Development of a Simulation Model for Active Chilled Beam Systems with an Integrative Experimental and Modeling Approach

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DEVELOPMENT OF A SIMULATION MODEL FOR ACTIVE CHILLED BEAM SYSTEMS WITH AN INTEGRATIVE EXPERIMENTAL AND MODELING APPROACH

by DAREUM NAM B.S., Seoul National University, 2013

A dissertation submitted to the Faculty of the Graduate School of the University of Colorado in partial fulfillment of the requirement for the degree of Doctor of Philosophy Department of Civil, Environmental and Architectural Engineering University of Colorado Boulder 2019

This dissertation entitled: Development of a Simulation Model for Active Chilled Beam Systems with an Integrative Experimental and Modeling Approach written by Dareum Nam has been approved for the Department of Civil, Environmental and Architectural Engineering

______John Z. Zhai, Ph.D.

______Moncef Krarti, Ph.D.

______Gregor P. Henze, Ph.D.

______Wangda Zuo, Ph.D.

______Jaeyoon Koh, Ph.D.

Date ______

The final copy of this dissertation has been examined by the signatories, and we find that both the content and the form meet acceptable presentation standards of scholarly work in the above mentioned discipline.

ABSTRACT

Nam, Dareum (Ph.D., Architectural Engineering)

Development of a Simulation Model for Active Chilled Beam Systems with an Integrative

Experimental and Modeling Approach

Dissertation directed by Professor John Z. Zhai

Active chilled beam systems have gained attention in the U.S. and Europe in recent years due to their superior thermal and energy performance. This gaining popularity has led to various experimental and simulation studies. Most of the experimental studies have focused on airflow patterns in spaces to evaluate indoor thermal comfort. For simulation studies, the Computational

Fluid Dynamics (CFD) simulations were widely used to study the thermal performance of active chilled beams. A limited number of studies were reported to verify the prediction of active chilled beams in commonly used building simulation tools with the experimental data. This dissertation is to test the performance of active chilled beam in a laboratory environment, and evaluate the existing models implemented in various simulation tools based on experimental data, and then develop the enhanced model by using an integrative experimental and modeling approach to fill the gap between simulation models and experimental data.

The experiment was planned under the guidance of the American Society of Heating,

Refrigerating and Air-Conditioning Engineers (ASHRAE) Standard 200 and the European

Standard (EN) 15116. Specifically, requirements for insulation, infiltration, experimental instrument settings are set to comply with the two standards. Most importantly, steady-state testing requirements and uncertainty requirements were satisfied. Six active chilled beam units

iii with different specifications provided by two manufacturers were used in the experiment: the six units are widely used in the field and serve as good representations of the application. The experiments were carried out in a full-scale building system laboratory located at the University of Colorado Boulder. The water-side cooling capacity, the air-side cooling capacity were calculated with the measurement under various ranges of the supply air flow rate and the supply water temperature. The study compared the experiment results of each active chilled beam with a focus on the water-side and the air-side cooling capacities as well as their correlations with the water and air supply conditions. The water-side cooling capacity increases when the primary air flow rate increases or when the water inlet temperature decreases. The air-side cooling capacity increases when the primary air flow rate increases or when the water inlet temperature increases.

The study further validated the impacts of air supply nozzle sizes of active chilled beams on the cooling performance, which is that smaller nozzle size provides a better transfer efficiency.

Induction ratio analysis with the Thermal Balance Method, which was included in the appendix of the new version of the ASHRAE Standard 200-2018, was also conducted by using experimentally measured parameters. It was found that the induction ratio varies even in one active chilled beam unit according to the supply (primary) air flow rate and the supply water temperature. The induction ratio tends to decrease as the water inlet temperature increases or the supply air flow increases. Besides, when the supply air flow rate is high enough (above 50 cfm approximately), the supply air flow rate has less impact on the induction ratio.

The performances of prevalent simulation tools were evaluated and compared to the experiment results. Among the existing eight energy simulation tools that are capable of modeling active chilled beams, three representative simulation programs were selected based on the calculation method: EnergyPlus, TRNSYS, and eQUEST. EnergyPlus represents the tools

iv that use the empirical model of active chilled beams. TRNSYS represents the tools that use energy balance equations to simulate an active chilled beam, which is based on the heat transfer process around the cooling coil in the active chilled beam. eQUEST represents the tools that use the induction unit model as a workaround to simulate an active chilled beam system; it does not have a dedicated model for active chilled beam units. The current active chilled beam simulation models cannot accurately predict the energy performance of the active chilled beams. The active chilled beam model of EnergyPlus uses empirical coefficients that are not easily accessible to the users and have to be provided by the manufacturer. The active chilled beam model of TRNSYS has several assumptions that make the result of the model inaccurate. The induction unit model of eQUEST cannot provide a reasonable result as the active chilled beam models of EnergyPlus and TRNSYS do.

Based on the experimental and the simulation analysis, the new model for the prediction of the energy performance of active chilled beams was developed and integrated with TRNSYS and EnergyPlus. The new active chilled beam model is grounded with the algorithm with the heat exchanger theory and easy-to-measure input parameters upon the findings from the experiment. By using both the Logarithmic Mean Temperature Difference (LMTD) Method and the Effectiveness – Number of Transfer Units (NTU) method, the new model allows the simulation of the energy performance of active chilled beams to be accurate under a wide range of system capacities and operating conditions with fewer input data. In terms of the water-side cooling capacity, the result of the new model shows the percentage differences that range from

0.04% to 8.4% from the experiment result, whereas the results with the original EnergyPlus and

TRNSYS model gives the percentage differences that range from 0.08% to 30.24%.

ACKNOWLEDGEMENTS

I would like to express my sincere gratitude to my advisor, Professor John Zhai. Over my

Ph.D. journey, he has been an amazing mentor to me in my academic and professional development. I have learned everything from him to be a better researcher. I am very grateful that I have met my Ph.D. advisor like him. I would also like to recognize my committee members, Dr. Krarti, Dr. Henze, Dr. Zuo, and Dr.Koh for their critical input and wise direction that enhanced the quality of this dissertation.

The research presented in this dissertation was sponsored by the American Society of

Heating, Refrigerating and Air-Conditioning Engineers (ASHRAE) through 1629-TRP “Testing and Modeling Energy Performance of Active Chilled Beam Systems.” I would like to thank the members of TC5.3 – Room Air Distribution for their support throughout this research. I also appreciate Halton and Price Industries for their support of this project. This research would not have been possible without their generosity and assistance. Special thanks go to Chris Lowell from Halton and Mikhail Koupriyanov from Price Industries.

The work by the author was also supported by the Department of Civil, Environmental and Architectural Engineering (CEAE) at the University of Colorado Boulder through a Doctoral

Assistantship for Completion of Dissertation. The author would like to thank the chair of the

CEAE department.

This study is dedicated to my family – my husband, parents, sister, brother-in-law, and lovely nieces - I love you guys so much. Lastly, I would like to say thank you to my husband who has been a constant source of support and encouragement during the challenges of this

Ph.D. journey. I am truly grateful to God every day for having you in my life.

TABLE OF CONTENTS

Chapter 1 Introduction ...... 1

1.1 Introduction and Background ...... 1

1.2 Significance of Research ...... 6

1.3 Research Objectives ...... 8

1.4 Research Method ...... 8

Chapter 2 Literature Review...... 10

2.1 Fundamentals of Active Chilled Beam ...... 10

2.2 Overall Literature Review ...... 12

2.3 Review on Simulation Models ...... 19

Chapter 3 Full-Scale Laboratory Experiment Analysis ...... 22

3.1 Experiment Objective ...... 22

3.2 Experimental Plan Development ...... 22

3.2.1 Experiment Location: Larson Building System Laboratory ...... 23

3.2.2 Selecting Active Chilled Beam Units ...... 25

3.2.3 Experiment Conditions ...... 26

3.2.4 Experiment Cases...... 27

3.2.5 Parameters to Measure ...... 28

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3.2.6 Experiment Procedure ...... 29

3.3 Experimental Quality Requirements ...... 30

3.3.1 Insulation Requirement ...... 30

3.3.2 Infiltration Requirement...... 32

3.3.3 Instrumental Requirement ...... 34

3.3.4 Steady-state Requirement ...... 36

3.3.5 Uncertainty Requirement ...... 37

3.4 Laboratory Setup and Systems ...... 38

3.4.1 Laboratory system overview ...... 38

3.4.2 Active Chilled Beam ...... 40

3.4.3 Internal Heat Gain ...... 46

3.4.4 Sensor Measurement Location ...... 47

3.4.5 Tests before the Experiment – Smoke Test ...... 48

3.4.6 Tests before the Experiment – Stability Test for the Experiment Equipment ...... 48

3.5 Experimental Result Analysis ...... 50

3.5.1 Steady-State Requirement ...... 51

3.5.2 Experiment Result and Analysis: Impacts of each parameter on the water-side

cooling capacity and the air-side cooling capacity ...... 51

3.5.3 Uncertainty Analysis ...... 66

3.5.4 Induction Ratio Analysis – Thermal Balance Method ...... 66

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Chapter 4 Simulation Analysis ...... 71

4.1 Overview of Simulation Programs ...... 71

4.1.1 Pros and Cons ...... 72

4.1.2 Inputs and Outputs ...... 76

4.1.3 Modeling ...... 79

4.1.4 Building Thermal Load Comparison ...... 80

4.2 Active Chilled Beam Simulation in EnergyPlus ...... 82

4.2.1 Mechanism of the Active Chilled Beam Model in EnergyPlus ...... 82

4.2.2 Empirical Coefficients ...... 83

4.2.3 Induction Ratio Analysis...... 84

4.2.4 Experiment Cases Simulation Results ...... 85

4.3 Active Chilled Beam Simulation in TRNSYS ...... 105

4.3.1 Mechanism of the Active Chilled Beam Model ...... 105

4.3.2 Experiment Cases Simulation Result ...... 110

4.4 Active Chilled Beam Simulation in eQUEST ...... 124

4.4.1 Mechanism of the Active Chilled Beam Model ...... 124

4.4.2 Experiment Cases Simulation Result ...... 125

Chapter 5 Sensitivity Analysis of the Existing Simulation Models ...... 128

5.1 Sensitivity Analysis of EnergyPlus ...... 128

5.1.1 Impact of Calculated Induction Ratio ...... 128

5.1.2 Impact of Empirical Coefficients ...... 129

5.2 Sensitivity Analysis of TRNSYS ...... 136

5.2.1 Impact of Calculated Induction Ratio ...... 136

5.2.2 Impact of Bypass Fraction ...... 138

5.2.3 Impact of Inputs Regarding Pressure of Primary Air and Secondary Air ...... 139

5.3 Necessity of Developing New Simulation Model for Active Chilled Beam ...... 142

Chapter 6 Development of a New Simulation Model ...... 143

6.1 Background: Heat Exchanger...... 143

6.1.1 The Logarithmic Mean Temperature Difference (LMTD) Method ...... 145

6.1.2 The Effectiveness – Number of Transfer Units (NTU) Method ...... 148

6.1.3 Active Chilled Beam Experimental Cases ...... 153

6.2 Development of New TRNSYS Model ...... 155

6.2.1 Algorithm of the new TRNSYS model ...... 155

6.2.2 Simulation Results with the new TRNSYS model ...... 158

6.3 Development of New EnergyPlus Model...... 163

6.3.1 Algorithm of the new EnergyPlus model...... 163

6.3.2 Simulation Results with the new EnergyPlus model ...... 164

6.4 Comparison of the Results from the New Models ...... 169

6.5 Validation with other experiment cases ...... 173

6.6 Discussions ...... 175

Chapter 7 Conclusion and Future Work ...... 182

7.1 Conclusion ...... 182

7.2 Contributions ...... 184

7.3 Future Work ...... 185

BIBLIOGRAPHY ...... 186

A. APPENDIX ...... 189

1) Steady-state Requirement for the Experiment Cases ...... 189

2) Uncertainty Requirement for the Experiment Cases ...... 191

3) Thermal Balance Method ...... 192

4) Code script: the main algorithm of the new TRNSYS model ...... 196

5) Code script: the main algorithm of the new EnergyPlus model ...... 198

LIST OF TABLES

Table 3.1 Selected active chilled beam units (Manufacturer And model names are blinded) ...... 26

Table 3.2 Experiment conditions ...... 27

Table 3.3 Experiment cases details ...... 28

Table 3.4 Experimental parameters ...... 29

Table 3.5 Environmental parameters ...... 29

Table 3.6 Envelope construction of the test chamber ...... 31

Table 3.7 Test chamber insulation requirement ...... 31

Table 3.8 Test chamber infiltration requirement ...... 34

Table 3.9 The sensor accuracy requirement of the ASHRAE Standard 200 for key parameters . 34

Table 3.10 Calibration result of thermistors for air temperature ...... 36

Table 3.11 Steady-state requirements ...... 37

Table 3.12 Numbers of nozzles and nozzle diameters of each active chilled beam unit ...... 43

Table 3.13 Numbers of fins on cooling coil of each active chilled beam units ...... 44

Table 3.14 Calculated Bypass Fraction of each active chilled beam unit ...... 46

Table 3.15 Example of standard deviation data for the steady-state requirement ...... 51

Table 3.16 Nozzle diameter of each active chilled beam (PA means primary air) ...... 63

Table 3.17 Water-side cooling capacity of each experiment case ...... 65

Table 3.18 Example of Uncertainty requirement data ...... 66

Table 3.19 Induction ratio calculation result by Thermal Balance Method ...... 69

Table 4.1 Key parameters and three simulation programs ...... 76

Table 4.2 Inputs and outputs in EnergyPlus ...... 77

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Table 4.3 Inputs and outputs in TRNSYS ...... 78

Table 4.4 Inputs and outputs in eQUEST ...... 78

Table 4.5 Building cooling load comparison ...... 81

Table 4.6 The percentage difference of the water-side cooling capacity between the EnergyPlus model vs experimental results (Default Coeff.) ...... 105

Table 4.7 The percentage difference of the water-side cooling capacity between the EnergyPlus model vs experimental results (Empirical Coeff.) ...... 105

Table 4.8 the percentage difference of the water-side cooling capacity between the TRNSYS model vs experimental results ...... 123

Table 5.1 Default coefficients in EnergyPlus and Empirical coefficient of product AH ...... 130

Table 5.2 Sensitivity analysis for empirical coefficients ...... 132

Table 5.3 Randomly fixed inputs ...... 140

Table 6.1 Effectiveness relations for heat exchangers (Kays & London, 1984) ...... 150

Table 6.2 NTU relations for heat exchangers (Kays & London, 1984) ...... 151

Table 6.3 Calculated Correction factor F for each experiment case ...... 154

Table 6.4 Effectiveness ε for each experiment case ...... 154

Table 6.5 the percentage difference of the water-side cooling capacity between the new TRNSYS model vs experimental results ...... 162

Table 6.6 the percentage difference of the water-side cooling capacity between the original

TRNSYS model vs experimental results ...... 162

Table 6.7 the percentage difference of the water-side cooling capacity between the new

EnergyPlus model vs experimental results ...... 169

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Table 6.8 the percentage difference of the water-side cooling capacity between the original

EnergyPlus model (default coeff.) vs experimental results ...... 169

Table 6.9 New experiment cases for the validation of the new model ...... 173

Table 6.10 Unit AL, Case 7: supply water temp. = 55°F (12.8°C), supply air flow rate = 40 cfm

(18.9 L/s) ...... 175

Table 6.11 Unit BH, Case 9: supply water temp. = 61°F (16.1°C), supply air flow rate = 80 cfm

(37.8 L/s) ...... 176

Table 6.12 Simulation results of the models with three convergence conditions ...... 178

Table 6.13 End use annual energy consumption in transient state ...... 180

Table A.1 Steady-state requirement for Manufacturer A’s Low flow range unit ...... 189

Table A.2 Steady-state requirement for Manufacturer A's Middle flow range unit ...... 189

Table A.3 Steady-state requirement for Manufacturer A's High flow range unit ...... 189

Table A.4 Steady-state requirement for Manufacturer B's Low flow range unit...... 190

Table A.5 Steady-state requirement for Manufacturer B's Middle flow range unit ...... 190

Table A.6 Steady-state requirement for Manufacturer B's High flow range unit ...... 190

Table A.7 Uncertainty requirement for Manufacturer A's Low flow range unit ...... 191

Table A.8 Uncertainty requirement for Manufacturer A's Middle flow range unit ...... 191

Table A.9 Uncertainty requirement for Manufacturer A's High flow range unit ...... 191

Table A.10 Uncertainty requirement for Manufacturer B's Low flow range unit ...... 191

Table A.11 Uncertainty requirement for Manufacturer B's Middle flow range unit ...... 192

Table A.12 Uncertainty requirement for Manufacturer B's High flow range unit ...... 192

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LIST OF FIGURES

Figure 1.1 Illustration of passive and active chilled systems...... 2

Figure 1.2 Illustration of active chilled beams ...... 3

Figure 1.3 Chilled beams market share by types in Europe, 2010 ...... 5

Figure 3.1 3-D view of the full-scale HVAC laboratory at the University of Colorado Boulder 24

Figure 3.2 Photos of the test chamber at the beginning ...... 25

Figure 3.3 Test chamber before/after applying 3M air and vapor barrier 3015 tape ...... 33

Figure 3.4 Blower door test on test chamber ...... 33

Figure 3.5 Radiant shielded sensors - surface/air temperature sensors ...... 36

Figure 3.6 Laboratory setup and systems ...... 38

Figure 3.7 The air-side System ...... 38

Figure 3.8 The water-side system and the data acquisition system ...... 39

Figure 3.9 Insulation of the chilled water tank - fiberglass and bubble foil insulations ...... 39

Figure 3.10 Installation of an active chilled beam ...... 40

Figure 3.11 Primary air inlets – the opposite side ...... 41

Figure 3.12 Different positions for hanger rod deck...... 42

Figure 3.13 Cooling coils in active chilled beam units – Manufacturer A (Left) and Manufacturer

B (Right) ...... 42

Figure 3.14 Nozzles in an active chilled beam unit (after the hinged grille face was opened) .... 43

Figure 3.15 ACB Frame that surrounds the cooling coil inside – BL, BM, and BH ...... 45

Figure 3.16 ACB Frame that surrounds the cooling coil inside – AL, AM, and AH ...... 45

Figure 3.17 Electrical blankets placement, heat level setting, and the electricity usage monitor 47

Figure 3.18 Sensor locations around the active chilled beam unit ...... 47

Figure 3.19 Sensor locations for surfaces/room conditions ...... 48

Figure 3.20 View from the chamber's door ...... 49

Figure 3.21 Reference temperature sensors – less than 1 inch below the active chilled beam unit

...... 49

Figure 3.22 Sensors around the active chilled beam unit ...... 50

Figure 3.23 Manufacturer A, Low flow range unit ...... 52

Figure 3.24 Manufacturer A, Middle flow range unit ...... 53

Figure 3.25 Manufacturer A, High flow range unit ...... 54

Figure 3.26 Manufacturer B, Low flow range unit ...... 55

Figure 3.27 Manufacturer B, Middle flow range unit ...... 56

Figure 3.28 Manufacturer B, High flow range unit ...... 57

Figure 3.29 Total cooling capacity of Manufacturer A's active chilled beams’ cases ...... 59

Figure 3.30 The water-side/air-side cooling capacity result of Manufacturer A's ACB cases ..... 60

Figure 3.31 Total cooling capacity of Manufacturer B's ACB cases ...... 61

Figure 3.32 The water-side/air-side cooling capacity result of Manufacturer B's ACB’ cases .... 62

Figure 3.33 Water-side performance of a typical active chilled beam versus primary air flow

(PRICE, 2011)...... 64

Figure 3.34 Temperature sensors positioned on the coil-fin-face ...... 68

Figure 3.35 Coil-fin-face temperature sensor positioning layout ...... 68

Figure 4.1 Building geometry configuration in simulations ...... 79

Figure 4.2 Building cooling load comparison...... 81

Figure 4.3 Induction ratio analysis result sample - AL unit's cases ...... 85

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Figure 4.4 Manufacturer A, Low flow range unit’s EnergyPlus result (Default Coeff.) ...... 86

Figure 4.5 Manufacturer A, Middle flow range unit’s EnergyPlus result (Default Coeff.) ...... 87

Figure 4.6 Manufacturer A, High flow range unit’s EnergyPlus result (Default Coeff.) ...... 88

Figure 4.7 Manufacturer B, Low flow range unit’s EnergyPlus result (Default Coeff.) ...... 89

Figure 4.8 Manufacturer B, Middle flow range unit’s EnergyPlus result (Default Coeff.) ...... 90

Figure 4.9 Manufacturer B, High flow range unit’s EnergyPlus result (Default Coeff.) ...... 91

Figure 4.10 Manufacturer A, Low flow range unit’s EnergyPlus result (Empirical Coeff.) ...... 93

Figure 4.11 Manufacturer A, Middle flow range unit’s EnergyPlus result (Empirical Coeff.) .... 95

Figure 4.12 Manufacturer A, High flow range unit’s EnergyPlus result (Empirical Coeff.) ...... 97

Figure 4.13 Manufacturer B, Low flow range unit’s EnergyPlus result (Empirical Coeff.) ...... 99

Figure 4.14 Manufacturer B, Middle flow range unit’s EnergyPlus result (Empirical Coeff.) .. 101

Figure 4.15 Manufacturer B, High flow range unit’s EnergyPlus result (Empirical Coeff.) ..... 103

Figure 4.16 Manufacturer A, Low flow range unit’s TRNSYS result...... 112

Figure 4.17 Manufacturer A, Middle flow range unit’s TRNSYS result ...... 114

Figure 4.18 Manufacturer A, High flow range unit’s TRNSYS result ...... 116

Figure 4.19 Manufacturer B, Low flow range unit’s TRNSYS result ...... 118

Figure 4.20 Manufacturer B, Middle flow range unit’s TRNSYS result ...... 120

Figure 4.21 Manufacturer B, High flow range unit’s TRNSYS result ...... 122

Figure 4.22 How to separate the water-side/air-side cooling capacity ...... 125

Figure 4.23 Manufacturer A’s Low flow rate unit’s cases eQUEST result ...... 127

Figure 5.1 AH's result with the rated/calculated induction ratio ...... 129

Figure 5.2 AH's result with the default /manufacturer-provided empirical coefficients ...... 131

Figure 5.3 Water-side cooling capacity calculator based on EnergyPlus equations ...... 132

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Figure 5.4 Sensitivity analysis for empirical coefficients ...... 134

Figure 5.5 AM's cases - when n2 = 0.6 instead of 0.7 ...... 135

Figure 5.6 AM's cases - when n3 = 0 instead of 0.17 ...... 135

Figure 5.7 AH's water-side cooling capacity with the rated/calculated induction ratio ...... 137

Figure 5.8 BM's water-side cooling capacity with rated/calculated induction ratio ...... 138

Figure 5.9 AH's water-side cooling capacity with Bypass Fraction = 0 and 0.055 ...... 139

Figure 5.10 AH’s result with Base case, Case 1, and Case 2: primary/secondary air inlet pressure change ...... 140

Figure 5.11 AH’s result with Base case, Case 3, and Case 4: primary/secondary air pressure drop change ...... 141

Figure 6.1 Flow configuration of Parallel flow and Counter flow in a double-pipe heat exchanger

(Cengel et al., 2016) ...... 144

Figure 6.2 Flow configuration of Cross-flow: Both fluids unmixed and one fluid mixed/one fluid unmixed (Cengel et al., 2016) ...... 144

Figure 6.3 ∆T1 and ∆T2 expressions in a parallel-flow and a counter flow heat exchanger (Cengel et al., 2016) ...... 146

Figure 6.4 The configuration of ∆T1 and ∆T2 in a cross-flow heat exchanger (Cengel et al., 2016)

...... 147

Figure 6.5 Correction factor F charts for cross-flow heat exchanger, both fluid unmixed (Cengel et al., 2016) ...... 147

Figure 6.6 Correction factor F charts for cross-flow heat exchanger, one fluid mixed and the other unmixed (Cengel et al., 2016)...... 148

Figure 6.7 Effectiveness for heat exchangers (Cengel et al., 2016; Kays & London, 1984) ...... 152

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Figure 6.8 Flow chart of the algorithm of the new ACB model ...... 158

Figure 6.9 The water-side cooling capacity result of AL unit with the new TRNSYS model ... 159

Figure 6.10 The water-side cooling capacity result of AM unit with the new TRNSYS model 159

Figure 6.11 The water-side cooling capacity result of AH unit with the new TRNSYS model . 160

Figure 6.12 The water-side cooling capacity result of BL unit with the new TRNSYS model . 160

Figure 6.13 The water-side cooling capacity result of BM unit with the new TRNSYS model 161

Figure 6.14 The water-side cooling capacity result of BH unit with the new TRNSYS model . 161

Figure 6.15 The number of reduced input values: the original model → the new model ...... 163

Figure 6.16 The water-side cooling capacity result of AL unit with the new EnergyPlus model

...... 165

Figure 6.17 The water-side cooling capacity result of AM unit with the new EnergyPlus model

...... 166

Figure 6.18 The water-side cooling capacity result of AH unit with the new EnergyPlus model

...... 166

Figure 6.19 The water-side cooling capacity result of BL unit with the new EnergyPlus model

...... 167

Figure 6.20 The water-side cooling capacity result of BM unit with the new EnergyPlus model

...... 167

Figure 6.21 The water-side cooling capacity result of BH unit with the new EnergyPlus model

...... 168

Figure 6.22 Total simulation results comparison of AL unit ...... 170

Figure 6.23 Total simulation results comparison of AM unit ...... 170

Figure 6.24 Total simulation results comparison of AH unit ...... 171

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Figure 6.25 Total simulation results comparison of BL unit ...... 171

Figure 6.26 Total simulation results comparison of BM unit ...... 172

Figure 6.27 Total simulation results comparison of BH unit ...... 172

Figure 6.28 Water-side cooling capacity experiment/simulation results of new cases ...... 174

Figure 6.29 Percentage difference of the water-side cooling capacity of each simulation results

...... 174

Figure 6.30 The flow chart of the new calculation procedure ...... 177

Figure 6.31 Fluctuating results of the water outlet temperature by timesteps ...... 178

Figure 6.32 Water-side cooling capacity results with transient state ...... 180

Figure A.1 Temperature sensors positioned on the coil-fin-face ...... 194

Figure A.2 Temperature sensors positioning layout ...... 194

Chapter 1 Introduction

1.1 Introduction and Background

Active chilled beams are cooling terminal units that integrate hydronic cooling systems with the primary ventilation system. Active chilled beam systems have gained attention due to their better thermal and energy performance, high quality of indoor air conditions, thermal comfort, and a low noise level with a reasonable life-cycle cost (Kosonen et al., 2000). The use of water-based heat transfer gives the system an inherent efficiency boost when compared with conventional fan-powered forced-air systems. With these advantages, active chilled beam systems are widely used in various building types in Europe and the U.S.

Active chilled beam systems are not a completely new technology but rather a proper combination of the two existing technologies: induction units and radiant cooling panels. The concept of the active chilled beam technology can date back to the late 1930s when Willis

Carrier introduced high-pressure perimeter induction units (Int-Hout & Wilbar, 2014). It was an advancement to use water that provides more energy capacity as a method of cooling instead of cooled air. It also showed an air-water terminal can be located and used in the space. This technology became a stepping stone to the chilled beam technologies. In the 1970s, engineers in

Scandinavia adjusted this technology to radiant cooling panels to take advantage of natural ventilation, which became a starting point of passive chilled beams. Mechanical ventilation was also integrated into chilled beams instead of natural ventilation so that it could be used in the climate zone where the natural ventilation could not be effective. Active chilled beams were the outcome of this effort (Titus, 2013).

Figure 1.1 shows the illustration of two types of chilled beams: passive and active chilled beams. Both passive and active chilled beams are zonal systems. Passive chilled beams consist of a heat exchange coil where chilled water circulates. Active chilled beams contain ventilation air supply duct and air nozzle in addition to a heat exchange coil. Passive and active chilled beam systems operate by inducing room air to a heat exchanger coil, reducing the temperature of the induced room air. Passive chilled beams recirculate this induced air into a space. Active chilled beams mix the induced air with air supplied by a ventilation system and then supply the mixed air into a space.

Figure 1.1 Illustration of passive and active chilled systems [https://www.archtoolbox.com/materials-systems/hvac/chilled-beam-ceiling.html]

Figure 1.2 Illustration of active chilled beams In the case of passive chilled beams, the room air is induced to the coil through the natural convection of warm room air. There is no fan pushing air to passive chilled beams, and ventilation air can be provided through other means.

In the case of active chilled beams, the chilled beam unit is connected to primary air

(supply air) stream as opposed to passive chilled beams. Active chilled beams mix the induced room air with the primary air and then supply to a space. Active chilled beams are capable of meeting the ventilation load. Taking the primary air stream can be beneficial because t primary air stream can provide additional sensible cooling capacity to the room. Because of these advantages, active chilled beams are more widely used compared with passive chilled beams.

The minimum spacing for chilled beams to work properly is six feet for passive chilled beams and ten feet for active chilled beams.

As a coil surface of active chilled beams are exposed to a space, it is important to prevent condensation on the coil surface. It is important to ensure that the chilled water temperature is

3 above the dew point temperature of induced air (i.e. room air) to prevent condensation. A proper active chilled beam design will often use a tertiary chilled water loop to provide chilled water to beams at a temperature range of 50 - 60 °F. Chilled beam systems are recommended to operate with the relative humidity of a space at 55% or less (at 72°F dry bulb temperature). Additional dehumidification can be used to adjust the dew point temperature, which may reduce the energy savings due to the additional system use in lowing the humidity. Hence, active chilled beams appear to achieve maximum energy savings in dryer climates compared to conventional air systems.

Chilled beam systems are estimated to have an energy reduction of 20-50% compared with all-air conditioning systems, depending on the types of chilled beams, building, and the climate (Butler et al., 2004). Specifically, chilled beam systems save fan energy and pump power over all-air systems. Passive chilled beams can supply about six to seven watts per square foot while active chilled beams are normally twice as effective. The estimated annual energy cost per square foot for HVAC systems is $0.90 for active chilled beams, $0.74 for passive chilled beams, and $1.50 for variable air volume (VAV) systems. (Butler et al., 2004)

Active chilled beams are more popular than passive chilled beams because of their benefit of providing the ventilation and the additional sensible cooling capacity to the zone.

According to the Building Services Research and Information Association (BSRIA, UK) website, “Active beams remain the dominant product in the European beams market, with the exception of the Swiss market, where both passive and active chilled beams are not very popular.”1

1 https://www.bsria.com/uk/news/article/chilled-ceilings-and-beams-continued-growth-in-the-european-market/ 4

Figure 1.3 Chilled beams market share by types in Europe, 2010 Overall, active chilled beams have 89% of market share (including the multi-service active chilled beams), and passive chilled beams have 11% of market share in Europe, 2010.2

Several building energy simulation tools provide models or models that can simulate active chilled beam systems. However, the accuracies of these models are rarely evaluated nor presented. Evaluations using comparisons between simulations and physical experiments are particularly limited. Active chilled beam system modeling requires all of the aforementioned items to be modeled appropriately. An active chilled beam model should be able to predict the impact of increasing the flow rates of the primary air or the hydronic flow rates through the coil.

This requires the active chilled beam models to include the impact of flow rates and temperatures for the primary air, the secondary air, and hydronic connections to an active chilled beam on the overall capacity and leaving air conditions. Proper control options should be included so as to ensure the condensation prevention on beam surfaces.

2 https://www.bsria.com/uk/, “Chilled beams and ceilings” 5

Most existing building energy simulation software lack capability of modeling chilled beams, and they, if ever, cannot adequately represent the current configurations that are used in industry. EnergyPlus currently uses the CooledBeam model originally implemented in DOE2.1E

(LBNL, 2000; United States Department of Energy, 2018). This model was developed by Halton

Oy in the late 1990s using a semi-empirical formulation. It utilizes outdated empirical coefficients that do not realistically represent the current technology available on the market.

Furthermore, it is difficult to produce new coefficients from manufacturer provided performance data.

There is not much current consensus on modeling methods for active chilled beams among the existing software tools. The current modeling methods being employed contain a host of modeling deficiencies, including the inability to model VAV chilled beams, variable induction ratios, and humidity control.

Moreover, the ability of existing energy modeling tools to connect active chilled beams to air handling units are limited. The tools do not provide the ability to model a tertiary loop that is modulated via a control valve per standard design procedures. A chilled water reset control may need to be allowed, which is widely used to prevent condensation at chilled beam coils. In this case, chilled water temperature needs to be reset based on the zone dew point temperature as opposed to the typical chilled water temperature reset control case where the water temperature resets based on the ambient temperature.

1.2 Significance of Research

In recent years, active chilled beam systems are gaining popularity. Most of the existing experimental studies regarding active chilled beam systems pertain to the measurement of the

6 airflow distribution inside zones with active chilled beams, or the analysis of specific airflow patterns caused by active chilled beams. Some studies analyze the relationship between the heat distributions in the zones of interest to the airflow patterns caused by active chilled beams. Also, several authors expanded on this relationship by studying the zone occupants’ thermal comfort with respect to the changed airflow patterns. Several papers that focus on these airflow patterns experimented and validated the airflow patterns using Computational Fluid Dynamics (CFD) simulations. However, fewer studies have examined the energy and thermal performance of active chilled beams under various operating conditions. Particularly, there is no reported experimental study that attempts to show the impact of supply air flow rate and supply water temperature on energy performance of actual chilled beam products. Also, studies on the validation of the performance of active chilled beams between experiments and commonly used energy simulation tools are barely found. Most energy modeling tools are limited in their ability to model active chilled beam systems accurately as designed.

Therefore, this study examines the energy and thermal performance of the representative active chilled beam products under various operating conditions. This study explores an experimental method to measure the cooling capacity of each product and examines the factors that potentially affect the determination of the cooling capacity of active chilled beams. This research will provide first-hand knowledge of thermal and energy performance of selected active chilled beam systems considered in both a full-scale laboratory and a computer-simulation under given test conditions. The findings from experiments and simulations have a potential to improve the foundation and quality of the current design and experiment guidelines for active chilled beams. To-be-developed active chilled beams model will be able to calculate the performance of active chilled beams under various designs and operating conditions more realistically.

1.3 Research Objectives

The overall objective of the research is to evaluate existing simulation models based on the experiment with the series of the active chilled beams and develop a simulation model that can predict the energy performance of active chilled beam systems accurately under a wide range of system capacities and operating conditions.

1.4 Research Method

First, previous research studies were reviewed. In particular, the studies regarding the field experiment with active chilled beam systems, and existing simulation tools implementing a model for active chilled beam systems were reviewed.

Under the guidance of the ASHRAE Standard 200 and EN 15116, a detailed experimental plan was developed. Air-conditioning, Heating, and Refrigeration Institute (AHRI) Standard 1240 was also referenced. Active chilled beam units with six different specifications were used in the experiments. The selected active chilled beam units represent the typical two-sided discharge active chilled beams used in practice. The performance of the six active chilled beam units was evaluated with various primary airflow rates and water inlet temperatures. Every parameter in all experiment cases fulfilled the requirements of ASHRAE Standard 200 and European standard EN

15116.

After the experiments, the performances of prevalent simulation tools, EnergyPlus,

TRNSYS, and eQUEST, were evaluated and compared to the experiment results.

Based on the findings from the experiments and simulations analysis, the to-be-developed models will be formulated with mathematical and physics ground with only inevitable and easy-

8 to-measure inputs. It is anticipated that the new models can be integrated with prevalent building energy simulation tools.

Chapter 2 Literature Review

2.1 Fundamentals of Active Chilled Beam

Total cooling capacity of an active chilled beam is a combination of the air-side cooling capacity and the water-side cooling capacity. The air-side cooling capacity is provided by primary air, and the water-side cooling capacity means the cooling capacity of water cooling coil

(ASHRAE, 2013).

푃푡표푡푎푙 = 푃푎𝑖푟−푠𝑖푑푒 + 푃푤푎푡푒푟−푠𝑖푑푒 (2.1)

The following equations is used to calculate the air-side cooling capacity.

푃푎𝑖푟−푠𝑖푑푒 = 푚̇ 푝푎 ∗ 퐶푝,푎 ∗ (푇푟푎 − 푇푝푎) (2.2) Where

푚̇ 푝푎 = mass flow rate of primary air

퐶푝,푎 = specific heat of air

푇푟푎 = average induced air temperature (room temperature)

푇푝푎 = primary air temperature

The water-side cooling capacity, which is the active chilled beam coil cooling capacity, can be calculated by coil heat transfer under steady-state conditions assuming no condensation on the cooling coil surface.

푃푤푎푡푒푟−푠𝑖푑푒 = 푚̇ 푤 ∗ 퐶푝,푤 ∗ (푇푤,표푢푡 − 푇w,in) (2.3) Where

푚̇ 푤 = mass flow rate of water (liquid media)

퐶푝,푤 = specific heat of water

푇푤,표푢푡 = water temperature exiting cooling coil

푇푤,𝑖푛 = water temperature entering cooling coil

The following equations also describes the water-cooling capacity under steady-state conditions assuming no condensation on the cooling coil surface.

푃푤푎푡푒푟−푠𝑖푑푒 = 퐾 ∗ 퐴 ∗ ∆푇 (2.4)

푃푤푎푡푒푟−푠𝑖푑푒 = 푚̇ 𝑖푎 ∗ 퐶푝,푎 ∗ (푇푎,표푢푡 − 푇푎,𝑖푛) (2.5) Where

퐾 = coil heat transfer coefficient

퐴 = coil surface area

∆푇 = temperature differences between induced air and mean water temperature

푚̇ 𝑖푎 = mass flow rate of induced air

퐶푝,푎 = specific heat of air

푇푎,표푢푡 = induced air temperature exiting cooling coil

푇푎,𝑖푛 = induced air temperature entering cooling coil

Since it is not easy to measure the cooling coil’s surface area (2.4), temperature of air around the coil (2.4), mass flow rate of induced air (2.5) and induced air temperature exiting cooling coil (2.5), it is common to use the equation (2.3) to estimate the water-side cooling capacity of an active chilled beam design practice instead of using (2.4) or (2.5) (ASHRAE,

2013).

2.2 Overall Literature Review

Active chilled beam had various names in its early stages – it was called “ventilated cooled beam”, “ventilated beam”, “active cooled beam”, “ceiling cooling convector”, and

“chilled beam cooling convector.” It was not until 2007 that it became more popularly known as

“Active chilled beam.”

Many studies on chilled ceiling were done in the early 1990s, but most of them were focused on the radiant cooling panel or a combination of displacement ventilation and chilled ceiling rather than active/passive chilled beams. The active/passive chilled beams were only introduced with their basic concepts or theoretical backgrounds, or they were used to be introduced as a good combination with displacement ventilation. Bunn and Wyatt introduced

"the chill beam" in 1991 as a next generation of the radiant cooling panel that overcomes its drawbacks (Bunn & Wyatt, 1991). Based on their laboratory tests and CFD modeling, the study explained the combination of chilled beams and displacement ventilation can be the alternative option for conventional air conditioning system because chilled beam can enhance the buoyance effect produced by the displacement process.

In the 2000s, active chilled beam systems are becoming more popular in the U.S. and

Europe in recent years. The gaining popularity has led to plenty of research including experimental studies. For actual and reliable testing with active chilled beams, Europe has a standard related to active chilled beams. EN 15116 Ventilation in buildings – Chilled Beams –

Testing and Rating of Active Chilled Beams are used as a standard for active chilled beam testing by the manufacturers in Europe. Europe also has REHVA – Chilled Beam Application

Guidebook is a chilled beam guidance that is widely used. In the U.S., ASHRAE Standard 200 –

Methods of Testing Chilled Beams and AHRI Standard 1240 - Performance Rating of Active

Chilled Beams are commonly used. Besides, several manufacturers such as Halton Oy, Price industries, Titus, and TROX provide detailed information about active chilled beams including the design consideration, system design strategies, and control logic.

Most of the existing experimental studies regarding active chilled beam systems pertain to measuring either the airflow distribution inside of the zone with active chilled beams, or specific airflow patterns caused by active chilled beams. Some studies expanded the airflow pattern caused by active chilled beams to the heat distributions in the zone and the occupants’ thermal comfort. Several papers that focus on those perspectives experimented and validated the airflow patterns by CFD simulation. To figure out the influence of heat sources in the zone with active chilled beams, field experiments were done with several combinations of heat sources and active chilled beams arrangement.

Zbořil et al. (Zbořil et al., 2006) investigated the impact of heat sources on the air flow pattern and thermal comfort in the zone ventilated with active chilled beams by conduction their experiments. They concluded that concentration of heat sources can make the air flow turned and cause local discomfort in the other side of the zone. As the heat load increased, the draught risk and the temperature differences between the middle and the low height of the room increased.

Also, the placement of active chilled beams has impact on thermal comfort, particularly the arrangement of active chilled beams that can provide lower induction can lower the draught risk.

Melikov et al. (Melikov et al., 2007) studied the influence of heat loads and airflow patterns to occupants’ thermal comfort by a full-scale test and questionnaires. Draught discomfort was defined from the experiment subjects’ responses – the subject felt body part cooler than normal or local cooling/air movement uncomfortable. They concluded that the airflow interaction in the zone is the major reason for occupants’ draught discomfort. The higher heat loads in the zone,

13 the higher draught discomfort for the occupants. For reducing the draught discomfort, induction control can be an efficient way to decrease the discomfort. Cao et al. (Cao et al., 2008) measured and predicted air distribution in a typical office room with active chilled beams. The author developed a free convection model that superposes a free convection and an isothermal jet velocity to analyze a downward air velocity along the wall in a typical office space. The calculated the maximum air velocities at the wall height of 0.6 to 2.0 meter was well matched with the experiment results. However, the model was not capable of calculating the velocity at the corner of the wall. Koskela et al. (Koskela et al., 2010) studied about air flow patterns including mean air speed in laboratory condition. Koskela found that the heat sources have a noticeable impact on the air flow pattern in the laboratory by causing large scale of circulation and interrupting the direction of air inlet jets. Upon this research, Koskela et al. (Koskela et al.,

2012) studied air flow patterns and draft risk for the active chilled beam systems in a typical office environment. CFD results were compared with the measurement in a laboratory mock-up room. The authors found that air flow pattern and velocity was greatly influenced near heat sources or surfaces having surface temperatures notably different from the air. Also, they reported that a large-scale circulation can result in high air velocity at the floor level increasing draft risk.

There are some other studies that conducted experiments for analyzing the performance of the active chilled beams. Le Dréau et al. (Le Dréau et al., 2015) conducted a full-scale experiment to analyze the energy performance of an active chilled beam under steady-state and dynamic conditions. The authors mainly focused on comparing energy consumptions of an active chilled beam with radiant wall cooling system rather than the other active chilled beam system.

Chen et al. (Chen, Cai, Wang, et al., 2014) evaluated the impact of various circuitry

14 arrangements on the thermal and hydraulic characteristics of the water-side of an active chilled beam system. The authors compared non-dimensional parameters, heat exchanger effectiveness, and performance index, for the analysis. It was found that the 2-circuits arrangement performed better than the 1-circuit arrangement that is widely used in the active chilled beam applications.

Recently, Filipsson et al. (Filipsson et al., 2016) proposed a thermal model of an active chilled beam. The model integrated NTU method and a hybrid of first principles and empirical data. The model was capable of capturing buoyancy forces of induced air. The unknown characteristic constants in the equation of the convective heat transfer on the air-side and induction ratio were determined from measurements in a full-scale mock-up of an office room. The authors reported that the calculated cooling capacity, supply air temperature, and flow rate agreed well with measured data.

During the experiment process, the induction ratio plays an important role in the calculation of cooling capacity of active chilled beams. Therefore, several papers focused on calculating the accurate induction ratio. Ruponen and Tinker (2009) proposed a simplified venturi method to measure the air entrainment ratio of an active chilled beam (Ruponen &

Tinker, 2009). The authors placed additional rectangular venturi along the edge of the face plate of the active chilled beam to improve the reliability of secondary air volume flow rate measurement. The measurement obtained from a simplified method was compared to results obtained according to European Standard EN 15116. The measured air entrainment ratio obtained from a simplified method was different by up to 40% compared with the measurement according to EN 15116. The authors argued that more detailed measurements of the velocity profile across an irregular coil’s face would be needed to improve the accuracy of the air entrainment ratio measurement. Meanwhile, while many studies focus on the effect of primary

15 air flow on the induction ratio calculation, Filipsson et al. (2016) investigated other variables besides primary air flow that could possibly affect the induction ratio calculation through a full- scale experiment (Filipsson et al., 2016). They reported that buoyant forces created by the chilled water temperature can considerably influence the induction ratio.

Studies about the energy simulation mostly compared active chilled beam unit system to other systems. variable air volume (VAV) system is the most popular system for comparison.

However, it is hard to compare each system to the active chilled beam system at once since different authors mentioned different systems in the different studies. From their results, some papers supported the application of active chilled beams but others not – the advantages and disadvantages of active chilled beams were focused differently case by case. Stein and Taylor compared energy efficiency, and the initial cost of dedicated outdoor air systems (DOAS) plus active chilled beam systems versus the variable air volume reheat (VAVR) system installed in

UC Davis Medical Center Graduate Studies Building (Stein & Taylor, 2013). Reportedly, VAVR had much lower initial costs, and energy costs compared to DOAS with the active chilled beam system. Reducing primary airflow rate and circulating medium temperature chilled water plus a water-side economizer might enhance the energy performance of DOAS with the active chilled beam system. However, the design may not be applicable in the extreme climate zone. Also,

Setty discussed potential limitations of the active chilled beam systems (Setty, 2011). Setty simulated active chilled beam systems in a typical five-zone-office space and reported that the active chilled beam systems consume more energy compared with a VAV system under certain conditions. A major concern for the active chilled beam systems is that the system is vulnerable to high latent heat gains. Active chilled beam systems may not be capable of handling the latent load due to high-occupancy or high infiltration rates. Condensation may occur at the cooling coil

16 surfaces in an active chilled beam if the system is not properly operated. The author also pointed out that there is a lack of standard programs for rating the chilled beams. Wahed et al. also used

TRNSYS to analyze the performance of the active chilled beam systems integrated with the thermally regenerated desiccant system in very hot and humid climate (Wahed et al., 2010). This integrated system was able to meet the sensible and latent cooling load of a typical office building in Singapore. The operating cost of the integrated system was reduced by 1.5 to 2.5 times compared to the conventional VAV system under very hot and humid climate.

Regarding the active chilled beam models in the simulation programs, Betz et. al summarized the differences between actual active chilled beam systems design and how they are implemented in various building energy simulation programs such as eQUEST, EnergyPlus,

IES-VE, Trane TRACE, and TRNSYS (Betz et al., 2012). The authors highlighted the limitations of each program and emphasized the need for more accurate, easy-to-use, and cost- effective active chilled beam calculation models. However, their research was based on testing the model against the basic concepts and principles and there were no practical experiments for validation. Other than this, very few papers did some comparisons of active chilled beam performance in different energy simulation tools. It is barely found any published study on the validation of the performance of active chilled beams between field experiment and commonly used energy simulation tools.

Meanwhile, there are some papers that developed their own modeling. Chen et al. developed a hybrid dynamic model of an active chilled beam terminal unit. The developed a dynamic model combined two models: the air jet model and the cooling coil model(Chen, Cai,

Giridharan, et al., 2014). The authors derived an empirical relationship to define the entrainment ratio as a function of the primary air plenum pressure. Least-Squares Method (LSM) was used to

17 estimate unknown parameters in the air jet model. For the cooling coil model, the authors used the simplified finite element analysis method by treating the whole cooling coil as the only control volume. Chen et al. also developed a simulation model based on the energy and mass conservation fundamentals for analyzing the operating characteristics of the variable air volume active chilled beam systems with a perspective that the sensible heat ratio (SHR) of the system should match the SHR of the space (Chen et al., 2015). The correlation between the total cooling output capacity of the active chilled beam system is then carried out. The authors reported that dehumidification capability of the active chilled beam systems can be enhanced by integrating any liquid desiccant dehumidifier with the active chilled beam systems. Maccarini et al. developed an empirical model of an active chilled beam for building energy simulation by using

Modelica (Maccarini et al., 2015). The authors compared the water outlet temperature predicted by the developed empirical model with temperature measurement from a full-scale experiment in order to evaluate conformity of the model. The empirical equation for the heat transfer coefficient being used in coil heat transfer model was derived based on the performance data sheet of the particular active chilled beam unit. The heat transfer coefficient was a function of the primary air mass flow rate, the water mass flow rate, and water inlet/outlet temperature, length and area of the active chilled beam unit. Condensation on the coil surface was ignored. Most recently, Filipsson et al. proposed a thermal model of an active chilled beam (Filipsson et al.,

2017). The model integrated NTU method for heat exchanger and a hybrid of first principles and empirical data. The model was capable of capturing buoyancy forces of the induced air. The unknown characteristic constants in the equation of the convective heat transfer on the air-side and induction ratio were determined from the measurements in a full-scale mock-up test of an

18 office room. The authors reported that the calculated cooling capacity, the primary air temperature, and flow rate agreed well with measured data.

Although there have been numerous experimental studies on active chilled beams, studies examining the energy and thermal performance of active chilled beams under various parameters are limited. Particularly, there are no papers that discuss the experimental method for figuring out the energy performance of existing active chilled beam products, and the impact of several parameters – such as the supply air flow rate and supply water temperature – on the energy performance of an active chilled beam. Also, there are no studies validating the energy simulation predictions using actual practical measurements.

2.3 Review on Simulation Models

Eight energy simulation tools that are capable of modeling active chilled beam systems were reviewed: EnergyPlus, DOE-2, IDA-ICE, eQUEST, EE4, Carrier HAP, Trane Trace 700, and TRNSYS. There are some similarities among these simulation tools – the whole system is composed of two parts; the central supply air system for the primary air and the active chilled beam (or induction unit) terminal. The primary air deals with the ventilation requirement and dehumidification. The active chilled beam coils deal with another load.

After detailed exploration of these simulation tools, the tools were classified into three categories: empirical-based model, heat balance model, and induction unit – based on how they treat the active chilled beam models in the whole HVAC system. The detailed classifications are as follows:

- Empirical-based model: EnergyPlus, DOE-2, IDA-ICE

- Heat balance model: TRNSYS

- Induction unit model: eQUEST, EE4, Carrier HAP, Trane Trace 700

Empirical-based models are based on the energy equations provided by specific manufacturers to calculate the cooling and heating capacity of active chilled beams. The equations are empirically determined. The inputs regarding active chilled beam units need to be provided by the manufacturer. EnergyPlus and DOE-2 use basically the same equations for the cooling capacity calculation. For those tools, Halton Oy provided the model source for these simulation tools. In the dynamic simulation process, energy balance (room load satisfied by primary air flow load plus cooling capacity provided by active chilled beams) is achieved by iteratively solving relative parameters in the energy equations. In IDA-ICE, there are two options: the manufacturer’s option and the simplified option. In the manufacturer’s option, the capacity of active chilled beams is calculated based on the convective heat transfer between the air and the cooling coil, given specific coefficients from manufacturers. While in the simplified option, the heating and cooling capacity should be given manually.

Heat balance models calculate the capacity of active chilled beam systems based on the heat transfer process around the cooling/heating coil in active chilled beams. In case of

TRNSYS, the active chilled beam model was developed in 2007. TRNSYS specifically introduces a bypass ratio in their active chilled beam model to describe the characteristic of airflow of active chilled beams and thus has a little difference in calculating the heating and cooling capacity.

Except for the four simulation tools mentioned above, all other tools regard an active chilled beam system as an induction unit with airflow inside the room by known induction ratios.

The workflows of these models are basically the same to each other: the primary (supply) air is connected to a central air system (mostly models define it as DOAS); the primary (supply) air

20 induces the room air by a certain induction ratio that forms a secondary air loop inside the rooms; a water loop inside an active chilled beam coil is connected to a central water loop or secondary water loop of central water system. The differences among the tools in this category are different capabilities in active chilled beam models: whether or not controlling zone level or system level humidity, constant or variable control of an induction ratio at zone level or system level, and whether or not supporting variable supply air in the primary air system.

Three representative simulation programs were selected from each category based on versatility – EnergyPlus from the empirical-based model, TRNSYS from the heat balance model and eQUEST from the induction unit model.

Chapter 3 Full-Scale Laboratory Experiment Analysis

3.1 Experiment Objective

The experiments were carried out to evaluate the performance of various active chilled beam products provided by manufacturers. The water-side cooling capacity and the air-side cooling capacity of active chilled beam units were measured under various conditions. The total capacities of each active chilled beam unit were calculated according to ASHRAE Standard 200 and EN 15116. The total capacity of chilled beam units is defined as follows:

퐴퐶퐵 푤푎푡푒푟 푠𝑖푑푒 푐표표푙𝑖푛푔 푐푎푝푎푐𝑖푡푦 + 퐴퐶퐵 푎𝑖푟 푠𝑖푑푒 푐표표푙𝑖푛푔 푐푎푝푎푐𝑖푡푦 = 푡표푡푎푙 ℎ푒푎푡 푔푎𝑖푛

The water side cooling capacity was measured by the temperature difference between the entering water and the leaving water that circulate through a cooling coil inside a chilled beam unit. The air side cooling capacity was measured by the temperature difference between the entering air and the leaving air. In this experimental study, heat blankets were installed in the testing chamber to regulate internal heat gains consistently.

These measured data were then used to evaluate the accuracy of existing simulation programs.

(ASHRAE, 2013; CEN, 2008).

3.2 Experimental Plan Development

A detailed experimental plan was developed under the guidance of ASHRAE Standard

200 and EN 15116. The plan throughout the experimental study was discussed in the following subsections including testing chamber selection, active chilled beam unit selection, testing case determination, and environmental variable setting.

3.2.1 Experiment Location: Larson Building System Laboratory

The experimental study was carried out in the Larson Building System Laboratory, a full- scale building system laboratory. Figure 3.1 shows the 3-D modeling view of the Larson

Building System Laboratory. The Larson Building System Laboratory is located in Department of Civil, Environmental and Architectural Engineering at the University of Colorado Boulder.

This laboratory permits the study of entire building systems with full-scale test rooms in a controlled dynamic environment, providing repeatable test condition. It is designed for dynamic testing of complete and full-scale HVAC systems. The facility consists of three full-size HVAC systems, two representative thermal zones, systems for producing controllable thermal loads on the HVAC system, data acquisition systems, and control systems.

Activities at the laboratory include evaluation and testing of control algorithms and hardware for HVAC components and systems, interactions between multiple control functions of

HVAC systems, the dynamic interactions between building thermal response and HVAC system controls, ventilation control for indoor air quality, and HVAC system diagnostics. The most recent ASHRAE experimental projects conducted in this lab include: (1) Experimental

Investigation of Hospital Operating Room Air Distribution (ASHRAE Project, 1397-TRP) (PI:

Dr. Zhai); (2) The Development of Simplified Rack Boundary Conditions for Numerical Data

Center Models (ASHRAE Project, 1487-TRP) (PI: Dr. Zhai); (3) Role of HVAC Systems in the

Transmission of Infectious Agents in Buildings and Intermodal Transportation (ASHRAE

Project, 1603-TRP) (Co-PI: Dr. Zhai).

The laboratory was designed for maximum flexibility to encourage a wide variety of research and testing programs. In particular, both the HVAC and control systems in the laboratory are re-configurable in that components, subsystems or entire systems installed, tested, and

23 modified. Both the water systems and air systems were connected to the test chamber to be used in this experiment. The entire laboratory was well instrumented with most of the testing equipment required for the proposed investigations.

Figure 3.1 3-D view of the full-scale HVAC laboratory at the University of Colorado Boulder

One of the test chambers was selected and remodeled for this research – the East Zone. The chamber is 20 feet long, 20 feet wide and 12 feet high. There is a 2 feet high plenum above the ceiling where a chilled beam unit can be mounted and connected with a ductwork of an air handling unit. The roof and walls of the test chamber is insulated with R-50 insulation. The floor is insulated with R-30 insulation. Each surface is sealed with films to minimize the infiltration through the envelope. Figure 3.2 shows inside of the East Zone at the beginning of the experiment.

Figure 3.2 Photos of the test chamber at the beginning

3.2.2 Selecting Active Chilled Beam Units

At the experimental planning stage, we were consulted by three manufacturers to find active chilled beam products suitable for the experimental study: Halton Oy, Price Industries, and Swegon. Active chilled beam products were selected depending on whether they met the three criteria we considered: (a) a product should be two-sided discharge active chilled beams,

(b) a product should be commonly used in the current market, and (c) a product should be able to operate under various test conditions considered. Three types of active chilled beams manufactured by Halton Oy and Price Industries were selected, as listed below.

Halton - ASB/B-R-48-42-N - ASB/D-R-48-42-N - ASB/E-R-48-42-N

Price Industries - ACBL-HE/48"L Size 20 nozzle - ACBL-HE/48"L Size 30 nozzle - ACBL-HE/48"L Size 40 nozzle

The vendors adjusted the system specifications to ensure each active chilled beam operates properly in high elevation area of Boulder, Colorado. Table 3.1 shows detailed specifications about the selected products.

Table 3.1 Selected active chilled beam units (Manufacturer And model names are blinded)

Air Chilled Beam System Manufacturer Manufacturer A Manufacturer B Active Chilled Beam Product Product Product Product Product Product Model AL AM AH BL BM BH Active Chilled Beam Specifications Design Chilled Water Flow 1.5 gpm 1.5 gpm 1.5 gpm 1.5 gpm 1.5 gpm 1.5 gpm flow (gpm) Rate 55.4 - Inlet: 57.2F Inlet: 57.2F Inlet: 57F Inlet: 57F Inlet: 57F CHW 62.6 F Inlet: 57.2F Outlet: Outlet: Outlet: Outlet: Outlet: (13-17 Outlet: 67F Temperature 64.1F 67.5F 59.3F/60.7F 60.3F/61.5F 60.8F/61.7F C) 5-20 Primary Air cfm/ft 20 -31 CFM 32-65 CFM 66-80 CFM 20-40 CFM 40-60 CFM 60-80 CFM Flow Rate (7.7-31 L/s-m ) CHW Pipe 1/2" NPT 1/2" NPT 1/2" NPT 1/2" NPT 1/2" NPT 1/2" NPT Dimension 5" diameter 5" diameter 5" diameter 5" diameter 5" diameter 5" diameter Supply Duct duct duct duct duct duct duct Dimension connection connection connection connection connection connection Total Total Total Supply Air Pressure .22” WG .45” WG .43” WG Pressure Pressure Drop .34" - .70” WG to .99” WG to .74” WG Static Pressure Drop .8" WC Drop .5" WC WC Cooling Coil 1,648-2,232 2,034-2,833 2,157-3,031 1,696-2,772 2,449-3,372 2,887-3,504 Capacity Btu/hr Btu/hr Btu/hr Btu/hr Btu/hr Btu/hr (Water-side) Coil Length 42" – (3.5') 42" – (3.5') 42" – (3.5') 48" – (4') 48" – (4') 48" – (4') Beam Length 48" – (4') 48" – (4') 48" – (4') 48" – (4') 48" – (4') 48" – (4') Total Width 24" – (2') 24" – (2') 24" – (2') 24" – (2') 24" – (2') 24" – (2') Total Height 7" 7" 7" 10" 10" 10"

3.2.3 Experiment Conditions

Experiment conditions were expected to be summer conditions. The details are shown in the Table 3.2.

Table 3.2 Experiment conditions

Test Parameter Range Room Temperature 67°F - 75 °F (19.4°C - 24°C) Chilled Water Supply Temperature 55°F / 58°F / 61°F (12.8°C / 14.4°C / 16.1°C) Chilled Water Flow Rate 1.0 GPM (0.063 L/s) Primary Air Temperature 64.4 °F (18 °C) Primary Air Flow Rate 20 - 80 cfm (9.44 – 37.76 L/s )

Note that the designed water flow rate of each active chilled beam unit is 1.5 gpm.

According to the AHRI test standard 1240, water flow rate shall be greater than or equal to 95% of the rating. (AHRI, 2017) However, it was figured out the water pump only performed up to

1.2 gpm while setting up the experiment chamber. The ASHRAE committees and the two manufactures confirmed that the water flow rate of 1 gpm still suitable for the experiment.

3.2.4 Experiment Cases

The specifications of all chilled beam units were carefully reviewed to design experiment cases. After reviewing specifications of all selected chilled beam units, the range of primary air flow was divided into three categories: Low (20 – 40 cfm), Middle (40 – 60 cfm), and High (60 –

80 cfm). For each category, each manufacturer’s active chilled beam unit were classified by the range of design primary air flow. Each category was then divided into nine cases by different combinations of chilled water supply temperature and primary air flow rate. 27 experiment cases were determined for each manufacturer, as shown in Table 3.3. Thus, 54 experiment cases in total were designed for this experimental study.

Table 3.3 Experiment cases details

Fixed Value Parameters Chilled Water Flow Rate 1.0 GPM (0.063 L/s) Primary Air Temperature 64.4 °F (18 °C) LOW Primary Air Flow Range (20 – 40 cfm) Product AL Chilled Water Supply Temperature Product BL 55 °F (12.8 °C) 58 °F (14.4 °C) 61 °F (16.1 °C) 20 CFM (9.4 L/s) Case 1 Case 2 Case 3 Primary Air 30 CFM (14.2 L/s) Case 4 Case 5 Case 6 Flow Rate 40 CFM (18.9 L/s) Case 7 Case 8 Case 9 MIDDLE Primary Air Flow Range (40 – 60 cfm) Product AM Chilled Water Supply Temperature Product BM 55 °F (12.8 °C) 58 °F (14.4 °C) 61 °F (16.1 °C) 40 CFM (18.9 L/s) Case 1 Case 2 Case 3 Primary Air 50 CFM (23.6 L/s) Case 4 Case 5 Case 6 Flow Rate 60 CFM (28.3 L/s) Case 7 Case 8 Case 9 HIGH Primary Air Flow Range (60 – 80 cfm) Product AH Chilled Water Supply Temperature Product BH 55 °F (12.8 °C) 58 °F (14.4 °C) 61 °F (16.1 °C) 60 CFM (28.3 L/s) Case 1 Case 2 Case 3 Primary Air 70 CFM (33.0 L/s) Case 4 Case 5 Case 6 Flow Rate 80 CFM (37.8 L/s) Case 7 Case 8 Case 9

3.2.5 Parameters to Measure

Parameters that need to be measured were decided based on the requirements of the

ASHRAE standard and the European testing standard. In the actual experiment, additional sensors were placed to obtain supplemental information that may be useful for the validation of energy simulation tools. Table 3.4 shows the selected parameters. These parameters can be classified into two types; experimental and environmental parameters. The experimental parameters are related to the performance of active chilled beams. The environmental parameters control the experimental environment to ensure the result accurately and convincible.

Table 3.4 Experimental parameters

Number of measurement Measurement point location points Supply water temperature 1 Supply water pipe Water-side Return water temperature 1 Return water pipe Water flow rate 1 Supply water pipe Primary airflow rate 1 Primary air side Primary air temperature 1 Primary air side Air-side Average air temperature of induced air Reference air temperature 3 along the length of the active chilled beam’s induction slots Room Distributed evenly in a vertical profile Room temperature 3 condition cross the center of the room

Table 3.5 Environmental parameters

Number of measurement points Measurement point location Room surface temperature 4 Center of 4 wall surfaces Relative humidity 1 Center of the room Readings from the power Internal heat gain 496 Watts (1692 Btu/h) meter of room heating blankets on the floor

3.2.6 Experiment Procedure

The procedure of each experimental case was as follows:

(1) Product setup: Installation/adjustment of the active chilled beam product on the ceiling

(2) Sensor setup: Instrument calibration and placement

(3) Internal load set up: Installation of heat blankets for generating uniform internal heat gain.

The test chamber was considered as an adiabatic condition because all surfaces enclose the space were highly insulated.

푇표푡푎푙 ℎ푒푎푡 푔푎𝑖푛 = 퐶표표푙𝑖푛푔 푐푎푝푎푐𝑖푡푦 푓푟표푚 푡ℎ푒 푝푟𝑖푚푎푟푦 푎𝑖푟 + 퐶표표푙𝑖푛푔 푐푎푝푎푐𝑖푡푦 푓푟표푚 푡ℎ푒 퐴퐶퐵 푤푎푡푒푟 푐표𝑖푙

= 푃푟𝑖푚푎푟푦 푎𝑖푟 푓푙표푤 푟푎푡푒 ∗ (𝑖푛푑푢푐푒푑 푎𝑖푟 푡푒푚푝푒푟푎푡푢푟푒 − 푝푟𝑖푚푎푟푦 푎𝑖푟 푡푒푚푝푒푟푎푡푢푟푒)

+ 푐ℎ𝑖푙푙푒푑 푤푎푡푒푟 푓푙표푤 푟푎푡푒 ∗ (푟푒푡푢푟푛 푤푎푡푒푟 푡푒푚푝푒푟푎푡푢푟푒 − 푠푢푝푝푙푦 푤푎푡푒푟 푡푒푚푝푒푟푎푡푢푟푒

In each case during the experiment, the return water temperature was expected to change accordingly.

(4) Run the system until it reaches steady-state, meeting steady-state criteria defined by

ASHRAE standard and the European testing standard. It took approximately 8 to 12 hours to reach the steady-state.

(5) Collect more than one hour of measured data for each case

3.3 Experimental Quality Requirements

The experimental requirements are clearly stated in two standards to ensure the quality of a series of experiments: the European standards EN 15116 Ventilation in Buildings – Chilled

Beams and Testing and Rating of Active Chilled Beam and the ASHRAE Standard 200 – Method of Testing Chilled Beams. Specifically, there are requirements for the test chamber, potential products and instruments in both standards. The requirements of both standards are very similar.

Each experiment was carefully monitored to ensure it meets the requirements.

3.3.1 Insulation Requirement

The ASHRAE Standard 200 requires that the perimeter of the chamber shall be insulated and have negligible heat flow through the surface – less than 0.13 Btu/h/ft2. The constructions of enclosing surfaces were investigated to examine the U-value of each surface. The heat transfer rates through all surfaces were then measured. The average heat transfer rate through all surfaces was 0.11 Btu/h/ft², which was less than 0.13 Btu/h/ft², meeting the ASHRAE requirement.

Table 3.6 Envelope construction of the test chamber

Specific Thermal Thickness Conductivity Density U-value Heat Resistance Exterior Roof [ft] [Btu/h.ft.F] [lb/ft³] [BTU/F.hr.ft²] [Btu/lb.F] [F.hr.ft²/BTU] 22-Gauge Sheet Metal 0.002 9.3948 501.12 0.120 0.000 5/8IN Gypsum 0.052 0.0924 49.00 0.198 0.563 3/4IN Tongue and 0.063 0.0751 40.58 0.406 0.839 Groove OSB R-19 Batt Insulation 0.521 0.0266 16.54 0.230 19.579 R-11 Batt Insulation 0.292 0.0266 16.54 0.230 10.977 5/8IN Gypsum 0.052 0.0924 49.00 0.198 0.563 Total R-Value 32.521 0.031

Specific Thermal Thickness Conductivity Density U-value Heat Resistance Interior Ceiling [ft] [Btu/h.ft.F] [lb/ft³] [BTU/F.hr.ft²] [Btu/lb.F] [F.hr.ft²/BTU] AcoustClgTile_05in 0.042 0.0329 18 0.32 1.277 Total R-Value 1.277 0.783

Specific Thermal Thickness Conductivity Density U-value Heat Resistance Raised Floor [ft] [Btu/h.ft.F] [lb/ft³] [BTU/F.hr.ft²] [Btu/lb.F] [F.hr.ft²/BTU] Access Floor Panel 0.134 0.1363 74 0.16 0.983 F05 Ceiling air space 0.491 1.022 resistance R-19 Batt Insulation 0.521 0.0266 16.54 0.230 19.579 R-19 Batt Insulation 0.521 0.0266 16.54 0.230 19.579 7/16IN Chipboard 0.036 0.1156 75.00 0.310 0.315 4IN PolyISO insulation 0.333 0.0116 2.00 0.220 28.846 Total R-Value 70.324 0.014

Specific Thermal Thickness Conductivity Density U-value Heat Resistance Wall [ft] [Btu/h.ft.F] [lb/ft³] [BTU/F.hr.ft²] [Btu/lb.F] [F.hr.ft²/BTU] 5/8IN Gypsum 0.052 0.0924 49 0.198 0.563 2 3/4IN PolyISO 0.229 0.0116 2 0.22 19.741 Insulation 24-Gauge Sheet Metal 0.002 9.3948 501.12 0.12 0.000 R-11 Batt Insulation 0.292 0.0266 16.54 0.23 10.977 5/8IN Gypsum 0.052 0.0924 49 0.198 0.563 Total R-Value 31.845 0.031

Table 3.7 Test chamber insulation requirement

U-value Delta T1 Heat flow ASHRAE Standard 200-2015 [Btu/F.hr.ft²] [F] [Btu/h/ft²] Wall 0.031 4.4 0.14 Floor 0.014 4.4 0.06 Roof 0.031 4.4 0.14 Average 0.11 Less than 0.13 Btu/h/ft² 1 o o : one experiment case – Ambient Temp: 74 F / Room Temp: 69.6 F

3.3.2 Infiltration Requirement

The chamber shall be air-tight sufficiently for minimizing the infiltration from the ambient air (or the exfiltration to the ambient air). The ASHRAE Standard 200 requires it shall not exceed 0.16 cfm/ft2 (0.81/s m2) at a pressure difference of 0.2 in.wc (50 Pa). The blower door test was conducted to affirm that it met the infiltration requirement (0.15 cfm/ft2 at a pressure difference of 50 Pa).

To meet the insulation requirement, the test chamber was sealed with the 3M Air and

Vapor Barrier 3015 Tape that was provided by the 3M company. All gaps in the space were covered with the 3M tape except surrounding area for replacement of future active chilled beam units.

Figure 3.3 Test chamber before/after applying 3M air and vapor barrier 3015 tape

Figure 3.4 Blower door test on test chamber

After the applying 3M Air and Vapor Barrier 3015 tape, the blower door test was conducted. The blower door test result is in Table 3.8. It was checked that the air-tightness of the test chamber meets the requirements of the EN 15116 and the ASHRAE Standard 200.

ASHRAE Standard 200-2015 requirement states that the infiltration shall be less than:

Table 3.8 Test chamber infiltration requirement

CFM50Pa CFMnatural(4Pa) Requirement 221.44 11.07 Test room 10.57

3.3.3 Instrumental Requirement

Both the EN 15116 and the ASHRAE Standard 200 require that the instruments shall meet some level of precision for the calibration of instruments, reading accuracy, and uncertainty requirement. Table 3.9 shows the sensor accuracy requirements of the ASHRAE Standard 200 for key variables.

Table 3.9 The sensor accuracy requirement of the ASHRAE Standard 200 for key parameters

Air temperature ± 0.2°F (± 0.1°C) Water temperature ± 0.2°F (± 0.1°C) Air flow meter ± 5% Water flow meter ± 0.5% Global temperature ± 0.4°F (± 0.2°C)

- Air temperature sensors:

The standard requires an accuracy of ± 0.2°F (± 0.1°C) for air temperature sensors.

Thermocouples cannot be used in this experiment since an accuracy of thermocouples are ±

0.4°F (± 0.2°C). Therefore, thermistors were used for the experiment to meet the accuracy requirement of ± 0.2°F (± 0.1°C).

- Water temperature sensors:

Since the water temperature is critical to decide water cooling capacity, the team used RTD sensor that has an accuracy of ± 0.05°C. It has much more stringent accuracy than the standard requirement, ± 0.1°C.

- Air flow sensor:

The airflow sensor with a reading accuracy of ± 3% was used. The standard requirement is ± 5%.

- Water flow sensor

The water flow sensor was selected with properly designed flow according to our experimental water flow that is 1 gpm. The water flow meter that has a reading accuracy of ± 0.5% was used.

The standard requirement is ± 0.5%.

- Surface temperature sensor:

The surface temperature sensors have a reading accuracy of ± 0.1°C, the same with the requirement.

All sensors were calibrated to meet the ASHRAE Standard 200-2015 requirements. For measuring the air temperatures, radiant shields were attached to the thermistors to minimize the effect of radiant heat transfer. The radiant shields for sensors met the requirement of ASHRAE

Standard 200-2015.

Figure 3.5 Radiant shielded sensors - surface/air temperature sensors Table 3.10 Calibration result of thermistors for air temperature

Reference Temp. Temperature (˚F) ΔT STDVE Satisfy? 63.77˚F (Avg. of 20 minutes) (˚F) (K) Room high 63.87 0.09 0.03 Yes Room mid 63.81 0.04 0.02 Yes Room low 63.74 -0.03 0.02 Yes Reference left 63.80 0.03 0.03 Yes Reference center 63.93 0.16 0.04 Yes Reference right 63.83 0.06 0.03 Yes Exhausted air 63.78 0.01 0.02 Yes Outdoor air 1 63.8 0.03 0.03 Yes Discharged air S 63.84 0.07 0.03 Yes Discharged air N 63.88 0.1 0.02 Yes Leaving air S coil 63.83 0.06 0.11 Yes Leaving air N coil 63.95 0.18 0.04 Yes Global temp. 63.76 -0.02 0.03 Yes Outdoor air 2 63.79 0.01 0.04 Yes Supplement 1 63.85 0.08 0.25 Yes Supplement 2 63.93 0.16 0.25 Yes

3.3.4 Steady-state Requirement

According to the EN 15116 standard, “Steady-state conditions are considered to have been obtained when the standard deviations of the recorded measurements (sampling rate

36 approximately 1 per minute) are within the following limits at least 60 minutes.” (CEN, 2008)

Table 3.11 shows steady-state requirements for each parameter in detail.

Table 3.11 Steady-state requirements

Standard deviations of the Measured parameters recorded measurements Reference temperature 0.05 K Interior room surface temperatures 0.5 K Mean water temperature 0.05 K Cooling water flow rate 1 % Primary air temperature 0.2 K Primary air flow rate 1.5 %

3.3.5 Uncertainty Requirement

The uncertainty of each test can be calculated according to the cumulative error law

(CEN, 2008). The uncertainty requirement is:

2 2 2 2 2 푟 = √푟1 + 푟2 + 푟3 + 푟4 + 푟5 푎푛푑 푟 ≤ ±6% (3.1)

Where r1 = uncertainty (%) of cooling water flow rate r2 = uncertainty (%) of cooling water temperature rise (water temperature difference between inlet and outlet) r3 = uncertainty (%) of a temperature difference between the reference temperature and mean cooling water temperature r4 = uncertainty (%) of primary air flow rate r5 = uncertainty (%) of test room configuration (internal/external heat sources)

The EN 15116 requires uncertainty r shall be less than ± 6%.

3.4 Laboratory Setup and Systems

3.4.1 Laboratory system overview

The DOAS delivers the cooled supply air at 64.4°F (18°C). The cooling to the blower coil unit is provided by a small packaged air cooled chiller.

Figure 3.6 Laboratory setup and systems

Figure 3.7 The air-side System

The chilled water to the active chilled beam was provided from a chilled water storage tank that is charged by an air-cooled electrical chiller. The chilled water storage tank was insulated with fiberglass and bubble foil insulation to minimize the heat loss.

Figure 3.8 The water-side system and the data acquisition system

Figure 3.9 Insulation of the chilled water tank - fiberglass and bubble foil insulations

Various experiment measurement sensors, as indicated in the experimental plan, were installed in the test chamber. To collect and record data from the sensors, the National

Instrument data acquisition system was used.

3.4.2 Active Chilled Beam

3.4.2.1 Active Chilled Beam Installation

Each active chilled beam unit was installed in the ceiling plenum, as shown in Figure

3.10 Installation of an active chilled beam. After proper installation of a chilled beam unit, the ceiling panels were mounted on the ceiling, separating the conditioned space from the plenum.

Figure 3.10 Installation of an active chilled beam

3.4.2.2 Physical Difference of Active Chilled Beams Between Manufacturer A and

Manufacturer B

While installing each active chilled beam unit, a few physical differences between the active chilled beam products from two manufacturers were found. This was important because

40 these differences required a different ceiling setup for the installation. The physical differences between the manufacturers are:

1) Primary air supply inlet placements

Each active chilled beam unit from both manufacturers has primary air inlet side placed

on opposite side. It means that different geometries of the primary air supply duct are

needed. Or, the geometries of the water pipes that are connected to the active chilled

beam coil need to be adjusted.

Figure 3.11 Primary air inlets – the opposite side

2) The positions of the deck for hanger rods

ACB units from both manufacturers have different positions of the deck for hanger rods.

It means different hangers with different lengths are needed to suspend the active chilled

beam units from the ceiling.

Figure 3.12 Different positions for hanger rod deck 3) Heat Exchange coil inside active chilled beam unit

After opening the hinged grille face, it was possible to observe the inside of active chilled

beams. Cooling coils in active chilled beam units of Manufacturer A are partially blocked

by thin metal plates. Cooling coils in active chilled beam units of Manufacturer B do not

have these metal plates.

Figure 3.13 Cooling coils in active chilled beam units – Manufacturer A (Left) and Manufacturer B (Right)

4) Nozzle numbers and nozzle spacing

It was already found that the nozzle diameters are different among each active chilled

beam. After opening the hinged grille face, it was able to count the number of the nozzle.

Table 3.12 shows the number of nozzles, nozzle spacing, and nozzle diameter of each

active chilled beam.

Table 3.12 Numbers of nozzles and nozzle diameters of each active chilled beam unit

AL AM AH BL BM BH Numbers of Nozzles on 72 72 72 32 38 44 Both Side Nozzle Spacing (in) 1 3/16 1 3/16 1 3/16 2 1/2 2 1/8 1 15/16 Nozzle Diameter (in) 0.134 0.232 0.295 0.25 0.25 0.3

Figure 3.14 Nozzles in an active chilled beam unit (after the hinged grille face was opened) 5) Density of cooling coil fins

The number of coil fins on the cooling coil was also able to be counted after opening the

hinged grille face. It was found that the fin numbers of each active chilled beam unit were

different based on the manufacturer.

Table 3.13 Numbers of fins on cooling coil of each active chilled beam units

AL AM AH BL BM BH Numbers of Fins on Cooling Coil 307 307 286 414 414 414

6) Shape of active chilled beam body frame

The shape of body frame of active chilled beam was different between Manufacturer A

and Manufacturer B. For Manufacturer A, the body frame does not surround the cooling

coil inside the active chilled beam unit. Active chilled beam units from Manufacturer B

have the body frame that surrounds the cooling coil. This can be the clue for estimating

the bypass fraction as described in 3.4.2.3.

3.4.2.3 Possibility of Bypass Fraction

Based on the frame of each active chilled beam unit, bypass fraction of induced air entering the coil can be assumed. For active chilled beams of Manufacturer B, the body frame of the active chilled beam tightly surrounds the side of cooling coil and blocks all the way except the grille-side, so that every induced air enters to the cooling coil after penetrating the grille face.

Therefore, it can be said that the coil air bypass fraction is 0 for BL, BM, and BH.

Figure 3.15 ACB Frame that surrounds the cooling coil inside – BL, BM, and BH Unlike Manufacturer B’s units, the frame of the active chilled beam from Manufacturer A does not surround the cooling coil inside. This brings the possibility of the coil air bypass fraction.

Figure 3.16 ACB Frame that surrounds the cooling coil inside – AL, AM, and AH

Based on the area of the potential path that the induced air after penetrating the grille face can enter, which is not inside of the cooling coil, the bypass fraction of each active chilled beam was calculated.

퐴푟푒푎 푡ℎ푎푡 𝑖푛푑푢푐푒푑 푎𝑖푟 푐푎푛 푝푒푛푒푡푟푎푡푒 푤𝑖푡ℎ표푢푡 푡표푢푐ℎ𝑖푛푔 푡ℎ푒 푐표𝑖푙 퐵푦푝푎푠푠 퐹푟푎푐푡𝑖표푛 = (3.2) 퐴푟푒푎 표푓 푔푟𝑖푙푙푒 푓푎푐푒

Table 3.14 Calculated Bypass Fraction of each active chilled beam unit

AL AM AH BL BM BH Calculated Bypass Fraction 0.04 0.04 0.055 0 0 0

3.4.3 Internal Heat Gain

For the uniformity, internal heat was simulated with the heat blankets on the floor rather than a centralized heat source – in the case of an electric heater, the heat source effect would be focused and it would affect the air temperature. The uniform internal heat would minimize any effect from another factor than key parameters.

According to the AHRI standard, the internal heat source is supplied to the test room evenly distributed through the walls and the floor. (AHRI, 2017) Twenty heat blankets were uniformly placed. The internal load of 496.6W was confirmed by using an electricity usage monitor iteratively. Each blanket provides six heat levels. In the experiment, level 2 heat settings for each electric blanket were used. According to the electrical specification of the blanket, each blanket shows 130 watts. If the heat blanket were set to have heat level 6, the total wattage of the

20 blankets will be 2600W. All lightings were shut down during the experiment. Therefore, the total internal load during the experiment is 496.6W.

Figure 3.17 Electrical blankets placement, heat level setting, and the electricity usage monitor

3.4.4 Sensor Measurement Location

All the sensors were placed based on the ASHRAE 200-2015 requirements. Figure 3.18 shows the locations of the sensors around an active chilled beam unit.

Figure 3.18 Sensor locations around the active chilled beam unit Surface temperatures were recorded on walls except the ceiling and the floor because of the convective effect from the active child beam unit to the ceiling, and the radiant effect from the heat blankets to the floor.

Figure 3.19 Sensor locations for surfaces/room conditions

3.4.5 Tests before the Experiment – Smoke Test

After installing the active chilled beam unit, smoke tests were conducted according to various primary air flow rates (20 cfm to 80 cfm) to see if there was any unexpected air stream. It was checked that there is no short-cut airstream and no bounced-back airstream on the north wall that is closer to the active chilled beam unit.

3.4.6 Tests before the Experiment – Stability Test for the Experiment Equipment

In order to make the experiment result reliable, the data acquisition system was rearranged. Electrical interferences among the voltage input channels were minimized so that stability requirements for reference temperatures were satisfied – which is 0.05K of standard deviation.

All the preparation for the experiment were finished. Figure 3.20 shows the overall view of the experiment chamber.

Figure 3.20 View from the chamber's door

Figure 3.21 Reference temperature sensors – less than 1 inch below the active chilled beam unit 49

Figure 3.22 Sensors around the active chilled beam unit

3.5 Experimental Result Analysis

All 54 experiment cases were finished after ensuring the quality of the testing instrument and meeting the standard requirements. The experiments of six active chilled beam units (54 cases) in the contract were conducted with a solid lab system established. Continuous measurements and improvements also ensured the stability and repeatability of the experiment.

3.5.1 Steady-State Requirement

Steady-state testing requirements were satisfied for all 54 cases before the data was formally collected and analyzed. It typically requires 8 to 12 hours to enter a steady-state after changing the supply conditions such as the primary air flow rate or the water inlet temperature.

Before each experiment, the test chamber was maintained the steady-state for at least three hours.

The following Table 3.15 shows an example of the steady-state status of each key parameter with an active chilled beam unit. The entire data for steady-state status with all active chilled beam units are in the Appendix section.

Table 3.15 Example of standard deviation data for the steady-state requirement

Measured parameters Standard 55F 58F 61F Requires 60cfm 70cfm 80cfm 60cfm 70cfm 80cfm 60cfm 70cfm 80cfm Reference temperature 0.05 K 0.034 0.037 0.040 0.042 0.039 0.036 0.044 0.033 0.038 Room inside surface temp. 0.5 K 0.31 0.30 0.32 0.30 0.29 0.31 0.29 0.30 0.29 Mean water temperature 0.05 K 0.018 0.028 0.021 0.019 0.030 0.026 0.019 0.024 0.021 Cooling water flow rate 1% 0.79% 0.91% 0.86% 0.86% 0.91% 0.85% 0.84% 0.84% 0.93% Primary air temperature 0.2 K 0.11 0.13 0.12 0.14 0.11 0.12 0.15 0.17 0.12 Primary air flow rate 1.5% 0.41% 0.42% 0.34% 0.44% 0.37% 0.29% 0.42% 0.37% 0.33%

3.5.2 Experiment Result and Analysis: Impacts of each parameter on the water-side

cooling capacity and the air-side cooling capacity

The experiment data were confirmed and aggregated. The separate experiment result of each active chilled beam unit and total comparison experiment results are shown below.

- Experiment results: Manufacturer A, Low flow range unit’s nine cases

Figure 3.23 Manufacturer A, Low flow range unit

- Experiment results: Manufacturer A, Middle flow range unit’s nine cases

Figure 3.24 Manufacturer A, Middle flow range unit

- Experiment results: Manufacturer A, High flow range unit’s nine cases

Figure 3.25 Manufacturer A, High flow range unit

- Experiment results: Manufacturer B, Low flow range unit’s nine cases

Figure 3.26 Manufacturer B, Low flow range unit

- Experiment results: Manufacturer B, Middle flow range unit’s nine cases

Figure 3.27 Manufacturer B, Middle flow range unit

- Experiment results: Manufacturer B, High flow range unit’s nine cases

Figure 3.28 Manufacturer B, High flow range unit

The overall findings from the experiments are summarized below:

• The total cooling capacity shape follows the water-side cooling capacity pattern, since the

water-side cooling capacity is dominant compared to the air-side cooling capacity.

• The water-side cooling capacity increases when the primary air cfm increases or the

water inlet temperature decreases.

• The air-side cooling capacity increases when the primary air cfm increases or the water

inlet temperature increases.

• The water-side cooling capacity trend follows the trend obtained by taking the

temperature difference between the water inlet temperature and the water outlet

temperature.

• The water-side cooling capacity ratio increases when the water inlet temperature

decreases or when the primary air flow rate decreases.

• The air-side cooling capacity ratio increases when the primary air cfm increases or the

water inlet temperature increases.

• The increasing gradient of the air-side cooling capacity ratio diminishes when the

primary air cfm exceeds approximately 60 cfm. Eventually, the air-side cooling capacity

ratio seems to almost plateau at a stationary value.

The air-side cooling capacity increases as the primary air cfm increases, or when the water inlet temperature increases. The correlation between the air-side cooling capacity and the water inlet temperature is a byproduct of the definition of the air-side cooling capacity. Water inlet temperature does not directly affect the air-side cooling capacity, but the two parameters have a correlation. The air-side cooling capacity is proportional to the temperature difference

58 between the induced air temperature and the primary supply air temperature of 64.4°F. When only the water inlet temperature increases, the room and the induced air temperature increase.

Since the primary air temperature is maintained at 64.4°F, the temperature difference between the reference air and the primary air (64.4F) increases when the water inlet temperature increases. Therefore, the air-side cooling capacity increases when the water inlet temperature increases.

Figure 3.29 Total cooling capacity of Manufacturer A's active chilled beams’ cases

Figure 3.30 The water-side/air-side cooling capacity result of Manufacturer A's ACB cases

Figure 3.31 Total cooling capacity of Manufacturer B's ACB cases

Figure 3.32 The water-side/air-side cooling capacity result of Manufacturer B's ACB’ cases

• AL: Manufacturer A’s ACB unit that is appropriate for a low primary air flow rate range • AM: Manufacturer A’s ACB unit that is appropriate for a middle primary air flow rate range • AH: Manufacturer A’s ACB unit that is appropriate for a high primary air flow rate range • Nozzle diameter: AL (0.134”) < AM (0.232”) < AH (0.295”) • BL: Manufacturer B’s ACB unit that is appropriate for a low primary air flow rate range • BM: Manufacturer B’s ACB unit that is appropriate for a middle primary air flow rate range • BH: Manufacturer B’s ACB unit that is appropriate for a high primary air flow rate range • Nozzle diameter: BL (0.250”) = BM (0.250”) < BH (0.300”) • Nozzle size and nozzle spacing are the key factors that differentiate each active chilled beam unit.

In Figure 3.32, the experimental result pattern of the Manufacturer B’s active chilled beam units shows some dissimilarities from that of the Manufacturer A’s in Figure 3.30. When the water-side cooling capacity for the products of Manufacturer B are observed at 40 cfm, we can notice that the water-side cooling capacity of BM (ACB units that covers a middle primary air flow range) is greater than BL (ACB units that covers a low primary air flow range).

However, when we observe the water-side cooling capacity for the products of Manufacturer A at 40 cfm, we can see that the water-side cooling capacity of AM is smaller than AL. These results are dependent on the nozzle size of each active chilled beam unit. Table 3.16 shows the nozzle diameters of the active chilled beam units used for the experiment.

Table 3.16 Nozzle diameter of each active chilled beam (PA means primary air)

Nozzle Diameter Manufacturer A’s ACB units Manufacturer B’s ACB units Low-PA-flow ACB Unit 0.134" 0.250" (20/30/40 cfm) Mid-PA-flow ACB Unit 0.232" 0.250" (40/50/60 cfm) High-PA-flow ACB Unit 0.295" 0.300" (60/70/80 cfm)

The nozzle size of an active chilled beam has an impact on the water-side cooling capacity: smaller nozzle size provides better transfer efficiency. For a fixed primary air flow rate, smaller nozzles in an active chilled beam result in higher plenum pressures. Larger nozzles will allow more primary air to be supplied, but reduce the induction ratio and the transfer efficiency

(PRICE, 2011). Figure 3.33 explains the relationship between the air flow rate and the water-side capacity. The curves illustrate the changing capacities with varying nozzle diameters. (PRICE,

2011).

Figure 3.33 Water-side performance of a typical active chilled beam versus primary air flow (PRICE, 2011)

It is noticeable that AM has a bigger nozzle size than AL, but BM has the same nozzle size as BL. Thus, it is reasonable that the water-side cooling capacities of BL and BM follow the same trend, maintaining a continuously increasing trend like one curve on the graph above.

Contrastively, the water-side cooling capacities of AL and AM show discontinuity, like two curves on the graph above.

The experiment results of the water-side cooling capacity were compared to the manufacturer-rated water-coil cooling capacity. The compared result is shown in Table 3.17.

Table 3.17 Water-side cooling capacity of each experiment case

Unit AL Unit BL Rated cooling coil capacity (water-side) = Rated cooling coil capacity (water-side) = 483 - 654 W (1648 - 2232 Btu/h) 497 - 812 W (1696 - 2772 Btu/h) 55 °F 58 °F 61 °F 55 °F 58 °F 61 °F 20 cfm 463 410 361 20 cfm 389 370 308 30 cfm 546 509 427 30 cfm 479 444 345 40 cfm 575 545 479 40 cfm 553 497 392 Unit AM Unit BM Rated cooling coil capacity (water-side) = Rated cooling coil capacity (water-side) = 596 - 830 W (2034 - 2833 Btu/h) 718 - 988 W (2449 - 3372 Btu/h) 55 °F 58 °F 61 °F 55 °F 58 °F 61 °F 40 cfm 496 439 394 40 cfm 589 530 484 50 cfm 566 534 452 50 cfm 658 572 491 60 cfm 611 570 474 60 cfm 700 603 533 Unit AH Unit BH Rated cooling coil capacity (water-side) = Rated cooling coil capacity (water-side) = 632 - 888 W (2157 - 3031 Btu/h) 846 - 1027 W (2887 - 3504 Btu/h) 55 °F 58 °F 61 °F 55 °F 58 °F 61 °F 60 cfm 459 395 316 60 cfm 599 517 431 70 cfm 484 417 337 70 cfm 638 548 441 80 cfm 520 438 354 80 cfm 663 563 449

The results show that the overall water-side cooling capacity values from the experiment were lower than the rated cooling coil capacity. It means that the thermal load in the chamber was low. For thermal load from envelope or infiltration load are not flexible because of the chamber requirements from the standard, the thermal load that can be changed is the internal load. The internal load in the chamber was around 496.6W for all experiment cases.

3.5.3 Uncertainty Analysis

According to the EN 15116 Standard, the uncertainty of each test was calculated based on the cumulative error law: (CEN, 2008)

2 2 2 2 2 푟 = √푟1 + 푟2 + 푟3 + 푟4 + 푟5 푎푛푑 푟 ≤ ±6% (3.1)

r1 = uncertainty of cooling water flow rate, % r2 = uncertainty of cooling water temperature rise (water temperature difference), % r3 = uncertainty of the temperature difference between the reference temperature and mean cooling water temperature, % r4 = uncertainty of primary air flow rate, % r5 = uncertainty of test room configuration (internal/external heat sources), %

EN 15116 requires the uncertainty r shall be less than ±6%. All result of 54 cases satisfied the uncertainty requirement. The entire data for the uncertainty requirement with all active chilled beam units are in the Appendix section.

Table 3.18 Example of Uncertainty requirement data

Individual Standard 55F 58F 61F Uncertainty (%) Requires 20cfm 30cfm 40cfm 20cfm 30cfm 40cfm 20cfm 30cfm 40cfm r1 3% 1.00% 1.00% 1.00% 1.00% 1.00% 1.00% 1.00% 1.00% 1.00% r2 2% 1.16% 0.99% 0.92% 1.31% 1.08% 0.99% 1.46% 1.23% 1.10% r3 2.5% 1.11% 1.24% 1.32% 1.23% 1.30% 1.44% 1.38% 1.51% 1.62% r4 3% 3.00% 3.00% 3.00% 3.00% 3.00% 3.00% 3.00% 3.00% 3.00% r5 3% 2.00% 2.00% 2.00% 2.00% 2.00% 2.00% 2.00% 2.00% 2.00% r 6% 4.07% 4.06% 4.07% 4.15% 4.11% 4.13% 4.25% 4.22% 4.22%

3.5.4 Induction Ratio Analysis – Thermal Balance Method

As mentioned in the literature review chapter, the induction ratio plays an important role in the calculation of the cooling capacity as well as the supply air temperature or the supply air

66 flow rate of an active chilled beam. But measuring the induced air flow rate or the induction ratio is not simple.

The method to calculate the induction ratio and the induced air mass flow was theorized based on several experimental data and named Thermal Balance Method, which will become the appendix of the new version of the ASHRAE Standard 200. Induction ratio can be calculated by the temperature measurement of the induced air after leaving the cooling coil inside an active chilled beam. It needs to be measured under a steady-state condition while the active chilled beam system is operating.

The mechanism used in Thermal Balance Method is straightforward. The energy balance between a chilled-beam cooling coil and the induced air allows for solving the induced air flow rate. The following equations explain the heat transfer procedure when the induced air passes through the chilled-beam coil under a steady-state condition assuming no condensation is on the chilled-beam coil surface. The measurement details are explained in the Appendix.

푚̇ 푤 ∗ 퐶푝푤 ∗ (푇푤2 − 푇푤1) = 푚̇ 푖 ∗ 퐶푝푎 ∗ (푇푟 − 푇𝑖푛푑푢푐푒푑 푎𝑖푟 푎푓푡푒푟 푙푒푎푣𝑖푛푔 푡h푒 푐표𝑖푙) (3.3)

푚̇ 푤 ∗ 퐶푝푤 ∗ (푇푤2 − 푇푤1) 푚̇ 푖 = (3.4) 퐶푝푎 ∗ (푇푟 − 푇𝑖푛푑푢푐푒푑 푎𝑖푟 푎푓푡푒푟 푙푒푎푣𝑖푛푔 푡h푒 푐표𝑖푙)

푚̇ 푖 = 휌𝑖 ∗ 푞𝑖 (3.5)

푚푤̇ ∗ 퐶푝푤 ∗ (푇푤2 − 푇푤1) 푞𝑖 = (3.6) 휌𝑖 ∗ 퐶푝푎 ∗ (푇푟 − 푇𝑖푛푑푢푐푒푑 푎𝑖푟 푎푓푡푒푟 푙푒푎푣𝑖푛푔 푡h푒 푐표𝑖푙)

Cp and ρ are calculated based on the temperature (functions of the temperature).

Figure 3.34 Temperature sensors positioned on the coil-fin-face

Figure 3.35 Coil-fin-face temperature sensor positioning layout

Table 3.19 shows he result of induction ratio calculation with Thermal Balance Method

Table 3.19 Induction ratio calculation result by Thermal Balance Method

BL (Rated induction ratio = 5.7) AL (Rated induction ratio = 4.99) 55 °F 58 °F 61 °F 55 °F 58 °F 61 °F 20 cfm 6.19 5.98 5.74 20 cfm 6.68 6.54 6.39 30 cfm 5.99 5.79 5.57 30 cfm 6.13 6.04 5.92 40 cfm 5.78 5.68 5.47 40 cfm 5.6 5.77 5.73 Average = 5.80 Average = 6.09 BM (Rated induction ratio = 5.3) AM (Rated induction ratio = 3.43) 55 °F 58 °F 61 °F 55 °F 58 °F 61 °F 40 cfm 5.25 5.15 5.07 40 cfm 4.46 4.37 4.40 50 cfm 5.17 5.13 4.86 50 cfm 4.32 4.46 4.28 60 cfm 5.19 5.14 4.88 60 cfm 4.64 4.36 4.20 Average = 5.09 Average = 4.39 BH (Rated induction ratio = 4.5) AH (Rated induction ratio = 2.5) 55 °F 58 °F 61 °F 55 °F 58 °F 61 °F 60 cfm 4.61 4.46 4.32 60 cfm 3.85 3.58 3.29 70 cfm 4.68 4.48 4.34 70 cfm 3.85 3.58 3.23 80 cfm 4.66 4.52 4.39 80 cfm 3.84 3.53 3.21 Average = 4.50 Average = 3.55 *Water inlet temp. = 55°F, 58°F, 61°F. Supply air flow rates were decided based on the specifications *2 sensors were used for the temp. of the air after leaving the coil of AL, AM, BL, and BM *4 sensors were used for the temp. of the air after leaving the coil of AH and BH

As the result shows, it was found that the induction ratio varies even in one active chilled beam unit according to the supply (primary) air flow rate and the supply water temperature. The results showed that the induction ratio tends to decrease as the water inlet temperature increases or the supply air flow increases. In addition, when the supply air flow rate is high enough (above

50 cfm approximately), the supply air flow rate has less impact on the induction ratio.

It is noticed that the induction ratio decreases as the supply air flow rate increases even though the induced air flow rate increases. This is caused by the definition of the induction ratio, which is the supply air flow rate divided by the induced air flow rate. In other words, the absolute quantity of the induced air increases as the supply air flow rate increases, but the induction ratio decreases because the increase in the induced air volume flow rate is less than the increase in the supply air volume flow rate quantitatively.

The result of induction ratio calculations by the Thermal Balance Method proves that the

Thermal Balance Method is reliable to find the induction ratio of an active chilled beam product.

Particularly in cases of Manufacturer B’s active chilled beam units, the result shows very similar to manufacturer’s rated induction ratio. The result of Manufacturer A’s active chilled beam units shows a higher induction ratio value than the manufacturer’s rated induction ratio.

It will be more reasonable if there are more sensors to measure the temperature of the air leaving the chilled-beam cooling coil. In this experiment, two sensors were used for experiment cases of AL, AM, BL, BM and four sensors were used for experiment cases of AH and BH. It is recommended to use 8 sensors to measure the temperature of the air leaving the chilled-beam cooling coil to get an accurate result. The detailed instruction about Thermal Balance Method is in the appendix of newly revised ASHRAE Standard 200.

Chapter 4 Simulation Analysis

In Chapter 2, calculation methods of various simulation programs were reviewed and compared. After the review, the calculation methods were categorized in three different ways: empirical method, heat balance method, and induction unit calculation method. Representative simulation programs were chosen for each category. In this section, an overview of each calculation method is provided including the pros and cons of each simulation program. Then, the accuracy of each simulation program is compared with the experimental data.

4.1 Overview of Simulation Programs

EnergyPlus is chosen as a representative tool of simulation programs that use empirical calculation method. EnergyPlus requires a set of coefficients for the empirical model that are eventually used to calculate the heat transfer rate of induced air. EnergyPlus iterates until it meets convergence criteria to calculate an energy balance between induced air and chilled water coils in ACB units. An EnergyPlus object called

“AirTerminal:SingleDuct:ConstantVolume:CooledBeam” was used for this study.

TRNSYS is chosen as a representative tool of simulation programs that use the energy balance method. TRNSYS calculates the energy balance between air- and water-side heat transfer in active chilled beam units. TRNSYS specifically introduces a coil bypass ratio in their active chilled beam model to describe the characteristic of airflow in the active chilled beam.

The first active chilled beam model in TRNSYS was developed in 2007 and is constantly being updated. TRNSYS 17, which is the latest version, was purchased for the research.

TRNSYS 17 allows to define the emissivity of the envelope resulting in the simulation result accuracy improvement.

eQUEST is chosen as a representative tool of simulation programs that use the induction unit calculation method. Such programs do not explicitly have active chilled beam models. One way to model active chilled beam systems on such programs is by using an induction unit model with minor changes to mimic the performance of active chilled beam systems. For example, eQUEST has an induction unit model called “Two-Pipe Induction Unit” and this model can be used to model active chilled beams by connecting it to a central air handling unit (usually defined as DOAS). The supply (primary) air induces room air by a certain induction ratio that forms a secondary air loop inside the rooms; The water loop inside an active chilled beam coils is connected to the central water loop or the secondary water loop of the central water system.

eQUEST has limited variables that user can control, and it has a disadvantage that HVAC system configuration is not as flexible as other simulation programs. The cooling coil in the chilled beam and the primary air-side cooling coil can only be supplied by one water loop, which means both cooling coils have the same water inlet and outlet temperature with overall water flow rate. For this reason, the program cannot report the chiller beam water flow rate and inlet/outlet water temperatures separately.

4.1.1 Pros and Cons

4.1.1.1 EnergyPlus

The empirical model to calculate the cooling capacity of active chilled beams in

EnergyPlus is based on the energy equations provided by a specific manufacturer, which is

Halton Oy. EnergyPlus allows the most flexible modeling of HVAC systems among three

72 representative simulation tools. EnergyPlus provides the most detailed and largest number of output variables including the water-side cooling capacity and the air-side cooling capacity of an active chilled beam units.

As mentioned above, EnergyPlus has several advantages, but at the same time there are disadvantages. The first disadvantage is that the empirical model requires empirical coefficients as input variables that most companies other than Halton Oy may not provide. This makes users hard to model active chilled beam systems accurately. For example, Price Industry provides a different set of coefficients using different equations than those being used in EnergyPlus. Upon our request, Price Industry were asked to provide new set of coefficients that can be used for

EnergyPlus in order to accurately model active chilled beam units produced by Price Industry.

EnergyPlus also has another limitation; the maximum induction ratio that user can input is

4. In other words, the user cannot input the induction ratio more than 4. Since most of the active chilled beam products have the induction ratio greater than 4, this limitation needed to be fixed.

Even our experiment cases, four out of six active chilled beam units have the induction ratio that is over 4. In this study, the maximum limit of the induction ratio was removed from EnergyPlus source code.

EnergyPlus does not calculate the state of the induced air after leaving the ACB cooling coil or the change of primary air state when it is mixed with the induced air. Also, EnergyPlus is not capable of modeling variable airflow active chilled beam systems and active chilled/heated beam systems.

4.1.1.2 TRNSYS

TRNSYS solves energy balance equations of the heat transfer process to calculate the air- side and the water-side cooling capacity. TRNSYS is capable of calculating the temperature and humidity ratio of the mixed air of the primary air supplied from the central air handling unit and the induced air after leaving cooling coils in an active chilled beam unit.

TRNSYS requires relatively small number of input variables for active chilled beam units compared with EnergyPlus. The induction ratio and the coil bypass fraction are the only required input variables regarding the active chilled beam unit in TRNSYS. This makes TRNSYS impossible to distinguish between active chilled beam units with the same induction ratio. In fact, active chilled beam units with different dimensions, nozzle sizes, and nozzle spacing may have the same induction ratio. The two active chilled beam units, which have the same induction ratio and different dimensions, nozzle sizes, and nozzle spacing, are expected to perform differently but TRNSYS cannot distinguish these two active chilled beam units.

TRNSYS requires some input variables that are not easy for users to find, including the primary air stream inlet absolute pressure, the primary air stream pressure drop, the secondary air stream inlet pressure, and the secondary air stream pressure drop as inputs. Using inaccurate values for those input variables may result in inaccurate simulation results and analysis.

4.1.1.3 eQUEST

eQUEST does not provide active chilled beam units explicitly and a workaround to model such system is to use the two-pipe induction unit. The two-pipe induction unit model in eQUEST has the only one input regarding the induction unit – the induction ratio. eQUEST allows only one system per one zone. The most critical limitation to model active chilled beam

74 systems with the induction unit model of eQUEST is that the chilled water coils of a central air handling unit and the cooling coils of induction units have to be served by one chilled water loop. It means that both cooling coils must have the same water loop temperature.

eQUEST provides the average water temperature of the water loop for the supply side and the return side of a plant loop as an output. However, these values cover both the main

HVAC system (typically an air handling unit) and the induction unit. It is not possible to separate the fraction of the chilled water goes to the main cooling coil and the chilled water goes to the cooling coil of the induction unit. Thus, it is impossible to obtain cooling coil water flow rate and inlet/outlet water temperature of an induction unit from eQUEST output results. The water-side cooling capacity result can be figured out by indirect calculation but accuracy of the calculation is not guaranteed by this process. Due to the limitation of water-side plant system configuration, users are not likely to be able to model active chilled beam systems accurately in eQUEST. This is likely to result in inaccurate predictions of energy performance active chilled beam systems.

Table 4.1 shows the ability of each program pertains to the key parameters decisive to the water-side and the air-side cooling capacity.

Table 4.1 Key parameters and three simulation programs

4.1.2 Inputs and Outputs

Table 4.2, Table 4.3, and Table 4.4 explain the input and output variables for each program to simulate the active chilled beam system or the induction unit system.

Table 4.2 Inputs and outputs in EnergyPlus

Table 4.3 Inputs and outputs in TRNSYS

Table 4.4 Inputs and outputs in eQUEST

4.1.3 Modeling

Upon acquisition of the all experiment data, the test chamber was modeled for the three simulation programs according to the experimental configuration of the test chamber and set up

HVAC system in the modeling.

Figure 4.1 Building geometry configuration in simulations

The list below describes the common settings for each program.

Schedule - People: No - Infiltration: 11.07 cfm - Equipment load: 496.6 W (1694.50 Btu/hr) - System: Always on Constructions of the test chamber - Roof: R-31.7 (IP) - Wall: R-31.8 (IP) - Floor: R-70.3 (IP) 79

- Interior Ceiling: R-1.3 (IP) HVAC setting - Zone Cooling set-point temperature: Each room temperature of each experiment case - Chilled beam water temperature: 55°F / 58°F / 61°F - Chilled beam supply Air Temperature: 64.4°F (18°C) - Chilled beam supply Air flow rate: 20 cfm / 30 cfm / 40 cfm / 50 cfm / 60 cfm / 70 cfm / 80 cfm

Weather File - Adjusted weather file: constant temperature and relative humidity throughout the year A few of detailed settings were added for EnergyPlus only: - Pre-heating set-point temperature (only for EnergyPlus).: 65°F (18.3 °C) - Options for no wind and no sun’s effect were checked Also, several detailed settings were added for TRNSYS only: - While inputting the building envelope materials, “Thickness or Conductivity/Density Ratio problem” occurred. R-19 Batt insulation was adjusted to meet the total R-value is the same. - External envelope emissivity was set to 0 to remove the long-wave radiation with sky temperature - Primary/Secondary air pressure: 0.83 atm (based on the Boulder’s elevation) - Primary/Secondary air pressure drop: 0 atm (default value) eQUEST also needed a detailed setting for the same condition to EnergyPlus and TRNSYS: - Since eQUEST cannot have the options for no wind, no solar, and no rain, constant wind speed (65m/s) was added to the weather file so that the effect of long-wave radiation with sky temperature can be minimized.

4.1.4 Building Thermal Load Comparison

Before analyzing the simulations in parallel to the experiment data, the simulation models were evaluated and confirmed in detail for each simulation tool against the experiment settings.

Without all the main and terminal (the active chilled beam/the induction unit) HVAC system, the building cooling load of three models were compared. The graph below shows the building cooling load of three models.

Figure 4.2 Building cooling load comparison Table 4.5 Building cooling load comparison

(Unit=Watt) Exp. E+ TRNSYS eQUEST Envelope 101.4 101.6 99.6 91 Infiltration 32.0 32.2 32.0 32 Internal heat gain 496.6 496.6 496.6 496.6 TOTAL 630.0 630.5 628.2 619.6 % Diff 0.0% 0.1% -0.3% -1.7% *According to one of experiment cases’ weather condition – 84F, 39%

Most of the cooling load difference is from the envelope. This is caused by the way each program deals with the short-wave radiation and the long-wave radiation. In the experiment cases, the outdoor temperature is maintained constant since the chamber is placed inside the laboratory. Even though the temperature and relative humidity of each weather file were adjusted, the effect of long-wave radiation cannot be controlled. For example, only EnergyPlus has options for no effect of the sun and the wind. Since TRNSYS and eQUEST do not have

81 those options, a few adjustments to the weather files were done, such as removing the emissivity or adding the wind speed, to avoid the long-wave radiation effect with the sky temperature.

After the evaluation and confirmation of the parameters and details of three simulation models against the experiment settings, modeling analysis of the three programs was carried out with the obtained experimental data for all the cases. The detailed and comprehensive data analysis of the simulation results was done with a focus on explaining the disparities between the simulations and experimental results.

4.2 Active Chilled Beam Simulation in EnergyPlus

4.2.1 Mechanism of the Active Chilled Beam Model in EnergyPlus

The mechanism of EnergyPlus’ active chilled beam model is based on the comparison between the two calculations about the water-side cooling capacities with different equations.

The first calculation (푄̇푤) uses the same equation as the definition in the experiment, which is calculated from the water mass flow rate and the water inlet/outlet temperature difference. The second calculation (푄̇푎𝑖푟) is based on the energy transfer between the water in the coil and the air around the cooling coil. The coefficients provided from the manufacturer are used for the convective heat transfer coefficient. The equation is:

푄̇푤 = 푚̇ 푤,푏푒푎푚 ∗ 퐶푝푤 ∗ (푇푤,표푢푡 − 푇푤,𝑖푛) (4.1)

푄̇푎𝑖푟 = 퐾 ∗ 퐴 ∗ ∆푇 ∗ 퐿푏푒푎푚 (4.2) 퐾 = 훼 ∆푇푛1 (푣휌)푛2 휔푛3 (4.3)

푣휌 = (푣𝑖푛/훼0) 휌푎𝑖푟 (4.4) 푛 푣𝑖푛 = 퐾1 ∆푇 + 퐾𝑖푛 푣푝푟 (4.5) Where

3 푚̇ 푤,푏푒푎푚: supply water flow rate per unit length, m /s-m

퐾: coil heat transfer coefficient, W/m2-K

푣휌: room air mass flow rate across coil, kg/m2-s

3 푣𝑖푛: room air volumetric flow rate across coil per unit length, m /s-m

∆푇: temperature difference between room air temperature and average water temperature, °C

휔: water velocity, m/s

3 푣푝푟: supply air volumetric flow rate per unit length, m /s-m

The subroutine terminates the iteration when the Pw and the Pair match to within 0.1 W.

Since the model equations are nonlinear, they must be solved iteratively (United States

Department of Energy, 2018).

At the first time, it was unable to meet the zone temperature in each case. By some iterative process, it was figured out that internal load had a conductive loss that is proportional to the supply air flow rate – since the heat sources (heat blankets) directly contacted to the chamber’s floor broadly and uniformly. After adjusting the proper internal load loss fraction that ranges from 0 to 0.3 according to the supply air flow rate, the zone temperature became the same as the setpoint temperature.

4.2.2 Empirical Coefficients

The active chilled beam model of EnergyPlus requires the empirical coefficients as inputs for the active chilled beam simulation. This empirical model was developed by Halton Oy, and the empirical coefficients for active chilled beam products from Halton are available from the manufacturer. However, active chilled beam products from other manufacturers do not have the empirical coefficients for the EnergyPlus empirical model. In our experiment cases, for example,

Price Industry uses its own empirical equations and coefficients that are different from that of

EnergyPlus. Upon request, new empirical coefficients for EnergyPlus’ equation were provided from Price Industry after their own calculation.

EnergyPlus requires 10 empirical coefficients as shown in Table 4.2. Each active chilled beam has a different set of empirical coefficients. The set of empirical coefficients for an active chilled beam has to be provided by the manufacturer.

4.2.3 Induction Ratio Analysis

Before simulating the experimental cases, the induction ratio analysis was performed. In

EnergyPlus, the maximum input value for induction ratio was 4. At the first time, the maximum input value of 4 was used instead of the real induction ratio value that is over 4, in most of the active chilled beam unit cases. To avoid this constraint, the EnergyPlus source code was edited and re-compiled so that more accurate analysis can be performed by using accurate induction ratio values. The input range for induction ratio (Kin) was revised to be 0 ≤ Kin ≤10.

As the graph shows below, there were overall changes in both the air cooling capacity and the water cooling capacity after editing the source code and inputting the proper induction ratio value. Meanwhile, the overall patterns of the capacity according to the cases rarely changed. The experiment cases with AL active chilled beam unit, as an example unit, show that the water-side cooling capacity increased when the proper induction ratio of 4.99 was used. This is because the temperature difference between the water inlet temperature and the water outlet temperature was increased. The air-side cooling capacity decreased because the proper induction ratio made the reference temperature lower than the previous result that was with the induction ratio of 4. This is because the air-side cooling capacity is proportional to the difference between the reference

84 temperature and the primary air temperature that is maintained to 64.4°F. When the induction ratio is 4 rather than 4.99, the zone temperature is higher because it cannot meet the setpoint temperature.

Figure 4.3 Induction ratio analysis result sample - AL unit's cases 4.2.4 Experiment Cases Simulation Results

The EnergyPlus simulation result data were confirmed and aggregated. The simulation result of each active chilled beam unit and comparison to the experiment results are shown below.

4.2.4.1 EnergyPlus Simulation Result with Default Empirical Coefficients

- EnergyPlus simulation results: Manufacturer A, Low flow range unit’s nine cases

Figure 4.4 Manufacturer A, Low flow range unit’s EnergyPlus result (Default Coeff.)

- EnergyPlus simulation results: Manufacturer A, Middle flow range unit’s nine cases

Figure 4.5 Manufacturer A, Middle flow range unit’s EnergyPlus result (Default Coeff.)

- EnergyPlus simulation results: Manufacturer A, High flow range unit’s nine cases

Figure 4.6 Manufacturer A, High flow range unit’s EnergyPlus result (Default Coeff.)

- EnergyPlus simulation results: Manufacturer B, Low flow range unit’s nine cases

Figure 4.7 Manufacturer B, Low flow range unit’s EnergyPlus result (Default Coeff.)

- EnergyPlus simulation results: Manufacturer B, Middle flow range unit’s nine cases

Figure 4.8 Manufacturer B, Middle flow range unit’s EnergyPlus result (Default Coeff.)

- EnergyPlus simulation results: Manufacturer B, High flow range unit’s nine cases

Figure 4.9 Manufacturer B, High flow range unit’s EnergyPlus result (Default Coeff.)

4.2.4.2 EnergyPlus Simulation Result with Manufacturer-provided Empirical Coefficients

- EnergyPlus simulation results: Manufacturer A, Low flow range unit’s nine cases

Figure 4.10 Manufacturer A, Low flow range unit’s EnergyPlus result (Empirical Coeff.)

- EnergyPlus simulation results: Manufacturer A, Middle flow range unit’s nine cases

Figure 4.11 Manufacturer A, Middle flow range unit’s EnergyPlus result (Empirical Coeff.)

- EnergyPlus simulation results: Manufacturer A, High flow range unit’s nine cases

Figure 4.12 Manufacturer A, High flow range unit’s EnergyPlus result (Empirical Coeff.) 97

- EnergyPlus simulation results: Manufacturer B, Low flow range unit’s nine cases

Figure 4.13 Manufacturer B, Low flow range unit’s EnergyPlus result (Empirical Coeff.)

- EnergyPlus simulation results: Manufacturer B, Middle flow range unit’s nine cases

100

Figure 4.14 Manufacturer B, Middle flow range unit’s EnergyPlus result (Empirical Coeff.)

101

- EnergyPlus simulation results: Manufacturer B, High flow range unit’s nine cases

102

Figure 4.15 Manufacturer B, High flow range unit’s EnergyPlus result (Empirical Coeff.) 103

Overall findings are below:

• Overall patterns show similar patterns to the experiment results. Particularly in the cases

of Manufacturer B’s Low flow range unit, the simulation result predicts very similar

result to the experiment result.

• The air-side cooling capacity shows a good match with the experiment result because the

zone temperature can be met with the setpoint temperature.

• The water-side cooling capacity shows differences from the experiment data. This is

because the water outlet temperature (and the water temperature difference) is different

from the experiment data. EnergyPlus tends to estimate the water side cooling capacity

less than the experiment.

• Empirical Coefficients are critical to simulate the water-side cooling capacity.

• The air-side cooling capacity ratio is a little bigger than the experiment result.

• In terms of the percentage difference of the water-side cooling capacity results between

the experimental results and the simulation results, the results of EnergyPlus simulation

with the default coefficients give the percentage differences that range from 0.73% to

27.94% and the averaged percentage difference is 17.08%. The results of EnergyPlus

simulation with the manufacturer-provided coefficients give the percentage differences

that range from 0.08% to 14.37% and the averaged percentage difference is 8.17%. Table

4.7 and Table 4.7 represents the percentage difference of the water-side cooling capacity

results between the original EnergyPlus model and the experimental results for each case.

104

Table 4.6 The percentage difference of the water-side cooling capacity between the EnergyPlus model vs experimental results (Default Coeff.)

(%) Case1 Case 2 Case 3 Case4 Case 5 Case 6 Case7 Case 8 Case 9 AL 11.0 9.9 8.7 18.0 17.2 15.6 15.7 19.1 18.4 AM 12.5 10.2 12.3 13.6 15.9 15.2 18.1 17.2 15.9 AH 11.6 10.6 9.9 13.9 13.0 11.1 14.9 13.9 12.5 BL 3.7 1.7 0.7 14.8 13.5 12.5 18.1 17.6 15.7 BM 25.7 25.8 26.1 27.0 27.9 24.6 26.9 27.9 25.0 BH 23.3 21.8 21.4 24.5 24.0 22.6 25.3 24.5 23.8

Table 4.7 The percentage difference of the water-side cooling capacity between the EnergyPlus model vs experimental results (Empirical Coeff.)

(%) Case1 Case 2 Case 3 Case4 Case 5 Case 6 Case7 Case 8 Case 9 AL 10.38 8.52 6.30 10.89 9.59 6.98 4.81 7.58 6.07 AM 10.21 7.59 8.90 9.11 11.00 9.27 11.57 10.39 8.15 AH 10.45 8.73 6.86 11.42 9.77 6.91 11.42 9.68 7.40 BL 2.91 5.30 8.11 0.08 1.81 4.59 0.98 2.16 5.67 BM 14.37 13.85 13.56 11.74 11.63 7.60 8.47 8.58 5.14 BH 10.83 8.65 7.24 10.00 8.78 6.36 9.15 7.74 6.04

4.3 Active Chilled Beam Simulation in TRNSYS

4.3.1 Mechanism of the Active Chilled Beam Model

TRNSYS’s active chilled beam model is based on the heat transfer process around the cooling coil in the active chilled beam. Because it does not use any heat exchanger model or empirical model, it contains several assumptions and uncertainty. Using their own psychrometric function, TRNSYS calculates psychrometric values of each state of air; such as the induced air before leaving/after the coil, the mixed air, and the primary air before meeting the induced air.

The following equations explain the logic of TRNSYS’s active chilled beam model.

1) Read all input values:

푇푙𝑖푞 𝑖푛 = liquid inlet temperature

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푚̇ 푙𝑖푞 𝑖푛 = liquid mass flow rate

푇푝푟𝑖푚 푎𝑖푟 𝑖푛 = primary air inlet tempearture

푤푝푟𝑖푚 푎𝑖푟 i푛 = primary air inlet humidity ratio

푅퐻푝푟𝑖푚 푎𝑖푟 𝑖푛 = primary air inlet relative humidity

푚̇ 푝푟𝑖푚 푎𝑖푟 𝑖푛 = primary air inlet mass flow rate

푃푝푟𝑖푚 푎𝑖푟 𝑖푛 = primary air inlet pressure

푑푃푝푟𝑖푚 푎𝑖푟 𝑖푛 = primary air pressure drop

푇푠푒푐 푎𝑖푟 𝑖푛 = secondary air inlet tempearture

푤푠푒푐 푎𝑖푟 i푛 = secondary air inlet humidity ratio

푅퐻푠푒푐 푎𝑖푟 𝑖푛 = secondary air inlet relative humidity

푃푠푒푐 푎𝑖푟 𝑖푛 = secondary air inlet pressure

푑푃푠푒푐 푎𝑖푟 𝑖푛 = secondary air pressure drop 퐼푛푑푢푐푡𝑖표푛푅푎푡𝑖표 = induction ratio of an active chilled beam 퐵푦푝푎푠푠퐹푟푎푐 = coil air bypass fraction of an active chilled beam

2) Check inputs for validity Check if the input values are less than 0

3) Get enthalpy and dry air density

푃푝푟𝑖푚 푎𝑖푟 𝑖푛, 푇푝푟𝑖푚 푎𝑖푟 𝑖푛, 푅퐻푝푟𝑖푚 푎𝑖푟 𝑖푛, 푤푝푟𝑖푚 푎𝑖푟 𝑖푛

→ ℎ푝푟𝑖푚 푎𝑖푟 𝑖푛 푏푦 푃푆푌퐶퐻푅푂푀퐸푇푅퐼퐶푆 푠푢푏푟표푢푡𝑖푛푒

푃푠푒푐 푎𝑖푟 𝑖푛, 푇푠푒푐 푎𝑖푟 𝑖푛, 푅퐻푠푒푐 푎𝑖푟 𝑖푛, 푤푠푒푐 푎𝑖푟 i푛 → ℎsec 푎𝑖푟 𝑖푛 푏푦 푃푆푌퐶퐻푅푂푀퐸푇푅퐼퐶푆 푠푢푏푟표푢푡𝑖푛푒

4) Determine the flow rate of secondary air

푚̇ 푝푟𝑖푚 푎𝑖푟 𝑖푛 푉푝푟𝑖푚 푎𝑖푟 𝑖푛 = 휌푝푟𝑖푚 푎𝑖푟 𝑖푛

푰풏풅풖풄풕풊풐풏 풓풂풕풊풐 푎푛푑 푉푝푟𝑖푚 푎𝑖푟 𝑖푛 → 푉푠푒푐 푎𝑖푟 𝑖푛

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5) Primary air in → Primary air out (from the nozzle)

푃푝푟𝑖푚 푎𝑖푟 표푢푡 = 푃푝푟𝑖푚 푎𝑖푟 𝑖푛 − 푑푃푝푟𝑖푚 푎𝑖푟

푃푝푟𝑖푚 푎𝑖푟 표푢푡 = 휔푝푟𝑖푚 푎𝑖푟 𝑖푛

ℎ푝푟𝑖푚 푎𝑖푟 표푢푡 = ℎ푝푟𝑖푚 푎𝑖푟 𝑖푛 푈푠푒 푃푆푌퐶퐻푅푂푀퐸푇푅퐼퐶푆 푠푢푏푟표푢푡𝑖푛푒

→ 푇푝푟𝑖푚 푎𝑖푟 표푢푡 푎푛푑 푅퐻푝푟𝑖푚 푎𝑖푟 표푢푡

6) Secondary air in → Secondary air out (from the coil)

푃sec 푎𝑖푟 표푢푡 = 푃sec 푎𝑖푟 𝑖푛 − 푑푃sec 푎𝑖푟

푃sec 푎𝑖푟 표푢푡 = 휔sec 푎𝑖푟 𝑖푛

ℎsec 푎𝑖푟 표푢푡 = ℎsec 푎𝑖푟 𝑖푛 푈푠푒 푃푆푌퐶퐻푅푂푀퐸푇푅퐼퐶푆 푠푢푏푟표푢푡𝑖푛푒

→ 푇sec 푎𝑖푟 표푢푡 푎푛푑 푅퐻sec 푎𝑖푟 표푢푡

7) Mix two air streams

푚̇ 푎𝑖푟 표푢푡 = 푚̇ 푝푟𝑖푚 푎𝑖푟 𝑖푛 + 푚̇ 푠푒푐 푎𝑖푟 𝑖푛

ℎ푝푟𝑖푚 푎𝑖푟 표푢푡 ∗ 푚̇ 푝푟𝑖푚 푎𝑖푟 𝑖푛 + ℎ푠푒푐 푎𝑖푟 표푢푡 ∗ 푚̇ 푠푒푐 푎𝑖푟 𝑖푛 ℎ푎𝑖푟 표푢푡 = 푚̇ 푎𝑖푟 표푢푡

휔푝푟𝑖푚 푎𝑖푟 표푢푡 ∗ 푚̇ 푝푟𝑖푚 푎𝑖푟 𝑖푛 + 휔푠푒푐 푎𝑖푟 표푢푡 ∗ 푚̇ 푠푒푐 푎𝑖푟 𝑖푛 휔푚𝑖푥 2 푎𝑖푟 표푢푡 = 푚̇ 푎𝑖푟 표푢푡

푃푎𝑖푟 표푢푡 = 푃푠푒푐 푎𝑖푟 표푢푡

푈푠푒 푃푆푌퐶퐻푅푂푀퐸푇푅퐼퐶푆 푠푢푏푟표푢푡𝑖푛푒 → 푇푎𝑖푟 표푢푡 , 휔푎𝑖푟 표푢푡 푎푛푑 푅퐻푎𝑖푟 표푢푡

8) In case the condensation has occurred 퐶표푛푑푒푛푠푎푡푒 푠푡푒푎푚 푐표푛푑𝑖푡𝑖표푛푠

푚̇ 푐표푛푑 푝푟𝑖푚 = 푚̇ 푝푟𝑖푚 푎𝑖푟 𝑖푛 ∗ (휔푝푟𝑖푚 푎𝑖푟 𝑖푛 − 휔푝푟𝑖푚 푎𝑖푟 표푢푡)

푚̇ 푐표푛푑 푠푒푐 = 푚̇ 푠푒푐 푎𝑖푟 𝑖푛 ∗ (휔푠푒푐 푎𝑖푟 𝑖푛 − 휔푠푒푐 푎𝑖푟 표푢푡)

푚̇ 푐표푛푑 푚𝑖푥2 = (푚̇ 푝푟𝑖푚 푎𝑖푟 𝑖푛 + 푚̇ 푠푒푐 푎𝑖푟 𝑖푛) ∗ (휔푎𝑖푟 표푢푡 − 휔푚𝑖푥 2 푎𝑖푟 표푢푡)

푚̇ 푐표푛푑 = 푚̇ 푐표푛푑 푝푟𝑖푚 + 푚̇ 푐표푛푑 푠푒푐 + 푚̇ 푐표푛푑 푚𝑖푥 2

푚̇ 푐표푛푑 푠푒푐 ∗ 푇푠푒푐 푎𝑖푟 표푢푡 + 푚̇ 푐표푛푑 푝푟𝑖푚 ∗ 푇푝𝑖푚 푎𝑖푟 표푢푡 + 푚̇ 푐표푛푑 푚𝑖푥 2 ∗ 푇푎𝑖푟 표푢푡 푇푐표푛푑 = 푚̇ 푐표푛푑

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푇푐표푛푑 푝푟𝑖푚 = 푇푝푟𝑖푚 푎𝑖푟 표푢푡

9) Secondary air stream passes over the coil Iteration start: Guess a liquid outlet temperature 푆푒푡 푠표푚푒 𝑖푛𝑖푡𝑖푎푙 푐표푛푑𝑖푡𝑖표푛푠 푓표푟 푡ℎ푒 𝑖푡푒푟푎푡𝑖표푛 푝푟표푐푒푠푠

푇푙𝑖푞 𝑖푛 + 푇푙𝑖푞 표푢푡 푇 = 푙𝑖푞 푎푣푒 2

푇푐표𝑖푙 푎𝑖푟 표푢푡 = 푇푙𝑖푞 푎푣푒

휔푐표𝑖푙 푎𝑖푟 표푢푡 = 휔푠푒푐 푎𝑖푟 𝑖푛

푃푐표𝑖푙 푎𝑖푟 표푢푡 = 푃푠푒푐 푎𝑖푟 𝑖푛 − 푑푃푠푒푐 푎𝑖푟

푈푠푒 푃푆푌퐶퐻푅푂푀퐸푇푅퐼퐶푆 푠푢푏푟표푢푡𝑖푛푒 → ℎ푐표𝑖푙 푎𝑖푟 표푢푡

10) Calculate condensate conditions

푚̇ 푐표푛푑 푐표𝑖푙 = 푚̇ 푠푒푐 푎𝑖푟 𝑖푛 ∗ (1 − 퐵푦푝푎푠푠퐹푟푎푐) ∗ (휔푠푒푐 푎𝑖푟 𝑖푛 − 휔푐표𝑖푙 푎𝑖푟 표푢푡)

푇푐표푛푑 푐표𝑖푙 = 푇푐표𝑖푙 푎𝑖푟 표푢푡

ℎ푐표푛푑 푐표𝑖푙𝑖푠 푐푎푙푐푢푙푎푡푒푑 푏푦 푎푛표푡ℎ푒푟 푇푅푁푆푌푆 푠푢푏푟표푡𝑖푛푒

11) Calculate the energy added to the liquid stream During the Secondary air in → Secondary air out, bypassed + non-bypassed air mixing

푄̇푙𝑖푞 = (푚̇ 푠푒푐 푎𝑖푟 𝑖푛 ∗ (1 − 퐵푦푝푎푠푠퐹푟푎푐) ∗ (ℎ푠푒푐 푎𝑖푟 𝑖푛 − ℎ푐표𝑖푙 푎𝑖푟 표푢푡) − 푚̇ 푐표푛푑 푐표𝑖푙 ∗ ℎ푐표푛푑 푐표𝑖푙)

12) Recalculate the liquid outlet temperature based on the energy added to it

푄̇푙𝑖푞 푇푙𝑖푞 표푢푡푛푒푤 = Tliq in + 푚̇ 푙𝑖푞 𝑖푛 ∗ 퐶푝푙𝑖푞

13) Check if

푇푙𝑖푞 표푢푡푛푒푤 > 푇푠푒푐 푎𝑖푟 𝑖푛 Then limit the fluid outlet temperature to the air temperature

푇푙𝑖푞 표푢푡푛푒푤 = 푇푠푒푐 푎𝑖푟 𝑖푛

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푇푙𝑖푞 𝑖푛 + 푇푙𝑖푞 표푢푡 푇 = 푙𝑖푞 푎푣푒 2 ̇ 푄푙𝑖푞 = 푚̇ 푙𝑖푞 𝑖푛 ∗ 퐶푝푙𝑖푞푢𝑖푑 ∗ (푇푙𝑖푞 표푢푡푛푒푤 − 푇푙𝑖푞 𝑖푛 )

14) Knowing that the same amount of energy must be removed from the air stream, get the air outlet condition

푄̇푙𝑖푞 ℎ푐표𝑖푙 푎𝑖푟 표푢푡 = ℎ푠푒푐 푎𝑖푟 𝑖푛 − 푚̇ 푠푒푐 푎𝑖푟 𝑖푛 ∗ (1 − 퐵푦푝푎푠푠퐹푟푎푐)

휔푐표𝑖푙 푎𝑖푟 표푢푡 = 휔푠푒푐 푎𝑖푟 𝑖푛

푃푐표𝑖푙 푎𝑖푟 표푢푡 = 푃푠푒푐 푎𝑖푟 𝑖푛 − 푑푃푠푒푐 푎𝑖푟

푈푠푒 푃푆푌퐶퐻푅푂푀퐸푇푅퐼퐶푆 푠푢푏푟표푢푡𝑖푛푒 → 푇푐표𝑖푙 푎𝑖푟 표푢푡

15) Iterate and see if the outlet liquid temperature has converged

푇푙𝑖푞 표푢푡푛푒푤 − 푇푙𝑖푞 표푢푡 < 푇푐표푛푣 (DATA TConv/0.01/)

16) Mix two air streams – one bypassed and one at coil conditions

ℎ푠푒푐 푎𝑖푟 표푢푡 = (1 − 퐵푦푝푎푠푠퐹푟푎푐) ∗ ℎ푐표𝑖푙 푎𝑖푟 표푢푡 + 퐵푦푝푎푠푠퐹푟푎푐 ∗ ℎ푠푒푐 푎𝑖푟 𝑖푛

휔푠푒푐 푎𝑖푟 표푢푡 = (1 − 퐵푦푝푎푠푠퐹푟푎푐) ∗ 휔푐표𝑖푙 푎𝑖푟 표푢푡 + 퐵푦푝푎푠푠퐹푟푎푐 ∗ 휔푠푒푐 푎𝑖푟 𝑖푛

푃푠푒푐 푎𝑖푟 표푢푡 = 푃푠푒푐 푎𝑖푟 𝑖푛 − 푑푃푠푒푐 푎𝑖푟

푈푠푒 푃푆푌퐶퐻푅푂푀퐸푇푅퐼퐶푆 푠푢푏푟표푢푡𝑖푛푒 → 푇푠푒푐 푎𝑖푟 표푢푡 푎푛푑 푅퐻푠푒푐 푎𝑖푟 표푢푡

17) Mix the primary air and secondary air streams

푚̇ 푎𝑖푟 표푢푡 = 푚̇ 푝푟𝑖푚 푎𝑖푟 𝑖푛 + 푚̇ 푠푒푐 푎𝑖푟 𝑖푛

푃푎𝑖푟 표푢푡 = 푃푠푒푐 푎𝑖푟 표푢푡

푈푠푒 푃푆푌퐶퐻푅푂푀퐸푇푅퐼퐶푆 푠푢푏푟표푢푡𝑖푛푒 → 푇푎𝑖푟 표푢푡 , 휔푎𝑖푟 표푢푡 푎푛푑 푅퐻푎𝑖푟 표푢푡

18) Check for condensation during mixing

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푚̇ 푐표푛푑 푚𝑖푥 2 = (푚̇ 푝푟𝑖푚 푎𝑖푟 𝑖푛 + 푚̇ 푠푒푐 푎𝑖푟 𝑖푛) ∗ (휔푎𝑖푟 표푢푡 − 휔푚𝑖푥 2 푎𝑖푟 표푢푡)

푇푐표푛푑 푚𝑖푥 2 = 푇푎𝑖푟 표푢푡

19) Recalculate the total amount of energy removed from the secondary air stream

푄̇푠푒푐 푎𝑖푟𝑖 푠푡푟푒푎푚 = 푚̇ 푠푒푐 푎𝑖푟 𝑖푛 ∗ (ℎ푠푒푐 푎𝑖푟 𝑖푛 − ℎ푠푒푐 푎𝑖푟 표푢푡)

푄̇푝푟𝑖푚 푎𝑖푟 푠푡푟푒푎푚 = 푚̇ 푝푟𝑖푚 푎𝑖푟 𝑖푛 ∗ (ℎ푝푟𝑖푚 푎𝑖푟 𝑖푛 − ℎ푎𝑖푟 표푢푡)

4.3.2 Experiment Cases Simulation Result

The TRNSYS simulation data were confirmed and aggregated. The simulation result of each active chilled beam unit and comparison to the experiment results are shown below.

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- TRNSYS simulation results: Manufacturer A, Low flow range unit’s nine cases

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Figure 4.16 Manufacturer A, Low flow range unit’s TRNSYS result

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- TRNSYS simulation results: Manufacturer A, Middle flow range unit’s nine cases

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Figure 4.17 Manufacturer A, Middle flow range unit’s TRNSYS result

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- TRNSYS simulation results: Manufacturer A, High flow range unit’s nine cases

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Figure 4.18 Manufacturer A, High flow range unit’s TRNSYS result

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- TRNSYS simulation results: Manufacturer B, Low flow range unit’s nine cases

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Figure 4.19 Manufacturer B, Low flow range unit’s TRNSYS result

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- TRNSYS simulation results: Manufacturer B, Middle flow range unit’s nine cases

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Figure 4.20 Manufacturer B, Middle flow range unit’s TRNSYS result 120

- TRNSYS simulation results: Manufacturer B, High flow range unit’s nine cases

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Figure 4.21 Manufacturer B, High flow range unit’s TRNSYS result

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Overall findings are below:

• Like EnergyPlus, the air-side cooling capacity shows a good match with the experiment

result, because the zone temperature is able to be met to the setpoint temperature.

• For the water-side cooling capacity, TRNSYS shows a similar result when the supply air

flow is 20 cfm. As the supply air flow rate increases, the water-side cooling capacity

difference between the result of TRNSYS simulation and the experiment also increases.

It can be solved if we can adjust the convective coefficient concept to the cooling coil in

the active chilled beam. Unfortunately, TRNSYS do not use the heat transfer process

between the water inside the cooling coil and the air around the cooling coil. TRNSYS

calculates the water-side cooling capacity with the psychrometric values of the primary

air and the induced air. TRNSYS tends to estimate the water side cooling capacity more

than the experiment.

• In terms of the percentage difference of the water-side cooling capacity results between

the experimental results and the simulation results, the results with the original TRNSYS

model gives the percentage differences that range from 2.79% to 30.24%. Table 4.8

represents the percentage difference of the water-side cooling capacity results between

the original TRNSYS model and the experimental results for each case.

Table 4.8 the percentage difference of the water-side cooling capacity between the TRNSYS model vs experimental results

(%) Case1 Case 2 Case 3 Case4 Case 5 Case 6 Case7 Case 8 Case 9 AL 8.79 7.50 5.72 4.09 5.29 7.08 20.08 16.67 17.37 AM 2.98 3.92 2.79 9.91 7.17 9.28 12.70 13.87 16.14 AH 4.45 6.18 8.68 8.35 10.03 13.20 12.11 14.14 17.00 BL 4.71 6.53 9.71 10.54 12.18 15.57 17.33 18.13 21.99 BM 15.27 15.91 16.29 20.02 20.49 25.36 25.05 25.48 29.65 BH 17.95 20.41 22.69 21.49 24.03 26.78 25.38 27.51 30.24

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4.4 Active Chilled Beam Simulation in eQUEST

4.4.1 Mechanism of the Active Chilled Beam Model

The most critical limitation of eQUEST is that water for the coil in the induction unit and the primary air-side cooling coil in a main HVAC system (air handling unit) can be supplied by only one water loop, which means the same water inlet and outlet temperature with the same overall water flow rate. The program cannot report the water flow rate of the main HVAC system cooling coil and the water flow rate of the cooling coil of the induction unit separately. It also cannot report the water inlet/outlet temperatures of the induction unit cooling coil. Thus, the air-side cooling capacity has to be figured out by the calculation that has the uncertainty of separating the water and the air side cooling capacity. Those features make that eQUEST cannot be expected to provide accurate results for the active chilled beam system’s cooling capacity.

The following Figure 4.22 explains how to separate the water-side cooling capacity and the air- side cooling capacity in eQUEST with an example case.

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Figure 4.22 How to separate the water-side/air-side cooling capacity

4.4.2 Experiment Cases Simulation Result

The eQUEST simulation data were confirmed and aggregated. The simulation result of one representative active chilled beam unit and comparison to the experiment results are shown below.

- eQUEST simulation results: Manufacturer A’s Low flow rate unit’s nine cases

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Figure 4.23 Manufacturer A’s Low flow rate unit’s cases eQUEST result

Overall findings are below:

• Like EnergyPlus and TRNSYS, the air-side cooling capacity shows a good match with

the experiment result because the zone temperature can be met to the setpoint

temperature.

• It is barely showing a pattern or tendency. Since eQUEST cannot provide the water

temperature difference or water outlet temperature, it is barely possible to analyze how to

improve the result.

As the results show, it seems necessary to develop a new model that predicts the accurate water/air cooling capacity of an active chilled beam system for eQUEST.

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Chapter 5 Sensitivity Analysis of the Existing Simulation Models

The prediction of the water-cooling capacity from EnergyPlus and TRNSYS showed some differences from the actual experimental results. In this chapter, factors that have the possibility of improving the water-cooling capacity simulation of EnergyPlus and TRNSYS were analyzed.

5.1 Sensitivity Analysis of EnergyPlus

The simulation results show that EnergyPlus provides the overall precise performance to predict the cooling capacity of the active chilled beam systems, but the water-side cooling capacity needs to be improved.

5.1.1 Impact of Calculated Induction Ratio

To figure out how to improve the water-side cooling capacity simulation more accurate, the simulations using calculated induction ratio values were tested. Instead of using the manufacturer’s rated induction ratio values as inputs, experimentally calculated induction ratio values were used as inputs for each case. In AH’s cases, for example, the rated induction ratio was 2.5 but experimentally calculated induction ratios are ranging from 3.21 to 3.85 according to the water inlet temperature and the supply air flow rate.

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Figure 5.1 AH's result with the rated/calculated induction ratio Since all calculated induction ratio values have greater values than the rated induction ratio, the result shows that the bigger water-side cooling capacities were achieved when the experimentally calculated induction ratios were used. But it is hard to say that the calculated induction ratios provide the better-matched result because the empirical coefficients for the active chilled beam model are already set to be proper to the manufacturer’s rated induction ratio.

5.1.2 Impact of Empirical Coefficients

Most of the inputs for active chilled beam units in EnergyPlus are empirical coefficients.

Empirical coefficients play a key role to calculate the water-side cooling capacity of an active chilled beam. Therefore, the impact of empirical coefficients was analyzed.

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5.1.2.1 Default Coefficients and Manufacturer-provided Coefficients

Firstly, the simulation results with the default coefficients in EnergyPlus ACB model and the results with the empirical coefficients provided by the manufacturer were compared. The

EnergyPlus user normally does not know the exact empirical coefficient of specific active chilled beam unit because these coefficients are typically not provided in the manufacture’s catalog.

Table 5.1 Default coefficients in EnergyPlus and Empirical coefficient of product AH

Test model: Product AH E+ Default coeff. Manufacturer’s coeff. Coil Surface Area per Coil Length 5.422 8.15 (m2/m) Model Parameter a 15.3 18.0 Model Parameter n1 0 0.0 Model Parameter n2 0.84 0.7 Model Parameter n3 0.12 0.17 Model Parameter a0 (m2/m) 0.171 0.246 Model Parameter K1 0.0057 0.0020 Model Parameter n 0.4 0.34 Coefficient of Induction Kin Auto-calculate 2.5 Leaving Pipe Inside Diameter (m) 0.0145 0.013

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Figure 5.2 AH's result with the default /manufacturer-provided empirical coefficients As shown in Figure 5.2, it is obvious that the water-side cooling capacity result with empirical coefficients provided by the manufacturer shows better prediction of the experiment result. The result with the default coefficients shows less water-side cooling capacity.

5.1.2.2 Sensitivity Analysis for Empirical Coefficients

To figure out the other factors that have an influence on the water-side cooling capacity simulation, the water-side capacity calculator was developed. It follows the same equations as

EnergyPlus’ equations and it is MS Excel-based tool so that the inputs can be changed easily. It was validated that the water-side cooling capacity calculator provides the very same result with

EnergyPlus’ result in all six active chilled beam cases.

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Figure 5.3 Water-side cooling capacity calculator based on EnergyPlus equations

Since it was already checked that using the accurate induction ratio with the empirical coefficients provided by the manufacturer is not the solution, each coefficient was analyzed. The sensitivity analysis of the empirical coefficients was conducted to figure out each coefficients’ impact on the water-side cooling capacity.

Table 5.2 Sensitivity analysis for empirical coefficients

Manufacturer- Sensitivity Test model : Product AL E+ Default provided test Coil Surface Area per Coil Length (m2/m) 5.422 8.15 - Model Parameter a 15.3 18.0 - +0.01, Model Parameter n1 0 0 +0.02 Model Parameter n2 0.84 0.7 ± 10% Model Parameter n3 0.12 0.17 ± 10% Model Parameter a0 0.171 0.246 ± 10% Model Parameter K1 0.0057 0.0020 ± 10% Model Parameter n 0.4 0.34 ± 10%

Coefficient of Induction Kin Auto-calculate 4 - Leaving Pipe Inside Diameter 0.0145 0.013 -

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Figure 5.4 Sensitivity analysis for empirical coefficients

Figure 5.4 shows the sensitivity analysis result. It is noticed that the empirical coefficient n2 and a0 are decisive coefficients to calculate the water-side cooling capacity.

Based on the sensitivity analysis, other values were also applied to each coefficient and figured out if they can match the experiment result of water-side cooling capacity. For example, in AM’s cases, if n2 is 0.6 instead of 0.7, or if n3 is 0 instead of 0.17, the water-side cooling

134 capacity calculated by EnergyPlus’ equation shows the accurate match with the experiment result.

Figure 5.5 AM's cases - when n2 = 0.6 instead of 0.7

Figure 5.6 AM's cases - when n3 = 0 instead of 0.17

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Therefore, it can be said that the key factor for simulating precisely with the current

EnergyPlus ACB model is the accurate empirical coefficients.

5.2 Sensitivity Analysis of TRNSYS

As mentioned above, an induction ratio and a coil bypass fraction are the only information that TRNSYS needs regarding an active chilled beam product. In this situation, an induction ratio and a bypass fraction will be the key variables that can change the water-side cooling capacity simulation. Also, TRNSYS has a few inputs regarding the pressure of the primary air and the secondary air: primary air stream inlet pressure, secondary air stream inlet pressure, primary air stream pressure drop, and secondary air stream pressure drop.

5.2.1 Impact of Calculated Induction Ratio

Firstly, the experimentally calculated induction ratio values were applied to the TRNSYS simulation instead of the manufacturer-rated induction ratio to see if the calculated ration can improve the water-side cooling capacity prediction of TRNSYS.

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Figure 5.7 AH's water-side cooling capacity with the rated/calculated induction ratio Since the calculated induction ratio values in every case of Manufacturer A’s ACB units have greater values than the rated induction ratio value, the water-side cooling capacity prediction with the calculated induction ratio also shows greater values than the water-side cooling capacity with the rated induction ratio. It is also assumed that the water-side cooling capacity calculation in TRNSYS is sensitive to the induction ratio – in cases of AH, the rated induction ratio is 2.5 and the calculated induction ratio ranges from 3.21 to 3.85.

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Figure 5.8 BM's water-side cooling capacity with rated/calculated induction ratio

Compared to Manufacturer A’s products, the calculated induction ratio values in every case of Manufacturer B’s ACB units showed similar value to the rated induction ratio. After the calculated induction ratio values for each case were applied to the Manufacturer B’s units, the water-side cooling capacity result became slightly closer to the experiment result in terms of the pattern. But the result with the calculated induction ratio still shows a big difference from the experiment result.

5.2.2 Impact of Bypass Fraction

In Table 3.14, bypass fractions of each active chilled beam were calculated based on the possible area that induced air can penetrate without touching the cooling coil. To investigate the impact of the bypass fraction on TRNSYS’ water-side cooling capacity simulation, the simulation results with the calculated bypass fraction and bypass fraction of 0 were compared.

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Figure 5.9 AH's water-side cooling capacity with Bypass Fraction = 0 and 0.055

Figure 5.9 shows the simulation results with the calculated bypass fraction and bypass fraction of 0. Bypass fraction lessons the water-side cooling capacity prediction of TRNSYS proportionally, but it is hard to say that the result with the calculated bypass fraction brings accurate prediction of the water-side cooling capacity in TRNSYS.

5.2.3 Impact of Inputs Regarding Pressure of Primary Air and Secondary Air

Primary air stream inlet pressure, secondary air stream inlet pressure, primary air stream pressure drop, and secondary air stream pressure drop were regarded as default inputs that are 0.

There might be a possibility of increasing the accuracy of the water-side cooling capacity prediction of TRNSYS if the accurate pressure and pressure drop values of the primary air and

139 secondary air were used. Experimentally, several fixed input values for these four inputs were applied to the TRNSYS simulation.

Table 5.3 Randomly fixed inputs

Base Case 1 Case 2 Case 3 Case 4 Primary air stream inlet pressure 0.84 atm 0.87 atm 0.84 atm 0.84 atm 0.84 atm Secondary air stream inlet pressure 0.84 atm 0.84 atm 0.87 atm 0.84 atm 0.84 atm Primary air stream pressure drop 0 atm 0 atm 0 atm 0.02 atm 0 atm Secondary air stream pressure drop 0 atm 0 atm 0 atm 0 atm 0.02 atm

Figure 5.10 AH’s result with Base case, Case 1, and Case 2: primary/secondary air inlet pressure change

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Figure 5.11 AH’s result with Base case, Case 3, and Case 4: primary/secondary air pressure drop change For Case 1 and Case 2, the water-side cooling capacity decreased by 3% because of the increased primary input pressure value and the water-side cooling capacity increased by 3% because of the increased secondary input pressure value. For Case 3 and Case 4, the water-side cooling capacity result of both cases barely shows any difference from the result of Base case.

According to this result, it is hard to say that the change of the input values of the primary air stream inlet pressure, the primary air stream pressure drop, the secondary air stream inlet pressure, the and secondary air stream pressure drop plays a key role to predict the accurate water-side cooling capacity. Also, the measurement of these four inputs is not easy in a real experiment. Therefore, it can be reasonable to develop a TRNSYS’ active chilled beam model without using those four input values.

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5.3 Necessity of Developing New Simulation Model for Active Chilled Beam

Most of the critical factors that affect the water-side cooling capacity simulation with the current model were investigated for both EnergyPlus and TRNSYS. However, none of these factors, including manufacturer-provided empirical coefficients, calculated induction ratio, calculated bypass fraction, and inputs regarding air pressure, allowed the water-side cooling capacity simulation of TRNSYS to be accurate compared to the experiment result. This investigation tells that the current active chilled beam models in EnergyPlus and TRNSYS need to be developed – not just by adding a few more inputs or equation, but by changing the key calculation logic or algorithm.

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Chapter 6 Development of a New Simulation Model

In Chapter 5, the necessity of developing a new model for active chilled beams was addressed. The empirical model like EnergyPlus or heat transfer model like TRNSYS have limitation to predict the water-side cooling capacity simulation of active chilled beam. The objective of this chapter is to develop a new model with a new algorithm, such as applying the concept of the heat exchanger.

6.1 Background: Heat Exchanger

Heat exchangers facilitate the heat exchange between two fluids that have different temperatures while avoiding mixing with each other. The heat transfer process in heat exchangers normally includes conduction through the surface of the coil separating the two fluids and convection in each fluid. (Cengel et al., 2016)

In a heat exchanger, three types of flow arrangements are available: parallel flow, counter flow, and cross-flow. In parallel flow, both a cold fluid and a hot fluid enter a heat exchanger at the same side and flow in the same direction. In counter flow, a cold fluid and a hot fluid enter a heat exchanger at the opposite side and move in opposite direction. These two arrangements are used in double-pipe heat exchangers. Figure 6.1 shows the flow regimes in parallel flow and counter flow. (Cengel et al., 2016)

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Figure 6.1 Flow configuration of Parallel flow and Counter flow in a double-pipe heat exchanger (Cengel et al., 2016) In a compact heat exchanger, another type of heat exchanger, a cold fluid and a hot fluid move perpendicular to each other. This type of flow configuration is cross-flow. Depending on the flow configuration, the cross-flow can be further classified as mixed flow and unmixed flow as shown in Figure 6.2. (Cengel et al., 2016)

Figure 6.2 Flow configuration of Cross-flow: Both fluids unmixed and one fluid mixed/one fluid unmixed (Cengel et al., 2016) 144

When the two fluids are prevented to move in the transverse direction of each other, with the plate fins forcing the fluid to move through interfin spacing like Figure 6.2 (a), this cross- flow is called unmixed. In Figure 6.2 (b), the fluid is free to flow in the transverse direction. This type of cross-flow is called to be mixed. (Cengel et al., 2016) One fluid is considered to move through separate tubes or fins so that cross mixing cannot happen, while the other fluid is cross mixed. (Kays & London, 1984)

In the case of an active chilled beam, the cold fluid is water and the hot fluid is room induced air. The induced air (hot fluid) is unmixed since the room air is induced and entering the interfin spaces of the cooling coil. The water (cold fluid) is mixed since the water is free to move in the transverse direction as following the coil direction. Therefore, an active chilled beam can be said that a heat exchanger with single-pass cross-flow with one fluid mixed and one fluid unmixed.

6.1.1 The Logarithmic Mean Temperature Difference (LMTD) Method

Heat transfer rate in a heat exchange can be expressed in Newton’s law of cooling as

푄̇ = 푈퐴푠∆푇푚 (6.1) Where

U: overall heat transfer coefficient

As: heat transfer area

∆Tm: appropriate average temperature difference between the two fluids

In general, the overall heat transfer coefficient U and the average temperature difference

between the two fluids ∆푇푚 vary along the heat exchanger and are not constant. It turns out that logarithmic in nature is the appropriate form on the average temperature difference between the

145 cold and the hot fluids. (Cengel et al., 2016). Logarithmic mean temperature difference (LMTD) method for parallel-flow heat exchangers and counter-flow heat exchangers can be described in the following equations.

푄̇ = 푈퐴푠∆푇푙푚 (6.2) ∆푇 − ∆푇 ∆푇 = 1 2 푙푚 ∆푇 (6.3) 푙푛 ( 1) ∆푇2

(6.3) is the log mean temperature difference that is the appropriate form of the mean temperature difference between the two fluids for the heat exchanger analysis. ∆T1 and ∆T2 are decided according to the flow configuration. Figure 6.3 shows the expressions of ∆T1 and ∆T2 in a parallel-flow heat exchanger and a counter-flow heat exchanger.

Figure 6.3 ∆T1 and ∆T2 expressions in a parallel-flow and a counter flow heat exchanger (Cengel et al., 2016)

For cross-flow heat exchangers, similar relations are also developed with the concept of correction factor F.

푄̇ = 푈퐴푠퐹∆푇푙푚,퐶퐹 (6.4)

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∆푇 − ∆푇 ∆푇 = 1 2 푙푚,퐶퐹 ∆푇 (6.5) 푙푛 ( 1) ∆푇2

∆푇1 = 푇ℎ,𝑖푛 − 푇푐,표푢푡 (6.6)

∆푇2 = 푇ℎ,표푢푡 − 푇푐,𝑖푛 (6.7)

Figure 6.4 shows the expressions of ∆T1 and ∆T2 in cross-flow heat exchanger. The correction factor F for cross-flow heat exchanger can be decided by the given charts in Figure

6.5 and Figure 6.6, where the subscripts 1 and 2 represent the inlet and outlet. (Cengel et al.,

2016)

Figure 6.4 The configuration of ∆T1 and ∆T2 in a cross-flow heat exchanger (Cengel et al., 2016)

Figure 6.5 Correction factor F charts for cross-flow heat exchanger, both fluid unmixed (Cengel et al., 2016)

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Figure 6.6 Correction factor F charts for cross-flow heat exchanger, one fluid mixed and the other unmixed (Cengel et al., 2016)

6.1.2 The Effectiveness – Number of Transfer Units (NTU) Method

When the outlet temperature of the two fluids are unknown, using the LMTD method requires tedious iterations, which is not practical. To eliminate this iteration problem, the effectiveness-NTU method was devised by Kays and London in 1955. (Cengel et al., 2016; Kays

& London, 1984) Heat transfer effectiveness ε, a dimensionless parameter, plays a key role in the effectiveness-NTU method.

푄̇ 퐴푐푡푢푎푙 ℎ푒푎푡 푡푟푎푛푠푓푒푟 푟푎푡푒 휀 = = (6.8) 푄푚푎푥 푀푎푥𝑖푚푢푚 푝표푠푠𝑖푏푙푒 ℎ푒푎푡 푡푟푎푛푠푓푒푟 푟푎푡푒

푄̇ = 퐶푐(푇푐,표푢푡 − 푇푐,𝑖푛) = 퐶ℎ(푇ℎ,𝑖푛 − 푇ℎ,표푢푡) (6.9)

퐶푐 = 푚̇ 푐퐶푝푐 (6.10)

퐶ℎ = 푚̇ ℎ퐶푝ℎ (6.11)

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The maximum possible heat transfer rate can be determined when the maximum temperature difference occurs in a heat exchanger, which is the temperature difference between the inlet of the hot fluid and the inlet of the cold fluid.

∆푇푚푎푥 = 푇ℎ,𝑖푛 − 푇푐,𝑖푛 (6.12)

푄̇푚푎푥 = 퐶푚𝑖푛(푇ℎ,𝑖푛 − 푇푐,𝑖푛) (6.13)

Where 퐶푚𝑖푛 is the smaller of 퐶ℎ = 푚̇ ℎ퐶푝ℎ and 퐶푐 = 푚̇ 푐퐶푝푐.

The inlet temperature and the mass flow rate of the two fluids are required to the determination of the maximum possible heat transfer rate 푄̇푚푎푥, and they are usually specified. If the effectiveness 휀 of the heat exchanger is known, it will be possible to determine the actual heat transfer rate 푄̇ without the outlet temperatures of the two fluids.

푄̇ = 휀푄̇푚푎푥 = 휀퐶푚𝑖푛(푇ℎ,𝑖푛 − 푇푐,𝑖푛) (6.14)

The effectiveness 휀 of a heat exchanger relies on the geometry or the flow arrangement of the heat exchanger. This means that different effectiveness relations come from different types of heat exchangers. These effectiveness relations of heat exchangers usually include two dimensionless groups: 푈퐴푠/퐶푚𝑖푛, the number of transfer units (NTU) and the capacity ratio 푐.

(Cengel et al., 2016)

푈퐴 푈퐴 푁푇푈 = 푠 = 푠 (6.15) 퐶푚𝑖푛 (푚̇ 퐶푝)푚𝑖푛

퐶 푐 = 푚𝑖푛 (6.16) 퐶푚푎푥

푈퐴 퐶 휀 = 푓 ( 푠 , 푚𝑖푛 ) = 푓(푁푇푈, 푐) 퐶푚𝑖푛 퐶푚푎푥 (6.17)

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Where

푈 = overall heat transfer coefficient

퐴푠 = heat transfer area of the heat exchanger

퐶푚𝑖푛 = the smaller of 퐶ℎ = 푚̇ ℎ퐶푝ℎ and 퐶푐 = 푚̇ 푐퐶푝푐

퐶푚푎푥 = the larger of 퐶ℎ = 푚̇ ℎ퐶푝ℎ and 퐶푐 = 푚̇ 푐퐶푝푐

Effectiveness relations were developed for each type of heat exchanger. The detailed relations are shown in Table 6.1 and Table 6.2.

Table 6.1 Effectiveness relations for heat exchangers (Kays & London, 1984)

Heat exchanger type Effectiveness relation 1. Double pipe: 1 − 푒푥푝 [−푁푇푈(1 + 푐)] 휀 = Parallel-flow 1 + 푐 1 − 푒푥푝 [−푁푇푈(1 − 푐)] Counter-flow 휀 = 1 − 푐 푒푥푝 [−푁푇푈(1 − 푐)] 2. Shell-and-tube: −1 1 + 푒푥푝[−푁푇푈√1 + 푐2] One-shell pass 2, 4, … 휀 = 2 {1 + 푐 + √1 + 푐2 } tube passes 1 − 푒푥푝[−푁푇푈√1 + 푐2] 3. Cross-flow (single-pass) 푁푇푈0.22 휀 = 1 − 푒푥푝 { [푒푥푝(−푐 푁푇푈0.78) − 1]} Both fluids unmixed 푐 Cmax mixed, 1 휀 = (1 − 푒푥푝 {−푐[1 − 푒푥푝(−푁푇푈)]}) Cmin unmixed 푐 Cmin mixed, 1 휀 = 1 − 푒푥푝 {− [1 − 푒푥푝 (−푐 푁푇푈)]} Cmax unmixed 푐

4. All heat exchangers with c = 0 휀 = 1 − 푒푥푝 (−푁푇푈)

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Table 6.2 NTU relations for heat exchangers (Kays & London, 1984)

Heat exchanger type NTU relation 1. Double pipe: 푙푛 [1 − 휀(1 + 푐)] 푁푇푈 = − Parallel-flow 1 + 푐 1 휀 − 1 Counter-flow 푁푇푈 = 푙푛 ( ) 푐 − 1 휀푐 − 1 2 2. Shell-and-tube: 1 − 1 − 푐 − √1 + 푐2 One-shell pass 2, 4, … 푁푇푈 = − 푙푛 (휀 ) √1 + 푐2 2 2 tube passes 휀 − 1 − 푐 + √1 + 푐 3. Cross-flow (single-pass) 푙푛 (1 − 휀푐) Cmax mixed, 푁푇푈 = −푙푛 [1 + ] Cmin unmixed 푐 Cmin mixed, 푙푛 [푐 푙푛(1 − 휀) + 1] 푁푇푈 = − Cmax unmixed 푐

4. All heat exchangers with c = 0 푁푇푈 = −푙푛 (1 − 휀)

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Figure 6.7 Effectiveness for heat exchangers (Cengel et al., 2016; Kays & London, 1984)

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The fluid with 퐶min may be the cold fluid or the hot fluid. According to Equations (6.8),

(6.9), (6.10), and (6.11), the effectiveness can be expressed only with the temperature of the two fluids.

푄̇ 퐶ℎ(푇ℎ,𝑖푛 − 푇ℎ,표푢푡) (푇ℎ,𝑖푛 − 푇ℎ,표푢푡) 퐹표푟 퐶ℎ = 퐶푚𝑖푛, 휀 = = = (6.18) 푄푚푎푥 퐶푚𝑖푛(푇ℎ,𝑖푛 − 푇푐,𝑖푛) (푇ℎ,𝑖푛 − 푇푐,𝑖푛)

푄̇ 퐶푐(푇푐,표푢푡 − 푇푐,𝑖푛) (푇푐,표푢푡 − 푇푐,𝑖푛) 퐹표푟 퐶푐 = 퐶푚𝑖푛, 휀 = = = (6.19) 푄푚푎푥 퐶푚𝑖푛(푇ℎ,𝑖푛 − 푇푐,𝑖푛) (푇ℎ,𝑖푛 − 푇푐,𝑖푛)

6.1.3 Active Chilled Beam Experimental Cases

As described in 6.1, the cold fluid is the chilled water and the hot fluid is the induced air.

The heat exchanger (cooling coil) in an active chilled beam unit. The induced air (hot fluid) is unmixed since the room air is induced and entering the interfin spaces of the cooling coil. The water (cold fluid) is mixed since the water is free to move in the transverse direction as following the coil direction. Therefore, an active chilled beam can be said that a heat exchanger with single-pass cross-flow with one fluid mixed and one fluid unmixed. For every experiment case, the fluid with 퐶max is the chilled water that is the cold fluid, and the fluid with 퐶min is the induced air (the air entering the coil) that is the hot fluid.

6.1.3.1 Correction Factor F in LMTD Method

With the results of 54 experiment cases, correction factors F for each experiment case were calculated according to Figure 6.7 (f). Table 6.3 shows the correction factor F for each experiment case.

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Table 6.3 Calculated Correction factor F for each experiment case

Case 1 Case 2 Case 3 Case 4 Case 5 Case 6 Case 7 Case 8 Case 9 AL 0.88 0.88 0.87 0.91 0.90 0.90 0.91 0.91 0.90 AM 0.92 0.92 0.92 0.92 0.92 0.92 0.92 0.92 0.91 AH 0.94 0.94 0.94 0.94 0.94 0.94 0.94 0.94 0.94 BL 0.85 0.85 0.84 0.88 0.87 0.86 0.86 0.85 0.86 BM 0.85 0.85 0.85 0.85 0.85 0.83 0.86 0.86 0.87 BH 0.86 0.88 0.85 0.86 0.87 0.86 0.86 0.86 0.86

According to the result, it is interesting that the correction factor F depends on the active chilled beam geometry regardless of the supply condition including the supply air mass flow rate or the supply water temperature.

6.1.3.2 Effectiveness ε in ε-NTU Method

With the results of 54 experiment cases, effectiveness ε were also calculated according to the equation in Table 6.1. Table 6.4 shows the effectiveness ε for each experiment case.

Table 6.4 Effectiveness ε for each experiment case

Case 1 Case 2 Case 3 Case 4 Case 5 Case 6 Case 7 Case 8 Case 9 AL 0.74 0.75 0.75 0.68 0.69 0.69 0.62 0.62 0.62 AM 0.66 0.66 0.67 0.62 0.62 0.63 0.55 0.58 0.59 AH 0.54 0.57 0.61 0.51 0.54 0.58 0.49 0.52 0.56 BL 0.79 0.80 0.81 0.74 0.75 0.76 0.69 0.70 0.70 BM 0.73 0.74 0.75 0.69 0.69 0.70 0.63 0.64 0.65 BH 0.66 0.67 0.68 0.61 0.62 0.63 0.57 0.58 0.59

According to the result, effectiveness ε depends on not only the geometry of the active chilled beam unit but also the supply condition including the supply air mass flow rate or the supply water temperature.

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6.2 Development of New TRNSYS Model

6.2.1 Algorithm of the new TRNSYS model

The current TRNSYS model has a possibility to be developed based on the heat exchanger theory mentioned above since the algorithm of the current TRNSYS model is a simple heat balance model and it contains several assumptions and uncertainty. The main algorithm that calculates the water-side cooling capacity of the active chilled beam is described below.

• Iterative process until 푇푙𝑖푞 out and 푇푙𝑖푞 표푢푡푛푒푤 become the same value

푄̇푙𝑖푞 = 푚푠푒푐 푎𝑖푟 𝑖푛 ∗ (1 − 퐵푦푝푎푠푠퐹푟푎푐) ∗ (ℎ푠푒푐 푎𝑖푟 𝑖푛 − ℎ푐표𝑖푙 푎𝑖푟 표푢푡) − 푚푐표푛푑 푐표𝑖푙 ∗ ℎ푐표푛푑 푐표𝑖푙

푇푠푒푐 푎𝑖푟 𝑖푛 − 푇푙𝑖푞 𝑖푛 푇푙𝑖푞 out = [ 푇푙𝑖푞 𝑖푛 + ] ≫ 푤𝑖푙푙 푏푒 푢푝푑푎푡푒푑 표푛 푡ℎ푒 푛푒푥푡 𝑖푡푒푟푎푡𝑖표푛 푝푟표푐푒푠푠 퐶푝푙𝑖푞푢𝑖푑

푻풄풐풊풍 풂풊풓 풐풖풕 = 푻풍풊풒 풂풗풆

푄̇푙𝑖푞 = 푚푠푒푐 푎𝑖푟 𝑖푛 ∗ (1 − 퐵푦푝푎푠푠퐹푟푎푐) ∗ (h푠푒푐 푎𝑖푟 𝑖푛 − h푐표𝑖푙 푎𝑖푟 표푢푡) − 푚푐표푛푑 푐표𝑖푙 ∗ h푐표푛푑 푐표𝑖푙

̇ 푄푙𝑖푞 푇푙𝑖푞 표푢푡푛푒푤 = 푇푙𝑖푞 𝑖푛 + 푚푙𝑖푞 𝑖푛 ∗ 퐶푝푙𝑖푞

̇ → 푄푙𝑖푞 = 푚푙𝑖푞 𝑖푛 ∗ 퐶푝푙𝑖푞 ∗ (푇푙𝑖푞 표푢푡푛푒푤 − 푇푙𝑖푞 𝑖푛)

Check if |푇푙𝑖푞 out − 푇푙𝑖푞 표푢푡푛푒푤| < 0.1

If not, replace 푇푙𝑖푞 표푢푡

푇푙𝑖푞 표푢푡 + 푇푙𝑖푞 표푢푡 푇 = 푛푒푤 ≫ 푈푝푑푎푡푒 푇 푎푛푑 푑표 푡ℎ푒 𝑖푡푒푟푎푡𝑖표푛 푎푔푎𝑖푛 푙𝑖푞 표푢푡 2 푙𝑖푞 표푢푡

The main iteration algorithm has a critical assumption of 푇푐표𝑖푙 푎𝑖푟 표푢푡 = 푇푙𝑖푞 푎푣푒. The starting point to develop a new model is to replace this assumption with an equation that contains the heat exchanger concept. By using both the LMTD method and the 휀-NTU method, it can be updated.

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The effectiveness 휀 of active chilled beam coil can be decided with 휀-NTU method. It can be expressed in two ways. Firstly, the effectiveness of each active chilled beam can be expressed with the experiment measurements according to (6.18), since the fluid with 퐶min is the induced air (the hot fluid).

(푇 − 푇 ) (푇 − 푇 ) 휀 = ℎ,𝑖푛 ℎ,표푢푡 = 푎𝑖푟,𝑖푛 푎𝑖푟,표푢푡 (6.20) (푇ℎ,𝑖푛 − 푇푐,𝑖푛) (푇푎𝑖푟,𝑖푛 − 푇푤푎푡푒푟,𝑖푛)

(6.20) can be rearranged to explain the relationship between 푇푐표𝑖푙 푎𝑖푟 표푢푡 and the experimental measurements. Note that 푇푐표𝑖푙 푎𝑖푟 표푢푡 in TRNSYS code is the same value as the temperature of the air leaving the cooling coil, 푇푎𝑖푟,표푢푡.

푇푐표𝑖푙 푎𝑖푟 표푢푡 = (1 − 휀)푇푎𝑖푟,𝑖푛 + 휀 푇푤푎푡푒푟,𝑖푛 (6.21)

Secondly, 휀 can also be expressed with NTU and c according to Table 6.1.

1 휀 = (1 − 푒푥푝 {−푐[1 − 푒푥푝(−푁푇푈)]}) (6.22) 푐

Where

푈퐴 푈퐴 퐶 푁푇푈 = 푠 = 푠 푎푛푑 푐 = 푚𝑖푛 퐶푚𝑖푛 (푚̇ 퐶푝)푚𝑖푛 퐶푚푎푥

Calculation on NTU requires the overall heat transfer coefficient U and the average temperature difference between the two fluids As. These two values are hard to be decided as one specific value. Still, it is possible to calculate (UAs) if the calculation circumvents the way using the LMTD method. (6.23) is the equation that (6.4) was rearranged.

푄̇ 푈퐴푠 = (6.23) 퐹∆푇푙푚,퐶퐹

푄̇ is the heat transfer rate in a heat exchanger, which is the same rate as the water-side heat transfer rate in (2.3). Therefore, 푄̇ can be substituted by (2.3).

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푚푤 퐶푝푤 (푇푤,표푢푡 − 푇푤𝑖푛) 푈퐴푠 = (6.24) 퐹 ∆푇푙푚,퐶퐹

According to (6.24), UAs can be calculated if the correction factor F and the log mean temperature difference ∆푇푙푚,퐶퐹 are known. It was verified in 6.1.3.1 that the Correction Factor F is not affected by the supply air and water conditions and it ranges from 0.84 to 0.94. In an effort to reduce the number of inputs, the correction factor F can be set a specific constant-coefficient for all active chilled beam units. This F value was decided with the value that gives the most accurate prediction of the water-side cooling capacity of ACB for all experiment cases, which is

0.85. ∆푇푙푚,퐶퐹 can be expressed the function of 푇ℎ,𝑖푛, 푇ℎ,표푢푡, 푇푐,𝑖푛, and 푇푐,표푢푡 according to

(6.5),(6.6), and (6.7).

∆푇 − ∆푇 (푇 − 푇 ) − (푇 − 푇 ) ∆푇 = 1 2 = ℎ,𝑖푛 푐,표푢푡 ℎ,표푢푡 푐,𝑖푛 푙푚,퐶퐹 ∆푇 푇 − 푇 (6.25) 푙푛 ( 1) 푙푛 ( ℎ,𝑖푛 푐,표푢푡) ∆푇2 푇ℎ,표푢푡 − 푇푐,𝑖푛

푇ℎ,𝑖푛 and 푇푐,𝑖푛 are the input values. 푇ℎ,𝑖푛 is the temperature of induced air that is room in the experiment and the zone setpoint temperature in the simulation. 푇푐,𝑖푛 is the chilled water supply temperature. 푇ℎ,표푢푡 and 푇푐,표푢푡 will be the output values. 푇ℎ,표푢푡 is the variable of

푇푐표𝑖푙 푎𝑖푟 표푢푡 in TRNSYS code, and 푇푐,표푢푡 is the chilled water outlet temperature that is 푇푙𝑖푞 표푢푡 in

TRNSYS code. These outputs will be calculated as the program takes iterations until when

푇푙𝑖푞 표푢푡 and 푇푙𝑖푞 표푢푡,푛푒푤 match to within 0.1.

Therefore, the new algorithm that links the LMTD Method and the effectiveness-NTU

Method is formed by connecting these equations (6.20) to (6.25). Figure 6.8 describes the flow chart of the new model’s algorithm. The main part of the code script for the new TRNSYS model is in A. Appendix, 4).

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Figure 6.8 Flow chart of the algorithm of the new ACB model

6.2.2 Simulation Results with the new TRNSYS model

The simulation results with the new TRNSYS model for each experimental case were confirmed and aggregated. Basically, all input data are the same. The calculated induction ratio values and bypass fraction values were used. The water-side cooling capacity simulation results with the new TRNSYS model, the experiment results, and the simulation results with the original

TRNSYS model are shown below.

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Figure 6.9 The water-side cooling capacity result of AL unit with the new TRNSYS model

Figure 6.10 The water-side cooling capacity result of AM unit with the new TRNSYS model

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Figure 6.11 The water-side cooling capacity result of AH unit with the new TRNSYS model

Figure 6.12 The water-side cooling capacity result of BL unit with the new TRNSYS model

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Figure 6.13 The water-side cooling capacity result of BM unit with the new TRNSYS model

Figure 6.14 The water-side cooling capacity result of BH unit with the new TRNSYS model

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It is noticeable that the prediction of water-side cooling capacity simulation with the new

TRNSYS model shows accurate results compared to the results with the original TRNSYS model. In terms of the percentage difference of the water-side cooling capacity results between the experimental results and the simulation results, the new TRNSYS model shows the percentage differences that range from 0.06% to 6.60%, where the results with the original

TRNSYS model gives the percentage differences that range from 2.79% to 30.24%.

Table 6.5 represents the percentage difference of the water-side cooling capacity results between the new TRNSYS model and the experimental results for each case, and Table 6.6 represents the percentage difference of the water-side cooling capacity results between the original TRNSYS model and the experimental results for each case.

Table 6.5 the percentage difference of the water-side cooling capacity between the new TRNSYS model vs experimental results

(%) Case1 Case 2 Case 3 Case4 Case 5 Case 6 Case7 Case 8 Case 9 AL 4.69 4.81 5.20 5.14 4.62 5.14 2.56 3.20 3.48 AM 1.09 1.20 1.77 0.63 1.02 1.90 2.61 0.06 0.49 AH 4.60 1.54 1.89 4.89 1.53 2.34 4.84 1.08 2.44 BL 0.07 1.38 2.03 2.00 3.72 3.62 2.43 3.64 3.06 BM 4.71 6.04 6.60 4.75 5.06 4.82 2.80 3.11 3.31 BH 3.07 3.34 3.76 1.04 2.20 1.83 0.06 0.24 0.17

Table 6.6 the percentage difference of the water-side cooling capacity between the original TRNSYS model vs experimental results

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6.3 Development of New EnergyPlus Model

6.3.1 Algorithm of the new EnergyPlus model

The current EnergyPlus model has a backbone of comparison between the two equations that calculate the water-side cooling capacities with different equations. One equation (Pw) is the heat transfer rate from the waterside, which is the same as the definition of the water-side cooling capacity describes in (2.3). The other equation (Pair) is the heat transfer rate from the airside. The subroutine terminates the iteration when the Pw and the Pair match to within 0.1W.

For the new EnergyPlus model, the same equations in the new TRNSYS model were applied without changing the backbone of the EnergyPlus code. Like the new TRNSYS model, the new EnergyPlus algorithm also links the LMTD Method and the effectiveness-NTU based on the equations (6.20) to (6.25). Therefore, the number of inputs is decreased significantly. The empirical coefficients of active chilled beam is no longer necessary. Figure 6.15 shows the number of the decreased inputs in the new EnergyPlus model.

Figure 6.15 The number of reduced input values: the original model → the new model

The only difference between the new EnergyPlus model and the new TRNSYS model is the way to terminate their iteration. The main algorithm of the new EnergyPlus active chilled

163 beam model is described below. It is a series of equations that are the same as the logic of the new TRNSYS model. The main part of the code script for the new EnergyPlus model is in

Appendix.

푃푤 = 푚̇ 푤퐶푝푤(푇푤,표푢푡 − 푇푤,𝑖푛) (6.26)

푃푎𝑖푟 = 푚̇ 푎𝑖푟,𝑖푛푑푢푐푒푑(ℎ푎𝑖푟,𝑖푛푑푢푐푒푑 − ℎ푎𝑖푟,푐표𝑖푙 표푢푡)(1 − 퐵푦푝푎푠푠퐹푟푎푐푡𝑖표푛) (6.27)

퐶푚𝑖푛 = 푚̇ 푎𝑖푟,𝑖푛푑푢푐푒푑퐶푝푎𝑖푟 (6.28)

퐶푚푎푥 = 푚̇ 푤퐶푝푤 (6.29)

퐶 푐 = 푚𝑖푛 (6.30) 퐶푚푎푥

(푇 − 푇 ) − (푇 − 푇 ) ∆푇 = 푎𝑖푟,𝑖푛푑푢푐푒푑 푤,표푢푡 푎𝑖푟,푐표𝑖푙 표푢푡 푤,𝑖푛 푙푚,퐶퐹 푇 − 푇 (6.25) 푙푛 ( 푎𝑖푟,𝑖푛푑푢푐푒푑 푤,표푢푡) 푇푎𝑖푟,푐표𝑖푙 표푢푡 − 푇푤,𝑖푛

퐶푚푎푥(푇푤,표푢푡 − 푇푤,𝑖푛) 푈퐴푠 = (6.31) 퐹 ∆푇푙푚,퐶퐹

푈퐴 푁푇푈 = 푠 (6.32) 퐶푚𝑖푛

1 휀 = (1 − 푒푥푝 {−푐[1 − 푒푥푝(−푁푇푈)]}) (6.22) 푐

푇푎𝑖푟,푐표𝑖푙 표푢푡 = (1 − 휀)푇푎𝑖푟,𝑖푛푑푢푐푒푑 + 휀 푇푤,𝑖푛 (6.33)

푊푎𝑖푟,푐표𝑖푙 표푢푡 = 푊푎𝑖푟,𝑖푛푑푢푐푒푑 (6.34)

ℎ푎𝑖푟,푐표𝑖푙 표푢푡 = 푃푠푦푐ℎ푟표푚푒푡푟𝑖푐 푠푢푏푟표푢푡𝑖푛푒 푤𝑖푡ℎ 푇푎𝑖푟,푐표𝑖푙 표푢푡 푎푛푑 푊푎𝑖푟,푐표𝑖푙 표푢푡 (6.35)

ℎ푎𝑖푟,𝑖푛푑푢푐푒푑 = 푃푠푦푐ℎ푟표푚푒푡푟𝑖푐 푠푢푏푟표푢푡𝑖푛푒 푤𝑖푡ℎ 푇푎𝑖푟,𝑖푛푑푢푐푒푑 푎푛푑 푊푎𝑖푟,𝑖푛푑푢푐푒푑 (6.36)

6.3.2 Simulation Results with the new EnergyPlus model

The simulation results with the new EnergyPlus model for each experimental case were confirmed and aggregated. The calculated induction ratio values and bypass fraction values were

164 used. The water-side cooling capacity simulation results with the new EnergyPlus model, the experiment results, and the simulation results with the original EnergyPlus model with default coefficients are shown below.

Figure 6.16 The water-side cooling capacity result of AL unit with the new EnergyPlus model

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Figure 6.17 The water-side cooling capacity result of AM unit with the new EnergyPlus model

Figure 6.18 The water-side cooling capacity result of AH unit with the new EnergyPlus model

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Figure 6.19 The water-side cooling capacity result of BL unit with the new EnergyPlus model

Figure 6.20 The water-side cooling capacity result of BM unit with the new EnergyPlus model

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Figure 6.21 The water-side cooling capacity result of BH unit with the new EnergyPlus model The results above show that the prediction of water-side cooling capacity simulation with the new EnergyPlus model shows accurate results compared to the results with the original

EnergyPlus model. In terms of the percentage difference of the water-side cooling capacity results between the experimental results and the simulation results, the new EnergyPlus model shows the percentage differences that range from 0.04% to 8.4%, where the results with the original EnergyPlus model gives the percentage differences that range from 0.73% to 27.94%.

Table 6.7 represents the percentage difference of the water-side cooling capacity results between the new EnergyPlus model and the experimental results for each case, and Table 6.8 represents the percentage difference of the water-side cooling capacity results between the original EnergyPlus model and the experimental results for each case.

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Table 6.7 the percentage difference of the water-side cooling capacity between the new EnergyPlus model vs experimental results

(%) Case1 Case 2 Case 3 Case4 Case 5 Case 6 Case7 Case 8 Case 9 AL 1.31 0.91 0.29 2.07 1.93 0.50 0.97 0.04 1.25 AM 1.73 2.03 1.97 2.26 2.38 1.78 6.08 3.15 3.11 AH 8.40 5.60 2.24 8.27 5.59 1.72 8.26 5.17 1.65 BL 2.56 1.94 1.62 0.40 0.24 0.18 0.13 0.05 0.47 BM 2.36 2.74 3.35 2.00 2.25 1.68 0.08 0.41 0.27 BH 1.32 0.63 0.73 1.50 0.64 0.81 2.32 2.30 2.71

Table 6.8 the percentage difference of the water-side cooling capacity between the original EnergyPlus model (default coeff.) vs experimental results

6.4 Comparison of the Results from the New Models

The results from the new TRNSYS model and the new EnergyPlus model were compared to the experiment result, the results of the original TRNSYS model and the original EnergyPlus model with default coefficients altogether. The darker colors represent the new models, and the lighter colors represent the original models. The red colors represent TRNSYS models, and the blue colors represent EnergyPlus models.

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Figure 6.22 Total simulation results comparison of AL unit

Figure 6.23 Total simulation results comparison of AM unit

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Figure 6.24 Total simulation results comparison of AH unit

Figure 6.25 Total simulation results comparison of BL unit

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Figure 6.26 Total simulation results comparison of BM unit

Figure 6.27 Total simulation results comparison of BH unit

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The results show that the new TRNSYS model and the new EnergyPlus model provide the precise prediction of the water-side cooling capacity of an active chilled beam. The cooling capacity differences between the results of the new TRNSYS model and the new EnergyPlus model originate in the convergence condition, the way how these two models terminate their iteration. As mentioned in 6.2.1 and 6.3.1, the condition that decides the termination of the iteration is the water outlet temperature for TRNSYS, and the cooling coil heat transfer rate from the water and the air side for EnergyPlus. The water-side cooing capacity difference between the results of the new TRNSYS model and the new EnergyPlus model range from 8 to 20 W and the percentage difference ranges from 1.71% to 4.78%.

6.5 Validation with other experiment cases

The new model was validated with the new experimental measurements that were not used while the new model was developed. The supply conditions of the new experiment cases are shown in Table 6.9.

Table 6.9 New experiment cases for the validation of the new model

New Case 1 New Case 2 ACB Unit AH BH Chilled Water Flow Rate 1.0 GPM (0.063 L/s) 1.0 GPM (0.063 L/s) Primary Air Temperature 64.4 °F (18 °C) 64.4 °F (18 °C) Chilled Water Supply Temperature 61 °F (16.1 °C) 55 °F (12.8 °C) Primary Air Flow Rate 65 CFM (30.7 L/s) 65 CFM (30.7 L/s)

The new experiment case 1 and case 2 was simulated in EnergyPlus, TRNSYS, the new

EnergyPlus model, and the new TRNSYS model and the water-side cooling capacity result was evaluated. Figure 6.28 shows the water-side cooling capacity simulation results and Figure 6.29 shows the percentage difference between the experiment result and each simulation result.

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Figure 6.28 Water-side cooling capacity experiment/simulation results of new cases

Figure 6.29 Percentage difference of the water-side cooling capacity of each simulation results

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The simulation results show that the prediction of water-side cooling capacity simulation with the new EnergyPlus/TRNSYS model shows better results compared to the results with the original EnergyPlus/TRNSYS model. In terms of the percentage difference of the water-side cooling capacity results between the experimental results and the simulation results, the new

EnergyPlus/TRNSYS model shows the percentage differences of range from 0.55% to 2.14%, where the results with the original EnergyPlus/TRNSYS model gives the percentage differences that range from 9.82 % to 12.65%.

6.6 Discussions

6.6.1.1 UA Calculation in the TRNSYS New Model

The variance in the UA values of the new model were checked from the whole year simulation. The UA value and iteration numbers of two example cases were figured out as outputs in the TRNSYS new model.

Table 6.10 Unit AL, Case 7: supply water temp. = 55°F (12.8°C), supply air flow rate = 40 cfm (18.9 L/s)

Number of Time UA value [W/K] iterations 1/1, 01:00:00 143.12083 80 1/1, 02:00:00 142.92514 95 1/1, 03:00:00 142.92652 72 1/1, 04:00:00 142.92653 72 1/1, 05:00:00 142.92653 72 ⁞ ⁞ ⁞ 12/31, 23:00:00 142.92653 72 12/31, 24:00:00 142.92653 72 *Timestep: 1 hour *Calculated UA value with the experiment data = 143.979

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In Case 7 of Unit AL, the UA values of 8754 timesteps (out of 8760) shows the same value after the first 6 timesteps. The converged UA values show the percentage difference of

0.7% from the calculated UA value with the experiment data.

Table 6.11 Unit BH, Case 9: supply water temp. = 61°F (16.1°C), supply air flow rate = 80 cfm (37.8 L/s)

Number of Time UA value [W/K] iterations 1/1, 01:00:00 229.3597 95 1/1, 02:00:00 229.20426 73 1/1, 03:00:00 229.20422 51 1/1, 04:00:00 229.20422 51 1/1, 05:00:00 229.20422 51 ⁞ ⁞ ⁞ 12/31, 23:00:00 229.20422 51 12/31, 24:00:00 229.20422 51 *Timestep: 1 hour * Calculated UA value with the experiment data = 226.345

In Case 9 of Unit BH, the UA values of 8755 timesteps (out of 8760) shows the same value after the first 5 timesteps. The converged UA values show the percentage difference of

1.2% from the calculated UA value with the experiment data.

The UA values in both example cases converge within the sixth timestep. For example, if it is given to have some spare timestep such as the 10th timestep rather than the 6th timestep, the

UA value is already converged. It makes the UAF can be a combined coefficient. The fixed UA value brings the fixed NTU value, and it brings the fixed effectiveness value by a chain reaction.

Finally, the water outlet temperature and the coil air out temperature are also decided within the

10th timestep. Therefore, it can be said that there is a possibility that the UA value can be fixed with the UA value of the sixth timestep in the code structure of the new model. Figure 6.30 describes the new calculation procedure.

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Figure 6.30 The flow chart of the new calculation procedure

6.6.1.2 Convergence Criteria in the New TRNSYS Model

The original TRNSYS model has a convergence condition is that their water outlet temperature variable and their new water outlet temperature variable converge to each other. The

exact convergence condition is |푇푙𝑖푞 표푢푡 − 푇푙𝑖푞 표푢푡푛푒푤 | < 0.01.

At the first step of development of the new model, the same convergence condition with the original model was applied to the new model. With the convergence condition of |푇푙𝑖푞 표푢푡 −

푇푙𝑖푞 표푢푡푛푒푤 | < 0.01, the simulation time was around 31 seconds. But the result of the water outlet temperature of several experiment cases showed fluctuating value according to timesteps. Figure

6.31 shows the result of one example case that showed fluctuating results of water outlet temperature by timesteps.

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Figure 6.31 Fluctuating results of the water outlet temperature by timesteps

To avoid the fluctuating results of the water outlet temperature by timesteps, the more

stringent convergence conditions of |푇푙𝑖푞 표푢푡 − 푇푙𝑖푞 표푢푡푛푒푤 | < 0.0001 was used in the new model.

The new model was also tested with the convergence criteria of |푇푙𝑖푞 표푢푡 − 푇푙𝑖푞 표푢푡푛푒푤 | < 0.001.

Table 6.12 Simulation results of the models with three convergence conditions

Original TRNSYS New TRNSYS Model Model

Simulation |푇 − 푇 | |푇 − 푇 | |푇 − 푇 | |푇 − 푇 | Convergence 푙𝑖푞 표푢푡 푙𝑖푞 표푢푡푛푒푤 푙𝑖푞 표푢푡 푙𝑖푞 표푢푡푛푒푤 푙𝑖푞 표푢푡 푙𝑖푞 표푢푡푛푒푤 푙𝑖푞 표푢푡 푙𝑖푞 표푢푡푛푒푤 < ퟎ. ퟎퟏ < ퟎ. ퟎퟏ < ퟎ. ퟎퟎퟏ < ퟎ. ퟎퟎퟎퟏ Condition Simulation time 11 seconds 31 seconds 44 seconds 79 seconds # of iterations 14, 9, 1, 1, 1 42, 1, 1, 1, 1 69, 47, 26, 26, 26 95, 73, 51, 51, 51 on timestep … … … … 1, 2, 3, 4, 5 … 푻풘풂풕풆풓 풐풖풕 result * Experiment 18.3 °C 17.7 °C 17.84 °C 17.84 °C Result = 17.84 °C

푻풘풂풕풆풓 풐풖풕 % diff. 2.5 % 0.9 % 0 % 0 % Water-side cooing capacity 584.8 W 406.3 W 448.74 W 449.01 W result

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* Experiment Result = 448.93 W Water-side cooling capacity 30.2 % 10.0 % 0 % 0 % % diff. 푻풘풂풕풆풓 풐풖풕 of The optimal several cases criteria with a Current new Note were fluctuated reasonable TRNSYS model by timestep simulation time

The results show that the new model with the convergence condition of |푇푙𝑖푞 표푢푡 −

푇푙𝑖푞 표푢푡푛푒푤 | < 0.001 provides the most reasonable result with a reasonable simulation time. The

result with the convergence condition of |푇푙𝑖푞 표푢푡 − 푇푙𝑖푞 표푢푡푛푒푤 | < 0.0001 provides more accurate result in terms of 0.01 W unit, but the simulation time become almost doubled when compared to

the result with the convergence condition of |푇푙𝑖푞 표푢푡 − 푇푙𝑖푞 표푢푡푛푒푤 | < 0.001. Therefore, it is

reasonable that the new TRNSYS model has the convergence criteria of |푇푙𝑖푞 표푢푡 − 푇푙𝑖푞 표푢푡푛푒푤 | <

0.001.

6.6.1.3 ACB Simulation under Transient State

To see the water-side cooling capacity of an active chilled beam under transient state, one simple building model of a small one-zone house under Golden, CO weather condition was simulated with both the original EnergyPlus model and the new EnergyPlus model.

The simulation test condition is the same with Case 9, the supply water temperature of

61°F (16.1°C) and the primary air flow rate of 80 cfm (37.8 L/s). The zone cooling setpoint temperature is 74°F (23.3°C). Since it turned out that the internal load was small when compared to the active chilled beam’s rated cooling capacity, the internal load was increased to 800 W.

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• Simulation time:

The original EnergyPlus model: 28 seconds

The new EnergyPlus model: 122 seconds

• The water-side cooling capacity of ACB

Rated water-side cooling capacity of ACB unit = 632 - 888 W (2157 - 3031 Btu/h)

Figure 6.32 Water-side cooling capacity results with transient state

• End use annual energy consumption: Electricity [GJ]

Table 6.13 End use annual energy consumption in transient state

Annual [GJ] Original model New model Cooling 3.95 4.44 Cooling energy consumption % diff. 11.68 % Fan 3.66 3.66 Pump 0.41 0.34

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Like the experiment cases which are steady-state simulations, the original EnergyPlus model tends to estimate the water-side cooling capacity less than the new EnergyPlus model.

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Chapter 7 Conclusion and Future Work

7.1 Conclusion

This research is to verify and enhance the simulation modeling of an active chilled beam system, via a combined and iterative experimental and modeling approach.

Previous research papers regarding the active chilled beam system were reviewed and documented in several categories. After that, the research focused on specifications of existing active chilled beam products and the simulation tools implementing these models, as well as existing experimentations and results. Based on the research, a detailed experimental plan was developed under the guidance of ASHRAE Standard 200P. 54 experiment tests were conducted with six active chilled beams to investigate critical impact factors on the cooling capacity of active chilled beams. Standard-required testing conditions were met before the data was formally collected and analyzed. The key parameters in all experiment cases fulfill the requirements of

EN 15116 and ASHRAE Standard 200P. Uncertainty analysis was done for the experiment results. The study compared the experiment results of each active chilled beam with a focus on the total, water-side and air-side cooling capacities as well as their correlations with the water and air supply conditions. As expected, the total cooling capacity profile follows the water-side cooling capacity profile, since the water-side cooling capacity is dominant compared to the air- side cooling capacity. The water-side cooling capacity increases when the primary air flow rate increases or when the water inlet temperature decreases. The air-side cooling capacity increases when the primary air flow rate increases or when the water inlet temperature increases. The water-side cooling capacity ratio increases when the water inlet temperature decreases or when the primary air flow rate decreases, as anticipated. By contrast, the air-side cooling capacity ratio

182 increases when the primary air flow rate increases or when the water inlet temperature increases as explained above. Interestingly, the increasing gradient of the air-side cooling capacity ratio diminishes when the primary air flow rate exceeds approximately 60 cfm (28.3 L/s). Ultimately, the air-side cooling capacity ratio approaches to almost a stationary value for a given water inlet temperature. The study further validated the impacts of air supply nozzle sizes of active chilled beams on the cooling performance, which is that smaller nozzle size provides a better transfer efficiency.

While conducting the experiment, the induction ratio was analyzed with the Thermal

Balance Method and the validation of the method with separate experiment data was conducted.

It was found that the induction ratio varies even in one active chilled beam unit according to the supply (primary) air flow rate and the supply water temperature. The induction ratio tends to decrease as the water inlet temperature increases or the supply air flow increases. Besides, when the supply air flow rate is high enough (above 50 cfm approximately), the supply air flow rate has less impact on the induction ratio.

On the basis of the findings, a systematic assessment of the simulation capacities of existing active chilled beam models was done with three building energy simulation programs –

EnergyPlus, TRNSYS, and eQUEST. The current active chilled beam simulation models cannot accurately predict the energy performance of the active chilled beams. The active chilled beam model of EnergyPlus uses empirical coefficients that are not easily accessible to the users and have to be provided by the manufacturer. The active chilled beam model of TRNSYS has several assumptions that make the result of the model inaccurate. The induction unit model of eQUEST cannot provide a reasonable result as the active chilled beam models of EnergyPlus and

TRNSYS do.

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Upon the research findings, the new active chilled beam model was developed and integrated with EnergyPlus and TRNSYS. The new active chilled beam model is grounded with the algorithm with the heat exchanger theory and easy-to-measure input parameters upon the findings from the experiment. The new model allows the simulation of the energy performance of active chilled beams to be accurate under a wide range of system capacities and operating conditions without using any empirical coefficients for an active chilled beam.

7.2 Contributions

The contributions of this dissertation is highlighted as below.

• The energy performances of the series of active chilled beams units under various supply

conditions were measured.

• The influence of the water supply temperature, the supply air mass flow rate of an active

chilled beam on the energy performance of active chilled beams were analyzed.

• The impact of the air supply nozzle size of an active chilled beam on the cooling

performance was validated.

• Induction ratio analysis was conducted. The experimental method to measure the

induction ratio of active chilled beams was established, which led to the appendix of

ASHRAE Standard 200-2018.

• The correlation between the induction ratio and the supply water temperature, and the

correlation between the induction ratio and the supply air flow rate were investigated.

• The energy performance prediction of the current simulation ACB models such as

EnergyPlus, TRNSYS, and eQUEST was evaluated and analyzed.

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• The influence of the empirical coefficients of active chilled beams and the induction ratio

of active chilled beams on the EnergyPlus simulation performance was analyzed.

• The influence of the induction ratio of active chilled beams, the bypass fraction of an

active chilled beam, and inputs regarding the air pressure on the TRNSYS simulation

performance were analyzed.

• The new active chilled beam model was developed. The algorithm of the new model is

based on the LMTD method and the effectiveness-NTU method. The model was

integrated with EnergyPlus and TRNSYS. The model shows the accurate prediction of

the energy performance of active chilled beams.

7.3 Future Work

The experiment conducted on this research was to measure the performance of the active chilled beams, which means no controller was used on the experiment. Hence, the room temperatures of some cases are too low to be called a thermal comfort zone. In the future, the active chilled beam experiment with a controller can be investigated in the aspect of thermal comfort.

185

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A. APPENDIX

1) Steady-state Requirement for the Experiment Cases

Table A.1 Steady-state requirement for Manufacturer A’s Low flow range unit

Standard 55F 58F 61F Measured parameters Requires 20cfm 30cfm 40cfm 20cfm 30cfm 40cfm 20cfm 30cfm 40cfm Reference temperature 0.05 K 0.046 0.035 0.031 0.048 0.034 0.042 0.037 0.033 0.047 Room inside surface temp. 0.5 K 0.24 0.27 0.27 0.23 0.25 0.27 0.23 0.25 0.25 Mean water temperature 0.05 K 0.008 0.006 0.016 0.008 0.007 0.004 0.008 0.022 0.011 Cooling water flow rate 1% 0.80% 0.82% 0.95% 0.84% 0.80% 0.88% 0.85% 0.85% 0.89% Primary air temperature 0.2 K 0.106 0.177 0.178 0.105 0.160 0.110 0.110 0.115 0.159 Primary air flow rate 1.5% 0.65% 0.52% 0.35% 0.68% 0.55% 0.37% 0.68% 0.54% 0.41%

Table A.2 Steady-state requirement for Manufacturer A's Middle flow range unit

Standard 55F 58F 61F Measured parameters Requires 40cfm 50cfm 60cfm 40cfm 50cfm 60cfm 40cfm 50cfm 60cfm Reference temperature 0.05 K 0.048 0.047 0.041 0.043 0.049 0.035 0.045 0.034 0.034 Room inside surface temp. 0.5 K 0.29 0.27 0.27 0.28 0.26 0.26 0.25 0.26 0.26 Mean water temperature 0.05 K 0.009 0.007 0.011 0.030 0.004 0.005 0.011 0.007 0.008 Cooling water flow rate 1% 0.97% 0.79% 0.87% 0.98% 0.76% 0.85% 0.81% 0.82% 0.92% Primary air temperature 0.2 K 0.117 0.115 0.106 0.167 0.153 0.139 0.144 0.182 0.177 Primary air flow rate 1.5% 0.72% 0.32% 0.23% 0.72% 0.36% 0.27% 0.76% 0.35% 0.31%

Table A.3 Steady-state requirement for Manufacturer A's High flow range unit

Standard 55F 58F 61F Measured parameters Requires 60cfm 70cfm 80cfm 60cfm 70cfm 80cfm 60cfm 70cfm 80cfm Reference temperature 0.05 K 0.034 0.037 0.040 0.042 0.039 0.036 0.044 0.033 0.038 Room inside surface temp. 0.5 K 0.31 0.30 0.32 0.30 0.29 0.31 0.29 0.30 0.29 Mean water temperature 0.05 K 0.018 0.028 0.021 0.019 0.030 0.026 0.019 0.024 0.021 Cooling water flow rate 1% 0.79% 0.91% 0.86% 0.86% 0.91% 0.85% 0.84% 0.84% 0.93% Primary air temperature 0.2 K 0.11 0.13 0.12 0.14 0.11 0.12 0.15 0.17 0.12 Primary air flow rate 1.5% 0.41% 0.42% 0.34% 0.44% 0.37% 0.29% 0.42% 0.37% 0.33%

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Table A.4 Steady-state requirement for Manufacturer B's Low flow range unit

Standard 55F 58F 61F Measured parameters Requires 20cfm 30cfm 40cfm 20cfm 30cfm 40cfm 20cfm 30cfm 40cfm Reference temperature 0.05 K 0.046 0.045 0.039 0.049 0.037 0.039 0.043 0.039 0.039 Room inside surface temp. 0.5 K 0.30 0.30 0.32 0.27 0.29 0.30 0.27 0.29 0.28 Mean water temperature 0.05 K 0.012 0.013 0.011 0.027 0.020 0.011 0.008 0.012 0.009 Cooling water flow rate 1% 0.46% 0.46% 0.48% 0.48% 0.43% 0.46% 0.44% 0.42% 0.47% Primary air temperature 0.2 K 0.137 0.110 0.149 0.110 0.132 0.118 0.112 0.155 0.124 Primary air flow rate 1.5% 1.14% 0.53% 0.35% 1.15% 0.54% 0.35% 1.22% 0.56% 0.33%

Table A.5 Steady-state requirement for Manufacturer B's Middle flow range unit

Standard 55F 58F 61F Measured parameters Requires 40cfm 50cfm 60cfm 40cfm 50cfm 60cfm 40cfm 50cfm 60cfm Reference temperature 0.05 K 0.048 0.037 0.043 0.038 0.041 0.042 0.048 0.036 0.039 Room inside surface temp. 0.5 K 0.27 0.29 0.29 0.27 0.28 0.29 0.26 0.25 0.27 Mean water temperature 0.05 K 0.011 0.032 0.015 0.010 0.007 0.020 0.006 0.004 0.016 Cooling water flow rate 1% 0.85% 0.82% 0.80% 0.81% 0.78% 0.84% 0.77% 0.99% 0.99% Primary air temperature 0.2 K 0.133 0.119 0.110 0.111 0.140 0.121 0.122 0.104 0.106 Primary air flow rate 1.5% 0.39% 0.33% 0.25% 0.36% 0.29% 0.27% 0.41% 0.28% 0.27%

Table A.6 Steady-state requirement for Manufacturer B's High flow range unit

Standard 55F 58F 61F Measured parameters Requires 60cfm 70cfm 80cfm 60cfm 70cfm 80cfm 60cfm 70cfm 80cfm Reference temperature 0.05 K 0.029 0.039 0.037 0.030 0.037 0.042 0.044 0.040 0.048 Room inside surface temp. 0.5 K 0.30 0.33 0.34 0.30 0.32 0.32 0.30 0.32 0.31 Mean water temperature 0.05 K 0.007 0.005 0.006 0.004 0.006 0.005 0.005 0.008 0.004 Cooling water flow rate 1% 0.82% 0.98% 0.87% 0.93% 0.85% 0.89% 0.93% 0.98% 0.85% Primary air temperature 0.2 K 0.137 0.148 0.122 0.179 0.189 0.106 0.147 0.196 0.119 Primary air flow rate 1.5% 0.41% 0.34% 0.32% 0.39% 0.35% 0.27% 0.44% 0.40% 0.32%

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2) Uncertainty Requirement for the Experiment Cases

Table A.7 Uncertainty requirement for Manufacturer A's Low flow range unit

Standard 55F 58F 61F Individual Uncertainty (%) Requires 20cfm 30cfm 40cfm 20cfm 30cfm 40cfm 20cfm 30cfm 40cfm r1 3% 1.00% 1.00% 1.00% 1.00% 1.00% 1.00% 1.00% 1.00% 1.00% r2 2% 1.16% 0.99% 0.92% 1.31% 1.08% 0.99% 1.46% 1.23% 1.10% r3 2.5% 1.11% 1.24% 1.32% 1.23% 1.30% 1.44% 1.38% 1.51% 1.62% r4 3% 3.00% 3.00% 3.00% 3.00% 3.00% 3.00% 3.00% 3.00% 3.00% r5 3% 2.00% 2.00% 2.00% 2.00% 2.00% 2.00% 2.00% 2.00% 2.00% r 6% 4.07% 4.06% 4.07% 4.15% 4.11% 4.13% 4.25% 4.22% 4.22%

Table A.8 Uncertainty requirement for Manufacturer A's Middle flow range unit

Standard 55F 58F 61F Individual Uncertainty (%) Requires 40cfm 50cfm 60cfm 40cfm 50cfm 60cfm 40cfm 50cfm 60cfm r1 3% 1.00% 1.00% 1.00% 1.00% 1.00% 1.00% 1.00% 1.00% 1.00% r2 2% 1.07% 0.94% 0.86% 1.21% 0.99% 0.93% 1.35% 1.17% 1.12% r3 2.5% 1.28% 1.28% 1.38% 1.40% 1.39% 1.45% 1.57% 1.61% 1.70% r4 3% 3.00% 3.00% 3.00% 3.00% 3.00% 3.00% 3.00% 3.00% 3.00% r5 3% 2.00% 2.00% 2.00% 2.00% 2.00% 2.00% 2.00% 2.00% 2.00% r 6% 4.10% 4.07% 4.08% 4.17% 4.11% 4.12% 4.27% 4.24% 4.26%

Table A.9 Uncertainty requirement for Manufacturer A's High flow range unit

Standard 55F 58F 61F Individual Uncertainty (%) Requires 60cfm 70cfm 80cfm 60cfm 70cfm 80cfm 60cfm 70cfm 80cfm r1 3% 1.00% 1.00% 1.00% 1.00% 1.00% 1.00% 1.00% 1.00% 1.00% r2 2% 1.15% 1.10% 1.03% 1.34% 1.27% 1.21% 1.68% 1.57% 1.50% r3 2.5% 1.47% 1.55% 1.58% 1.66% 1.76% 1.84% 2.04% 2.12% 2.21% r4 3% 3.00% 3.00% 3.00% 3.00% 3.00% 3.00% 3.00% 3.00% 3.00% r5 3% 2.00% 2.00% 2.00% 2.00% 2.00% 2.00% 2.00% 2.00% 2.00% r 6% 4.18% 4.20% 4.19% 4.31% 4.33% 4.34% 4.58% 4.58% 4.59%

Table A.10 Uncertainty requirement for Manufacturer B's Low flow range unit

Standard 55F 58F 61F Individual Uncertainty (%) Requires 20cfm 30cfm 40cfm 20cfm 30cfm 40cfm 20cfm 30cfm 40cfm r1 3% 1.00% 1.00% 1.00% 1.00% 1.00% 1.00% 1.00% 1.00% 1.00% r2 2% 1.36% 1.10% 0.95% 1.42% 1.19% 1.06% 1.71% 1.53% 1.35% r3 2.5% 1.31% 1.51% 1.62% 1.34% 1.60% 1.77% 1.59% 1.98% 2.18% r4 3% 3.00% 3.00% 3.00% 3.00% 3.00% 3.00% 3.00% 3.00% 3.00% r5 3% 2.00% 2.00% 2.00% 2.00% 2.00% 2.00% 2.00% 2.00% 2.00% r 6% 4.19% 4.18% 4.19% 4.22% 4.24% 4.27% 4.41% 4.50% 4.53%

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Table A.11 Uncertainty requirement for Manufacturer B's Middle flow range unit

Standard 55F 58F 61F Individual Uncertainty (%) Requires 40cfm 50cfm 60cfm 40cfm 50cfm 60cfm 40cfm 50cfm 60cfm r1 3% 1.00% 1.00% 1.00% 1.00% 1.00% 1.00% 1.00% 1.00% 1.00% r2 2% 0.91% 0.81% 0.76% 1.01% 0.94% 0.89% 1.10% 1.09% 1.00% r3 2.5% 1.45% 1.52% 1.63% 1.59% 1.76% 1.89% 1.73% 1.94% 2.04% r4 3% 3.00% 3.00% 3.00% 3.00% 3.00% 3.00% 3.00% 3.00% 3.00% r5 3% 2.00% 2.00% 2.00% 2.00% 2.00% 2.00% 2.00% 2.00% 2.00% r 6% 4.11% 4.12% 4.15% 4.19% 4.24% 4.28% 4.27% 4.35% 4.38%

Table A.12 Uncertainty requirement for Manufacturer B's High flow range unit

Standard 55F 58F 61F Individual Uncertainty (%) Requires 60cfm 70cfm 80cfm 60cfm 70cfm 80cfm 60cfm 70cfm 80cfm r1 3% 1.00% 1.00% 1.00% 1.00% 1.00% 1.00% 1.00% 1.00% 1.00% r2 2% 0.88% 0.82% 0.80% 1.02% 0.96% 0.94% 1.23% 1.20% 1.18% r3 2.5% 1.74% 1.83% 1.93% 1.96% 2.07% 2.20% 2.30% 2.51% 2.69% r4 3% 3.00% 3.00% 3.00% 3.00% 3.00% 3.00% 3.00% 3.00% 3.00% r5 3% 2.00% 2.00% 2.00% 2.00% 2.00% 2.00% 2.00% 2.00% 2.00% r 6% 4.22% 4.25% 4.28% 4.35% 4.39% 4.44% 4.56% 4.66% 4.76%

3) Thermal Balance Method

1. Temperature Measurement of Induced Air after Leaving the Cooling Coil 1) Induction ratio shall be calculated by measuring the temperature of induced air after leaving the cooling coil under steady-state while the active chilled beam system is operating. 2) Energy balance between the chilled-beam coil and induced air allows for solving the induced air flow rate. The following equations explain heat transfer procedure while the induced air passes the chilled-beam coil under steady-state condition assuming no condensation on the chilled-beam coil surface.

푚̇ 푤 ∗ 퐶푝푤 ∗ (푇푤2 − 푇푤1) = 푚̇ 𝑖 ∗ 퐶푝푎 ∗ (푇푟 − 푇𝑖푛푑푢푐푒푑 푎𝑖푟 푎푓푡푒푟 푙푒푎푣𝑖푛푔 푡h푒 푐표𝑖푙)

푚̇ 푤 ∗ 퐶푝푤 ∗ (푇푤2 − 푇푤1) 푚̇ 𝑖 = 퐶푝푎 ∗ (푇푟 − 푇𝑖푛푑푢푐푒푑 푎𝑖푟 푎푓푡푒푟 푙푒푎푣𝑖푛푔 푡h푒 푐표𝑖푙)

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푚̇ 𝑖 = 휌𝑖 ∗ 푞𝑖

푚̇ 푤 ∗ 퐶푝푤 ∗ (푇푤2 − 푇푤1) 푞𝑖 = 휌𝑖 ∗ 퐶푝푎 ∗ (푇푟 − 푇𝑖푛푑푢푐푒푑 푎𝑖푟 푎푓푡푒푟 푙푒푎푣𝑖푛푔 푡h푒 푐표𝑖푙)

2. Measurement Requirements and Locations 1) Temperature measuring instruments shall meet the requirements of ASHRAE Standard 41.1 and ASHRAE Standard 41.6 and the following subsections. a. Accuracy of the temperature-measuring instruments shall be within the following limits: i. Air temperature ±0.2˚F (0.1˚C) ii. Water Temperature ±0.2˚F (0.1˚C) 2) The temperature of the water into and out of the test object shall be measured by sensors placed in the water flow immediately before and after the water passes through the test object. 3) Air temperatures shall be measured by radiant shielded sensors in order to mitigate radiant effects on the sensors. 4) Water flow rate shall be measured as per section 4.5. a. The water flow rate through the test object shall be measured with a flowmeter calibrated to an uncertainty of ±0.5% or less. 5) Primary air flow shall be measured as per section 4.4 a. Primary air flow measurement shall have an accuracy of ±5.0% of primary airflow reading. b. Primary air temperature shall be maintained to within ±0.4˚F (0.2˚C) during airflow measurement

6) Reference temperature (tr, induced air temperature before entering the cooling coil) shall be measured as per section 5.3.5.1.

a. Reference temperature (tr) is an average of air temperatures of the induced air on the inlet side of the cooling coils, measured with radiation-shielded sensors in three positions in each induced air opening, two centrally at the quarter points and one at the central point of the opening, at 1 in. (25mm) from the induced air opening. 7) The temperature measurements of induced air after leaving the cooling coil shall be taken at eight (8) points located 1.5 in. (38mm) from the outside edge of the coil-fin-face with sensors positioned on the coil-fin-face, 0.5 in (13mm) from above the coil-fin-face according to the following layout:

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Figure A.1 Temperature sensors positioned on the coil-fin-face

Figure A.2 Temperature sensors positioning layout

4. Testing Requirements 1) Testing shall be conducted after the system reaches steady-state as per section 5.6. The standard establishes steady-state conditions as those in the standard deviation of all of the measurement parameters is within the ranges of the following subsections.

a) Reference temperature, Tr (std. dev.), is less than 0.09˚F (0.05˚C). b) Inlet water temperature, Tw1 (std. dev.), is less than 0.09˚F (0.05˚C). c) Cooling water flow rate, qw (std. dev), is less than 1% of the specified value. d) Primary air temperature, Tp (std. dev), is within 0.4˚F (0.2˚C) of the specified value.

e) Primary air flow rate, qp (std. dev), is within 1.5% of the specified value.

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5. Test Procedures 1) Test measurements shall only be made when all of following conditions are satisfied. a) Reference temperature is between and 67˚F and 82˚F (19˚C and 27˚C). b) The primary air flow rate shall not vary by more than ±5% or 1 cfm (0.5 L/s), whichever is greater. c) The primary air temperature shall not vary by more than ±1˚F (0.5˚C). 2) Test measurements shall include at least 300 temperature recordings (1 per second, at least 5 minutes) at each measurement point.

6. Calculation and Expression of Results 1) The temperature of induced air after leaving the cooling coil shall be calculated as the arithmetic average of the temperature readings at all eight (8) measurement points. 2) The temperature of induced air after leaving the cooling coil shall be calculated according to the following equation:

푚̇ 푤 ∗ 퐶푝푤 ∗ (푇푤2 − 푇푤1) 푞𝑖 = 휌𝑖 ∗ 퐶푝푎 ∗ (푇푟 − 푇𝑖푛푑푢푐푒푑 푎𝑖푟 푎푓푡푒푟 푙푒푎푣𝑖푛푔 푡h푒 푐표𝑖푙)

3) The induction ratio shall be calculated by dividing the induced air flow rate by the primary air flow rate.

푞𝑖 퐾𝑖푛 = 푞푝

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4) Code script: the main algorithm of the new TRNSYS model

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5) Code script: the main algorithm of the new EnergyPlus model

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