ASHRAE 95-P

SOLAR DOMESTIC

HOT WATER

TEST FACILITY

NATL INST OF STANDARDS & TECH R.I.C.

A1 11 021 54499

/NBS building science series _

TA435 U58 ^140;1982 C.2 NBS-PUB-C 1974

NBS BUILDING SCIENCE SERIES 140

Analytical and Experimental Analysis

of Procedures for Testing Solar

Domestic Hot Water Systems

U.S. DEPARTMENT OF COMMERCE • NATIONAL BUREAU OF STANDARDS
NATIONAL BUREAU OF STANDARDS

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HATlOtJAL BUREAU OP STA»DAr-OS

LI (.'>Apy NBS BUILDING SCIENCE SERIES 140 may io 1933

Analytical and Experimental Analysis

of Procedures for Testing Solar

Domestic Hot Water Systems

A. H. Fanneyi

W. C. Thomas^

C. A. Scarbrough^

C. P. Terlizzii

^Center for Building Technology

National Engineering Laboratory

National Bureau of Standards

Washington, DC 20234

-Department of Mechanical Engineering

Virginia Polytechnic Institute and State University

Blacksburg, VA 24061

Sponsored by the Office of Solar Applications for Buildings

U.S. Department of Energy

Washington, DC 20585

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

NATIONAL BUREAU OF STANDARDS, Ernest Ambler, Director

Issued February 1982
Library of Congress Catalog Card Number: 81-600191

National Bureau of Standards Building Science Series 140

Nat. Bur. Stand. (U.S.), Bldg. Sci. Ser. 140, 158 pages (Feb. 1982)

CODEN: BSSNBV

U.S. GOVERNMENT PRINTING OFFICE

WASHINGTON: 1982

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

A repeatable test method independent of outdoor environmental conditions and

laboratory geographical location is required in order to provide a means by

which solar domestic hot water systems may be rated and compared. Three

experimental techniques which allow the net thermal output of an irradiated

solar collector array to be reproduced indoors without the use of a solar

simulator are investigated. These techniques include use of an electric heat source only, use of a nonirradiated collector array in series with an electric

heat source, and the use of electric strip heaters which are attached to the

back of nonirradiated absorber plates. Expressions are developed to compute

the input power required for each experimental technique. Solar collectors

connected in parallel and series combinations are considered.

All three test techniques were shown to reproduce the outdoor daily collector

array thermal output within four percent. Two of the techniques allow the

actions of the circulator controller for an outdoor irradiated system to be

duplicated indoors. One technique applies to solar hot water systems which

operate on the thermosyphon principle.

Experiments conducted to determine the effect of storage tank temperature

stratification on system performance for a single-tank direct solar hot water

system are described. Several return tube designs, which introduce the solarheated water into the storage tank, were fabricated and tested to determine the

Influence of thermal stratification on system performance. The best return

tube design increased the performance of a single-tank direct system approximately ten percent compared to a conventional return tube design.

An analytical model for a single-tank direct hot water system is developed.

The model is used to support parametric studies for the thermal performance

characteristics which result from the use of each test method to duplicate the

net thermal output of an irradiated array. The model is also used to assess thermal performance differences which occur due to indoor versus outdoor

environmental conditions.

Key words: ASHRAE Standard 95; collectors in parallel; electric strip heaters;

environmental conditions; indoor testing; modeling; NBS; solar;

solar domestic hot water system; stratification; test method.

iii
PREFACE

This reseach was sponsored by the Office of Solar Applications for Buildings,

U.S. Department of Energy, Washington, D.C. 20585.

DISCLAIMER

Certain commercial equipment, instruments, or materials are identified in this

paper in order to adequately specify the experimental procedure. Such identification does not imply recommendation or endorsement by the National Bureau

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

iv
TABLE OF CONTENTS

Page

LIST OF FIGURES vi

LIST OF TABLES x

NOMENCLATURE xii

Chapter

1. INTRODUCTION AND BACKGROUND 1

2. ANALYSIS 5

3. EXPERIMENTAL APPARATUS AND INSTRUMENTATION 25

4. EXPERIMENTAL PROCEDURE, RESULTS, AND DISCUSSION 57

5. ANALYTICAL RESULTS AND DISCUSSION 99

6. CONCLUSIONS AND RECOMMENDATIONS 123

REFERENCES 126

APPEtroiX A A-1

ASHRAE STANDARD 95-P COLLECTOR HEAT REMOVAL FACTOR

MODIFICATION PROCEDURE

APPENDIX B B-1

HEAT LOSS ANALYSIS

CALCULATION OF FILM CONDUCTANCE BETWEEN THE TEST FLUID

AND FLOW CHANNEL WALL

APPENDIX C C-1

DIMENSIONS AND PROPERTIES OF SOLAR COLLECTOR

INSTRUMENTATION SPECIFICATIONS

DATA CHANNEL ASSIGNMENT

V
LIST OF FIGURES

Page

Fig. 1 Nonirradiated solar collector with downstream conventional

energy source 9

Fig. 2 Nonirradiated solar collector with upstream conventional

energy source 10

Fig. 3 M parallel rows of N collectors connected in series 12

Fig. 4 Solar storage tank model 16

Fig. 5 Single-tank direct system schematic 27

Fig. 6 Flat-plate liquid collector 28

Fig. 7 Hot water storage tank 29

Fig. 8 Circulator performance characteristics 30

Fig. 9 Solar collector plumbing schematic 33

Fig. 10 Solar hot water system with the irradiated collector

array replaced with an electric heat source 34

Fig. 11 Electric heat source 35

Fig. 12 Pressure drop across the irradiated collector array and across the electric heat source 36

Fig. 13 Solar hot water system with the irradiated collector

array replaced with a nonirradiated collector array with

downstream heat source 37

Fig. 14 Location of strip heaters on the back of solar collector

absorber plate 41

Fig. 15 Schematic of supply water conditioning loop 42

Fig. 16 Vertical thermocouple array construction details 46

Fig. 17 Location of absorber plate thermocouples 47

Fig. 18 Volumetric flowmeter schematic 48

Fig. 19 Typical calibration curve for volumetric flowmeter 51

Fig. 20 Thermal output comparison of irradiated collector array for

Aug. 7, 1979, and electric heat source for Sept. 4, 1979 59

vi
LIST OF FIGURES (cont.)

Page

Fig. 21 Thermal output comparison of irradiated collector array for

Aug. 8, 1979, and electric heat source for Sept. 5, 1979 60

Fig. 22 Thermal output comparison of irradiated collector array for

Aug. 10, 1979, and electric heat source for Sept. 7, 1979 61

Fig. 23 Thermal output comparison of irradiated collector array for

Aug. 13, 1979, and electric heat source for Sept. 8, 1979 62

Fig. 24 Thermal output comparison of irradiated collector array for

Oct. 6, 1980, and nonirradiated collector array with downstream heat source for Nov. 17 , 1980 67

Fig. 25 Thermal output comparison of irradiated collector array for

Oct. 8, 1980, and nonirradiated collector array with downstream heat source for Nov. 25 , 1980 68

Fig. 26 Controller operation comparison of irradiated collector array

for Oct. 6, 1980, and nonirradiated collector array with downstream heat source for Nov. 17 , 1980 69

Fig. 27 Controller operation comparison of irradiated collector array

for Oct. 8, 1980, and nonirradiated collector array with downstream heat source for Nov. 25 , 1980 70

Fig. 28 Thermal output comparison of irradiated collector array for

Oct. 6, 1980. and nonirradiated collector array with attached

strip heaters for April 1, 1981 73

Fig. 29 Thermal output comparison of irradiated collector array for

Oct. 8, 1980, and nonirradiated collector array with attached

strip heaters for April 2, 1981 74

Fig. 30 Controller operation comparison of irradiated collector array

for Oct. 6, 1980, and nonirradiated collector array with

attached strip heaters for April 1, 1981 76

Fig. 31 Controller operation comparison of irradiated collector array

for Oct. 8, 1980, and nonirradiated collector array with

attached strip heaters for April 2, 1981 77

Fig. 32 Return tube construction details 79

Fig. 33 Storage tank temperature distribution for the return tube

shown in figure 32 80

Fig. 34 Return tube construction details 81

vii
LIST OF FIGURES (cont.)

Page

Fig. 35 Storage tank temperature distribution for the return tube

shown in figure 34 82

Fig. 36 Construction details of a commercially available return tube .. 83

Fig. 37 Storage tank temperature distribution for the commercially

available return tube 84

Fig. 38 Return tube construction details 85

Fig. 39 Return tube construction details 86

Fig. 40 Return tube construction details 87

Fig. 41 Storage tank temperature distribution for the return tube

shown in figure 38 88

Fig. 42 Storage tank temperature distribution for the return tube

shown in figure 39 89

Fig. 43 Storage tank temperature distribution for the return tube

shown in figure 40 90

Fig. 44 Return tube construction details 91

Fig. 45 Storage tank temperature distribution for the return tube

shown in figure 44 92

Fig. 46 Storage tank temperature distribution for stratification

sensitivity test, Feb. 3, 1981 94

Fig. 47 Storage tank temperature distribution for stratification

sensitivity test, Jan. 29, 1981 95

Fig. 48 Solar collector efficiency curve 101

Fig. 49 Experimental and two-node analytical model predicted storage

tank temperature distribution for Oct. 8, 1980 107

Fig. 50 Experimental and five-node analytical model predicted storage

tank temperature distribution for Oct. 8, 1980 108

Fig. 51 Recorded wind speed for Oct. 6, 1980 113

Fig. 52 Recorded ambient temperature and effective sky temperature for

Oct. 6, 1980 114

vlii
LIST OF FIGURES (cont.)

Page

Fig. 53 Recorded wind speed for Oct. 8, 1980 116

Fig. 54 Recorded ambient temperature and effective sky temperature for

Oct. 8, 1980 117

Fig. 55 Thermal efficiency of electric heat source 119

Fig. 56 Solar and infrared wavelength radiosity distributions in a solar collector with one cover B-2

ix
LIST OF TABLES

Page

Table 1 Comparison of System Performance Using an Irradiated

Collector Array to System Performance Using an Electric

Heat Source 64

Table 2 Comparison of System Performance Using an Irradiated

Array to System Performance Using a Nonirradiated Array

with Downstream Heat Source 66

Table 3 Comparison of System Performance Using an Irradiated

Array to System Performance Using a Nonirradiated Array

with Attached Strip Heaters 72

Table 4 Comparison of System Performance for Two Return Tube

Designs 96

Table 5 Repeatability of Test Results Using ASHRAE STANDARD 95 97

Table 6 Comparison of Experimental Results with Analytical Model

Predicted Results for August 7 , 1979 102

Table 7 Comparison of Experimental Results with Analytical Model

Predicted Results for August 13, 1979 103

Table 8 Comparison of Experimental Results with Analytical Model

Predicted Results for October 6, 1980 104

Table 9 Comparison of Experimental Results with Analytical Model

Predicted Results for October 8, 1980 105

Table 10 Effect of Test Method on SDHW System Performance; Analytical

Model Results Using October 6, 1980, Meteorological

Conditions 110

Table 11 Effect of Test Method on SDHW System Performance; Analytical

Model Results Using October 8, 1980^ Meteorological

Conditions Ill

Table 12 Effect of Wind Speed and Effective Sky Temperature on SDHW

System Performance; Analytical Model Results for Various

Test Methods, October 6, 1980, Meteorological Conditions 115

Table 13 Effect of Wind Speed and Effective Sky Temperature on SDHW

System Performance; Analytical Model Results for Various

Test Methods, October 8, 1980, Meteorological Conditions 118

X
LIST OF TABLES (cont.)

Page

Table lA Effect of Heat Source Thermal Loss on SDHW System Performance

for Nonirradiated Array with Downstream Heat Source;

Analytical Model Results Using Two-Node Storage Tank for

October 6, 1980 121

Table 15 Effect of Heat Source Thermal Loss on SDHW System Performance

for Nonirradiated Array with Downstream Heat Source;

Analytical Model Results Using Two-Node Storage Tank for

October 8, 1980 122

xi
A- Collector aperture area, m

NOMENCLATURE

2

2 A^ Cross-sectional area of storage tank, m

Ag Collector edge area,

Ag Collector gross area, m

Agj^ Surface area of storage tank segment i which experiences heat

loss to the environment, —^ F (ra) Intercept of the collector efficiency curve determined in

Ag R n accordance with ASHRAE 93-77, dimensionless

^a -p jj Slope of the collector efficiency curve determined in

Ag R L accordance with ASHRAE 93-77, W/(m2c)

'-''-'CON Daily combustion heat loss from a conventional water heater,

kJ

Cp Specific heat, kJ/(kg C)

Cp ^ Specific heat of the transfer fluid used in the collector

during the ASHRAE Standard 93-77 tests, kj/(kg C)

Cp g Specific heat of the transfer fluid used in the collector

during the solar hot water system test, kj/(kg C)

dj^ Hydraulic diameter of flow tube, m

D Flow tube outside diameter, m

e Fractional energy savings, dimensionless

E^^ Emissive power of cover 1, W/m'^

Ep Emissive power of absorber plate, W/m

Eg Emissive power of sky, W/m

F Fin efficiency, dimensionless

F^'^ Collector control function for node i, dimensionless

F' Collector efficiency factor, dimensionless

Fjj^ Collector heat removal factor, dimensionless

xii
Fj^ Collector heat removal factor for two collectors connected in

series, dimensionless

G Total global irradiance incident upon the aperture plane of

the collector, W/m^

G Fluid mass flow rate per unit collector aperture area, kg/Cm^s)

h£ Heat transfer coefficient between the fluid and tube wall, W/(m C)

hp(,]^ Conductance, natural convection, between absorber and cover 1,

W/(m2c)

Conductance, outside cover to ambient air, W/(m C)

^CON Daily heat loss from a conventional water heater, kJ

J Heating element control logic, dimensionless

Jjjj Solar beam radiosity, jth surface, dimensionless

J^j Solar diffuse radiosity, jth surface, dimensionless

Jj Longwave diffuse radiosity, jth surface, dimensionless

kg Thermal conductivity, edge insulation, W/(m C)

k^ Thermal conductivity, back insulation, W/(m C)

kp Thermal conductivity, absorber plate, W/(m C)

k^^^ Thermal conductivity, water, W/(m C)

Kq^j Incident angle modifier, dimensionless

L Number of time steps in test period, dimensionless; length of

absorber plate, m

Lg Thickness of edge insulation, m

Lj^ Thickness of back insulation, m

Lj^'^ Control logic function for node i, dimensionless

m Parameter defined by equation 34 , m~l

m Mass flow rate, kg/s

m^. Mass flow rate of the transfer fluid through the collector

during the ASHRAE Standard 93-77 tests, kg/s

xiii
Mass flow rate of hot water draw, kg/

nig Mass flow rate of the transfer fluid through the collector array

during the solar hot water system test, kg/

M Number of parallel rows of collectors, dimensionless

M-[ Mass of water contained in storage tank node i, kg

n Number of time steps during which hot water draws occur,

dimensionless

N Number of collectors connected in series, dimensionless;

number of segments into which the storage tank is divided,

dimensionless

Nu Nusselt number, dimensionless

Pr Prandtl number, dimensionless

Wetted perimeter, m

q -, Fraction of total solar irradiance absorbed by cover, W/m

o- C X

•Q Power supplied to storage tank, W

QCON Daily energy supplied to a conventional water heater, kJ

Qjlj; Thermal output of storage tank heating element, W

Qjjg Daily energy consumed by storage tank heating element, kJ

Q]^g Thermal output of conventional heat source, W

.* Q^g Thermal output of conventional heat source when located

upstream of a nonirradiated array, W

Qj^ Daily energy extracted from storage tank, kJ

Qg^ Daily thermal output of irradiated collector array, kJ

Qg^ Thermal output of electric strip heaters, W

Qgj Heat loss from the section of pipe located in the laboratory

which supplies the collector array, W

Qg-p Change in storage tank internal energy during the test day,

kJ

Daily energy supplied to storage tank from solar collector

array, kJ

xiv
Heat loss per unit time and aperture area throuugh absorber-cover

assembly, W/m^

Net thermal output of collector array, W

Reynolds number, dimensionless

Quantity of solar irradiance absorbed by the absorber plate of a single collector, W

Ambient temperature, C

Laboratory ambient temperature, C

Fluid temperature leaving the collector array, C

Temperature, mean, cover 1, C

Fluid temperature entering the collector array, C

Fluid temperature entering the nonirradiated collector array with

an upstream heat source, C

Mean fluid temperature, C

Upper temperature limit for the heating element thermostat, C

Average fluid temperature within the heat source, C

Fluid temperature entering the heat source, C

Fluid temperature leaving the heat source , C

Temperature of storage tank segment i, C

Predicted future temperature of storage tank segment i, C

Lower temperature limit for the heating element thermostat, C

Temperature of supply water entering the storage tank, C

Mean absorber plate temperature , C

Mean absorber plate temperature of a nonirradiated collector, C

Mean absorber plate temperature of a nonirradiated collector with

upstream heat source, C

Temperature of the solar-heated fluid as it enters the storage tank,

C

XV
tgg|. Ultimate desired hot water delivery temperature, C

tg^ Temperature of the collector supply fluid as it leaves the Interior

of the laboratory, C

Temperature, tube wall, C

Tg]^y Effective sky temperature, K

Atj^ Log-mean temperature difference, C

o

U|j Collector back and edge loss coefficient, W/(m C)

o

U-|^ Collector overall loss coefficient, W/(m C)

Uj^^ Storage tank loss coefficient, W/(m^C)

Uip Collector top loss coefficient, W/Cm'^C)

(UA)^g Heat loss coefficient-area product for heat source, W/C

(UA)j^j Heat loss coefficient-area product for interior portion of solar

collector array return pipe, W/

(UA)j^Q Heat loss coefficient-area product for exterior portion of solar

collector array return pipe, W/C

(UA)gj- Heat loss coefficient-area product for interior portion of solar

collector array supply pipe, W/C

(UA)gQ Heat loss coefficient-area product for exterior portion of solar

collector array supply pipe, W/

V Volume of storage tank segment, m^

Wind velocity, m/s

W Tube spacing, ra

ttjj Absorptance of the collector absorber coating to the solar

spectrum at the normal incidence, dimensionless

Op Absorptance, plate, dimensionless

0 Angle of incidence between the direct solar beam and the normal to

the collector aperture, deg

6 Absorber plate thickness, m

Ti Collector efficiency based on the net area, dimensionless

xvi
n^s Heat source efficiency, dimensionless

y Fluid viscosity, kg/(m s)

p Density, kg/m^

fy^l Reflectance, solar beam irradlance, cover 1, dimensionless

p^l Reflectance, solar diffuse irradiance, cover 1, dimensionless

p^l Reflectance, longwave diffuse irradiance, cover 1, dimensionless

Pp Reflectance, longwave diffuse irradiance, absorber, dimensionless

a Stefan-Boltzmann constant, W/Cm^ K^)

1^2. Transmittance , solar beam irradiance, cover 1, dimensionless

T(j]^ Transmittance, solar diffuse irradiance, cover 1, dimensionless

Xf^l Transmittance, longwave diffuse irradiance, cover 1, dimensionless

Tj^ Effective transmittance at normal incidence, dimensionless

(Ta)j^ Effective transmittance-absorptance product for the collector

at normal incidence, dimensionless

At Length of time step, s

xvii

ASHRAE95P

1. INTRODUCTION AND BACKGROUND

1.1 PURPOSE

The purpose of this analytical and experimental investigation is to develop a test method for solar domestic hot water (SDHW) systems. Although the primary

components of a SDHW system are the collector modules and storage tanks, for

which standards have been adopted, the interaction between components has not

been established. Consequently, a standard test method for a SDHW system is

needed. Once a test method is adopted, an industry rating association such as the Air Conditioning and Refrigeration Institute (ARI) or the Solar Energy

Industries Association (SEIA) may develop and adopt a companion rating standard

that specifies the number of tests to be conducted, the test conditions to be

used, and the manner in which performance results are to be reported.

1
A repeatable test method, independent of outdoor environmental conditions and

laboratory geographical location, is required in order to provide a means by

which SDHW systems may be rated and compared. One possible method employs a solar simulator to irradiate the solar panels which are part of the SDHW system

being tested. Another method, which is the subject of this research, requires

that the energy normally furnished by an outdoor irradiated solar collector

array be replaced by an equivalent amount of energy supplied by a conventional

energy source.

Three experimental techniques which allow the net thermal output of an irradiated collector array to be duplicated without the use of a solar simulator are described in detail. Governing equations required for each technique

are developed. Experiments performed at the National Bureau of Standards to

validate each technique are described and the results are presented. The

influence of temperature stratification on system performance and the repeatability of test results are investigated using one selected test method. A

numerical model of a single-tank direct solar hot water system is developed.

The model is used to predict the effects of different test methods and

meteorological conditions on system performance.

1.2 BACKGROUND

The Steering Committee of the American National Standards Institute on Solar

Energy Standards designated the American Society of Heating, Refrigerating and

Air-conditioning Engineers (ASHRAE) to develop a standard test method for SDHW

systems. The ASHRAE then formed Standards Project Committee 95-P in February

1977 to begin this task.

During the early deliberations of the Committee, alternative methods of testing

a complete SDHW system were proposed. One method considered would require

installation of each system to be tested in a particular location. The solar

collector array of the system would be subjected to normal outdoor meteorological conditions. System performance would be monitored during the test interval.

This method was eliminated due to the nonrepeating nature of outdoor environmental conditions.

Analytical methods based on mathematical models to predict system performance

were also considered. The programs TRNSYS [1], f-CHART [2], and Solcost [3] are widely used in predicting the thermal performance of solar hot water systems

since, for some systems, these computer codes have been shown to predict longterm performance generally in good agreement with experimental data [4]. The

credibility of these codes for general rating purposes has not been established.

Additionally, the system input parameters including the heat exchanger effectiveness, the overall heat loss coefficient for the storage tanks, and the

performance characteristics of the solar collector modules have to be measured

in order to achieve acceptable agreement with experimental results.

A third method considered [5] uses a combination of experimental results and

analytical modeling. This concept is based on conducting single-day outdoor

tests of the system and a "baseline" SDHW system. The ratio of the performance

of the test system to the baseline system is called the daily relative solar

2
rating. Long-term performance for the test SDHW system would be determined

from the daily relative solar rating, a correlation procedure based on computer

simulations, and long-term measured performance results for the baseline system.

Preliminary results for single-tank direct systems in sunny climatic regions are encouraging. Comparison of test results from different laboratories may prove

difficult as a result of the variability in the performance of baseline systems.

Difficulties may also be encountered in attempting to use a computer program,

such as TRNSYS, to model all types of SDHW systems which may be encountered.

The Standard Projects Committee considered a test method which employs a solar

simulator to irradiate the solar panels which are part of the SDHW system being

tested. The irradiance on the collectors is varied throughout the test day as specified by an industry rating association. Specified quantities of energy

are withdrawn from the storage tanks at selected times during the test period.

The test continues until system performance is nearly identical for two successive days. This method of testing is included in the final ASHRAE Standard 95,

Method of Testing to Determine the Thermal Performance of Solar Domestic Water

Heating Systems [6]. The advantages of this method include repeatability of

test conditions, independence from outdoor meteorological conditions, and

elimination of analytical modeling. Additionally, a separate test to determine

the thermal performance characteristics of an individual collector module [7]

is not required. The use of a solar simulator enables testing of thermosyphon

and integral storage systems. These systems pose formidable difficulties for some other test methods. However, among the disadvantages of using a solar

simulator are high costs, limited availability, and the physical constraint

that most solar simulators are designed to irradiate a single panel rather than

an array consisting of several panels.

Two test methods are considered employing a conventional heat source such as an in-line electrical resistance heater. One method employs a conventional heat source instead of an irradiated collector array to supply the net quantity of

power normally supplied by an irradiated array. The use of a heat source only

to supply the net thermal output of an irradiated array results in excellent

agreement between outdoor and laboratory system performance [8]. The influence

of collector performance on total system performance may be evaluated without

changing the experimental apparatus. Kotas and Wood [9] describe an apparatus

and control system to automate this testing sequence. However, there are several disadvantages to using only an electric heat source to supply the

quantity of power normally delivered by an irradiated array: 1) outdoor stagnation (no flow) conditions cannot be duplicated; 2) thermosyphon systems cannot be tested using this technique since the power input is localized rather

than distributed along the length of the absorber plates contained within the

solar collector modules; and 3) the solar industry may object to a system test

in which the collector array is not part of tne experimental apparatus.

The second heat source method employs a conventional heat source in series with

the nonirradiated SDHW system collector array. The combination of a heat source

and nonirradiated array supplies the net quantity of thermal power normally

supplied by the irradiated array. The use of a conventional heat source in

series with a nonirradiated collector array and a storage tank bypass loop

allows outdoor stagnation conditions to be simulated. The heat source may be

3
located upstream or downstream of the nonirradlated array when the proper

governing equation is used to compute the quantity of power supplied to the

heat source [10]. As before, this technique does not apply to thermosyphon

systems because it uses a localized heat source. The ASHRAE Standards Project

Committee incorporated this method of testing into ASHRAE Standard 95. Validation of this test procedure is part of the present investigation.

In another method, the problem of the localized heat source is eliminated.

Solar energy absorbed by irradiated solar collector absorber plates is replaced

with energy supplied by electric strip heaters attached to the back side of non- irradiated absorber plates. Thermosyphon systems may be tested using this

technique. Also, outdoor stagnation conditions can be duplicated within the

laboratory using this technique. The calculation procedure required to determine the quantity of power supplied to the strip heaters is independent of mass

flow rate and fluid composition. Additionally, testing of an individual collector module is not necessary if the optical properties of the collector are known. A disadvantage of this method is the requirement that the collector

modules must be disassembled to install the heaters. Since this concept was advanced by NBS after the proposed Standard 95 was submitted for approval, this

method has not yet been incorporated into the test standard. The present

investigation is concerned with establishing the feasibility and validity of

this method, which is important for testing thermosyphon systems.

4
2-2

2. ANALYSIS

Three methods were investigated which allowed the net thermal output of an irradiated collector array to be duplicated rnder nonirradiated conditions. The

first method uses an electric heat source only. The second method uses a non- irradiated collector array in series with an electric heat source. The third

method employs electric strip heaters attached to the back of nonirradiated

absorber plates. Governing equations used in determining the power supplied to

the electric heat source and strip heaters are developed in this chapter.

An analytical model is developed to simulate the performance of an entire solar

domestic hot water system. The model is used to support parametric studies of

5
the thermal performance characteristics which result from the use of each method

to duplicate the net thermal output of an irradiated array. The model is also

used to assess thermal performance differences which occur as a result of indoor versus outdoor environmental conditions.

2.1 ANALYSIS OF IRRADIATED COLLECTOR ARRAYS

Under steady-state conditions, the useful energy from a single irradiated solar

collector can be expressed in accordance with the Hottel-Whillier-Bliss (H-W-B)

analysis as

Qu = Ag[Kc^T4^FR (Ta)n G - ^

FrUl (tfi- t^) ] (1)

^g

The useful energy output of an array consisting of M collectors connected in

parallel is

Qu = MA [KecT Y- ^ ''"^n G - ^ FrUl (tfi" t^) ] (2)

^g ^g

The useful energy output of an array which includes collectors connected in

series can be represented similarly, taking into account modification of

the heat removal factor, F|^. For a single solar collector module

mC^

Fr = UL^a

1 - exp ( - mC_

(3)

Consider the case where two collectors are connected in series. Each solar

collector module experiences the same mass flow rate. However, the inlet temperature to the downstream collector is higher. The collector efficiency

factor, F' , is weakly influenced by the fluid temperature through its effect on

the convective heat transfer coefficient. The overall heat loss coefficient,

U^, is somewhat more dependent on fluid temperature, but, in this analysis,

both F' and Ul are assumed constant. With these assumptions, eq (3) becomes

mC_

Fr2 = UL2Aa

1 - exp I - mC„

(4)

Algebraic manipulation of eqs (3) and (4) gives

Fr2 = Fr 1 - 2mC.(5)

P

-»

Thus for two collectors in series

(FrUl)2 = (FrUl) 1 - 2mC

P J

(6)
and

(FR(Ta)n)2 = (FR(Ta)n)

2mC.

P

Generalizing to any number, N, of identical collectors in series

(7)

1 -

( 1 - ^^A)Aa

\ mC

and

(8)

(Fp(Ta)„)mC

(9)

For N irradiated collectors connected in series, the useful energy output is

^a . . Qu = NAg[K(,^ ^ (FR(Ta)n)N G - (FRUL)N(tfi - t^) ] (10)

Substitution of eqs (8) and (9) into eq (2) yields the useful energy output

of a solar collector array consisting of M rows of N collectors connected in

series

A A

Qu = MNAg [Ko,^ -f (FR(Ta)n)N G - -f (FRUL)N(tfi- ta)] (11)

g g

2.2 REPLACEMENT OF THE IRRADIATED ARRAY WITH A CONVENTIONAL HEAT SOURCE ONLY

A conventional heat source is used to replace the irradiated solar collector

array. This method requires that the solar collector be tested according to

ASHRAE Standard 93-77 [7] and values of Fj^Ul, Fj^CTa)^, and K^^ be determined.

The inlet fluid temperature to the heat source, t^-j^, is measured. The required

thermal output of the heat source is calculated using the above information in

conjunction with specified meteorological conditions and eq (11). It should be

noted with reference to eq (3) that the value of Fg^ is dependent upon mass flow

rate of the transfer fluid and the properties of the heat transfer fluid.

ASHRAE Standard 95-P requires that the fluid used in the system test be identical to the one used in the required collector test. If the flow rate of the

transfer fluid through the collector during the system test is different from

the value used in the collector test, the value of Fj^ must be modified.

Appendix A gives the ASHRAE Standard 95-P modification procedure.

7
2.3 REPLACEMENT OF THE IRRADIATED ARRAY WITH A CONVENTIONAL HEAT SOURCE IN

SERIES WITH A NONIRRADIATED ARRAY

This method employs a conventional heat source upstream or downstream of a nonirradiated solar collector array. The useful energy output of a single

collector may be expressed as

Q

"T" = Kai ( Ta)n G - UL(tp - ta) (12)

Equating eqs (1) and (12) and solving for tp

tp = Fr (tfi - ta)

^

(1 - Fr) + tca (13)

Xj For a nonirradiated solar collector, eq (13) reduces to

= \ ^'fi - 'aP + ^a£ d^)

where tpj^ is the mean plate temperature of a nonirradiated collector. The

above equations can be used to calculate the necessary thermal output of a heat

source used in series with a nonirradiated array..

Consider a conventional heat source located downstream of a single nonirradiated solar collector, as shown in figure 1. In order for the net thermal output of a nonirradiated solar collector and heat source to be equivalent to the

output of a normally irradiated array, the following is required

Qhs

I

v-pn - Ul (tpn - ta£) = Fr( Ta)^ G - FrUl (tfi - tg) (15)

a The left side of eq (15) represents the net energy output from the collector

simulator. The right side of eq (15) represents the net output that would

occur from an irradiated array. Since the heat source is downstream of the

nonirradiated array, the inlet fluid temperature to the nonirradiated array

is identical to the inlet fluid temperature for the corresponding irradiated

array. Consequently, eq (14) can be introduced into the left side of eq (15)

and the resulting equation solved for Q|^g .A A

Qhs =

"T" Fr ( Ta)n G Ag - FrUl (ta£ - ta)Ag (16)

Another possibility is to locate the conventional heat source upstream of the

nonirradiated collector as depicted in figure 2. Equating the net energy

output from the nonirradiated collector with upstream heat source to that of an irradiated collector yields

• A - UL(tpn - ta£) = Fr( Ta)^ G - FrUl (tfi - ta) (17)

^a

8

NONIRRADIATED

SOLAR COLLECTOR

fi

CONVENTIONAL

HEAT SOURCE

Figure 2. Nonirradiated solar collector with upstream

conventional energy source

10
The asterisk Indicates that tp„ is different from the previous case because the

collector loop heater is now upstream of the irradiated collector. Equation

(14) is still valid except that the inlet fluid temperature to the nonirradiated

collector array is different from the previous case as a result of the electric

heat source location

^pn = \ ^'*fi - ^aP 'ai (18)

An energy balance on the heat source results in the following expression for

the exit fluid temperature from the heater in terms of the inlet fluid temperature to the heater

Solving eqs (17), (18), and (19) simultaneously

A A„

^"-^ aJ FR('^«)n G Ag - ^ (FrUl) (t^^- t^) Ag

Qhs= (20)

A

1 -

Ag R L g

mCp

This analysis is readily extended to combinations of M parallel rows of N

collectors connected in series as shown in figure 3. Consider the situation

where there are M collectors connected in parallel and no collectors connected in

series, (N=l). Equation (16) becomes

Qhs = Kax 7^ Fr( ^a)n GMAg - ^

FrUl (^aZ " ta) MAg (21)

^g

and eq (20) becomes

^"^T^ Fr( Ta)n GMAg - ^ (FrUl) (tail" ta)MAg

Qhs = g 8 (22)

Aa

1 -

Ag ^R^L^g

mCp

Collectors connected in series require that the heat removal factor, Fr, be

modified according to eqs (8) and (9). Therefore, in the most general case where the heat source is located downstream of M rows of N collectors connected

in series,

Qhs = Kax

^

(Fr( '^«)n)N GMNAg - ^ (FRUL)N(ta£ " ta) MNAg (23)

^g ^g

11
Figure 3. M parallel rows of N collectors connected in series

12
With the heat source located upstream of M rows of N collectors connected in

series, eq (22) becomes

K^^^

(p^(^^)^)^ GMNAg - (FRUL)N(ta£- t^)mAg

%s = i g (24)

1 - ^

(FrUl)n MNAg

mCp

Therefore, the proper quantity of energy to be supplied to a conventional heat source in series with a nonirradiated array may be computed using eq (23) or

eq (24).

2.4 REPLACEMENT OF AN IRRADIATED ARRAY WITH STRIP HEATERS ATTACHED TO

NONIRRADIATED SOLAR COLLECTORS

Solar energy absorbed by an irradiated collector array is replaced by energy

supplied by electric strip heaters attached to the back of nonirradiated absorber

plates. The quantity of solar energy absorbed by the absorber plate of a single

collector is

S = Aa Kc(T ( Ta)n G (25)

Thus, the amount of energy to be supplied by the strip heaters can be expressed

as

Qsh = Kar (Ta)n G (26)

Simulation of a collector array consisting of M parallel rows of N collectors

connected in series requires that

Qsh = MNAa K„x (Ta)n G (27)

Input values of energy into the strip heaters are calculated using the

prescribed radiation values, G, and the optical characteristics of the collectors. Thus, if the values of ( Ta)^ and K^tx are known, an ASHRAE 93-77 test is

not required

.

2.5 SOLAR DOMESTIC HOT WATER SYSTEM MODEL DEVELOPMENT

An analytical model has been developed for a solar domestic hot water system.

A discussion of each model component follows.

2.5.1 Solar Collectors

The collector model is based on the Hottel-Whillier-Bliss (H-W-B) analysis.

The following summary of analytical expressions for a single-cover flat-plate

collector is directly from reference 11.

13
Collector efficiency (based on net area) is given by

n = —^ X 100% (28)

G Aa

where G is the total solar irradiance on the outer surface of the cover. The

useful energy, Q^, is given by

Qu = AaFR {K^T (Ta)r^ G-Ul [tfi - t^] } (29)

The loss coefficient, U^, is the sum of the top loss coefficient U|- and the back

(and edge) loss coefficient U^. An initial estimate of the back conductance \J\y IS

L-r LA 1 e a (30)

The expressions and procedure for determining U^- are given in Appendix B. Since

depends on the mean absorber plate temperature, the following expression is

used in an iterative approach

tp = Fr [tfi - ta + (Ta)n G/Ul] + - (Ta)n G/Ul (31)

It should be noted that, in the H-W-B analysis, Uj^ is assumed constant over the

temperature range tp(0) to tp(L) and accounts for heat losses by convection and

longwave radiation. The expression for the collector flow factor Fr depends on

the flow tube-absorber plate configuration. The tube-plate configuration is

reflected in the expressions for the efficiency factor F' which accounts for

tubes below, above, and integral with the plate. The values of F' for these

three configurations differ only if the thermal bond between the tube and plate

is poor. Since the bond resistance is negligible for the collectors under con- sideration, the expression for the efficiency factor is taken as

P^hf D + [W-D]F

where

Y

= tanh (m[W-D]/2) m [W-D]/2

m = (UL/[kp6])l/2 (34)

For the parallel-flow collectors which are being modeled,

Fr = [GCp/UL] [1 - exp (-F'UL/{GCp})] (35)

14
and

tfm = [Fr/F'] [tfi - + (Ta)n G/Uj^] + K^^ ( xa) ^ G/Ul (36)

The heat transfer coefficient, hf, calculated using the procedure outlined in

Appendix B, is based on the mean fluid temperature which is determined by using

eq (36) in an iterative procedure.

The algorithm for calculating the thermal performance of a collector is organized

as follows:

Step 1 . The operating conditions, design parameters and heat transfer properties,

and fluid inlet temperature are specified.

Step 2 . Collector and material parameters which do not vary with temperature

are calculated. An initial estimate of tp, in terms of tfj[, is made.

Step 3 . With the current estimate of tp, the heat loss through the cover assembly and U^- are determined ( in a subroutine). Next, Ul and Fr are calculated. An estimate of tjjjj, in terms of tp, is made.

Step 4 . With the current estimate for tfjjj, hf is calculated in a separate

subroutine using a supporting fluids properties routine. The factors

F' and Fj^ are calculated from eqs (32) and (35) to obtain an improved

estimate of tfjjj, and Step 4 is repeated until tf^ converges.

Step 5 . Equation (31) is used to improve the estimate of tp . Steps 3, 4, and

5 are repeated in an iterative procedure until tp converges.

Step 6. The outputs Q^^ and n are calculated from eqs (29) and (28), respectively,

along with other parameters of interest.

2.5.2 Storage Tank

A one-dimensional heat transfer analysis is performed on the storage tank in

the manner outlined in reference 12. The tank is divided vertically into N

elements as illustrated in figure 4. An energy balance on the element i yields

the following differential equation

(PV^i IT = (^Vi'^i

^

^^i

^

(37)

- Ult ^si (ti - ta£) + Q

The left hand side of eq (37) represents the increase in the internal energy of

the element. The first term on the right hand side represents the convective

heat transfer - within the tank, the second terra heat transfer due to conduction,

the third term heat loss to the surroundings, and the fourth term external

energy addition to the element. The conduction term was found to be negligible

and is ignored in this analysis. Thus eq (37) becomes

15
COLLECTOR

RETURN

HEAT

LOSS

COLD WATER

SUPPLY

m.

^ HOT WATER

SUPPLY

1 O

2 o

3 o

i- 1 o

i o

i+l o

o

N-l o

N o

HE

COLLECTOR

SUPPLY

Figure 4. Solar storage tank model

16
^^S^^i JT ^ (^Vi^^^i

^- "lT ^si (^i - tail) + Q (38)

Expressing the partial derivatives in eq (38) in terms of finite difference

approximations results in the following finite-difference equation

(PSV)i-\:^ = ('^Vl '';^y^'' - UltAAsI (ti - t^O + Q (39)

The volume of the i*^'^ node may be expressed as

Vi = Ay^ Ac (40)

where A,, is the cross-sectional area of the tank.

Substitution of eq (40) into eq (39) gives

(pCpAe Ay)i = (m Cp) i Ay^ -^^^^ " ^si (^i " t^^) + Q (41)

Mass flow rate through a node can be expressed as the sum of two components,

• • • "1 = (^L - ^cH (42)

• • where m^ is the mass flow rate circulating through the collector array, and m-^ is the hot water load mass flow rate. Substitution of eq (42) into (41) and

subsequent rearrangement yields the general equation

(^S>c + ('"Cp^L ULTAAgi ^^-—^ + ^ (43)

t."-t.

= (PC A ). ^ ^ - P c 1

This generalized analysis is applied to the particular storage tank used in

this investigation. A thermostatically controlled electric heating element

located in the storage tank maintains the water above it within a specified

temperature range. Hot water is withdrawn from the tank upper node, i = 1, and

makeup water enters the bottom node, i = N. Supply water to the collector

array is withdrawn from the bottom node. It is assumed that the returning

heated water enters the tank at a location such that

ti-1 > tR > ti (44)

where t^ is the temperature of the returning fluid as it enters the storage

tank.

17
Applying eq (43) to the upper portion of the storage tank, i = 1, yields

(mC )^ F^^ (t^ - t^) + (mC )l (t^ - t^) - U^tAA^. (t^ - t^p

t t + JQpp = (pC V) ^

(45)

^HE ^^"p^a At

where F'^ and J are control logic functions.

The collector control function for this node is defined as

1 if tj^ >^ ti

F^^ =

I (46)

0 if tR < ti

The heating control element function is defined as

1 if ti < tL

J = \ (47)

0 if ti > tji

where t-^ and tjj are the "turn-on" and "turn-off" temperature levels of the

thermostat. Applying eq (43) to nodes 2 through N-1 results in

(mCp)^ F.^ (t^ - t.) + (i;Cp)^ L.^ (ti_i-ti) + (mCp)^ (ti+i'ti)

^ n_ ^ (48) - ULT^siC^i - t^l) = (PV^i \,

^

The control logic functions are

1 if ti_l > tR > ti

c

i

F,.- = \ (49)

0 otherwise

and

1 if tR > ti_l

^i" = \ (50)

0 otherwise

18
Applying eq (43) to the bottom node of the tank, i = N, yields

(liCp)^ V - t^) + (icp)^ L/(tN-i-tN) + (^^ph (^m -

^n)

t.^-t, (51)

- Ult^SN - ta£) = (PV^N-V^

where the control logic functions are

1 if tN_i > tR > tN

= (52)

and

0 otherwise

1 if tR > txT- R ^ i^N-l

=

{ (53)

0 otherwise

Modeling of the entire storage tank is achieved with the system of equations

developed above.

2.5.3 Heat Source

The solar domestic hot water system model includes subroutines which allow the

modeling of a hot water system where the irradiated array has been replaced

with an electric heat source or an electric heat source used in series with a nonirradiated array. This section deals with the analytical model for the heat source.

Energy supplied to a conventional heat source used in series with nonirradiated

collectors is computed using eq (23) or eq (24). An energy balance for the

heat source under steady-state conditions yields

Qhs = (tho - tt,i) + (UA)^s (T^ - t^s) (54)

where

\ - ^^^^-^

(55)

Defining the efficiency of a heat source as

%s = ^ A (56)

^hs

19
gives

tho = \c^^ + thi (57)

where r^^g = 1 corresponds to an adiabatic heat source. The product (UA)^g was determined experimentally. Equation (57) is used in the simulation model for

tho2.5.4 Pipe Thermal Losses

The model accounts for heat losses from the collector array supply and return

pipes. The supply pipe is divided into two sections, indoors and outdoors, for

the heat loss analysis. Thermal losses for the section located within the test

facility may be expressed as

QSI = (UA)si At^ (58)

The log-mean temperature difference is defined as

li r ^sw " ta£ "|

Substitution of eq (59) into eq (58) yields

Qsi = (UA)si rtl,"-'t,.-i (60)

In

tsw ~ ^al

The thermal losses may also be expressed as

QSI = '^S (tN - ts„) (61)

Combining eqs (60) and (61) yields

/ (UA)_sj\

'P sw = ^aZ + (tN - ta£) exp I -

) (62)

Equation (62) predicts the supply fluid temperature as it leaves the interior

of the test facility. With this temperature, the temperature entering the solar

collector array is predicted in an identical manner, i.e.

/ (UA),,^ \

tfi = ta + (tsw - ta) exp ( - 1 (63)

Substitution of eq (62) into eq (63) allows the collector array inlet fluid

temperature to be calculated in terms of the fluid temperature entering the pipe

and the indoor and outdoor surrounding ambient temperatures. Thus,

20
tfi = t, 1 - exp ( - (UA) SO

P /J

tfj exp -

(UA)g;^ + (UA)so

mC_

+ t exp - (UA)

mC

1 - exp

(UA)

mC.

P U

(64)

The return pipe from the outlet of the solar collector array to the storage

tank also experiences heat loss. The fluid temperature returning to the storage tank can be expressed in terms of the exit fluid temperature of the array

and the interior and exterior ambient temperatures

tp = t ai

+ t.

1 - exp

exp

(UA) RI

(UA)

mC.

RI

+ t CO exp

(UA)^Q + (UA) KL

AC.

mC

1 - exp

(UA) RO

mC

P / J J

(65)

The overall heat loss coefficients used were calculated using the measured

length and diameter of each section, the thickness of the insulating material,

and the values of thermal conductivity as supplied by the manufacturer. The use of an electric heat source upstream or downstream of the nonirradiated

array introduces a small amount of additional interconnecting pipe. However,

calculations show that the additional thermal pipe losses are negligible.

2.5.5 Computation Sequence For An Irradiated System

The various component models discussed above, in conjunction with recorded

meteorological data, allow simulation of the solar domestic hot water system.

Solar collector, storage tank, and piping model components are used when

simulating an irradiated system. Nodes within the storage tank are set to an initial temperature. Collector array flow rate and circulator pump status

recorded during operation of the actual system are supplied to the model.

Equation (64) predicts the fluid temperature entering the collector array using

recorded indoor and outdoor ambient temperatures. The collector model, using

measured meteorological data, calculates the fluid temperature leaving the

array. The fluid temperature entering the tank is predicted using eq (65).

Storage tank nodal temperatures are recalculated every minute with eqs (45),

(48), and (51).

In the actual system, a hot water withdrawal occurs at specific time intervals.

This situation is simulated as follows

Ql =^ (^L Cp (ti - t^^))-j At (66)

j=l

Values of m^ and n are selected such that the amount of energy extracted is

equivalent to that withdrawn from the actual system.

The computer model predicts the auxiliary energy needed to maintain the upper

node of the storage tank within the control temperature range of the thermostat

Daily energy consumption may be expressed as

21
Qhe = ^^^^

For comparison with experimental data, the quantity of energy transferred from

the solar collector array to the storage tank is calculated using the computer

model. The daily thermal output of the irradiated collector array may be

calculated as

L

QSA =^ (^S (tco - tfi))j At (68)

Energy supplied to the storage tank may be expressed as

L

Qt =^ (TMS (tR - tN))j At (69)

j=l

Subtraction of eq (69) from eq (68) yields the amount of energy lost during

transport to and from the collector array. The change in stored energy within

the storage tank during the simulation period is

2_](MiCpti) ^ QST = Z^(MiCpti) - ^> (MiCpti; (MiCpti) (70)

t = LAt T = 0

Once the above quantities have been predicted, the daily performance of the

solar system is compared to that of a conventional system. An energy balance

performed on a conventional system yields

QCON = Ql + ^^^coN + CLcoN + QsT C^^)

For an electric hot water heater, CL(]ON ~ ^» ^1 (71) reduces to

QcON = Ql + HLcoN + QsT (72)

A comparison between the performance of a solar hot water system and that of

a conventional system meeting the same load requirements is possible by

defining a parameter called the "fractional energy savings," i.e.

e = 1 - She (73^

QcoN

The rationale behind this definition is as follows. If the solar system met

the required load with no auxiliary energy required, Qjjg = 0, the fractional

energy savings would be 1.0. If the same amount of energy is required to

operate the solar system as the conventional system, the fractional energy

savings would be zero.

22
2.5.6 Computation Sequence For A Nonirradiated System In Series With A

Conventional Heat Source

The computer model can accommodate a heat source located upstream or downstream

of a nonirradiated solar collector array. The fluid temperature entering the

heat source is predicted using the storage tank and pipe component models when

the heat source is upstream of a nonirradiated array. Fluid temperature

leaving the heat source is predicted using eq (57). This temperature is

subsequently supplied to the collector model in which the incident radiation

has been set to zero. The remainder of the system simulation is identical to

that for an irradiated system. When the heat source is downstream of the

nonirradiated array, the tank and pipe component models predict the entrance

temperature to the nonirradiated array. The predicted collector outlet temperature is used to calculate the heat source outlet temperature. The remainder

of the system simulation is again identical to that described earlier.

2.5.7 Computation Sequence For Nonirradiated Collectors With Attached Strip

Heaters

The computation sequence is identical to that for an irradiated system.

However, the collector component model must be modified slightly. Under

irradiated conditions, the collector model computes the amount of energy

absorbed by the absorber plate using the known incident radiation conditions

and the optical properties of the cover system. The calculation is not

required when modeling collectors with strip heaters attached to the absorber

plates. The absorbed energy is given directly by eq (27).

23

3. EXPERIMENTAL APPARATUS AND INSTRUMENTATION

A test facility has been fabricated at the National Bureau of Standards to aid

in the development of a test method for solar hot water systems. Two solar

domestic hot water systems are located within the test facility. The solar

collector array of one system is subjected to outdoor meteorological conditions,

while the second system uses a thermal simulator. Thermal simulation is

achieved by using an electric heat source only, an electric heat source in

series with nonirradiated collectors, or by attaching strip heaters to nonirradiated absorber plates. Instrumentation and data recording equipment for monitoring the thermal performance of each system and environmental conditions are included within the facility. A water conditioning loop supplies constant

25
temperature water to both systems. The test facility is located at Gaithersburg

,

Maryland, where the latitude is 39.1° and the longitude is 77.2°. Altitude of

the facility is 130 m above sea level.

3.1 SOLAR HOT WATER SYSTEM WITH IRRADIATED COLLECTOR ARRAY

A schematic of the single-tank direct system with irradiated collectors is

given in figure 5. This system consists of two solar collectors connected in

parallel, one water storage tank, an "on-off" differential controller, circulating pumps, piping, and insulation. Two flat-plate, liquid-type collectors

(as shown in figure 6), set at a 39° tilt angle, are used in this system.

Collector dimensions and properties are summarized in Appendix C. A black

chrome coating applied to the steel absorber plate enhances short wavelength

(solar) radiation absorptance and reduces longwave (infrared) radiation losses.

Silicone rubber pads thermally isolate the absorber plate from the galvanized

sheet steel enclosure.

A State Industries conventional electric water heater, as shown in figure 7,

is used for storage. This tank has a nominal 0.310 m^ capacity and outside

dimensions of 1.57 m in height by 0.61 m in diameter. The particular tank used

in this experiment has a capacity of 0.290 m^ and a working pressure of 1034 kPa.

The inner surface of the tank is lined with vitrified glass to minimize corrosion. Glass fiber insulation, thickness 51 mm, surrounds the actual storage

tank, which in turn is covered by a thin metal jacket. A temperature and pressure relief valve rated at 1034 kPA and 99 C is located at the top of the storage tank. The lower heating element was not used in this experiment. The upper

heating element, located 1.27 m above the tank bottom, is a 208 volt, 3.5 kW,

direct immersion heating element. The heating element is controlled by a thermostat which senses the temperature of the storage tank immediately above it.

A solid-state controller, Hawthorne Industry Model H-1504-A "Fixflow," activates

the circulator pumps when an 8.9 C temperature difference exists between the

collector sensor and the storage tank sensor. A temperature difference of less

than 1.7 C deactivates the pumps. The temperature sensors are hermetically

sealed thermistors. The collector thermistor senses the temperature of the

copper outlet pipe at a distance of 38 mm external to the collector enclosure.

The sensor is clamped to the outside surface of the pipe and covered with glass

fiber insulation. The storage tank sensor is located on the exterior tank surface, underneath the insulation, at an elevation of 153 mm above the bottom of

the tank.

Two Grundfos UPS 20-42 pumps connected in series circulate the water through the

solar collector array. These two three-speed pumps in series with flow control

valves provide the capability of varying the mass flow rate through the collector array. Performance characteristics, i.e., head vs flow rate, are given in

figure 8. Isolation valves are installed on each side of the pumps to

facilitate easy removal.

The water supply line to the storage tank consists of 12 . 7-mm-diameter hard

copper tubing. The inlet water supply flows through a gate valve, an integrating volumetric meter, and a temperature-measuring well before entering the

tank. A drain valve, located 92 mm above the bottom of the storage tank, was

26
27
UJ<

enUJ>oo<

UJm3O

_J

O

UJzOo o

4-1 ooH

cr

01

4-1 iH

O-I i-i

P4

QL

Q_

u

•H
AIR VENT^

FROM SOLAR

COLLECTORS

HOT WATER

OUTLET

THERMOCOUPLE

ARRAY

TEMPERATURE a PRESSURE

RELIEF VALVE

HEATING ELEMENT

RETURN TUBE

t TO SOLAR COLLECTORS

COLD WATER

SUPPLY

i TANK DRAIN

Figure 7 . Hot water storage tank

29
0 12 3 4 5

FLOW RATE (m^/h)

Figure 8. Circulator performance characteristic

30

replaced with a tee fitting. The water supply enters through one port of

the tee, while the second port supplies water to the solar collector array.

The supply and return lines to the collector array consist of 12 . 7-iinn-diameter

copper tubing. The supply includes two temperature-measurement wells, one at

the storage tank and one at the collector array, a turbine flowmeter, valves,

and two Grundfos UPS 20-42 pumps. The two solar collectors are connected in

parallel using 25 . 4-mm-diameter copper tubing. A reverse "Z" pattern (figure 9)

is used in order to ensure equal flow through both collectors. The return line

includes a check valve to prevent back flow, a temperature measurement well at

the array, a rotameter, and a temperature well at the storage tank. The solarheated water enters the tank through a 12 . 7-mm-diameter dip tube extending

597 mm from the upper storage tank surface. An air vent is provided at the

top of the storage tank to release trapped air.

The hot water supplied from the storage tank passes by a temperature-measurement

well and through a needle valve before entering an open weigh tank.

Rubatex closed-cell insulation, 12.7 mm thickness, provides internal pipe

insulation. Exterior insulation consists of 31 .8-mm-thick glass fiber insulation covering the 12.7 mm piping, while 51.0 mm glass fiber insulation encases

the 25 .4-mm-diameter collector headers.

3.2 SOLAR HOT WATER SYSTEM WITH THERMALLY SIMULATED COLLECTOR ARRAY

All components of the indoor nonirradiated system are identical to those of the

irradiated system discussed above, with the exception of the solar collector

array. During this experimental investigation, three different techniques were used to supply the quantity of energy normally supplied by the irradiated

array. A description of each is given below.

3.2.1 Electric Heat Source

This system (as shown in figure 10) employs an electric heat source to supply

the net energy gain of a solar collector array. The heat source (figure 11)

consists of three Chromalox Immersion Heaters. Each immersion heater, having a maximum capacity of 1 kW, is encased within a 19. 1-mm-diameter copper tube. A

mixing chamber is located at the entrance and exit of the heat source. The

three rod heaters and associated piping are located within an insulated vacuum

jar. The jar is suspended in polyethylene foam inside a box constructed of

12 . 7-mm-thick plywood. Power input to the electric heaters is controlled by a Powerstat Variable Autotransformer . Electrical specifications are given in

Appendix C. The pressure drop across the simulated collector array has been

matched to that of the actual solar collector array, as shown in figure 12.

In addition, a copper canister and copper tubing were placed in series with

the electric heat source and sized so that the time constant of the simulated

array is approximately equal to that of the actual array (304 s).

3.2.2 Electric Heat Source In Series With Nonirradiated Array

This experimental technique employs nonirradiated solar collectors in series

with the electric heat source described above, as shown in figure 13. A

32
33
34
ROD-TYPE

HEATING ELEMENT

FLOW

MIXING CHAMBER

DEWAR FLASK

FOAM INSULATION

Figure 11. Electric heat source

35
-ac

CO

COUu

COUouooo

01 w

CO •H

13

COU

•H

CJ

01 o <u

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36
tr. UJ

I-UJI

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37
Electric heat source consisting of three immersion heaters

38
39
collector array identical to that used for the irradiated array was constructed

inside the test facility. All piping and insulation is equivalent to that of

the outdoor system. The collector tilt angle is the same (39°). A 6.4-mmthick pljrwood shield mounted parallel and 203 mm above the nonirradiated array

provides a uniform "sky" temperature. The shield extends 203 mm beyond the

outer boundary of the collector array. As discussed in the analysis section,

the heat source may be located upstream or downstream of the nonirradiated

array. During this experimental investigation, the heat source was located

downstream. A storage tank bypass loop allows outdoor collector stagnation

conditions to be duplicated indoors. When an 8.9 C temperature difference

exists between the pipe leaving the heat source and the storage tank sensor, the

controller positions two three-way zone valves such that the flow path is identical to that of an irradiated system. When the temperature difference becomes

less than 1.7 C, the storage tank is bypassed. In this mode of operation, power

supplied by the electric heat source is partially dissipated as heat from the

nonirradiated array.

3.3 NONIRRADIATED COLLECTORS WITH ATTACHED STRIP HEATERS

Energy normally absorbed by the irradiated solar collector is supplied by strip

heaters attached to the back side of the absorber plates. The quantity of

energy supplied to the strip heaters is dependent on the simulated outdoor

radiation conditions and the optical characteristics of the collector.

Strip heaters having a maximum power density of 0.0031 W/cm^ were used in this

series of experiments. Figure 14 shows the strip heater layout on one absorber

plate. Three different sizes were attached to each absorber plate. The combination consists of fifteen 50.8 mm by 1.524 m strip heaters, three 50.8 mm by

609 mm strip heaters, and three 50.8 mm by 228.6 ram heaters. Use of three

different sizes avoids the need for specially sized strip heaters. Brush-on

silicone adhesive was used to bond the heaters to the absorber plates.

The strip heaters were electrically connected in parallel using 14-gage

Tef lon^-coated stranded wire. Power regulation was provided by the autotransformer previously described. The controller sensor locations are identical to

the locations for the irradiated system.

3.4 INLET TEMPERATURE CONDITIONING SYSTEM

The inlet water temperature to both solar hot water systems is held at a preset

temperature by a water-circulating loop, as shown in figure 15. The system consists of two storage tanks, 0.310 m-^ and 0.159 m-^ , a water chiller, a circulating pump, and a temperature controller. Supply water enters the 0.310 m-^ tank.

This tank has a 4.5 kW heating element located near the bottom. Water leaves

the top of the tank and enters the 0.159 m-^ tank which provides additional storage capacity. A constant displacement pump draws water from the bottom of the

smaller tank. The water circulates through a closed loop, which contains a Temprite Model R-30 Water Chiller, before returning to the bottom of the large

storage tank. A Honeywell Model R7350 Digital Set Point Temperature Controller,

with a Type T thermocouple sensor located in the 0.310 m^ storage tank, senses

the temperature deviation from the set point. The controller then activates the

40
CENTER

UJo

E

E

m

CO

(T

Ld

UJX

a.

a:

• • 229mm-»

-6IC mm-

•813mm-

.524 m

Figure 14. Location of strip heaters on the back of solar collector

absorber plate

41

(D

CO3

to(D>

'-^fa

1^

to<u

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:ji o

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chiller or heating element as necessary. The fluid temperature in the loop is

generally maintained within + 2 C of the set point.

3.5 INSTRUMENTATION

Volumetric flow, temperature, power, and environmental conditions are monitored

during the operation of each solar system. These measurements allow determination of system thermal performance. The irradiated and nonirradiated solar hot

water systems are identically instrumented, with two exceptions. Pertinent outdoor meteorological conditions are measured for the irradiated system, while the

input power to the electric heat source or strip heaters is measured for the

nonirradiated system.

Premium grade thermocouple wire which has a nominal accuracy of + 0.5 C was used

in all thermocouple and thermopile construction. All wire runs are continuous

from the temperature sensors to the data acquisition system. All thermopiles

were calibrated using constant temperature baths and NBS-certif led thermometers.

A vertical array of type T (copper-constantan) thermocouples, with the junctions

at 152 mm intervals, is located inside each water storage tank. Construction

details are shown in figure 16. Thermocouples are used to measure the inlet and

outlet potable water temperatures and a three- junction thermopile measures the

temperature difference when a hot water load occurs. Two shielded thermocouples,

supported by a wooden rod, located 1.5 m from the tank, are used to measure the

surrounding ambient temperature. The thermocouples have elevations of 0.53 m

and 1.05 m, respectively, above the floor. Thermocouples and a six-junction

thermopile are used to monitor the irradiated or simulated solar collector array

supply and return temperatures at the storage tank. Thermocouples are located

in the inlet and outlet of the collector array. A six- junction thermopile is

used to measure the temperature difference across the array.

The apparatus used to supply the quantity of energy supplied by an irradiated

array may consist of an electric heat source, strip heaters attached to non- irradiated collector absorber plates, or nonirradiated collectors in series with

an electric heat source. Thermocouples are placed at the inlet and outlet of

each component while a six-junction thermopile measures the temperature difference across each component. An additional thermopile is placed across the non- irradiated array-heat source combination. Two thermocouples measure the skyshield temperature above the nonirradiated array. The absorber plate temperatures are monitored in four locations along the center line of the plate as shown in figure 17.

Flow rate through the collector array is measured with a Flow Technology, Model

FTM-NIO-LB, Omniflo Turbine Flowmeter. This volumetric flowmeter uses a bladed

rotor to generate a digital signal. An orifice directs the flow tangentially

past the underside of a rotor, figure 18. The turbine rotor rotates at a speed

proportional to the velocity of the fluid. A variable reluctance transducer

located adjacent to the rotor has generated within it an electrical current

whose frequency is proportional to the rotational speed of the rotor. The

pulses generated are fed to a Flow Technology Series 7050 digital processor.

The turbine flowmeter and digital processor are calibrated, as a unit, before

each experiment using the dead weight method. A typical resulting calibration

45
TYPE T

( COPPER - CONSTANTAN)

THERMOCOUPLE WIRE^

I/4-20NC

THREAD

EPOXY FILL —

-9.5 mm HARD- COPPER

TUBING

-EPOXY FILLED

1/8-NPT THREAD

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1/4-20 NC

15.9 mm TEFLON "TF" ROD

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SILICONE

ADHESIVE.

9.5 mm POLYCARBONATE

ROD

TYPE T /

THERMOCOUPLE'

f

1.677m.

TANK

BOTTOM

7152 mm

4

152mm

Figure 16. Vertical thermocouple array construction details

46
TYPE T

THERMCOUPLES

Figure 17. Location of absorber plate thermocouples

47
48

curve is shown in figure 19. In addition to the turbine flowmeter, a Brooks

Model 1110 Full-View rotameter was used to monitor the flow rate.

The energy supplied to the heating element in the storage tank is measured

using a General Electric 240 volt Type IW-70-S Watthour Meter equipped with a D20-1 contact closure device. Each contact closure is equivalent to 20 watthours.

An electrical counter sums the number of contact closures. Connected in series

with the watthour meter is a General Electric Type IB-10 Portable Watthour

Meter Standard. Use of the watthour standard provides 0.12 watthour resolution.

The use of both instruments provides two independent measurements. Instantaneous power consumption is measured with a Scientific Columbus XL5C5-A2-1-7 Watt

Transducer. This instrument provides a 0 to 100-mV-output signal directly

proportional to a 0 to 3500 watt power input. The output signal is fed to the

data acquisition system. Specifications are given in Appendix C. Additionally,

a 1.5 Vdc signal is fed to the data acquisition system whenever the heating

element is energized.

Operation of the circulator pumps is monitored by a Cramer Elapsed-Time Meter.

Resolution of 0.1 minutes is provided over a range of 0 to 9999.9 minutes.

When the circulators are operating, a signal is sensed by the data acquisition

system. This arrangement allows the operational status of the circulators to

be continuously monitored.

A specified amount of energy is withdrawn from the solar storage tank during

each test. An integrating volumetric water meter and weigh tank are used to measure the amount of water withdrawn from the storage tank. The water meter

is a Model 15 Badger Recordall. Water withdrawn is collected in an insulated

0.114 m3 tank which rests on a Fairbanks, Morse, and Co. balance scale. A

thermometer suspended in the tank measures the mixed temperature of the withdrawn water. When a draw occurs, the three- junction thermopile signal is fed

to an analog integrator. The integrator, manufactured by AGM Electronics,

integrates the analog input signal with expired time and provides an analog

output signal proportional to the accumulated integrated quantity. The input

and output signals to the integrator are fed to the data acquisition system

and to a Westronics MT-22 Strip Chart Recorder.

The instantaneous power supplied to the simulated solar collector array is

measured during operation of the nonirradiated solar hot water system. This

measurement is accomplished by using a Scientific Columbus watt transducer

identical to the one described earlier. In addition, a Scientific Columbus

Model 6335FF Digital Wattmeter is used. This wattmeter provides a visual

digital readout of instantaneous power supplied to the simulated array.

Specifications are given in Appendix C.

Outdoor meteorological conditions are measured during operation of the

irradiated system. The net total global radiation is measured with an Eppley

Model PSP Radiometer. The radiometer specifications are given in Appendix C. The longwave radiation, both indoors and outdoors, is measured using an Eppley

Precision Infrared Radiometer (Pyrgeometer) whose specifications are given in

Appendix C. Both instruments are located on the collector array support stand

at a tilt angle of 39°. A Weather Measure W103/B three-cup anemometer is used

50
o

CO
31

52
53
54
to measure the wind speed. Calibration was performed by the NBS Fluid

Engineering Division. Outdoor ambient temperature is measured using a Type T

thermocouple housed within a vented weather shelter.

Analog signals from various sensors are connected to a Leeds & Northrup

Trendscan 1000 Data Acquisition System. This unit scans all instruments in

one-minute intervals and is capable of accommodating 100 inputs. Appendix C shows the data recorded and channel assignments. The Trendscan 1000 has three

ranges, which include a + 400 mV range, a + 10 V range and a -200 C to +400 C

Type T thermocouple range. Any of the 100 input points may be addressed and

assigned to any one of the three available ranges or omitted. The unit has an electronic discharge printer, and an internal clock, provides real time display

and initiates periodic logs at specified time intervals. Reference junction

thermocouple compensation, using a bridge network, is located on the range input cards. Specifications are given in Appendix C for this data acquisition

system.

A Kennedy Model 1642 Interface Unit interfaces the data acquisition system with

a Kennedy Model 9832 Magnetic Tape Recorder. After recording the data, the

magnetic tapes are read on the NBS UNIVAC 1108 Computer.

55

4. EXPERIMENTAL PROCEDURE, RESULTS, AND DISCUSSION

A number of experiments have been performed at the NBS to assess proposed test

methods for solar domestic hot water systems. Three different experimental

techniques of supplying the quantity of energy normally supplied by an outdoor

irradiated array have been investigated. Tests were conducted to determine if

the control functions of a system located entirely within a test facility are similar to those of a system subjected to outdoor meteorological conditions.

An experimental study was performed to determine if one proposed test method

(ASHRAE Standard 95) is sensitive to the amount of thermal stratification in

the solar storage tank. Repeatability of test results was also investigated.

A discussion of each experimental investigation follows.

57
4.1 REPLACEMENT OF THE IRRADIATED ARRAY WITH AN ELECTRIC HEAT SOURCE

An electric heat source is employed to supply the net energy gain of a solar

collector array. The outdoor irradiated system was allowed to operate in a normal manner for several days in order to serve as a basis of comparison to

the system performance using the electric heat source. During operation of the

irradiated system, both clear and intermittently cloudy sky conditions existed.

A description of the test sequence for the irradiated system follows.

At the beginning of each test day, the storage tank was filled with water at

20 C. The environment surrounding the tank was held at 20 C throughout the

test. At 09:15 the heating element in the storage tank and differential

controller were energized. Two hot water draws, one at 13:15 and one at 17:15,

occurred during the test days. Each draw consisted of water entering the

storage tank at 20 C at a flow rate of 2 x 10~^ m^/s. During withdrawal

periods, the incoming water temperature and temperature of the water withdrawn

were measured. At the conclusion of each test day, water was withdrawn from

the storage tank until the outlet temperature was within 1 C of the inlet

temperature. The quantity of energy withdrawn was computed. Power consumed by

the storage tank heating element was measured during each test day.

The testing procedure described for the irradiated system was followed by tests

of the system which employed an electric heat source. Identical measurements

were made. Using the incident irradiance and ambient temperature which were

recorded during operation of the irradiated system, the power supplied to the

heat source was determined using eq (2) and the flow rate correction procedure

presented in Appendix A. The power supplied to the electric heat source was updated every five minutes.

Use of the identical test procedure for both systems allowed a direct comparison

to be made between the performance of a solar domestic hot water system using

irradiated collectors and one which employed an electric heat source. Figures

20 through 23 show the variation in solar irradiance for the four August test

days. The test period included an almost clear day (August 10, figure 22) as well as intermittently cloudy days. The figures also show a comparison between

the collector output for the irradiated array and the electric heat source. At

the start of each test day for the irradiated system, the thermal output of the

collector array is initially large and quickly decreases. This behavior is

because the irradiated collectors experience stagnation conditions until the

circulators are energized. The system using the electric heat source shows a rapidly increasing thermal output at the beginning of each test period. Outdoor stagnation conditions cannot be simulated accurately using an electric

heat source only. Thus, the power output of the heat source is initially zero.

Large thermal output transients also occur during the hot water draw. During a hot water draw, cold makeup water enters the simulated and irradiated collector

arrays. During the short time period between the time the cold water enters

the arrays, but before the outlet temperature changes, the thermal output of

the arrays suddenly increase. After the draw has terminated, water supplied

from the storage tank enters the arrays causing the inlet temperature to

increase sharply, resulting in a sudden decrease of thermal output.

58
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62
Table 1 shows the results of running the solar domestic hot water system with

the irradiated array on four different days and repeating the tests for each

day using the system with the electric heat source. It was possible to reproduce the outdoor daily collector output in the laboratory to within four percent even on the intermittently cloudy day of August 8 (figure 21). The daily

fractional energy savings for the system with the electric heat source was within two percentage points for all the days tested. The fractional energy savings were calculated based on comparing the solar system with a 0.2 m^ conventional

electric water heater having thermal losses of 11,160 kJ per day.

The use of an electric heat source to simulate the thermal output of an irradiated array results in excellent agreement between outdoor and laboratory

system performance. Additionally, a SDHW system may be tested with various

numbers and types of collectors. These tests can be accomplished using corre- sponding values of Fj^(Ta)^, Fj^Ul, and Ag without hardware changes. Outdoor

stagnation conditions, however, cannot be accurately simulated using this

technique. Thus, outdoor control functions cannot be duplicated indoors using

an electric heat source only. A second disadvantage is the complexity of the

calculation sequence. The calculation procedure requires that the real time

measured quantities t^^^ and A be used in addition to the simulated meteorological conditions.

4.2 REPLACEMENT OF THE IRRADIATED ARRAY WITH AN ELECTRIC HEAT SOURCE DOWNSTREAM

OF A NONIRRADIATED ARRAY

An electric heat source downstream of a nonirradiated array is used to supply

the net thermal output of an irradiated array. A fluid loop to bypass the

storage tank allows stagnation conditions to be simulated within the laboratory.

The irradiated collector array system was allowed to operate in a normal manner

for two days in October 1980 to provide a comparison basis for the system using

an electric heat source downstream of a nonirradiated array. During the outdoor

tests performed in August 1979, the pumps ran continuously during each test day

except for a brief period on August 10. This continuous operation is attributed

to the 20 C initial storage tank temperature. Initial storage tank temperatures

of 35 C and 55 C were selected for October 6 and October 8, respectively, so that the controller would operate in a more typical manner. The test sequence

for the irradiated system is identical to that followed in August 1979 until

17:15. Instead of deactivating the heating element and controller at 17:15,

the system was allowed to operate until the controller cut the pumps off. At

this time, water was withdrawn from the storage tank until the outlet

temperature was within 1 C of the inlet temperature.

Tests were conducted in November 1980 using an electric heat source downstream

of a nonirradiated array. The recorded incident irradiance during the comparison outdoor tests in conjunction with the mass flow rate correction procedure

(Appendix A) allowed the quantity of power supplied to the electric heat source

to be computed using eq (21). Input power was adjusted every three minutes

throughout the test day. The ambient temperature surrounding the indoor non- irradiated collectors was controlled such that it was equivalent to the outdoor

temperature during the irradiated system tests.

63
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64
Table 2 compares the results obtained by running the irradiated system for the

two selected days during October 1980 and repeating the tests using the system

employing an electric heat source downstream of a nonirradiated array. The

fractional energy savings was identical for the October 6 - November 17 comparison and within 3 percentage points for the October 8 - November 25 comparison.

For both comparisons, the daily energy supplied by the electric heat source- nonirradiated array combination was higher than that delivered by the irradiated

outdoor array, 2.8 percent for the October 6 - November 17 comparison and 3.4

percent for the October 8 - November 25 comparison. The combination of a higher

effective sky temperature and a zero wind speed for the indoor system would

tend to account for this difference.

Figures 24 and 25 show the incident tilt radiation for October 6 and 8,

respectively. A comparison of the thermal output of the irradiated array with

that of the nonirradiated array with downstream heat source is also shown. The

large transients which occurred at the beginning of each test day resulted from

the pumps cycling on and off. The pumps were initially energized when the systems were under stagnation conditions. When the incident irradiance on the outdoor collector system was not sufficient to maintain a 1.8 C temperature difference between the collector controller sensor and the storage tank sensor, the

pumps were deactivated by the controller. Similarly, when the power input to

the electric heat source was insufficient to maintain a 1.8 C temperature difference between the fluid leaving the heat source and the storage tank sensor,

the controller deactivated the loop-bypass valves which stopped flow through the

storage tank.

During the hot water draws, large transients in the thermal output occurred for

both the irradiated collector system and the system using an electric heat source downstream of a nonirradiated array. The explanation for this behavior

is identical to that given in the previous section. Figure 24 also shows an additional pump shutdown for both systems, shortly after the hot water draw, for

the experiments with initial tank temperatures of 55 C. An explanation for this

behavior for the irradiated collector array system follows. The reasoning is

also valid for the nonirradiated collector array-heat source combination system.

The pump restarted after the hot water draw and circulated the cold water in

the bottom of the storage tank to the irradiated array. The water returning

to the storage tank was at a lower temperature than the water in the upper portion of the tank. Mixing occurred within the storage tank. As the pump continued to operate, warm water reached the controller sensor location. The

controller tank sensor temperature increased sufficiently to deactivate the

pump. Under stagnation conditions, the temperature of the collector sensor increased and reactivated the pump. Due to the lower initial storage tank

temperature during the October 8 - November 25 tests, this sequence of events

did not occur after the hot water draw.

Figures 26 and 27 show the controller operation for October 6 and 8,

respectively. There is good controller operation agreement between the irradiated system and the nonirradiated array with downstream heat source system.

The total elapsed pump operation time for the irradiated system for October 6 was 424.6 minutes as compared to 427.5 minutes for the nonirradiated array-heat

source system. On October 8, the circulators on the irradiated system operated

65
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for 479.4 minutes compared to 489.7 minutes for the nonirradiated array-heat

source system. In addition to the good agreement between total circulator

operation time, there is excellent agreement among the times when the circulator operated and the times at which stagnation conditions existed.

Using a nonirradiated collector array with a downstream heat source to supply

the thermal output of an irradiated array results in excellent agreement

between outdoor and laboratory system performance. Outdoor stagnation conditions can be simulated and thus outdoor control functions can be duplicated.

With the heat source located downstream of a nonirradiated array, the calculation procedure is simpler than the calculation sequence for the electric heat source only. The ASHRAE 95 Committee extended this method to allow either an upstream or a downstream heat source, and incorporated it into the ASHRAE 95

Test Procedure. No experiments were performed at NBS using an upstream heat

source.

4.3 REPLACEMENT OF AN IRRADIATED ARRAY USING STRIP HEATERS ATTACHED TO

NONIRRADIATED SOLAR COLLECTORS

Solar energy absorbed by the irradiated absorber plates is replaced by energy

supplied by electric strip heaters attached to the backside of nonirradiated

absorber plates.

The irradiated collector array system operated on October 6 and 8, 1980 provided

a basis for comparison with the system using electric strip heaters. These two

test days have been described in the previous section.

Tests were conducted in April 1981 using electric strip heaters attached to the

back of nonirradiated absorber plates. The recorded incident irradiance during

the comparison outdoor tdsts allowed the quantity of power supplied to the strip

heaters to be computed using eq (27). Power supplied to the strip heaters was updated every three minutes. The ambient temperature surrounding the indoor

nonirradiated collectors was controlled such that it was equivalent to the outdoor temperature during the irradiated system tests. The test procedure was identical to that followed during the irradiated system tests in October 1980.

Table 3 compares the results of operating the solar domestic hot water system

with an irradiated array and repeating the tests for each day using the system

with electric strip heaters attached to the nonirradiated absorber plates.

The daily fractional energy savings for the system using strip heaters was 2.4

percent higher than the irradiated array system for the October 6 - April 1 comparison, and 1.9 percent higher for the October 8 - April 2 comparison. For

both comparisons, the daily quantity of energy supplied by the system using

the nonirradiated array with attached strip heaters was within 0.9 percent of

that supplied by the irradiated array. Figures 28 and 29 show the incident

tilt radiation for October 6 and 8, respectively. A comparison between the

thermal output of the irradiated array and that of the nonirradiated array

with strip heaters is also shown. Examination of figures 28 and 29 reveals

that the thermal output of the irradiated array is greater than the thermal

output of the nonirradiated array with attached strip heaters before solar

noon. After solar noon, the thermal output of the nonirradiated array with

71
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74
attached strip heaters is greater than the thermal output of the irradiated

array. This behavior may be attributed to the increased effective thermal

capacity of the collector array which results from attaching strip heaters to

the absorber plates. The thermal capacity of an individual collector to which

strip heaters have been added is approximately 50 percent greater than an unmodified collector.

Figure 30 shows the circulator pump status for October 6, 1980 and April 1,

1981. There is good agreement between the two test days among the times at

which the circulators operated and the times at which stagnation conditions

existed. The system using the electric strip heaters experienced one additional "off" period as compared to the irradiated system. The total elapsed

pump operation time for the irradiated system was 424.6 minutes as compared to

436.1 minutes for the system using nonirradiated collectors with strip heaters

attached to the absorber plates. The circulator pump status for October 8,

1980 and for April 2, 1981 is shown in figure 31. The circulators for the

irradiated system operated for a total of 479.4 minutes as compared to 501.8

minutes for the system using nonirradiated collectors with attached strip

heaters. Agreement between the times at which stagnation conditions existed

and times at which the circulators operated is excellent.

Thermal performance agreement between the system using an irradiated array and

the system which uses electric strip heaters attached to nonirradiated collector absorber plates is excellent. Outdoor stagnation conditions can be duplicated without the use of a storage tank bypass loop. The calculation procedure

required to determine the amount of power supplied to the strip heaters is

independent of mass flow rate and fluid type. Testing of an individual collector module, according to the ASHRAE 93-77 Test Procedure, is not necessary if

the optical properties of the collector are known. Thermosyphon systems may be

tested using this technique. This method, however, requires that the collectors

be partially disassembled. Also, the absorber must be reasonably flat and sized

to be compatible with commercially available strip heaters.

4.4 TEST PROCEDURE SENSITIVITY TO STORAGE TANK STRATIFICATION

Experiments have been conducted to determine if the ASHRAE 95 Test Procedure is

sensitive to temperature stratification in the storage tank. Storage tank temperature stratification is influenced by the design and dimensions of the storage

tank, velocity at which fluids enter the tank, and the manner in which fluids

are introduced into the storage tank. During this experimental investigation,

several designs of return tubes, which introduce the solar heated water into

the storage tank, were fabricated and tested. The storage tank was not altered

and the mass flow rate through all return tubes was maintained at 0.055 kg/s.

An electric heat source downstream of a nonirradiated array was used in the

studies to determine the influence of each return tube on storage tank stratification. The recorded outdoor irradiance and associated incident angle for

October 6, 1980 were used to calculate the quantity of power supplied to the

electric heat source. The input power was updated every five minutes. Ambient

temperature surrounding the nonirradiated array was maintained at 20 C. Each

test began with an initial 20 C storage tank temperature. The circulator pumps

75
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77
and storage tank heating element were energized at 09:00. Power was supplied

to the electric heat source from 09:00 to 11:30.

Construction details for each return tube and the resulting storage tank

temperature distributions are shown in figures 32 through 45. Details of the

return tube which was used in the previously described experiments are given in

figure 32. Figure 33 shows the resulting storage tank temperature distribution.

The upper thermocouple shows the influence of the heating element. The temperature of this thermocouple increases rapidly until the thermostat deactivates the

heating element. The upper thermocouple temperature subsequently decreases as

a result of heat loss from the storage tank and mixing within the tank. The

second thermocouple from the top, located 25.4 mm below the heating element,

is also influenced by the heating element. The bottom portion of the storage

tank is well mixed as shown by the temperatures of the bottom seven thermocouple

Figure 34 shows a return tube which redirects the downward axial fluid flow into

a number of radially directed fluid streams. Two concentric tubes were used in

an effort to eliminate any downward momentum. The resulting temperature distribution, figure 35, shows improved stratification as compared with figure 33.

Figure 36 shows a commercially available return tube which also redirects the

downward axial fluid flow. Only one tube, with holes drilled in the radial

direction, is used. The resulting temperature distribution is displayed in

figure 37. Stratification in the lower portion of the tank is almost identical

to that displayed in figure 35. The temperature decay in the upper portion of

the tank is slightly greater than that shown in figure 35.

Three return tubes, as shown in figures 38, 39, and 40, were constructed such

that fluid enters the storage tank in a single radially directed fluid stream.

The corresponding temperature profiles, as shown in figures 41, 42, and 43, are essentially identical. The bottom seven thermocouples display a temperature

profile equivalent to the profile shown in figures 35 and 37. The temperature

decay rate for the two uppermost thermocouples is greater than the decay rate

displayed in figures 35 and 37 but less than the decay rate exhibited in figure

33.

A return tube consisting of a number of radially directed openings at a fixed

elevation is shown in figure 44. The temperature decay rate for the two uppermost thermocouples, shown in figure 45, is less than the rate displayed in

figures 33, 41, 42, and 43 but greater than that exhibited in figures 35 and 37.

The return tube shown in figure 34 was selected as the best among those

considered for promoting stratification within the storage tank. Figure 32

shows the return tube which provided virtually no stratification in the storage

tank and a rapid temperature decay rate for the two uppermost thermocouples.

All-day tests were performed using each return tube to determine if the ASHRAE

95 Test Procedure is sensitive to temperature stratification in the storage tank

The recorded outdoor irradiance and associated incident angle of October 6, 1980 were used to calculate the quantity of power supplied to the electric heat source. Input power was updated every three minutes. Ambient temperature sur- rounding the nonirradiated collectors was maintained at 20 C. Storage tank initial temperature was 35 C. The controller and heating element were energized

78
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597 mm

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92
were used to calculate the quantity of power supplied to the electric heat

source. Input power was updated every three minutes. Ambient temperature

surrounding the nonirradiated collectors was maintained at 20 C. Storage tank

initial temperature was 35 C. The controller and heating element were ener- gized at 09:00. One hot water draw occurred during each test day at 13:00.

The draw consisted of water entering the storage tank at 20 C at a flow rate of

2 X 10"^ m^/s. During withdrawal periods, the incoming water temperature and

water temperature leaving the tank were measured. Near the end of the test day,

the system was allowed to operate until the controller deactivated the circulator pumps. At the conclusion of each test day, water was withdrawn from the

storage tank until the outlet temperature was within 1 C of the inlet temperature. The quantity of energy withdrawn was then computed. The power used by

the storage tank heating element was measured during each test.

Figure 46 displays the tank temperature distribution for February 3, 1981,

during which, the return tube which best promoted stratification was used.

The tank temperature distribution for January 29, 1981, as shown in figure 47,

resulted when the return tube shown in figure 32 was used. Table 4 gives the

results of running the hot water system for January 29, 1981 and February 3,

1981. The fractional energy savings which results from a stratified tank is

8.7 percent higher than the fractional energy savings which occurs using a tank which exhibits minimal stratification. The slower temperature decay rate

in the upper portion of the tank results in lower auxiliary energy usage for

the system where the tank is stratified. The stratified storage tank temperature profile also results in a lower inlet fluid temperature to the collector

array as compared to the mixed storage tank temperature profile. The lower

inlet fluid temperature allows the collector array to operate at a higher

efficiency and thus a greater quantity of energy is delivered to the storage

tank.

Based on these results, it is concluded that the ASHRAE 95 Test Procedure is

sensitive to the degree of temperature stratification which occurs in the

storage tank.

4.5 REPEATABILITY OF TEST RESULTS

The test conducted on January 29, 1981 was repeated on February 4, 1981 to

determine the repeatability of the ASHRAE 95 Test Procedure. The same return

tube, as shown in figure 32, was used in both tests. Table 5 gives the test

results for January 29, 1981 and February 4, 1981. The daily fractional energy

savings was within 0.1 percent for the two test days.

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97

Thermal Test Methods for Solar Hot Water Systems

Sponsor Office of Solar Applications lor Buildings, DoE

Objective: To develop standard test procedures lor solar domestic hot water systems.

Results: • ASHRAE Standard 95 adopted In eady 1981

• Validation experiments completed using a system

tested outdoors (Figure 1) and one tested Indoors

(Figure 2). Results are shown below — Using a conventional heat source, Figures 3 and 4. — Using nonlrradlated collectors In series with a

conventional heat source, Figure 5. — Using strip fieaters attached to nonlrradlated

solar collectors.

• Controller operation documented to be nearly

Identical to that achieved during an outdoor test. Figures 6 and 7.

5. ANALYTICAL RESULTS AND DISCUSSION

Effects on solar hot water system performance resulting from the use of various

test methods and different environmental conditions are considered. Results

are presented for a system having normally irradiated solar collectors,

nonlrradlated collectors in series with an electric heat source, and nonirradiated collectors with strip heaters attached to the absorber plates.

5.1 SDHW SYSTEM ANALYTICAL MODEL VALIDATION

The thermal performance of a solar hot water system is predicted using the

analytical model. The predicted performance is compared to the measured

performance of the single-tank direct system for four selected test days.

99
The first step in validation of the total system model is to verify the routine

used to predict the thermal performance of the solar collector modules. Figure

48 shows the solar collector efficiency curve which was predicted using the

analytical model. The absorptance of the absorber plate and the edge heat

loss coefficient were selected such that the linearized calculated efficiency

curve for the solar collector was identical to the measured efficiency curve. Both of these parameters, however, are well within the range of those expected

from handbook tabulations of property data. Following reference 7, the linearization is based on least-squares fits to the points shown which are at approximately 10, 30, 50 and 70 percent of the absorber stagnation temperature rise

above ambient. The linearized efficiency curve slightly overestimates the

ordinate intercept. Also, the negative slope of the linear curve is somewhat

greater than that of the corresponding nonlinear curve over the range of practical importance for domestic water heating.

Experimental data collected during operation of the single-tank direct SDHW

system during August 7 and 13, 1979 and October 6 and 8, 1980 provide a comparison basis for the predicted results of the analytical model. The initial tank

temperature for the analytical model is set equal to the measured value for

each test day. The meteorological data required for the analytical model are

the values recorded during operation of the SDHW system for each test day. As

discussed previously, the analytical model can simulate a storage tank which

is composed of N volume segments or "nodes." Selection of a single- temperature

tank corresponds to a "well mixed" condition and neglects any thermal stratification occurring within the storage tank. Increasing the number of nodes in

the storage tank model corresponds to increasing the degree of temperature

stratification within the tank. In order to determine the number of temperature nodes which would provide the closest agreement between analytical and

experimental results, three computer simulations were made for each test day.

These three computer simulations are based on

(a) taking the tank to be completely mixed (one nodal temperature),

(b) considering partial stratification with two tank volume elements (two

nodal temperatures) , and

(c) considering more complete thermal stratification with five tank

volume elements (five nodal temperatures).

The upper volume element for the two- and five-node storage tanks corresponds

to the volume of water between the storage tank heating element and the upper

surface of the tank. The four bottom elements of the five- node tank contain

equal volumes of water.

Tables 6 through 9 show the results of the analytical SDHW system model using

a one-node, two- nod e , and five-node storage tank for August 7, August 13,

October 6, and October 8, respectively. The fractional energy savings predicted

by the analytical model using a one-node storage tank is significantly less than

the fractional energy savings predicted when a two- or five-node storage tank

is used. This large difference in predicted performance can be explained as follows. The storage tank heating element maintains the upper node (in which

100
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105
it is located) within a specified temperature range. A single-node storage

tank requires that the heating element supply a quantity of energy sufficient

to maintain the entire storage tank within this temperature range. Thus, the

electrical energy used by the heating element is greatly increased over the

amount required for only the upper node of a two- or five-segment tank. The

predicted system performance using a five-node tank is slightly greater than

the predicted performance when a two-node tank is used. The fluid which enters

the solar collector array is supplied from the bottom volume element. As the

number of storage tank nodes increases, the average fluid temperature entering

the collector array decreases. The lower fluid temperature results in a higher

collector efficiency, so the higher system performance is expected.

Analytical model results for both the two-node and five-node storage tank are

in close agreement with the experimental results for all four test days. However, whereas the analytical model predicted a small increase in system performance from increased storage tank stratification (five-node as compared to

two-node) , the experimental results indicated a significant increase in fractional energy savings (see Table 4). This discrepancy is due to the internal

mixing within the storage tank which the analytical model fails to take into

account

Figures 49 and 50 show a comparison of the experimental and the analytically

predicted storage tank temperature distributions for October 8, 1980. Figure 49

shows the storage tank distribution for a two-node storage tank, while figure 50

shows the distribution for a five-node storage tank. The other three test days

exhibited similar temperature profile comparisons. In the experimental tank,

the top thermocouple displays a greater temperature decay than predicted by the

analytical model, while the lower seven thermocouples show a greater increase

in temperature than predicted. The second thermocouple from the top of the

experimental storage tank is influenced, although to a lesser degree than the

upper thermocouple, by the operation of the heating element. This thermocouple,

located 25.4 mm below the heating element, increases rapidly in temperature

until the heating element is de-energized by the thermostat. Then, the temperature of this thermocouple decreases rapidly because of internal mixing

Internal mixing also affects the bottom seven thermocouples, which show no stratification in the portion of the tank they monitor. The numerical results

show that the five-node tank model slightly overestimates the effects of thermal

stratification, while the two-node tank gives results which agree more consistently with experimental results. Due to this lack of stratification in the

experimental storage tank, a two-node storage tank model was used in all subsequent simulations. Thus, the upper node represents the region which contains

the heating element, and the lower node represents the remaining "well mixed"

lower portion of the tank.

5.2 TEST METHOD INFLUENCE ON SYSTEM PERFORMANCE

The thermal performance of a SDHW system may be influenced by the technique

used to supply the net themal output of an irradiated array. In order to

quantitatively determine the effect of various test methods, the SDHW system

model was used to predict the performance of a system using an irradiated collector array, a nonirradiated array with a downstream heat source, a non- irradiated array with an upstream heat source, and strip heaters attached to

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108
nonlrradlated absorber plates. The recorded environmental and initial

conditions for October 6 and 8, 1980 provided the required input parameters.

The use of the same ambient temperature, effective sky temperature, and wind

velocity for the simulation of each test method allowed the effect of various

test methods (as opposed to different environmental conditions) on system

performance to be determined.

Tables 10 and 11 summarize the results of the test method comparison. The

values of Fj^( Ta)^ and Fj^Ul used in the expressions for calculating the power

input to the heat source are based on the linearized efficiency curve 6hown in

figure 48. Use of the various test methods resulted in almost identical fractional energy savings. For the October 6, 1980 comparisons, the daily thermal

output of the nonirradiated array with downstream heat source and the nonirradiated array with strip heaters attached to the absorber plates are 1.3 percent

less than the daily thermal output of the irradiated array. Using the environmental conditions for October 8, 1980, the dally thermal output of the nonirradiated array with downstream heat source and the nonirradiated array with strip

heaters attached to the absorber plates are 1.4 percent and 1.1 percent less,

respectively, than that for the Irradiated array. The lower net thermal output

is attributed to the lack of solar absorption and the resulting heating in the cover which has also been analyzed by Gillett [13]. The lower cover temperature under nonirradiated conditions results in a higher value of FrUl as com- pared to irradiated conditions. The higher value of F^U^ results in greater

thermal losses for the nonirradiated collectors and thus a decrease in the net

thermal output. The daily thermal output of the nonirradiated array with

upstream heat source is 0.1 percent and 0.3 percent less than that of the

irradiated array for October 6 and October 8, 1980, respectively. This close

agreement is attributed to two phenomena. The nonirradiated conditions result

in a higher value of F^Ul than for irradiated conditions, thereby decreasing

the net thermal output of the nonirradiated array with an upstream heat source.

The other contributing factor can be explained by considering the denominator

in eq (20) . This term accounts for the incremental heat loss resulting from

the higher inlet temperature to the downstream nonirradiated collector array.

Since Fj^Ul is based on the linearized measured efficiency curve and is larger

than the Fj^U^ value for both the nonirradiated collector and the irradiated

collector, at the inlet temperature conditions encountered during October 6

and 8, 1980, the denominator term in eq (20) allows for a greater heat loss

than is actually dissipated from the downstream nonirradiated collector array.

For example, for an irradiance of 900 W/m^, a 25 C ambient temperature, an effective sky temperature of 11 C, a 2.0 m/s wind velocity and an inlet temperature of 40 C, the Fj^U-j^ value is 4.19 W/(m C). For an irradiance of zero, all

other conditions identical, the Fj^Uj^ value increases to 4.31 W/(m C) due to the

decreased co.ver temperature, thus accounting for the higher heat loss from the

nonirradiated collector. However, the 4.31 W/(m^C) value of Fj^Ux is still con- siderably less than the value of 4.73 W/(m2c) obtained from the linear efficiency curve. This condition results in a greater power input to the upstream

heat source than is actually required. Thus, the greater thermal input to the

heat source tends to compensate for the greater thermal loss from the

nonirradiated array.

109
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111
5.3 INFLUENCE OF ENVIRONMENTAL CONDITIONS ON SYSTEM PERFORMANCE

Environmental conditions surrounding the nonirradiated array located in the

laboratory influence the performance of the SDHW system. The analytical model

is used to predict the performance of a SDHW system subjected to zero wind

speed and an effective sky temperature equivalent to ambient temperature using

each indoor test method. The results are compared to an outdoor system

experiencing the environmental conditions recorded during October 6 and 8,

1980. The indoor ambient temperature was set equivalent to the recorded outdoor

ambient temperature. By this approach, the combined effects of testing the

system indoors and the use of various test methods can be simulated.

Figure 51 displays the recorded wind speed for October 6, 1980, and the ambient

temperature and effective sky temperature recorded for October 6, 1980 are shown in figure 52. Table 12 gives the results for each test method for the

October 6, 1980 comparison. As expected, the fractional energy savings

increased for all test methods as compared to the results given in Table 10.

This result is due to a lower heat loss from the nonirradiated collector array

because of a zero wind speed and higher effective sky temperature. The

increase in fractional energy savings is less than 0.5 percent for all methods.

Under indoor environmental conditions, the daily thermal output of a nonirradiated array with downstream heat source best matches the daily thermal output

of an irradiated array subjected to outdoor environmental conditions. The

daily thermal output of the nonirradiated array with an upstream heat source exceeds the outdoor irradiated array by 1.8 percent. The daily thermal output

of the nonirradiated array with a downstream heat source and the nonirradiated

array with attached strip heaters exceeds the daily thermal output of the

irradiated array by 1.1 percent and 1.6 percent, respectively.

Figure 53 displays the wind speed and figure 54 shows the ambient temperature

and effective sky temperature recorded for October 8, 1980. Table 13 gives the

results for each test method for the October 8, 1980 comparison. The increase

in fractional energy savings is less than 0.8 percent for all methods. The

test method using the nonirradiated collector array with a downstream heat

source matches the performance of the irradiated system better than the two

other test methods. The thermal output of the nonirradiated array with an upstream heat source and the nonirradiated array with attached strip heaters

are almost identical; both are approximately 1.4 percent higher than the

thermal output of the normally irradiated array.

5.4 INFLUENCE OF THERMAL LOSSES FROM HEAT SOURCE

The electric heat source, even when highly insulated, will experience some heat

transfer to the surrounding environment. The numerical SDHW system model, in

conjunction with the heat source efficiency curve shown in figure 55, allows

the effect of a nonadiabatic heat source to be determined.

The test method using a nonirradiated collector array with a downstream heat source is used to generate the results shown in Tables 14 and 15. Each table

includes the predicted performance results of a system using an irradiated

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119
array, a system using a nonlrradlated collector array with an adiabatic

downstream heat source, and a system using a nonirradiated array with a non- adiabatic downstream heat source. Inclusion of the thermal losses based on

the measured heat source effficiency curve resulted in a fractional energy sav- ings decrease of 0.3 percent for both test days. The net thermal output of the

nonirradiated collector array with a downstream heat source decreased by 2.0

percent for November 17 conditions and 1.5 percent for November 25 conditions.

5.5 COMPARISON OF ANALYTICAL MODEL RESULTS WITH EXPERIMENTALLY MEASURED

RESULTS

The analytical model results can be compared to the experimental results for

the irradiated system and the system which uses a nonirradiated collector

array with a nonadiabatic downstream heat source. As shown in Table 14, the

analytical model predicts identical fractional energy savings for the irradiated system using October 6, 1980 environmental conditions, and for the system

which employs a nonirradiated collector array with a nonadiabatic downstream

heat source using November 17, 1980 environmental conditions. Identical fractional energy savings were measured experimentally for the outdoor irradiated

system on October 6, 1980 and the indoor nonirradiated collector array with

downstream heat source on November 17, 1980 (see Table 2). However, the measured fractional energy savings is 1.4 percent higher than that predicted by

the analytical model. The measured thermal output of the nonirradiated array

with the downstream heat source was 2.8 percent higher than the measured thermal output of the irradiated array. However, Table 14 shows that the analytical model predicts a decrease of 0.2 percent, assuming a nonadiabatic heat

source

The analytically predicted fractional energy savings (Table 15) for the system

which uses a nonirradiated collector array with a downstream nonadiabatic heat source subjected to November 25, 1980 environmental conditions is 0.1 percent

less than the predicted fractional energy savings for the irradiated system

subjected to October 8, 1980 environmental conditions. However, an increase of

3.0 percent in fractional energy savings was observed experimentally. Table 2 shows that the measured thermal output of the nonirradiated array with a downstream heat source was 1.6 percent higher than the measured thermal output of

the irradiated array. However, the analytical model predicts a decrease of 0.6

percent, as shown in Table 15.

120
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122
6. CONCLUSIONS AND RECOMMENDATIONS

Conclusions which may be drawn from this analytical and experimental

investigation of test methods for SDHW systems are as follows

(1) The thermal performance of a SDHW system can be duplicated indoors,

without the use of a solar simulator, by replacing the irradiated array

with an electric heat source alone, an electric heat source in series

with a .nonirradiated collector array, or by attaching electric strip

heaters to the back of nonirradiated absorber plates. All three indoor

test techniques reproduced the outdoor daily collector array thermal

output within 4 percent.

123
(2) Outdoor stagnation conditions cannot be duplicated in the laboratory using

an electric heat source only. Stagnation conditions and outdoor pump controller actions can be reproduced indoors using an electric heat source in

series with a nonirradiated collector array or by attaching strip heaters

to the back of the absorber plates contained within a nonirradiated

collector array.

(3) The ASHRAE Proposed Standard 95, which specifies the use of an electric

heat source in series with a nonirradiated collector array, is sensitive

to temperature stratification in the storage tank. The use of a return

tube which reduced internal mixing in the storage tank increased, for a selected test day, the fractional energy savings 74 percent to 83 percent.

The analytical model predicted a smaller increase in fractional energy

savings, approximately three percent. The analytical tank model does not

accurately simulate internal mixing within the storage tank, which accounts

for this discrepancy.

(4) Test result repeatability is excellent for the experimental technique of

using a nonirradiated collector array in series with a downstream heat source.

(5) Computer simulations indicate that the particular technique used to

supply the quantity of energy normally supplied by an irradiated array

has an insignificant effect on system performance for the SDHW system

used during this experimental investigation. Techniques considered during

this analytical investigation included a nonirradiated collector array

with upstream heat source, a nonirradiated array with downstream heat

source, and a nonirradiated array with strip heaters attached to the back

of the absorber plates. The thermal output of an irradiated collector

array can best be reproduced by using a nonirradiated collector array with

upstream heat source if outdoor environmental conditions are duplicated

indoors. For easily obtainable indoor environmental conditions, V^, = 0

and Tg^y = t^ , the use of a nonirradiated collector array with a downstream neat source gave the best agreement with the irradiated collector

array thermal output.

(6) The electric heat source should be constructed such that thermal losses

to the surrounding environment are minimized.

The following recommendations should be considered for future research;

(1) Correlations should be developed which relate ASHRAE Standard 95 test

results to the input meteorological conditions. These correlations would

enable the thermal performance of a SDHW system to be predicted for

environmental conditions which differ from the prescribed test conditions.

(2) Experiments should be conducted to validate the suggested test method for

SDHW systems which rely on thermosyphoning action for circulating fluid

through the collector array. The thermal performance of such systems,

using normally irradiated collector arrays, should be compared to an identical system where the array consists of collectors equipped with electric

strip heaters attached to the absorber plates.

124
Further investigations should be undertaken to improve the currently

proposed method of accounting for the effects of liquid flow rate and

composition of the absorber plate efficiency factor. An improved method

would eliminate the requirement that the same test fluid be used in

testing the collector and the entire SDHW system.

125
REFERENCES

1. "TRNSYS, A Transient Simulation Program," Engineering Experiment Station

Report 38-10, University of Wisconsin-Madison, Solar Energy Laboratory,

June, 1979.

2. Beckman, W. A., and Duffie, J. A., Solar Heating Design by the f-CHART

Method , John Wiley & Sons Inc., 1977.

3. Connolly, M., Giellis, R. , Jensen, C, and McMordie, R. , "Solar Heating

and Cooling Computer Analysis - A Simplified Design Method for Non-Thermal

Specialists," Proceedings of the Joint Conference of the American Section

of the International Solar Energy Society and the Solar Energy Society of

Canada, held in Winnipeg, Canada, August 15-20, 1976.

4. Fanney, A. H., and Liu, S. T., "Comparison of Experimental and ComputerPredicted Performance for Six Solar Domestic Hot Water Systems," ASHRAE

Transactions , Vol. 86, Part I, 1980, pp. 823-835.

5. Chandra, Subrato, and Khatter, Mukesh K. , "Analytical Investigations of

the Relative Solar Rating Concept," Draft Final Report, Florida Solar

Energy Center, Report No. FSEC-TT-80-6 , June 20, 1980.

6. "Method of Testing to Determine Thermal Performance of Solar Domestic

Water Heating Systems," ASHRAE Proposed Standard 95-P , February, 1981.

7. "Methods of Testing to Determine the Thermal Performance of Solar

Collectors," ASHRAE Standard 93-77 , 1977.

8. Hill, J. E., and Fanney, A. H., "A Proposed Procedure of Testing for

Rating Solar Domestic Hot Water Systems," ASHRAE Transactions , Vol. 86,

Part 1, 1980, pp. 805-822.

9. Kotas, J. F., and Wood, B. D., "Design and Evaluation of a Computer

Controlled Solar Collector Simulator," ASME Paper 8O-WA/S0I-6, November,

1980

10. Fanney, A. H. , and Thomas, W. C, "Simulation of Thermal Performance of

Solar Collector Arrays," ASME Transactions, Journal of Solar Energy

Engineering , submitted for publication, 1981.

11. Thomas, W. C, "Solar Collector Test Procedures: Development of a Method

to Refer Measured Efficiencies to Standardized Test Conditions," Report

VPI-E-80.23, Virginia Polytechnic Institute and State University,

November, 1980.

12. Duffie, J. A., and Beckman, W. A., Solar Engineering of Thermal Processes,

John Wiley & Sons Inc., 1980.

13. Gillett, W. B., "The Equivalence of Outdoor and Mixed Indoor/Outdoor Solar

Collector Testing," Solar Energy , Vol. 25, No. 6, 1980, pp. 543-548.

126
14. Buchberg , H. I., Catton, I., and Edwards, D. K. , "Natural Convection in

Enclosed Spaces — A Review of Application to Solar Energy Collection,"

Trans., ASME , Vol. 98C, May, 1976, pp. 184-188.

15. U.S. Standard Atmosphere, 1962 , U.S. Government Printing Office,

Washington, D.C., December, 1962, p. 14-15.

16. Streed, E. R., Thomas, W. C, Dawson, A. G., Ill, Wood, B. D., and

Hill, J. E., "Results and Analysis of a Round-Robin Test Program for

Liquid-Heating Flat-Plate Solar Collectors," NBS TN 975, August, 1978.

17. Thomas, W. C. , "Effects of Test Fluid, Flow Rate, and Flow Regime on Solar

Collector Thermal Performance Measurements," Modeling, Simulation, Testing,

and Measurements Symposum , Vol. H00138, ASME, 1978, pp. 85-89.

18. Thomas, W. C, "Effects of Test Fluid Composition and Flow Rates on the

Thermal Efficiency of Solar Collectors," Report ENG 80-03, Virginia

Polytechnic Institute and State University, Blacksburg, Virginia 24061.

19. Eckert, E. R. G., and Drake, R. M. , Analysis of Heat and Mass Transfer , McGraw-Hill, 1972, p. 351.

20. Seider, E. N., and Tate, C. E., "Heat Transfer and Pressure Drop of Liquids

in Tubes," Industrial Engineering Chemistry , Vol. 28, 1936, p. 1429.

21. Kreith, F., Principles of Heat Transfer , Third Edition, International

Textbook Co., 1973, p. 433.

127

APPENDIX A

ASHRAE STANDARD 95-P COLLECTOR HEAT REMOVAL FACTOR MODIFICATION PROCEDURE

The value of Fj^ is dependent upon mass flow rate and the specific heat of the

transfer fluid. Therefore, if the flow rate of transfer fluid through the

collector in the operation of the solar hot water system is different from the

value used during the ASHRAE 93-77 tests on the collector, then the value of

must be modified from the value obtained from the ASHRAE 93-77 tests. This can

be done utilizing the following procedure:

1. Calculate (Ta)^^ for the solar collector. For an ordinary flat-plate

collector

(Ta)n = ^^L^ (A.l)

1-(1 -

where

= 0.16 for a one-cover glass system

P(j = 0.24 for a two-cover glass system

= 0.29 for a three-cover glass system

For concentrating or other types of collectors, the value of (Ta)^^ must be

determined based on the geometrical and optical properties of the collector.

2. Calculate the value of Fr from the ASHRAE 93-77 tests by

A3

A ( ^«)n

Fr (A.2)

A/-(-)n

3. Calculate the value of the Ul from the ASHRAE 93-77 tests by -^a %Ul

Ul (A. 3)

Fr

where Fj^Ul is the absolute value of the slope of the efficiency curve

at a value of tggj- - t^ at solar noon during the system test.

G

A-1
Calculate the collector efficiency factor, F' by

F' = '"cS^c In

Aa Ul

|_ n»cCp,cJ

(A.4)

Once F' is determined, Fj^ can be calculated for any system's transfer fluid,

flow rate and specific heat by using m C

Fo = ^ P >^ 1 - exp ( - °^sS,s ,

(A.5)

It should be noted that the above correction technique is based on the

assumption that the collector absorber plate efficiency factor, F' , is not a function of flow rate.

A-2
APPENDIX B

TOP HEAT LOSS ANALYSIS

The purpose of this analysis, taken directly from reference 11, Is to develop

mathematical expressions for heat loss through the absorber-cover assembly and

to define a top loss coefficient. The conventional analysis summarized in

reference 12 applies only to materials with vanishingly small transmittance

for longwave (infrared) radiation. The limitation is a result of the expression

used for radiation heat transfer between a cover and either a parallel absorber

plate or another cover. The present analysis for the heat loss through the cover assembly accounts for strong (as well as weak) infrared transmittance.

A two-step iterative approach is used. The distribution of nodal (infrared)

radiosities is determined for specified absorber, ambient, and sky temperatures

and an estimated set of cover temperatures. Energy balances on the covers are then used to calculate improved estimates of the cover temperature distribution.

The analysis is developed for a one-cover collector.

Referring to figure 56, the top heat loss and loss coefficients are

Qt = Vl - 'cJ + [^1 - ^2]

Ut = Qt/tS - tj (B.2)

The usefulness of the latter coefficient is limited since it varies strongly

with temperature and includes the effect of radiant interchange with the sky at an effective temperature generally different from t^ . Nevertheless, IS^ (and

Ul) are convenient concepts and widely used in simplified, but adequate, analyses to predict the thermal performance of collectors.

Longwave radiation is assumed diffuse. The radiation properties of the upper

and lower surfaces are assumed equal. The temperature gradient in the covers

is neglected. Referring to figure 56, the nodal equations for determining the

infrared radiosity distribution are

^p ^cl

J2 = Eel + Pel Jl + ^cl Eg (B.4)

^3 = Eel + Tel Jl + Pel Eg (B.5)

The above equations may be solved simultaneously using an estimated cover temperature. Improved estimates of the cover temperature follow from an energy balance on the cover which yields

= Vl ^n + ^a - Jg] + [Jl - J3] + G qacl w + ^pcl h„ h

(B.6)

The result is used to improve the estimate of the radiosity distribution until

the cover temperature converges. The possibility exists, since tgky ^ t^, that

B-1
Solar (shortwave)

Radiation

1

Longwave Beam Diffuse

Radiation Connponent Component

1 ( -I r

COVER I

1^,

d3

'CI

ABSORBER

hpci

•J.

> >

^ 7

Fig. 56 Solar and infrared wavelength radiosity distributions in collectors with

1 cover

B-2
= 0 with tp = tg; in this case, eq (B.l) applies but eq (B.2) is meaningless.

Fu the oiTG y it \^ rioted ths. '^sky ^ ^a ^'^^ '^p '^^^^ tg , there can be a net

heat gain, i.e. Q^- < 0; this condition is conceivable for indoor tests with

solar simulators.

The (diffuse) longwave radiation properties of many commercially important

cover materials are not as well established as corresponding properties for

solar (shortwave) radiation. Nevertheless, reliable values of emittance and

longwave transmittance are needed in the analysis for accurate results.

Sky temperature and h^^ are estimated [12], respectively, from

tg^y = 0.0552 (t^/°K)l-5°K

h^ = [5.7 + 3.8 VJ W/(m2 C)

The free-convection coefficients between covers and plates are calculated using

the three-region correlation by Buchberg et al . [14] using temperature-dependent

air properties at one atmosphere [15].

(B.7)

(B.8)

CALCULATION OF FILM CONDUCTANCE BETWEEN THE TEST FLUID AND FLOW CHANNEL WALL

The following analysis is taken directly from reference 11. Suitable

expressions for calculating the film conductance between the test fluid and wall

of the flow channel are required. The flow regimes of practical importance

include laminar, transitional, and "fully turbulent" flow. In previous related

analyses [16-18] , an abrupt change from laminar to fully turbulent flow was assumed at a critical Reynolds number (of 2100). This previous procedure may

cause problems in analytical approaches for estimating collector parameters.

For the specified flow rate [7] of aqueous fluids, the strong dependence of

viscosity on temperature frequently results in Reynolds numbers well above the

practical upper limit for laminar flow.

Following a critical review of the available Nusselt number correlations [18],

the practical significance of uncertainties in calculated results, and the

benefit of increased computational time requirements, the following expressions

were selected for laminar [19] and fully turbulent [20] flows, respectively.

Nu(RePrdt^/L) = 3.66 +

0.0668 Re Pr d^/L

1 +0.04 (Re Pr d ^/L)0.667 [M(t„)_

0.14

(B.9)

where Re Pr d^/L < 1000, and

[^(j. ) 1

0.14

where 0.7 <. Pr £ 120 and 10,000 <. Re <^ 120,000. The latter expression is

essentially the Seider-Tate correlation. The coefficient of 0.021 was

B-3
selected to better match the data [20] at lower Reynolds numbers than the value

of 0.027 reported in many textbooks. Considering current designs, the L/d

ratio effect is strong for laminar flow but weak for turbulent flow, i.e., less

than 10 percent for L/d greater than 60 [21]. Consequently, entrance effects

are neglected for turbulent flow.

In the transition region Re^i ^ ^^c2 > ^ parabola is used for interpolation,

i .e.

Nu = a + b [Re - Re^^^] + c [Re - Re^2]

where a, b, and c are selected to satisfy

Nu (Re^i Pr d|^/L) = Nu]^ (from eq (B.9))

Nu (Rec2P^) = Nu2 (from eq (B.IO))

d Re

Nu _ d

Rec2

d Re

0.021 Pr1/3 ReO-

(B.ll)

(B.12)

(B.13)

Rec2

0.8

NU2_

Rec2

(B.14)

As supported by reference 20, the slope

turbulent correlation are matched by eq

is taken as 2100 < Re < 4000.

of the interpolating curve and fully

(B.14) at Re(,2 ^rid the transition range

B-4
APPENDIX C

DIMENSIONS AND PROPERTIES OF SOLAR COLLECTOR

INSTRUMENTATION SPECIFICATIONS

DATA CHANNEL ASSIGNMENT
DIMENSIONS AND PROPERTIES OF SOLAR COLLECTOR

Manufacturer and Model: Lennox LSC18-15

Component /Property

Absorber

Material

Flow configuration

Effective length

Effective width

Flow tubes : number

Material

OD

Hyd diameter

Wetted perimeter

Thickness

Thermal conductivity

Emittance

Solar absorptance (est)

Steel

Parallel

1.730 m

0.813 m

10

Copper

6 . 35 mm

4.93 mm

15.5 mm

0.90 mm

47.6 W/m-C

0.10

0.94

Cover

Material

Number of covers

Air space under cover Emittance

Transmittance

Index of refraction, solar

Extinction coeff, solar

Thickness

Glass

1 65.2 mm

0.88 mm

0.04

1.51

0.010 ram"l

3.18 mm

Insulation

Material

Thickness: back

edge

Conductivity: back

edge

Estimated

Gross Area

Glass Fiber

89.0 mm

25 .4 mm

0.040 W/m-C

0.040 W/m-C

1.4 W/m^-C

1.662 m2

C-2
SUPERIOR ELECTRIC COMPANY

AUTOTRANSFORMER SPECIFICATIONS

Voltage input

Voltage output

Maximum amperage output

Maximum output

Phase

240 volts ac

0 - 280 volts ac

28

7.8 kVa

Single

SCIENTIFIC COLUMBUS XL5C5-A2-1-7

WATT TRANSDUCER SPECIFICATIONS

Connection

Input voltage

Current input , nominal

Potential input

Potential burden

Watt input range

Output signal

Accuracy

Output ripple peak

Response time (to 99 percent)

Power factor range

Temperature range

Temperature influence (max.)

Operating humidity

Stability (per year)

Single Phase

0 - 240 Vac

0 - 15 a 120 Vac

0.1 Va at 120 Vac

0 - 3.5 kW

0 - 100 mv + 0.25 percent of indicated reading

< 0.5 percent of output

< 400 ms Unity to zero lead or lag

-20 C to +70 C

+ 0.005 percent/C

0-95 percent

+0.1 percent of output

C-3
Function

Full scale watts

Input voltage

Input current

Frequency

Power factor range

Operating temperature

effect on accuracy

Accuracy at 25 C

Digital readout

SCIENTIFIC COLUMBUS DIGITAL WATTMETER

MODEL 6335 PF SPECIFICATIONS

Three-phase, two-element wattmeter

One-phase, one-element wattmeter

Three phase: 0.4, 0.8, 1.2, 2, 2.4, 4, 6, 8,

12, 20, 24, 60 kW total

Single Phase: 0.2, 0.4, 0.6, 1, 1.2 2, 3, 4,

6, 10, 12, 30 kW per element

0-100, 200, 500 Vac, push button select

Max. with linearity 125 percent of range

Max. without linearity 150 percent of range

Burden 0.5 Va per element

2, 6, 20, 60 amps, push button select

Max. with linearity 124 percent of range

Max. without damage 200 percent of range

60+5 Hz

Unity to zero lead or lag

0 C to 50 C max. + 0.015 percent/G

Digital Readout + 0»25 percent of reading ~ Power factor

+ 1 digit

Analog Output + .2 percent of reading ~ Power factor

+ 0.03 percent of R.O

Direct reading in kW

C-4
EPPLEY PRECISION SPECTRAL PYRANOMETER

MODEL PSP SPECIFICATIONS

Sensitivity 9 microvolts per watt m~2 approx

Impedance 650 ohms approx.

Receiver circular, coated with Parsons' black

optical lacquer

Temperature dependence + 1 percent over ambient temperature range

-20 to +40 C

Linearity + 0.5 percent

Response time 1 second (1/e signal)

Cosine + 1 percent from normalization, 0-70° zenith angle + 3 percent, 70-80° zenith angle

EPPLEY PRECISION INFRARED RADIOMETER

(PYRGEOMETER) MODEL PIR SPECIFICATIONS

Sensitivity 3 microvolts/W m"^ approx.

Impedance 700 ohms approx.

Temperature dependence + 1 percent, -20 to 40 C (nominal)

Linearity + 1 percent

Response time 2 seconds (1/e signal)

Cosine response better than 5 percent from normalization,

insignificant for a diffuse source

C-5
NBS SOLAR DOMESTIC HOT WATER TEST FACILITY

DATA CHANNEL ASSIGNMENT

SYSTEM EMPLOYING ACTUAL SOLAR COLLECTORS EXPOSED

TO OUTSIDE METEOROLOGICAL CONDITIONS

CHANNEL NO. MEASUREMENT

0 Temperature Tank 0.15 m (6 in.) Elevation

1 Temperature Tank 0.30 m (12 in.) Elevation

2 Temperature Tank 0.46 m (18 in.) Elevation

3 Temperature Tank 0.61 m (24 in.) Elevation

4 Temperature Tank 0.76 m (30 in.) Elevation

5 Temperature Tank 0.91 m (36 in.) Elevation

6 Temperature Tank 1.07 m (42 in.) Elevation

7 Temperature Tank 1.22 m (48 in.) Elevation

8 Temperature Tank 1.37 m (54 in.) Elevation

9 Temperature Tank 1.52 m (60 in.) Elevation

10 None

11 None

12 Temperature Collector Supply at Tank

13 Temperature Collector Return at Tank

14 Temperature Collector Supply at Solar Array

15 Temperature Collector Return at Solar Array

16 Voltage (mV) Temperature Difference Across Solar

Storage Tank, Six-Junction Thermopile

17 Voltage (V) Flovnneter Analog Signal Output

18 ' Voltage (V) Circulator Pump Status, 1.5 Vdc On - 0.0 Vdc Off

19 Voltage (V) Wind Speed Anemometer

20 Voltage (mV) Storage Tank Heating Element Watt

Transducer Output

21 Voltage (mV) Temperature Difference Across Collector

Array, Six-Junction Thermopile

22 Voltage (mV) Radiation Level on Collector Surface

23 Voltage (mV) Effective Sky Temperature

24 Temperature Controller Tank Temperature Sensor

25 Temperature Controller Plate Temperature Sensor

26 Temperature Controller Collector Exit Pipe

Temperature Sensor

27 Temperature Collector Plate Temperature 1/4 Length

28 Temperature Collector Plate Temperature 1/2 Length

29 Temperature Collector Plate Temperature 3/4 Length

30 Temperature Supply Circulation Loop Temperature

31 Temperature Calibration Input Channel 1 32 None

33 Voltage (V) Storage Tank Heating Element Status

1.5 Vdc On - 0.0 Vdc Off

C-6
(CONTINUED)

CHANNEL NO. MEASUREMENT

34 None

35 Temperature Ambient Temperature Surrounding Tank

1.14 m Elevation

36 Temperature Ambient Temperature Surrounding Tank

0.62 m Elevation

37 None

38 None

39 Temperature Ambient Temperature Outdoors

40 Voltage (mV) Temperature Difference Across Hot

Water Load Three-Junction Thermopile

41 None

42 None

43 None

44 Voltage (mV) Integrated Temperature Difference

Across Hot Water Load

45 Temperature Cold Water Supply to Storage Tank

46 Temperature Hot Water Outlet From Storage Tank

47 Voltage (mV) Calibration Input Channel 2 48 None

49 None

50 None

C-7
NBS SOLAR DOMESTIC HOT WATER TEST FACILITY

DATA CHANNEL ASSIGNMENT

SYSTEM EMPLOYING NONIRRADIATED SOLAR COLLECTORS

CHANNEL MEASUREMENT

31 Temperature Calibration Channel No. 1 32 None

33 None

34 Voltage (V) Storage Tank Heating Element Status

1.5 Vdc On - 0.0 Vdc Off

35 None

36 None

37 Temperature Ambient Temperature Surrounding Tank

1.14 m Elevation

38 Temperature Ambient Temperature Surrounding Tank

0.62 m Elevation

39 None

40 Voltage (mV) Temperature Difference Across Hot Water

Load, Three-Junction Thermopile

41 Temperature Collector 1 Sky Cover

42 Temperature Collector 2 Sky Cover

43 Voltage (mV) Temperature Difference Across Collector

Inlet and Heat Source Outlet, Six-Junction

Thermopile

44 Voltage (mV) Integrated Temperature Difference Across

Hot Water Load, Three-Junction Thermopile

45 None

46 None

47 Temperature Calibration Channel No. 2

48 None

49 None

50 Temperature Tank 0.15 m (6 in.) Elevation

. 51 Temperature Tank 0.30 m (12 in.) Elevation

52 Temperature Tank 0.46 m (18 in.) Elevation

53 Temperature Tank 0.61 m (24 in.) Elevation

54 Temperature Tank 0.76 m (30 in.) Elevation

55 Temperature Tank 0.91 m (36 in.) Elevation

56 Temperature Tank 1.07 m (42 in.) Elevation

57 Temperature Tank 1.22 m (48 in.) Elevation

58 Temperature Tank 1.37 m (54 in.) Elevation

59 Temperature Tank 1.52 m (60 in.) Elevation

60 Temperature Cold Water Supply to Storage Tank

61 Temperature Hot Water Supply from Storage Tank

62 None

63 None

64 Temperature Collector Supply at Solar Array

C-8
CHANNEL MEASUREMENT

65 Temperature

66 Temperature

67 Voltage (mV)

68 None

69 Temperature

70 Voltage (mV)

71 Temperature

72 Temperature

73 Voltage (mV)

74 Temperature

75 Temperature

76 Voltage (mV)

77 Temperature

78 Temperature

79 Temperature

80 Temperature

81 Temperature

82 Voltage (mV)

83 None

84 Voltage (mV)

85 Voltage (mV)

86 Voltage (mV)

87 Temperature

88 Voltage (V)

89 Temperature

Collector Outlet at Solar Array

Ambient Temperature Above Collectors

Temperature Difference Between Inlet

to Solar Collector Array and Ambient

Above Solar Collector Array, SixJunction Thermopile

Heat Source Outlet

Temperature Difference Across Solar

Collector Array, Six-Junction Thermopile

Inlet to Heat Source

Outlet from Heat Source

Temperature Difference Across Heat

Source, Six-Junction Thermopile

Supply at Tank

Return at Tank

Temperature Difference Across Solar

Storage Tank, Six-Junction Thermopile

Controller Tank Temperature Sensor

Controller Exit Pipe Temperature Sensor

Controller Plate Temperature 1/4 Length

Controller Plate Temperature 1/2 Length

Controller Plate Temperature 3/4 Length

Flowmeter

Heat Source Watt Transducer Output

Storage Tank Watt Transducer Output

Effective Sky Temperature

Collector Plate Temperature 1/8 Length

Zone Valve Status

1.5 Vdc Collector-Heat Source-Tank Loop

0.0 Vdc Collector-Heat Source Loop

Inlet Temperature to Solar Collector Array

C-9
TRENDSCAN 1000 PERFORMANCE SPECIFICATIONS

Measurement technique : Dual slope integrating auto-zeroing measuring circuit,

fully-guarded for high common mode rejection.

Linearization : Variable-length segment digital linearization.

Reference junction compensation (T.C. Ranges) ; Electrically compensated to 0

reference junction. A temperature sensitive component for each T.C. range is

located on the range card for each range in the Trendscan 1000.

Temperature Coefficient : (5 to 20 C and 30 to 45 C)

Normal mode rejection : (At range card input terminals)

Common mode rejection :

Self -test voltage accuracy :

Max. common-mode voltage :

Max. source resistance :

Input impedance

Repeatability ;

Range Total

Description Range

Copper-Constantan -200 C to + 400 C

EMF + 4 V Linear

EMF + 10 V Linear

+ (0.015 percent of reading +0.001

percent of range/C)

70 dB at 50-60 Hz, 50 dB at 100/120

Hz increasing 12 dB/octave beyond

100/120 Hz

120 db at DC, 130 dB at 50/60 Hz with

1000 Q, line unbalance

+ (0.05 percent of reading)

250 Vdc or Vac Peak

2500 ^

100 Megohms (all ranges)

+ (0.05 percent of reading)

Display

Resolution System Accuracy

1 Year

25 + 10 C

0.1 C 0.9 C

0.1 mV + (0.02 percent

+0.4 mV)

1 mV + (0.02 percent + 4 mV)

C-10
NBS-n4A IREV. 2-BC) U S DEPT. OF COMM. 1. PUBLICATION OR 2. Performing Organ, Report No, 3. Publication Date BIBLIOGRAPHIC DATA

REPORT NO. February 1982 SHEET (See instructions) NBS BSS 140

4, TITLE AND SUBTITLE

Analytical and Experimental Analysis of Procedures for Testing Solar Domestic Hot

Water Systems

5. AUTHOR(S)

A.H. Fanney, W.C. Thomas, C.A. Scarbrough, CP. Terlizzi

6. PERFORMING ORGANIZATION (If joint or other than NBS. see /n struct/on sj NATIONAL BUREAU OF STANDARDS

DEPARTMENT OF COMMERCE

WASHINGTON, D C 20234

7. Contract/Grant No,

8. Type of Report & Period Covered Final

9. SPONSORING ORGANIZATION NAME AND COMPLETE ADDRESS (Street, City. State. ZIP)

Office of Solar Applications for Buildings

U.S. Department of Energy

Washington, D C 20585

Library of Congress Catalog Card Number: 81-600191

10. SUPPLEMENTARY NOTES

11. ABSTRACT

A repeatable test method independent of outdoor environmental conditions and laboratory geographical location is required in order to provide a means by which solar domestic hot water systems may be rated and compared. Three experimental techniques which allow the net thermal output of an irradiated solar collector array to be reproduced indoors

without the use of a solar simulator are investigated. These techniques include use of an electric heat source only, use of a nonirradiated collector array in series with an electric heat source, and the use of electric strip heaters which are attached to the

back of nonirradiated absorber plates. Expressions are developed to compute the input

power required for each experimental technique. Solar collectors connected in parallel

and series combinations are considered.

All three test techniques were shown to reproduce the outdoor daily collector array

thermal output within four percent. Two of the techniques allow the actions of the circulator controller for an outdoor irradiated system to be duplicated indoors. One technique applies to solar hot water systems which operate on the thermosyphon principle.

Experiments conducted to determine the effect of storage tank temperature stratification on system performance for a single-tank direct solar hot water system are described. Several return tube designs, which introduce the solar heated water into the

storage tank, were fabricated and tested to determine the influence of thermal stratification on system performance. The best return tube design increased the performance of

the single-tank direct system approximately ten percent compared to a conventional re- turn tube design.

An analytical model for a single-tank direct hot water system is developed. The

model is used to support parametric studies for the thermal performance characteristics

which result from the use of each test method to duplicate the net thermal output of an irradiated array. The model is also used to assess thermal performance differences

which occur due to indoor versus outdoor environmental conditions.

12. KEY WORDS (Six to twelve entries; alphabetical order; capitalize only proper names; and separate key words by semico/ons;

ASHRAE Standard 95; collectors in parallel; electric strip heaters; environmental con- ditions; indoor testing; modeling; NBS; solar; solar domestic hot water system; stratification; test method . 13. AVAILABILITY

Unlimited For Official Distribution. Do Not Release to NTIS

Order From Superintendent of Documents, U.S. Government Printing Office, Washington, D C

20402.

Order From National Technical Information Service (NTIS), Springfield, VA 22 16 I

14, NO, OF

PRINTED PAGES

158

15, Price$6 .00

O U.S. GOVERNMEOT PRINTING OmCE : 1982 0 - 366-393 USCOMM-DC 6043-PeO

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