SUPERCONDUCTING

POWER TRANSMISSION

ANNUAL REPORT

October 1979 - September 1980

>Q0-

100

PREPARED DY Thermophysical Properties Division Motional Bureau of Standards

Boulder, Colorado 80303

81-1649

1081

c,2

PREPARED FOR Brookhaven Narionol Loborarory

Power Transmission Project

Upton, New York 1 1973

OCT 2 2 1981

NBSIR 81-1649

HELIUM RESEARCH IN SUPPORT OF

SUPERCONDUCTING

POWER TRANSMISSION

rto, n

Wit

Annual Report for the period

October 1, 1979 - September 30, 1960

PREPARED DY Thermophysical Properties Division Center for Chemical Engineering

National Bureau of Standards

Boulder, Colorado 80003

D. E. Daney, Editor

Conrracr No 433475-S

PREPARED FOR Brookhaven National Laboratory

Power Transmission Project

Upton, New York 1 1 973

S<,»tAU Of

U.S. DEPARTMENT OF COMMERCE, Malcolm Baldrige, Secretary

NATIONAL BUREAU OF STANDARDS, Ernesr Ambler Direcror

CONTENTS

PAGE

1.0 COOL-DOWN STRATEGIES 1

1.1 Introduction 1 1.2 Description of the Transmission Line and Refrigeration System 3 1.3 Computational Method 8 1.3.0 Governing Equations 8 1.3.1 Material Properties 9 1.3.2 Numerical Method 9 1.3.3 Boundary Conditions 10

1.4 Results and Discussion 10

1.5 Fast Cool-Down Prototype 19

1.6 Summary 19

1 .7 References 22 2.0 HEAT- FLOW METERS 23

2.1 Introduction 23

2.2 Pipe Section and Meter Assembly 24

2.3 Field Burial, Pipe Assembly 24

2.4 Test Sequence 28

2.5 Test Results 28

2.6 Field Performance of the "A" and "B" Pair of Heat-Flow Meters 29

2.7 Field Performance of the "C" Pair of Meters 32 2.8 Summation of the emf from the "C" Pair of Meters 32 2.9 Time Response of the Heat-Flow Meters 35

2.10 Comparison of a Meter Thermocouple to an Earth Thermocouple During a Snow-Melt Period 35 2.11 Performance of the Heat-Flow Meters Above Ground 35

2.11

.1 Non-Shaded 35

2.11.2

Shaded Cover 38

2.12 Conclusions 38 2.13 References 39

ii i
ABSTRACT

During FY 80, the MBS Thermophysical Properties Division program of research for

superconducting power transmission line (SPTL) development focused on two tasks:

1) Development of SPTL cool-down strategies.

2) Experimental evaluation of thermal flux meters as a possible technique for determining enclosure heat leak.

In developing SPTL cool-down strategies, we used a computer code (previously developed

at MBS) to explore cool-down times for a wide variety of realistic boundary conditions.

Cool-down times of 10 days are feasible with only minor modification to previously proposed

ref ri geration systems if cooling channel lengths are reduced to two thirds or one half those

dictated by steady-state considerations.

Evaluation of thermal flux meters was concluded. Below ground field tests revealed

large effects due to seasonal variation in the soil heat flux. These effects can be largely

cancelled by algebraic addition of the signals from a pair of horizontally opposed sensors. A brief above ground field evaluation indicated that diurnal variations in the heat

flux completely mask heat fluxes typical of the anticipated enclosure heat flux.

Key words: Cable cool-down; cool-down; heat flux meters; helium; superconducting power

transmi ssion.

iv
1.0 COOL-DOWN STRATEGIES

(D. E. Daney)

1.1 Introduction

Cool-down of superconducting power transmission lines (SPTL's) has been the subject of

investigation at N3S for the past several years. Analytical, experimental, and numerical

modeling of SPTL cool-down have been reported previously [1, 2, 3]. The subject of this

report is application of the numerical model developed at NBS by Jones [2, 4] to perfect a rapid cool-down strategy for a particular SPTL application. The application chosen is for a Philadelphia Electric Company (PECo)* hypothetical 10,000 MVA transmission line between a power station on the Susquehanna River and Philadelphia - a distance of 106 km. This

3-p'nase ac, 230 kV SPTL has previously been the subject of an extensive economic and technological evaluation comparing a number of different power transmission systems [5].

Rapid cool-down is essential to a practical superconducting power transmission line

since total down-time for repairs includes both the cool-down and warm-up times. Although

the more rapid the better, a week to ten day cool-down is taken as a reasonable compromise

between short repair down-times and the design modifications required to achieve them.

The PECo, SPTL cooling arrangement is based on one proposed at Brookhaven National

Laboratory ( BML ) . The BNL design [6, 7], figure 1.1, uses a counter-flow arrangement wherein supercri tical helium flows down the hollow core of the cable, is expanded to a lower

pressure (and temperature) at the far end, and then is returned in the space between the

cable and the cryogenic enclosure. Although this counter-flow arrangement eliminates the

expense of a separate insulated helium return line, it poses cool-down difficulties. The

cooling channel is, in effect, a rather efficient counter-flow heat exchanger, and at the

limit of perfect heat exchange a counter-flow heat exchange will never cool down. This

situation arises because as the heat exchange effectiveness increases, the temperature difference between the go and return streams (which determines the rate of refrigeration) de- creases. Fortunately, there are several practical ways to alleviate this situation. Their

development requires a fundamental understanding of the problem.

Preliminary calculations for this PECo SPTL gave a cool-down of 104 days. Clearly

significant changes were required if acceptable cool-down times of a week to ten days were

to be achieved.

The general approach to be taken was indicated by our previous analytical work for the

counter-flow arrangement which gave the expression for cool-down time, t cd , as:

*The hypothetical transmission line chosen for the purpose of the study was established by

the sponsor. The use of this hypothetical case in no way constitutes an endorsement by the

National Bureau of Standards of the described transmission line or the Philadelphia Electric

Company PECo)."
2

Figure

1.1

SPTL

cooling

system

schematic,

unmodified

for

fast

cool-down
where i is the cooling channel length and m is the mass flow rate. When the flow is limited

by the cooling channel pressure drop the cool-down time expression becomes:

*cd

For cooling both ends, the effective length is one-half that for cooling from one end

only. Our numerical studies support the same general conclusions, although the exponents

differ slightly. Minimum cool-down times thus require:

a. Minumum economically practical cooling channel length,

b. Maximum refrigeration possible (minimum practical temperature) applied to both

ends, and

c. Maximum driving pressure (i.e., maximum flow).

By appropriate application of these principles we reduced the calculated cool-down time

to the required 10 days. Among the changes in design that are required is a reduction in

cooling channel length from 28 km to 19 km or 14 km.

During the course of the study, 24 cases were calculated to evaluate the effect of

length, pressure drop, far-end expander, far-end liquid nitrogen precooling, near-end re- frigerator configurations, and heat leak for the counterflow cool-down arrangement. Cocurrent cool-down and counter-flow warm-up were also studied.

1.2 Description of the Transmission Line and Refrigeration System

Refrigeration for the Philadelphia Electric Company (PECo) hypothetical superconducting

power transmission line (SPTL) is supplied by the arrangement in figure 1.2. Two refrigeration stations plus three far-end turboexpander stations cool the entire 106 km length, so that the length of a single cooling channel is 26.6 km. Because it may not be convenient to position the stations at the exact intervals shown, we assume a 28 km cooling channel length

as a base case. Closer spacings considered in the cool-down calculations are 18.7 km (3

refrigeration stations), and 14 km (4 refrigeration stations).

The refrigerator uses the hybrid Claude-Brayton cycle shown in figure 1.3, which gives

typical steady state operating conditions. Some cool-down calculations assume a 20 bar go- stream inlet rather than the 15 bar given in the figure.

Although the far-end refrigeration supplied by the far-end turboexpander (figure 1.1)

is adequate for steady state operation, it is inadequate for rapid cool-down. A simple and

inexpensive means of improving the far-end refrigeration is addition of a far-end liquid

nitrogen precooler as shown in figure 1.4. With this arrangement, the return stream may be

cooled to 80 K by the LN

2 bath and returned directly to the SPTL or it may also pass through

the turboexpander before returning. The lowest far-end return-stream temperature is, of course, achieved when both the turbine and the LN2 bath are used. However, the mass flow is

greater if the turbine is bypassed since the pressure drop that was previously taken up across the turbine is now used to drive the flow.

The cable construction and dimensions assumed for the cool-down calculations are given

in figure 1.5 and table 1.1.

3
Id-site s©

33

<0

*>» m 4

Figure

1.2

Hydraulic

schematic

for

one

line

of

230

kV

superconducting

cable
300 K 295 K

1 5.4 bar I T 2.8 bar

H.X. I S >

58.6 k 153.2

Turboexpander

H.X.H <

39.2 I 38.8

H.X. III < <

21.2 K l

19- 8

s —Turboexpander

H.X. IV < VTs

13.8 13.7

H.X.V «> >6.6 K

7.3 K 1 3.0 bar

WTurboexpander

6.1 bar 1—4 Return Stream

1 5.2 bar

Figure 1.3 Refrigerator schematic

5
6

Figure

1.4

SPTL

cooling

system

schematic

with

far-end

liquid

nitrogen

precooler
Bronze Core, 2 Helices

5 Strips

Aluminum Stabilizer,

2 Helices, 5 strips

Copper Superconductor,

2 Hel ices 5 Strips

Inner Oielectric Screen (intercalated), 2 Tapes Metallized Polyimide

Dielectric, Polyethylene

93 Layers

Outer Dielectric Screen (intercalated), 2 Tapes

Copper Superconductor,

2 Hel ices

Alumi num Stabi 1 izer,

2 Helices

Outer Conductor Insulation

(kaDton), 4 Layers Stainless Steel Compression

Layer

Taped Fiberglass Thermal Insulation, 2 Layers

Lead Gas Jacket

Armor, mylar-Bronze Layment

Lay Angle of 80° to 85°

Dimensions in Millimeters

O.D. Thickness

A 90. L4 1.77

B 93.68 1.77

C 94.44 .38

D 95.20 .38

E 95.40 .10

F 95.60 .10

G 96.00 .20

H 126.10 15.05

I 126.20 .05

J 126.40 .10

K 126.60 .10

L 127.10 .25 M 127.60 .25

N 128.60 .51 0 130.14 .76

P 131.66 .76 0 136.76 2.55

R 139.73 1.49

S 142.70 1.49

Figure 1.5 Cable cross section, cut back view

7
Table 1.1

Summary of Cable and Enclosure Details

CABLE

Core inside diameter (mm) 90.1

Insulation inside diameter (mm) 95.6

Insulation outside diameter (mm) 126.2

Cable outside diameter (mm) 142.7

Go stream flow area (3 cables) (m^) 17.7 (10)"^

Weight/meter (3 cables) (kg/m) 121.7

ENCLOSURE

Inside diameter (mm) 340.0

Wall thickness (mm) 2.2

Steady state heat leak (W/m) 0.45

Cooling channel length (base case) (km) 28.0

1.3 Computational Method

1.3.0 Governing Equations

Because the computational method has previously been described in detail [2, 4] we give

only a brief outline here. For each of the helium streams the one-dimensional continuity,

energy, and momentum equations are:

Vdp +

dx

pdV

5x

= 0 (l.D

«)*

pVo / ax v

’ H+V^ + + —

J

at a (1.2)

m - dx

_

1P2Pax fv | v

!

ZT

p

as' ax (1.3)

Here, the density p is regarded as a known function of P and H, the potential energy § is a known function of x, accounting for changes in elevation, and A is the heat transfer to the

fluid stream per unit length. The friction factor f could be calculated from the local

fluid Reynolds number or, as we have done here, set equal to a constant for typical conditions. p and a are wetted perimeter and cross-sectional area of the appropriate helium

channel

.

We can arrive at a set of equations in which only one time derivative -Is-, -IS- or

at at at occurs by taking linear combinations of (1.1), (1.2), and 1.3), and by use of thermodynamic

identities. This form is required by the numerical method we use. By retaining P and H as thermodynamic variables rather than P and T we avoid thermodynamic derivatives such as specific heat which are sharply peaked at the transposed critical temperature. The

resulting equations are:

8
(1.4) aP _ vaP pc

2

_av §£ " ax ax

aH = - vH_c2 j^ + !±?U + py

2f | v | p M ~ ax ax p a “2a

av = _

i ep _ vav _ f v|v|p _ a®' tl p ax ax “2a ” ax

a + Pv

2f I v I p

a 2a

(1.5)

( 1 . 6 )

where $ = ^the Gruneisen parameter and c

2 =^j , where c is the acoustic velocity.

There is oni such set of three equations for each

s fluid stream and the independent

variables and coefficients can be subscripted 1 and 2 for the inner and outer streams re- spectively. They are coupled to appropriate energy equations for the cable and outer pipe

through the heat transfer terms A and a

2

. For the cable composite the appropriate one-dimensional energy conservation equation is

of the form

r , 8T cable V d/ L , aT cable\ . cable . cable M 7

v 2^ pi C

i a t~M 2—t &x (^i ai -A2 U - 7)

where pn-, C,-, kj, and a

n

* are the density, specific heat, thermal conductivity and cross

sectional area of the -1^1 cable component. It is easy to show that the heat conduction term

on the right hand side is quite negligible in all cases considered. A similar equation can

be written for the outer pipe:

( pCa) pipe (ka) • ST + Q

- A2P

1‘Pe

( 1 . 8 )

where again for practical purposes the first term on the right hand side can be neglected.

Q is the radiant heat leakage per unit length to the outer pipe from the ambient temperature

vacuum pipe through a multilayer insulation which is assumed to cover the outer pipe.

1.3.1 Material Properties

In eqs. 1.4 through 1.6 the coefficients p, C

2

, $, and, through A, T are regarded or known functions of the dependent variables P and H. For fast computer evaluation it was

found advantageous to express these as polynomials fitted to the more exact equation of state of McCarty [8]. In eqs 1.7 and 1.8 the specific heats Cj of the cable components were

fitted to selected data from the literature.

1.3.2 Numerical Method

Integration of the set of eight partial differential equations was accomplished using

the computer program PDECOL developed by Madsen and Sincovec [9], This program is based on the method of lines and uses finite element collocation for the discretization of the

spatial variable. Using piecewise polynomials for the trial function space, the procedure

reduces the system to a set of first order ordinary differential equations with time as the

independent variable. The program then uses established procedures for stiff systems to

9
perform the time integration. PDECOL has been tested by its authors and found to be up to

an order of magnitude faster than a similar finite difference technique for the same accuracy.

Since PDECOL was designed for non-linear partial differential equations it was relatively straightforward to introduce a properties subroutine to calculate the properties

discussed above for every new value of the solution vector and supply the solution-dependent

coefficients in the differential equation. This adds significantly to the run time, but the

increase was still less than 10% in our computations.

1.3.3 Boundary Conditions

In the counter-flow cool-down calculations the flow is governed by the pressure drop

which gives the boundary conditions

at x

P

P

1

2

= 0

= constant

= constant

at x = L

P

1 = constant

P

i A

i

V

2 = p 2 a

2 V2

( contl

’ n uity)

(1.9)

( 1 . 10 )

where subscript 1 refers to the go stream and 2 refers to the return stream. For constant temperature refri geration the enthalpy boundary condition is H = H(t)

where H(t) is a linear or exponential decay to a fixed yalue. The far-end expander gives

H

2 = H 2 (Pi,y,P 2 ,Hi) and a more realistic near-end refrigerator gives H, = H

] (P

2 ,H ). In co-current cool-down the flow is limited by the available mass flow and the boundary

conditions become

at x

( PAV)

1

( PAV

)

2

at x

Pi

P 2

= 0

= constant

= constant

= L = constant

= constant

( 1 . 11 )

( 1 . 12 )

Specification of both pressure and flow at the same end is numerically unstable, and

physically impossible without feed-back control at the far-end.

1.4 Results and Discussion

The effect of length, pressure drop, far-end expander, liquid nitrogen precooling,

precooling temperature, heat leak, and refrigerator configuration were evaluated in twenty10
four cool-down calculations for the PECo SPTL. Table 1.2 summarizes these results. One

warm-up calculation (run 1015) is also included.

The shortest cool-down time, 3.8 days, (run 519) is given by the co-current arrangement

wherein streams 1 and 2 both flow in the same direction. Increasing the flow to 1.5 kg/s

would give a 2.5 day cool-down. Although the co-current arrangement gives the fastest cooldown, it requires a separate return line and a refrigerator much larger than that required

for steady-state operation. Thus it is not an economically attractive method of SPTL cooldown.

The next fastest cool-down time is 7.9 days for a 14 km length (run 225) assuming cooling from both ends with ref rigerators supplying a constant 10 K helium stream. Although

constant temperature ref rigerators at both ends are not particularly practical, this case

gives the minimum cool-down time for the counterflow arrangement.

In considering the far-end arrangement, the preferred mode of operation (referring to figure 1.4) is to bypass the turboexpander and use all the available pressure drop to in- crease the flow rate. When the go-stream exit temperature is greater than the liquid

nitrogen bath temperature, the flow is passed through the LN

2 bath before returning. After

the go-stream exit temperature falls below the bath temperature the go-stream connects

directly to the return stream inlet, bypassing the LN

2 pot. A reduction in cool-down time

of about 10% can be achieved by pumping the far-end LN bath to 65 K, but the added com- plexity and capital cost appear undesirable.

Although a constant temperature near-end refrigerator is a convenient device for com- paring the effect of length, pressure drop, heat leak, and various far-end arrangements, it

is not a realistic model of actual refrigerator behavior. In practice, the refrigerator

output temperature (go-stream inlet temperature) depends on the input temperature (returnstream exit temperature). If we assume that the refrigerator passes through a series of

quasi -steady states during cool-down (i.e., that the refrigerator time constant is short

compared to the transmission line time constant) then the go-stream inlet temperature may be

calculated in a fairly straightforward manner for a particular ref rigeration scheme. With

no modification to the proposed refrigerator (figure 1.3) we obtained a 28 day cool-down for

a 18.7 km length (run 605) compared to 12.2 days for a near-end constant temperature refrigerator (run 506). In this case, the go-stream inlet is cooled to the temperature of the

turboexpander discharge.

Two schemes which provide much faster cool-down are given in figure 1.6. Both schemes

are modified versions of the steady-state refrigeration scheme shown in figure 1.3. Both

schemes use liquid nitrogen precooling. The refrigerator output temperature (go-stream

inlet temperature) for these schemes is given as a function of input temperature (returnstream exit temperature) in figure 1.7.

In Scheme A, (figure 1.6) the return stream exit is cooled to 80 K by the LN

2 bath if

T

2 is greater than 80 K. If T

2 is less than 80 K, the return stream goes directly to the

expander. In both cases the gas is expanded from 6 atm to 3 atm to give temperature T^. We assume that the go-stream enters the heat exchanger at temperature T

a

= 2 T

2 or that

11
Table 1.2 Summary of Calculations

Run 1015, warm-up

Near-end, 300 K supply

Far-end, 300 K heater

P. = 15.1 - 6.8 bar

P, = 3.0 - 6.9 bar

Length = 14 km

Warm-up time = 23 days to 257 K at

mi dpoi nt

Run 724

Near-end, refrigerator Scheme B Far-end, 80 K LN, cooling

P, = 15.2 - 6.8 Bar

P, = 3.0 - 6.9 bar

Length = 18.7 km

Cool-down time = 19.1 days

Run 627b

Near-end, refrigerator Scheme B Far-end, 80 K LN cooling

P. = 20.3 - 8.8 Bar

P, = 3.0 - 8.8 bar

Length = 18.7 km

Cool-down time = 11.6

Run 605

Near-end, turboexpander only

Far-end, 80 K LN cooling

P, = 20.3 - 10.6 bar

P = 6.1 - 9.9 bar

Length = 18.7 km

Cool-down time = 28.4 days

Run 528

Heat leak, linear function of

temperature

Near-end, 10 K refrigerator

Far-end, turboexpander with 80 K LN

2 precool i ng

P = 20.3 - 10.6 bar

P: = 6.1 - 9.9 bar

Length = 18.7 km

Cool-down time = 12.2 days

Run 506

Near-end, 10 K refrigerator

Far-end, turboexpander with 80 K LN

2 precool i ng

P = 20.3 - 10.6 bar

P, = 6.1 - 10.0 bar

Length = 18.7 km

Cool-down time = 12.2 days

Run 904

Near-end, refrigerator Scheme B Far-end, 80 K LN cooling

P, = 15.2 - 6.8 Bar

P

2 = 3.0 - 6.9 bar

Length = 14 km

Cool-down time =9.8 days

precool i ng

Run 627

Near-end, refrigerator Scheme B Far-end, 80 K LN, cooling

P = 20.3 - 10.6 bar

P 9 = 6.1 - 9.9 bar

Length = 18.7 km

Cool-down time = 12.7 days

RUN 624

Near-end, refrigerator SCHEME A Far-end, 80 K LN cooling

P, = 20.3 - 10.6 bar

P = 6.1 - 9.9 bar

Length = 18.7 km

Cool-down time = 23 days

Run 529

Heat leak = 2 x nominal value

Near-end, 10 K refrigerator

Far-end, Turboexpander with 80 K LN,

precooling

P, = 20.3 - 10.6 bar

P = 6.1 - 9.9 bar

Length = 18.7 km

Cool-down time = 12.2 days

Run 519

recurrent cool-down

Near-end, 10 K refrigerator

in = 1.0 kg/s

Length = 28 km

Cool-down time = 3.8 days

Run 505

Near-end, 10 K refrigerator

Far-end, turboexpander with 80 K LN,

precooling

P = 20.3 - 10.6 bar

P = 6.1 - 10.0 bar

Length = 28 km

Cool-down time = 38 days

12
Run 325

Near-end, 10 K refrigerator

Far-end, turboexpander with 65 K LN

2 precooling

P = 15.2 - 11.1 bar

PI = 6.1 - 7.7 bar

Length = 28 km

Cool-down time = 90 days

Run 321

"Rear-end, 10 K refrigerator

Far-end, turboexpander without

precooling

p = 15.2 - 11.1 bar

P, = 6.1 - 7.7 bar

Length = 28 km

Cool-down time = 179 days

Run 319

Rear-end, 10 K refrigerator

Far-end, turboexpander without

precool i ng

P = 15.2 - 10.1 bar

P

2 = 6.1 - 8.0 bar

Length - 14 km

Cool-down time = 17 days

Run 314

Rear-end, 10 K refrigerator

Far-end, turboexpander with 65 K precool i ng

P = 15.2 - 10.1 bar

P

2 = 6.1 - 7.9 bar

Length - 14 km

Cool-down time = 10.1 days

Run 313

Rear-end, 10 K refrigerator

Far-end, turboexpander with 80 K LR

2 precooling

P = 15.2 -'10.1 bar

P

2 = 6.1 - 8.0 bar

Length = 14 km

Cool-down time =11 days

Run 312b Near-encT, 10 K refrigerator

Far-end, turboexpander with 80 K LN

? precooling 1

P. = 15.2 - 10.1 bar

P

2 = 6.1 - 7.5 bar

Length = 28 km

Cool-down time = 84 days

Run 126

Near-end, 10 K refrigerator

Far-end, turboexpander without

precooling

P = 15.2 - 10.1 bar

P

? = 6.1 - 8.0 bar

Length = 28 km

Cool-down time = 104 days

Run 324

Near-end, 10 K refrigerator

Far-end, turboexpander with 65 K LN

2 precooling

P = 15.2 - 9.1 bar PI = 6.1 - 8.2 bar

Length = 28 km

Cool-down time = 69 days

Run 320

Near-end, 10 K refrigerator

Far-end, turboexpander without

precooling

p = 15.2 - 9.1 bar PI = 6.1 - 8.2 bar

Length = 28 km

Cool-down time = 131 days

Run 318

Rear-end, 10 K refrigerator

Far-end, turboexpander with 65 K LN

2 precool i ng

p = 15.2 - 10.1 bar PI = 6.1 - 7.3 bar

Length = 23 km

Cool-down time = 77 days

Run 314b

Rear-end, 10 K refrigerator

Far-end, turboexpander with 65 K LR

2 precool i ng

P = 15.2 - 10.1 bar

= 6.1 - 7.9 bar

Length = 18.7 km

Cool-down time = 22.5 days

Run 312

Near-end, 10 K refrigerator

Far-end, turboexpander with 80 K LR

2 precool i ng

P = 15.2 - 10.1 bar

P, = 6.1 - 8.1 bar

Length = 18.7 km

Cool-down time = 26 days

Run 225

Near-end, 10 K refrigerator

Far-end, 10 K refrigerator

P, = 15.2 - 10.1 bar

P

2 = 6.1 - 7.7 bar

Length = 14 km

Cool-down time = 7.9 days

13
SCHEME A*

SCHEME &

14

Fast

cool-down

refrigeration

schemes.

The

heat

exchanger

numbers

refer

to

equivalent

positions

in

figure

1.3.
o

>i ‘aaruvaadwax iaiNi wvaaxs oo

15

RETURN

STREAM

EXIT

TEMPERATURE
T

a = 80 K, whichever is less. Knowing T

a

, T

d

, and the heat exchanger effectiveness, the

heat exchanger is completely specified and T

c (T^ is obtained by iteration. That is, a value of T

b is assumed and then the effectiveness is calculated using real helium properties. If the calculated effectiveness does not agree with the assumed effectiveness, then

a new value of T

b is chosen and the calculation is repeated. The choice of Tg

= 2 T

2 is not completely arbitrary, since this is approximately the condition for normal operation.

The evaluation of Scheme E proceeds in an analogous fashion. The return stream exit

temperature is cooled by the LN , bath when T

? is greater than 80 K. Temperature T

b

, the

heat exchanger effectiveness, the expander efficiency, P

d and P d determine the upper heat

exchanger operating condition. A value of T

a is estimated, and T

^

is then calculated from

P

fl

, P d, and the expander efficiency. Temperature T

c and the heat exchanger effectiveness

are then computed. If the calculated effectiveness does not agree with the chosen value,

then a new value of T

a is assumed and the calculation is repeated. The lower heat exchanger

is now completely determined since T

c

, T

d , P c

, Pf and the expander efficiency are known.

Scheme A reduces the cool-down time from 28 days (run 605) to 23 days (run 624) for a 13.7 km length, and Scheme B reduces it to 12.7 days (run 627) or 11.6 days (run 627b). The

shorter time for run 627b results from the 3.0 bar return stream exit pressure. These times

for Scheme B are close to those for constant temperature near-end refrigeration (run 506)

and about 50% greater than the 7.9 days attained for counter-flow cool-down using constant

temperature ref ri gerators at both ends (run 225).

Although we explored 24 different cool-down cases, there still remain many more conditions of possible interest. We may extrapolate these results over a limited range by using

our previous analytical results [3].

tcd ~ (pAp)

n (1.13)

where m = 3 and n = -1. A somewhat more accurate extrapolation uses exponents evaluated

from the numerical results. In figure 1.8 and 1.9, we plot cool-down time versus length and

the product of average pressure p times the pressure drop Ap for the up-stream. The

exponents evaluated from these figures are m = 2.8 and n

= -0.85.

Warm-up (run 1015) takes considerably longer than cool-down. For this case it required

23 days to warm the mid-point to 257 K, and we estimate approximately 30 days to warm to 0°C. This aysmmetry in the warm-up and cool-down times is not due to any fundamental difference in these processes. Rather it is due to the strong temperature dependence of the

cable and enclosure heat capacities. During warm-up the heat capacity is largest during the

final stages when both the flow and the temperature difference between the go and return

streams are at a minimum.

The slow warm-up is not necessarily as disadvantageous as it might seem, however,

because the time reported in table 1.2 refers only to the mid-point temperature for a 14 km

length. Repairs requiring warm-up would probably not be made at the mid-point, and a section of cable exposed for repair would quickly warm-up without counter-flow circulation

of the helium.

16
TIME,

days

LENGTH, km

Figure 1.8 Cool-down times as a function of length for several cooling arrangements

17
TIME,

days

PAP, bar2

Figuer 1.9 Cool-down times versus the product of the average go-steam pressure and pressure drop. The 28 km curve is for a 10 K near-end refrigerator, and a 65 K

far-end precooler. The 18.7 km curve is for Scheme B with an 80 K far-end

precooler.

18

V
1.5 Fast Cool-Down Prototype

A prototype of a fast cool-down SPTL refrigerator for the PECo system is the configuration assumed for run 904. Figure 1.10 shows this arrangement which uses the Scheme B refrigerator at the near-end and liquid nitrogen cooling at the far-end. A 10 day cool-down

time is computed for a 14 km cable length. Note that Scheme B is essentially the same arrangement as the midportion of the standard refrigerator (figure 1.3) proposed for steady

state cooling of the SPTL. The principle modification is the addition of the LN

2 cooler.

The go-stream inlet pressure is 15.2 bar and the cooling channel length is 14 km. The

liquid nitrogen baths are bypassed once the temperature for the associated helium stream

falls below 80 K, and the far-end expander is bypassed during the entire cool-down so that

all the available pressure drop may be used to increase the mass flow rate. Although Scheme B is a conceptually simplified version of the refrigerator shown in

figure 1.3, there is no practical requirement for the elimination of heat exchangers III and

V. The elimination of these heat exchangers is merely a device to simply the analysis.

Thus the practical implementation of Scheme B would be the refrigeration of figure 1.3 with

an added LN

2 cooler.

The operating conditions are given in figure 1.11 which shows the liquid nitrogen con- sumption rates, helium inlet temperatures to the LN

2 precooling and go-stream helium mass flow rate. The cool-down history for the entire length is given in figure 1.12.

The total liquid nitrogen consuption at the far-end is 4 . 5 ( 10

)

5x ( 1 . 19 ( 10

)

5 gal)

without counter-flow heat exchange to the nitrogen boil off gas and 3.68(10) 5 £ (0.97(10) 5

gal) with counter-flow heat exchange. The near-end liquid nitrogen consumption is

0.95(10) 5Jt ( . 25 ( 10) 5 gal). An example of a suitable far-end LN

2 precooler would be a coil

of tubing 50 m long by 10 cm i.d. tubing immersed in a LN

2 dewar.

1.6 Summary

The work presented in this report demonstrates that the original refrigerator scheme

for the PECo SPTL (designed for steady state operating conditions) must be modified if

acceptable cool-down times are to be achieved. Without modification a 104 day cool-down

time is expected.

Two principal modifications are required if short cool-down times are to be obtained.

The refrigerator spacing must be reduced from 56 km to 28 km (14 km cooling channel length),

and liquid nitrogen cooling must be used both at the refrigerator and the far-end. With

those modifications the computed cool-down time for a 14 km cable length is 10 days. For

fastest cool-down, the far-end expander is by-passed so that all the available pressure drop

is used to increase the mass flow rate in the cooling channel.

Finally, we note that the trends of these computer calculations agree well with our previous numerical work so that (1.13) may be used to extrapolate the results of the 24

cool-down cases studied here.

19
Ul

20

Figure

1.10

Fast

cool-down

refrigerator

prototype
TEMPERATURE,

K

80.0

160.0

240.0

320.0

£

O

J

Un

(/}

c/> <

2

Figure 1.11 Liquid nitrogen consumption and operating conditions for fast cool-down prototype

Figure 1.12 Cool-down history for SPTL with fast cool-down refrigerator prototype, run 904.

21
1.7 References

[1] Daney, D. E., Cool-down of cryogenic power transmission lines, Progress in

Refrigeration Science and Technology, Proceedings of XVth International Congress of

Refrigeration, Venice (Italy), 1979 (1980).

[2] Daney, D. E., Helium research in support of superconducting power transmission. Annual

Report, NBSIR 79-1618, National Bureau of Standards (U.S.) (1979).

[3] Daney, D. E., Helium research in support of superconducting power transmission. Annual

Report, NBSIR 80-1637, National Bureau of Standards (U.S.) (1980).

[4] Jones, M. C., Cool-down of superconducting power transmission lines with single phase

helium. Cryogenics 20^, 139-145 (1980).

[5] Evaluation of the economical and technological viability of various underground

transmission systems for long feeds to urban load areas. Final Report, Department of

Energy (U.S.), Report HCP/T-2055/1 (December 1977).

[6] Dean, J. W., and Jensen, J. E., Supercri tical helium refrigeration for superconducting

power transmission cable studies. Advances in Cryogenic Engineering, 21_, Proceedings of

1975 Cryogenic Engineering Conference, Kingston, Ontario, Plenum Press, New York

(1976).

[7] Morgan, G. H. and Jensen, J. E., Counter-flow cooling of a transmission line by

supercritical helium. Cryogenics, JL7_, 259-267 ( 1977).

[8] McCarty, R. D., Thermodynamic properties of Helium 4 from 2 to 1500 K at pressures to 10(8) Pa, J. Phys. Chem. Ref. Data 2, 923-1041 (1973).

[9] Madsen, N. K., and Sincovec, R. F., General software for partial differential systems,

in Numerical Methods for Differential Systems, L. Lapidus and W. E. Schiesser, ed.

Academic Press, Inc., New York (1976).

22
2.0 HEAT- FLOW METERS

(P. R. Ludtke)

2.1 Introduction

The feasibility of using heat-flow (thermal flux) meters on the outer surface of a cryogenic envelope as non-i ntrusi ve indicators of insulating vacuum degradation is the

subject of this section.

A passive, non-i ntrusi ve indicator of the insulation quality has the desirable characteristics of allowing one to completely seal (weld) the vacuum enclosure and to forget

about o-ring seal problems and human errors associated with instrusive type vacuum sensors

and gauges. An external sensor becomes even more desirable if the transfer line is

buried. With this in mind, an investigation was conducted to determine the feasibility of

using heat-flow meters for this application.

One of the major concerns when using heat-flow meters underground is the magnitude of

the thermal flux within the earth compared to the thermal flux of interest. Wear the

surface of the earth, the thermal gradient is affected by four different mechanisms:

a. The surface of the earth is continuously heated by thermal energy from the earth's

core; the average heat flux near the surface is .032

W- a^

2

S (0.01 ^

-^r?) , producing

a geothermal gradient of about l°C/24.4 meters (l°C/80 feet). This constant source

of energy is quite small compared to the average energy from surface normal solar

radiation of 915 W/m 2 (290 on a clear day.

b. The varying thermal energy at the earth's surface due to diurnal radiation is

another consideration. It has been established that variations of the surface

temperature due to heating by day and cooling at night do not penetrate below the

surface more than 90-120 cm (3-4 feet) [2].

c. The annual or seasonal temperature changes near the surface of the earth are of a larger magnitude; the changes from the cold of winter to the heat of summer may

affect the temperature of the earth to a depth of 18-21 meters (60-70 feet) [2].

d. Local weather conditions usually do not affect the thermal gradient at a depth of

1.5 meters, however, it will be shown that the moisture from a heavy snow melt will

disturb a stable thermal gradient at this depth.

Thus, thermal energy from three of the above mechanisms contribute to the net thermal

gradient at a depth of 1.5 meters (axis of the pipe-section); the effects of diurnal solar

radiation become negligible at this depth.

During the first phase of the project, four different pairs of commercially available

meters were mounted on a short pipe section, tested and calibrated under controlled conditions. The meters were found to have good sensitivity, and the calibrations were linear and

repeatable with small deviations. Details of the work in phase 1 are given in NBSIR 80-1637

Cl].

During the seond phase of the investigation, the pipe section was buried for an eleven

month field test; the results of the undergound testing are reported here.

23
2.2 Pipe Section and Meter Assembly

A section of steel pipe (61 cm long x 50.8 cm O.D. x 0.64 cm wall) was selected to simulate a cryogenic envelope. Eight heat-flow meters from four different manufacturers were epoxied to the outside of the pipe. The meters will be referred to as "A", "B", "C",

and "D" pairs to eliminate bias toward any one manufacturer. Table 2.1 lists the characteristics of each meter. Six of the meters have flat surfaces, and the other two have surfaces contoured to fit

the outer surface of the pipe. Copper adapters with one side flat and the other side curved were machined from high purity copper. The adapters were designed such that the curved

surface area was equal to the flat surface area of the heat-flow meters. A moderately highthermal -conductivity epoxy was used to bond the adapters and meters to the steel pipe.

Sensors from each pair were mounted 180° from each other on a circumferential centerline,

equidistant from each end of the pipe. The outer surface of the pipe was divided by four

quadrant lines parallel to the pipe axis. Two sensors were placed near each quadrant line;

the edge of each sensor was 2 cm from the quadrant line.

Thin stainless steel strip heaters were bonded to the inside of the pipesection to provide a source of energy for calibration and testing. The stainless steel strips were

59.7 cm long x 2.46 cm wide x .005 cm thick, wired in series, and bonded to the pipe inner

surface with RTV silicone. The inside of the pipe was filled with pour-type polyurethane

foam, and the ends were squared-off and hermetically sealed with fiberglass caps. A thin

coat of RTV silicone coating was applied onto the sensors with sandwich construction for an additional hermetic seal.

2.3 Field Burial, Pipe Assembly

The pipe assembly was buried to center-line depth of 1.5 meters and backfilled with

clean, dry mortar sand. Type E thermocouples were placed in the sand during backfill to determine the thermal gradient of the earth in the vicinity of the pipe section. A cross- section showing the burial configuration is shown in figure 2.1. The pipe was laid in the

ground horizontally, with the quadrant lines in the vertical and horizontal planes. This

configuration places 2 sensors on the bottom of the pipe, 2 on the top, and 4 on the

sides. A photograph of the pipe placement in the hole just prior to backfilling with sand

is shown in figure 2.2.

The pipe was buried 7 meters from the side of the laboratory; wiring from the sensors

and earth thermocouples was routed inside the building over a length of approximately 30 meters using pair shielded cables. The heat-flow meter signals were wired directly into a multi -range, multi -point recorder, and the thermocouple signals were connected to a solidstate ice reference, and then to the multi-point recorder. DC power leads were connected to

the internal strip heaters, providing the option of heating the inside of the pipe.

24
Table 2.1 Heat-Flow Meter Characteristics

Manufacturer's

Sensiti vity

Brand

Size

(cm)

Controured

Surface

pW cm 2-mV Type of

Contruction

Type of

Sensor

A 11. 4x11. 4x. 48 No 300 Integral, Epoxy

Glass

Thermopile

B 5. 08x15. 2x. 64 Yes 135 Integral, Dial lyl

Phthalate

Thermopile

C 15. 2x15. 2x. 47 No 113 Sandwich Copperpoly i mi de

Solid state

D 5 .08x5 .Q8x.25 No 160 Sandwich AluminumphenolicSolid state

25
T

ype E

thermocouples in

sand-.

E

E

o>c

o

• u

co

in

26

Figure

2.1

Burial

configurations

for

the

pipe-meter

assembly.
Figure 2.2 Photograph of pipe before back-filling.

27
2.4 Test Sequence

A chronological listing of the test sequence is given below:

8-20-79 Buried the pipe section

8-

20-79 to 9-16-79 Allowed the burial area to assume an equilibrium

temperature

9-

16-79 to 1-3-80 Started data recording of earth thermocouples and heat1-3-80

flow meter emf. Zerr power input to pipe section. Just

observed the heatflow meter emf due to the thermal

gradient of the earth.

Applied 275 mW power to the pipe section. This is the

power necessary to generate a heat flux of 28 pW/cm 2 --

the predicted heat flux of the BNL-3 cryogenic envelope.

Maintained this power level until 1-22-80.

1-22-80 Increased the internal power to 1380 mW; this power level

produces a heat flux 5 times the predicted level. This

power level was maintained until 2-29-80.

2-29-80 Decreased the power into the pipe section from 1380 to 275

mW; this decreases the heat flux from 5 times to 1 times

the predicted value. This power level was maintained

until 3-14-80.

3-14-80 Increased the power level from 275 to 824 mW; a change in

heat flux from predicted to 3 times predicted value. This

power level was maintained until 5-30-80.

5-30-80 Decreased the power level from 824 ftW to zero. Observed

the heat-flow meter emf's at zero power level from 5-30-80

until 7-29-80.

7-29-80 Excavated the pipe section to a horizontal position above

ground with the axis 76 cm above the surface. Observed

the heat-flow meter emf's with zero internal power and no cover over the pipe section.

8-4-80 A completely shaded enclosure consisting of one layer of

double aluminized Mylar of 635 pm thickness was built

around the pipe section. The heat-flow meter emf, with

zero power input, was observed until 8-11-80.

8-11-80 Terminated the test.

2.5 Test Results

The pipe section was buried for eleven months (8-29-79 to 7-29-80). During that

period, two of the heat-flow meters malfunctioned — the first 28 days after the burial and

the second after approximately 7 months. Upon excavation both of the "D" meters were

seriously delaminated, apparently due to moisture. The other meters were all intact and

working properly.

28
The vertical temperature profile of the sand backfill during the test period is shown

in figure 2.3. The significant aspects of the earth temperature plots are:

a. The upper thermocouple (33 cm below the surface) is quite susceptible to local

weather conditions, whereas the lower thermocouple (below the pipe section) is affected very little by day-to-day or local weather patterns.

b. The vertical temperature difference across the pipe holds fairly steady (at 4°C)

during the winter months, decreases to zero in April, and starts to reverse in

May.

c. Changes in electrical power into the pipe section have negligible effect on the

thermocouples above and below the pipe section.

d. A fast heavy snow melt at the surface, caused by Chinook winds peculiar to Boulder,

Colorado results in water near 0°C seeping down through the sand the altering the

temperature gradient near the pipe. This occurred on 12-4-79 and 1-13-80. The

sand temperature directly above and below the pipe section decreased by 1.0 to 1.2°C.

e. The temperature excursion of the thermocouple 20 cm below the pipe section varied

from a low of 7°C in February to a high of 24°C in July.

2.6 Field Performance of the "A" and "B" Pair of Heat-Flow Meters

Meter pairs "A" and "B" were mounted just off the vertical plane of the pipe section,

as shown in figure 2.4. The behavior of the "A" and "B" pairs was nearly the same; however,

the "B" pair is approximately 2.2 times more sensitive than the "A" pair. After examining

the meter emf's for the period from 9-16-79 to 1-30-80 (at zero power level), it necame evident that the meters were experiencing a thermal field of considerable magnitude. As illustrated by figure 2.4, the soil heat flux does not bypass the pipe. Instead it enters

the pipe wall at the bottom and exits at the top. In doing so the heat flux passes through

the top and bottom meters in opposite directions (the meters are wired to give a positive

signal for a radial outward heat flux) and produces the mirror image signals of figure

2.4. The thermal conductance of the pipe wall (thermal conductivity times the cross section

area) is the same order of magnitude as the thermal conductance of the soil, so the heat

flux measured by the "A" and "B" meter pairs is also of the same order as the soil heat flux — even though they are mounted on the pipe surface.

After establishing the behavior of the system with zero heater input, the heater was turned on (1-3-80) to simulate the predicted heat flux of BNL-3 cryogenic envelope (28

pVI/cin 2 heat flux or 275 mW total power). During the remainder of the test the power was stepped up and down by as much as five times the nominal value to determine the minimum

detectable change in heat flux. An interesting but disturbing aspect of the data plots is the effect of a heavy snow melt on the heat-flow meter emf. Heavy snow melts occurred on 12-4-79 and 1-13-80, and on both occasions, the meter signals experienced large excursions of about twice the previous

29
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30

Figure

2.3

Soil

temperatures

during

heat

flux

meter

field

evaluation.
Heat-Flow

Meter

emt,

fj.V

Heat

Flow

Meter

emf,

6600

-1800

-2400

-3000

-3600

-4200 -4800

-6400

-6000 -6600

-7200

O°°°°o

° ° o o ® ° ©

J L J I I L

o°o uo

u

o

_!_1 I L

Time. Date

Figure 2.4 Field response of meters B1 and B2.

31
The data indicate that it is not possible to clearly distinguish a gradual change in

heat flux of even 5 times the nominal value for individual vertically mounted meters such as the "B" pair. However, the mirror image behavior for the pair of sensors suggests that

adding the output from a pair of sensors would cancel much of the effect of the soil heat

flux. This idea is discussed in detail in section 2.8.

2.7 Field Performance of the "C" Pair of Meters

The "C" pair of heat-flow meters were mounted just off the horizontal plane of the pipe

section. The "C" meters were the largest of all; the center of the meters was 9.6 cm from

the horizontal plane and 180° from each other. The companion pair of "D" meters near the

horizontal plane malfunctioned during the burial period.

The individual plots of "Cl" and "C2" emf's, figure 2.5, were in most aspects about the same as the plots for the "A" and "B" meter pairs even though the "C" meters were a mirror

image of each other, especially with zero power input. The same perturbations occurred

during the heavy snow melts as with meters "A" and "B", (signal shifts approximately twice

current values). A step power increase of 5 times predicted value gives a very distinct

shift in signal level from each meter, but again this would not be nearly as distinct if this increased heat flux occurred over a period of time.

2.8 Summation of the emf from the "C" Pair of Meters

After observing the mirror image signals from each pair of sensors, it became apparent

that the best way to minimize the effect of the earth's thermal gradient would be to arithmetically sum the emf from a meter pair. The "C" meter pair near the sides of the pipe section was selected, since this pair is less susceptible to the earth's radial thermal gradient because of its nearly horizontal orientation.

Summing the pair signals has the net effect of minimizing the output due to a varying

external thermal gradient, and of doubling the emf due to an increase in thermal flux

through the pipe wall. The emf sum from the meters "Cl" and "C2" were plotted as shown in

figure 2.6. The net effect of summing the emf's is easily illustrated by considering the

period from 9-16-79 to 12-1-79. Meter "Cl" had an excursion of 2000 pV and meter "C2" had

an excursion of 1900 pV, whereas for the same period, the sum of "Cl" + "C2" gives an ex- cursion of 500 pV. Also, compare the effect of a step power increase. When the power was increased from 275 to 1380 mW on 1-22-80, "Cl" changed -710 pV and "C2" changed -850 pV

whereas the sum of "Cl" + "C2" =-1560 pV. Thus, the technique of summing the emf from a pair of heat-flow meters results in a signal change of twice the value for one sensor. It also gives a sizable cancelling effect of the emf due to external thermal fields.

In looking at the response of "Cl" + "C2" to step power changes, the summed signal from

a step increase in power from 1 times to 5 times predicted value is much more distinct.

Similarly, a significant decrease in signal level was experienced when the power was decreased the same amount on 3-1-80. Changes in power by a factor of 3 are not as perceivable, and could be masked by the daily variations if the thermal flux increased over a period of several days. However, if averages for several weeks are considered, changes of

32
Heat-Flow

Meter

emf,

/jlV

Time, Date

Time, Date

Figure 2.5 Field response of meters Cl and C2.

33
EMF,

MetersCt

+C

2

,

/xV

600r

400-

200 -

0 —

-2°0 -q -400-

-600 -

25 cm

snow melted

2400 r2000 -

1600-

1200 -

800- 275 mW

applied &

to pipe {

0*-

Increased power

to 1382 mW

| 25 cm snow melted -4?0

Chinook wind

- 5 cm snow melted

Increased power

\ from 275

J to 824 mW

I Decreased

[ power from I

( 1382 to 0 °

t 275 mW °o

2400r

2000 -

1600

1200

800

400

2 day rainy period considerable rain.

Very wet

0

spring snow

storm; ~ 5 cm of wet slush on surface

Power level = 824 mW

Heat flux = 3x

Normal predicted value of 28/xW/cm2

J

Power decreased

Pipe section excavated

DATE

Figure 2.6 Effect of summing signals of horizontally opposed sensors.

34
one or two times the nominal heat flux are detectable. The summing technique does not de- crease the perturbations in meter signals when a heavy snow melt occurs. The technique of summing the emf from all six functional meters was also investigated,

but this was not nearly as well behaved as the plot of "Cl" + "C2". The sum of the six

heat-flow meters ranged from +700 pVolts to 1700 pV for the period 9-16-79 to 1-4-80. This

is considerably more than the 500 pV for ("Cl" + "C2") for the same period.

In the summary, it appears feasible that a pair (or several pairs) of heat-flow meters

mounted on the sides of an undergound cryogenic envelope could be used as a insulation degradation sensor for thermal flux increase on the order of once or twice nominal values. In any field application one should bear in mind that the meters are very sensitive to extreme

weather conditions capable of disturbing the thermal gradient of the earth near the pipe

such as a heavy snow melt.

2.9 Time Response of the Heat-Flow Meters

The response time of the heat-flow meters was monitored during the step power increase

to the pipe section. Figure 2.7 shows the time response of the "C" meter pair to a power

increase from 275 to 1380 mW (predicted to 5 times predicted heat flux). The time constant

is 18 minutes. So in spite of the large mass of the steel pipe section, the time response

is more than adequate for sensing vacuum degradation over a relatively short time period.

2.10 Comparison of a Meter Thermocouple to an Earth Thermocouple During a Snow-Melt Period

After observing the snow melt perturbations, we looked more closely at thermocouple

temperatures of the earth and meters. Figure 2.8 is a plot of the temperature of meter "Dl"

and of the sand backfill at the pipe section axis. The temperatures are the same prior to and after snow melt, but during the snow melt, the 0°C water from the surface seeps down

through the sand and decreases the sand temperature at the pipe axis 1.4°C below that of the

meter on the pipe section. Thus the meters are experiencing a 1.4°C temperature change at the outer surface which is responsible for the large signal perturbations occurring during

heavy snow melts.

2.11 Performance of the Heat-Flow Meters Above Ground

2.11.1 Non-Shaded

The pipe section was excavated on 7-29-80 and placed on a wooden stand with the pipe

axis 75 cm above the surface. Sensors "Dl" and "D2" were both inoperative and badly delaminated, evidently due to moisture in the earth during the burial period.

The emf's from the meter pairs "A", "B", and "C", were monitored (with zero power input) for several days to determine the magnitude of the spurious thermal energy subjected to

a pipe section above ground.

A plot of emf from the "C" pair of sensors on the side of the pipe section is shown in

figure 2.9. This plot is typical for all three pair of sensors. The variation in sensor

emf during the (non-daylight) hours is considerably less than the variation during the

35
2400

2000

^ 1600

CJ O

"t 1200

o

u.2

w 800

400

Increase Power Level

to 1375 mW

O O O O o° °o 00

o o °qOO

1 h/cm

J L 1

<o

V“

I

egeg

l

co sr eg eg

i t o

eo

0 1 2 3 4 5 6 24

Figuer 2.7 Transient response of meters Cl and C2.

10

8

©k.3

OSh.® - a 6

E©

b~

Jo

c

toE

o

ineg

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®

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4 —

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1 0-

1 i

(O

y—

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00 O

eg

eg

eg

Date

Figure 2.8 Comparison of earth and sensor temperatures during a snow melt.

36
Heat-Flow

Meter

emf,

mV

Heat-Flow

Meter

emf,

mV

200

A Meter C t O Meter C 2

8 pm 12 4 8 12 4 8 12 4 8 12 4 8 12 4 8 am

midnight noon midnight noon midnight

7-31-80 8-1

Time

8-2 8-3

Figure 2.9 Above ground response of meters Cl and C2 - unshaded arrangement

8pm 12 4 8 12 4 8 12 4 8 12 4 8 12 4 8 am

midnight noon midnight noon midnight

8-8-80 8-9 8-10 8-11

Time

Figure 2.10 Above ground reponse of meters Cl and C2 - shaded arrangement.

37
daylight hours when the pipe section is subject to the radiant energy of the sun. The

variation of the "B" and "C" pairs of sensors was approximately + 100 mV during the day, and

approximately 20mV during the night. For a predicted sensor emf of 250 |iV, the above variations give noise to signal ratios of 400 and 80 respectively.

2.11.2 Shaded Cover

Because of the large variations in sensor emf during the daylight hours, a shade was placed almost completely around the pipe section in order to decrease the amount of solar

radiation to the heat-flow meters. The shade consisted of one layer of 635 |jm thick double

aluminized Mylar. Once again, the power input was zero. The emf from the "C" pair of sensors is shown in figure 2.10.

The shade decreased the sensor variation by approximately an order of magnitude during

the day. There was only a slight decrease in variation during the night. If we neglect the

two spikes (due to short term weather disturbance) that occur in figure 2.10, and assign an approximate signal variation of ± 10 mV to the sensors, we still obtain a noise to signal

ratio of 40 for a predicted sensor emf of 250 ^V. Thus, even with shading, and considering

only the non-daylight variations, the noise to signal ratio is still excessive for detecting

increases in envelope heat flux in the range up to 40 times predicted value.

2.12 Conclusions

a. All of the heat-flow meters tested had adequate sensitivity to sense changes in thermal flux in the predicted range of 28 [iW/cm 2

.

b. We experienced no physical problems with the meters of solid plastic construction. However, for meters of laminar construction, one pair failed completely, and

the other pair suffered no damage during the 11 month undergound test period.

c. Both above and below ground, the meters are sensitive to anything that disturbs or changes the thermal gradient of the nearby surroundings; such as solar radiation

above ground or moisture below ground.

d. The instrumentation necessary to measure the emf from the meters must be high

quality and capable of measurements in the microvolt range. The instrumentation

leads between the meters and the data acquisition equipment must be carefully

shiel ded.

e. Heat-flow meters can be used as non-intrusive thermal flux meters for sensing

insulation degradation within an undergound cryogenic envelope. The optimum

configuration is to place a pair of meters 180° opposed on a horizontal line thru

the mid-section of the pipe. Our tests indicate that thermal flux increases of one

or two times the nominal 28^iW/cm 2 should be detectable.

f. The successful use of heat-flow meters to sense insulation degradation in a cryogenic envelope outside, above ground, appears doubtful. Spurious thermal

energy is significantly larger (40 times) than predicted heat leak through a high

quality, vacuum insulated cryogenic envelope.

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2.13 References

[1] Daney, D. E., Helium Research in Support of Superconducting Power Transmission, NBSIR

80-1637, National Bureau of Standards (U.S.), 1980.

[2] Carslaw, H. S. and Jaeger, J. C., Conduction of Heat in Solids, 2nd Edition, Clarendon

Press, 1959.

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