NUREG/CR-4819

ORNL/TM-12038

Vol. 2

Aging and Service Wear of

Solenoid-Operated Valves

Used in Safety Systems of

Nuclear Power Plants

Evaluation of Monitoring Methods

Prepared by

R. C. Kryter

Oak Ridge National Laboratory

Prepared for

U.S. Nuclear Regulatory Commission e
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NUREG/CR-4819

ORNL/TM-12038

Vol. 2

RV

Aging and Service Wear of

Solenoid-Operated Valves

Used in Safety Systems of

Nuclear Power Plants

Evaluation of Monitoring Methods

Manuscript Completed: September 1991

Date Published: July 1992

Prepared by

R. C. Kryter

Oak Ridge National Laboratory

Managed by Martin Marietta Energy Systems, Inc.

Oak Ridge National Laboratory

Oak Ridge, TN 37831-6285

Prepared for

Division of Engineering

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, DC 20555

NRC FIN B0828

Under Contract No. DE-AC05-840R21400
ABSTRACT

Solenoid-operated valves (SOVs) were studied at Oak

Ridge National Laboratory as part of the U.S. Nuclear

Regulatory Commission Nuclear Plant Aging Research

(NPAR) Program. The primary objective of the study was

to identify, evaluate, and recommend methods for

inspection, surveillance, monitoring, and maintenance of

SOVs that can help ensure their operational readiness-that

is, their ability to perform required safety functions under

all anticipated operating conditions-because, under certain

circumstances, failure of one of these small and relatively

inexpensive devices could have serious consequences.

An earlier (Phase I) NPAR Program study described SOV

failure modes and causes and identified measurable

parameters thought to be linked to the progress of

degradations that may ultimately result in functional failure

of the valve. Using this earlier work as a guide, the present

(Phase II) study focuses on devising and then

demonstrating the effectiveness of performance-measuring

techniques and equipment that show promise for detecting

and trending the progress of such degradations before they

reach a critical stage.

Intrusive techniques requiring the addition of magnetic or

acoustic sensors or the application of special test signals

were investigated briefly, but major emphasis was placed

on the examination of condition-indicating techniques that

can be applied with minimal cost and impact on plant

operation. These include monitoring coil mean temperature

remotely by means of coil dc resistance or ac impedance,

determining valve plunger position by means of coil ac

impedance, verifying unrestricted SOV plunger movement

by measuring current and voltage at their critical bistable

(pull-in and dropout) values, and detecting the presence of

shorted turns or insulation breakdown within the solenoid

coil using interrupted-current test methods. The first of

these techniques, though perhaps the simplest conceptually,

will probably benefit the nuclear industry most because

SOVs have a history of failure in service as a result of

unwitting operation at excessive temperatures.

Experimental results are presented that demonstrate the

technical feasibility and practicality of the monitoring

techniques assessed in the study, and recommendations for

further work are provided.

iii
CONTENTS

Abstract.i. ... i

Acknowledgments ........................ ............. ix

Acronyms .xi

1. Background. 1

1.1 Identification of the SOV as a Component for Study. 1

1.2 Research Directions and Ideas Drawn from Preceding Reports. 4

2. Evaluation of Diagnostic Methods Applicable to SOVs .7

2.1 General Considerations; Hardware Studied. 7

2.2 Nonintrusive Measurement of Solenoid-Coil Operating Temperature ..... . . . . . . . . . . . . . . . . 12

2.2.1 Principles of Operation ....... . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 13

2.2.2 Practical Field Application ....... . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 19

2.3 Indication of Valve Position and Change of State ...... . . . . . ........... . . . . . . . . . 22

2.3.1 Static Tests (Nonintrusive to Plant Operations) . . . . . . . . . . . . . . . . . . . . . . . . . . . . . 22

2.3.2 Dynamic Tests (Intrusive to Plant Operations) ....... . . . . . . . . . . . . . . . . . . . . . . 22

2.4 Verification of Unrestricted Plunger Movement ........ . . . . .......... . . . . . . . . . . 26

2.4.1 Sensitivity of Results to Presence of Internal Contamination . . . . . . . . . . . . . . . . . . . . . . 26

2.4.2 Consistency of Results for Clean Valves in Good Operating Condition ..... . . . . . . . . . . . 28

2.5 Indication of Shorted Turns or Insulation Breakdown Within the Coil ...... . . . . . . . . . . . . . . . 31

2.6 Other Methods Examined or Proposed . . . ............. . . . ..... ... . . .. . . . . . . 39

3. Closing Remarks and Recommendations ..... . .... 43

4. Publications Resulting from This Work. . .................... 45

References . . ....... ... .......... 47

APPENDICES

Appendix A: Dc Resistances of Nine Test Valves over the Temperature Range of 20 to 1650C ..... . . . . . . A-I

Appendix B: Ac Impedances of Five Selected Test Valves over the Temperature Range of 20 to 1550C

at Three Widely Separated Frequencies: 5, 60, and 500 Hz ...... . . . . . . . . . . . . . . . . B-i

Appendix C: Ac Inductances (L) and Quality Factors (Q) of Five Selected Test Valves over the

Temperature Range of 20 to 1550C at Three Widely Separated Frequencies: 5, 60, and 500 Hz . . . . C-1

Appendix D: Frequency Dependence of Impedance for dc- and ac-Powered Solenoid Coils .. ....... . . . . D-1

Appendix E: Variation of Coil Resistance and Impedance with Applied Voltage ..... . . . . . . . . . . . . . . E- I

Appendix F: Verification of Absolute Temperature Calibration by Approach-toIsothermal-Condition Tests . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . . Fv
FIGURES

1.1 Exploded view of a typical enclosed-type, direct-acting, three-way SOV ....... . .. .. . . . . . .. 3

2.1 The nine SOVs used in this study, fully assembled .8.. . . . .... .. . .. . . . . . . . . ....... 8

2.2 Photograph of disassembled SOVs "A" and "C" showing interior parts and differing

construction details. 9

2.3 Photograph of disassembled SOVs "D" and "H" .1.0.......... .. ... .. .. .. .. .. ... . lo

2.4 Linear variation of coil dc resistance with temperature for two different solenoid coils ..... . . . . . . 14

2.5 Changes in solenoid coil temperature as inferred from coil dc resistance, brought about by altered

electrical excitation, fluid flow, and environmental conditions .14

2.6 Approximately linear variation of solenoid inductance and quality factor with temperature for a typical

SOV excited at 60 and 500 Hz .16

2.7 Linear variation of coil de resistance and 60-Hz ac impedance with temperature for a typical

ac-powered SOV .17

2.8 Variation of ac impedance with impressed voltage for an SOV held at a constant temperature of 250C,

showing the abrupt impedance change accompanying valve state change .18

2.9 Changes in solenoid coil temperature as inferred from coil ac impedance, brought about by altered

electrical excitation, fluid flow, and environmental conditions .19

2.10 Weekend performance of an ac-powered SOV in a refrigeration system (first 27 h) . . . . . . . . . . . . . 20

2.11(a) Weekend performance of an ac-powered SOV in a refrigeration system (second 27 h) . .. . . .. . . .. 21

2.11(b) Expanded view of a typical 2-h period from Fig. 2.11(a) during which the compressor was cycling . . . . 21

2.12 Dc resistance of SOV from open to closed state ........... .. .. .. .. .. .. .. .. .. .. . 23

2.13 Nonlinear variation of SOV current during a linear rampdown of voltage over a 50-s period ..... . . . 24

2.14 Current and impedance discontinuities during linear voltage ramp-up over a period

of 0.2 s-an indication of plunger movement in a normally operating ac-powered SOV .25

2.15 Lack of impedance change and current discontinuity during linear voltage ramp-up over a period

of 0.2 s-an indication of lack of plunger movement in an ac-powered SOV in which the plunger

was purposely jammed .25

2.16 Instantaneous current through a solenoid coil immediately following application of 60-Hz ac power

compared with rms voltage across the coil calculated by a true rms meter having an averaging time

constant of 7.7 ms. .......................................... 26

2.17 Ac current and voltage characteristics obtained during a slow excitation ramp-up ..... . . . . . . . . . 27

2.18 Ac current and voltage characteristics obtained during a slow excitation rampdown ..... . . . . . . . 27

2.19 Short-term variability of pull-in and dropout voltages and currents for a clean and normally

assembled ac-powered SOV .29

2.20 Short-term variability of pull-in and dropout voltages and currents for a clean and normally

assembled dc-powered SOV .29

2.21 Large "flyback" voltages generated when current to 125-V dc-powered SOV "C" was

interrupted abruptly .............................................. . 33

2.22 Circuit used for measuring transients generated by interruption of coil excitation to a

dc-powered SOV ............................................... . 33

2.23(a) High-voltage transient suppression provided by two varieties of metal-oxide varistors placed

directly across the terminals of SOV "C" .................... ... .. .. ... .. ... . . 35

2.23(b) Time expansion of the early period illustrated in Fig. 2.23(a), showing limitation of peak transient

voltage to -500 V by the 10-A suppressor (upper curve) and to -600 V by the 1-A suppressor

(lower curve) ................................................. 35

2.24 Solenoid coil with no shorted turns, deenergized from 30 Vdc by a wetted-contact relay ..... . . . . . 36

2.25 Solenoid coil having -176 of its -3042 turns shorted, deenergized from 30 Vdc by a

wetted-contact relay ............................................. 36

2.26 Equivalent circuit for an SOV ............ .. .. .. .. .. .. .. .. .. .. .. .. .. .. . 37

2.27 Coil current time waveforms for a 60-Hz ac-powered SOV: loose assembly vs normal ..... . . . . . . 41

2.28 Coil current frequency spectra derived from the waveforms shown in Fig. 2.27 .............. . 41

vi
A.1

A.2

A.3

A.4

A.5

A.6

A.7

A.8

A.9

B.1

B.2

B.3

B.4

B.5

B.6

B.7

B.8

B.9

B.10

B.11

B.12

B.13

B.14

B.15

C.

C.2

C.3

C.4

C.5

C.6

C.7

C.8

C.9

C.10

D.1

D.2

D.3

D.4

D.5

E.1

E.2

Dc resistance vs temperature: SOV "A". . .................... A-2

Dc resistance vs temperature: SOV "B". . . .................. A-2

Dc resistance vs temperature: S OV "C" .. . ... ..... ... ... . .... .. ... . .... .. . . . . A-3

Dc resistance vs temperature: SOV "D". . .................... A-3

Dc resistance vs temperature: SOV ".. ......................... A-4

Dc resistance vs temperature: SOV "F .......................... A-4

Dc resistance vs temperature: SOV "G". . ..................... A-5

Dc resistance vs temperature: SOV "H" .. . A-5

Dc resistance vs temperature: SOV "I". . ..................... A-6

Dc resistance, 5-Hz impedance vs temperature: SOV "A" ........ . ... .. .. ... .. .. .. .. . B-2

Dc resistance, 60-Hz impedance vs temperature: SOV "A" ........ . ... .. .. .. ... .. .. .. . B-2

Dc resistance, 500-Hz impedance vs temperature: SOV "A" ........ .. .. .. .. .. . .. .. .. . . B-3

Dc resistance, 5-Hz impedance vs temperature: SOV "C" ........ . . . . . . . . . . . . . . . . . . . . B-3

Dc resistance, 60-Hz impedance vs temperature: SOV "C" ........ . . . . . . . . . . . . . . . . . . . . B4

Dc resistance, 500-Hz impedance vs temperature: SOV "C" ........ . . . . . . . . . . . . . . . . . . . B4

Dc resistance, 5-Hz impedance vs temperature: SOVY"E" ........ . .. ... .. .. ... .. .. .. . B-5

Dc resistance, 60-Hz impedance vs temperature: SOV "E" ...... .. .. ... .. .. ... .. .. .. . B-5

Dc resistance, 500-Hz impedance vs temperature: SOV "E" ...... . .. .. .. .. . .. .. .. .. . . B-6

Dc resistance, 5-Hz impedance vs temperature: SOV "G" ........ . .. ... .. .. .. ... .. .. . B-6

Dc resistance, 60-Hz impedance vs temperature: SOV "G" ........................... .B-7

Dc resistance, 500-Hz impedance vs temperature: SOV "G" .......................... .B-7

Dc resistance, 5-Hz impedance vs temperature: SOV "I" ............................ .B-8

Dc resistance, 60-Hz impedance vs temperature: SOV "I" ........ . .. ... .. .. .. ... .. .. . B-8

Dc resistance, 500-Hz impedance vs temperature: SOV "I" ........ . . . . . . . . . . . . . . . . . . . B-9

Inductance vs temperature at three excitation frequencies: SOV "A" ....... . ... .. .. .. ... .. . C-2

Inductance vs temperature at three excitation frequencies: SOV "C" ....... . . . . . . . . . . . . . . . . C-2

Inductance vs temperature at three excitation frequencies: SOV "E" ..... .. ... .. .. .. ... .. . C-3

Inductance vs temperature at three excitation frequencies: SOV "G" ...................... .C-3

Inductance vs temperature at three excitation frequencies: SOVY . ..... ... .... ... ... .. . C4

Quality factor vs temperature at three excitation frequencies: SOV "A" ................... .C4

Quality factor vs temperature at three excitation frequencies: SOV "C" ....... . . . . . . . . . . . . . . C-5

Quality factor vs temperature at three excitation frequencies: SOV "E" ....... . .. . .. ... .. .. . C-5

Quality factor vs temperature at three excitation frequencies: SOV "G" ....... . .. . .. ... ... . . C-6

Quality factor vs temperature at three excitation frequencies: SOV "I" ....... . .. . .. ... ... . . C-6

Variation of impedance components with excitation frequency: SOV "A" ....... .. . .. .. .. .. . . D-2

Variation of impedance components with excitation frequency: SOV "D" ................... .D-2

Variation of impedance components with excitation frequency: SOV "G" ................... .D-3

Variation of inductance and impedance with frequency: bare solenoid coils "C" and "I" ...... . . . . . . D-3

Variation of inductance and impedance with frequency: bare solenoid coils "E" and "G" .. . D.A........ 4

Dc resistance vs level of excitation: SOV "D" ....... . ... .. . . .. ... . .. .. .. .. .. . . E-3

Ac impedance vs level of excitation: SOV "D" ......... ... .. . . ... .. . .. ... . .. .. . . E-3

F.1 Comparison of temperature derived from coil resistance to that indicated by a thermometer external

to the coil under (1) essentially equilibrium (electrically powered) and (2) transient (electrically

unpowered) thermal conditions 

F.2 Two more approach-to-isothermal-condition tests for SOV "B" placed in a thermally well

insulated container .

.... F-2

.... F-2

vii
TABLES

1.1 Aging and malfunction of SOVs-Phase I study findings ...... . . . . . . . . . . . . . . . . . . . . 5

1.2 Measurable performance parameters, testing, and surveillance methods identified as potentially

useful by the Phase I study .5............... ... ... ... ... ... ... ... ... .. . 5

2.1 Solenoid-operated valves used in this study .......... .. . .. . .. . .. . .. . .. . . .. . . . 8

2.2 Solenoid-operated valve monitoring methods evaluated in this study . . . . . . . . . . . . . . . . . . . . 11

2.3 Variation of 60-ziz impedance with position of plunger ........ . . . . . . . . . . . . . . . . . . . 23

2.4 Variability of SOV actuation levels (pull-in and dropout voltages and currents) and its relationship

to solenoid design and power source (ac vs dc) ......... . .. . . . . .. . . . . .. . . . .. . . . 30

2.5 Room-temperature electrical properties of faulted and normal bare coils ...... . . . . . . . . . . . . . 39

E. 1 Dc resistance/ac impedance of SOV "" at 230

C . ............... . . ... . ... . .. . . . E-2

E.2 Dc resistance/ac impedance of 850-4 resistor .............................. E-2

viii
ACKNOWLEDGMENTS

The author wishes to acknowledge the contributions of the

following people and organizations, without whose

assistance and cooperation this study could not have been

done:

* John Shank and Frank Fry of Automatic Switch

Company, Florham Park, New Jersey, for helpful

discussions regarding solenoid-operated valve

(SOV) testing techniques, the loan of two new

nuclear-grade SOVs for testing, and the loan of

solenoid coils that had developed shorted turns as a

result of environmental qualification testing at

extreme conditions of temperature and humidity.

* Guy Vacca of Franklin Research Center,

Philadelphia, Pennsylvania, for the loan of two

naturally aged SOVs.

* Jack McElroy of Philadelphia Electric Company,

Conshohocken, Pennsylvania, for helpful

discussions regarding acoustic means for verifying

proper opening and closing of SOVs and the

application of this principle to the monitoring of

boiling-water reactor control-rod-drive scram pilot

valves.

* Dave Eissenberg of Oak Ridge National Laboratory

(ORNL) for encouragement and technical guidance

throughout the study, and for constructive criticism

of the various technical papers generated as a result

of the work.

* Howard Haynes of ORNL for his generous

assistance and loan of equipment necessary to

acquire and process portions of the data, and for his

comradeship throughout the study.

* John Anderson of ORNL for many elucidative

discussions concerning measurement and

interpretation of the electromagnetic properties of

SOV solenoids.

* Bob Shepard of ORNL for the long-term loan of the

high-quality impedance analyzer used to ascertain

the basic electrical properties of the solenoid coils

under study.

* Kathy Sharp, Kathy McIntyre, and Maria Williams

of ORNL for expert typing and editing of the

manuscript.

Finally, the financial support provided by the Division of

Engineering, Office of Nuclear Regulatory Research,

U.S. Nuclear Regulatory Commission (NRC), is gratefully

acknowledged, along with the patience and personal

encouragement of the NRC Technical Monitor, Bill

Fanner, throughout the Phase II investigation.

ix
ACRONYMS

AEOD (Office for) Analysis and Evaluation of Operational Data

ASCO Automatic Switch Company

BWR boiling-water reactor

EQ environmental qualification (test)

IEEE Institute of Electrical and Electronics Engineers

LWR light-water reactor

MOV metal-oxide varistor

NASA National Aeronautics and Space Administration

NPAR Nuclear Plant Aging Research (Program)

NPP nuclear power plant

NRC Nuclear Regulatory Commission

ORNL Oak Ridge National Laboratory

PWR pressurized-water reactor

SOV solenoid-operated valve

xi
Background

1. BACKGROUND

In the context of this report, "aging" is defined as

degradation that occurs with the passage of time. This

degradation is associated with the alteration of physical

properties brought about by the action of environmental

and operational stressors. Aging affects all reactor

structures, systems, and components to some degree and

may potentially increase the risk to the public health and

safety if its effects are not recognized and controlled.

Therefore, to ensure continuous safe operation of a nuclear

power plant as it ages, measures must be taken to monitor

its systems, components, and interfaces in order to detect

the presence of degradation and, if necessary, to restore

integrity through maintenance, repair, or replacement.

Solenoid-operated valves (SOVs) were studied at Oak

Ridge National Laboratory (ORNL) as part of the U.S.

Nuclear Regulatory Commission (NRC) Nuclear Plant

Aging Research (NPAR) Program, which was established

to help resolve technical safety issues related to the aging

of electrical and mechanical components, safety systems,

support systems, and civil structures used in commercial

nuclear power plants. Specifically, the NPAR Program is

designed

* to identify and characterize aging effects that, if

unchecked, could cause degradation of components,

systems, and civil structures and thereby impair

plant safety;

* to identify methods of inspection, surveillance, and

monitoring that will ensure timely detection of

significant aging effects before loss of safety

function and also evaluate residual life of

components, systems, and civil structures; and

* to evaluate the effectiveness of storage,

maintenance, repair, and replacement practices in

mitigating the rate and extent of degradation caused

by aging.

The current Phase II report is intended to satisfy the second

of these three goals.

1.1 Identification of the SOV as a

Component for Study

In accordance with NPAR Program strategy, a component,

system, or structure is identified for study by considering

information from several sources. Criteria used in the

selection process include

* the potential risk from failures of components,

systems, and structures;

* experience obtained from operating plants;

* surveys of expert opinion; and

* user needs (including the resolution of generic

issues, plant performance indicators, and plant

maintenance and surveillance).

Information relevant to the selection of the SOV as a

component meriting study on the basis of these criteria was

developed and is documented in an NPAR Phase I report

authored by Bacanskas et al.1 A short summary of the

findings of the Phase I work is given in Subsect. 1.2, but

prior to addressing the directions provided by the Phase I

study, additional background may be beneficial to the

reader.

An SOV is a valve that is opened and closed by an

electrically actuated solenoid coil that, in most designs, lifts

a plunger to open or close the valve port(s). The process

fluid that is thus controlled is most often instrument air, but

nitrogen or process water may be encountered in some

plant systems. SOVs may be direct-acting (where the

solenoid coil provides the motive force for opening and

closing the valve) or may be pilot-assisted (where the

solenoid coil causes the opening of a pilot orifice, thereby

allowing the process fluid to force the main orifice open).

Only the direct-acting variety was employed in this study.

SOVs are available in a variety of sizes and constructions,

both with and without nuclear qualification, from a number

of different manufacturers. SOVs are found throughout

nuclear power plants in relatively large numbers, often as a

-subcomponent of larger, more complex, and clearly

safety-related systems such as containment isolation valve

actuators, boiling-water-reactor (BWR) control-rod scram

systems, and pressurized-water-reactor (PWR) safety

injection systems.* A listing of safety- and

*Reference I estimates that the population of SOVs used in safety-related

systems at U.S. light-water reactors (LWRs) lies between 1000 and 3000

per plant. BWRs generally employ a greater number of SOVs than

PWRs.

1 NUREG/CR-4819, Vol. 2
Background

non-safety-related systems that use SOVs at U.S.

light-water reactors includes

a reactor building and auxiliary building (ventilation

and isolation);

* BWR scram; 9 main feedwater, and

* reactor coolant pump seal;

* safety injection;

* auxiliary feedwater,

* primary containment isolation;

* high-pressure coolant injection and reactor core

isolation cooling;

* high-pressure injection;

* automatic depressurization;

* emergency diesel generator;

* instrument air,

* chemical volume control, charging and letdown, and

boration;

* pressurizer control;

• steam generator relief (power-operated relief valves

and atmospheric dump valves);

* low-temperature overpressurization protection;

* decay heat removal and residual heat removal;

* component cooling water;

* service water;

* reactor head vent;

* reactor cavity, spent fuel, and fuel handling;

* torus, drywell, vent, and vacuum;

* emergency dc power;

* main steam (main steam isolation valves and

auxiliary boiler);

* condensate.2

They are relatively simple devices (Figure 1.1), with a long

history of satisfactory operation in a variety of both nuclear

and nonnuclear industrial applications. However, their

presence in systems important to safety requires an

especially high degree of assurance that they are ready to

perform their required function under all anticipated

operating conditions, because failure of one of these small

and relatively inexpensive devices could have serious

consequences under certain circumstances. 3

For this reason the NRC Office for Analysis and

Evaluation of Operational Data (AEOD) has had a

continuing interest in SOV performance and failure

characteristics and mechanisms and has documented its

assessment of the vulnerability of safety-related equipment

to common-mode failures or degradations of SOVs in a

recently published operating experience feedback report.2

This study cites over 20 representative operational events

in which SOV failures or degradations affected or had the

potential to affect multiple safety systems or multiple trains

of individual safety systems. Although such common-mode

SOV failures and degradations are often beyond the

conditions analyzed in plant final safety analysis reports

and are not ordinarily modeled explicitly in present-day

probabilistic risk assessments, operating experience

indicates that such failures and degradations have indeed

compromised front-line safety systems in the past2 and will

probably continue to do so in the future. Events in which

safety systems have been adversely affected by

degradations or failures of SOVs are considered

"legitimate precursors to more significant events" by

AEOD. The operating experience feedback report notes

that SOV problems permeate almost all U.S. nuclear power

plants and that they encompass many aspects of the SOV's

design, maintenance, and operation. AEOD analysis of

operating data indicates that the underlying cause of many

SOV failures is the licensees' lack of information or

understanding of SOV requirements and capabilities.

Compounding the problem is the fact that some SOV

manufacturers fail to provide users with adequate guidance

regarding proper SOV selection and maintenance.2

Moreover, most plant technical specifications do not

require periodic testing of SOVs per se, although the

valves' performances may come into play in the regular

exercise of systems and devices covered by technical

specification testing requirements.

NUREG/CR-4819, Vol. 2 2
Backgiound

ER GASo

ORNL-DWG 86-5315 ETD

SCREWS (31 AND

ATTACHED.

'ET

~>)-SPRING WASHER

SOLENOID BASE

SUB-ASSEMBLY

UPPER SOLENOID

BASE GASKET

HOUSING

RETAINER GASKETS

IL

(OMITTED WHEN MOLDED

COIL IS USED)

ASSEMBLY

GASKET

SPRING

-END CAP

Figure 1.1 Exploded view of a typical enclosed-type, direct-acting, three-way SOV, illustrating relationships of

internal parts and positions of the solenoid coil and elastomeric components (i.e., degradation sites). Source:

Reprinted with permission from Automatic Switch Company, Bulletin 8320, Florham Park, New Jersey, 1978.

3 NUREG/CR-4819, Vol. 2
Background

The AEOD report concludes that" ... SOV problems

represent a significant safety concern," and that " . . . the

SOV problems outlined in this report need to be addressed

to ensure that the margins of safety for U.S. LWRs remain

at the levels perceived during original plant licensing.

Generic and plant-specific actions are needed to correct the

SOV problems in order to restore the plants' safety

margins to their original perceived values."4

The findings and conclusions summarized above-plus

continuing occurrences at plants resulting in the issuance

over the past few years of some 36 NRC communications

alerting licensees to generic problems with SOVs3-justify

the inclusion of SOVs in the list of components to be

studied by the NPAR Program.

1.2 Research Directions and Ideas

Drawn from Preceding Reports

The previously referenced Phase I NPAR study described

proximate and root causes of SOV failure or malfunction,

recognized SOV-aging mechanisms, and listed principal

SOV degradation sites (Table 1.1, Figure 1.1).' The study

also identified measurable indicators of performance

(Table 1.2) as well as nonquantitative testing and

surveillance methods that might be used to detect the

presence of or monitor the progression of various

degradations that could ultimately lead to functional failure

of the valve.* Those findings and suggestions provided

valuable insight and direction for the ensuing Phase II

studies reported here. Listed here is a summary of the

conclusions of the Phase I study:

* The reported incidence of SOV failure in

safety-related applications has been reasonably low.

* The most prevalent causes of SOV failure have been

open-circuited coils, short-circuited coils, worn or

degraded mechanical parts, and contamination by

foreign materials.

* Review of technical literature did not reveal any

degradation-monitoring techniques, either in use or

under development, that are oriented specifically to

SOVs. However, several potentially useful methods

were identified.

*The findings summarized in Tables 1.1 and 1.2 are reinforced by similar

conclusions reported by an Institute of Electrical and Electronics

Engineers (IEEE) task group composed of representatives from utility,

consulting, and manufacturing sectors of the nuclear power industry.5

* SOV failures caused by air system contamination

and high dc voltages result from stressors that need

to be addressed at the system level rather than at the

component (i.e., SOV) level.

* Emphasis should be placed on development of

degradation-monitoring techniques for the solenoid

coil.

a Considering the large number of SOVs used in

nuclear power plants and the low incidence of

failure, degradation monitoring on a sampling basis

should be considered.

a As required to maintain environmental qualification

of safety-related SOVs, periodic replacement of

elastomers and solenoid coils also aids in

minimizing the incidence of SOV failure.

a Because of the importance of SOVs in controlling

equipment such as control-rod drives (in BWRs) and

isolation valves (in BWRs and PWRs), SOV failures

can initiate unwanted challenges to plant safety

systems and can be responsible for unanticipated

transients. Development of a methodology for

detection of incipient failures will reduce the

frequency of these challenges and thereby enhance

plant safety and operability.

The second, fourth, and fifth conclusions provided the

driving force and direction of the present research, while

the third conclusion dictated that the Phase II study would

probably be inventive or developmental in nature rather

than evaluative.

Specific recommendations of the Bacanskas report were to

' . . . [conduct] research and development to devise test

methods suitable for evaluation of the electrical and

mechanical portions of SOVs" and to " .. . test the

proposed monitoring techniques with new and used SOVs

to evaluate the suitability, cost effectiveness, and

practicality of the techniques."' Within the budgetary and

sponsor-imposed limitations of this segment of the NPAR

Program, we have attempted to carry out these two

recommendations.*

*Several naturally aged SOVs were obtained, with the help of Pacific

Northwest Laboratory, from the now-decommissioned Shippingport

Atomic Power Station (operational from 1957 to 1982), but they were

found to be radioactively contaminated to a degree that made their use

in this study impractical.

NUREG/CR-4819, Vol. 2 4
Background

Table 1.1. Aging and malfunction of SOVs-Phase I study findings

Topics of study Findings

Proximate causes of malfunction

Root causes of malfunction

Degradation sites

Electrical failure (open- or short-circuited coil)

Mechanical failure (valve will not change state)

Mechanical binding (sluggish operation)

Prolonged operation at excessive temperatures

High-voltage transients produced by solenoid turnoff

Improper choice of valve for operating environment

Improper valve installation

Lack of maintenance; incorrect replacement of parts

Chemical attack of elastomers by oil

Contamination by dirt, thread sealant, polymerized lubricants

Elastomeric components (core seat, seals)

Solenoid coil (insulation)

Core spring

Sliding surfaces

Aging mechanisms Changes in mechanical and electrical properties of materials

Conductor burnout

Source: NUREG/CR-4819, Vol. 1.

Table 1.2. Measurable performance parameters, testing, and surveillance

methods identified as potentially useful by the Phase I study

Type Performance indicators

Quantitative measures

Nonquantitative indicators

Coil resistance

Coil power consumption

SOV temperature

Valve actuation time

Valve flow coefficient, Cv

Internal (through-valve) leakage rate

Relative level of hum or chattering

Visual check:

* General appearance of valve exterior

* No evidence of accumulated moisture, burned paint, cracked or frayed wiring

Aural check.

* Absence of buzzing or rattling while in energized state

Operational check:

* For rapid, smooth change of state

* Should be performed at extremes of rated valve pressure range

Inspection of internal components:

. Not recommended on a periodic basis

* Necessary to establish replacement intervals for limited-life components

Source: NUREO/CR-4819, Vol. 1.

5 NUREG/CR-4819, Vol. 2
Evaluation of Diagnostic Methods Applicable to SOVs

2. EVALUATION OF DIAGNOSTIC METHODS APPLICABLE TO SOVs

2.1 General Considerations; Hardware

Studied

Any proposal for implementation of surveillance and

diagnostics in nuclear power plants must address the issues

of practicality and cost-effectiveness. For example, it must

be recognized that the SOVs installed in present-day

nuclear power plants are uninstrumented and that

backfitting them with instrumentation would probably be

quite expensive. It is also necessary to understand that

many SOVs important to plant safety are inaccessible

during plant operation and that some requiring verification

of operational readiness will change state only rarely

during normal operations, thereby offering little

opportunity for measuring SOV dynamic performance

parameters. Because of these circumstances, many utilities

have elected to periodically replace degradable components

or subassemblies within SOVs that are embodied in

safety-related systems-that is, to use preventive

maintenance-rather than to attempt to practice predictive

maintenance using one or more of the approaches listed in

Table 1.2.1 In the extreme, some utilities replace the entire

SOV as a precautionary measure at predetermined intervals

based on lifetime expectations derived from environmental

qualification (EQ) tests, even though no malfunction has

been observed. 1

'

2

In carrying out these Phase II studies, we chose to view the

practical implementation considerations discussed

previously as challenges rather than insurmountable

obstacles. Therefore, while acknowledging that

implementation of diagnostic capabilities must be

cost-effective relative to the alternative of SOV

replacement, we continued to search for techniques and

equipment with which to measure some of the performance

parameters identified in the Phase I study and thereby

detect and trend the progress of any degradation that might

eventually compromise the ability of an SOV to perform its

intended function. In recognition of the cost-effectiveness

issue, attention was focused on remotely applied,

completely nonintrusive techniques (i.e., ones that do not

require physical access to the SOV, the addition of sensors

or signal wires, the removal of power to the solenoid, or

application of a special test signal). However, we found it

necessary to depart from these restrictions in order to be

able to detect some well-recognized modes of SOV

degradation. Clearly, tradeoffs must usually be made

between the degree of disruptiveness to plant operations

and the amount of information obtained, and this aspect of

the surveillance and diagnostic methods investigated here

is in need of further attention.

Nine small enclosed-type SOVs from various

manufacturers were obtained for use in the study (see

Table 2.1 and Figures 2.1-2.3). Three were nuclear-grade

valves; one of these ("C) was used, and the others were

new. All of the remaining six valves except one ("I") were

new. The valve pressure ratings (60 to 2200 psi) are

indicative of valve wall thickness and material of

construction (i.e., brass or stainless steel), while electrical

ratings (115 to 120 Vac or 125 Vdc) are established by

solenoid coil construction (i.e., wire gage and number of

turns). The nominal power consumption for all nine valves

was in the range of 10 to 20 W at rated voltage.

The study began with measurement of every conceivably

useful electrical property of an SOV because, in many

nuclear plant applications, the solenoid lead wires provide

the only available link to the outside world where

measurements can be performed. Full details of these

fundamental electrical measurements, along with the

conclusions and possibilities for monitoring schemes that

resulted from thoughtful consideration of each measurable

property, are documented in Appendices A through E.

However, for the sake of brevity, only a few representative

examples are extracted for presentation in the main body of

the report, with occasional reference to the appendices for

additional detail.

The surveillance or diagnostic techniques studied

(summarized in Table 2.2 and described in the following

five subsections) were a natural outgrowth of the

investigations of fundamental electrical properties of SOVs

just described. Each technique was proposed to be

specifically responsive to one or more of the proximate or

root causes of SOV failure or malfunction cited in

Table 1.1. The strengths and weaknesses of each technique

as a means for detecting or tracking the progression of

age-related degradation were then evaluated. The four

techniques receiving the most attention were

measurement of coil mean temperature during

operation by means of in situ measurement of coil

electrical dc resistance or ac impedance, combined

with an experimentally established temperature

coefficient of resistivity for the copper winding (i.e.,

the same principle as a resistance thermometer);

7 NUREG/CR4819, Vol. 2
Evaluation of Diagnostic Methods Applicable to SOVs

Table 2.1. Solenoid-operated valves used in this study

Designation Manufacturer valve Design power Comments identification

A ASCONP8320A 185VSINS14 125 Vdc Nucleargrade, rated for 115 psi

B ASCO NP8320A 185V SIN S 15 125 Vdc Nuclear grade, rated for 11 5 psi

C ASCO NP8314C29E S/N K-62 125 Vdc Used; nuclear grade; rated for 60 psi air

D Skinner V5H30650 CTN 120 Vac Rated for 150 psi

E Skinner V5H38880 120 Vac Rated for 100 psi

F Skinner (no type number available) 120 Vac

G ASCO 8262A214 S/N S94804 115 Vac Rated for 2200 psi

H ASCO 8210B26 SIN 91634S 115 Vac Rated for 350 psi air

I ASCO HTX831429 S/N 2445A 125 Vdc Used; rated for 70 psi air

3⁄4 01

I

',100

Figure 2.1 The nine SOVs used in this study, fully assembled. (See Table 2.1 for manufacturer and type.)

o indication of plunger movement upon application of

ac power, plunger static position within the solenoid

coil, and freedom of plunger movement within the

guide tube by means of measurement of coil ac

impedance (inductance) changes resulting from

movement of the iron plunger relative to the coil;

* indication of mechanical binding by tracking

changes in the electric current (or voltage) at the

SOV critical bistable (pull-in and dropout) points,

because these define conditions of balance between

electrical (magnetic), spring (restoring), and friction

forces within the SOV assembly;

NUREG/CR-4819, Vol. 2 8
Evaluation of Diagnostic Methods Applicable to SOVs

N

I

I - X I

I % . .; i .. 4

i r CEN tcE yj

,I 0 

C-I

-w

Id-l

*I

Figure 2.2 Photograph of disassembled SOVs "A" and "C" showing interior parts and differing construction

details (Both are three-way valves, but the placement of the third port is different.)

9 NUREG/CR-4819, Vol. 2
Evaluation of Diagnostic Methods Applicable to SOVs

I

*1

CENTIMETERS

4 iii I ii £ ii

U

,3

end

i

w w

1'

I A

413B

Q!]M

. ,,,

Figure 23 Photograph of disassembled SOVs "D" and "H." (These are three-way and two-way valves,

respectively.)

NUREG/CR-4819, Vol. 2 10
Table 2.2. Solenoid-operated valve monitoring methods evaluated in this study

Method Degracheron(i or nnslfunctad (s ) Strengths Weaknesses Promise for in-plant use

Measurement of SOV temperature via

coil resistance or impedance

Indication of valve position and change

of state upon application of power,

via change in coil impedance

Indication of mechanical binding by

tracking changes in current/voltage at

SOV pull-in and dropout points

Indication of shorted coil turns or

insulation breakdown based on

characteristics of the electrical

transient generated upon deenergizing

a dc SOV

Indication of mechanical binding by

analyzing the time-varying

characteristics of the inrush current

accompanying application of

electrical power to the SOV

Indication of mechanical looseness

within ac-powered valves via

electrical detection of humming or

chattering of the plunger assembly

(frequency decomposition of

steady-state coil current)

_ . .. . . . . . .

Electrical failure of coil and degradation

of elastomers resulting from

prolonged operation at excessively

high temperatures

Mechanical binding, sluggishness, or

failure to shift as a result of worn or

improper parts or the presence of

foreign materials inside the valve

Mechanical binding and sluggish

shifting caused by worn, swollen, or

improper parts or the presence of

foreign materials inside the valve

Electrical failure of solenoid coil, caused

by high-voltage turn-off transients in

combination with insulation

weakened by prolonged operation at

high temperatures

Mechanical binding and sluggish

shifting caused by worn, swollen, or

improper parts or the presence of

foreign materials inside the valve

Wear of internal valve parts, improper

assembly, or replacement with

incorrect parts

* Nonperturbative to plant operations

* No new sensors or signal cables are

required

* No pemanent instrumentation

required; can be applied as

needed from a remote location

* Applicable to ac- and dcpowered SOVs

* No need for add-on sensors or signal

cables

* Valve position readout is obtained

from a remote location

* Static method does not disturb SOV

* Simultaneously detects degradation

of magnetic and spring forces, increase

in frictional forces

* No need for add-on sensors or cables or

access to SOV

* Applicable to ac- and dc-powered

SOVs

* Detects presence of defects within coil

that cannot be revealed by other means

* No need for add-on sensors, signal

cables, or access to SOV

* Information could be obtained as a

result of everyday valve operation

* No need for add-on sensors, signal

cables, or access to SOV

* Nonperturbative to plant operations

* Measures mean coil temperature, not High; ready for immediate use

coil hot spot or valve body temperature

Dynamic method requires introduction High; some additional developmental

of special ramp-voltage power supply work required

Applicable only to ac-powered valves

* Requires introduction of a special

variable-voltage power supply

* Adequate sensitivity to degraded valve

condition is presently undemonstrated

* Requires access to coil leads at the

SOV and removal of normal

valve-control electrical cable

* Installation of transient suppression

devices effectively removes threat of

insulation puncture

* Apparent lack of sensitivity to valve

degradations

* Short stroke time of small SOVs

creates instrumentation problems for

ac-powered valves

* Apparent lack of sensitivity to valve

degradations

* Requires sophisticated signal analysis

equipment

Medium: further testing needed to

ascertain cause of poor repeatability

of test results

Low useful for laboratory postmortem

tests

Minimal; investigation of method

abandoned early in the study

Minimal; investigation of method

abandoned early in the study.

Addition of miniature acoustic sensor

to SOV might prove worthwhile

ti1

0

0

01

CC

z

00

Ž0

0

of
Evaluation of Diagnostic Methods Applicable to SOVs

* indication of shorted turns and/or insulation

breakdown within the solenoid coil based on the

characteristics of the "flyback" transient generated

when a dc SOV is deenergized abruptly.

Two other techniques were examined briefly and then

abandoned because initial results showed little promise for

eventual development into effective degradation detection

and diagnosis tools. These were

* indication of mechanical binding by analysis of the

detailed time "signature" of inrushing current that

accompanies initial plunger movement upon

application of electrical power to the SOV; and

* indication of mechanical looseness (resulting from

wear of internal valve parts, improper assembly, or

replacement with incorrect internal parts) by

electrical detection of the presence of humming or

wear-producing chattering of the plunger assembly

within ac-powered SOVs, using frequency-domain

analysis of the coil current signal.

The operational principles, selected test results, and

perceived merits and shortcomings of each of these

techniques are described in the following subsections. For

most techniques, additional details are available in the

appendices.

2.2 Nonintrusive Measurement of

Solenoid-Coil Operating

Temperature

[Malfunctions addressed: electrical failure of coil and

degradation of elastomers, caused by prolonged operation

at excessive temperatures]

As a result of the strong temperature dependence of

chemical reactions (Arrhenius theory6) that are constantly

in progress and that tend to break down organic

compounds, excessive operating temperature can be

expected to shorten the service life of both the solenoid coil

insulation and critical elastomeric parts (0-ring seals and

valve seats, in particular) within the valve. Arrhenius

theory is widely used in extrapolating results obtained in

equipment qualification tests performed to industry

standards7 '

8 to more realistic (as well as abnormal) plant

conditions in order to obtain an estimate of qualified

service life for the component. To be meaningful, the

Arrhenius calculations must employ correct time and

temperature profiles, especially the temperatures actually

encountered during plant operation. Measurement of true

coil temperature using the method described here may help

to improve the accuracy of such qualified service life

predictions.

Undesirably high operating temperatures may be caused by

a number of circumstances such as an application of

higher-than-normal voltage to the solenoid coil (as when

station batteries are being charged), restricted air flow or an

elevated ambient air temperature at the valve location,

elevated temperature of the fluid being controlled by the

valve, or an insulation breakdown within the solenoid coil

winding (resulting in shorted turns, hence lowered

electrical resistance and increased power consumption).

Therefore, one major cause of premature valve failure

might be vastly curtailed if a simple, reliable, and

inexpensive means were available to monitor (at will) the

actual operating temperatures* of critical SOVs that may

be inaccessible and thereby unmonitorable by traditional

methods such as infrared pyrometry.

The test results presented below indicate that a promising

means of fulfilling this need is to use the copper winding of

the solenoid as a self-indicating, permanently available

resistance thermometer. To do so requires only in situ

measurement of coil dc resistance (or ac impedance)

combined with prior knowledge of the temperature

coefficient of resistivity for the copper winding and the coil

resistance (or impedance) at a single known temperature

(most conveniently, normal room temperature).** The dc

resistance or ac impedance may be obtained

nonperturbatively by measuring voltage applied to and

current drawn by the solenoid coil, the former by a

voltmeter and the latter by a clamp-on current transformer

or a Hall-effect probe (neither of which requires

disturbance of the valve control circuitry). The resistance

of the electrical leads connecting the SOV to the power

source (and the possible variation of this lead resistance

Monitoring actual SOV operating temperature is, of course, only one

approach to increasing life; others are to increase heat dissipation

capability (e.g., by the addition of cooling fins) or to curtail heat

generation within the solenoid. Devices are available to automatically

reduce solenoid operating voltage to a "holding" level (-70 V) once the

solenoid has achieved pull-in on full voltage (125 V), the reduction

taking place -5 s after valve actuation. One manufacturer of such

devices is Enertech Sollex of Brea, California the availability of

nuclear-qualified units is unknown to the author.

**The temperature inferred by this method will be an average value for

the volume occupied by the copper winding rather than for the hottest

spot within the solenoid coil. However, unless the coil has a localized

fault (such as a shorted turn), this difference may not be very large:

measurements by Bacanskas et al. indicate that the 'hot spot increment'

is in the range of 3 to 5'C.9

NUREG/CR-4819, Vol. 2 12
Evaluation of Diagnostic Methods Applicable to SOVs

with changes in ambient temperature) is normally so small

relative to the resistance of the solenoid coil that it can be

ignored, but it could be measured separately and subtracted

from the field-acquired data if such correction is thought to

be warranted.

2.2.1 Principles of Operation

To serve as an accurate and a remote sensor of local

temperature, some electrical property of the copper

solenoid coil must be demonstrated to have a stable

dependence on temperature. Though not essential, it is

desirable that the relationship of the electrical property and

temperature be linear. It is also desirable that the

relationship not be affected by changes in other conditions

that may not be controllable in a real plant environment. As

the data of Appendices A, B, and C illustrate, several

candidate parameters exist, each having advantages and

disadvantages both in theory and in practice. Because

SOVs are designed for two fundamentally distinct

electrical environments, alternating current (ac) and direct

current (dc), the two types are treated separately below.

2.2.1.1 Dc-Powered SOVs

Coil de resistance is an easily measured quantity satisfying

the given criteria. Similar to platinum (widely used as a

resistance thermometer), copper has a stable and

sufficiently large temperature coefficient of resistance

(-0.2 to 0.3% of value per 'C, depending on copper

purity10) to permit its use as the resistance element of a

practical industrial thermometer. Figure 2.4 shows the

resistance-temperature relationships obtained for two SOV

coils having quite different dc resistances as a result of

differences in wire size and in the number of turns

employed in their construction. These data, which are

typical of results measured on nine separate valves

designed for both ac and dc operation and produced by

different manufacturers (see Appendix A), were obtained

by placing all nine valves in a thermostatically controlled

oven and measuring their dc resistances at selected

elevated temperatures. For each of the nine valves an

extremely linear relationship (correlation coefficient

0.9997) was obtained over the temperature range of interest

(20 to 170'C). It must be stressed that, regardless of the

actual numerical value of the coil resistance, the

temperature coefficient of resistance (expressed as a

percentage of value) is sufficiently large (0.3% per 0C) to

permit temperature measurement with better than ±100 C

accuracy using resistance measurement equipment of only

modest (2 to 3%) accuracy. Temperature measurement

accuracy of this order is surely adequate for indicating coil

temperatures that exceed accepted operating limits

established by qualification tests7 ' 8 or service life

prediction curves based on Arrhenius reaction rate theory.

In this context it should also be recognized that the trend of

indicated coil temperature over a time span of months may

be as important to assessing the impact of operating

environment on SOV condition and operability as a highly

accurate measurement of the coil's temperature would be.

Figure 2.5 provides a laboratory demonstration of this

technique. When the 125-Vdc SOV was first energized at

the start of the test (when ambient temperature was known

to be 250C), a coil resistance of 793.9 n was established by

means of Ohm's law from applied electrical potential and

current readings. This single calibration point, in

conjunction with the empirically determined slope of the

resistance-vs-temperature relationship for the copper wire

used in this solenoid (3.41 WPC), allows establishment of a

temperature scale for the right-hand ordinate of the graph

that exactly matches the directly measured resistance scale

on the left-hand ordinate; namely,

T(°C) = 34Q) - 793 +25.0*(1

3.41 K2/ 0 C(1

Once established, this linear scaling relationship permits

direct interpretation of changes in SOV coil resistance

accompanying altered test conditions in terms of their

temperature equivalents. (Such dual scales are used in

figures throughout this section as a reminder that resistance

or impedance is the quantity directly measured.)

Placed inside a 2-ft3 open-ended enclosure with only

natural convection for cooling and with no instrument air

flowing through it, the SOV was first allowed to approach

thermal equilibrium at a dc excitation voltage of 117 V.

The coil temperature inferred from periodic measurement

of dc resistance (Figure 2.5) reached about 113'C (as a

result of the deposition of about 14 W of resistive heat

within the coil) 65 min after initial power-up, whereas the

temperature indicated at this time by a mercury-in-glass

stem thermometer placed in contact with the periphery of

the potted solenoid coil was only -70'C (Figure 2.5). This

rather large (>40'C) difference is not indicative of a

calibration problem but, as has been shown by data

acquired under more nearly isothermal conditions (see

Appendix F), results from a combination of high thermal

resistance at the thermometer-coil contact point and the

existence of a large temperature gradient between the

center and the outer periphery of the coil, which is

encapsulated in a material having poor thermal

conductivity. The point here is that, despite the

13 NUREG/CR-4819. Vol. 2
Evaluation of Diagnostic Methods Applicable to SOVs

a

0

z

F

C.)

C3

1300

1200

1100

1000

900

800

700

2300

2100

C

1 900 F- z

co

1700 wL

0

C)

1500

1300

20 40 60 80 100 120 140 160

COIL TEMPERATURE (C)

Figure 2.4 Linear variation of coil dc resistance with temperature for two different solenoid coils.

1200 | ~ . . ., . . . . . . . . . . . . . I . . .aI no 117 V still ir, no Excitation

fluid

1150 <

a

w

0

z

fC/31

Cr

0

a

-j

0

1 100

1050 

1000 C

0

flow thru SOV increased 0 00

- - -- 5 lo 134 V 000

00 ~~ . 00 00~~~~~~~0

0 

0

0

: : : ++

: : : + +

+ ~ ~ .4 ~ ~ 

+ ~~~~Process ir turned off;

other conditions sonne

-00

Forced air

oling initiated

4 . *

+ +

+

00 0

/ dB

0 : lb

0e,

0

140

120 0

100 )

I--

ItU

80 aL

60 93

z

20

950 0

0

+

4

+

+

44444 +

900 

850 F 0 4

800 4 _! ~r

Process air allowed to

flow through open SOV

o ,es;stomce-deeiaed

tempeotues

+ thermometer-dersed

temperatures

I I . I I

750

0 40 80 120 160 200

ELAPSED TIME (minutes)

Figure 2.5 Changes in solenoid coil temperature as inferred from coil dc resistance, brought about by altered

electrical excitation, fluid flow, and environmental conditions.

NUREG/CR-4819, Vol. 2 14
Evaluation of Diagnostic Methods Applicable to SOVs

considerable differences in the absolute temperatures

indicated by the coil resistance and the mercury bulb, the

two curves in Figure 2.5 clearly track each other in detail

throughout the 220-min duration of the test.

At = 65 minutes instrument air was allowed to flow

through the already-open SOV. The cooling effect of

room-temperature air flowing through the brass valve body

is evident in both curves-somewhat more so for the

mercury bulb than for the coil, which is understandable

since the bulb was positioned closer to the valve body than

is the bulk of the coil. At t = 85 minutes the dc supply

voltage driving the solenoid was increased to 134 V to

simulate a condition that might be encountered on a

nominal 125-Vdc power bus during times when station

batteries are being charged. Note that the resulting

temperature increase was immediate in the coil but slower

to appear at the mercury bulb; this phenomenon results

because the additional heat produced by the increased

excitation voltage is generated instantaneously within the

wire of the coil, whereas it must be transported to the coil

periphery (where the mercury thermometer was located) by

the relatively slow process of thermal conduction.

The 20-psi instrument air was turned off at t = 100 min, but

the SOV remained powered at 134 Vdc and cooled only by

natural convection within its enclosure. Over the next

70 min the indicated coil temperature rose to an asymptotic

value of about 135oC,* and a commensurate rise was

recorded on the mercury bulb thermometer. At t = 170 min

a muffin fan positioned below the enclosure surrounding

the SOV was turned on to draw air down past the valve at

low velocity. The effect of this forced-air cooling was seen

to be a prompt reduction of both indicated

temperatures-somewhat more rapid for the bulb

thermometer than for the coil, which is explainable by the

same reasoning offered for the difference in time response

of the two curves when the electrical excitation was

increased at I = 70 min (i.e., time lag due to conductive heat

transport). However, the effect on the curve is in the

opposite direction because, in the experiment, heat was

being removed from the outside of the coil rather than

added from the inside.

*Even this relatively high temperature, representing an increase relative to

ambient of about I 10'C, is still well below what is considered

acceptable for the Class H coil insulation that has been wed in recent

years in nuclear-grade SOVs. (Class H insulation is rated for continuous

operation at 1 60'C; the newer Cass N is rated for 175'C.) However, had

the ambient been 50C-a condition that might be encountered in some

areas within containmentClass 11 insulation would be at its operational

lirnit.

In summary, the results of the test described previously and

depicted in Figure 2.5 illustrate that using the copper

winding of a dc-powered SOV as a resistance thermometer

can provide meaningful real-time indication of altered

excitation, environmental, and fluid flow conditions likely

to occur in power plants from time to time that could result

in unacceptably high SOV operating temperatures and

therefore in shortened service life. Moreover, the

temperatures so obtained are better indicators of

temperature stress imposed on the coil materials than might

be obtained from conventional, externally applied

temperature-indicating devices such as thermocouples or

thermistors.

2.2.1.2 Ac-Powered SOVs

Ac-powered valves offer many more electrical

measurement possibilities, some of which may be useful

for the inference of coil temperature. These include

inductance, quality factor, and impedance (expressed either

as a vector magnitude or in terms of its real and imaginary

orthogonal components). However, some measurable

parameters prove more suitable than others: take for

example inductance and quality factor,* whose variations

with temperature are shown in Figure 2.6 for a typical ac

SOV excited at two widely separated frequencies, 60 and

500 Hz. (See Appendix C for additional data.) As a

measure of temperature, quality factor is easily dismissed,

owing to the poor quality of the correlation at 60 Hz (as

evidenced by the considerable data scatter) and almost zero

slope at 500-Hz excitation frequency. Inductance is not so

easily dismissed. its relationship to temperature is quite

linear and its temperature coefficient (somewhat less than

0.1% of value per IC), although not large, is sufficient to be

useful, especially at 500-Hz excitation frequency.

However, application of a 500-Hz test signal for the

purpose of temperature measurement would require

removal of normal of 60-Hz power from the SOV, thereby

making the measurement perturbative. Moreover, the

measurement of inductance requires fairly sophisticated

electronic circuitry. For all these reasons, temperature

inference via inductance and quality factor was not pursued

further, instead, attention was turned to the obvious analog

of dc resistance: ac impedance.

*Quality factor-often designated Q in electrical engineering exts-is a

dimensionless attribute of an inductor or a capacitor that quantifies its

ability to store power (as inductance or capacitance) as opposed to its

ability to dissipate power (as resistance); that is, it is the ratio (volt-amps

stored)/(watts dissipated) = reactance/resistance.1 Therefore, the greater

Q is, the more nearly the inductor or capacitor approaches an ideal,

purely reactive (zero-resistance) device.

15 NUREG/CR-4819, Vol. 2
Evaluation of Diagnostic Methods Applicable to SOVs

C Inductance

at 60 Hz

0 at

80 Hz

1 Inductance 0 d at

at 500 Hz W50Hz

900

850 

800 F

I

E

"C

-

1.,

0

z

750

700 

n5+U'C (0.O5%' aai

- *

siop~~O5 5 rr~~J' (Oi8%' at 2

2.0

1.8

a

1.6 -

C.)

U1.4 r

"C

a

1.2

1.0

650 I

600 F

550 

500 F

450 . . . . . . . . . . . . . . .

20 40 60 80 100 120 l140 160

COIL TEMPERATURE (C)

Figure 2.6 Approximately linear variation of solenoid inductance and quality factor with temperature for a typical

SOV excited at 60 and 500 Hz. (Data for other coils and excitation frequencies are given in Appendix C.)

By analogy to the measurement procedure previously

illustrated for a dc-powered SOV, the inference of

operating temperature for an ac-powered SOV would

follow the same path except that impedance (measured at

the power line frequency of 60 Hz in the United States but

often 50 Hz abroad) would replace resistance as the

quantity obtained by applying Ohm's Law to the

root-mean-square (rms) voltage and current measured at the

SOV's electrical leads. However, two complicating factors

arise that were not troublesome when dealing with dc

valves. While neither poses an insurmountable obstacle to

temperature inference, these troublesome factors need to be

appreciated and so will be described in some detail in the

following two subsections.

Diminished Temperature Coefficient

Figure 2.7 illustrates the manner in which the magnitude of

a typical solenoid's electrical impedance, I Z |, varies with

temperature for a power line frequency of 60 Hz. (Data for

five of the nine valves at three different frequencies are

given in Appendix B for the interested reader.) The

orthogonal components of the vector Z-Re(Z) and

Im(Z)-are also shown, along with the dc resistance, R, for

comparison. Note that, like R, each of the representations of

ac impedance shows a highly linear variation with

temperature and that, with the exception of Im(Z), they all

display approximately the same absolute slope, expressed

in the figure as ohmslC. The linear relationship is, of

course, highly desirable, and shows that in principle any of

these quantities might be used as a measure of temperature.

However, because the absolute magnitudes of the measured

quantities increase substantially as one moves from dc

resistance through Re(Z) and Im(Z) to I Z I while retaining

essentially the same slope, the temperature coefficients for

these four measurable quantities, expressed as percentage

of value, decrease markedly as one moves from the bottom

to the top of the graph. The practical implication of this

observation is that because the temperature coefficient of

I Z I (0.118%P/C) is only about 30% as large as the

temperature coefficient of R (0.3%%/ 0C), the accuracy of

I Z I measurements would have to be three times greater

than the accuracy of R measurements to achieve an equally

accurate inference of temperature. Given a typical

field-instrumentation accuracy of +2%, the above numbers

suggest that coil temperatures inferred from dc resistance

measurements may be expected to be accurate to within

perhaps ±50C, whereas ac impedance measurements of

comparable accuracy could not be expected to yield coil

temperatures more accurate than within ±17'C.

NUREG/CR-4819, Vol. 2 16
Evaluation of Diagnostic Methods Applicable to SOYs

450

a

z

60

0

coN

To

0

w

z

lco

wa cc

400 

350 

300 F

| .a * ~ ~ ' ~~~I I I I I I 

f.C at 20 'C) ~~IZI0.175 0o'C (00.5B %/C at 20 C)

Re (Z)

R -

I I * ~I. at I * I

250 F

200

150

100

20 40 60 80 100 120 140 160

COIL TEMPERATURE (0C)

Figure 2.7 Linear variation of coil dc resistance and 60-Hz ac impedance with temperature for a typical

ac-powered SOV.

One apparent solution to this problem would be to measure

Re(Z) rather than I Z , since the temperature coefficient of

the former (0.261%/C) is only 35% smaller than that of R.

The difficulty in doing so is the more complex and

therefore costly instrumentation necessary to measure

Re(Z) as opposed to I Z I.

Variation of Impedance with Excitation

In contrast to the dc resistance, which for all practical

purposes does not change with the amount of voltage

applied to the coil, the measured ac impedance of a

solenoid in its actuated state is strongly dependent on the

level of electrical excitation. This is illustrated in

Figure 2.8, which shows the manner in which the 60-Hz

impedance of a typical ac-powered SOV held at constant

temperature changes with impressed voltage (see also

Appendix E). Although it is true that the solenoid

impedance is very nearly constant at excitation levels

between 20 and 60 Vims (i.e., at voltages insufficient to

make the valve shift), once the solenoid changes state, the

impedance drops by approximately 40% as impressed

voltage is further increased from 70 to 135 Vms.

Of course, this variation of impedance with level of

excitation occurs simultaneously with (and is

indistinguishable from) the variation of impedance with

temperature that is the basis for its intended use as an

impedance thermometer, and must therefore be accounted

for when translating the measured impedances into

temperatures. Unfortunately, the excitation dependency of

impedance (-6.5 KVV at the operating point of 117 Vff5 for

the SOV whose characteristics are shown in Figure 2.8)

may be even stronger than the temperature dependency

(+0.72 WV0

C for this same SOV). Therefore, if the voltage

applied to the solenoid is not reasonably constant over

time, corrections for excitation changes may be quite large

and obviously would introduce additional uncertainty into

the inferred temperatures. C(Tis will be illustrated in a later

figure).

Four possible solutions to this problem come to mind:

(1) perform temperature inference measurements at a low

impressed voltage where the impedance is invariant with

respect to excitation level, (2) power the SOV from a

constant-voltage power supply, (3) make the

temperature-determining measurement with dc rather than

ac excitation, or (4) correct for the change in impedance

introduced by varying excitation voltage. Assuming that

the SOV is normally energized, the first solution requires

that the valve be allowed to change state momentarily,

which may or may not be permissible depending on the

plant situation. The second option is completely

17 NUREG/CR-4819, Vol. 2
Evaluation of Diagnostic Methods Applicable to SOVs

950 . . . . . . . . . . . . . . . . . . . . . . . .

, . .

i S . c l . d , . f 

I SOV closed * , 

.

. . I . . I  . . .

C' I I i I I I i , , . Ir

Pho1

0

850 

a

N

0

C

(U

0

Ili

I

0)

750

650

550

0I

I

a)I

p .

0,

O.

n,

lI

\ \ O0 ,

"6' -:I

'O

II450

350

mutual inductance saturates 

00o000000000 o 0 0 0 0 

0 I

10,I OVopn7

. . . . . . . . . . . . . . . . . . . . . . .

0 25 50 75 100 125

AC EXCITATION VOLTAGE (V)

Figure 2.8 Variation of ac impedance with impressed voltage for an SOV held at a constant temperature of 250C,

showing the abrupt impedance change accompanying valve state change.

straightforward but may be difficult to backfit into an

existing plant. The third solution, though technically most

attractive (because it solves both of the problems cited),

may be hardest to implement because special circuitry

would be needed either to switch the SOV momentarily

from ac to dc power or to superimpose a dc resistance

interrogation signal on the normal ac valve actuation

voltage. The fourth option, though inherently inexact

because of the large corrections required, is perhaps the

easiest to implement. Its workability is illustrated in the

laboratory demonstration described next.

Using a test sequence very similar to that described in

detail for a dc-powered SOV (Figure 2.5), the temperature

response of an ac-powered SOV to changes in fluid flow

through the valve, excitation level, and air flow in the

vicinity of the valve was demonstrated (Figure 2.9). The

same general features are present here as in the test

previously presented (including a difference in

temperatures indicated by the coil impedance vs the

mercury bulb thermometer reading that approaches 40 0C)

except that the scatter of the coil temperatures inferred

from ac impedance is noticeably larger than was observed

from the dc-derived data. This is understandable in view of

the preceding discussion regarding the reduced magnitude

of the temperature coefficient for ac impedance and the

necessity of large corrections to the data.

The magnitude of the corrections applied for excitation

voltage changes (specifically the changes at t = 70 and

115 min) can be seen by comparing the crosses and the

filled circles in Figure 2.9. The corrections were performed

by using a third-order polynomial fit to the postclosure

impedance data of Figure 2.8 to estimate the impedance

that might have been measured had the excitation been

exactly 117 Vns (the excitation used at the calibration

point for this valve) via the relation

25°C,117

Corr iZ|meas | Z 25°CV 

fit (2)

where r z I C represents the impedance appropriate to

temperature T, corrected to standard excitation;

I Z I "V . represents the impedance actually measured at

temperature Tand impressed voltage V; and the remaining

quantities represent impedances that would be obtained at

standard (117 Vrm,) and nonstandard (V) excitation

voltages at a temperature of 25 0C, as determined from the

polynomial fit to the data of Figure 2.8. Despite the large

corrections, for the purpose of detecting far-off-normal

operating temperatures that pose a serious problem in

regard to accelerated aging of coil insulation and valve

NUREO/CR4819, Vol. 2 18
Evaluation of Diagnostic Methods Applicable to SOVs

uncorrected ipedonce-derived

+ temperolur.e

Corrected impedonce-derived

* temperotes the.noeoter-deri.ed

terperotres

685 . . . . . . . . . . . . . . .

C

z

a

w

0.

C.)

-J

0

668

651

634

617

117 V still air , no

fluid flow thru SOV

.

0

.0

*'°0

o

* 0

*o t Process air allowed to

T flow through open SOV

Cah/ba//on point

itation Forced air cooling initiated;

eased excitation also decreased

to 126 V to 121 V 

0 0 a 

. An' r

.

* . S

000 0

C)0

,00 *Ooo°o o+

Process ir turned off.

other conditions sme

+4

00* 

* ,+4,+ 0

* a * 

0 0

4

110

0

LU

90 a:

70 u

50 n

Lu

30 I30 ()

S

z

10

+ 0 0 

2 .00

4.A

600 . . . . . . . . . .

0 30 60 90 120 150

ELAPSED TIME (minutes)

Figure 2.9 Changes in solenoid coil temperature as inferred from coil ac impedance, brought about by altered

electrical excitation, fluid flow, and environmental conditions.

elastomerics, the level of accuracy implied by the results of

Figure 2.9 is probably acceptable.

2.2.2 Practical Field Application

To illustrate the workability of this coil resistance/

impedance method of SOV operating temperature

inference in a true field environment, the rms voltage

impressed upon and the rms current drawn by an ac SOV

controlling the flow of refrigerant in a large chilled-water

air-conditioning system were recorded at 100-s intervals

over a 55-h period of system operation. The recorded

voltages were divided point by point by the recorded

currents to yield a curve of I Z I vs time. Knowing the

60-Hz I Z I of the solenoid at a known temperature (593.9

fQ at 25.5 0

C) and assuming a value of 0.115% C (the

median value for the five ac-powered SOVs tested) for the

temperature coefficient of impedance, an equivalent

temperature scale was affixed to the data plot; namely,

(2) were required for impedance change caused by altered

excitation level.

The test results, split into two roughly 27-h time periods for

clarity, are presented in Figure 2.10 and Figure 2.11(a) and

(b). The entire 55-h time period encompassed in these data

was characterized by generally rising outside air

temperatures (necessitating continuous compressor

operation to maintain the chilled-water temperature

setpoint during the initial 28 h of the test) as well as by

rising ambient temperature at the SOV location (which is

near the compressor, the essentially continuous operation

of which caused the entire equipment room in which it is

located to heat up). These weather and local environment

temperature trends explain the generally rising SOV

temperatures seen in the two figures. An environmental

perturbation was introduced 24 h into the test just to see its

effect a blower that had been aiding the cooling of the

finned SOV was turned off. This change resulted in a

prompt rise in SOV coil temperature (by about 10C),

followed by a slow fall as nighttime brought cooler

temperatures to the equipment room.

The second 27 h of operation [Figure 2.1 1(a) and (b)]

illustrate a new data feature; namely, cyclic compressor

operation. At first glance, the inferred temperatures during

these periods appear to have a great deal of scatter

T(OC) = IZ I (l)-5)9399)+ 25.5 (0.001 15)(593.9) -_' (3)

Since the impressed voltage changed very little over the

duration of the test, no corrections of the form given by Eq.

19 NUREG/CR-4819, Vol. 2
Evaluation of Diagnostic Methods Applicable to SOVs

635

630

I X . . I . . * I . . I T . T T }

period

C

LU

C)

z

w

0

U3

625

620

615

o SS

hX~~

~~~~~~~~~~~~AndP

I

. :j.;W

I

F

h

85

a

80 w

75 -

M

LU 70 0.

65 F-

-j

60 0

0 55 co

0

w 50 G

w

LL 45 z

Blower directed at SOV

turned off at this time

610

605 , . . . . . . L. . . . . . . . I

0 4 8 12 16 20 24 28

TIME (hours)

Figure 2.10 Weekend performance of an ac-powered SOV in a refrigeration system (first 27 h).

[Figure 2.1 1(a)], but when examined on an expanded time

scale [Figure 2. 11(b)], what had appeared to be data scatter

is revealed as repeated SOV heat-up during each cycle of

compressor operation followed by cool-down after each

compressor shutoff.

The temperatures inferred for the SOV installed on the

chilled-water system during the monitored period were

never so high as to pose a threat to its continued operation.

However, the data illustrate that, had some operational

anomaly occurred that would have resulted in a dangerous

rise in SOV temperature, it almost certainly would have

been detected by this relatively simple, nonintrusive

temperature-measurement technique, which required only

the connection of a voltmeter and a clamp-on current

transformer.

In conclusion, laboratory- and field-acquired test data

illustrate that the true operating temperature of an SOV can

be inferred simply, nondisruptively, and with satisfactory

accuracy for detecting temperature conditions that exceed

accepted operating limits by using the copper winding of

the solenoid coil as a self-indicating, permanently available

resistance thernometer. This approach to monitoring the

temperatures actually "seen" by SOVs during the course of

plant operation and the temperature changes introduced by

altered environmental or process conditions might be used

to advantage by those concerned with prediction of

qualified life. The method has a number of merits, among

which are the following: (1) there is no need for an add-on

temperature sensor, (2) the true volume-averaged

temperature of a critical-and probably the hottest-part of

the SOV (i.e., the electrical coil) is measured directly,

(3) temperature readout can be provided at any location at

which the SOV electrical lead wires are accessible (even

though remote from the valve), (4) the SOV need not be

disturbed (whether normally energized or deenergized) to

measure its temperature in situ, and (5) the method is

applicable to all types of SOVs-large and small, ac- and

dc-powered. From a standpoint of prediction of qualified

life, the principal shortcoming of the method is that the

coil's volume-averaged temperature-rather than its peak

temperature-is the quantity measured, although this

difference may not be substantial. 9

This method is usable in its current form, although

additional developmental work could improve the accuracy

of temperature indications derived from impedance

measurements on ac-powered SOVs. The hardware

implementation would depend on plant needs but could

take the form of a permanently installed, computerized data

logger or hand-held, walk-around instrumentation for use

on an as-needed basis.

NUREG/CR-4819, Vol. 2 20
Evaluation of Diagnostic Methods Applicable to SOVs

644

640

. . . . . . . . . . . . . . . . . .

a

z

D3

w

EL

0

Uo

636

632

628

, .,

compressor .. *-

: .-- -. * - _-' WA~d * 2 * 

.~~~~~~~~~~~~~~: I* ... 

compressor cyclng

compressor

* runnig * 

I I , | , |~~~I 

t.)

a

95

w

(C

90 :

a:

w

IL

85 m

w

-

80 0

75 0

Co

w

70 a:

w

U.

65 Z

624

620 .........................

28 32 36 40 44 48 52

TIME (hours)

Figure 2.11(a) Weekend performance of an ac-powered SOV in a refrigeration system (second 27 h).

642

640 

1-

w

z

ul

IL

-

0

co

638

636

634

632

630

628

"0

0 *I,0 O

*0 0 ft 0*

0~~~~

011~~~~~~~~

* 0

6 0~~~~~~~

0 0*~~

0 

0

95 U

a

Iwa:

9 c

go" wIL

0

to

80 0

w

M.

a:

wU- IL

z

75

45 46 47

TIME (hours)

Figure 2.11(b) Expanded view of a typical 2-h period from Fig. 2.11(a) during which the compressor was cycling.

21 NUREGfCR4819, Vol. 2
Evaluation of Diagnostic Methods Applicable to SOVs

2.3 Indication of Valve Position and

Change of State

[Ma jfunctions addressed: mechanical binding,

sluggishness, or failure to shift caused by worn or improper

parts or he presence offoreign materials inside the valve]

The technique presented in the preceding section addresses

the issue of reduced solenoid coil and elastomer life

resulting from prolonged operation at excessive

temperature, an acknowledged problem in the nuclear

industry. Here we address another well-known problem,

that of the solenoid plunger (core) "frozen" in position so

that it refuses to move when the SOV is called upon to

change states, whether normally energized or normally

deenergized. Such a condition may arise from age-related

changes in the physical properties of the elastomeric seats

(the seat material may become sticky, thus holding the

plunger fast), thermally degraded and deformed shaft seal

0-rings, faulty valve assembly, the use of incorrect

replacement parts, or the presence of internal contaminants

such as metal chips, dirt, paint, thread sealant, desiccant,

and hardened lubricant that can interfere with free

movement of the core within the core tube.1'

3 12 Depending

on the nature of the information to be obtained, this general

technique can be implemented either as a nonintrusive

static measurement of coil impedance or as a dynamic

measurement, the latter requiring application of a special

voltage profile to the SOV. The two types of tests are

treated in separate subsections below.

2.3.1 Static Tests (Nonintrusive to Plant

Operations)

Absolute plunger position within the solenoid coil or

simple indication of plunger movement upon application of

ac power can be obtained in situ by measuring the ac

impedance presented by the SOV electrical lead wires. As

shown in Figure 2.8, if SOV terminal impedance at 60 Hz

is plotted as a function of excitation level, it is found first

to increase about 30% as turn-to-turn flux linkages within

the coil saturate in the current region of 0 to 50 mA, then to

remain about constant until 130 mA is reached, at which

point the plunger shifts position, and is drawn abruptly into

the solenoid. This change of valve state is seen to produce

a near doubling of the coil's electrical impedance, which

then decreases fairly rapidly as excitation is further

increased, owing to increased eddy current losses and

magnetic saturation of the core iron. Accordingly, if the

temperature of the SOV is known (at least approximately),

a simple rms voltmeter/ammeter measurement of ac

impedance of the powered SOV (or an impedance meter

measurement of the unpowered SOV) will provide a good

indication of whether the valve is truly open or closed.

Carrying this idea one step further, one can try to quantify

the absolute position of the core within the solenoid coil.

This might be important if internal contamination by

foreign objects or substances were suspected of holding the

plunger slightly off its seat, thus causing the valve to

leak. 1.

12 Static measurements were performed by placing

movement-restricting shims of various thicknesses between

one or the other valve orifice and the corresponding

elastomer seat. Referring to Table 2.3, an impedance

sensitivity figure of I Z I 8d - 7 f)/mil of plunger

displacement (1.4% change in I Z /mil) was obtained for

small movements of the valve core in the vicinity of the

valve's unenergized seat position. This implies that an

impedance measurement performed on an SOV at a known

temperature and electrical excitation level could provide

indication of plunger position in the vicinity of the valve's

unenergized seat to within a few thousandths of an inch,

making detection of anomalous conditions causing

displacement of the plunger from its unenergized,

spring-assisted rest point a relatively simple matter.

However, as Table 2.3 also illustrates, a corresponding

sensitivity to plunger displacement in the vicinity of the

other (energized) seat was not found, probably because

after the iron core is drawn substantially into the

solenoid-thus becoming a part of the magnetic

circuit-further movement within the coil has little or no

effect on the resulting coil inductance (and hence the

impedance).

It must be clearly understood that the preceding discussion applies only to SOVs powered-permanently

or temporarily-by ac. Dc-powered valves show no

corresponding change in their terminal resistance as the

plunger changes its position within the solenoid coil, as

illustrated in Figure 2.12.

2.3.2 Dynamic Tests (Intrusive to Plant

Operations)

For ac-powered SOVs, dynamic tests that verify both the

presence of plunger movement upon application of power

and also the absence of binding throughout the transition

from an unpowered to a powered state, are also possible if

one is able to temporarily remove the 117-V, 60-Hz ac

power applied to the SOV and substitute an ac power

supply capable of ramping up and down more or less

uniformly with time. To see how this is possible,

Figure 2.13 shows the manner in which the current drawn

NUREG/CR-4819, Vol. 2 22
Evaluation of Diagnostic Methods Applicable to SOVs

Table 2.3 Variation of 60-Hz impedance with position or the plunger

SOV condition Z I (Q) 8 I Z I /d

Normal assembly, unenergized state 483.37 (base condition)

Shimmed 0.0035 in. off unenergized 510.97 7.89 Q/mil (1.6%/mil)

valve seata

Shimmed 0.0065 in. off unenergized 529.15 7.04 f/mil (1.4%/mil)

valve seat

Shimmed 0.0115 in. off unenergized 551.37 5.91 Wmil (1.2%k/mil)

valve seat

Normal assembly, energized state 635.5 (base condition)

Shimmed 0.0035 in. off energized 635.2 -0.09 flmil

valve seat b

Shimmed 0.0065 in. off energized 637.9 0.37 QWmil RESULT

valve seat

Shimmed 0.01 15 in. off energized 636.4 0.08 f/mil

valve seat

'The seat position assumed by the plunger, with spring assistance, when the SOV coil is unenergized.

bmhe seat position assumed by the plunger when the SOV coil is energized.

140

139 .1%

138

a 137 _ , _

w 136 SOV closed

< 135 0,

w 134-

EC 133-

soy

132 - State

Change

131

130

0 5 10 15 20 25 30 35

DC EXCITATION VOLTAGE (VdJ)

Figure 2.12 D resistance of SOV from open to closed state. Unlike ac impedance (Fig. 2.8), the dc resistance of an

SOV does not change when the valve closes.

23 NUREG/CR.4819, Vol. 2
Evaluation of Diagnostic Methods Applicable to SOVs

140

0

co

CO

0

cc

C.,

-J

1-

z

w

l'-

0

a.

C.)

120

100

80

60

40

250

6

200

0

co

X

150 CD

0

100 iz

W

50 C

C)

\

-0

50

20

0

0 10 20 30 40

ELAPSED TIME (s)

Figure 2.13 The nonlinear variation of SOV current during a linear rampdown of voltage over a 50-s period is a

reflection of the solenoid's nonconstant impedance at all but low levels of excitation.

by the solenoid changes with time when the voltage

applied to the coil leads is slowly ramped down in a linear

fashion from about 135 V ns 5 to 5 Vims over a period of

50 s.* The nonlinear behavior of the current seen in

Figure 2.13 over the initial 33 s (equivalent to the

right-hand side of Figure 2.8) reflects the variation of

impedance caused by eddy current and hysteresis losses,

whereas the behavior over the final 15 s (t = 35 to 50 s)

reflects the approximately constant and much lower

impedance produced by the plunger being somewhat

outside the solenoid coil (equivalent to the left-hand side of

Figure 2.8).

The shift of the plunger in its guide tube is clearly seen at

about t = 33 s as an abrupt change in the current drawn by

the SOV as a result of the change in impedance that

accompanies plunger movement. It is the current/voltage

(i.e., impedance) characteristics obtained during this brief

time of transition that reveal not only the overall movement

of the solenoid core but also any tendency it may have to

*The downward voltage ramp was produced by driving the shaft of a

Variac autotransfomier at a constant rotational speed (I rev/min) with

the aid of a gear motor, thus producing a ramp rate of -2.75 V/s. The

gear motor's direction of rotation could be electrically reversed to

produce an upward voltage ramp having the same rate of change with

time.

bind within its guide tube during the valve state transition,

as is illustrated at some length in Subsect. 2.4.

The ramp rate can be increased considerably in order to

improve both time definition and thereby the capture of

detail. For example, if we take an ac-powered SOV in good

condition and make provision to ramp its excitation voltage

over the full operational range ( to 135 V) in a rather short

time interval (200 ins) as shown in Figures 2.14 and 2.15,

then a marked rise in impedance, I Z I, and a

corresponding fall in current will be evident when the

valve changes state (t = 125 rns in Figure 2.14). On the

other hand, if the plunger is jammed in position and cannot

move (the condition purposely created in Figure 2.15), the

lack of impedance change and the continued monotonic

rise of current throughout the ramp-up will clearly signal a

lack of plunger movement.

If I Z I can be measured on a time scale of to

1.5 electrical cycles (from 17 to 25 ms for 60-Hz power)

following contact closure, it might even be possible to

obtain indication of plunger movement by using normal

switch-on actuation rather than a ramp-up, thus obviating

the need for electrical disconnections and a special power

source. An attempt was made to do so (see Figure 2.16) but

without much success, owing to the difficulty of

calculating rms quantities over time periods as short as

NUREG/CR4819, Vol. 2 24
Evaluation of Diagnostic Methods Applicable to SOVs

1000 . . . . . . . . . . . . . . . . .

800

0

-J

5

600 

Plunger

Begins

to Move

J000oo0ao0o 0 o0°°0 0 0 0 @+

0

4*4 4 .40* * * 

00 0 000 00 C

0

0)o

I-300

) 250

200 E

z

W

150 ¢

D

WfI ~~~J 100 0

50

200

400

200

0

- Impedance

- Current

0 50 100 150

ELAPSED TIME (ns)

Figure 2.14 Current and impedance discontinuities during linear voltage ramp-up over a period of 0.2 s-an

indication of plunger movement in a normally operating ac-powered SOV.

1000 . . . . 300

No Impedance

800 Change Means 250

800 No Plunger *

Movement

4 ~200 E

U 600 Z z 

< o

o 5 150 a: III 0o0 o000o0o 0 o00000o00oW

400 ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ ~ 0

E5 100~~~~~~~~~~

20

200 444 ~0 - Impedance 50

4 -Current

0 50 100 150 200

ELAPSED TIME (ns)

Figure 2.15 Lack of impedance change and current discontinuity during linear voltage ramp-up over a period of

0.2 s-an indication of lack of plunger movement in an ac-powered SOV in which the plunger was purposely jammed.

25 NUREGICR4819, Vol. 2
Evaluation of Diagnostic Methods Applicable to SOVs

100 ,> 300

E ' ' ~~~rm:/

80 200 uZ

CA 60 100 °

0

: 0 4 0 ---------- 0~~e D

Z <~~~~~~~~~~~~~~~~~~~~ 

Ji'iiT ,e Constont o/ Q mAdv

RAdS A.'q. CocudI W f\ 

w 20 I , 100

Z

,° / ' I ,'' I,

0 '' 'x -200 z

0 5 10 15 20 25 30 35 40 45

TIME AFTER CIRCUIT SWITCH-ON (ms)

Figure 2.16 Instantaneous current through a solenoid coil immediately following application of 60-Hz ac power

compared with rms voltage across the coil calculated by a true rms meter having an averaging time constant of 7.7 ms.

I to 1.5 electrical cycles. As seen in the figure, the rms

calculation simply cannot keep pace with the plunger's

rapid movement, which appears to be complete about

20 ms after power is applied.

In conclusion, positive indication of current SOV state and

ability to change state (as a dynamic response to an "open"

or "close" command) is often possible using either static

(nonintrusive) or dynamic (intrusive) ac measurement

techniques. The merits are similar to those cited for

temperature inference through coil resistance/impedance;

namely, that (1) there is no need for an add-on sensor,

(2) valve state readout can be provided at any location at

which the SOV electrical lead wires are accessible (even

though remote from the valve), and (3) the SOV need not

be disturbed (whether normally energized or deenergized)

if the static method is used. The shortcomings are that

(1) the method is applicable only to ac-powered SOVs and

(2) arrangements must be made to power the valve from a

special ramp-voltage power supply if the dynamic method

is used.

The methods, although useable in their present form, would

need additional development to provide a means of

obtaining desired dynamic information from instantaneous

switch-on tests rather than from the 200-ms ramps

described here. In any case, the preferred hardware

implementation would be determined by plant needs.

2.4 Verification of Unrestricted

Plunger Movement

[Malfunctions addressed: mechanical binding and sluggish

shifting caused by worn, swollen, or improper parts or the

presence offoreign materials inside the valve]

2.4.1 Sensitivity of Results to Presence of

Internal Contamination

The ramp-up technique described in Subsect. 2.3.2 is also

useful in diagnosing mechanical abnormalities less severe

than a completely immobile plunger (e.g., a plunger whose

movement within its guide tube is impeded by the presence

of foreign substances such as dirt or oil).3"12 Figure 2.17

illustrates the current/voltage relationship obtained for an

ac-powered SOV when its excitation is ramped up over its

full voltage range within a time period of a few tens of

seconds. Curve A (solid) was obtained with a clean,

normally assembled valve, whereas curve B (broken) was

obtained from a valve of identical construction in which

NUREG/CR4819, Vol. 2 26
Evaluation of Diagnostic Methods Applicable to SOVs

250

E

0

Co

X

0

a:

I

z

a:

a:

03

200

150

100

50

0

0 20 40 60 80 100 120

AC VOLTAGE ACROSS SOV (V)

Figure 2.17 Ac current and voltage characteristics obtainedTuring a slow excitation ramp-up, showing sensitivity

to the presence of mechanical binding. (A: clean, normally assembled SOV; B: valve treated to produce binding.)

250

E

0

U,

W

CA0

a: 0

I

z

w

C.)

200

150

100

50

0

0 20 40 60 80 100 120

AC VOLTAGE ACROSS SOV (V)

Figure 2.18 Ac current and voltage characteristics obtained during a slow excitation rampdown. (A: clean,

normally assembled SOV; B: valve treated to produce binding.)

27 NUREG)CR4819, Vol. 2
Evaluation of Diagnostic Methods Applicable to SOVs

the core and the interior of the guide tube had been coated

with a thin layer of thickened, sticky, liquid shellac during

assembly to simulate the mechanical binding that has been

found to result from polymerized lubricants, use of

excessive thread sealant, or unfiltered air supplies." 3"12

The curves are seen to track each other extremely well in

the low- and high-excitation regions that represent static

plunger positions (i.e., the purely electrical properties are

quite reproducible), but they differ markedly in the middle

region of the graph at which the valve shifts its state. The

"normal" SOV was seen not only to shift at a lower level of

excitation than its "gummy" counterpart but also to exhibit

some instability as it began to execute the change of state.

(The valve was audibly buzzing at this time during the

ramp-up.) However, the SOV having internal

contamination exhibited a smoother (damped) transition

after sufficient magnetic force was developed to overcome

the resistance to movement produced by the sticky shellac.

Similar results are obtained if the excitation is ramped in

the decreasing (Figure 2.18) rather than increasing

direction-namely a pronounced difference between the

voltage/current transition points for the normal and

sluggish SOVs in addition to lower V and critical values

than were obtained for the ramp-up. Both upward and

downward ramp tests thus establish a set of four pull-in and

dropout critical points [identified by "TP" (trendable

parameter) symbols in the figures] that can be trended over

time to provide early indication of mechanical binding,

should it occur.

2.4.2 Consistency of Results for Clean Valves

in Good Operating Condition

The degree to which the pull-in and dropout voltages and

currents provide reliable, consistent performance indicators

was studied only briefly and with mixed results. The

reproducibility of the voltages and currents first required to

make the plunger draw into the solenoid (i.e., the pull-in

point) and then to allow the spring to withdraw the plunger

(i.e., the dropout point) was measured for both ac and dc

SOVs, as illustrated in Figures 2.19 and 2.20 respectively.*

For all these baseline tests the valves were clean and

assembled normally. The test duration for the ac-powered

valve "D" (Figure 2.19) was -2 h for 17 trials; for the

*Completion of the SOV stroke (i.e., the pull-in or dropout point) was

ascertained by the sound produced by the core shifting its position. For

dc-powered valves, the critical points were clearly audible as a distinct

metallic click; for ac-powered valves, the critical points were less

precisely determinable, owing to a continuous buzzing of the core in its

guide tube as the transition point was approached from either direction

(but especially for the upward-going ramp).

dc-powered valve "B" (Figure 2.20) the duration was 3

days, each day consisting of 10 trials subdivided into 2 sets

of 5 consecutive actuations separated by -1 h.

Despite careful and consistent test execution, the data

scatter is seen to be rather large (±10%) for the ac-powered

SOV. The gradual merging of the pull-in and dropout

voltage data throughout the tests is thought to be the result

of gradual heating of the valve assembly produced by

repeated actuations over the relatively short testing period,

even though power was applied during each trial only long

enough to obtain an accurate measure of the transition

points.

Data scatter is substantially reduced (±3%) for the

dc-powered SOV, making possible the observation of

deviant behavior at each occurrence of first valve operation

following a period of inactivity. The anomalous data points

in the pull-in voltage and current plots of Figure 2.20 at

trial numbers 1, 11, and 21 each represent the first time that

the valve was actuated on successive test days. Close

examination of these data will reveal a much smaller but

consistent aberration on each fifth trial, resulting from the

1-h pause between test sets.

To determine the reason for the more consistent

performance of the dc-powered SOV, further tests were

conducted on additional valves, yielding the statistical

results summarized in Table 2.4. The data plotted in

Figures 2.19 and 2.20 correspond to the first and third rows

of the table respectively. (Figure 2.19 shows the

trial-by-trial results for SOV "D," a valve intended for ac

operation and powered by ac for this test, and Figure 2.20

displays corresponding results for SOV "B," a valve

intended for dc operation and powered by dc for this test.)

In order to sort out performance differences based on

different valve constructions from those based on the use

of ac vs dc power sources, cross-combination tests were

also included in the test matrix (i.e., valves designed for ac

operation were tested on dc power as well as on ac).*

The following conclusions may be drawn from the data of

Table 2.4 (primarily from the mean values and the

variabilities expressed as percent of mean value):

*It was not possible to power the SOVs designed for dc operation on ac, however,

since their 60-Hz impedances (of about 3 to 5 kW) were much too high to permit

currents sufficient to actuate the valves to be produced at reasonable applied

voltages (<200 Vac). Likewise, SOVs intended for ac operation proved

troublesome to test on dc using the motor-driven autotransfonner because their

coil dc resistances were so small that large currents flowed at quite low excitation

voltages, thereby reducing the accuracy achievable with the relatively rapid ramp

rate of 2.75 V/s.

NUREG/CR-4819, Vol. 2 28
Evaluation of Diagnostic Methods Applicable to SOVs

A Pull-in -- 0 Dropout ... V-.Pull-in -0-Dropout U 

voltage voltage current current

125 -j 80

E

a.

"a

z

C.)

z

I

O

115

105

95

85

75

65

R..d LEFT-HAND I,l*I

.V V

v vA

_-12 %

IV'V

.A R..d RIOI4T44AND WOW

6 A. A A. -: A .A A oA A

I0 K~~~~~~~~~~~.d RIGHT-HAND We

75

70 1-:

65 (.'

z

60 

55 

00

-j

0 A :; 

* 

T .

R..j LEFT HAND S-..E U . M

in * IM

I . I . I . I  I . I . I  I 55 40

1 3 5 7 9 1 1 13 15 1 7

TRIAL NUMBER

Figure 2.19 Short-term variability of pull-in and dropout voltages and currents for a clean and normally

assembled ac-powered SOV. [The gradual merging of the pull-in and dropout voltages (middle curves) with trial

number is associated with temperature rise of the SOY resulting from application of power over the testing period

(-2 h).]

..A.. Pull-In ----0.... Dropout - --V. Pull-in .... M.... Dropout voltage voltage current current

100

E

F:

z

C,

z

cc

C.

90I

80

70 

A

- V6VVVV vVV V-VV.Vvv, fVVVVVVV-VV

POW LIFT-HANG erni

ReeW LFT-HAND 0csi.

Reed RIGHT-HWW so. 

. ..... .... . . 00 . *.*.. I .. .. 

70

b5

60 IL

z

50 I

w

45 

0 40 >

60 

50 

40 

30 35

1 6 1 1 16 2 1 26 3 1

TRIAL NUMBER

Figure 2.20 Short-term variability of pull-in and dropout voltages and currents for a clean and normally

assembled dc-powered SOV. (The discontinuities at trials 1, 11, and 21 result from first operation after a period of

inactivity.)

29 29 ~~~~~NUREG/CR-4819, Vol. 2
Evaluation of Diagnostic Methods Applicable to SOVs

Table 2.4. Variability of SOV actuation levels (pull-in and dropout voltages and currents) and

its relationship to solenoid design and power source (ac vs dc)

Pull-in volts Dropout volts Pull-in mA Dropout mA

Min No. of trials Min No. of trials Min No. of trials Min No. of trials Valve ID Power Mean Std. dev. Mean Std. dev. Mean Std. dev. Mean Std. dev.

(type) source Max (% of mean) Max (% of mean) Max (% of mean) Max (% of mean)

Da ac 58.7 16 51.5 17 106.8 16 58.5 17

(ac) 62.5 2.16 58.0 3.1 115.1 4.6 64.4 3.8

66.3 (3.5) 63.4 (5.4) 119.9 (4.0) 72.0 (5.8)

Ab dc 71.5 31 37.3 31 84.8 31 45.2 31

(de) 72.2 0.51 38.0 0.10 85.8 0.83 46.1 0.36

73.4 (0.7) 38.7 (1.1) 88.1 (1.0) 46.7 (0.8)

Bb dc 65.3 31 37.9 31 80.2 31 47.5 31

(dc) 66.1 0.68 38.5 0.31 81.0 0.98 48.1 0.21

68.6 (1.0) 39.0 (0.8) 84.6 (1.2) 48.4 (0.4)

Cb dc 48.7 31 6.2 31 60.1 31 7.5 31

(dc) 49.2 0.54 6.71 0.21 60.8 0.75 7.93 0.33

51.7 (1.1) 7.2 (3.2) 64.5 (1.2) 8.7 (4.2)

De ac 65.9 28 46.7 28 83.5 28 53.2 28

(ac) 67.8 0.96 48.3 1.04 108 13 70.2 12

69.9 (1.4) 50.5 (2.2) 123 (12.0) 93.2 (17.3)

El ac 72.1 28 50.4 28 81.6 28 53.9 28

(ac) 74.9 1.9 56.3 3.1 142.1 42 109.1 38

78.9 (2.5) 62.4 (5.5) 191.8 (29.5) 155.1 (34.7)

He ao 64.5 28 37.3 28 116 28 67.9 28

(ac) 66.8 1.52 40.3 1.85 196 69 169 64

69.0 (2.3) 43.9 (4.6) 304 (35.0) 229 (37.8)

De de 20.6 28 2.37 28 150 28 19.8 28

(a) 21.0 0.45 2.68 0.19 152 2.9 21.0 0.67

23.1 (2.2) 3.03 (7.0) 166 (1.9) 22.4 (3.2)

E' do 30.0 28 2.66 28 214 28 21.4 28

(ac) 30.8 0.64 3.09 0.29 219 5.5 23.8 1.8

33.3 (2.1) 3.68 (9.5) 245 (2.5) 26.6 (7.7)

H do 16.3 28 1.10 28 296 28 26.1 28

(ac) 16.6 0.25 1.40 0.18 302 3.6 30.3 1.7

17.2 (1.5) 1.67 (12.6) 310 (1.2) 32.8 (5.5)

a[)ata aoquired 7/88 using instrumentation having 500-ms response (2 readings per second).

bData acquired 2/89 using instrumentation having 500-ms response (2 readings per second).

cData acquired 7/91 using instnmentation having 100-ms response (10 readings per second).

1. For ac SOVs powered on ac, pull-in and dropout voltages

appear to be considerably more reproducible from test

to test than pull-in and dropout currents.

2. For ac SOVs powered on dc, pull-in voltages and currents

appear to be more reproducible than dropout voltages

and currents. (This apparent difference may be

attributable to imprecise measurement of small

numerical values of current and voltage rather than to

true valve performance differences, and so must be

regarded with some skepticism.)

3. For dc SOVs powered on dc, pull-in and dropout voltages

and currents appear to be about equally reproducible,

provided that the dropout values are not so small in

magnitude that least-count errors in the instrumentation

artificially inflate the variance.

4. For the single SOV ("D") for which ac critical-point

measurements were repeated after a considerable lapse

of time (3 years), the repeatability was only as good as

could be expected from the standard deviations

estimated from the repeated trials: mean pull-in voltage

increased over this period from 62.5 to 67.8 V, whereas

NUREG/CR4819, Vol. 2 30
Evaluation of Diagnostic Methods Applicable to SOVs

mean dropout voltage decreased from 58.0 to 48.3 V;

mean pull-in current decreased from 115.1 to 108 mA,

whereas mean dropout current increased from 64.4 to

70.2 mA. (These changes are not consistent with an

assumption that valve "D" developed some degree of

internal binding over the 3-year idle period because, as

is demonstrated by Figures 2.17 and 2.18, binding

would be characterized by simultaneous increases in

pull-in voltage and current plus decreases in dropout

voltage and current, and this is clearly not the case

here.)

5. There is some evidence that the detailed properties of the

measuring instruments may influence the results

obtained. (Citing again the repeat measurements on

SOV "D" in 1988 and 1991, it appears that-in the

absence of other unknown factors-the use in 1991 of

instrumentation having more rapid response caused the

standard deviations of the voltage measurements to

decrease but caused the standard deviations of the

current measurements to increase. This may result from

the manner in which the different meters used in 1988

and 1991 compute the rmis value of a nonstationary,

alternating-polarity voltage or current; this possibility

was not investigated further.)

6. Critical voltage and current values are sensitive to valve

construction details. For example, the electrical

properties of SOVs "A" and "C" are virtually identical

(see Appendices A and B), the coil housings look

substantially the same (Figure 2.1), and the valve core

assembly, core spring, and solenoid base subassembly*

are of similar construction, although the actual valve

bodies are noticeably different in shape and although

two-way SOV "C" lacks the disc holder subassembly

incorporated in three-way SOV "A" (see Figure 2.2).

The pull-in voltages and currents are about 45% greater

for SOV "A" than for "C" able 2.4), whereas the

dropout voltages and currents are greater for SOV "A"

than for "C" by a factor of almost 6. This leads to the

conclusion that the magnetic attraction and spring

restoration forces must be quite different for the two

SOVs, perhaps as a result of different stroke lengths,

degrees of core insertion into the solenoid, or materials.

Therefore, despite early results (Figures 2.19 and 2.20)

which gave indication that the critical excitation levels

were less repeatable for a particular ac-powered SOV than

for a particular dc-powered SOV, further investigations

with an increased number of samples revealed that a broad

*Refer to Fig. 1.1 for terninology and location of parts.

generalization of this initial finding cannot be supported

because valve construction differences and greater

measurement uncertainties for the ac SOVs were shown to

affect experimental variance estimates.

Because we had no service-aged SOVs available for test

nor any means by which to realistically simulate gradual

valve degradation from the binding of internal parts, no

data from a degraded SOV are available for comparison.

This leaves unproven our contention that measurement of

pull-in and dropout voltages and currents should provide

reliable indicators of unrestricted plunger movement.

In summary, measurements performed on valves in which

plunger movement had been somewhat impeded by

artificial (but moderately realistic) means illustrate that

measurements of coil voltage and current at valve

transition points may be useful as early indicators of SOV

contamination by foreign substances, inadvertent use of

incorrect internal replacement parts, swollen elastomers,

and other deteriorations that can cause sluggishness of the

valve plunger. However, further testing will be required to

establish the sensitivity and limitations of this

performance-monitoring technique. Even though the

technique is applicable to both ac and de SOVs and has no

need for add-on sensors or for physical access to the valve,

it is also true that measurement of pull-in and dropout

critical points is inherently intrusive, necessitating

disconnection of the SOV from its normal source of

excitation and temporary substitution of a variable-voltage

power supply. There is also some indication that the critical

points for ac-powered valves-even those in good

operating condition-display greater variability than do

dc-powered SOVs, which may limit the application of this

technique.

2.5 Indication of Shorted Turns or

Insulation Breakdown Within the

Coil

[MaIfinction addressed: electricalfailure of solenoid coil,

caused by high-voltage turnoff transients in combination

with insulation weakened by prolonged operation at high

temperatures]

On the basis of recorded operational experience and

opinions cited in the Phase I study,1 a considerable number

of coil open- and short-circuit SOV failures may be

attributable to high-voltage transients generated by the

abrupt collapse of a dc solenoid's magnetic field as a result

of circuit interruption. SOVs operating in elevated31 NUREG/CR4819, Vol. 2
Evaluation of Diagnostic Methods Applicable to SOVs

temperature, high-humidity environments are particularly

susceptible to failure from this stressor because he

insulating properties of the coil varnish, encapsulating

material, and electrical lead wires are weakened under

these conditions. The phenomenon is quite similar to the

generation of high-voltage pulses of short duration that

feed the spark plugs in an automobile's ignition system.

Such an ignition system is composed of an ignition coil

whose primary winding current is turned on and off rapidly

as a result of periodic closing and opening of the breaker

points within the distributor. In both the automobile and the

SOV, energy that is stored in the magnetic field of an

inductor (i.e., the coil) as a result of being energized with

dc current is converted at the instant of circuit interruption

to an electrostatic potential (i.e., voltage) that appears

across the distributed capacitance of the coil (and, in the

case of the automobile, across the secondary winding).

It stands to reason that such high-voltage transients within

the solenoid coil may produce damage by puncturing the

insulation that covers the copper wire of the winding at any

existing weak points. When this has happened, the area

surrounding the point of puncture becomes carbonized,

thereby lowering the insulation resistance and so assuring

that the same spot will be the site of electrical arcs upon

future valve deexcitations. Once begun, this process

produces rapid degradation of the insulation, which will

ultimately bum away and leave a permanent tum-to-tum or

layer-to-layer short circuit. According to industry sources*

the consequences of such a short circuit within the coil are

not especially serious for dc solenoids, which will continue

to function normally with a considerable number of shorted

turns. However, ac solenoids are not so tolerant in this

regard because, in effect, a step-down transformer is

created by the shorted turn. The result is a large current

flow through the shorted turn, intense localized heating,

and eventual thermal destruction (followed by electrical

destruction) of adjacent portions of the coil (i.e., coil

burnout).

An example will serve to illustrate that very high voltages

can be generated at the coil terminals at power turnoff,

provided that the energy transfer from electromagnetic

field to electrostatic potential is accomplished

efficiently.** The energy stored in the electromagnetic

field of an inductor of inductance L whose field is being

sustained by a continuous current I is given 13 by

*Telephone discussion, John Shank and Frank Fry, Automatic Switch

Company (ASCO), with Robert Ksyter (ORNL), November 17, 1988.

**Without going into detail, the rapidity with which the current

interruption takes place, dlldt, is a major determining factor.

Eem1⁄4= /LI2 ; (4)

taking L = 6.5 H for a typical dc-powered SOV

(Appendix C) whose 8002- resistance will draw -155 mA

at the rated excitation voltage of 125 Vdc, we obtain

Eem = 78 mJ. Likewise, the energy stored in the

electrostatic field of a capacitor of capacitance C charged

to a voltage V is given13 by

Ees=/2CV2 ; (5)

therefore, if all the energy stored in the inductor were to be

transferred to the capacitor, it would charge to a voltage

V, = i'2fEe /C . (6)

Substituting an approximate value of 45 pF for the

distributed capacitance of a typical fully assembled SOV

yields an estimate for Vc of -59 kV.

This estimated "flyback" voltage is clearly an upper bound

since in practice the energy transfer between the

electromagnetic and electrostatic fields cannot be achieved

without loss. Nonetheless, as illustrated in Figure 2.21, it is

indeed possible to generate large inductive surges at the

SOV coil terminals (in this example, almost 25 kV for a

duration of -100 ps) by abruptly deenergizing a 125-Vdc

solenoid using a relay having mercury-wetted contacts

(Figure 2.22) so as to obtain an extremely rapid

interruption of current flow (because the magnitude of the

flyback voltage is in large measure controlled by the time

derivative of the current, dl/dt). Despite this finding we do

not believe that inductive surge need be a major problem in

connection with SOV service ife, for the following reasons:

a Coil insulation is tough. Hundreds of transients of

the magnitude shown in Figure 2.21 were produced

during the course of this study, but they failed to

cause any shorted turns in a variety of coil types

having Class H insulation. (It must be

acknowledged, however, that our tests were

performed in a mild operating environment rather

than under loss-of-coolant accident conditions.) The

new Class N insulation is claimed to have electrical

properties superior to Class H, particularly in high

humidity.* Its temperature rating for continuous

operation is also 15C higher than for Class H.

Telephone discussion. John Shank and Frank Fry (ASCO) with Robert

Kryter (ORNL). November 17, 1988.

NUREG/CR-4819, Vol. 2 32
Evaluation of Diagnostic Methods Applicable to SOVs

-j

6

0co

co

Z:U

0

I0M-a

a0

0

-5

-10

-15

-20

-25

0.0 0.2 0.4 0.6

TIME AFTER CIRCUIT INTERRUPTION (ms)

0.8

Figure 2.21 Large "flyback" voltages generated when current to 125-V dc-powered SOV "C" was interrupted

abruptly. [A mercury-wetted contact relay was employed to obtain clean, extremely rapid circuit-breaking action (with

no contact bounce). The test circuit is shown in Figure 2.22.]

The RC (0. Of jiF f /) Jf network across the

181 relay o contacts IS for arc sppress/ol.

0.01 

pF T C.P. Claire mercury-wetted relay, type HGP 2009.

Variable frequency

pulse or squarewave

generator

(freq-1-10 Hz)

Figure 2.22 Circuit used for measuring transients generated by interruption of coil excitation to a dc-powered

SOV.

33 NUREG/CR-4819, Vol. 2
Evaluation of Diagnostic Methods Applicable to SOVs

Transient suppression devices are readily

available. Inexpensive metal-oxide varistors

(MOVs) can be placed across the coil leads to

effectively absorb the energy released by the sudden

collapse of the solenoid's magnetic field upon

turnoff. [See Figures 2.23(a) and 2.23(b), in

comparison to Figure 2.21.] The magnitude of the

high-voltage transient produced upon deenergizing

the SOV is thereby greatly reduced (e.g., -600 V

rather than -23,000 V), although the transient's

duration is correspondingly extended (e.g., -20 ms

rather than -0.8 ns). However, the positiveness and

rapidity of solenoid release are not noticeably

affected by the presence of the transient-suppressing

device. (Tests demonstrated an ability to cycle an

MOV-protected SOV at a rate of at least

5 actuations per second.) Therefore, no major

change in SOV operating characteristics would be

expected to result from the addition of such a

protective device.* The extent to which transient

suppression devices are used for SOV coil

protection in U.S. nuclear power plants is unknown

to the author.

* Potentially damaging voltage differences are not

as great as they might at first appear. The voltage

differences appearing between adjacent turns and

even between successive coil layers are only a small

fraction of the total inductive surge voltage

appearing across the coil terminals as a result of

magnetic field collapse. Industry sources** state

that although turn-to-turn short circuits can occur as

a result of imperfections in the lacquer insulation

coating the copper coil wire, it is quite rare that this

causes any problem, because the turn-to-turn

voltages are too small to jump the interconductor air

gap. It is also stated that layer-to-layer short circuits

are practically unheard of because most coil

manufacturers place additional dielectric material

between successive coil layers during the winding

process.**

* Circuit interruption is likely to be slower in

practice. The manual switch or relay normally

employed in nuclear plants will probably not break

This conclusion is in agreement with the authors of ref. 1, who state (on

p. 26) 'lAlthough] testing has shown that the operation of a relay

actuated by a coil with surge suppression usually takes twice as long as

one without this protective feature. .generally this time delay is small

enough not to be a practical concern in control circuits.'

**Telephone discussion, John Shank and Frank Fry (ASCO) with Robert

Kryter (ORNL), November 17. 1988.

the circuit as rapidly as the mercury-wetted contact

relay used in these laboratory tests; therefore, the

flyback voltage generated in practice is unlikely to

be as high as that shown in Figure 2.21.

Dc SOVs are tolerant of shorted turns. The

previously cited industry sources state that shorted

turns present a serious operational problem only for

ac-powered SOVs, whereas the production of a large

flyback voltage upon circuit interruption is a

characteristic of dc-powered SOVs. (Ac-powered

SOVs have magnetic fields that continually change

polarity, and circuit interruption occurs at a random

point with respect to the phase of the 60-Hz power

line.)

Changing our point of view, it is conceivable that the

inductive surge phenomenon can be turned to

advantage as a means for measuring certain critical

solenoid coil electrical parameters that would

probably be affected by shorted turns, insulation

breakdown, and similar degradations. Following up

on work performed by others for NASA,14 this

possibility was investigated by examining the free

oscillation ("ring-down") behavior obtained upon

deenergization of two bare solenoid coils of identical

construction-one undamaged and the other having

-176 of its -3042 original turns shorted out as a

result of having undergone severe EQ tests.* A

mercury-wetted contact relay was used to provide a

clean circuit interruption, but to avoid the occurrence

of dissipative phenomena (e.g., saturation losses in

the core iron), only 30 Vdc was applied to the coil

instead of the normal operating voltage of 125 Vdc.

The undamaged coil was observed to behave like a

damped oscillator when deenergized from this lower

voltage (Figure 2.24), whereas the waveform

produced by the coil with shorted turns was highly

asymmetric and died out very rapidly (Figure 2.25).

Some insight into the marked differences between

these two ring-down traces can be obtained by

examining the solenoid's lumped equivalent circuit

(Figure 2.26). Basically, the inductance of the coil,

L, plus some inevitable series resistance, RL, are

seen to be shunted by the distributed (turn-to-turn

and layer-to-layer) capacitance of the solenoid, C. In

principle, there is also some series resistance, Rcl,

*These coils were loaned to us by John Shank and Frank Fty of ASCO,

Florham Park, New Jersey.

NUREG/CR4819, Vol. 2 34
Evaluation of Diagnostic Methods Applicable to SOVs

0 III' I . I . . I I . I . I

5

o

0

C),

tr

D

z

2

AIL

.- c0

0o

-100

-200 F

-300 F

-400 

-500 F

-600

. . . .. . . . . . . . . . . . .

0 2 4 B 8 10 12 14 16 18

TIME AFTER CIRCUIT INTERRUPTION (ms)

Figure 2.23(a) Same SOY and experimental setup as for Fig. 2.21, except that two varieties of transientsuppressing metal-oxide varistors were placed directly across the terminals of SOV IC" to absorb the energy released

by the coil's collapsing magnetic field. (The voltage decay characteristics of the two types of suppressor are essentially

the same for times greater than 6 ms. The experimental setup is the same as for Fig. 2.21.)

0

-j

Cd,

0

aZ,

Ii

0

-100

-200

-300 F

-400 1

-500 

-600

I I 

0 1 2 3 4 5

TIME AFTER CIRCUIT INTERRUPTION (ms)

Figure 2.23(b) Time expansion of the early period illustrated in Fig. 2.23(a), showing limitation of peak transient

voltage to -500 V by the 10-A suppressor (upper curve) and to -600 V by the 1-A suppressor (lower curve). (The

suppressor with the higher current rating achieves its somewhat better peak limiting at the expense of slower voltage

decay.)

35 NUREG/CR-4819, Vol. 2
Evaluation of Diagnostic Methods Applicable to SOVs

1500

_1I

co

co

0

a:

z

I- a:

0

1000

500

0

-500

-1000

-1500

-2000

0.0 0.4 0.8 1.2 1.6

TIME AFTER CIRCUIT INTERRUPTION (ms)

Figure 2.24 The solenoid coil with no shorted turns behaves like a classical damped oscillator when deenergized

from 30 Vdc by a wetted-contact relay.

200

-g

to0

t

CA) 0

cc

z

a:

IL0.

0

I0

0

-200

-400

-600

-800

0.0 0.4 0.8 1.2

TIME AFTER CIRCUIT INTERRUPTION (ms)

Figure 225 The solenoid coil having -176 of its -3042 turns shorted behaves quite differently from its healthy

counterpart when deenergized from 30 Vdc by a wetted-contact relay.

NUREG/CR4819, Vol. 2 36
Evaluation of Diagnostic Methods Applicable to SOVs

- -- 1

C

L RC2 0

- \ RL RC,

Figure 2.26 Equivalent circuit for an SOV. (As a first approximation, the possible presence of parasitic resistances

Rci and RC2 may be ignored for simplicity of mathematical analysis.)

associated with the coil's distributed capacitance,

but this is probably small in value and can therefore

be ignored in first approximation. Likewise, in

principle there is also some shunt resistance, Rc2,

associated with the coil's distributed capacitance,

but it is probably very large in value and can

therefore be ignored in first approximation. With

these simplifying assumptions, the free-oscillation

resonance frequency,fr, for the series/parallel

combination of L, Ri, and C is found to be'5 given

by

f 1 I1 Rt r f 2i 4L2

Because we can easily measured, L, and RL, it is

convenient to rearrange Eq. (7) and solve for the

unknown parallel distributed capacitance, C:

the protective coil housing), yields estimates of

distributed parallel capacitance in the range of 150 to

210 pF (for bare coils) and 30 to 45 pF (for fully

assembled SOVs).

Two observations may be drawn immediately from this

equivalent circuit analysis:

* The natural resonance frequency associated with

circuit interruptions of SOVs (i.e., the ring-down

frequency) depends on the value of the inherent

distributed capacitance of the solenoid coil that

appears in parallel, electrically, with the coil's

inductance. Because the numerical value of this

distributed capacitance is likely to change

substantially if the coil develops faults such as

shorted turns or layer-to-layer insulation breakdown,

measurement of natural resonant frequency would

appear at first glance to be an excellent means of

detecting SOV coil electrical faults (as borne out by

the strikingly different characteristics of

Figures 2.24 and 2.25).

* However, because the capacitance of the electrical

cable connecting the SOV to its power source

(which may approach a value of tens of nF in

practice) also appears in parallel with the SOV coil,

(7)

C= 4L

4(Lf,2j)2 + RL (8)

where C is in farads, RL in ohms,f, in hertz, and L in

henrys. Applying Eq. (8) to transient data obtained

on several solenoid coils, both bare and fully

assembled (with plunger present and placed inside

37 NUREGJCR4819, Vol. 2
Evaluation of Diagnostic Methods Applicable to SOVs

this cable capacitance is likely to completely

overshadow the distributed capacitance of the coil

alone (which is about 50 pF) and thereby fix the

ring-down characteristics of the in-situ SOV,

regardless of the true condition of the solenoid coil.

Although providing general insight, the simplified

equivalent circuit analysis presented does not appear to

explain adequately the marked differences between the

ring-down data obtained for the normal SOV solenoid coil

(Figure 2.24) vs the SOV coil known to have -6% of its

original turns shorted (Figure 2.25).* First of all, if one

calculates the approximate distributed capacitance for the

normal and faulted coils from deexcitation transient data

using Eq. (8), substantially the same values (210 pF for the

normal coil vs 215 pF for the faulted coil) are obtained in

both cases.** This finding is contrary to the seemingly

reasonable assumption that the presence of shorted turns

would alter the coil's distributed capacitance, C. (Perhaps

this assumption is true for layer-to-layer but not for

turn-to-turn faults; we have no idea which type of fault is

present in the defective coil, which is of molded

construction and therefore cannot be unwound or examined

visually without destroying it.) Secondly, we must deal

with the apparent contradiction that although the extremely

rapid damping (or die-away) of the faulted coil's transient

can be explained in terms of the simplified equivalent

circuit model only by an increase in the parameter

a=RL/2L (9)

-because the magnitude of a establishes the rate at which

the oscillatory transient dies away with timel5

V) = 2 ea sin(2frt) (10) 2ir fr L (0

-no such large increase in a is actually measurable with a

low-excitation-voltage impedance bridge for the defective

coil as opposed to the normal one:

defective (297.0 484 1 1

(2)(100.2 mH) (1

,vhereas

tnormal = 113 9 =441 s 1 . (2X 113.9 mH) (12)

In fact, other than the vastly different free-oscillation

characteristics, about the only electrical properties that are

significantly different for the faulted and unfaulted coils

(Table 2.5) are the in-phase component of impedance,

Re(Z), measured at an excitation frequency of 1000 Hz

(231.2 vs 103.8 [I), and the closely related quality factor,

Q, also measured at 1000 Hz (2.7 vs 6.9).

We have no explanation for the discrepancies described;

they are presented for completeness and in the hope that

some more knowledgeable reader may be able to use these

results to obtain a better understanding of the phenomena

involved. Perhaps the lumped equivalent circuit model is

just too simple to account for the changes introduced by

shorted turns within the solenoid coil, or perhaps the shorts

are not "hard"; that is, their presence is unmeasurable at the

low levels of excitation applied by the impedance bridge

and detectable only when the coil is subjected to

high-voltage transients.

In summary, it is clear that additional research will be

needed to advance the understanding of the various

electrical phenomena that occur in coils having defective

windings before a practical SOV coil diagnostic system

can be devised, should it be decided that one is truly

needed. Results also show that it may be impractical to

perform coil diagnosis in situ from a distant location using

any sort of measurement technique based on free

oscillation upon deexcitation, because the capacitance of

the power leads connecting the valve to its power source

l) may be several times larger than the distributed capacitance

of the coil itself (thereby clouding interpretation of the

measurement). However, if the coil power leads can be

accessed directly (with the capacitive burden of the cable

removed), the flyback transient characteristics may provide

a far more sensitive indication of coil insulation breakdown

than does simple change in dc resistance.

I

*Telephone discussion, John Shank and Frank Fry (ASCO) with Robert

Kryter (ORNL), November 17, 1988.

**The data used to estimate C are not those shown in Figures 2.24 and

125, since the latter data were obtained at 30 Vdc where R-C relay

contact protection is required, thereby making the test results more

difficult to interpret. The transient curves used to estimate C were

obtained without relay contact protection at 10 Vdc, where resonant

frequencies of -34 kHz were observed.

Note also that if transient suppression devices (MOVs)

were used consistently to protect dc-powered SOVs a

major contributor to solenoid coil failure would be

eliminated, leaving only operation at excessive

temperatures (and perhaps radiation damage in a very few

plant locations) as a significant stressor for coil insulation.

NUREG/CR-4819, Vol. 2 38
Evaluation of Diagnostic Methods Applicable to SOVs

Table 2.5. Room-temperature electrical properties or faulted

and normal bare coils

Coil 17 (3042 eff. turns before, Coil 18 (3039 eff. turns

Property 2866 eff. turns after EQ tests) before and after EQ tests)a

Rde 97.0 L2 100.4 Q

Re(Z) (5 Hz) 97.66 101.13

Im(Z) (5 Hz) 3.09 3.32

Re(Z) (60 Hz) 98.05 101.02

Im(Z) (60 Hz) 40.02 42.92

Re(Z) ( kHz) 231.2 103.8

Im(Z) (1 kHz) 629.8 715.7

L (5 Hz) 98.4 mH 105.8 mH

Q(5 Hz) 0 0

L(60Hz) 106.15mH 113.86mH

Q (60 Hz) 0.4 0.4

L ( kHz) 100.22 113.92 mH

Q ( kHz) 2.7 6.9

'EQ-environmental qualification.

2.6 Other Methods Examined or

Proposed

[Degradations or mlfunctions addressed: mechanical

looseness or binding and valve leakage caused by worn,

degraded, or improper parts or the presence offoreign

materials inside the valve]

Four additional potentially useful monitoring and

diagnostic measurement techniques identified for study

midway through this investigation were

* analysis of the detailed time "signature" of inrushing

current that produces SOV actuation;

* detection of the presence of humming or chattering

of the plunger assembly within ac-powered SOVs

(indicative of mechanical wear) using either

frequency analysis of the coil current signal or an

acoustic sensor placed in close proximity to the

valve;

* verification of complete and normal SOV opening

and closure (including the detection of valve

leakage) based on the acoustic "signature" sensed by

a miniature microphone placed in close proximity to

the SOV; and

* inference of plunger position within the valve body

based on the magnetic field strength external to the

SOV as sensed by a Hall effect device placed on the

valve exterior.

In principle, these techniques should permit the detection

of worn or improper parts and extraneous materials, valve

misassembly, degraded elastomers, and misalignment of

the solenoid core. Although the purely electrical techniques

would require no special sensors to be fitted to the valve,

affixing an acoustic or magnetic sensor to the valve would

make possible as well the detection of through-valve and

external leakage of process fluid, the identification of

plunger mechanical problems, and positive verification of

plunger movement.

These first two measurement techniques were examined

briefly but were soon abandoned as unpromising. The

signature of the inrush current accompanying normal SOV

switch-on had been studied earlier by Meininger and Weir

with somewhat inconclusive results.16 Their data analysis

had focused almost exclusively on the frequency domain

(where they were unable to demonstrate marked sensitivity

to the presence of valve problems), so we naturally turned

to further exploration of waveform properties in time.

However, after acquiring the inrushing-current data shown

in Figure 2.16 (demonstrating by the variable impedance

properties of an ac-powered SOV that its plunger

movement is very rapid---complete in -20 ms following

the application of power-and therefore difficult to track

electrically) and performing a few additional

inrushing-current tests on dc-powered SOVs with no

positive results, it was decided not to pursue this idea

further.

39 NUREG/CR-4819, Vol.2
Evaluation of Diagnostic Methods Applicable to SOVs

The technique of identifying looseness and vibration by the

detection of perturbations to the coil current waveform

[presumed to result from the humming or chattering of

SOV internals at the frequency of the magnetization

reversals (60 Hz) or its harmonics] was also examined

without much success. Current waveforms typical of both

normal and loosely assembled (i.e., audibly buzzing)

valves are shown in Figure 2.27; very little difference can

be seen,* indicating that feedback of the audible

mechanical vibration into the time-dependent current

drawn by the solenoid is almost nonexistent. This

conclusion is reinforced by the spectra of these waveforms

(Figure 2.28)-both assembly conditions are characterized

by the presence of strong odd harmonics (n = 1,3,5 ... ) of

the 60-Hz fundamental but, except for a slightly higher

current amplitude** when the valve is correctly assembled,

the spectra are substantially identical-in particular, no

new frequencies [with the possible exception of the

appearance of even harmonics (n = 2,4.. .)] appear to arise

as a result of the buzzing internal parts. Given these

unpromising results, detailed analysis of coil current

waveform as a means of detecting looseness and vibration

in ac-powered SOVs was not pursued further. However,

this is not to imply that a miniature acoustic sensor placed

on or near the valve would have a similar difficulty in

detecting excessive internal vibration; in fact, judging from

the clearly audible low-frequency buzzing produced by

loosely assembled valves, such a locally positioned sensor

would probably be quite effective in detection of wear or

misassembly. Of course, the extra complexities of add-on

instrumentation may affect the practicality of such an

approach, but the basic principle seems to be sound.

The latter two measurement techniques listed at the

beginning of this section were never actually tried owing to

the pressure of time and programmatic limitations.

However, others have reported considerable success with

the use of a miniature microphone placed near the SOV to

sense, through differences in time- and frequency-domain

signatures, incomplete opening and closure of SOVs used

in a BWR control-rod-drive scram system. 17

It seems reasonable to assume that similar information

could be derived using a Hall effect device placed on or

near the SOV to detect magnetic field changes

accompanying correct execution of plunger stroke, as an

alternative to monitoring the acoustic signature

characteristic of correct valve opening and closure. Hall

effect devices are tiny and thereby provide good spatial

resolution but, being solid-state devices, they are not

known for their robustness in high-temperature,

radiation-containing environments. However, their

usefulness in this application merits study if further

research is performed in this area.

Finally, there would be added benefit if the findings of this

SOV investigation could be applied and utilized creatively,

wherever practical, in surveillance and testing programs

directed at safety-related electrical components of nuclear

power plants other than SOVs.

*The two waveforms have been time-shifted slightly to facilitate shape

comparison.

**This causes the mean value of the spectral ratio plotted in Figure

2.28(c) to be >1, which is consistent with the greater peak amplitude

seen in Figure 2.27.

NUREG/CR-4819, Vol. 2 40
Evaluation of Diagnostic Methods Applicable to SOVs

2

-j

z

u90

Cl: Er

w

0

LI

C',

z

.4:

z

cca:

1

0

-1

-2

0 5 10 15 20 25

TIME (ms)

Figure 2.27 Coil current time waveforms for a 60-Hz ac-powered SOV. (A: normally assembled, no audible

buzzing; B: loosely assembled, loud buzzing indicating vibration of internal parts. The waveforms are distorted

because an iron-cored current transformer was used to acquire the data without breaking into the circuit.)

41 NUREG/CR4819, Vol. 2
Evaluation of Diagnostic Methods Applicable to SOVs

ao-2 

10° -3' '

_a 4

10-5

3

2

2

2

1

a

a

0 .

0 .

0 50 100 150 200 250

FREQUENCY IN Hz

300 350 400

Figure 2.28 Coil current frequency spectra derived from Ihe waveforms shown in Fig. 2.27. (A: normally

assembled SOV, no audible buzzing; B: loosely assembled SOV, loud buzzing indicating vibration of internal parts;

C: line-by-line ratio, spectrum A divided by spectrum B.)

NUREGCR-4819, Vol. 2 42
Closing Remarks and Recommendations

3. CLOSING REMARKS AND RECOMMENDATIONS

Because our review of the technical literature did not

reveal any degradation-monitoring techniques (either

presently in widespread industrial use or under

development) that are oriented specifically to SOVs, this

Phase II study was necessarily more inventive than

evaluative in nature. However, the work has revealed a

number of SOV-monitoring methods based on electrical

properties that are potentially useful for ascertaining

operational readiness through the measurement of

degradation-sensitive performance parameters that may be

trended over component life. In so doing, early indications

of poor performance that may foreshadow more serious

malfunctions or failures of these devices can be obtained,

thus aiding in the scheduling of maintenance activities or

timely valve replacement.

The four techniques receiving the most attention were

* measurement of coil mean temperature during

operation by in situ measurement of coil electrical

dc resistance or ac impedance, combined with an

experimentally established temperature coefficient

of resistivity for the copper winding (i.e., the same

principle as a resistance thermometer);

* indication of plunger movement upon application of

ac power, plunger static position within the solenoid

coil, and freedom of plunger movement within the

guide tube by means of measurement of coil ac

impedance (inductance) changes resulting from

movement of the iron plunger relative to the coil;

* detection of mechanical binding by tracking changes

in the electric current (or voltage) at the SOV critical

bistable (pull-in and dropout) points, because these

define conditions of balance between electrical

(magnetic), spring (restoring), and friction forces

within the SOV assembly; and

* detection of shorted turns or insulation breakdown

within the solenoid coil based on the characteristics

of the "flyback" transient generated when a dc SOV

is deenergized abruptly.

Two other techniques were examined briefly and then

abandoned because initial results showed little promise for

eventual development into effective degradation detection

and diagnosis tools. They were

* detection of mechanical binding by analysis of the

detailed time "signature" of inrushing current that

accompanies initial plunger movement upon

application of electrical power to the SOV, and

* indication of mechanical looseness (resulting from

wear of internal valve parts, improper assembly, or

replacement with incorrect internal parts) by

electrical detection of the presence of humming or

wear-producing chattering of the plunger assembly

within ac-powered SOVs, using frequency-domain

analysis of the coil current signal.

Strengths shared by most of these monitoring techniques

include the following:

* The measured parameters have been demonstrated

to be sensitive to historically important modes and

causes of SOV failure.

* Most of the measurement techniques employed are

minimally disruptive to plant operations.

* The diagnostic tests can be performed at a location

remote from the SOV with no attendant degradation

in sensitivity.

* The SOV can often be made to serve as its own

sensor, thereby eliminating the need for additional

instrumentation and signal wires.

Weaknesses of the examined monitoring techniques must

also be acknowledged and include the following:

* Those techniques requiring a momentary change of

valve state or short-term substitution of a special

electrical power supply are disruptive to plant

operations.

* The monitoring techniques described may not cover

all conceivable types of aging-related degradation

and may not be universally applicable to SOVs of

all type, size, and construction.

Also, as a result of programmatic priorities and limited

resources, some planned tests whose results might have

yielded additional insights were never performed but are

43 NUREG/CR4819, Vol. 2
Closing Remarks and Recommendations

recommended for serious consideration in any future

investigation in this area. Among these are

* further study of the reproducibility of SOV pull-in

and dropout critical points and their sensitivity to

realistic, progressively worsening problems.

Compare baseline measurements with those

obtained after introducing different sources and

degrees of mechanical binding to the valve (using

thread sealant, lubricating compound used to aid the

assembly of internal valve parts, metal or paint

chips, etc., in combination with heat). Try

introducing internal contamination more naturally

by repeated cycling of valves held at elevated

temperature while connected to an oil-contaminated

compressed air supply.

* further study of the coil impedance change

accompanying plunger motion. Determine the

sensitivity with which the following can be verified

electrically: (1) obstructions (i.e., dirt, metal slivers

and paint chips) that restrict total plunger travel and

(2) permanently indented or chemically degraded

(i.e., gummy) elastometric valve seats.

A major criticism that can be leveled at the overall

investigation is that it does not go far enough-that is, that

the monitoring techniques studied were demonstrated only

in a controlled laboratory environment, using a small

population of unaged SOVs of substantially similar

construction and a limited number of implanted (rather

Than naturally occurring) defects. Accordingly, it is

recommended that in a future study,

performance-monitoring techniques be field-tested

using a larger population of both new and naturally

aged SOVs that would be likely to display one or

more varieties of degraded performance, and

techniques be refined and adapted as is necessary to

permit their use in a real plant environment (for

example, devise a means for applying a dc

interrogation signal to an ac-powered SOV so that

its temperature could be measured accurately even

in a normally energized state, or devise a means for

ascertaining free plunger movement upon SOV

turn-on without need for a special ramped-voltage

power supply).

Finally, there would be added benefit if the findings of this

SOV investigation could be applied and utilized creatively,

wherever practical, in surveillance and testing programs

directed at safety-related electrical components of nuclear

power plants other than SOVs.

NUREG/CR4819, Vol. 2 44
Publications Resulting from this Work

4. PUBLICATIONS RESULTING FROM THIS WORK

1. Kryter, R. C., "Assessment of Nonintrusive Methods for

Monitoring the Operational Readiness of

Solenoid-Operated Valves," Proceedings of the 16th

Water Reactor Safety Information Meeting, October

24-27, 1988, Gaithersburg, Maryland,

NUREG/CP-0097, Vol. 3, March 1989.

2. Kryter, R C., "Nonintrusive Methods for Monitoring the

Operational Readiness of Solenoid-Operated Valves,"

Nucl. Eng. Des. 118, 409-17 (1990).

3. Kryter, R C., "A Nonintrusive Method for Measuring the

Operating Temperature of a Solenoid-Operated Valve,"

p. -3-3-1 in Guidefor Monitoring Equipment

Environments During Nuclear Plant Operation,

NP-7399, proceedings of the EPRI Workshop on

Monitoring Equipment Environments During Nuclear

Plant Operation, April 10-11, 1990, Baltimore,

Maryland, Electric Power Research Institute, Palo Alto,

California, June 1991.

45 NUREG/CR-4819, Vol. 2
References

REFERENCES

1. Bacanskas, V. P., G. C. Roberts, and G. J. Toman, "Aging

and Service Wear of Solenoid-Operated Valves Used in

Safety Systems of Nuclear Power Plants, Vol. 1:

Operating Experience and Failure Identifications,"

NUREG/CR-4819, Vol. 1, March 1987.

2. Ornstein, H. L., "Operating Experience Feedback

Report-Solenoid-Operated Valve Problems,"

NUREG-1275, Vol. 6, February 1991.

3. Verna, B. J., "Commendable Operator Performance and

Solenoid-Operated Valve Problems," Nucl. News,

32-33 (July 1991).

4. Ref. 2, Sect. 8.1, p. 32, and Sect. 8.2, p. 32.

5. Institute of Electrical and Electronics Engineers,

"Evaluation of Maintenance and Related Practices for

Solenoid Operated Valves in Nuclear Power

Generating Stations," 89TH0248-5 PWR, New York,

March 1991.

6. Arrhenius, S., Z. Physik. Chem., 4, 226 (1899). Available

in public technical libraries. Simplified descriptions of

the so-called Arrhenius Law may be found in most

university-level physical chemistry textbooks (e.g.,

Laidler, K. J., Chemical Kinetics, 2nd ed.,

McGraw-Hill, New York, 1965, pp. 50-54; also Frost,

A. A., and R. G. Pearson, Kinetics and Mechanism, 2nd

ed., Wiley, New York, 1961, pp. 22-25).

7. Institute of Electrical and Electronics Engineers, Inc.,

"IEEE Standard for Qualifying Class IE Equipment for

Nuclear Power Generating Stations," IEEE 323-1974,

New York, September 1983.

8. Institute of Electrical and Electronics Engineers, Inc.,

"IEEEF Standard for Qualification of Safety-Related

Valve Actuators," EEE 382-1985, New York, 1985.

9. Bacanskas, V. P., G. J. Toman, and S. P. Carfagno,

"Aging and Qualification Research on Solenoid

Operated Valves," NUREG/CR-5141, August 1988,

Sect. 6.4, p. 31 .

10. Chemical Rubber Publishing Company, Handbook of

Chemistry and Physics, 37th ed., Cleveland, Ohio,

Sept. 1, 1955, pp. 2355-61.

11. Henney, K., ed., Radio Engineering Handbook, 5th ed.,

McGraw-Hill, New York, 1959, Chap. 5, p. 11.

12. Ref. 2, Sects. 4, 5.2.3, and 5.2.4.

13. Ref. 11, Chap. 1, p. 105.

14. George C. Marshall Space Flight Center, "Pulse coil

tester," MFS-29301, NASA Tech Briefs, Huntsville,

Alabama, March 1988.

15. Ref. 11, Chap. 5, pp. 18-21.

16. Meininger, R. D., and T. J. Weir, "Development of a

Testing and Analysis Methodology to Determine the

Functional Condition of Solenoid-Operated Valves,"

NUREG/CR-5008, September 1987.

17. McElroy, J. W., "BWR Control Rod Drive Scram Pilot

Valve Monitoring System," Proceedings of EPRI

Conference 'Incipient Failure Detection for Power

Plants,' October 10-12, 1984, Orlando, Florida,

Electric Power Research Institute, Palo Alto, California.

47 NUREG/CR-4819, Vol. 2
---

Appendix A

APPENDIX A

DC RESISTANCES OF NINE TEST VALVES OVER THE TEMPERATURE

RANGE OF 20 TO 165 0C

The nine fuly assembled SOVs used in this study

(Table 2.1) were placed simultaneously in a

thermostatically controlled, electrically heated oven whose

interior temperature could be held constant (to within

-±10C) at any desired value from ambient to 1650C, the

highest temperature we believe likely to be achieved by

SOV self-heating alone (i.e., ignoring the possibility of fire

or a steam leak). The electrical leads from the solenoids

were passed through a hole at the top of the oven, which

was then sealed with thermal insulation. The oven

temperature was read externally from a mercury-filled

capillary stem thermometer protruding into the oven cavity

and in close proximity to the valves.

After each temperature change, resistance measurements

were delayed for a minimum of one hour in order to ensure

isothermal conditions within the valve bodies and coils.

Measurements were made with a digital electronic

ohmmeter of -0.1% repeatability that employed a low

excitation level (1.5 V) so as to minimize heat deposition

within the coil as a result of the measurement itself, which

required -3 s to perform on each valve.

The test results are shown in Figures A.1 through A.9,

where coil dc resistance is plotted against oven (and

therefore coil) temperature on linear scales.

The following observations and conclusions were drawn

from these measurements:

1. The variation of dc resistance, Rdc, with temperature over

the range of interest to this study (20 to 165C) is

almost exactly linear (i.e., the coefficient of correlation

= 0.9997 or better for all nine SOVs). Thus, Rdc is a

highly accurate indicator of coil mean temperature, as

long as the resistivity of the copper winding remains

constant with thermal cycling and the passage of time.

2. The magnitude of the temperature coefficient of coil

resistance is sufficiently large to permit accurate

temperature inference using resistance-measuring

equipment of only modest accuracy:

(change in Rdc)PC 0.3 - 0.4% at 20'C

(change in Rdc)/C | 0.27 - 0.35% at 70'C

Thus, if measurements were to be made with

industrial-grade equipment accurate to within only 2 to

3% of reading, inferred coil temperature would

nonetheless be accurate to about ±100 C (assuming that

the temperature coefficient of resistance is known to

somewhat greater accuracy, presumably through

laboratory measurements). Such temperature accuracy

is probably quite sufficient for purposes of component

life prediction.

3. The measured value of the temperature coefficient of

resistance compares favorably with the magnitude

predicted by typical handbook values of p and a for

annealed copper I

Resistivity (p) of pure annealed Cu at 200C = 1.692 E6

D-cm

Temperature coefficient of resistivity (a) of pure annealed

Cu over the range 0 to 1000C = 0.00433 E-6 Q cmPC.

0,00433 Therefore, (AP/p) 6 = 0.256%/OC at 200C.

REFERENCE

1. Chemical Rubber Publishing Company, Handbook of

Chemistry and Physics, 37th ed., Cleveland, Ohio,

Sept. 1, 1955, pp. 2355-61.

A-1 NUREG/CR-4819, Vol. 2
Appendix A

a

w

0

z

co

U)

w

0

1300

1200

1100

1000

900

800

700

20 40 60 80 100 120 140 160

COIL TEMPERATURE (0C)

Figure A.1 Dc resistance vs temperature: SOV "A."

a

w

C

z

1-

co

F5

C.)

C.,

1300

1200

1100

1000

900

800

700

20 40 60 80 100 120 140 160

COIL TEMPERATURE (C)

Figure A.2 Dc resistance vs temperature: SOV "B."

NUREG/CR-4819, Vol. 2 A-2
Appendix A

C

w

u,

z

CO)

w

EC

0

1300

1200

1100

1000

900

800

700

20 40 60 80 100 120 140 160

COIL TEMPERATURE (C)

Figure A.3 D resistance vs temperature: SOV "C."

a

w

z

Co

w

CO0u

220

210

200

190

180

170

160

150

140

130

20 40 60 80 100 120 140 160

COIL TEMPERATURE (C)

Figure A.4 Dc resistance vs temperature: SOV "D."

A-3 NUREG/CR4819, Vol. 2
Appendix A

220

210

_200

a

at 190

0

z

< 180

c5 170

= 160 0

150

140

130

20 40 60 80 100 120 140 160

COIL TEMPERATURE (C)

Figure A.5 Dc resistance vs temperature: SOV SE."

220

210

200

Lw 190

C.)

z

< 180

I6i 170

'C 160 0

0

150

140

130

20 40 60 80 100 120 140 160

COIL TEMPERATURE (C)

Figure A.6 D resistance vs temperature: SOV "F."

NUREG/CR-4819, Vol. 2 A-4
Appendix A

85

80

a

w

0

z

Ce

111 w

C,

0

75

70

65

55

5D)

20 40 60 80 100 120 140 160

COIL TEMPERATURE (

0C)

Figure A.7 Dc resistance vs temperature: SOV "G."

80

75

C

w

0

z

Cf,

w

cc

0

u0

70

65

60

55

50

20 40 60 80 100 120 140 160

COIL TEMPERATURE (C)

Figure A.8 D resistance vs temperature: SOV "H."

A-5 NUREG/CR-4819, Vol. 2
Appendix A

2300 X l l-l T l

2100

w

o 1900

z

w 1700

1500

1300 .

20 40 60 80 100 120

COIL TEMPERATURE (C)

Figure A.9 Dc resistance vs temperature: SOV "L"

140 160

NUREG/CR4819, Vol. 2 A-6
Appendix B

APPENDIX B

AC IMPEDANCES OF FIVE SELECTED TEST VALVES OVER THE

TEMPERATURE RANGE OF 20 TO 1550C AT THREE WIDELY SEPARATED

FREQUENCIES: 5, 60, AND 500 Hz

Using exactly the same setup and procedures as described

in Appendix A for dc resistance measurements, ac

impedances were measured for five fully assembled SOVs

(i.e., "A", "C", E", "G', and "I") as a function of

temperature at three fixed excitation frequencies.* To

minimize self-heating, the measurements were performed

at a low level of excitation (1.1 Vs) using a

Hewlett-Packard model 4192A low-frequency impedance

analyzer. The five SOVs were selected to be representative

of various types of physical construction (e.g., solenoid

coil, valve body, plunger/guide tube arrangement) and

electrical design (i.e., ac and dc operation).

The results of these tests are shown in Figures B.1 through

B.15, where the real (in-phase), imaginary (quadrature),

and vector magnitude (absolute value) representations of

the complex impedance are plotted against oven (and

therefore coil) temperature on linear scales. Also shown for

reference is the dc resistance (shown by a dashed line),

which should be equivalent to the ac impedance

extrapolated to zero frequency.

The following conclusions and observations were drawn

from these data:

1. As would be expected, the complex impedance at an

excitation frequency of 5 Hz is nearly all real [i.e., the

imaginary component, Im(Z), is comparatively small],

and the real component, Re(Z), is very nearly equal to

Rdc. Also, both Re(Z) and R& display the same

temperature coefficient (i.e., the same magnitude of

positive slope with increasing temperature). In addition,

the imaginary component of the impedance is

essentially independent of temperature at 5 Hz.

2. Both Re(Z) and I Z I display highly linear dependencies

on temperature. At an excitation frequency of 5 Hz,

they also have essentially the same slope.

*With the exception of 60 Hz-which was selected, of course, because it

is the frequency of the power mains throughout the United States-the

selection of the other two test frequencies was arbitrary. However, 5 and

500 Hz represent reasonable lower and upper limits for test signals that

would be practical to apply to plant wiring.

3. Coils designed for 125 Vdc operation (i.e., "A", "C", and

ar) display quite similar variation of impedance with

temperature and frequency as those designed for

117 Vac operation (i.e., "E' and "G), although the

actual impedances are much greater for the dc coils,

because they are wound with many more turns of wire

than the coils intended for operation at 60 Hz ac.

4. Re(Z) is -50% larger than Rd at 60-Hz excitation

frequency (and 5 to 10 times larger at 500 Hz) but

continues to show excellent linearity with temperature

and almost as large a temperature coefficient (expressed

as ohms per degree C) as Rdc. However, since Re(Z) >

Rdc at frequencies higher than 60 Hz, the percentage

change in value per degree C (which provides a

yardstick of difficulty in performing a measurement) is

less (namely, -0. Ito 0.3%/PC) for Re(Z) at these

frequencies than for Rdc, although still in the range of

field measurement practicability.

5. At excitation frequencies of 60 and 500 Hz, Im(Z) is no

longer independent of temperature, but the magnitude

of the temperature coefficient of this quadrature

component of impedance is significantly smaller than

the temperature coefficients associated with Rdc and

Re(Z).

6. As a result of observations 4 and 5 above, the change

(expressed as a percent of value) per degree C is

considerably smaller at 60 or 500 Hz for I Z I than for

either Re(Z) or Rdc. This is unfortunate from a

standpoint of field measurement, because I Z I is the

quantity most easily measured, being merely the ratio

of the root-mean-square voltage applied to the coil

(Vlms) to the root-mean-square current drawn by the

coil (Inns). Therefore, obtaining a temperature estimate

accurate to ±100C would require measurements of Vns

and Inns (yielding I Z ) to accuracies considerably

better than ±1% if the coil's temperature coefficient is

only 0.1% of value per degree C. Note that such

measurement accuracy is 2 to 4 times greater than

would be required of an Rdc measurement yielding the

same ±100C uncertainty in coil temperature, because of

the larger temperature coefficient associated with Rdc.

B-1 NUREG/CR4819, Vol. 2
Appendix B

+ DC

Resistance

A Rs(Z)

at 5 Hz

A Im(Z)

at 5 Hz

* IZI

at 5 Hz

1200

C3

z

tLa

C.)

z

Ico

lU

Ci,

w

cc

1000

800

600

400

200

20 40 60 80 100 120 140 160

COIL TEMPERATURE (C)

Figure B.l Dc resistance, 5-Hz impedance vs temperature: SOV "A."

+ DS

Resistance

A Re(Z) A Im(ZI

at oo Hz at Hz

* IzI

at 60 Hz

3200

wC

w

0

-

0

w

l

0

w

0

U7

i<

a:

2800

2400

2000 F

Siope~3 06 7 

Slope=1.94 0 0/C (0082 %/°C at 200C) A

0224 6

0 .375 /' t+ -

20 40 60 80 100 120 140 160

1600

1200

800

COIL TEMPERATURE (C)

Figure B.2 Dc resistance, 60-Hz impedance vs temperature: SOV "A."

NUREG/CR-4819, Vol. 2 B-2
Appendix B

+ DC

Rolaltin.c

As(Z)

at 500 Hz A 50 zJ

at soo HZ .

16

lZI

at s0o Hz

C

0-

z

a 0

a D

V )

LU F

s -

Ct

i1 2

0

'00 120 140 1hin

Figu 3 c D B M i anc , - .... R~ F A IU R E (

0C)

Fgr11+ ceiae SOOsHz impedance vs temperature: SOt "A."

+ DC A -

RlsirttnC9 A Fla (Z)

at Hz

A Im(ZJ

at 5 Hz

.-1

a

C

lu

IL

X

0

Lu

z

i!CO2i

.4

I1250

1050

* az

at 

8 HZ~

850

650,

450

250

80 100 ~120 140 160

COIL TEMPERATURE (C)

resistance, S-nz impedance Vs temperature: SOV "C."

B-3 .REGCR4819, Vo. 2
Appendix B

+ DC A ROeZ)

Reaistanc, at 00 Hz

A Ir(Z)

at e Hz

0 iZI

at 60Hz

a

w

z

L

4

-

CI0

w

C.)

z

4i

C,,

C,,

w

3200

2800

2400

/~G (0.09 0Io C at 200) 

sope22 3 45 Aoi

Sope2O 0I 002 //Ca

-pe 2 6 n

I - , I . , I , I , I a , 0I

2000 I

1600

1200

800

20 40 60 80 100 120 140 160

COIL TEMPERATURE (C)

Figure B.5 Dc resistance, 60-Hz impedance vs temperature: SOV "C."

+ DC

Resostance

A RG(Z)

at 500 Hz

A Irn (Z)

at 500 Hz

0 hZI

at 500 Hz

16

a

wlU

z

IUz

Z

a

I.-i

w. _

14

12

10

8

6

4

2

0

20 40 60 80 100 120 140 160

COIL TEMPERATURE (C)

Figure B.6 Dc resistance, 500-Hz impedance vs temperature: SOV "C."

NUREG/CR4819, Vol. 2 B-4
Appendix B

+ DC

ReSistance

A RO(Z)

at 5 Hz

A Im(Z)

at 5 Hz

* IZI

at 5 Hz

220

wC,

z

0

0

w

c.

z

0

IwuX

200

180

160

140

120

100

80

60

40

20

20 40 60 80 100 120 140 160

COIL TEMPERATURE (C)

Figure B.7 Dc resistance, 5-Hz impedance vs temperature: SOV "E."

+ DC

Reslatance

A RG(Z)

at 80Hz

A tm(Z)

at 50 Hz

* IZI

at 50 Hz

450

0

w

z

C

w

a

I0

w

C.

z

tU,

a:l

400

350 

300

Sl~ope-O~ 00

Slope=0.17 4 5 0/0 (0.058 %/oC at 200C)

~~~~~~- - - +- Ct 9 * / A 

- . op .S .1 t ,.,. at 20'C)-

-I- - ~~~~pe~0.5? 9 0/00a

+ I-A Slp =05 .(I o * I I *

250 I

200

150

100

20 40 60 80 100 120 140 160

COIL TEMPERATURE (C)

Figure B.8 Dc resistance, 60-Hz impedance vs temperature: SOV "E."

B-5 NllREG/CR-4819, Vol. 2
Appendix B

+ DZ

Floefstanoo

A Re(Z)

at 500 Hz

A Im(Z)

of 500 Hz

* IZI

at 500 Hz

2000

w

C

z

a

Ul

a.

0

.

a

FcoLU

1800

1600

1400

1200

1000

800

600

400

200

20 40 60 80 100 120 140 160

COIL TEMPERATURE (C)

Figure B.9 Dc resistance, 500-Hz impedance vs temperature: SOV "E"

+ DC

Resistance

A Re(Z)

st5 Hz

A im(Z)

*t 5 Hz

0 IZI

*t 5 Hz

90

a

U'

C.,

z

CZ

U'

I0

CO

w

c

80

70

60

50

40

30

20

10

20 40 60 80 100 120 140 160

COIL TEMPERATURE (C)

Figure B.10 Dc resistance, S-Hz impedance vs temperature: SOV "."

NUREG/CR4819, Vol. 2 B-6
+ De 

Appendix B

R sac a (z) A jm(Z)

t so -* _ atOfz )Z

SUope=0.1725 /C (0.086 %/oC at 20°C)

a t0Sope=0.1091 

/C (0.059 0/C at 200C)

w0b 160

120 

-50 t Slope=0.1788 (lIC (0.241 %C at 20°C)

40 Slope-0.1914 /CC (0.360 %/oC at 20°C)

20 40 60 80 100 120 140 160

FigUre B.11 DCCOIL TEMPERATURE (C)

Fiue 11 De resistnce 60-Hz irape~lnce Vs temperatiure: S y".

ResistnceRn(zj 1 4 0 n e 0sb at 5 00 H z A tIm (Z ) IZ

400 

a *t5 00 HZ

1200 __. _. . _.. _- - _- - _* ___ _* _* _

0 100 A A A * ^. 

1 t S(ope'=1.0794 0/oC (0.102 %I'C at 20C)

LU(.> 800

IU 600 ( Siope=03 8 4 4 l/oC (0.077 %/OC at 20C C)

1- 400

cc 200

20 40 60 8 10 12 140 6

COIL TEMPERATURe (C)16

Figure B.12 Ilk reistanCe,500-)j) 'lPedance vs tewperature: SOV,"G."

B-7 NUREG/CR4819 Vol. 2
Appendix B

+ DC

Reaslitance

A Re(Z)

at 5 Hz

A Im(Z)

at 5 Hz

* IZI

at 5 Hz

2200

a

U

C.)

z

1-

co

Ul

IC

2000

1800

1600

1400

1200

1000

800

600

400

20 40 60 80 100 120 140 160

COIL TEMPERATURE (C)

Figure B.13 Dc resistance, 5-Hz impedance vs temperature: SOV 'q."

+ DC

Resistance

A Re(Z)

at 80Hz

A Im(Z)

at 80 Hz

a IZI

at 6 Hz

l~~~~~~~~V -

5400

z

a

LU

-

0

Lu

z

Lu

cc

4600

3800 

slp 5-( l ' l i - - -- - - r

at 7 0 C 1

Slope=3.054 O /-C

Slope~ k n

3000 

2200 e -4-p +-- 

( 0 .3 _-

_ - - -4 - -

- 4 

1400 - +- I I I I I

20 40 60 80 100 120 140 160

COIL TEMPERATURE (C)

Figure B.14 Dc resistance, 60-Hz impedance vs temperature: SOV "L'

NUREG/CR-4819, Vol. 2 B-8
Appendix B

+ DC

lhuRaencs

A R&(2)

at 500 Hz

A Im(Z)

at 500 Hz

* IZI

at 600 Hz

28

aDW-fl

(n

S-] o0

wz

2

cUl w

25

22

19

16

13

10

Slope-2 1.56 Q/0C (0.089 %/cC at 200C) -

Slope=2 1.56 0/00 (0.089 %lSC at 2000)

- p 

p p

Slope=18.1 0/CC (0.085 %C at 20°C)

Slope=11.86 0/0 (0.104 %1tC at 2000I

A .^.e._ .h_ _ 

7

4

1

Sope=5.652 W/OC (0.426 %/GC at 20C)

------- r--4- -------------------------

A 6 Q . 20 4(-

vv ou 100 120 140 160

COIL TEMPERATURE (C)

Fgure 11.15 De resstace, S0Hz impedance vs temperature: SOV 'ri."

B1-9 NUREG/CR4819, Vol. 2
Appendix C

APPENDIX C

AC INDUCTANCES (L) AND QUALITY FACTORS (Q) OF FIVE SELECTED

TEST VALVES OVER THE TEMPERATURE RANGE OF 20 TO 1550C AT

THREE WIDELY SEPARATED FREQUENCIES: 5,60, AND 500 HZ

As an adjunct to the measurement of the ac impedances of

five of the fully assembled test SOVs (i.e., "A", "C', 'E",

"G", and "I') as a function of temperature at three fixed

excitation frequencies (Appendix B), inductances and

quality factors (defined as the ratio of inductive reactance

to effective resistance) were also measured at a low level of

excitation (1.1 Vnns) using a different operating mode of

the same impedance analyzer. These measurements were

performed primarily to see if either L or Q were quantities

suitable for ascertaining coil temperature. The test

procedures were the same as described in Appendices A

and B, and consisted of placing the assembled SOVs in a

temperature-controlled oven.

The test results are displayed in Figures C.1 through C.10,

where the real quantities L and Q are plotted against oven

(and therefore coil) temperature on linear scales. The three

excitation frequencies chosen were again 5, 60, and

500 Hz.*

The following conclusions and observations were drawn

from these data:

1. Inductance of the fully assembled SOVs increases quite

linearly with temperature over the range of interest (20

to 155°C), but the temperature coefficient is not large

when expressed as a percentage of value, being

typically 0.03 to 0.13% of value per degree C. This

number is a factor of 2 to 4 smaller than the

percent-of-value temperature coefficients typically

measured for Re(Z) or Rdc (see Appendix B).

2. The temperature coefficient of L for these SOVs,

expressed as a percent change of value per degree C, is

larger at higher than at lower frequencies. This is

attributable to the diminished value of the inductance at

higher excitation frequencies, coupled with an absolute

slope (mH per degree C) that varies little with

frequency.

With the exception of 60 Hz-which was selected, of course, because it

is the frequency of the power mains throughout the United States-the

selection of the other two test frequencies was arbitrary. However, 5 and

500 Hz represent reasonable lower and upper limits for test signals that

would be practical to apply to plant wiring.

3. Inductance typically drops by almost a factor of 2 as the

excitation frequency rises from 5 to 500 Hz, but

solenoid coils designed for ac operation (i.e., "E' and

"G) retain a larger fraction of their low-frequency

inductances at 60 Hz than do coils intended for dc

operation (i.e., "A", "C", and "I').

4. SOVs intended for dc excitation typically have ten to

twenty times the inductance at any frequency of interest

to this study (i.e., <500 Hz) compared with SOVs

intended for ac excitation at about the same voltage.

This is attributable to the far greater number of turns of

wire in the coils designed for dc operation as compared

with their ac counterparts.

5. Quality factor decreases slightly with increasing

temperature for most of the SOVs tested (valve "G" at

500 Hz excitation is an exception), but the linearity of

the relationship is not exceptionally good. The Q of all

the solenoids shows a marked increase with frequency

of excitation (particularly in the intervals of 5 to

60 Hz); this is attributable to the rapid increase of

inductive reactance with frequency (see Appendix D)

compared with the much slower increase of effective

resistance with frequency.

6. Inductance is a much better indicator of solenoid

temperature than quality factor, but as a result of the

relatively small magnitude of the temperature

coefficient of inductance and the relatively poor

linearity of the Q vs temperature relationship, neither L

nor Q appears to be as well suited to temperature

indication as Rdc or Re(Z). However, inductance

measured at 500-Hz excitation frequency may be at

least as good an indicator of coil temperature as the

absolute value of the impedance, I Z I, although the

instrumentation necessary to perform the L

measurement is considerably more complicated than

the simple Vrus and Irms measurements needed to

compute I Z I at 60 Hz. Also, measurement of L at

500 Hz would require disruption of normal 60-Hz

power to the SOV, whereas the measurement of I Z

can be accomplished nonperturbatively.

C-1 NUREG/CR-4819, Vol. 2
Appendix C

* Inductance

at 5 Hz

A Inductance

at 80 Hz

El Inductance

at 500 Hz

U.

C

2

.4

IC

2

8.5 i . .* , I 

7.5

Slope=0.00396 HC (0.057 %t°C at 200C)

6.5 _ 20 C

Slope=0.00513 H 0 .8 /0a 00

5.5

4.5 . Sloe=0.0427H/ 0

C (0.109 %/C at 20'C)n

3.5 f .

20 40 60 80 100 120 140 160

COIL TEMPERATURE (C)

Figure C.1 Inductance vs temperature at three excitation frequencies: SOV "A."

0 Inductunce

at8 Hz

A Inductance

at M0 Hz

0 Inductance

at 500 Hz

8.5

7.5

LU

0C,

z

i T I I I I I . 7 - . I I I

Sl~ope0.0031 5 HC (0.041 %/°C at 200C)

=eS==~- 't20-C

0 ~icpe0.O059 6.5 2 HC 0.02/o0a

5.5

4.5

3.5

Sb e004B HIG (0o.132 % a 20C0l

I -i . . - . . I .I . , , . L . * 1 4 0 1 6 0

20 40 60 80 100 120 140 160

COIL TEMPERATURE (C)

Figure C.2 Inductance vs temperature at three excitation frequencies: SOV "C."

NUREG/CR-4819, Vol. 2 C-2
Appendix C

* ual. Factor

at S Hz

Oual. Factor

at O Hz

0 Qual. Factor

at 500 Hz

2.5

2.0

a

0

I.- 1.5

<t:

L

1.0

a

0.5

0.0

Figure C.'

2.5

2.0

3

- 1.5

L

'C

1.0

0.5

r I I I I I 

13 ~~~AE3E3 E 5 E a E _

I 0- S *_~~~~~~ * _ . a__* 

. . . . . .

. . . , . , . . . . , .

20 40 60 80 100 120 140 160

COIL TEMPERATURE (C)

7 Quality factor vs temperature at three excitation frequencies: SOV "C."

* Qust. Factor A Ouao. Factor O Ouao. Factor

at 5 Hz at O Hz at 500 Hz

I

I

I

I

I

i

I

LI A~~~~~~~~~~~~~~~~~~~~~~~~~~~~~~

. .3 . El . . . . .

A _

S Pe W.I 002 P err°C

. . _g *_ p * __ .* 

/.fI

%J.J

20 40 60 80 100 120 140 160

COIL TEMPERATURE (C)

Figure C.8 Quality factor vs temperature at three excitation frequencies: SOV "E."

C-S NUREG/CR-4819, Vol. 2
Appendix C

0 Qual. Factor

ot 5 Hz

A Qual. Factor

lot 0 Hz

El Qual. Factor

at 500 Hz

3.0

a

0

0

C.

-J

M

a

2.0

1.0

I I I . I I I I

SIp3.c 03 per °C

0 n

6

I~~~~~~

0.0

20 40 60 80 100 120 140 160

COIL TEMPERATURE (C)

Figure C.9 Quality factor vs temperature at three excitation frequencies: SOV 'G."

* Qual. Factor

at 5Hz

A Cual. Factor

gkt 0 Hz

i1 Qual. Factor

at 500 Hz

2.5

2.0 

a

0

0

U.

a

1.5

1.0

I I ~~~~ I ~~~I I I I I

- I I I I -

Slopse.o. 0 Per 02 

_ _ w_ @ @ 

0.5

0.0

20 40 60 80 100 120 140 160

COIL TEMPERATURE ('C)

Figure C.10 Quality factor vs temperature at three excitation frequencies: SOV "1."

NUREG/CR-4819, Vol. 2 C-6
Appendix C

* Inductance

at a Hz

A Inductance

at 60Hz

0 Inductance

at 500 Hz

X

E

LU

F0

z

950

Sope=0.2 815 H/oc (0.033 %/OC at 200C)

5 =ope=0.463 7 (0.05a

750

650

550 W C

Slope-'0.5 5 3 MdHiIC (0.1%ICa20)

450 I

20 40 60 80 100 120 140 160

COIL TEMPERATURE (C)

Figure C.3 Inductance vs temperature at three excitation frequencies: SOV "E."

* Inductance

at 8 Hz

A Inductance

at Hz

Indutanc

at 500 Hi

X

-E

C)

D

0

600 , . , . , * 

- =~~o pe 024 8

7 rr /OC (0.049 %/ b at 20 0

500 Siop~028 O rrJ~40C (.05 tO/ 00Ca 200

400

Sjp =03440 r-I 0 (0 0 1 0 a

300 * 

20 40 60 80 100 120 140 160

COIL TEMPERATURE (0C)

Figure C.4 Inductance vs temperature at three excitation frequencies: SOV "G."

C-3 NUREG/CR-4819, Vol. 2
Appendix C

* Itlducta,)c

at H Z2

a fnductnoo ,

ot 60 Hz

o Induct,,c.

At "a0 Hz

13.0

12.0

X

w

z

C)

a

1

10.0

9.01

8.0

7.0

Figure C.5

100 120 140 160

COIL TEMPERAritmr ..-.

.

Iductance Vs temperature at three excitation frequencies SO;,"L"

Quag. Factof A _

of a HZ

4U Qual. FmCo4

at 4o Hz

2.5

2.0

1.5 

o QulJ. Factor

at 500 Hz

0

i.

a:

I

z

1.0

0.5

0.0

100 120 140 1 6Fwr

COIL TIEFAPERATUAIE (C) Fgure C.6 Quality factor vs ternIerature at three excitaton 

NUREG/CR-4j 19, Vol. 2 C-4
Appendix D

APPENDIX D

FREQUENCY DEPENDENCE OF IMPEDANCE FOR DC- AND

AC-POWERED SOLENOID COILS

As a follow-up to the unexpected finding (see Appendix B)

that Re(Z) is -50% larger than Rdc at 60-Hz excitation

frequency (and 5 to 10 times larger at 500 Hz), a more

detailed investigation of the frequency dependence of

impedance and inductance was conducted for a few

selected solenoid coils. All measurements were made at

room temperature (23TC) and at a low level of excitation

(1.1 Vnmn), using a Hewlett-Packard model 4192A

low-frequency impedance analyzer, so the SOVs were in

their rest (unpowered) state. SOVs "A", "D", and "G" were

tested in a normally assembled condition, whereas the

solenoid coils of SOVs "C', "E', "G", and "r' were

measured bare (i.e., removed from the SOV housing and

devoid of plunger and guide tube).

The measurement results are shown in Figures D. 1 through

D.5, where the real (in-phase), imaginary

(quadrature-phase), and vector sum (absolute value) of

impedance, i Z , and the inductance, L, are plotted against

frequency on linear scales. Unless otherwise stated, the

solid curves have no theoretical significance; their function

is merely to join the measured data points smoothly so that

the trend is easily seen.

The following observations and conclusions were drawn

from these measurements:

1. Fully assembled SOVs designed for dc operation (i.e.,

"A", "C', and I') have impedance vs frequency

characteristics that are similar to SOVs designed for ac

operation (i.e., "D", "E"', and "G"), even though the

impedance is typically ten times greater for the dc coils

as a result of the larger number of turns used in winding

the dc solenoids.

2. The real (in-phase) component of the complex solenoid

impedance is not independent of frequency (as one

might expect for a nearly ideal inductor-that is, an

ideal inductor plus some series resistance) and is

always greater than the dc resistance, although Re(Z)

approaches RdC as the frequency of excitation

approaches zero. This was at first thought to be

attributable to "skin effect," but data obtained on bare

coils (Figures D.4 and D.5) later revealed that the noted

behavior is due to a combination of eddy current losses

in the copper winding and hysteresis losses in the iron

plunger and guide tube present in the fully assembled

SOV.

3. The imaginary (reactive) component of the complex

solenoid impedance does not grow linearly with

frequency, as one would expect for an ideal inductor. It

is thought that the downward curvature (negative

second derivative with respect to frequency) seen in

Figures D. through D.3 is attributable to the presence

of distributed stray parallel (i.e., turn-to-turn)

capacitance, the reactance of which tends to partially

cancel the inductive reactance of opposite algebraic

sign.

4. As a result of observations 2 and 3 above, the magnitude

of the complex impedance of fully assembled SOVs is

seen to increase fairly linearly with frequency over the

range of 0 to 120 Hz, although this relationship is of no

particular value to the study at hand.

5. The inductances of the bare solenoid coils (Figures D.4

and D.5) were found to be almost independent of

excitation frequency (the expclted result for an

air-cored coil), whereas I Z for the bare coils plots

nicely as a second-orderpolynomial in frequency-that

is, I = ao + alf+ af . The existence of thef 2 term

strongly suggests that eddy-current losses in the copper

coil winding are responsible for this nonideal behavior,

since such losses are known to grow as the second

power of frequency.

6. The presence of permeable iron within the bare solenoid

coil significantly increases its impedance (e.g., at

200-Hz excitation frequency, I Z I of the bare coil "G"

was found to increase from 104 Q to 575 Q as a result

of threading the coil onto the guide tube/plunger

assembly).

REFERENCE

1. Webber, F. G., and R. L. Sanford, "Inductance; Magnetic

Materials," p. 3-2 in Henney, K, ed., Radio

Engineering Handbook, 5th ed., McGraw-Hill, New

York, 1959.

D-1 NUREG/CR-4819, Vol. 2
Appendix D

+ Re(Z) A Im(Z) 0 lZ

a

0

z

D

a- Cl

7000 - -

6000

5000

4000

3000

2000

1000

0 - - I

0 30 60 90 120 150 180 210

EXCITATION FREQUENCY (Hz)

Figure D.1 Variation of impedance components with excitation frequency: SOV "A."

+ Re(Z) A Im(Z) 0 lZI

1000

800

. . . . . .. . . . . I .i . . . .

a

0

uLJ z

a- us

600

400

200

0 . . . . . . . . . . . . . . . . . . .

0 30 60 90 120 150 180 210

EXCITATION FREQUENCY (Hz)

Figure D.2 Variation of impedance components with excitation frequency: SOV "D."

NUREG/CR-4819, Vol. 2 D-2
Appendix D

+ Re(Z) 1i Im(Z) 0 JZI

600 Ii . . . . . . . . . . . . .

500 -

w

0

w

400

300

200

100 

0

0 30 60 90 120 150 180 210

EXCITATION FREQUENCY (Hz)

Figure D.3 Variation of impedance components with excitation frequency: SOV "G."

+ L A ZI

coil C coil C

O L,

coil I

* IZI,

coil I

2.0 I . I . . I I . I I I . . . .

1.5 

M=

w

0

z

0

0

Z

.-

3000

2000 a

w

0

z

0

w

a.

1000 2

1.0

0.5

..0 __

I~~~~~- 

^- 2 The two dashed curves are

second-order polynomials

in frequency. i.e.,

IZI=a0 +af+ a21

2

I . . I . . I . . I . . I . I

0.0 0

0 30 60 90 120 150 180 210

EXCITATION FREQUENCY (Hz)

Figure D.4 Variation of inductance and impedance with frequency: bare solenoid coils "C" and "."

D-3 NUREG/CR4819, Vol. 2
Appendix D

+ L A IZI o L,

coil G

° IzI,

coil E coil E coil G

0.15

0.12

I . I I I , . . I I I I

w

0

z

0

0

z

L6- ----------------

. c3 n B~~~~~~~~-

0.09

0.06

250

200

a

150 w

0

a

100 a

50

0

0.03

0.00

The two dashed curves are

second-order polynomials

in frequency, i.e.,

IZI =aO+alf +a2 f

2

..- I ................ I I .. 

0 30 60 90 120 150 180 210

EXCITATION FREQUENCY (Hz)

Figure D-5 Variation of inductance and impedance with frequency: bare solenoid coils "E" and 'G."

NUREG/CR-4819, Vol. 2 D-4
Appendix E

APPENDIX E

VARIATION OF COIL RESISTANCE AND IMPEDANCE WITH APPLIED

VOLTAGE

Appendices A, B, C, and D establish the variation of

several of the important electrical characteristics of SOVs

as a result of changes in temperature and frequency of

excitation. In all cases, however, the measurements

represent small-signal properties, since the level of

excitation applied by the measuring instrument (a digital

ohmmeter or an impedance analyzer) was only about 1 V.

We now address the large-signal behavior of two especially

important electrical characteristics of SOVs-dc resistance

and ac impedance at normal power line frequency (60 Hz).

Table E.1 shows measurements obtained from four

different instruments applied to SOV "D" at room

temperature. (Ac measurements were performed at 60 Hz

exclusively.) The nearly 3:1 difference between R& and

I Z I as measured by the Hewlett-Packard model 4192A

impedance analyzer is understandable in terms of the data

presented in Appendix B, but the source of the almost

2:1 difference in I Z I measured at high vs low excitation is

not readily apparent. The instrumentation was vindicated

by measuring a nonreactive load of about the same ohmic

value as SOV "D" (four nominal 3500-0 resistors

connected in parallel so as to yield a combined resistance

of approximately 850 Q). Table E.2 shows that the results

obtained using the same measurement techniques on the

resistor as had been used on the SOV are in substantial

agreement, so a different explanation for the low- vs

high-excitation discrepancy must be sought.

The key to this mystery is the realization that SOV "D"

was in its "rest" (open) state for the first two measurements

of Table E.1, whereas it was in its "energized" (closed)

state for the third and fourth measurements, the primary

difference being the position of the iron-cored plunger

within the solenoid coil. Electrical engineers will recognize

that the presence of magnetically permeable iron in or near

the coil will dramatically affect the inductance (and

therefore the impedance) of a solenoid, and also that the

magnitudes of certain magnetic and electrical loss

mechanisms that are present only with ac are strongly

dependent on the level of magnetization, which in turn

depends on the level of electrical excitation. These effects

are illustrated in Figures E.1 and E.2.

Figure E. I was obtained by gradually increasing the dc

potential applied to SOV "D" while simultaneously

measuring the voltage across and the current through the

SOV coil in order that the SOV's dc resistance, Rdc, could

be calculated by application of Ohm's law: Rdc = VdclIdc.

The SOV was observed to change state (from open to

closed) at about 14 Vdc,* but the computed resistance is

seen to be essentially unaffected by this state change or, in

fact, by the application of additional excitation voltage.

Figure E.2 was generated in the same manner as the

previous figure except that 60-Hz ac (rather than dc)

excitation was employed. In marked contrast to the dc

results, the magnitude of the ac impedance,

I Z I = VrdI., is seen to vary by almost a factor of 3 as

the voltage applied to the SOV is increased from zero to

135 Vffns. Three distinct regions are identifiable in the

figure: (1) a low-voltage region in which magnetic flux

linkages among coil turns and between the coil and the

plunger continually build (causing increased electrical

impedance as a result of increased inductance), a process

which eventually saturates, thereby yielding a plateau

throughout which I Z I is essentially independent of

excitation voltage; (2) an abrupt, almost 2:1 increase in

impedance resulting from the retraction of the SOV

plunger into the solenoid coil as the valve changes state;

and (3) a rapid decline in I Z I as parasitic effects (i.e.,

eddy currents in the coil copper, hysteresis losses in the

iron of the plunger) cause inductance to be reduced as

solenoid excitation is further increased.

The following observations and conclusions were drawn

from these data:

1. The value measured for dc resistance is, for all practical

purposes, independent of the level of excitation** as

well as the valve state [open or closed (i.e., plunger

withdrawn or retracted into the solenoid coil)].

*'his valve actuation voltage is very low because SOV 1-3" was

designed for ac rather than dc operation; therefore, it has relatively few

turns in its solenoid coil and a rather low dc resistance (135 0). which

causes it to draw a relatively large amount of current (100 mA) at this

low voltage.

*lt is assumed that measurements are performed in a matter of seconds

in order that coil temperature is not raised significantly owing to the

deposition of heat within the coil (resistive power loss = 12R).

E-1 NUREG/CR-4819, Vol. 2
Appendix E

Table E.1. Dc resistance/ac impedance of SOV "D" at 230C

Instrument used Technique Quantity obtained Result (')

Dc ohmmeter Direct reading Low-excitation Rdc 135.5

Impedance analyzer Direct reading Low-excitation I Z 374.1

Rms voltmeter and ammeter Vrms/Imnns High-excitation (112 Vm) Z | 648.1

True rms electronic module VrmsIrms High-excitation (112 Vlm) Z | 645.9

Table E.2. Dc resistance/ac impedance of 8so-0 resistor

Instrument used Technique Quantity obtained Result

Dc ohmmeter Direct reading Low-excitation R& 846

Impedance analyzer Direct reading Low-excitation I Z l 848

Rms voltmeter and ammeter Vfmfms/s High-excitation (112 V) Z | 833

True rms electronic module Vnndnrms High-excitation (112 V) Z 879

2. In contrast, the value measured for ac impedance depends

strongly on both (a) the valve state (because inductance

is materially increased when the iron plunger is drawn

into the solenoid) and (b) the level of ac excitation

(since parasitic mechanisms decrease inductance at

higher levels of applied voltage after the plunger is

drawn into the solenoid, while at very low applied

voltage mutual inductance may be unsaturated).

3. Observations 1 and 2 above indicate that if ac impedance

measurements (as opposed to dc resistance

measurements) are to provide correct indications of coil

temperature, they must be made at known

electromechanical conditions (i.e., at a known valve

state and level of excitation). It may, for example, be

necessary to standardize the voltage applied to the SOV

or, if this is impractical, to correct the impedance

measured at nonstandard operating voltage to what

would have been obtained at an established standard

excitation.

4. The strong dependence of I Z I on the position of the

plunger within the solenoid may provide a diverse and

possibly sensitive means for ascertaining valve

position, especially under unusual (e.g., accident)

conditions where valve position may not be

determinable by other means.

NUREG/CR-4819, Vol. 2 E-2
Appendix E

140

138

I I I I I I

a

w

0

z

co

UJ

:

136 

+1%

-r

°o ' * I

ISOV Closed I

SOV open 1 -1 %

SoV

Stote

Change

134

132[

130 

0

I I I I I

5 10 15 20 25 30 35

DC EXCITATION VOLTAGE (Vdc)

Figure E.1 Dc resistance vs level of excitation: SOV ."

950

850

a

N

0

Co

0

z

0

w

<- us

750

650

550

450

350

0 25 50 75 100 125

AC EXCITATION VOLTAGE (V)

Figure E.2 Ac impedance vs level of excitation: SOV "D."

E-3 NUREG/CR4819, Vol. 2
Appendix F

APPENDIX F

VERIFICATION OF ABSOLUTE TEMPERATURE CALIBRATION BY

APPROACH-TO-ISOTHERMAL-CONDITION TESTS

The rather large differences (>400C) between temperatures

derived from SOV coil resistance measurements and those

measured at the coil periphery with a stem thermometer

(see Sect. 2.2.1.1 and Figure 2.5) created a need to prove

that there was no systematic error in measuring

temperature by the electrical method. Measurements made

in a thermostatically controlled oven (Appendix A) had

shown no problem, so what was different about the

open-air environment of the laboratory bench? It was

reasoned that open-air measurements would never yield

correspondence between the electrically derived coil

temperatures (which are inherently volume-averaged) and

the mercury-in-glass temperatures (which are point

measurements at the coil's outer radius), since the

inevitability of a thermal gradient between the inner source

of heat (-14 W of power deposited within the coil) and the

outer convective heat transfer surface of the

coil-combined with thermal resistance at the point of

thermometer contact with the coil-would ensure that the

external readings would always lie well below any internal

readings made while the coil was being heated from within.

To test this hypothesis, a more-nearly-isothemal test

environment was created by placing a dc-powered SOV

("B") in a thick-walled styrofoam insulating enclosure,

passing the electrical leads and the stem thermometer

through holes in the side of the enclosure, and filling the

remaining empty volume of the box with styrofoam

"peanut hulls" so as to minimize convective air circulation

within.

Initially, full electrical power was applied to the valve for a

period of some tens of minutes in order to bring the SOV

and its surrounding stagnant airspace up to a temperature

well above room ambient (24.60C, at which temperature

the solenoid had an electrical resistance of 793.9 fl).*

Electrical power was then reduced to a "holding" level

necessary to overcome inevitable beat losses through the

enclosure walls in order that approximate thermal

equilibrium could be attained. This essentially steady-state

situation is illustrated by the initial 2.5 h of Figure F.1,

*For SOV "B," T() = R(t) -793.9 1W3.4095 WPC + 24.6, where T(t) is

the time-dependent temperature in C, R() is the coil dc resistance in

ohms, and the temperature coefficient of resistance is taken from the

slope of Figure A.2.

where the temperature indicated by the stem thermometer

is seen to be on the order of 92 to 1000C, while the

temperature derived from coil resistance is on the order of

123 to 130'C. The fact that, even under these

well-insulated conditions, there existed an unchanging AT

of -30TC between the two slowly rising readings indicates

that the thermal resistance of the coil-potting compound is

fairly high, and therefore a strong radial thermal gradient

within the coil is unavoidable so long as the coil is

powered.

Therefore, at t = 2.5 h the electrical "holding" power was

removed. Over the succeeding 32 min the mercury

thernometer reading dropped -22TC and the

resistance-derived temperature plummeted by -49TC

(Figure F.1) as the SOV continued to dissipate heat to its

surroundings. After an additional 8 min the two indicated

temperatures differed by only 1.5TC, even though the entire

SOV was still quite hot (-75C). This

approach-to-isothermal test provides convincing evidence

that, under the right conditions, temperatures inferred from

coil resistance will be in good agreement with temperatures

indicated by a glass stem thermometer or thermocouple

placed in contact with the coil's outer surface.

Further confirmation is provided by the data plotted in

Figure F.2, which illustrates two additional

approach-to-isothermal tests in which SOV "B" was

allowed to cool (unpowered) within the insulated enclosure

after having been brought to an elevated temperature by

application of normal electrical power (125 Vdc). In the

first test (left side of graph) the two temperature indications

came to within 2.1TC of agreement at an SOV temperature

of -92TC in about 16 min after termination of coil heating.

In the second test (right side of graph) even better

agreement (<0.50C) was obtained after a longer period

(-30 min) in the unpowered state allowed valve

temperatures to fall even further (45 to 500C) but

nevertheless remain well above normal ambient.

In summary, the apparent temperature discrepancies noted

in Sect. 2.2.1.1 are not indicative of a calibration problem

or a systematic measurement error but rather result from a

combination of relatively high thermal resistance at the

stem thermometer-coil contact point and the existence of a

F-1 NUREG/CR-4819, Vol. 2
Appendix F

from coil dc

resistance

from Hg

thermometer

140

a

0

tLZ

C, W

a

z

130

120

110 I

91".7°1C .

Coil excitation turned off 

I ............ F.......I........

100 I

90

80

70

0.0 0.5 1.0 1.5 2.0 2.5 3.0 3.5

ELAPSED TIME (hours)

Figure F.1 Comparison of temperature derived from coil resistance to that indicated by a thermometer external to

the coil under (1) essentially equilibrium (electrically powered) and (2) transient (electrically unpowered) thermal

conditions. (For the initial 2.5 h of the test, 98 Vdc was applied to the SOV to approximately balance its heat losses to

the environment.)

120 - . . . . . . . . . . . . . . . . . . . . . . . . . . . . .

.. ,.vr.,....,....j....

0

w

UJI W

IL

wLU

a

z

115

110

105

100

95

f from coil dc

resislonce

Q

" \

IQ.

READ 

SCALE 

. . .I .... ....... .

-T .I I'II ..

f from mercury

thermometer

60

fLu 0

to

55 M

to

a

0.

50 Lo

z

45

90 I

.... ......................................

0 10 20 30 40 50 60 70 80 90

ELAPSED TIME (minutes)

Figure F.2 Two more approach-to-isothermal-condition tests for SOV B" placed in a thermally well insulated

container. (For a prior period, the valve was brought to an elevated temperature by application of normal electrical

power to the solenoid coil; electrical excitation was terminated at = 7 min.)

NUREG/CR4819, Vol. 2 F-2
Appendix F

large radial temperature gradient within the potted solenoid

coil when the latter is electrically energized. The important

implications of this finding to real plant measurements are

that (1) coil-resistance-derived temperature inferences

should provide more realistic figures for the prediction of

coil life than would be obtained from a nonimbedded (i.e.,

surface-mounted) conventional temperature sensor, and

(2) coil-resistance-derived temperatures may be too high

(i.e., overly conservative) for estimating service life of

temperature-sensitive SOV components other than the coil

itself (e.g., valve seats, O-rings, gaskets and seals, etc.).

F-3 NUREG/CR4819, Vol. 2
Distribution

NUREG/CR-4819, Vol. 2

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70. A. C. Thadani, U.S. Nuclear Regulatory Commission, Office of Nuclear Reactor Regulation, Division of

System Technology, MS8 E2, 1 White Flint North, 11555 Rockville Pike, Rockville, MD 20852

71. H. L. Ornstein, U.S. Nuclear Regulatory Commission, Office for Analysis and Evaluation of Operational Data,

Reactor Operations Analysis Branch, MS 2104, Maryland National Bank Building, 7735 Old Georgetown

Road, Bethesda, MD 20814

72. V. P. Bacanskas, Gilbert/Commonwealth, Inc., P. 0. Box 1498, Reading, PA 19603-1498

73. John R. Shank, Automatic Switch Co., 50 Hanover Road, Florham Park, NJ 07932

74. Phil Holzman, Strategic Technology and Resources, 195 High St., Winchester, MA 01890

75. Jack Shortt, Stone and Webster Engineering Corporation, 245 Summer St., Boston, MA 02107

76-405. Given distribution as shown in NRC category RV (10-NTIS)

NUREG/CR-4819, Vol. 2
NRC ORv 335 U.S. NUCLEAR REGULATORY COMMISSION 1. REPORT NUMIBER A7 89. (A&41;"i dby 1RC. AOd Vol.. Swao.. Rev.. \,RCM I 02 ad Aod.ndum Nwnbers, I a&-V.)

3201 3202 BIBLIOGRAPHIC DATA SHEET

(See nsrpuctions on the revers£1 NUREG/CR-4819

2. TITLE AND SuBTITLE ORNL/T14-12038

Aging and Service Wear of Solenoid-Operated Valves Vol._2

Used in Safety Systems of Nuclear Power Plants 3. DATE REPORT PUBLISHED

MONTH EAA

Evaluation of Monitoring Methods July 1992

4. FIN OR GRANT NUMBER

B0828

5. AUTHORISi 6. TYPE OF REPORT

Technical

R. C. Kryter 7. PERIcD COVEREDclajiow-r8 PERFORMING ORGANIZATION - NAME AND ADDRESS .vAC. a'o.'. f ,O,,sot, core~so,. Ulf L'.S hcIle, Arw9Jrorv Coynn"'.Son fna nbdg aoddess l cont"ItcO' ro..o.

nae *fd Lana m ddreAwl

Oak Ridge National Laboratory

Oak Ridge, TN 37831

9. SPONSORtNG ORGANIZATION - NAME ANO ADODRESS It/ sac vSad at AMnO ' ,ntracro, oo,0i, NAC 0vion Oil ce o Regon US, Nuel Awe9W1egoV Co--".'oS ond milding addffll

Division of Engineering

Office of Nuclear Regulatory Research

U.S. Nuclear Regulatory Commission

Washington, DC 20555

10. SUPPLEMENTARY NOTES

11. ABSTRACT (200 .ored 0r 1Ii

Solenoid-operated valves (SOVs) were studied at Oak Ridge National Laboralory as part of the USNRC Nuclear Plant Aging

Research (NPAR) Program. The primary objective of the study was to identify, evaluate, and recommend methods for inspection,

surveillance, monitoring, and maintenance of SOVs that can help ensure their operational readiness-that is, their ability to perform

required safety functions under all anticipated operating conditions, since failure of one of these small and relatively inexpensive

devices could have serious consequences under certain circumstances.

An earlier (Phase I) NPAR program study described SOV failure modes and causes and identified measurable parameters thought to

be linked to the progression of ever-present degradation mechanisms that may ultimately result in functional failure of the valve.

Using this earlier work as a guide, Ihe present (hase 11) study focused on devising and len demonstrating the effectiveness of

techniques and equipment with which to measure performance parameters that show promise for detecting the presence and trending

the progress of such degradations before they reach a critical stage.

Intrusive techniques requiring the addition of magnetic or acoustic sensors or the application of special test signals were investigated

briefly, but major emphasis was placed on the examination of condition-indicating techniques that can be applied with minimal cost

and impact on plant operation. These include monitoring coil mean temperature remotely by means of coil de resistance or ac

impedance, determining valve plunger position by means of coil ac impedance, verifying unrestricted SOV plunger movement by

measuring current and voltage at their critical bistable (pull-in and drop-out) values, and detecting the presence of shorted turns or

insulation breakdown within the solenoid coil using interrupted-current test methods. The first of these techniques, though perhaps

the simplest conceptually, will likely benefit the nuclear industry most because SOVs have a history of failure in service as a result of

unwitting operation at excessive temperatures.

12. KEY WORDS/DESCR!PTORS tist wore or pheafseS that *.I assisr reatehlrsin locating the rtooJ I 13 AVAILABILlT STATE.JEN.T

Aging, degradation, diagnostics, maintenance practices, Unlimited

monitoring methods, nuclear power plants, operating experience, *& SECRuT'iCLASS%;CAT&ON

solenoid-operated valves, SOV, stressors, testing Unclassified

(rhj RoeorLr

Unclassified

IS. NUMBER OF PAGES

16 PRICE

l4RC FORM 35 2 891
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