulated in Table 14 and plotted in Fig. 30. It

should be added that the data in Table 13 are also applicable for calculating

the error in the measurement of no-load power loss. Instead of the current

interaction term, the voltage interaction term has to be used.

53
Table 10

Errors of Components of Standard System

Component

Magnitude

error

(ppm)

Phase Angle

error

(lirad)

Two-stage CT and shunt

1000-pF standard capacitor

100-pF HV capacitor

LV capacitor networks

Inductive voltage divider

RMS voltmeter

Power transducer

100

25 100 500

5 120 2000

8

5

5

510

50

Table 11

Errors in Current Measurement

Standard System

Source of error Magnitude

error

(ppm)

Phase angle

error

(urad)

CT and shunt

IV D RMS voltmeter

100

5 120

810

RSS error 156 13

Maximum error 225 18

54
8.3 Test System 1

The errors in power measurement of test system 1 are summarized in

Table 15. They include the contributions from the standard system as well as residual deviations of the test system after the alignment had been done. For

the magnitude and phase angle errors of the current and voltage channels, the

largest values from the Tables 4 and 5 are selected. For the errors of the

power transducer the largest three-phase average is selected from Table 6.

The errors for several typical low power factors are tabulated in Table 16 and

plotted in Fig. 31.

8.4 Test System 2

The errors of test system 2 are summarized in Tables 17 and 18 and

are plotted in Fig. 32. The method of selecting various error components

is the same as that for test system 1.

Table 12

Errors in Voltage Measurement

Standard System

Source of error Magnitudeerror

(ppm)

Phase angle error

(prad)

HV capacitor

LV capacitor networks

Active divider summing point

IV D RMS voltmeter

100 500

5 120

5

5

7 10

RSS error 524 14

Maximum error 725 27

55
Table 13

Errors In Power Measurement

Standard System

Source of error

RSS error Maximum error Magnitude (a)

(ppm)

Phase angle (b)

(urad)

Magnitude (a)

(ppm)

Phase angle (b)

(urad)

Current channel

Voltage channel

Power transducer

Current Interaction

effect

100

510

2000

240

13

14

50

105

605

2000

240

18

27

50

TOTAL 2080 54 2950 95

Table 14

Combined Errors in Power Measurement

Standard System

Power factor

RSS error Maximum error

a b/cos 9 Total a b/cos Total

CO80 (Z) (Z) (Z) (Z) (%) (Z)

0.01 0.21 0.54 0.58 0.30 0.95 1.25 0.02 0.21 0.27 0.34 0.30 0.48 0.78

0.05 0.21 0.11 0.24 0.30 0.19 0.49

0.10 0.21 0.05 0.22 0.30 0.10 0.40

0.20 0.21 0.03 0.21 0.30 0.05 0.35

Pi (l±a± b/cos e)

56
1.5

1.0

^ 0.5

^-» c

o

o

o3

GCO

GC

lu -0.5

-1.0

-1.5

_ \

1 1 T"

Error limits

Maximum error

RSS error

N

k /

I

I

I

J L

H \ \ h

0.02 0.04 0.06 0.08 0.10 0.12

POWER FACTOR

Figure 30. Standard System, Limits of Errors

57
Table 15

Errors in Power Measurement

Test System 1

Source of error

RSS error Maximum error Magnitude (a) (ppm)

Phase angle (b)

(yrad)

Magnitude (a) (ppm)Phase angle (b)

(urad)

Standard

Current channel

Voltage channel Power transducer

100 510

2000

13

1450

105 605

2000

1827

50

Test system

Current channel

Voltage channel Power transducer

3400

A 200 2200

96

106—

3400

4200

2200

96

106—

Current interaction 2000 — 2000 ~

TOTAL 6,505 153 14,510 297

*In calibrating the power transducers in the test system, the contributions from magnitude

and phase angle errors are conbined.

Table 16

Combined Errors in Power Measurement

Test System 1

Power factor

RSS error Maximum error

a b/co8 9 Total a b/cos e Total

cos© (%) (%) (%) (%) (%) (%)

0.01 0.65 1.53 1.66 1.45 2.97 4.42

0.02 0.65 0.765 1.00 1.45 1.48 2.93

0.05 0.65 0.31 0.72 1.45 0.59 2.04 0.10 0.65 0.15 0.67 1.45 0.30 1.75 0.20 0.65 0.08 0.65 1.45 0.15 1.60

Pt - Pi (l±a±b/co8 9)

58
5

4

3

2

¥ 1 o

3

O

trcc -1h

-2

-3

-4

-5

Error limits

Maximum error

RSS error

_ /

1 1

0.02 0.04 0.06 0.08 0.10 0.12

POWER FACTOR

Figure 31. Test System 1, Limits of Errors

59
Table 17

Errors in Power Measurement

Test System 2

Source of error

RSS error Maximum error Magnitude (a) (ppm)

Phase angle (b)

(prad)

Magnitude (a) (ppm)

Phase angle (b) (urad)

Standard

Current channel

Voltage channel Power transducer

100 510

2000

13

1450

105 605 20001827

50

Test System

Current channel

Voltage channel Power transducer

2000

2000 6500

30

20—

2000

2000 6500

30

20—

Current interaction 2000 — 2000 —

TOTAL 7,650 65 15,210 145

*In calibrating the power transducers in the test system, the contributions from

magnitude and phase angle errors are combined.

Table 18

Combined Errors in Power Measurement

Test System 2

Power factor

RSS error Maximum error

a b/cos e Total a b/cos e Total

C089 (Z) (Z) (%) (Z) (Z) (Z)

0.01 0.765 0.65 1.00 1.52 1.45 2.97 0.02 0.765 0.32 0.83 1.52 0.73 2.25 0.05 0.765 0.13 0.78 1.52 0.29 1.81 0.10 0.765 0.07 0.77 1.52 0.15 1.67 0.20 0.765 0.03 0.77 1.52 0.07 1.59

Pt - Pi (l±a± b/cos 9)

60
1 1 r

Error limits — Maximum error — RSS error

^

1 ^^c

O

a

o

LU -1

-2

4 h

-3 _ /

-4

0.02 0.04 0.06 0.08 0.10 0.12

POWER FACTOR

Figure 32. Test System 2, Limits of Errors

61
9. SYSTEM- BASED VERIFICATION

9.1 Experimental Method and Results

The alignment and calibration of the two test systems is described in

section 7. The calibration results for various system components are combined

in sections 8.3 and 8.4 to yield the overall measurement error for the test

systems.

This approach to verifying the principal components of the test system is

highly desirable, as it facilitates diagnosing potential malfunctions in the

test system and correcting them. It also reduces the number of possible

calibration points and allows an orderly analysis of errors.

Equally desirable, if not essential, is some verification of the test

equipment on a system basis, whereby the measurement results of the test

•system can be compared with those of another, preferably more accurate, system

when both measure the same transformer under test. A system based

verification may uncover error-producing interactions among the components of

the test systems. This step provides an assessment of the severity of

electromagnetic interference during the test operation, and verifies whether

the methodology of combining the uncertainties is reasonable.

Test system 1 was verified against the standard system on a single-phase

basis by measuring simultaneously with both systems, the losses of each phase

of a test transformer separately, and then summarizing the results.

Verification was done on twenty transformer configurations that cover the

principal current and power factor ranges of medium size power transformers

manufacturered at ASEA Electric.

The data of the system-based verification is tabulated in Table 19. To

illustrate how well the data compare with the estimated errors in section 8.3,

the differences in power as measured with the standard and test systems are plotted in Fig. 33. The rss error limits are also plotted, as well as the rss limits corrected for the rss contribution by the standard system.

9.2 Discussion of Experimental Results

Several items from the data in this and the previous two sections are noteworthy and can be summarized as follows:

(1) Both test systems have rss error limits that are ±1% or smaller for

power factors of 0.02 and larger.

(2) The system-based tests support a contention that the rss limits are very reasonable for estimating the system accuracy. Of the 20 test

points in Fig. 33, 19 are well within the rss limits; one test point

is at the limit. The data in Table 19 indicate that the 95%

confidence (2a) limit for the dispersion of the system errors for

power measurement is within the estimated rss errors. Also, the

dispersion of the errors in current and voltage measurement are within the figures used for estimating the contributions from the

current and voltage channels.

62
(3) The maximum error limits represent, as expected, an overly

pessimistic estimate.

(4) The phase angle errors of test system 2 are reduced after

alignment to a negligible value, "^30 yrad maximum. For test

system 1, the residual errors are larger, of the order of 100 yrad

maximum.

(5) A somewhat surprising result is that the magnitude errors for both

test systems are somewhat larger than expected, possibly because the

greatest emphasis in alignment and retrofitting was placed on the

phase angle error. In future alignments, the causes of magnitude

errors will be investigated.

(6) The accuracy goals for the standard system are well exceeded; the

current, voltage, and power calibrations meet the present and

foreseeable future requirements for medium size power transformers.

63
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64
2.0

1.5-

1.0-

^

0.5„ -

c

CDO

i_

S

O

cc^ -0.51-

-1.0-

-1.5 -

-2.0

1 1 1 1 Error limits (RSS)

\ - Test system

- Test system, without ~

\\ contribution from calibration

\\

\\

\\

system

—

£ •

°l °*l n ° 1 1

1 a

1 D

Dpi 1

D _ D

<^^' ^

//

it

ft Current ranges

1 1

1

• 10 A

1 If D 100 A

1

1

^ 1000 A

1 1 1 0.02 0.04 0.06 0.08

POWER FACTOR

0.10 0.12

Figure 33. System-Based Verification, Test System 1 Comparison of results as obtained by the

test system 1 and standard system

65
10. MAINTENANCE OF STANDARD INSTRUMENTS AND TEST SYSTEMS

10.1 General Comments

This section provides guidelines for the future calibration of standard

instruments and alignment and calibration-related maintenance of test

systems. These guidelines are based partly on published information on the

stability and failure modes of the devices, as well as on the authors'

experiences. The suggested calibration intervals may have to be adjusted as the history of various standard instruments and test systems is accumulated.

The fact that drifts in such devices as standard capacitors and resistors are

in the magnitude of the quantity, and not in the phase angle, alleviates the

calibration problem. In fact, barring significant failures, the phase angle

of all standard instruments should remain negligible indefinitely.

The process of calibrating various standard instruments has a number of

redundancies. Thus, a failure in an instrument is likely to become apparent

in in-house verifications. For example, the ratios and phase angles of both

the standard current transformer and the inductive voltage divider should

remain unchanged. As the two are compared, a failure in one will become

apparent by a relative change between the two. Recordkeeping and convenient

presentations of the measurement data are essential in the maintenance of

standards. Control charts, and plots of measured data versus time are convenient forms of recordkeeping. It is strongly recommended that the

calibration instruments be reserved only for the purpose of performing

calibrations and checks in order to minimize the possibility of failures due

to accidentally produced excessive voltages and currents and mechanical

abuses.

In the use and maintenance of the standard instruments, frequency and

harmonic influences have been considered. Most of the standard instruments —

current transformers, shunts, standard and high voltage capacitors, standard

power transducers — are unaffected by moderate frequency variations up to a few percent. These standard instruments also have adequate harmonic response

up to at least the tenth harmonic.

Low voltage capacitive networks used in conjunction with an active

divider, and the phase shifter are frequency-sensitive instruments. Effects

of their frequency dependence are summarized. In the capacitive network shown

in Fig. 21, the effective negative transconductance varies quadratically with

frequency. Thus, for a small change in frequency, the change in the

compensating transconductance is twice as large. However, this presents no problem if the dissipation factor of the capacitor is small. For example, if

the original dissipation factor is 100 ppm, a frequency change of one percent

Increases or decreases the effective negative transconductance by 2 ppm.

Typically, the dissipation factor of these capacitors is relatively

independent of frequency. Therefore, the overcompensation or undercompensation is only 1 ppm. In either case, the effect due to small

changes in the fundamental frequency is negligible.

In the case of the harmonics, the capacitive network is overcompensated

leading to an effective capacitance with negative dissipation factor.

However, for lower harmonics — those of interest in tests — the effective

dissipation factor levels will not exceed 1000 ppm. Again, the effect on

66
measuring the load losses is negligible, if the harmonic content in the supply

is moderate. Only if tests are performed at different fundamental frequency,

e.g., 50 Hz, is there a need to change the compensation networks.

In using the active phase shifter of Fig. 25(b) for the calibration of

power transducers, the applied power varies as the reciprocal of frequency.

Hence, a signal generator having frequency accurate to about to. 05% is

required.

10.2 Standard Current Transformer

As losntioned in section 4.4, in the absence of obvious failures, the

current transformer should remain stable indefinitely. Functional

verifications, whereby selected ratios of the current transformer are compared

against those of the standard inductive voltage divider, and the outputs of

the windings of the current transformer are compared, are recommended as in-house checks. Only Infrequent calibrations at NBS are needed.

A summary of suggested calibrations and verifications is as follows:

a. Comparison of all ten taps in the secondary winding of the current

transformer against the standard inductive voltage divider. Also a similar comparison using the primary and secondary windings to yield

10/1 ratio. Conduct these tests yearly, in-house.

b. Calibration of 1000/1, 100/1, and 10/1 ratios and the corresponding

trans impedances with the coaxial shunts. Conduct this calibration

every five years, at NBS.

10.3 Standard Shunts

As noted earlier, the phase angle is of primary importance, and since it

depends primarily on the geometry of the shunt, the phase angle can be

expected to remain stable indefinitely unless the shunt is physically

damaged. One of the shunts used by ASEA Electric was calibrated by NBS and

its phase angle was found to be negligible. Since the other shunt had

identical geometry, its phase angle was also assumed to be negligible. The

relative phase angle of both shunts can be monitored in-house by interchanging

them in the two-stage current transformer circuit of Fig. 6. Some long-term

resistance drifts can be expected, but they should not exceed ten ppm per year

and thus should be negligible. Monitoring of the dc resistance is recommended

to detect possible accidental abuses.

Here is a summary of suggested calibrations and verifications:

a. DC resistance measurement. Conducted yearly, in-house.

b. Relative phase angle verification using a substitution method.

Conducted yearly, in-house.

c. Trans impedance measurement in conjunction with the calibration

of the standard current transformer (as in 10.2).

67
10.4 Standard Capacitors

The standard 1000-pF unit is the primary reference. Its capacitance is

expected to remain stable within a few ppm per year. The dissipation

factor, assumed to be 5 ppm or smaller, should remain negligible

indefinitely. The high voltage capacitor is a completely shielded unit, that

is insulated with compressed gas. Two capacitances, 100 pF and 20 pF, are available; only the 100-pF is critical. The dissipation factor is

negligible and should remain such. The capacitance value is generally

affected by the mechanical rigidity of the device and by the amount of gas in

the device. Under normal circumstances, the capacitance should remain stable

within 0.01% per year. The solid-dielectric capacitors, O.Ol-yF, 0.01 yF, and

1.0-yF units, in the feedback loop of the active divider have a temperature

dependence of capacitance of approximately 100-ppra/°C. Long-term drifts of

the order of 0.05% per year are possible. The dissipation factor of the

solid-dielectric capacitors is in the 60 ppm range. The effective

dissipation factor of the transfer capacitance is compensated by the use of

the RC type T-networks. It is expected that the dissipation factor of the

compensated capacitive network will remain stable within 5 ppm/year. A

complete check of all capacitors against the 1000-pF standard is recommended

before each major realignment operation. Unexpected changes either in the

capacitance value or the dissipation factor could serve as indicators of

possible catastrophic failures.

Summary of suggested calibrations and verifications:

a. Relative capacitance and dissipation factor measurements —

1.0-yF, 0.1-yF, 0.01-yF, and 100-pF units against the standard

1000-pF capacitor. Use the bridge of Fig. 23, or, preferably,

that of Fig. 24. Conducted yearly, in-house.

b. Calibrate 1000-pF capacitor at NBS. Every two years.

c. Optional — consider acquiring another set of solid-dielectric

capacitor networks (1.0 yF, 0.1 yF, and 0.01 yF). Use these

for one-to-one comparison with the capacitors in the original set.

68
10.5 Resistors In the Phase Shifter

The most Important factor Is the phase angle of the 26,500-ohm

resistor. A miniature metal film device Is selected having negligible

inductance. The effective shunt capacitance is smaller than 0.5 pF, thus,

yielding a phase defect of 5 mlcroradians or smaller at 60 Hz. The changes

in geometry that would significantly affect the phase angle are gross and thus

improbable. Nevertheless acquistions of two input resistors is recommended In

order to facilitate checking by a substitution method. The phase angle of the

resistors in the feedback path is not important. The knowledge of resistances

to ±0.05% of all the resistors is desirable in order to determine the real

component of the product of input and output voltages.

Summary of suggested checks and calibrations:

a. Phase angle comparison of two input resistors by a substitution

method. Yearly, in-house.

b. DC resistance of all resistors. Yearly, in-house.

10.6 Inductive Voltage Divider

Barring catastrophic failures in transformers or switches, the ratio and

phase angle of the Inductive voltage divider should remain stable

indefinitely. The procedures for checking the standard current transformer

described in section 10.2 will also check out the Inductive voltage divider.

Additional calibration at NBS is recommended, partially to check the

integrity of switches, since some of the failures in switches can produce

subtle changes in the ratio of the divider.

Suggested calibration:

a. Ratio and phase angle. Every two years at NBS. As the stability

data are accumulated and if they are favorable, the frequency of

calibration at NBS can be reduced, e.g., to every five years.

10.

7

Standard Power Transducer

The phase shifter is designed to calibrate the power transducer at low

power factors. A high accuracy ac voltmeter can be used for calibrations at

unity power factor. Since the redundancy in this operation is not as strong

as for other calibrations, it is suggested that spot calibrations of the

transducer be performed at NBS until evidence is available to demonstrate the

reliability of the phase shifter and voltmeter approach.

Suggested calibrations and verification:

a. Calibrate the power transducer using the phase shifter and voltmeter.

Conducted yearly, in-house.

b. Perform spot calibrations, e.g., two per range, of the transducer at

NBS, every two years.

69
10.8 Maintenance of Test Systems

In order to ensure test system performance within the error limits

established in the earlier sections, periodic alignments and verification on system basis are essential. Design of both systems Incorporates built-in low

voltage generators (calibrators) that allow for routine verification for

proper system operation. These calibrators are signal sources with controlled

power factor and controlled signal amplitudes. They feed known low voltage

signals into voltage and current buffers. Thus, signal paths beyond primary

voltage and current scaling inputs can be readily verified for proper

operation. This type of routine functional checking assures that no major

drifts or misalignments have taken place. These checks, however, do not

provide traceable calibration of the total measurement system.

Experience gained over the years on the operation and stability of both

test systems indicate that these internal generators do not produce

calibration signals that are of adequate phase stability. The amplitude

stability, however, is quite good. Hence, the use of a built-in generator

should only be made to provide gross assurance of systems performance. Use of

the signals from calibrators for the alignment of current and voltage buffers

is not desirable.

Based on ASEA Electric' s experience, this gross system check by use of

built in generators should be done once a month. This check should be used to

provide evidence of system operation without major malfunctions. Data derived

from such checks can show trends in drifts and signal degradation. Such

checks can also be used as trouble shooting aids in localizing faults in test

systems. Over the years, such generators have identified faulty operational

amplifiers, rms transducers, and unstable power transducers. Alignment of

current channels, voltage channels, power transducers, as well as system-based

verification utilizing procedures described in previous sections or equivalent

procedures are required to obtain calibration data that can be defined as being traceable.

As discussed earlier, both systems were modified at ASEA Electric to

provide noninteracting adjustments in both the voltage and current channels.

Modifications in circuit designs were also necessary to limit the effect of

drifts in circuit components. In order to ensure that these modifications

yielded desired performance, current, voltage and power channel alignments

were checked three times within the past year using the procedure outlined

earlier. Results of these checks indicate that implemented modifications

were necessary to control and eliminate component drifts. System alignments

did not change significantly. Due to the critical nature of measurements

and their significant economic impact, it is felt that traceable system-based

verification and alignments of the subsystems be carried out at least once every year.

70
The necessity for a yearly alignment is based on the following

considerations

a. Although standard gas capacitors are extremely stable devices, gas

pressure changes due to leaks or loading, vibrations, or ambient

environmental changes can introduce a variation in active capacitance

that may have a significant impact on measurement accuracies,

b. Precision current transformers should provide stable long-term

performance. However, subtle failures of transformers and their

burdens are possible and could introduce measurement errors.

c. Electronic components do drift and age. Their performance may slowly

change over a period of time. In addition, the effect of factory

environment and electrical noise levels may create conditions that can adversely affect the system performance.

A long-term, (e.g., five-year) study of trends in the stability of

subsystems alignment should be made before system-based verification and

alignment requirements can be relaxed.

71
11. DISCUSSION AND CONCLUSIONS

Discussion

The authors hope that some of the technical ideas, calibration hardware,

measurement procedures, and methodology for evaluation of system errors described in this report will help lead to an industry-wide consensus standard

dealing with the requirements of accuracy and traceability in measurement of

transformer losses. By this, or an equivalent procedure, a uniform comparison

of losses in power transformers could be made. A complete statement of

measurement accuracy for a loss test system requires that the measurement

uncertainties of the system be shown through a calibration process that is

traceable.

In order to develop a diverse data base related to transformer loss

measurement uncertainties, a "round-robin" comparison of measurement results

on the same transformer would be necessary. Such a comparison would provide

an excellent indication of the state of overall measurement capability within

the industry.

Although the calibration (standard) system described in this report was specifically tailored to calibrate two test systems at ASEA Electric, the

concepts and many of the developments presented in this report are broadly

applicable for the calibration of all types of test systems shown in Fig. 1.

With relatively simple modifications, test systems that utilize bridge

techniques for loss measurements can be calibrated also.

To achieve high accuracies in the two test systems at ASEA Electric,

considerable effort and care were expended in retrofitting and modifying the

principle components. In the authors' judgment, this work should be performed

by a skilled technician under the guidance of a knowledgeable engineer. This

will ensure that future maintenance and alignment of the test system, as well

as on-line equipment calibrations would be carried out by an experienced and

qualified personnel.

Conclusions

The following conclusions summarize the developments reported in this

technical note:

(1) The calibration (standard) system described in this document

provides a means to achieve traceable measurements of transformer

losses. This system is utilized not only to perform calibration and

alignment of principle components in a test system, but also is

designed to provide a system-based verification for the entire loss

test system.

(2) Methodology for combining errors in principal components to arrive

at an estimate of total test system error at various power factors

is presented. System-based verification tests demonstrated the

validity and adequacy of this approach.

72
(3) The self-contained calibration plan detailed in his report provides

a simple means to achieve traceability. This approach requires

sending only a few small instruments for infrequent calibration to

outside organizations such as NBS.

(4) The necessary steps for improving measurement accuracy and for

establishing traceability are detailed in Appendix 2. It is

estimated that the cost of such actions may be only a fraction of

the cost of losses in a single medium-size power transformer.

(5) Suggestions are made regarding accuracy goals for loss test systems

of the future. Specifics are presented in Appendix 2.

73
12. REFERENCES

[1] Belanger, B. C, Traceability: An Evolving Concept, ASTM Standardization

News (January 1980).

[2] Boyd, E. L. and Brost, J. D. , Design Concepts For An Amorphous Metal

Core Distribution Transformer, IEEE Transactions on Power Apparatus and

Systems, Vol. PAS-130 (November 1984).

[3] Ponni, K. and Majuri, M. , The Measurement of Reactor Losses with a Schering Bridge, Kraft och Ljus (Finland), Vol. 33, No. 10, (October

1960) (In Swedish).

[4] Moore, W.J.M. and Raftis, F. A., Measurement of Shunt Reactor Loss at

High Voltage With An Alternating Current Comparator Bridge, IEEE

Transactions on Power Apparatus and Systems, Vol. PAS-92 (March/April

1973).

[5] Malewski, R. and Douville, J., Mesure Des Pertes Dans Les Reactances

Shunt EHT Par Un Pont C-tan6 Insenibillse Aux Fluctuations De La

Frequence Du Reseau, IEEE Canadian Communications and Power Conference,

Montreal, Ch 1373-0, Reg. 7, (1978).

[6] Malewski, R. , Anderson, W. E., and Moore, W.J.M. , Interlaboratory

Comparison of EHV Shunt Reactor Loss Measurements, IEEE Transactions on Power Apparatus and Systems, Vol. PAS-99, No. 4 (July/August 1980).

[7] Moore, W. J. M. , Love, G. , and Raftis, F. A., Measurement of Short

Circuit Load Losses in Large Three-Phase Power Transformers Using An

Alternating Current Comparator Bridge, IEEE Transactions on Power

Apparatus and Systems, Vol. PAS-94, No. 6, (November-December 1975).

[8] So, E., A Frequency-Compensated High-Voltage Current-Comparator-Based

Capacitance Bridge for Shunt Reactor Loss Measurement, IEEE Transactions

on Instrumentation and Measurement, Vol. IM-31, No. 1 (March 1982).

[9] So, E., An Improved Frequency-Compensated Capacitance Bridge for

Accurate Shunt Reactor Loss Measurements at Very Low Power Factors, IEEE

Transactions on Power Apparatus and Systems, Vol. PAS- 103, No. 5 (May 1984).

[10] Brooks, H. B. , and Holtz , F. C, The Two Stage Current Transformer, AIEE

Transactions, Vol. 41 (June 1922).

[11] Petersons, 0., A Self Balancing Current Comparator, IEEE Transactions on Instrumentation and Measurements, Vol. IM-15, No. 1,2 (March-June 1966).

[12] Souders, T. M. , An Audio-Frequency Four-Terminal Resistance Bridge, IEEE

Transactions on Instrumentation and Measurements, Vol. IM-23, No. 4

(December 19 74).

74
[13] Petersons, 0. and Anderson, W. E. , A Wide Range High Voltage Capacitance

Bridge With One PPM Accuracy, IEEE Transactions on Instrumentation and

Measurements, Vol. IM-24 , No. 4 (December 19 75).

[14] So, E., The Application of The Current Comparator In Instrumentation For

High Voltage Power Measurements at Very Low Power Factors, Paper No.

85SM317-3. Presented at IEEE/PES 1985 Summer Meeting.

[15] Cutkosky, R. D. , Active and Passive Direct-Reading Ratio Sets for the

Comparison of Audio-Frequency Admittances, Journal of Reseach NBS C.

Engineering and Instrumentation, Vol. 68C, No. 4, (October-December

1964).

[16] Tomota, M. , Sugiyama, T., and Yamaguchi, K., An Electronic Multiplier

for Accurate Power Measurements, IEEE Transactions on Instrumentation

and Measurements, Vol. IM-17, No. 4 (1968).

[17] Turgel, R. S. and Oldham, N. M. , High-Precision Audio-Frequency Phase

Calibration, IEEE Transactions on Instrumentation and Measurements,

Vol. IM-27, No. 4 (December 1978).

[18] Specht, T. R. , Rademacher, L. B., and Moore, H. R, , Measurement of Iron

and Copper Losses in Transformers, AIEE Transactions, Vol. 77, pt. Ill,

Aug. 1958, pp. 4 70-4 76.

[19] Malewski, R., A., Arseneau, R. , So, E., and Moore, W. J. M. , Comparison

of Instrumentation for Measuring the Losses of Large Power Transformers,

IEEE Transactions on Power Apparatus and Systems, Vol. PAS- 102, No. 6

(June 1983).

[20] Mehta, S. P., Measurements of Transformer Losses. Paper No. 84EHO

209-7-PWR. Part of IEEE/PES Tutorial presented at 1984 IEEE PES Winter

Meeting.

[21] Eisenhart, C. , Ku, H. H. , and Colle, R. , Expressions of the

Uncertainties of Final Measurement Results, NBS Special Publication

No. 644 (January 1983)

[22] Lentner, K. J., A Current Comparator System to Establish the Unit of

Electrical Energy at 60 Hz. IEEE Transactions on Instrumentation and

Measurements, Vol. IM-23, No. 4 (December 19 74).

[23] Moore, W, J. M. and So, E., A Current-Comparator-Based System for

Calibrating Active/Reactive Power and Energy Meters, IEEE Transactions

on Instrumentation and Measurements, Vol. IM-32, No. 1 (March 1983).

75
13. APPENDIX 1 ~ CALIBRATION REPORTS AND DATA

13.1 Introductory Comments

Calibration or test reports for six standard instruments are included in

this technical note. These reports are, for the most part, self explanatory.

However, several comments are in order. Report 13.4 on the 1000-pF standard

capacitor does not contain information on the dissipation factor since the

dissipation factor values are not provided with this calibration service.

Comparisons of high quality gas-dielectric capacitors indicate that the

dissipation factors do not differ among units by two or three parts per

million, Hence, it is assumed that the absolute value is within 5 ppm.

Report 13.6 on the inductive divider gives large estimated calibration

uncertainty for the quadrature component — ±5 x 10" of the input

quantity. At small ratios the uncertainty becomes quite large, e.g.,

± 50 x 10~ at the ratio of 0.1. The inductive divider was compared against

the two-stage transformer using the circuit of Fig. 14(a). The results are reported in section 13.6, Table 20. Combining on rss basis the largest phase

angles of ratio (b/A) in Tables 3 and 20 with the measurement uncertainty in

Table 20, yields ±6.0 yrad. It is impractical to apply phase angle correction

to the inductive divider. The residual corrections are treated as errors. On

basis of the above data an allowance of ±10 yrad is made for possible phase

angle errors of inductive divider.

76
13.2 Two-Stage Current Transformer

FORM N85— 259 (REV. t2-7t> U.S. Department of Commerce

NATIONAL BUREAU OF STANDARDS

WASHINGTON, D.C. 20234

REPORT OF TEST

TWO-STAGE CURRENT TRANSFORMER

10-1000 to 1 Ampere, 60 Hertz

Submitted by ASEA Electric Inc. 400 South Prairie Avenue

Waukesha, Wisconsin 53186

Secondary

Frequency Secondary Current Current Phase Angle

(hertz) Burden (amperes)

1.0 10

Ral

X 1

;io

000000

(microradians)

60 A -4

0.5 1 000000 -4 0.2 1. 000000 -3 0.1 0. 999999 -2

B 1.0 1. 000000 -1 0.5 1 000000

0.1 0. 999998 +3

A 1.0 100 X 999997 -3 0.5 999997 -3 0.2 999997 -3 0.1 999997 -4

B 1.0 999997

0.5 999997

0.1 999997 -1

A 1.0 1000 X 1 000000 -2 0.5 1 000000 -2 0.2 999999 -2 0.1 1 000000 -2

B 1.0 1 000000 +1 0.5 1 000001 +1 0.1 1 000000 +2

Date of Test: June 27, 1984

722/233345-84

77
TWO-STAGE CURRENT TRANSFORMER

10-1000 to 1 Ampere, 60 Hertz ASEA Electric Inc.

Secondary burden A consisted of separate resistances of 2 ohms each in both the secondary and tertiary circuits. Secondary burden B consisted of resistances less than 0.1 ohm each in both the secondary and tertiary

circuits.

It is very unlikely that the above values of current ratio and phase angle are

in error by more than 10 ppm and 10 microradians, respectively. These figures

include allowances for both the random and systematic errors of the calibration process.

For the Director

National Engineering Laboratory

Barry A. Bell, troup Leader

Electronic Instrumentation and Metrology

Electrosystems Division

Test No: 722/233345-84

Order No: 76625

Date: August 8, 1984

78
13.3 Two-Stage Current Transformer/ Shunt Combination

FORM NBS—1»

IREV. Il-7>l U.S. Department of Commerce

NATIONAL BUREAU OF STANDARDS

WASHINGTON. D.C. 20234

REPORT OF TEST

TWO-STAGE CURRENT TRANSFORMER/AC SHUNT COMBINATION

ASEA Transformer, 10-1000 to 1 Ampere, 60 Hertz

Submitted by ASEA Electric Inc. 400 South Prairie Avenue

Waukesha, Wisconsin 53186

The transfer resistance and phase angle of this transformer-shunt combination

were measured in June, 1984, with the shunt having serial no. 6829 connected

across the transformer's secondary winding, and the remaining shunt,

serial no. 6828, connected across the tertiary winding. The measured values

and test conditions are reported below.

Room Temperature

24°C

Frequency

(hertz)

6565

Transformer

Ratio

100/1

1000/1

Primary

Test Current

(amperes)

2040

Resistance

(ohms)

0.01999302

0.00199930

Phase Angle

(microradians)

-2.6

-1.8

It is unlikely that the above values of resistance and phase angle are in error by more than 0.01 percent and 5 microradians, respectively.

For the Director

National Engineering Laboratory

Barry A. BelU/aroup Leader

Electronic Instrumentation and Metrology

Electrosystems Division

Test No: 722/233345-84

Order No: 76625

Date: August 8, 1984

79
13.4 Standard Capacitor

U S. Department of ComtntTce NATIONAL BUREAU OF STANDARDS

National Measuvement Laboratory

Washington, DC 20234

REPORT OF CALIBRAIION

Description of Standard

Standard Capacitor

Submitted By:

PTE-ASEA

Waukesha, WI 53186

P No 60794

Temperature 23 Degrees Celsius

Date of Calibration: 1 March 19S4

FREQUENCY CAPACITANCE LIMITS OF UNCERTAI NT , (Hertz) (Picofarads) = Total

100 999 95 e ppm 15 p p

For additional information regarding the calibration of staridc.Td capiicitovs

the user should consult the information sheet<s) enclosed with this report.

For The Director

National MeasuTement Laboratoru

Norman B Belecki, Physicist

Center For Basic Standards

Electricity Division

Test No 232213

Date: i March 19S4

80
13.5 Capacitor Networks

°FORmNB*-44S

(MIV. I>-W)

U.S. DEPARTMENT OF COMMERCE

NATIONAL BUREAU OF STANDARDS

WASHINGTON, O.C. 2Sej4

REPORT OF CALIBRATION

CAPACITORS

0.01, 0.1, and 1 Microfarad

Submitted by

ASEA Electric Inc. 400 South Prairie Avenue

Waukesha, Wisconsin 53186

The capacitance and dissipation factor of the three capacitors were measured

on June 27, 1984, at a frequency of 60 Hz and a room temperature of 21.5°C. The results are tabulated below.

Voltage Capacitance

(volts) (farads) Dissipation Factor

0.01 Microfarad Nominal Standard

10 1.000332 X 10-^ +6 x 10"^ dissipation factor adjusted to zero

10 1.000346 X 10-^ X 10'^ 20 1.000358

35 1.000383

5 1.000345

0.1 Microfarad Nominal Standard

10 1.001018 X

10-''' -13 X lO-'' dissipation factor adjusted to zero

10 1.001020 X 10-^ X 10-^

20 1.001021

35 1.001027

5 1.001022

1 Microfarad Nominal Standard

10 0.999786 x 10"^ x 10"^ 20 0.999786

35 0.999792 +1

5 0.999788

722/233345-84

81
CAPACITORS

0.01, 0.1, and 1 microfarad

ASEA Electric Inc.

- 2

Estimated Uncertainties

Capacitance: ±0.003% of measured value

Dissipation Factor: ±5 x 10"^ Temperature: ±1*C

For the Director

National Engineering Laboratory

^N/ Robert E. Hebner, Group Leader

^ Applied Electrical Measurements Electrosystems Division

Test No: 722/233345-84

Order No: 76625

Date: August 7, 1984

82
13.6 Inductive Divider

U. S. Department of Commerce NATIONAL BUREAU OF STAtvlDARD

National Measurement Laboratory

Washington. D. C. 20234

REPORT OF CALIBRATION

Description of Standard:

Submitted By:

RTE/ASEA Corporation

Waukesha, WI 53186

P No 60794

Date of Calibration: 10 February 1984

Temperature: 23 Degrees Celsius

Input Voltage Distortion: Negligible

Output Burden: Negligible

Case connected to input low terminal (COMMON)

The in-phase and quadrature correctiona for each dial setting gi^en

below are the four-terminal transfer ratio corrections Any dials not

given in the dial setting were set to zero. To obtain the actual

transfer ratio for a dial setting, the corrections must be multiplied by

10 exp-6 and then added as the appropriate complex components to the

nominal value of the dial setting.

The standard deviations given below are those for an individual

measurement of the corrections To relate these to the actual transfer

ratios, they must be multiplied by 10 exp-6. The uncertainty will be

increased if the magnitude of the standard deviation of any individual

step is sufficiently large.

For further information regarding the use of the calibration results and

the calibration of this standard, the user should consult the

information sheet(s) enclosed with this report.

Test No 232213

83
Decade Transformer Divider Page

Fre quencu 100 Hertz Inp ut Voltage: 25 Volts

CORRECTIONS STANDARD DEVIATION NUMBER

DIAL SETTING IN-FHASE QUADRATURE IN-PHASE QUADRATURE MEASUREMENTS

•i 00 0. 0. 000 0. 00 3

. 9 23 -1. 5 0. 000 0. 00 3

. S o 25 -1 6 0. 000 00 3

. 7 2= -1 4 0. 000 0. 00 3

h 19 -1. 1 0. 000 00 3

5 13 -0 7 0. 000 00 3

4 11 -0. 3 0. 000 0. 00 3

3 09 0. 1 000 0. 00 3

2 0. 04 0. 3 0. 000 0. 00 3

1 03 0. 3 0. 000 0. 00 3

0. 02 000 0. 00 3

OX -0 ^-7

1. 4 0. 000 00 3

09 -0, 22 1. 1 0. 000 0. 00 3

OS -0 19 9 0. 000 0. 00 3

07 -0. 16 8 0. 000 0. 00 3

06 -0. 13 7 000 0. 00 3

05 -0. 11 6 0. 000 0. 00 3

04 -0. OB 5 0. 000 00 3

03 -0 05 0. 4 000 0. 00 3

02 -0. 03 3 0. 000 0. 00 3

01 0^ 00 2 000 0. 00 3

. 00

-0 05 3 000 00 3

009 -0 04 2 000 0. 00 3

OOP, -0. 04 0. 2 0. 000 0. 00 3

007 -0. 03 2 0. 000 0. 00 3

006 -0 02 2 0. 000 0. 00 3

005 -0 01 0. 1 000 0. 00 3

004 -0. 01 1 000 0. 00 3

003 0. 00 0. 1 0, 000 0. 00 3

002 0. 01 0. 1 000 0. 00 3

001 0. 02 0. 000 0. 00 3

OOOX 01 0. 2 0. 000 0. 00 3

699X -0. 07 -0. n 000 0. 00 3

29<?y -0. 20 1. 4 000 0. 00 3

The limits of uncertainty for the corrections above are estimated to be

0. 5 X 10 exp-6 for the in-phase corrections and 5. x 10 exp-6 for the

quadrature corrections.

For The Director

National Measurement Laboratory

^^^ . _^ y^- '^s^ui^

Norman B. Belecki, Physicist

Center For Basic Standards

Electricity Division

Test No 232213

Date 10 February 1984

8A
GENERAL INFORMATION ON STANDARD DECADE TRANSFORMER DIVIDERS

Standard decade transformer dividers are calibrated by comparison with

NBS reference standards. The correction given In the calibration reports

are in terms of the four- terminal transfer ratio at the terminals of the

divider. In the following, the definitions and nomenclature follow the

American National Standard for decade transformer dividers (voltage type),

ANSI ClOO.l - 1972.

The transfer ratio of a decade transformer divider is defined as V^i ' A„ * a . jb

where

E » output voltage

E^ •= input voltage A = nominal transfer ratio (dial setting of the divider)

n

g a = transfer ratio in-phase correction times 10

b = transfer ratio quadrature correction times 10"

As indicated above, the corrections given in the calibration report must be multiplied by 10' before use in the equations.

If in actual use the Input terminals are defined at the ends of connecting

leads, the voltage drop across these leads enters into the ratio corrections

in the following way.V^i ' ^\ * * * J'')/^^ "

h'''-2}

where

I E^ - E^ * voltage across input leads

Z-. > Impedance of input leads

Z2 input impedance of the divider

The transfer ratio angle correction is defined as

where

e - tan"^ tb/(A^ + a)]

e = transfer ratio angle correction (In radians)

The transfer ratio angle correction is included here because the previous

calibration may contain the quadrature component as a phase angle which

is the same as the transfer ratio angle correction. A useful approximation

to the above equation is

e » b/A„

85
Table 20

Comparison of Inductive Divider

against Two-Stage Transformer

Tap of Quadrature correction

two-stage of IVD relative to

transformer two-stage transformer

(A) (b, absolute value)

0.1 0.30 X

10"^

0.2 0.56

0.3 0.69

0.4 0.84

0.5 0.85

0.6 0.78

0.7 0.70

0.8 0.64

0.9 0.36

1.0 3.00

-I2/I1 = V1/V2 = A + a + jb

where A Is the nominal ratio,

a Is the In phase correction

b Is the quadrature correction

phase angle of ratio '"b/A radians.

Estimated uncertainty In measurement of b = 5A x 10-b

86
13.7 Power Transducer

FORM NBS— 2M

IREV. l2-7al U.S. Department of Commerce

NATIONAL BUREAU OF STANDARDS

WASHINGTON, DC. 20234

REPORT OF TEST

WATT CONVERTER

1.25, 2.5, 5 Volts, 0.01, 0.1, 1.0 Power Factors

Submitted by

ASEA Electric Inc. 400 South Prairie Avenue

Waukesha, Wisconsin 53186

The watt converter was tested as a voltage multiplier at 60 Hz with nearly

sinusoidal waveforms. The test was performed in June 1984 at a room temperature of approximately 23°C.

Measurements were made after a warm-up period of at least one hour by

applying voltages of known rms value and phase angle to the converter inputs

VI and V2. The power factor switch was adjusted to give maximum resolution

and the reading on the converter digital display was recorded. The voltage

waveforms were digitally synthesized by the NBS Phase Angle Standard and measured by a calibrated rms responding voltmeter.

To minimize interactions between the power line and the synthesized test signal, all of the instrumentation was powered from a 70-Hz supply.

The converter reading was adjusted to zero with VI and V2 disconnected,

however readings as large as 0.35 percent were observed with rated voltage

applied to V2; and with VI disconnected.

Test results are shown in Tables I and II where the watt converter errors

are computed in terms of full scale range by the following relationship:

Percentage error = reading - (VI x V2 x power factor) applied

^ ^ ^^^

(VI X V2 x power factor) range

At lagging power factor, the voltage applied at V2 lags the voltage applied

at VI.

The estimated uncertainty in the product of the applied voltage (VI and V2)

is 0.2 percent. The power factor is given by cos (e±4>) where e is the nominal phase angle between VI and V2 and is the estimated uncertaninty in this angle ( = 50 x

10"*" radians for the values in Table I and

* = 100 X

10"'' radians for the values in Table II). The lower uncertainity

assigned to the values in Table I is due to the measurement procedure which

allowed the exchange of the signals applied to VI and V2 without amplitude

adjustment.

722/233345-84

87
WATT CONVERTER

ASEA Electric Inc.

- 2 -

To increase resolution, some of the measurements (from Table I) were made on

a power factor range which was lower than the actual applied power factor.

For example on the 5-volt range with 2 volts applied at 0.1 power factor,

readings were taken on the 0.01 power factor switch position. The errors,

however, are reported in terms of the 0.1 power factor range (corrresponding

to the actual power factor). Measurements at zero power factor were made on the 0.01 power factor range and results are reported in terms of the 0.01 power factor range.

For the Director

National Engineering Laboratory

Barry A. Bell^ Group Leader

Electronic Instrumentation and Metrology

Electrosystems Division

Test No: 722/233345-84

Order No: 76625

Date: August 7, 1984
Power Factor

0.00 0.01 0.1 0.5 1.0

lead lag lead lag lead lag lead lag

WATT CONVERTER

ASEA Electric Inc.

- 3

-

Table I

Watt Converter Errors in Percent of Full Scale Range

Voltage Applied

Range Volts

(volts) (nominal

)

VI V2 VI V2

5 5 5 5 0.0 +1.1 -0.9 +0,2 -0.7 -0.6 0.0 0.0 0.0

5 5 4 4 0.0 +0.7 -0.6 +0.2 -0.4 -0.4 0.0 0.0 0.0

5 5 3 3 -0.1 +0.4 -0.4 0.0 -0.3 -0.2 -0.10 -0.10 0.0

5 5 2 2 -0.1 +0.1 -0.2 -0.1 -0.14 -0.13 -0.05 -0.04 -0.09

5 5 11 0.0 0.0 -0.1 0.0 -0.04 -0.03 -0.01 -0.01 -0.02

2.5 2.5 2.5 2.5 +0.9 +0.3 0.0 -0.6 -0.6 -0.7 0.0 0.0 0.0

1.25 1.25 1.25 1.25 +0.3 +0.6 -0.6 -0.2 -0.7 -0.6 0.0 0.0 0.0

Errors with 0.01 percent resolution indicate measurements were made on a power factor range lower than the applied power factor.

Table II

Watt Converter Errors in Percent of Full Scale Range

Voltage Applied Power Factor Range Volts

(volts) (nominal) 0.00 0.01 0.1 0.5 1.0

VI V2

2.5

VI

5

V2

2.5

lead lag lead lag lead lag lead lag

5 -0.1 +0.7 -0.7 -0.2 -0.7 -0.6 0.0 0.0 0.0

2.5 5 2.5 5 +0.3 +0.3 -0.3 -0.6 -0.6 -0.7 0.0 0.0 0.0

5 1.25 5 1.25 -0.2 +0.4 -0.8 -0.4 -0.7 -0.6 0,0 0.0 0.0

1.25 5 1.25 5 -0.4 +0.8 -1.0 0.0 -0.7 -0.6 0.0 0.0 0.0

2.5 1.25 2.5 1.25 +0.5 -0.1 -0.1 -1.0 -0.6 -0.7 0.0 0.0 0.0

1.25 2.5 1.25 2.5 +0.1 +0.4 -0.5 -0.5 -0.6 -0.6 0.0 0.0 0.0

722/233345-84

89
14. APPENDIX 2 — SUGGESTED CALIBRATION PROCEDURE AND ACCURACY GOAL

14 , 1 Calibration Procedure

Recommended specific steps for achieving an accurate and "traceable" test

system are summarized below:

(a) Analyze the existing test system for magnitude and phase errors in

principal components — current and voltage transformers, power

transducers, voltage dividers, etc.

(b) Evaluate the existing test environment in terms of the influences

of electromagnetic fields and electrical noise; shielding and

grounding practices.

(c) Determine the power source waveform purity (harmonic content),

frequency stability, and magnitude stability.

(d) Evaluate system errors at various load loss power factors based

on data obtained in (a), (b), and (c) utilizing technically sound

methodology. (Examples of methodology are presented in this

report)

(e) Establish goals for accuracy and traceability requirements for the

principal components of and for the whole test system.

(f) Construct or acquire calibration (standard) system components

such as those described in this report or equivalent. Determine

their accuracy.

(g) Align principal components in the test system to satisfy the

requirements established in (e). If necessary, modify and retrofit

the test system.

(h) Document the residual errors (magnitude and phase) in each of the

principal components of the test system.

(i) Combine the residual errors of the test system with those of the

calibration system, and thus determine the total test system error

at various power factors. A methodology for combining errors is

outlined in this report.

(j) Conduct system-based verificaton by comparing the results obtained

with the test and standard systems.

(k) Compare the results obtained in (j) against those in (i). If

significant discrepancies are found, investigate and correct the causes of these discrepancies.

(1) In general, strive for independent methods to verify the accuracy of

the measurements.

(m) Finally, implement a continuing calibration or Measurement Assurance

Program, such as the one recommended in this report, to ensure that

the traceability requirements in (e) are met for the critical

standard instruments and the test system.

90
14 .2 Accuracy Goal

In view of the instrumentation and measurement technology that is

available, the authors suggest the accuracy goals listed in Table 21 as challenging but achievable in future test systems where power losses have to

be measured at power factors of 0.01 and larger. In actual use of such a test

system, some degradation in performance may occur due to instabilities that are not caused by the test system itself. Hence, a safety factor of perhaps 2 may have to be applied to the numbers in Table 21 to arrive at realistic,

achievable measurement accuracies.

Table 21

Suggested Goal for Limits of Errors

of Future Test Systems

(a) Errors in principal components

Source of error

Error

Magnitude

(%) Phase Angle

(urad)

Current channel

Voltage channel Power transducer

Current measurement

interaction

0.1 0.1

0.2 0.2

50

5050

RSS 0.32 87

(b) Combined error

Power Factor Error in power measurement

(X)

0.01 0.02

0.05

0.10

0.93

0.54

0.36

0.33

91
NBS-IUA (REV. 2-8C) U.S. DEPT. OF COMM. BIBLIOGRAPHIC DATA

SHEET (See instructions)

1. PUBLICATION OR

REPORT NO,

NBS/TN-1204

2. Performing Organ. Report No. 3. Publication Date

September 1985

4. TITLE AND SUBTITLE

CALIBRATION OF TEST SYSTEMS FOR MEASURING POWER LOSSES OF TRANSFORMERS

5. AUTHOR(S)

Oskars Petersons and S. P. Mehta

6. PERFORMING ORGANIZATION (If joint or other than NBS, see instructions)

National Bureau of Standards

U.S. Department of Commerce

Gaithersburg, MD 20899

7. Contract/Grant No.

8. Type of Report & Period Covered

Final

9. SPONSORING ORGANIZATION NAME AND COMPLETE ADDRESS (Street, City. State, ZIP)

10. SUPPLEMENTARY NOTES

QJ Document describes a computer program; SF-185, FIPS Software Summary, is attached.

11. ABSTRACT (A 200-word or less factual summary of most significant information. If document includes a significant

bi bliography or literature survey, mention it liere)

A calibration systefn for accuracy verification and alignment of test systems for

measuring transformer losses is described. Methodologies are presented for

assessing measurement uncertainties and for evaluating overall accuracy of test

systems. Procedures are suggested for continuing maintenance and calibration of

standard instruments and test systems to ensure traceable measurements.

12. KEY WORDS (Six to twelve entries; alphabetical order; capitalize only proper names; and separate key words by semicolons)

calibration; loss measurement; low power factor; power transformers; traceability

13. AVAILABILITY

^jg Unl I mi ted 03] Po'' Official Distribution. Do Not Release to NTIS ^^ Order From Superintendent of Documents, U.S. Government Printing Office, Washington, D.C.

20402.

CZl Order From National Technical Information Service (NTIS), Springfield, VA. 2216 I

14. NO. OF

PRINTED PAGES

104

15. Price

USCOMM-DC 6043-P80
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