The High Strength Friction Grip (HSFG) bolts

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Chapter 4Experimental Studies on Belleville Springs use in the Sliding Hinge Joint Connection

Introduction

Both SFC and AFC may be subjected to the post-sliding bolt tension loss (e.g. [6, 7, 10, 42, 60-63]), with the improvements observed in the experiments using Belleville springs (BeSs). Consequently the BeSs have been considered as a potentially beneficial but optional component of the friction sliders, with no analytical and experimental research found in the literature, prior to the current research, focused on investigating their effects and optimum configuration. The analytical discussion is presented in chapter 3 and this chapter deals with the experimental aspects of the subject regarding the sliding hinge joint (SHJ) with the Asymmetric Friction Connections (AFCs). A downside of retaining the post sliding AFC clamping force could have been deteriorating the self-centring capability of the system through exhibiting resistance to slide while the system tends to come back to its original position. This issue has been overcome through the use of partially deflected BeSs as proposed by Ramhormozian et al. [35], explained in chapter 3 of this thesis, and experimentally tested in the research which is presented in this chapter.
This chapter describes an experimental research program comprising nine real-scale tests at quasi-static and dynamic rates of displacement controlled loading on the SHJ’s beam bottom flange AFC, to investigate the influence of using BeSs on the SHJ’s AFC sliding behaviour. The HSFG bolts were installed in the elastic range at ≈50% of the bolt proof load. Different AFC configurations were considered including having no BeSs and having BeSs at one side as well as both sides of the AFC with varying BeS system stiffness values.
This chapter provides the answers to the following questions through the results of the experimental research undertaken:
1- What is the influence of different configurations of BeSs in the AFC bolt assemblage on maintaining the post-sliding SHJ’s AFC clamping force?
2- What is the influence of different configurations of partially deflected BeSs in the AFC bolt assemblage on the self-centering capability of the SHJ’s AFC?
3- What is the influence of different configurations of BeSs in the AFC bolt assemblage on the SHJ’s AFC system coefficient of friction?
4- What is the influence of different configurations of partially deflected BeSs in the SHJ’s AFC bolt assemblage on the clamping force variations during sliding, mainly, due to moment, shear, and axial force (MVP) interaction and/or prying actions?
5- What is the influence of different configurations of BeSs in the AFC bolt assemblage on the AFC sliding surfaces wearing?

SHJ’s beam bottom flange AFC real-scale component tests description

Test rig and AFC plies

A test rig (Figure 4.1) was designed and used to simulate the sliding behaviour of the SHJ beam bottom flange AFC assemblage. It consists of a column hinged to a strong wall base plate at A, through an effectively friction less bearing, connected to a ±300 kN Shore Western (SW) dynamic actuator at B, and having four threaded holes and a circular shear key to bolt and hold the cleat at C. A beam flange plate was welded to the strong floor base plate to allow assembling the AFC component for each test. This test setup is an inverted representation of the SHJ, with the point of rotation at A, and the SHJ beam bottom flange AFC at C. The SW dynamic actuator applies the quasi-static as well as intermediate and main dynamic displacement controlled loads at B. The drawings of the test rig may be found in [6, 10].
The vertical distance between the hinge centre and the cleat top surface, where the AFC plies relative displacements are measured, is 478 mm, a figure which closely represents a beam section depth of a 460 UB section. The reaction arm amplifies the load from that applied by the SW dynamic actuator at B by a factor of 2.24, if the system is assumed as static with a perfectly friction-less bearing at A, generating the required AFC sliding force transferred to the cleat through the shear key at C. Assuming the system as dynamic, the test rig column can be considered as a generalized single degree of freedom (SDOF) dynamic system [91]. However, since the imposing loads are orders of magnitude way higher than the mass generated inertia forces, as a result of the absence of the dead (permanent) and live (imposed) contributing building mass, it is acceptable here to use the factor of the static system also for the dynamic system. The distance from the hinge centre and the centre of the actuator connection plate to the column is 1071mm. Hence, the load amplifying factor, or the displacement decreasing factor assuming a rigid body rotation (RBR) for the column, is (1071/478) = 2.24.
Each test specimen comprised a beam flange plate, a cleat, and a cap plate on top, all of which are schematically shown in Figure 4.1. The beam flange plate was 16 mm thick with 22 mm diameter plasma cut circular holes and was welded to the strong wall. The narrow end of the cleat was 243 mm × 150 mm × 16 mm with 57 mm long elongated plasma cut holes, with the wider end bolted to the top of the column and held by the circular shear key. The cap plate was 175 mm × 150 mm × 16 mm with 22 mm diameter plasma cut circular holes. These plates were made from Grade 350 steel (typical fy=350-440MPa and typical hardness=140-180BHN). The 5mm thick high hardness (wear or abrasion resistant) steel shims were made from Raex 450 grade plate (typical fy=1200MPa and hardness range=420-500HBW). The shims were 190 mm × 175 mm including 22 mm diameter plasma cut circular holes and were placed at both sides of the cleat. The plies surfaces were sweep blasted to St2 surface finish standard. A new cleat was used for each experiment. The shims used for each experiment were either new or reused from the previous experiments, with the unscratched main sliding surface from the previous experiment facing the cleat in the new experiment.

Bolts and Belleville springs

The M20 galvanized HSFG property class (PC) 8.8 steel structural bolts were supplied by a specialist New Zealand fasteners supplier and used for the experiments. New bolts were used for each experiment. Table 4.1 shows the nominal characteristics of the bolts.
The BeSs were supplied by a specialist American Belleville Spring manufacturing company. Table 4.2 shows the characteristics of the BeS “M20-52-6.0NF/S1” tested before being used in the experiments. The BeSs were all pre-set, meaning that they were loaded to a flat disc once after the production, hence behaved fully elastically in the experiments. Pre-setting is essential for BeS’s used in this system. Consequently, the BeSs were re used for different experiments.
The load-deflection values of the M20-52-6.0NF/S1 BeS are shown in Table 4.3 and the associated loading as well as unloading curves are demonstrated in Figure 4.2 along with a fitted linear line originated at (0,0) to represent both loading and unloading data points the best indicating the nominal stiffness of 135kN/mm for each BeS. A detailed discussion on the BeS’ characteristics may be found in chapter 3 and [35].

AFC Bolts and Belleville springs assemblage configurations

50% of the HSFG property class 8.8 M20 bolt proof load i.e. 0.5×147=73.5kN was considered as the target installed bolt tension for all of the experiments. This is to keep the bolts in the elastic range at installation, increasing their capacity to accommodate additional elastic stresses during stable sliding as described in chapter 3. Five different configurations were considered for the use of BeSs namely: 1) NS-Ti, 2) S1-Ti, 3) S2-Ti, S3-Ti, and 5) S4-Ti. These designations mean; 1) having no BeSs but ordinary HSFG hardened washers, and 2) to 5) having one, two, three, and four BeSs in series respectively. A detailed discussion on the BeS’ possible series, parallel, and series/parallel assemblage configurations may be found in chapter 3 and [35]. Ti=1, 2, and 3 represent the number of each specific configuration’s test repeat. The configurations S2 and S4 used one Belleville spring at nut side between a hardened washer under the nut and the cap plate while the configurations NS, S1, and S3 used only the HSFG hardened washer under the nut. Configurations S1 and S2, and S3 and S4 used 1 and 3 BeSs at head side of each bolt respectively. Considering the joint grip length, the values for the bolt lengths that were used for the experiments were 110mm, 120mm, 120mm, 130mm, and 140mm for NS-Ti, S1-Ti, S2-Ti, S3-Ti, and S4-Ti respectively, all excluding the bolt head thickness. This ensured that the full height of the nut was engaged with the bolt threaded part by showing at least one clear thread above the nut after tightening, as is recommended by [15]. Figure 4.3 shows different configurations of the bolts and BeSs assemblage.

SHJ’s AFCs assemblage and test measurements

The AFC bolts were tightened in sequence from the first one to the forth one, first, up to snug tight according to [15], then from the first one to the fourth one up to the desired level of tension using a V-RAD-16 electric torque multiplier. The bolts were numbered as is shown in Figure 4.3. The tension of each bolt was monitored and recorded continuously over the whole test time by a Transducer Techniques TT-LWO-60 load cell under each bolt (Figure 4.3). Each load cell was sandwiched between two M20 HSFG hardened washers, and each load cell’s capacity was 267kN.
The relative displacements between the AFC plies were recorded by five portal displacement gauges (Figure 4.3). The displacement and axial force of the actuator were recorded by the actuator internal displacement gauge and load cell. The actuator was connected to the test rig column by four pre tensioned structural bolts. The bracket supporting the actuator load and the strong wall base plate were installed on the strong floor and strong wall respectively, each one by seven post tensioned rods to minimize support movement. A number of trials were performed at first to determine the calibration factor of the test rig due to the sources of potential flexibility and potential slips. The aim was to evaluate the maximum necessary displacement of the actuator resulting in the maximum desired displacement of the AFC cleat. These trials were performed using the AFC bolts tightened up to 50% of the proof load as was the case for the main tests. The calibration factor was determined as 1.5. After these trials, the bolts of the test rig connections, especially the actuator connections, were firmly retightened to minimize the system slip during the main tests.
The length of each bolt was measured by an ultrasonic G5 bolt tension meter, before tightening, after tightening, after each test, and after untightening (Figure 4.3). The bolt head and the nut were marked so as to investigate if there has been any nut relative to bolt head rotation during the sliding tests. The temperatures of the bolts were measured before the test, right after the main dynamic loading, and after the test using a laser point infrared digital thermometer gun (Figure 4.3h). The height of the BeSs “at both sides if applicable” was measured by a depth micrometre before and after the test (Figure 4.3e).

Loading regime

The loading regime comprised a combination of displacement controlled load histories, namely a main dynamic loading pattern, four quasi-static, and four intermediate dynamic regimes. The loading regime was designed to simulate the pre-earthquake, severe earthquake, and post-earthquake conditions to represent a combination of SLS, ULS, and greater than ULS events. The loading regime displacement versus time is shown in Figure 4.4.
The quasi-static loading provides an AFC cleat-beam flange plate relative displacement of approximately 4.8 mm, corresponding to a SHJ beam-column relative rotation of 0.010 rad which was applied at a slow slip rate of 1 mm/minute. The displacement controlled quasi-static load was in both positive and negative directions, as is shown in Figure 4.4.
Following each quasi-static loading, the intermediate dynamic load was applied to ensure that the AFC cleat-beam flange plate relative displacement was gradually reduced and the AFC bolts ended up close to the straight vertical orientation. The number of cycles consisted of five cycles each to 3.6 mm and 2.4 mm, followed by nine cycles of decreasing amplitudes, as is shown in Figure 4.4.
The main dynamic load was applied after an intermediate dynamic loading, in the middle of the whole loading regime, as is shown in Figure 4.4, consisting of 3 AFC cleat-beam flange plate relative displacement cycles to 2.4 mm and 3.6 mm, 2 cycles to 4.8 mm, 1 cycle to 7.2 mm and 9.6 mm, 2 cycles to 14.3 mm, and back down again. This corresponds to SHJ beam-column relative rotations of 0.005 rad, 0.0075 rad, 0.010 rad, 0.015 rad, 0.020 rad and 0.030 rad. It is worth noting that the higher amplitudes mentioned above push the SHJ’s AFC beyond the limits associated with the 2.5% ULS inter-storey drift which is specified by [92].
For the amplitudes less than or equal to 3.6mm, the period of the loading was T=1sec/cycle, and for the amplitudes greater than 3.6mm, the period was T=1.5sec/cycle. This is because, the higher SHJ’s AFC sliding displacements, or the damage in a non-low damage MRSF building, are associated with the higher periods. The loading regime was modified in this research from that initially adapted by Khoo [6] from the recommended testing procedure for MRSFs by the SAC-Joint-Venture [93].
The displacement controlled (and not load controlled) loading regime is basically a modified version of the loading regime presented in recommended testing procedure for steel MRFs by the SAC Joint Venture [93] as is already mentioned. The original proposed loading regime comprises a series of stepwise increases in drift. The modifications made are as follows:
The regime was applied in terms of beam-column relative rotations and not storey drifts. Drifts are a combination of the deformations of column, beam, panel zone, and joint, which the test rig is unable to represent all of them.
The rotational amplitudes were reduced by 0.005 rad from the specified drifts to take the above mentioned deformations into account. This was based on the assumption that drifts under 0.5% are non-damaging [93]. This is elastic deformation in the overall conventional joint assembly, as opposed to inelastic beam-column relative rotation through slip in the AFC.
The loading regime was modified from the original stepwise increase to maximum drift, to a stepwise increase to maximum rotation in half the number of cycles and back to minimum with the other half, then reaching to zero rotation by a further nine cycles of linearly decreasing amplitudes. Given the aim here is to have no damage in the system, this better reflects an actual earthquake induced shaking where the peak is reached in the first half to middle of the shaking duration rather than at the end. This also ensures the condition of the specimen at the end of each dynamic load history part (i.e. intermediate and main) to be representative of post-earthquake conditions. If the joint came to rest immediately after the maximum amplitude, the bolts would be in the stable sliding state with different residual sliding strengths in each direction. If the AFC is then loaded in the same direction as sliding was taking place prior to stopping, the connection would develop the stable sliding resistance before sliding re-commences. If the specimen was loaded in the opposite direction, it would behave as it does under load reversal, and thus slide at very low levels of friction. Both of these cases do not necessarily represent the most probable conditions of the AFC under earthquake shaking. The stepwise decreases in cycles were applied to ensure the bolts were close to vertical when the AFC came to a complete stop.
The quasi-static loading ramp was introduced to investigate the point at which the sliding commences both before and after a severe earthquake.
The intermediate dynamic load was applied following each quasi-static loading to gradually reduce the beam-column relative rotation to zero so that the AFC bolts ended up close to the straight vertical orientation.
The frequencies of the dynamic load history parts are based on the numerical time history analysis of a typical building incorporating the SHJs [10].

Results and discussions

Five tests were first carried out for each configuration followed by two more repeats on each one of the NS and S4 configurations that showed the less-than and the most desirable seismic behaviour respectively, particularly in terms of post sliding clamping force retention. Hence a total of nine tests were carried out in the research presented in this chapter. The details of the test specimens are shown in Table 4.4. The values on the table are calculated as follows in which D is calculated based on linear interpolation using load-deflection data points presented in Table 4.3, E is the BeS system nominal reserved deflection up to a flat BeS system considering 0.84mm as the flat deflection of each BeS, and j is the number of BeSs in each test bolt assemblage configuration:

SHJ’s AFC elastic strength reduction

Table 4.5 shows the percentage of post-sliding AFC clamping force loss at the end of each test with respect to the installed clamping force as well as the AFC post sliding force threshold of losing stiffness normalized to the installed clamping force, for different configurations. These are named the normalized clamping force loss (CFL) and the normalized elastic strength limit (ESL) respectively. The force threshold of stiffness loss is defined as the point on the last quasi static loading line at which there is a significant reduction in the system stiffness, suggesting the sliding has occurred. This was obtained by finding out the point on the data points at which the rate of the load increase just becomes noticeably and considerably smaller suggesting having sliding. In some cases this was a sudden drop in the applied load. The post sliding=elastic) strength reduction factor (SRF) is defined as the ratio of CFL and ESL ( shown in Table 4.5. The smaller SRF means more retention of the post sliding elastic strength.

Table of content
Abstract
Acknowledgment 
Dedication 
Table of contents 
List of figures 
List of tables 
1. Chapter 1 Introduction
1.1 Moment Resisting Frames (MRFs)
1.2 Improvements in steel MRFs’ seismic performance
1.3 Objectives and scope of the PhD research and the thesis outline
2 Chapter 2 The High Strength Friction Grip (HSFG) bolts: Proposed Changes on Australia and New Zealand Standards and Bolting Practice 
2.1 Introduction
2.2 Importance of the bolt tension in friction-type connections
2.3 Existing NZS3404 recommended part turn method of tightening
2.4 Bolt Fracture and variable installed tension observed in the bolt tightening research.
2.5 Research to investigate the existing NZS3404 recommended part turn method of tightening
2.6 Calculating the required nut turn, based on a bolt length definition to reach the bolt proof load
2.7 Recommendations for Australia and New Zealand bolting practice and standards
3 Chapter 3 Stiffness-Based Approach for Belleville Springs use in Friction Sliding Structural Connections
3.1 Introduction
3.2 Asymmetric and symmetric friction connection sliding behaviour and bolt tension loss reasons
3.3 Bolt longitudinal stiffness:
3.4 Joint stiffness:
3.5 Belleville spring(s) stiffness: .
3.6 AFC or SFC bolt installation and post-sliding tension loss with and without BeSs:
3.7 Using not flattened Belleville springs to minimize the prying effects and to improve the AFC self-centring capability:
3.8 Other potential benefits of using Belleville springs in the AFC and/or SFC:
3.9 Proposed design procedure for using BeSs in the AFC and SFC:
3.10 Conclusions to chapter 3
4 Chapter 4 Experimental Studies on Belleville Springs use in the Sliding Hinge Joint Connection 
4.1 Introduction
4.2 SHJ’s beam bottom flange AFC real-scale component tests description
4.3 Results and discussions
4.4 Conclusions to chapter 4
5 Chapter 5 Experimental Studies on the Asymmetric Friction Connection (AFC) Optimum Installed Bolt Tension and Influence of the Customized Belleville Spring in Retaining the Bolt Tension 
5.1 Brief background to this research
5.2 The AFC experiments on the 500kN MTS machine
5.3 Results & Discussion
5.4 Conclusions to chapter 5 .
6 Chapter 6 Establishing the Method of Bolt Tightening with Belleville Springs in the Elastic Range of the Bolt Preload 
6.1 Introduction
6.2 Test components
6.3 The tightening tests
6.4 The test procedure
6.5 The measurements and calculations
6.6 The research results .
6.7 Conclusions to chapter 6 and recommendations
7 Chapter 7 Influence of the Surface Roughness on the AFC Sliding Behaviour
7.1 Introduction
7.2 Methodology
7.3 Results
7.4 Discussion
7.5 Conclusions to chapter 7 and recommendations
8 Chapter 8 Experimental Studies on three new potential Asymmetric Friction Connection (AFC) Ply Configurations, Surface Condition, and Material
8.1 Introduction
8.2 The test setup, loading regime, and measurements
8.3 The AFC with abrasion resistant cleat and no shims
8.4 The AFC with Titanium Nitride (TiN) coated shims
8.5 The AFC with abrasion resistant shims and cleat
8.6 Conclusions to chapter 8
9 Chapter 9 Self-Centering Capability of the Seismic Friction Dampers: A Conceptual Study on the Static and Dynamic Self-Centering Requirements for the Single Degree of Freedom (SDOF) Asymmetric and Symmetric Friction Connections (AFC and SFC) 
9.1 Introduction
9.2 Statically self-centering SFC and AFC
9.3 Available potential types of pre-compressed springs to be used
9.4 Potential use of Lurethane spring block in the SHJ
9.5 Numerical study on the dynamic self-centering capability of the AFC and SFC SDOF systems
9.6 Results
9.7 Discussion
9.8 Conclusions to chapter 9
10 Chapter 10 Ongoing Research
10.1 Finite Element Analysis (FEA) of the Asymmetric Friction Connection (AFC) in the Sliding Hinge Joint (SHJ)
10.2 Dynamic Performance Analysis and System Identification (SI) of a Low Damage Multi-Storey Structural Steel Building under two Moderately Severe Earthquake Events using Structural Health Monitoring (SHM) Data
10.3 Simplified Multi Degree of Freedom (MDOF) Model of the Buildings Incorporating the Seismic Friction Dampers
10.4 Developing a modified AFC bolt model
10.5 Developing an analytical model of the SHJ connection
10.6 Industry/practice guide on the design and installation of the Sliding Hinge Joint (SHJ) and/or Asymmetric Friction Connection (AFC) with Belleville springs (BeSs) along with the examples (to be used for the new and existing buildings)
11 References .
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Enhancement of the Sliding Hinge Joint Connection with Belleville Springs

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